ML11236A226

Partially Released Updated Final Safety Analysis Report for Indian Point Unit 2. (Redacted Information: Figures 5.1-2, 5.1-3, 5.1-4, 5.1-5, 5.1-6 & 5.1-7 Are All Plant Drawings). Table of Contents - Chapter 7, Page 18 of 108

ML11236A226
Person / Time
Site:	Indian Point
Issue date:	10/06/2010
From:	Brinkman D Office of Nuclear Reactor Regulation
To:
References
FOIA/PA-2011-0256, FOIA/PA2011-0256
Download: ML11236A226 (900)

Text

{{#Wiki_filter:June 2011 This version of the Indian Point Unit 2 Updated Final Safety Analysis Report (UFSAR) is the licensee's version 22, submitted to the NRC on October 6, 2010, with information current through April 2010, with certain redactions of sensitive information by staff of the Nuclear Regulatory Commission (NRC) to allow release to the public. The redactions are made under 10CFR2.390(d)(1). The material included within is classified as publicly available information. As of June 2011, this is the latest UFSAR revision submitted to the NRC. The redactions were made due to meeting the NRC's criteria on sensitive information, as specified in SECY-04-0191, "Withholding Sensitive Unclassified Information Concerning Nuclear Power Reactors from Public Disclosure", dated October 19, 2004, ADAMS ML042310663, as modified by the NRC Commissioners Staff Requirements Memorandum on SECY-04-0191, dated November 9, 2004, ADAMS ML043140175. The following information was redacted by NRC staff: Figures 5.1-2, 5.1-3, 5.1-4, 5.1-5, 5.1-6, 5.1-7, which were supplied as plant drawings 9321-2501, 2502, 2503, 2506, 2507, 2508. Any other material that is listed as "deleted" was deleted by the licensee as part of their continuous update process for the UFSAR.

Instructions and Key to Indian Point 2 UFSAR for Revision 22 Current Changes - Revision 22: The 2010 revision to the UFSAR is Revision 22. The current changes made in this update that have been sent to the NRC are highlighted in a gray shaded background. Historical Information: This UFSAR has marked some parts of the UFSAR as Historical Information following the guidance in NEI 98-03, Rev. 1, Guidelines for Updating Final Safety Analysis Reports. Information that is highlighted with a green background is designated as Historical Information and is not updated. The definition of Historical Information means that the information meets the following criteria:

• Information relating to initial plant licensing and start-up that was included in the original FSAR to meet the requirements of 10CFR50.34(b).
• Information that was accurate at the time the plant was originally built, but is not intended or expected to be updated for the life of the plant (unless required by the Commission).
• Information that is not affected by changes to the plant or its operation.
• Information that does not change with time.

Deleted Information and Figures: Deleted information will have the word Deleted in yellow highlighting. NRC Orders and License Conditions: Information relocated to the UFSAR from the Technical Specifications or included by NRC Order, become a Licensing Condition and cannot be removed from the UFSAR without NRC approval. Where these are identified, there will be a [Note:] with bold text and purple highlighting. Fission Product Barrier Design Basis Limits for IP2: Information contained in the UFSAR that represent fission product barrier design basis limits for the plant are highlighted with an aqua shaded background, or aqua lettering if the change is contained in a gray shaded background. These are fundamental to the fission product barrier and cannot be exceeded, changed or altered without NRC approval. UFSAR Figures and Plant Drawings: In an effort to improve the accuracy of the figures in the UFSAR and to reduce redundant work, Revision 19 replaced UFSAR Figures, where possible, with references to the current plant drawings. Each current revision has placed snapshots of the referenced plant drawings for use by the NRC and for user convenience. Searching the UFSAR: The UFSAR has been reformatted in Adobe Acrobat. Use the Adobe search feature to find the information you are looking for.

INDIAN POINT ENERGY CENTER (IPEC) UNIT 2 UPDATED FINAL SAFETY ANALYSIS REPORT (UFSAR) Table of Contents Chapter 1 - Introduction and Summary Chapter 2 - Site and Environment Chapter 3 - Reactor Chapter 4 - Reactor Coolant System Chapter 5 - Containment Spray Systems Chapter 6 - Engineered Safety Features Chapter 7 - Instrumentation and Control Chapter 8 - Electrical Systems Chapter 9 - Auxiliary and Energy Systems Chapter 10 - Steam and Power Conversion System Chapter 11 - Waste and Radiation Protection System Chapter 12 - Conduct of Operations Chapter 13 - Test and Operation Chapter 14 - Safety Analysis

IP2 FSAR UPDATE CHAPTER 1 INTRODUCTION AND

1.1 INTRODUCTION

[Historical Information] The Final Safety Analysis Report (FSAR) was submitted in support of an application by Consolidated Edison Company of New York, Inc. (Con Edison) for a license to operate the Indian Point Unit 2 nuclear power plant (AEC Docket 50-247, Permit No. CPPR-21). It provided pertinent technical information in accordance with Section 50.34 of 10 CFR 50 requirements to obtain a nuclear power plant operating license. Westinghouse Electric Corporation was the primary contractor and had turnkey responsibility for the design, construction, testing, and initial startup of the facility. Westinghouse had contracted with United Engineers and Constructors as architect-engineer to provide engineering assistance in the design of and construction of the structural and civil works. This document is the Updated FSAR and is submitted in accordance with Section 50.71(e) of 10 CFR Part 50. The revision history is summarized in Section 1.9.2. The following paragraphs contain a summary of the reports scope: The unit employs a pressurized water reactor nuclear steam supply system (NSSS) furnished by Westinghouse Electric Corporation. The reactor was originally licensed at a maximum thermal power of 2758 MW. By letter dated September 30, 1988 Con Edison initiated a request to authorize an increase in the licensed thermal power level to 3071.4 MW. The NRC documented their acceptance of this request in a Safety Evaluation Report dated March 7, 1990. By letter dated December 12, 2002, Entergy Nuclear Operations, Inc. initiated a request for a Measurement Uncertainty Recapture power uprate of 1.4%, for a licensed thermal power level of 3114.4 MW. The NRC documented their acceptance of this request in a Safety Evaluation Report dated May 22, 2003. By letter dated January 29, 2004, Entergy Nuclear Operations, Inc. initiated a request for a Stretch Power Uprate, for licensed Thermal power level of 3216MW. The NRC documented their acceptance if this request in a Safety Evaluation Report dated October 27, 2004 and issued License Amendment #241. The reactor is presently licensed to operate until September 28, 2013. The reactor power corresponds to an electric output from the turbine-generator of approximately 1078 MW. The plant heat removal systems have been designed for the equivalent guarantee rating of 3071.4 MWt; some of the portions of the safety analysis dependent on heat removal capacity of plant and safeguards systems have assumed the maximum calculated power of 3216 MWt as have the evaluations of activity release and radiation exposure. By NRC order dated August 27, 2001 (Reference 1), Con Edisons ownership/operation of Indian Point 1 and 2 was transferred to Entergy Nuclear Indian Point 2 (ENIP2), LLC, as the owner of Indian Point 1 and 2 plants, and Entergy Nuclear Operations (ENO), Inc. as the operator of Indian Point 2 and maintainer of Indian Point 1. Consequently, references to Con Edison (or derivatives thereof) in this document remain only when used in historical context. Chapter 1, Page 1 of 71 Revision 22, 2010

IP2 FSAR UPDATE The remainder of Chapter 1 of this report summarizes the principal design features and parameters of the plant, pointing out the similarities and differences with respect to other pressurized water nuclear power plants presently in operation. A general description of the plant is included as well as a statement and summary of all the General Design Criteria. Chapter 2 contains a description and evaluation of the site and environs, supporting the suitability of the site for a reactor of the size and type described. Chapters 3 and 4 describe and evaluate the reactor and the reactor coolant system; Chapter 5, the containment system; Chapter 6, the engineered safety features; Chapter 7, plant instrumentation and control; Chapter 8, the electrical system; Chapter 9, the auxiliary and emergency system; Chapter 10, the steam and power conversion system; Chapter 11, radioactive waste disposal and radiation protection. Chapter 12 and 13 are conduct of operations and initial test and operations, respectively. They describe plant organization, training programs, and startup administrative procedure. Chapter 14 is a safety evaluation summarizing the analyses, which demonstrate the adequacy of the reactor protection system and the containment and engineered safety features, and which show that the consequences of various postulated accidents are within applicable limits. REFERENCES FOR SECTION 1.1

1. NRC letter to Consolidated Edison, Indian Point Nuclear Generating Unit No.s 1 and 2 - Order Approving Transfer of Licenses from the Consolidated Edison Company of New York, Inc., to Entergy Nuclear Indian Point 2, LLC, and Entergy Nuclear Operations, Inc. and Approving Conforming Amendments (TAC Nos.

MB0743 and MB0744), August 27, 2001. 1.2

SUMMARY

PLANT DESCRIPTION 1.2.1 Site Indian Point Unit 2 is adjacent to and north of Unit 1 on a site of approximately 239 acres of land on the east bank of the Hudson River at Indian Point, Village of Buchanan in upper Westchester County, New York. Indian Point Unit 3 (owned and operated by Entergy Nuclear and Entergy Nuclear Operations, Inc.) is adjacent to and south of Unit 1. The site is about 24 miles north of the New York City boundary line. The nearest city is Peekskill, 2.5 miles northeast of Indian Point. An aerial photograph, [Historical] Figure 2.2-1, shows the site and about 58 miles2 of the surrounding area. 1.2.1.1 Meteorology Meteorological conditions in the area of the site were determined during a 2-year test program. These data were used in evaluating the effects of gaseous discharges from the plant during normal operations and during the postulated loss-of-coolant accident. In addition, data supplied by the U.S. Weather Bureau at the Bear Mountain Station, regarding the meteorological conditions during periods of precipitation, were used to evaluate the rainout of fission gases into surface water reservoirs following the postulated loss-of-coolant accident. The evaluations indicate that the site meteorology provides adequate diffusion and dilution of any released gases. Chapter 1, Page 2 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.2.1.2 Geology and Hydrology Geologically, the site consists of a hard limestone in a jointed condition, which provides a solid bed for the plant foundation. The bedrock is sufficiently sound to support any loads, which could be anticipated up to 50 tons per ft2, which is far in excess of any load, which may be imposed by the plant. Although it is hard, the jointed limestone formation is permeable to water. Thus, if water from the plant should enter the ground (an improbable event since the plant is designed to preclude any leakage into the ground) it would percolate to the river rather than enter any ground water supply. Additional studies by geology consultant, Thomas W. Fluhr, and examination of soil borings confirmed the above conclusions. In the Hudson River, about 80,000,000 gallons of water flow past the plant each minute during the average tidal flow. This flow provides additional mixing and dilution for liquid discharges from the facility. In fact, however, this aspect is superfluous since the assumption in the plant design is to treat the river water as if it were used for drinking and thus to reduce radioactive discharges, by dilution with ordinary plant effluent, to concentrations that would be tolerable for drinking water. There is minimal danger of flooding at the site as discussed in Section 2.5. 1.2.1.3 Seismology Seismic activity in the Indian Point area is limited to low-level microseismicity. Detailed field investigations (e.g., Ratcliffe, 1976, 1980; Dames and Moore, 1977) have been conducted in the immediate vicinity of Indian Point and along the major faults in the region. To date, no evidence has been found in the rocks exposed at the surface or sediments overlying fault traces or in cores obtained in the vicinity of Indian Point, that might support a conclusion that displacement has occurred along major fault systems within the New York Highlands, the Ramapo or its associated branches during Quaternary time (the last 1.5 million years). In the vicinity of Indian Point, evidence that no displacement has occurred in the last 65 million years (since the Mesozoic) along specific major structures has been observed. The plant is designed to withstand an earthquake of Modified Mercalli Intensity VII. The validity of the selection of an Intensity VII earthquake was adjudicated before the Atomic Safety and Licensing Appeal Board. The Appeal Board's decision (ALAB-436) verified Intensity VII as the design basis earthquake for the plant. 1.2.1.4 Environmental Radiation Monitoring Environmental radioactivity has been measured at the site and surrounding area for nearly 20 years in association with the operation of the three Indian Point Units. These measurements will be continued and reported. The radiation measurements of fallout, water samples, vegetation, marine life, etc., have shown no significant postoperative increase in activity. Noticeable increases in fallout have coincided with weapons-testing programs and appear to be related almost entirely to those programs. The New York State Department of Health in an independent 2-year postoperative study found that environmental radioactivity in the vicinity of the site is no higher than anywhere else in the State of New York. 1.2.1.5 Conclusions Consideration of all the items mentioned above, plus the containment design and the engineered safety features included in the plant design lead to the conclusion of appropriate suitability of the site for safe operation of the Indian Point Unit 2 nuclear power plant. Accident Chapter 1, Page 3 of 71 Revision 22, 2010

IP2 FSAR UPDATE analyses presented in Chapter 14 verify that the maximum expected doses at or beyond the site boundary are within applicable limits. 1.2.2 Plant Description The unit incorporates a closed-cycle pressurized water nuclear steam supply system, a turbine generator and their necessary auxiliaries. A radioactive waste disposal system, fuel handling system and all auxiliaries, structures, and other onsite facilities required for a complete and operable nuclear power plant are provided for the unit. The general arrangement of the plant is shown on historical Figures 1.2-1 and 1.2-4, and Figure 2.2-2. Other general plant arrangement drawings have been removed due to security reasons following September 11, 2001 and can be viewed as plant drawings 9321-2510, 9321-2511, 9321-2514, 9321-2517, 9321-3052, and 209812. 1.2.2.1 Nuclear Steam Supply System (NSSS) The nuclear steam supply system consists of a pressurized water reactor, reactor coolant system, and associated auxiliary fluid systems. The reactor coolant system is arranged as four closed reactor coolant loops, each containing a reactor coolant pump and a steam generator, connected in parallel to the reactor vessel. An electrically-heated pressurizer is connected to one of the loops. The reactor core is composed of uranium-dioxide pellets enclosed in zircaloy tubes with welded end plugs. The tubes are supported in assemblies by a spring clip grid structure. The mechanical control rods consist of clusters of stainless steel clad absorber rods and guide tubes located within the fuel assembly. The core was initially loaded in three regions of different enrichments with new fuel being introduced into the outer region at successive refuelings and discharged to spent fuel storage, following burnup. The steam generators are vertical U-tube units employing Inconel tubes. Integral separating equipment reduces the moisture content of the steam leaving the steam generators to 0.25-percent or less. The reactor coolant pumps are vertical, single stage, centrifugal pumps equipped with controlled leakage shaft seals. Auxiliary systems are provided to perform the following functions:

1. Charge the reactor coolant system.

2. Add makeup water.

3. Purify reactor coolant water.

4. Provide chemicals for corrosion inhibition and reactor control.

5. Cool system components.

6. Remove residual heat when the reactor is shut down.

7. Cool the spent fuel storage pool.

8. Sample reactor coolant water.

9. Provide for emergency core cooling.

10. Collect reactor coolant drains.

11. Provide containment spray.

12. Provide containment ventilation and cooling.

13. Dispose of liquid, gaseous and solid wastes.

Chapter 1, Page 4 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.2.2.2 Reactor and Plant Control The reactor is controlled by a coordinated combination of chemical shim and mechanical control rods. The control system allows the unit to accept step load changes of +/- l0-percent and ramp load changes of +/- 5-percent per min over the load range of 15-percent to, but not exceeding, 100-percent power under nominal operating conditions subject to xenon limitations. 1.2.2.3 Turbine and Auxiliaries The turbine is a tandem-compound, comprising one high pressure and three low pressure cylinders, 1800 rpm unit having 45-in. exhaust blading in the low pressure cylinders. There are four moisture preseparators located at the four high-pressure turbine exhaust lines and six combination moisture separator-reheater units that are employed to dry and superheat the steam between the high and low pressure turbine cylinders. The turbine generator is capable of a 50-percent loss of external electrical load without turbine or reactor trip. The turbine auxiliaries include deaerating surface condensers, steam jet air ejector, turbine-driven main feedwater pumps, motor-driven condensate pumps, and six stages of feedwater heating. The original turbine generator had a guaranteed capability of 1,021,793 kWe at 1.5-in. Hg absolute exhaust pressure with zero percent makeup and six stages of feedwater heating. 1.2.2.4 Electrical System The main generator feeds electrical power through an isolated phase bus to two half-sized main power transformers. Station auxiliaries receive power during normal operation from either the station auxiliary transformer (i.e., offsite power) or the unit auxiliary transformer (i.e., unit main power transformers). The auxiliary electrical system provides power to those auxiliary components required to operate during normal or emergency conditions of operation. Standby power required during plant startup, shutdown, and after reactor trip is supplied to the station auxiliary transformer from the Con Edison 138-kV system by either of two separate overhead lines from the Buchanan substation approximately 0.50 mile from the plant. Alternate feeds from the 13.8-kV system are also available for immediate manual connection to the auxiliary buses. [Deleted] Emergency power supply for vital instruments and controls is from four 125-V station batteries. The system design provides sufficient independence, isolation capability, and redundancy between the different power sources to avoid complete loss of auxiliary power. 1.2.2.5 Control Room The plant is provided with a reactor and turbine-generator control room containing all necessary instrumentation for the operation of the plant under normal and accident conditions. Adequate shielding and air conditioning facilities permit occupancy during all operating or accident conditions. Chapter 1, Page 5 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.2.2.6 Diesel Generators Three diesel-generator sets supply emergency power for shutdown or essential safeguards operation in the event of a loss of all other alternating current auxiliary power. 1.2.2.7 Waste Disposal System The waste disposal system collects and processes liquids, gaseous, and solid waste from plant operation for removal from plant site. All removals are made in accordance with government guidelines for the process. 1.2.2.8 Fuel Handling System The fuel handling system provides the ability to fuel and refuel the reactor core. Carefully established administrative procedures plus the design of the system minimizes the probability of potential fission product release during the refueling operation. The system also includes the following features:

1. Safe accessibility for operating personnel.

2. Provisions to prevent fuel storage criticality.

3. Visual monitoring of the refueling procedures at all times.

1.2.2.9 Engineered Safety Features The engineered safety features for this plant have sufficient redundancy of component and power sources such that under the conditions of a hypothetical loss-of-coolant, the system can, even when operating with partial effectiveness, maintain the integrity of the containment and keep the exposure of the public within applicable limits. The major engineered safeguards systems are as follows:

1. The containment system, which incorporates continuously pressurized and monitored penetrations and liner weld channels and a seal water injection system, which provides a highly reliable, essentially leaktight barrier against the escape of radioactivity, which might be released to the containment atmosphere.

2. The safety injection system (which constitutes the emergency core cooling system) provides borated water to cool the core in the event of a loss-of-coolant accident.

3. The containment air recirculation cooling system provides a heat sink to cool the containment atmosphere.

4. The containment spray system provides a spray of cool, borated water to the containment atmosphere that is a heat sink and also provides iodine removal capability.

1.2.2.10 Structures The major structures are the reactor containment building, the primary auxiliary building, the control building, the fuel storage building, the turbine building, and the maintenance and Chapter 1, Page 6 of 71 Revision 22, 2010

IP2 FSAR UPDATE operations building. General layouts of the reactor containment interior components arrangement are shown on Plant Drawings 9321-2501, 9321-2502, 9321-2503, 9321-2506, 9321-2507, 9321-2508 [formerly UFSAR Figures 5.1-2 through 5.2-7]. General layouts and interior components arrangement of the primary auxiliary building, control building, fuel storage building, and holdup tank building were removed from the UFSAR due to security reasons following September 11, 2001 and can be viewed on plant drawings. 1.2.2.11 Containment The reactor containment is a steel-lined reinforced concrete cylinder with a hemispherical dome and a flat base. The containment is designed to withstand the internal pressure accompanying a loss-of-coolant accident, is virtually leaktight, and provides adequate radiation shielding for both normal operation and accident conditions. When required, the containment isolation valve seal water system permits automatic rapid sealing of pipes, which penetrate the containment so that in the event of any loss-of-coolant accident, leakage from containment to the environment is minimal. Ground accelerations for the operational basis earthquake used for containment design purposes and all seismic Class I structures (Section 1.11) are 0.10g applied horizontally and 0.05g applied vertically. In addition, ground accelerations for the design basis earthquake of 0.15g horizontal and 0.10g vertical are used to analyze the no loss-of-function concept. 1.2 FIGURES Figure No. Title Figure 1.2-1 Indian Point Nuclear Generating Units 1 & 2 [Historical] Figure 1.2-2 Deleted Figure 1.2-3 Deleted Figure 1.2-4 Cross Section of Plant [Historical] Figure 1.2-5 Deleted Figure 1.2-6 Deleted Figure 1.2-7 Deleted Figure 1.2-7 Deleted Figure 1.2-8 Deleted Figure 1.2-9 Deleted 1.3 GENERAL DESIGN CRITERIA (GDC) The General Design Criteria define or describe safety objectives and approaches incorporated in the design of this plant. These General Design Criteria, tabulated explicitly in the pertinent systems sections in this report, comprised the proposed Atomic Industrial Forum versions of the criteria issued for comment by the AEC on July 11, 1967. Also included in this section, are brief descriptions of related plant features, which are provided to meet the design objectives reflected in the criteria at the time of the initial license application. The descriptions are more fully developed in those succeeding sections of the report indicated by the references. More recently, Con Edison completed a study of compliance with 10 CFR Parts 20 and 50 in accordance with the Commission's Confirmatory Order of February 11, 1980. The detailed Chapter 1, Page 7 of 71 Revision 22, 2010

IP2 FSAR UPDATE results of the evaluation of Indian Point Unit 2 compliance with the then current General Design Criteria established by the Nuclear Regulatory Commission (NRC) in 10 CFR 50 Appendix A, were submitted to the NRC by Con Edison on August 11, 1980 (Reference 1). Commission concurrence was received on January 19, 1982. The parenthetical numbers following the section headings indicate the numbers of their related proposed Atomic Industrial Forum versions of the General Design Criteria as described in the first paragraph of this section. 1.3.1 Overall Plant REQUIREMENTS (GDC 1 - GDC 5) All systems and components of the facility are classified according to their importance. Those items vital to safe shutdown and isolation of the reactor or whose failure might cause or increase the severity of a loss-of-coolant accident or result in an uncontrolled release of excessive amounts of radioactivity are designated Class I. Those items important to reactor operation but not essential to safe shutdown and isolation of the reactor or control of the release of substantial amounts of radioactivity are designated Class II. Those items not related to reactor operation or safety are designated Class III. Class I systems and components are essential to the protection of the health and safety of the public. Consequently, they are designed, fabricated, inspected, erected, and use materials selected to the applicable provisions of recognized codes, good nuclear practice and to quality standards that reflect their importance. All systems and components designated Class I are designed so that there is no loss of function in the event of the maximum potential ground acceleration acting in the horizontal and vertical directions simultaneously. The working stresses of both Class I and Class II items are kept within code allowable values for the operational basis earthquake. Similarly, measures are taken in the plant design to protect against high winds, sudden barometric pressure changes, flooding, and other natural phenomena. Reference sections: Section Title Section Site & Environment; Meteorology 2.6 Geology and Seismology 2.7 Reactor Coolant System; Design Bases 4.1 Containment System Structures; Design Bases 5.1.1 Electrical Systems; Design Bases 8.1 Introduction & Summary; Design Criteria for Structures and Equipment 1.11 Fire prevention in all areas of the nuclear electric plant is provided by structure and component design, which maximizes the use of fire-resistant materials, optimizes the containment of combustible materials and maintains exposed combustible materials below their ignition temperature in the design atmosphere. Fixed and portable fire fighting equipment is provided with capacities proportional to the energy that might credibly be released by fire. Chapter 1, Page 8 of 71 Revision 22, 2010

IP2 FSAR UPDATE Reference sections: Section Title Section Instrumentation & Control; Operating Control Stations 7.7 Auxiliary & Emergency Systems; Facility Service System 9.6 The only structures, systems, or components important to safety that are shared between Units 2 and 3 are:

1. The cooling water discharge channel, which carries the service water discharge to the river.

2. The Emergency Fuel Supply to the Emergency Diesel Generators.

Since the channel is designed to handle the discharge flow from both operating units, sharing of this structure will not impair the ability of Unit 2 safety systems to perform their safety functions. Units 2 and 3 are required by Technical Specification to maintain a designated amount of fuel reserve onsite or at the Buchanan substation, which is dedicated for emergency diesel generator use at that unit. Therefore sharing the Emergency Fuel Supply to the Emergency Diesel Generators will not impair the ability of Unit 2 safety systems to perform their safety functions. Reference sections: Section Title Section Electrical Systems; Emergency Fuel Supply 8.2.3.2 Auxiliary & Emergency Systems; Service Water System 9.6.1 A complete set of facility plant and system diagrams including arrangements, plans, and structural plans and records of initial tests and operation are maintained throughout the life of the reactor. A set of all the quality assurance data generated during fabrication and erection of the essential components of the plant, as defined by the quality assurance program, is retained. Reference sections: Section Title Section Conduct of Operations; Records 12.4 Initial Tests and Operation 13 Introduction & Summary; Quality Assurance Program 1.10 1.3.2 Protection By Multiple Fission Product Barriers (GDC 6 - GDC 10) The reactor core, with its related control and protection system, is designed to function throughout its design lifetime without exceeding acceptable fuel damage limits. The core design, together with reliable process and decay heat removal systems, provides for this capability under all expected conditions of normal operation with appropriate margins for uncertainties and anticipated transient situations. The reactor control and protection instrumentation is designed to actuate a reactor trip for any anticipated combination of plant conditions when necessary to ensure DNBR remains at or Chapter 1, Page 9 of 71 Revision 22, 2010

IP2 FSAR UPDATE above the applicable safety analysis DNBR limit and fuel center temperature below the melting point of uranium-dioxide. Reference sections: Section Title Section Reactor; Design Bases, 3.1 Reactor Design 3.2 Instrumentation and Control; Protective Systems 7.2 Safety Analysis 14 The design of the reactor core and related protection systems ensures that power oscillations, which could cause fuel damage in excess of acceptable limits are not possible or can be readily suppressed. Low frequency spatial xenon oscillations may occur in the axial dimension. However, the core is expected to be (and has proven to be) stable to xenon oscillations in the X-Y dimension. Ex-core instrumentation is provided to monitor any xenon induced oscillations. Incore instrumentation is used periodically to calibrate and verify the information provided by the ex-core instrumentation. The moderator temperature and overall power coefficients in the power operating range were maintained negative by inclusion of burnable poison shims in the first core loading. The overall power coefficient in the power operating range is always negative (as discussed in Section 14.1.11.2). Reference sections: Section Title Section Reactor; Design Bases, 3.1 Reactor Design 3.2 The reactor coolant system in conjunction with its control and protective provisions is designed to accommodate the system pressures and temperatures attained under all expected modes of plant operation or anticipated system interactions, and maintain the stresses within applicable code stress limits. The materials of construction of the pressure retaining boundary of the reactor coolant system are protected by control of coolant chemistry from corrosion phenomena, which might otherwise reduce the system structural integrity during its service lifetime. System conditions resulting from anticipated transients or malfunctions are monitored and appropriate action is initiated automatically to maintain the required cooling capability and to limit system conditions so that continued safe operation is achieved. The system is protected from overpressure by means of pressure relieving devices, as required by Section III of the ASME Boiler and Pressure Vessel Code and by an overpressure protection system intended to ensure compliance with 10 CFR 50, Appendix G. Isolable sections of the system are provided with overpressure relieving devices discharging to closed systems such that the system code allowable relief pressure within the protected section is not exceeded. Chapter 1, Page 10 of 71 Revision 22, 2010

IP2 FSAR UPDATE Reference section: Section Title Section Reactor Coolant System; Design Bases 4.1 The design pressure of the containment exceeds the peak pressure occurring as the result of the complete blowdown of the reactor coolant through any pipe rupture of the reactor coolant system up to and including the hypothetical severance of a reactor coolant system pipe. All piping systems, which penetrate the vapor barrier are anchored so that the penetration is structurally adequate to resist the piping loads and the vapor barrier will not be breached due to a hypothesized pipe rupture. The lines (with the exception of sample tubing) connected to the reactor coolant system that penetrate the vapor barrier are restrained near the secondary shield walls and are each provided with at least one valve between the shield wall and the reactor coolant system. These restraints are designed to withstand the thrust moment and torque resulting from a hypothesized rupture of the attached pipe. All isolation valves are supported to withstand, without impairment of valve operability, the combined loading of the design-basis accident and design seismic conditions. Reference section: Section Title Section Containment System Structures; Design Bases 5.1.1 1.3.3 Nuclear And Radiation Controls (GDC 11 - GDC 18) The plant is equipped with a control room, which contains all controls and instrumentation and facilities necessary for continuous operation of the reactor and turbine generator under normal and accident conditions. Sufficient shielding, ventilation control and filtration, and containment integrity are provided to ensure that control room personnel will not be subjected to doses under postulated accident conditions during occupancy of the control room, which in the aggregate, would exceed the applicable limits. Instrumentation and controls essential to avoid undue risk to the health and safety of the public are provided to monitor and maintain neutron flux, reactor coolant pressure, flow rate, temperature, and control rod positions within prescribed operating ranges. The non-nuclear regulating, process and containment instrumentation measures temperatures, pressures, flows, and levels in the reactor coolant system, steam systems, containment, and other auxiliary systems. The quantity and types of process instrumentation provided ensures safe and orderly operation of all systems and processes over the full operating range of the plant. The operational status of the reactor is monitored from the control room. When the reactor is subcritical, the neutron source multiplication is continuously monitored and indicated by proportional counters located in instrument wells in the primary shield adjacent to the reactor Chapter 1, Page 11 of 71 Revision 22, 2010

IP2 FSAR UPDATE vessel. The source detector channels are checked prior to operations in which criticality may be approached by the use of an incore source. Any appreciable increase in the neutron source multiplication, including that caused by the maximum physical boron dilution rate, is slow enough to give ample time to start corrective action (boron dilution stop and/or emergency boron injection) to prevent the core from becoming critical (as discussed in Sections 14.1.5.2.3 and 14.1.5.3). When the reactor is critical, means for showing the relative reactivity status of the reactor is provided by display of all rod bank positions in the control room. Periodic samples of the coolant boron concentration are taken, the variation of which provides a further check on the reactivity status of the reactor including core depletion during life. Instrumentation and controls provided for the protective systems are designed to trip the reactor when necessary to prevent or limit fission product release from the core, to limit energy release, to signal closure of containment isolation valves, and to control the operation of engineered safety features equipment. During reactor operation in the startup and power modes, redundant safety limit signals will automatically actuate two reactor trip breakers, which are in series with the rod drive mechanism coils. This action would interrupt power and initiate reactor trip. Reference sections: Section Title Section Instrumentation and Control; General Design Criteria, 7.1 Protective Systems, 7.2 Nuclear Instrumentation, 7.4 Operating Control Stations 7.7 The reactor protection system receives signals from plant instrumentation, which are indicative of an approach to an unsafe operating condition, actuates alarms, prevents control rod motion, initiates load cutback, and/or opens the reactor trip breakers, depending on the severity of the condition. The basic reactor tripping philosophy is to define a region of power and coolant temperature conditions allowed by the primary tripping functions (e.g., overpower T trip, overtemperature T trip, and nuclear flux trips). The allowable operating region within these trip settings is shown to prevent any combination of power, temperature, and pressure, which would result in DNB with all reactor coolant pumps in operation. Additional tripping functions such as low and high pressurizer pressure trips, high pressurizer level trip, loss of flow trip, steam and feedwater flow mismatch trip, steam generator low-low level trip, turbine trip, safety injection trip, nuclear flux trips (source, intermediate, and high range), and manual trip are provided as backup to the primary tripping functions for specific accident conditions and mechanical failures. Intermediate Range and Power Range rod stops, Overtemperature T and Overpower T rod stops, are provided to prevent abnormal power conditions which could result from excessive control rod withdrawal initiated by operator violation of administrative procedures. Reference sections: Chapter 1, Page 12 of 71 Revision 22, 2010

IP2 FSAR UPDATE Section Title Section Engineered Safety Features; Safety Injection System 6.2 Instrumentation and Control; Protective Systems 7.2 Positive indications in the control room of leakage of coolant from the reactor coolant system to the containment are provided by equipment, which permits continuous monitoring of containment air activity and humidity, and of runoff from the condensate collecting pans under the cooling coils of the containment air recirculation units. The basic design criterion is the detection of deviations from normal containment environmental conditions including air particulate activity, radiogas activity, humidity, condensate runoff and in addition, in the case of gross leakage, the liquid inventory in the process systems and containment sump. The containment atmosphere, the plant vent, the containment fan-coolers service water discharge, the waste disposal system liquid effluent, and the component cooling loop are monitored for radioactivity concentration during all normal operations, anticipated transients, and accident conditions. For the case of leakage from the reactor containment under accident conditions, the plant area radiation monitoring system supplemented by portable survey equipment to be kept in the control room provide adequate monitoring of accident releases. Monitoring and alarm instrumentation is provided for fuel and waste storage and handling areas to detect excessive radiation levels. The permanent record of activity releases is provided by radiochemical analysis of known quantities of waste. A controlled ventilation system removes gaseous radioactivity from various areas of the plant and discharges it to the atmosphere via the plant vent. Radiation monitors are in continuous service in these areas to actuate high activity alarms on the control board annunciator and initiate containment isolation. Reference sections: Section Title Section Engineered Safety Features; Safety Injection System, 6.2 Isolation Valve Seal Water System, 6.5 Leakage Detection and Provisions for the Primary and Auxiliary Coolant Loops 6.7 Auxiliary & Emergency Systems; Auxiliary Coolant System 9.3 Waste Disposal & Radiation Protections System; Radiation Protection 11.2 1.3.4 Reliability And Testability Of Protection Systems (GDC 19 - GDC 26) Upon a loss of power to the magnetic-type control rod drive mechanisms, the rod cluster control (RCC) assembly is released and falls by gravity into the core. The reactor internals, fuel assemblies, RCC assemblies and drive system components are designed as Class I equipment. The RCC assemblies are fully guided through the fuel assembly and for the maximum travel of Chapter 1, Page 13 of 71 Revision 22, 2010

IP2 FSAR UPDATE the control rod into the guide tube. Furthermore, the RCC assemblies are never fully withdrawn from their guide thimbles in the fuel assembly. Due to this and the flexibility designed into the RCC assemblies, abnormal loadings and misalignments can be sustained without impairing operation of the RCC assemblies. All reactor protection channels are supplied with sufficient redundancy to provide the capability for channel calibration and test at power. The analog channel trip bistables have the capability to be bypassed for surveillance testing. This test in bypass feature enables analog logic relays in a protective channel to be bypassed while testing actuation of the associated bistable. Bistable testing does not preclude the protective action provided by concurrent channels. Two reactor trip breakers are provided to interrupt power to the rod drive mechanisms. The breaker main contacts are connected in series with the power supply to the mechanism coils. Opening either breaker interrupts power to all full length rod mechanisms. Each breaker is opened through an undervoltage trip coil. Each protection channel actuates two separate trip logic trains, one for each reactor trip breaker undervoltage trip coil. The protection system is thus inherently safe in the event of a loss of rod control power. Channel independence is carried throughout the system extending from the sensor to the relay providing the logic. In most cases, the safety and control functions when combined are combined only at the sensor. A failure in the control circuitry does not affect the safety channel. This approach is used for pressurizer pressure and water level channels, steam generator water level, Tavg and T channels, steam flow, and nuclear power range channels. The power supplied to the channels is fed from four vital instrument buses. All four of the buses are supplied by static inverters. The initiation of the engineered safety features provided for loss-of-coolant accidents (e.g., high head safety injection, residual heat removal pumps, and containment spray systems) is accomplished from redundant signals derived from reactor coolant system and containment instrumentation. The initiation signal for containment spray comes from coincidence of two sets of two-out-of-three high-high containment pressure signals. On loss of voltage to the safety features equipment buses, the emergency diesel generators will be automatically started and connected to their respective buses. Trip signals for the containment isolation valves are derived from either a high-high containment pressure signal (phase B), and/or a safety injection signal (phase A), and/or a containment ventilation isolation signal. The components of the protection system are designed and laid out so that the mechanical and thermal environment accompanying any emergency situation in which the components are required to function does not interfere with that function. Each protection channel in service at power is capable of being calibrated and tripped independently by simulated signals for test purposes to verify its operation. Redundancy is provided in that there are three emergency diesel-generator sets capable of supplying separate 480-V buses. The minimum complement of safety features equipment is supplied from any two of the three emergency diesel generators. The ability of the emergency diesel-generator sets to start within the prescribed time and to carry load is periodically checked. Chapter 1, Page 14 of 71 Revision 22, 2010

IP2 FSAR UPDATE An open circuit or loss of reactor trip channel power causes the system to go into its trip mode. In a two-out-of-three circuit, the three channels are equipped with separate primary sensors and each channel is energized from independent electrical buses. The signal for containment isolation is developed from a two-out-of-three circuit in which each channel is separate and independent and which signals for containment isolation upon loss of power. The failure of any channel to de-energize when required does not interfere with the proper functioning of the isolation circuit. Diesel engine cranking is accomplished by a stored energy system supplied solely for the associated emergency diesel generator. The undervoltage relay scheme is designed so that loss of 480-V power does not prevent the relay scheme from functioning properly. Reference sections: Section Title Section Instrumentation and Control; Protection Systems 7.2 Electrical Systems 8 1.3.5 Reactivity Control (GDC 27 - GDC 32) Reactivity control is achieved by two independent systems, the rod cluster control assemblies and the chemical and volume control system, which regulates the concentration of boric acid solution neutron absorber in the reactor coolant system. The system is designed to prevent, under anticipated system malfunction, uncontrolled or inadvertent reactivity changes, which might stress the system beyond design limits. The reactivity control systems provided are capable of making and holding the core subcritical from any hot standby or hot operating condition, including those resulting from power changes. The maximum excess reactivity expected for the core occurs for the cold, clean condition at the beginning-of-life (as discussed in Section 14.1.11.2). The full length rod cluster control assemblies are divided into two categories comprising control banks and shutdown banks. The control bank of the rod cluster control assemblies is used to compensate for short-term reactivity changes at power produced due to variations in reactor power or in coolant temperature. The chemical shim control is used to compensate for the more slowly occurring changes in reactivity throughout core life such as those due to fuel depletion and fission product buildup and decay. The shutdown banks are provided to supplement the control banks of the rod cluster control assemblies to make the reactor at least 1-percent subcritical (keff = 0.99) following trip from any credible operating condition to the hot, zero power condition assuming the most reactive rod cluster control assembly remains in the fully withdrawn position. Boron injection from the safety injection system supplements rod insertion and prevents exceeding core safety limits in the event of the maximum credible steam break, namely opening of a safety valve. This is accomplished with maximum worth rod fully withdrawn. Chapter 1, Page 15 of 71 Revision 22, 2010

IP2 FSAR UPDATE Any time that the plant is at power, the quantity of boric acid retained in the boric acid storage tanks and ready for injection always exceeds that quantity required for the normal cold shutdown. The boric acid solution is transferred from the boric acid storage tanks into the reactor coolant by the boric acid transfer pumps and charging pumps, which can be operated from emergency diesel-generator power on loss of primary power. Boric acid can be injected by either boric acid transfer pumps and one of the three charging pumps to shut down the reactor from full power with no rods inserted. In addition, boric acid can be injected to compensate for xenon decay (xenon decay below the equilibrium operating level will not actually begin until approximately 20 hr after shutdown). Additional boric acid is employed if it is desired to bring the reactor to cold shutdown conditions. The reactor protection systems will limit reactivity transients such that DNBR remains at or above the applicable safety analysis DNBR limit due to any single malfunction in the reactor coolant deboration controls. Limits, which include considerable margin, are placed on the maximum reactivity worth of control rods by limiting position of insertion as a function of power and on rates at which reactivity can be increased to ensure that the potential effects of a sudden or large change of reactivity cannot (a) rupture the reactor coolant pressure boundary or (b) disrupt the core, its support structures, or other vessel internals so as to lose capability to cool the core. The rod cluster drive mechanisms are wired into preselected groups and are prevented from being withdrawn in other than their respective groups. The rod drive mechanism is of the magnetic latch type and the coil actuation is sequenced to provide variable speed rod travel. The maximum reactivity insertion rate is analyzed in the detailed plant analysis assuming two of the highest worth banks to be accidentally withdrawn at maximum speed, yielding reactivity insertion rates of the order of 8 x 10-4 k/sec, which is well within the capability of the overpower-overtemperature protection circuits to prevent core damage (as discussed in Section 14.1.1). Reference sections: Section Title Section Reactor; Design Bases 3.1 Instrumentation and Control; Protective Systems, 7.2 Regulating Systems 7.3 Auxiliary & Emergency Systems; Chemical and Volume Control System 9.2 1.3.6 Reactor Coolant Pressure Boundary (GDC 33 - GDC 36) The reactor coolant pressure boundary is shown to be capable of accommodating without further rupture, the static and dynamic loads imposed as a result of a sudden reactivity insertion such as rod ejection (as discussed in Section 14.2.6.10). The operation of the reactor is such that the severity of an ejection accident is inherently limited. Since the rod cluster control assemblies are used to control load variations only and core depletion is followed with boron dilution, only the rod cluster control assemblies in the controlling Chapter 1, Page 16 of 71 Revision 22, 2010

IP2 FSAR UPDATE groups are inserted in the core at power, and at full power these rods are only partially inserted. A rod insertion limit monitor is provided as an administrative aid to the operator to ensure that this condition is met. By using the flexibility in the selection of control rod groupings, radial locations and positions as a function of load, the design limits the maximum fuel temperature for the highest worth ejected rod to a value, which precludes any resultant damage to the primary system pressure boundary. The failure of a rod mechanism housing causing a rod cluster control assembly to be rapidly ejected from the core is evaluated as a theoretical, though not a credible, accident. The analysis is discussed in Section 14.2.6. In the core region of the reactor vessel, the V-notch toughness of the material will change during operation as a result of fast neutron exposure, which results in a shift in the nil ductility transition temperature (NDTT). This is factored into the operating procedures in such a manner that full operating pressure is not obtained until the affected vessel material is above the increased design transition temperature (DTT) and in the ductile material region. The pressure during startup and shutdown at the temperature below NDTT is maintained below the threshold of concern for safe operation. The DTT is a minimum NDTT plus 60oF and dictates the procedures to be followed in the hydrostatic test and in station operations to avoid excessive cold stress. The value of the DTT is increased during the life of the plant as required by the expected shift in the NDTT, and as confirmed by the experimental data obtained from irradiated specimens of reactor vessel materials during the plant lifetime. The design of the reactor vessel and its arrangement in the system provide the capability for accessibility during service life to the entire internal surfaces of the vessel and the external zones of the vessel including the nozzle to reactor coolant piping welds and the top and bottom heads. The reactor arrangement within the containment provides sufficient space for inspection of the external surfaces of the reactor coolant piping, except for the area of piping within the primary shielding concrete. Determination of the NDTT of the core region plate forgings, weldments, and associated heat treated zones is performed in accordance with ASTM E-185, Recommended Practice for Surveillance Tests on Structural Materials in Nuclear Reactors. Samples of reactor vessel plate material have been retained and catalogued in case future engineering development shows the need for further testing. The material properties surveillance program includes not only the conventional tensile and impact tests, but also fracture mechanics specimens. The observed shifts in NDTT of the core region materials with irradiation will be used to confirm the calculated limits to startup and shutdown transients. To define permissible operating conditions below DTT, a pressure range is established, which is bounded by a lower limit for pump operation and an upper limit, which satisfies reactor vessel stress criteria. Since the normal operating temperature of the reactor vessel is well above the maximum expected DTT, brittle fracture during normal operation is not considered to be a credible mode of failure. Reference sections: Chapter 1, Page 17 of 71 Revision 22, 2010

IP2 FSAR UPDATE Section Title Section Reactor Coolant System; System Design and Operation, 4.2 Safety Limits and Conditions, 4.4 Inspections and Tests, 4.5 Determination of Reactor Pressure Vessel NDTT Appendix 4A 1.3.7 Engineered Safety Features (GDC 37 - GDC 65) The design, fabrication, testing and inspection of the core, reactor coolant pressure boundary, and their protection systems give assurance of safe and reliable operation under all anticipated normal, transient, and accident conditions. However, engineered safety features are provided in the facility to back up the safety provided by these components. These engineered safety features have been designed to cope with any size reactor coolant pipe break up to and including the circumferential rupture of any pipe in that boundary assuming unobstructed discharge from both ends as discussed in Section 14.3.3.3. They are also designed to cope with any steam or feedwater line break up to and including the main steam or feedwater headers as discussed in Section 14.2.5. The total loss of all offsite power is assumed concurrent with these accidents. The release of fission products from the reactor fuel is limited by the safety injection system, which by cooling the core and limiting the fuel clad temperature, keeps the fuel in place and substantially intact with its essential heat transfer geometry preserved and limits the metal-water reaction to an insignificant amount. The basic design criteria for ensuring that the core geometry remains in place and substantially intact so that effective cooling of the core is not impaired following a loss-of-coolant accident:

1. The cladding temperature is to be less than:

a. The melting temperature of Zircaloy-4

b. The temperature at which gross core geometry distortion, including fragmentation, may be expected.

2. The total core metal-water reaction will be limited to less than 1-percent.

The safety injection system (which constitutes the emergency core cooling system) consists of high and low head centrifugal pumps driven by electric motors, and passive accumulator tanks, which are self-energized and which act independently of any actuation signal or power source. The release of fission products from the containment is limited in three ways:

1. Blocking the potential leakage paths from the containment. This is accomplished by:

a. A steel-lined concrete reactor containment with continuously pressurized double-sealed penetrations and liner weld channels, which form a virtually leaktight barrier to the escape of fission products should a loss-of-coolant accident occur.

b. Isolation of process lines by the containment isolation system, which imposes double barriers in each line penetrating the containment.

Chapter 1, Page 18 of 71 Revision 22, 2010

IP2 FSAR UPDATE

2. Reducing the fission product concentration in the containment atmosphere. This is accomplished by containment spray, which removes elemental iodine vapor and particulates from the containment atmosphere by washing action.

3. Reducing the containment pressure and thereby limiting the driving potential for fission product leakage. This is accomplished by cooling the containment atmosphere by the following independent systems, each with adequate heat removal capacity:

a. Containment spray system

b. Containment air recirculation cooling system.

A comprehensive program of plant testing is performed for all equipment systems and system control vital to the functioning of engineered safety features. The program consists of performance tests of individual pieces of equipment in the manufacturer's shop and integrated tests of the system as a whole, and periodic tests of the actuation circuitry and mechanical components to assure reliable performance, upon demand, throughout the plant lifetime. In the event that one of the components should require maintenance as a result of failure to perform during the test according to prescribed limits, the necessary corrections or minor maintenance will be made and the unit retested immediately. The plant is supplied with emergency power sources as follows:

1. Three independent emergency diesel generators, located in the Diesel Generator Building adjacent to the Primary Auxiliary Building, supply emergency power to the engineered safety features buses in the event of a loss of AC auxiliary power. There are no automatic bus ties associated with these buses. Each diesel generator is started automatically on a safety injection signal or upon the occurrence of an undervoltage condition on any vital 480-V switchgear bus. The system is sufficiently redundant such that any two diesels have adequate capacity to supply the engineered safety features for the design basis accident concurrent with a loss of offsite power. One diesel is adequate to provide power for a safe and orderly plant shutdown in the event of a loss-of-offsite electrical power.

2. Emergency power for vital instrumentation and control and for emergency lighting is supplied from the 125 VDC system via four independent DC channels. The station batteries supply emergency power to the instrumentation and control systems when their associated battery chargers are not available.

For such engineered safety features as are required to ensure safety in the event of an accident, protection from dynamic effects or missiles is considered in the layout of plant equipment and missile barriers. Layout and structural design specifically protect injection paths leading to unbroken reactor coolant loops against damage as a result of the maximum reactor coolant pipe rupture. Injection lines penetrate the main missile barrier, and the injection headers are located in the missile-protected area between the missile barrier and the containment outside wall. Individual injection lines, separated to the maximum extent practicable, are connected to the injection header, pass through the barrier and then connect to the loops. Movement of the injection line, associated with rupture of a reactor coolant loop, is accommodated by line flexibility and by the design of the pipe supports such that no damage outside the missile barrier is possible. Chapter 1, Page 19 of 71 Revision 22, 2010

IP2 FSAR UPDATE In 1989, the NRC approved changes to the design basis with respect to dynamic effects of postulated primary loop ruptures, as discussed in Section 4.1.2.4. Each engineered safety feature provides sufficient performance capability to accommodate any single failure of an active component and still function in a manner to avoid undue risk to the health and safety of the public. Under the hypothetical accident conditions, the containment air recirculation cooling system, and the containment spray system are designed and sized so that either system, operating alone at its rated capacity, is able to supply the necessary postaccident cooling capacity to reduce rapidly the containment pressure following blowdown and cooling of the core by safety injection. Together these two systems provide the single failure protection for the containment cooling function as analyzed in Chapter 14. All active components of the safety injection system (exception: injection line isolation valves) and the containment spray system are located outside the containment and not subject to containment accident conditions. Instrumentation, motors, cables, and penetrations located inside the containment are selected to meet the most adverse accident conditions to which they may be subjected. These items are either protected from containment accident conditions or are designed to withstand, without failure, exposure to the worst combination of temperature, pressure, radiation, and humidity expected during the required operational period. The reactor is maintained subcritical following a primary system pipe rupture accident. Introduction of borated cooling water into the core results in a net negative reactivity addition (as discussed in Section 14.3.2). The control rods insert and remain inserted, except for the large break loss of coolant accident analysis where it is conservatively assumed that the control rods do not insert as discussed in Section 14.3.3.2. The delivery of cold safety injection water to the reactor vessel following accidental expulsion of reactor coolant was evaluated to ensure that this does not cause further loss of integrity of the reactor coolant system boundary (as discussed in Section 14.3.4.3.3). Design provisions have been made to the extent practical to facilitate access to the critical parts of the reactor vessel internals, injection nozzles, pipes, valves, and safety injection pumps for visual or boroscopic inspection for erosion, corrosion, and vibration wear evidence, and for non-destructive test inspection where such techniques are desirable and appropriate. Design provisions are made so that active components of the safety injection system can be tested periodically for operability and functional performance. The safety injection pumps can be tested periodically during plant operation using the minimum flow recirculation lines provided. The residual heat removal pumps are used every time the residual heat removal loop is put into operation. An integrated safety injection system test is performed at refueling outage intervals. This test does not introduce flow into the reactor coolant system but demonstrates the operation of the valves, pump circuit breakers, and automatic circuitry upon initiation of safety injection. The accumulator tank pressure and level are continuously monitored during plant operation and flow from the tanks can be checked at any time using test lines. Chapter 1, Page 20 of 71 Revision 22, 2010

IP2 FSAR UPDATE The accumulators and the safety injection piping up to the final isolation valve are maintained sufficiently filled of borated water at refueling water concentration while the plant is in operation to ensure the system remains operable and performs properly. Flow in each of the high head injection headers and in the main flow line for the residual heat removal pumps is monitored by flow and pressure instrumentation. The design provided for capability to test initially, to the extent practical, the full operational sequence up to the design conditions for the safety injection system to demonstrate the state of readiness and capability of the system. These functional tests provided information to confirm valve operating times, pump motor starting times, the proper automatic sequencing of load addition to emergency diesel generators, and delivery rates of injection water to the reactor coolant system. The following general criteria were followed to ensure conservatism in computing the required containment structural load capacity:

1. In calculating the containment pressure, rupture sizes up to and including a double-ended severance of reactor coolant pipe were considered.

2. In considering postaccident pressure effects, various malfunctions of the emergency systems were evaluated. Contingent mechanical or electrical failures were assumed to disable one of the emergency diesel electric generators, two of the fan coolers and one of the containment spray pumps (as discussed in Section 14.3.5.3.7).

3. The pressure and temperature loading obtained by analyzing various loss-of-coolant accidents, when combined with operating loads and maximum wind or seismic forces, does not exceed the load-carrying capacity of the structure, its access opening or penetrations.

Discharge of reactor coolant through a double-ended rupture of the main loop piping, followed by operation of only those engineered safety features, which can run simultaneously with power from two of the three emergency onsite diesel generators, results in a sufficiently low radioactive materials leakage from the containment structure that there is no undue risk to the health and safety of the public. The concrete containment is not susceptible to a low temperature brittle fracture. The containment liner is enclosed within the containment and thus is not exposed to the temperature extremes of the environs. The containment ambient temperature during operation is between 90°F and 130°F, which is well above the NDTT + 30°F for the liner material. Containment penetrations, which can be exposed to the environment are also designed to the NDTT + 30°F criterion. The reactor coolant pressure boundary does not extend outside of the containment. Isolation valves for all fluid system lines penetrating the containment provide at least two barriers against leakage of radioactive fluids to the environment in the event of a loss-of-coolant accident. These barriers, in the form of isolation valves or closed systems, are defined on an individual line basis. In addition to satisfying containment isolation criteria, the valving is designed to facilitate normal operation and maintenance of the systems and to ensure reliable operation of other engineered safety features. Chapter 1, Page 21 of 71 Revision 22, 2010

IP2 FSAR UPDATE After completion of the containment structure and installation of all penetration and weld channels, an initial integrated leakage rate test was conducted at the peak calculated accident pressure, maintained for a minimum of 24 hr, to verify that the leakage rate was not greater than 0.1-percent by weight of the containment volume per day. This leakage rate test was performed using the reference vessel method. A leak rate test at the peak calculated accident pressure using the same method as the initial leak rate test can be performed at any time during the operational life of the plant, provided the plant is not in operation and precautions are taken to protect instruments and equipment from damage. Penetrations are designed with double seals so as to permit pressurization of the interior of the penetration whenever a leak test is required. The system utilizes a supply of clean, dry, compressed air which places all the penetrations under an internal pressure above the peak calculated accident pressure (Peak calculated accident pressure is discussed in Section 14.3.5.1.1). Leakage from the system is checked by measurement of the integrated makeup air flow. In the event excessive leakage is discovered, each penetration can then be checked separately. Capability is provided to the extent practical for testing the functional operability of valves and associated apparatus during periods of reactor shutdown. Initiation of containment isolation employs coincidence circuits, which allow checking of the operability and calibration of one channel at a time. The main steam and feedwater barriers and isolation valves in systems, which connect to the reactor coolant system are hydrostatically tested to measure leakage. Design provisions are made to the extent practical to facilitate access for periodic visual inspection of all important components of the containment air recirculation cooling and containment spray systems. The containment pressure-reducing systems are designed to the extent practical so that the spray pumps, spray injection valves, spray nozzles can be tested periodically and after any component maintenance for operability and functional performance. Permanent test lines for the containment spray loop are located so that all components up to the isolation valve at the spray nozzle may be tested. These isolation valves are checked separately. The air test lines for checking that spray nozzles are not obstructed connect upstream of the isolation valve. Air flow through the nozzles is verified by periodic testing in accordance with the Technical Specifications. Capability is provided to test initially to the extent practical the operational startup sequence beginning with transfer to alternate power sources and ending with near design conditions for the containment air recirculation cooling and containment spray systems. Chapter 1, Page 22 of 71 Revision 22, 2010

IP2 FSAR UPDATE Reference sections: Section Title Section Containment System 5 Engineered Safety Features 6 Electrical Systems; Design Bases, 8.1 Electrical Systems Design 8.2 1.3.8 Fuel And Waste Storage Systems (GDC 66 - GDC 69) The new and spent fuel storage racks are designed so that it is impossible to insert assemblies in other than the prescribed locations. The spent fuel storage pit is filled with borated water, normally at a similar concentration to that used in the reactor cavity and refueling canal during refueling operations. The fuel is stored vertically in an array with sufficient neutron absorbers and distance between assemblies to assure keff <1.0 even if unborated water were used to fill the pit and 0.95 when filled with water borated 2000 ppm boron. During reactor vessel head removal and while loading and unloading fuel from the reactor, the boron concentration is maintained at not less than that required to maintain the reactor subcritical by 5-percent k/k with all the rods inserted. The refueling water boron concentration is periodically checked to ensure the proper shutdown margin. The design of the fuel handling equipment incorporating built-in interlocks and safety features, the use of detailed refueling instructions and observance of minimum operating conditions provide assurance that no incident could occur during the refueling operations that would result in a hazard to public health and safety. The refueling system interlocks are verified to be functioning each refueling shutdown prior to refueling operations. The refueling water provides a reliable and adequate cooling medium for spent fuel transfer. Heat removal from the refueling water is provided by an auxiliary cooling system. Adequate shielding for radiation protection is provided during reactor refueling by conducting all spent fuel transfer and storage operations underwater. This permits visual control of the operation at all times while maintaining low radiation levels for periodic occupancy of the area by operating personnel. Pit water level is alarmed and water to be removed from the pit must be pumped out as there are no gravity drains when the pit is isolated from the refueling canal. Shielding is provided for waste handling and storage facilities to permit operation within requirements of 10 CFR 20. Gamma radiation is continuously monitored in the fuel storage building. A high level signal is alarmed locally and is annunciated in the control room. Auxiliary shielding for the waste disposal system and its storage components was also designed to limit the dose rate. All fuel and waste storage facilities are contained and equipment designed so that accidental releases of radioactivity directly to the atmosphere are monitored and will not exceed the applicable limits. Chapter 1, Page 23 of 71 Revision 22, 2010

IP2 FSAR UPDATE The refueling canal and spent fuel storage pit are reinforced concrete structures with a seam-welded stainless steel plate liner. These structures are designed to withstand any anticipated earthquake loadings as seismic Class I structures so that the liner should prevent leakage even in the event the reinforced concrete develops cracks. Reference sections: Section Title Section Auxiliary & Emergency Systems; Sampling System 9.4 Waste Disposal & Radiation Protection System; Waste Disposal System 11.1 Radiation Protection Systems 11.2 1.3.9 Plant Effluents (GDC 70) Liquid, gaseous, and solid waste disposal facilities are designed so that discharge of effluents and offsite shipments are in accordance with applicable governmental regulations. Radioactive fluids entering the waste disposal system are collected in sumps and tanks until determination of subsequent treatment can be made. They are sampled and analyzed to determine the quantity of radioactivity, with an isotopic breakdown if necessary. Before any attempt is made to discharge, they are processed as required and then released under controlled conditions. The system design and operation are characteristically directed toward minimizing releases to unrestricted areas. Discharge streams are appropriately monitored and safety features are incorporated to preclude excessive releases. Radioactive gases are pumped by compressors through a manifold to one of the gas decay tanks where they are held a suitable period of time for decay. Cover gases in the nitrogen blanketing system are reused to minimize gaseous wastes. During normal operation, gases are discharged intermittently at a controlled rate from these tanks through the monitored plant vent. The system is provided with discharge controls so that environmental conditions do not restrict the release of radioactive effluents to the atmosphere. Liquid wastes are processed to remove most of the radioactive materials. The spent resins from the demineralizers and the filter cartridges are packaged and stored onsite until shipment offsite for disposal. Reference section: Section Title Section Waste Disposal & Radiation Protection System; Waste Disposal System

11.1 REFERENCES

FOR SECTION 1.3

1. Letter from P. Zarakas, Con Edison, to H. Denton, NRC,

Subject:

Actions Taken to Comply with NRC Confirmatory Order of February 11, 1980, dated August 11, 1980.

2. Deleted Chapter 1, Page 24 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.4 DESIGN PARAMETERS AND PLANT COMPARISON 1.4.1 Design Highlights The original design of the plant is based upon proven concepts, which have been developed and successfully applied in the construction of pressurized water reactor systems. In subsequent paragraphs, the original design features of the plant are discussed. 1.4.1.1 Power Level The initial license application power level for Indian Point Unit 2 was 2758 MWt. The increase in this power rating over 1473 MWt for Connecticut Yankee was achieved by a 44-percent increase in heat transfer surface area and a 31-percent increase in average heat flux. The increased heat transfer surface area is due to 22-percent more fuel rods, each 20-percent longer. The increase in maximum heat flux and the 13.4 kW/ft linear heat generation rate (LHGR) resulting are justified by the results of incore experiments by Westinghouse and others. 1.4.1.2 Reactor Coolant Loops The reactor coolant system for the Indian Point Unit 2 consists of four loops as compared with three loops for San Onofre and four loops for Connecticut Yankee. The use of four loops in the Indian Point Unit 2 for the production of 2758 MWt requires an attendant increase in the size and capacity of the reactor coolant system components such as the reactor vessel, reactor coolant pumps, piping, and steam generators. These increases represent reasonable engineering extrapolations of existing and proven designs. 1.4.1.3 Peak Specific Power Based on values for hot channel factors, reactivity coefficients, and other design parameters, which were established in the PSAR and are supported by the previous experience with other plants of the same type, this reactor can be operated safely at power levels at least as high as the license application rating. 1.4.1.4 Fuel Cladding The fuel rod design for the plant employs zircaloy as a cladding material. This clad has proven successful in numerous operating facilities. The fuel rod dimensions are identical to those in Ginna, Salem, and Zion Station Units 1 and 2. 1.4.1.5 Fuel Assembly Design The fuel assembly incorporates the rod cluster control concept in a canless 15 x 15 fuel rod assembly using a spring clip grid to provide support for the fuel rods. Extensive out-of-pile and in-pile tests have been performed on this concept; operating experience is available from numerous facilities. Chapter 1, Page 25 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.4.1.6 Moderator Temperature Coefficient of Reactivity The reactor has a negative moderator temperature coefficient of reactivity at operating temperature at all times throughout core life (as discussed in Section 14.1.11.2). 1.4.2 IP2 - IP3 Design Differences An NRC Confirmatory Order (Reference 1) required the Licensees (Consolidated Edison and the Power Authority) to jointly review and identify the significant differences between Indian Point Units 2 and 3 and evaluate these differences in light of the regulatory standards and requirements in existence at the time. Consolidated Edison determined, evaluated, and provided justification for each design difference in submittals to the NRC for acceptance (References 2, 3). These design differences were found acceptable in an NRC Safety Evaluation Report (Reference 4). REFERENCES FOR SECTION 1.4

1. Letter from Harold R. Denton, NRC, to Consolidated Edison, "Confirmatory Order", dated February 11, 1980.

2. Letter from William J. Cahill, Consolidated Edison, to Harold R. Denton, NRC, "Confirmatory Order", dated May 9, 1980.

3. Letter from John D. OToole, Consolidated Edison, to Steven A. Varga, NRC, "Confirmatory Order", dated May 27, 1982.

4. Letter from Steven A. Varga, NRC, to John D. OToole, Consolidated Edison, "Confirmatory Order", dated December 1, 1982.

1.5 RESEARCH AND DEVELOPMENT REQUIREMENTS Research and development were conducted relating to finalization of core design details and parameters, air recirculation system halogen filters, failure of core cooling systems and means to ameliorate consequences, emergency core cooling system, control rod ejection analysis, and reactor coolant pump controlled leakage seals. The detailed final core design and thermal-hydraulics and physics parameters have been completed. The nuclear design including fuel configuration and enrichments, control rod pattern and worths, reactivity coefficients and boron requirements are presented in Section 3.2.1 and the final thermal-hydraulics design parameters are in Section 3.2.2. Section 3.2.3 presents the final fuel, fuel rod, fuel assembly, and control rod mechanical design. The core design incorporates fixed burnable poison rods1 in the initial loading to ensure a negative moderator reactivity temperature coefficient at operating temperature. This improves reactor stability and lessens the consequences of a rod ejection or loss-of-coolant accident. Core stability has been analyzed2-4 and design provisions for detection and control of potential xenon oscillations have been finalized3. The original core design incorporated part-length control rods for controlling these xenon oscillations and shaping the axial power distribution. Chapter 1, Page 26 of 71 Revision 22, 2010

IP2 FSAR UPDATE These have since been found unnecessary and removed from the reactor. X-Y control is not required and therefore not provided. Tests in operating reactors demonstrate the ability of the out-of-core instrumentation to give accurate indication of power redistribution and provide the operator information necessary to monitor redistributions and control axial oscillations by moving the rods in a prescribed pattern3. This capability was verified during startup tests in the Indian Point Unit 2 Plant. Full-size filter tests were conducted for the Connecticut Yankee Atomic Power Company to demonstrate the efficiency for iodine removal under the most extreme conditions anticipated in the postaccident containment environment. The results of these tests5 filed with the former U. S. Atomic Energy Commission (now the U. S. Nuclear Regulatory Commission) under Docket No. 50-213, are directly applicable to the charcoal filter system originally employed in this plant. The charcoal filters are no longer required by the Technical Specifications and the radiological consequences analysis presented in Section 14.3.6 does not credit the filters. A program for development of a crucible system design, which would contain the reactor core assuming failure of the core cooling system to prevent a core meltdown, was undertaken. A scheme for containing the molten core in a water submerged high melting point refractory lined steel crucible resulted. Refractory materials and crucible physical design were investigated along with analysis of the temperature distribution expected with the molten core and crucible refractory, and steam and water recirculation paths. As a result of uncertainties in material properties at the high application temperature, the lack of experimental proof that the boiling core mass would dissipate its heat upward through the water cover, and the possibility of violent liquid metal-water reactions, it became apparent that the proper emphasis for research and development on the loss-of-coolant accident should be increased emphasis placed on research and development for emergency core cooling system improvement in order to eliminate need for a crucible. This is supported by the conclusions of the Report of Advisory Task Force on Power Reactor Emergency Cooling, "Emergency Core Cooling," USAEC. This additional development effort on emergency core cooling system design resulted in the modification of the system to include pressurized accumulator tanks for rapid core reflooding. This increased flooding capability limits the clad temperature after a loss-of-coolant accident to well below the melting temperature of Zircaloy-4, minimizes metal-water reaction and ensures that the core remains in place and intact thereby ensuring preservation of essential heat transfer geometry. The system design incorporates redundancy of components such that the minimum required water addition rates can be met assuming any active component to fail concurrent with the loss-of-coolant accident or, over the long term period of postaccident core decay heat removal a passive or active component failure in either the safety injection or service water systems, or an active failure in the component cooling water system. Details of system design and operation are given in Chapters 6 and 9, and analysis of the loss-of-coolant accident is presented in Section 14.3. Because of the incorporation of this revised emergency core cooling system, the reactor pit crucible was deleted from the plant design. Although it is not required to provide cooling for molten fuel in the bottom of the reactor vessel with the upgraded emergency core cooling system performance, the clearance between the insulation and the instrumentation penetrations with the pressure relief holes in the insulation at the top of the vessel provide assurance that water in the flooded reactor vessel cavity will be in contact with the vessel. No other provisions for direct vessel cooling are provided or required. A control rod ejection analysis was performed for the final core design, rod worths, rod position limits, and moderator reactivity temperature coefficient. As mentioned above, the addition of the burnable poison rods eliminates power operation with a positive moderator temperature Chapter 1, Page 27 of 71 Revision 22, 2010

IP2 FSAR UPDATE coefficient and reduces the severity of the ejected rod accident, hence, lessening the need for research and development on this subject. The analysis is presented in Section 14.2.6. The reactor coolant pump controlled leakage seal design for this plant has been fully developed. A full scale mock-up of this seal was operated for over 100 hr to confirm that seal deflection and leak rate under load were acceptable. The full scale mock-up has been used during the development of the controlled leakage seal to provide information related to long-term performance. One of the two seals used in this plant was operated about 300 hr and the other about 100 hr, each in its pump motor unit. During hot functional testing in the plant, before the core was loaded, additional operation brought the total operating time for each seal to well over 500 hr. Successful operation of similar seals had previously been demonstrated with over 5000 hr total running time in San Onofre and over 3000 hr in Haddam Neck. REFERENCES FOR SECTION 1.5

1. P. M. Wood, E. A. Bassler, et al., Use of Burnable Poison Rods in Westinghouse Pressurized Water Reactors, WCAP-7113, October 1967.

2. P. M. Wood, J. M. Gallagher, R. M. Metz, R. A. Dean, Use of Part-Length Absorber Rods in Westinghouse Pressurized Water Reactors, WCAP-7072, June 1967.

3. Westinghouse Electric Corporation, Power Distribution Control of Westinghouse Pressurized Water Reactors, WCAP-7208, October 1968.

4. Westinghouse Electric Corporation, Power Maldistribution Investigations, WCAP-7407-L (proprietary), January 1970.

5. Connecticut Yankee Charcoal Filter Tests, CYAP-101, December 1966.

1.6 IDENTIFICATION OF CONTRACTORS [Historical Information Only] The Indian Point Unit 2 was designed and built by the Westinghouse Electric Corporation as prime contractor for Con Edison. Westinghouse undertook to provide a complete, safe, and operable nuclear power plant ready for commercial service. The project was directed by Westinghouse from the offices of its Atomic Power Division in Pittsburgh, Pennsylvania, and by Westinghouse representatives at the plant site during construction and plant startup. Westinghouse engaged United Engineers and Constructors of Philadelphia, Pennsylvania, to provide the design of certain portions of the plant. The plant construction was under the general direction of Westinghouse through United Engineers and Constructors, which was responsible for the management of all site construction activities and either performed or subcontracted the work of construction and equipment erection. Preoperational testing of equipment and systems at the site and initial plant operation was performed by Con Edison personnel under the technical direction of Westinghouse. Chapter 1, Page 28 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.7 PROJECT REORGANIZATION - DECEMBER 1969 [Historical Information Only] This section describes a reorganization in project management, which was implemented by Westinghouse in December 1969. Westinghouse formed a wholly-owned subsidiary corporation, called WEDCO Corporation, to perform certain functions at the Indian Point site of Con Edison. Westinghouse remained the prime contractor and continued to exercise overall control and to have full responsibility for the Indian Point 2 project. WEDCO performed, under Westinghouse, project management, engineering, quality assurance, construction, and procurement functions for Indian Point Unit 2. These functions were previously carried out by Westinghouse or United Engineers and Constructors (UE&C). The entire Westinghouse senior management organization, which prior to the advent of WEDCO, was responsible for the Westinghouse effort at Indian Point Unit 2, remained responsible. All other personnel within Westinghouse senior management who, prior to WEDCO, carried any responsibility in any area for Indian Point Unit 2 continued to carry those responsibilities, regardless of the formation of WEDCO or changes in title or designation. Furthermore, WEDCO had behind it the full organization and strength of the Westinghouse Electric Corporation: Westinghouse engineering, legal, and other personnel continued to be available for the project. The functional relationships among Con Edison, Westinghouse, WEDCO, and UE&C are shown in Historical Figure 1.7-1. Westinghouse WEDCO-UE&C Relationship. Westinghouse retained UE&C as its architect-engineer-constructor to perform certain work and services in connection with the plant. Initially, UE&C performed services within its scope in the following areas:

1. Design and Engineering

2. Procurement

3. Construction Management and Construction

4. Quality Assurance (including Home Office Quality Control Engineering, Vendor Surveillance and Onsite Quality Control)

Westinghouse removed items (2) and (3) from the scope of work to be performed by UE&C and assigned these functions to WEDCO. In these areas, however, UE&C provided qualified personnel to assist in effectuating the transition of work to Westinghouse and WEDCO. UE&C continued to have responsibility for all of the design and engineering functions and all of the quality assurance functions, including home office quality control engineering, vendor surveillance and onsite quality control, for which it had responsibility prior to the advent of WEDCO. UE&C continued to have direct corporate responsibility to Westinghouse for all of the work within its scope. Chapter 1, Page 29 of 71 Revision 22, 2010

IP2 FSAR UPDATE In its organizational structure, WEDCO exercised a high level quality and engineering reliability function. This function included the activities previously performed by the Nuclear Power Service Staff Resident Quality Assurance Engineer, and in addition included the centralization and overall management for quality assurance activities previously performed by various organizations. This function was carried out by a Reliability Manager who was located at the site. The Reliability Manager was responsible for surveillance visits to selected shops or suppliers. This function was previously delegated to the Westinghouse Nuclear Power Services Group. In addition, the Reliability Manager continually audited the quality assurance efforts of UE&C. In effect, a new reliability management function over and above those previously set forth was established while all existing organizational functions and responsibilities for quality assurance were maintained. The quality control functions previously performed at various Westinghouse organizational levels continued to be performed. At the Westinghouse headquarters level, the staff quality assurance audit team reviewed periodically the quality control program for Indian Point Unit 2 as it had done in the past. At the Westinghouse PWR Systems Division level, the quality control functions performed by that division for the nuclear steam supply system continued as before. Con Edison. The project reorganization described did not in any way alter the ultimate responsibility of Con Edison for the quality assurance program. There was no basic change in the Con Edison program. However, the following minor procedural changes were made in view of the existence of WEDCO:

1. Con Edison's monitoring function included monitoring the activities of WEDCO.

2. Con Edison forwarded the United States Testing Company quality assurance reports to Westinghouse and/or WEDCO.

3. Con Edison contacted Westinghouse and/or WEDCO for necessary corrective action.

1.7 FIGURES Figure No. Title Figure 1.7-1 Functional Relationships [Historical] 1.8 PROJECT REORGANIZATION - MARCH 1970 [Historical Information Only] This section describes a change, which was implemented by Westinghouse in the spring of 1970 in the project organization. The changes made in December 1969 (see Section 1.7) involved the creation of WEDCO and the delegation to WEDCO by Westinghouse of certain functions at the Indian Point site previously carried out by Westinghouse or United Engineers and Constructors (UE&C). Following the December 1969 reorganization, UE&C retained the following functions within its scope of work as architect-engineer-constructor: Chapter 1, Page 30 of 71 Revision 22, 2010

IP2 FSAR UPDATE

1. Design and Engineering.

2. Quality Assurance (including Home Office Quality Control Engineering, Vendor Surveillance, and Onsite Quality Control).

The change consisted of the removal of the vendor surveillance and onsite quality control portions of item (2) from the scope of work to be performed by UE&C, and the assigning of these functions to WEDCO. There was little change of personnel involved in the transfer of the onsite quality control function. WEDCO employed a Manager of Vendor Surveillance and other personnel for this work. The transition in this respect was gradual. New personnel were phased in and the UE&C personnel were used during the transition period to ensure continuity of the surveillance program. The transfer was made on a purchase-order-by-purchase-order basis, with UE&C personnel working with new personnel in performing the surveillance during the transition. To assure that a level of quality assurance review was not lost, the organization of the WEDCO reliability group was structured to provide for an independent, internal audit of the two quality assurance functions transferred to WEDCO. The Vendor Surveillance Group and Onsite Quality Control Group each reported directly to the Reliability Manager. The activities of both the Vendor Surveillance Group and the Onsite Quality Control Group were audited by a Systems Reliability Group. The Systems Reliability Group reported directly to the Reliability Manager to ensure its functional independence. Historical Figure 1.8-1 shows these organizational relationships in chart form. 1.8 FIGURES Figure No. Title Figure 1.8-1 Organization Chart WEDCO Reliability Group [Historical] 1.9 SUPPLEMENTS AND REVISIONS TO ORIGINAL FSAR 1.9.1 Supplements Supplement 1 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission as contained in two letters. The first letter from Peter A. Morris, Director of the Division of Reactor Licensing, on March 5, 1969, to Mr. Donham Crawford of Con Edison, requested additional information on the medical plans and facilities at Indian Point. The questions and responses are found following Tab I of Volume 5 of the original FSAR. These responses were incorporated into Section 11.2.5 of the original FSAR as page changes. The responses to the questions in Volume 5 indicate where the specific answer may be found in the page change. The second letter to Arthur N. Anderson of Con Edison from Peter A. Morris, dated August 4, 1969, requested additional information on Chapters 1, 2, 3, 4, 5, 6, 7, 8, 11, 12, and 14 of the original FSAR. Supplement 1 responded to several of the questions in the second letter found behind Tab II of Volume 5 of the original FSAR. The responses consisted of questions and Chapter 1, Page 31 of 71 Revision 22, 2010

IP2 FSAR UPDATE answers given in Volume 5 of the original FSAR and also of page changes to the original text of the FSAR in some instances. Supplement 2 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission and page changes to the report. The questions were contained in a letter to Arthur N. Anderson of Con Edison from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR in the proper order behind Tab II. Page changes for the FSAR were included with Supplement No. 2. Supplement 3 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission and page changes to the report. The questions were contained in a letter to Arthur N. Anderson of Con Edison from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR in the proper order behind Tab II. This supplement responded to several questions concerning Chapters 1, 4, 5, 7, 8, and 11 of the report. Supplement 4 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission and page changes to the report. The questions were contained in a letter to Arthur N. Anderson of Con Edison from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR in the proper order behind Tab II. Also included with this supplement was a description of the project reorganization within Westinghouse. This supplement also responded to several questions concerning Chapters 4, 5, 7, 11, and 14 of the report. Supplement 5 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission. The questions were contained in a letter to Arthur N. Anderson of Con Edison from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969, and a letter to William J. Cahill, Jr., of Con Edison from Peter A. Morris dated November 13, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR in the proper order behind Tab II. The supplement responded to several questions concerning Chapters 1, 4, 6, 11, 12, and 14 of the report. Supplement 6 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission. The questions were contained in a letter to Arthur N. Anderson of Con Edison from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969, and a letter to William J. Cahill, Jr., of Con Edison from Peter A. Morris dated November 13, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR in the proper order behind Tab II. The supplement responded to several questions concerning Chapters 1, 3, 4, 6, 9, and 14 of the report. Also included with this supplement was the Indian Point Unit 2 Containment Design Report. Supplement 7 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission and page changes to the report. The questions were contained in a letter to Arthur N. Anderson of Con Edison, from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969, and a letter to William J. Cahill, Jr., of Con Edison, from Peter A. Morris, dated November 13, 1969. This supplement responded to several questions concerning Chapters 4, 5, 6, 9, 13, and 14 of the report. Chapter 1, Page 32 of 71 Revision 22, 2010

IP2 FSAR UPDATE Supplement 8 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission and page changes to the report. The questions were contained in a letter to Arthur N. Anderson of Con Edison, from Peter A. Morris, Director of the Division of Reactor Licensing, dated August 4, 1969, and a letter to William J. Cahill, Jr., of Con Edison, from Peter A. Morris, dated November 13, 1969. The responses consisted of questions and answers added to Volume 5 of the original FSAR. The supplement responded to questions concerning Chapters 4, 6, 7, and 13 of the report. Supplement 9, 10, 12, 14, 20 and 21 to the Indian Point Unit 2 Final Safety Analysis Report consisted of corrections and additional information for the original FSAR in the form of page changes. Supplement No. 11 to the Indian Point Unit 2 Final Safety Analysis Report provided the proposed Technical Specifications for operation of the facility in accordance with the rules of practice, 10 CFR 50.36. Supplement 13 to the Indian Point Unit 2 Final Safety Analysis Report consisted of responses to questions from the Atomic Energy Commission contained in a letter from Peter A. Morris, Director of the Division of Reactor Licensing, on July 24, 1970, to William J. Cahill, Jr., of Con Edison. The letter requested additional information on Chapters 1, 4, 7, 8, 12, and 14 of the original FSAR. The responses consisted of questions and answers given in Volume 5 of the FSAR and also of page changes to the original text of the FSAR in some instances. Supplement 15 to the original Final Safety Analysis Report consisted of correction pages that updated certain areas where final design parameters were available and where design modifications had resulted from AEC review. In addition, a cross-reference index was submitted for each chapter of the FSAR where required. The index referenced the responses to questions in Volumes 5 and 6 where additional information could be found concerning specific sections. The proposed Technical Specifications were reissued in their entirety with this supplement. This issue superseded the specifications submitted in Supplement 11. Supplement 18 to the original Final Safety Analysis Report consisted of the relocation of information from the site Custom Technical Specifications into the UFSAR for items and topics that were no longer found in the Improved Technical Specifications. It also updated references to the new Technical Specification sections, to information relocated from the Technical Specifications into the Off Site Dose Calculation Manual (ODCM) and added cross references to the new Technical Requirements Manual (TRM). Supplement 19 to the original Final Safety Analysis Report consisted of corrections and additional information for the original FSAR in the form of changes to reflect several plant modifications, changes to reflect 10 CFR 100.11, the new fuel design and new core design for Cycle 17 and Cycle 16 Core Reload Design, the permanent increase in Tave to 565oF, and the approved alternate source term fuel handling accidents (FHB & VC) which take no credit for charcoal filtration. Changes were also included from NRC approved projects, including Appendix K Power Uprate [1.4% Power Uprate] with the re-analysis of some of the Chapter 14 accidents to account for the 1.4% power uprate, re-analysis of the Loss of Electrical Load transients and LONE/LOOP transients, and the re-analysis of the Feedwater System Malfunction with a step increase of 120% of nominal feedwater flow to one steam generator, and to reflect the approved Stretch Power Uprate to 3216 MWt. Chapter 1, Page 33 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.9.2 Revisions Pursuant to 10 CFR 50.71(e), Con Edison submitted an updated Final Safety Analysis Report for Indian Point Unit 2 on July 22, 1982, reflecting changes made up to a maximum of 6 months prior to the submittal date. In addition, the following revisions to the updated Final Safety Analysis Report have been submitted to date: Revision 1, July 1983 Revision 2, July 1984 Revision 3, July 1985 Revision 4 July 1986 Revision 5, June 1987 Revision 6, June 1988 Revision 7, June 1989 Revision 8, June 1990 Revision 9, June 1991 Revision 10, June 1992 Revision 11, June 1993 Revision 12, June 1994 Revision 13, December 1995 Revision 14, December 1997 Revision 15, December 1999 Revision 16, July 2001 Revision 1 May 2003 Revision 18, October 2003 Revision 19, May 2005 Revision 20, November 2006 Revision 21, October 2008 Revision 22, October 2010 Revision 2, in addition to reflecting required changes, incorporated a major editorial effort to standardize the format of the updated FSAR and to correct typographical, grammatical, and syntax errors, so that the material is presented in a more uniform and clear manner. The changes of technical content and some major editorial changes were marked in the margins with the numeral 2. The majority of editorial changes were minor and were not marked individually. A changed page, however, was indicated by the label Revision 2 at the lower right hand corner. 1.10 QUALITY ASSURANCE PROGRAM 1.10.1 General Entergys Quality Assurance Program (QAP) for Indian Point Unit 2 is in accordance with the quality assurance requirements of 10 CFR 50 Appendix B. The QAP is described in a Quality Assurance Program Manual, which satisfies the criteria of Appendix B. Changes to the program description are submitted to the NRC in accordance with the provisions of 10 CFR 50.54(a)(3). 1.10.2 Scope The Quality Assurance Program provides control for activities affecting the quality of structures, systems, and components of the plant and their operation to the extent consistent with their Chapter 1, Page 34 of 71 Revision 22, 2010

IP2 FSAR UPDATE importance to safety. Those structures, systems, and components of the plant that prevent or mitigate the consequences of postulated accidents that could cause undue risk to the health and safety of the public are designated "Class A" as described in the Quality Assurance Program. All items and activities affecting safety addressed in Regulatory Guide 1.29 Seismic Design Classification revision 3, September 1978, are also governed by the Quality Assurance Program. A list of Class A items is maintained. Elements of the Quality Assurance Program are also applicable to activities and items affecting safety as defined in Licensing commitments.2 It is recognized that not every portion of each of the listed systems and components affect the safety related function. Therefore, allowance is made for subcomponents of systems to be declassified. When such is the case, the agreement is appropriately documented identifying the parts or subcomponents concerned and showing appropriate concurrences. 1.10.3 Organization And Responsibilities The major organizations or groups participating in the Quality Assurance Program are: Nuclear Quality Assurance and Oversight, Nuclear Power, Nuclear Power Engineering, and the Safety Review Committee. The duties and responsibilities of the individuals participating in the Quality Assurance Program are described in procedures, or manuals. REFERENCES FOR SECTION 1.10

1. Deleted

2. Letter from John D. OToole, Con Edison, to Director of Nuclear Reactor Regulation, NRC,

Subject:

Response to NRC letter of September 23, 1980 to Mr. Zarakas requesting information on the Quality Assurance Program for Indian Point Unit 2 dated March 11, 1981. TABLE 1.10-1 DELETED 1.11 DESIGN CRITERIA FOR STRUCTURES AND COMPONENTS 1.11.1 Definition Of Seismic Design Classifications All structures and components are classified as seismic Class I, Class II, or Class III as recommended in:

1. TID-7024, "Nuclear Reactors and Earthquakes," August 1963 and,

2. G. W. Housner, "Design of Nuclear Power Reactors Against Earth-quakes,"

Proceedings of the Second World Conference on Earthquake Engineering, Volume I, Japan, 1960, Pages 133, 134 and 137. Class I Seismic Class I is defined as those structures and components including instruments and controls whose failure might cause or increase the severity of a loss-of-coolant accident or result in an uncontrolled release of radioactivity causing more than 10 rem to the thyroid or 10 Chapter 1, Page 35 of 71 Revision 22, 2010

IP2 FSAR UPDATE rem whole body to the average adult beyond the nearest site boundary. Also included are those structures and components vital to safe shutdown and isolation of the reactor. Class II Class II is defined as those structures and components, which are important to reactor operation but not essential to safe shutdown and isolation of the reactor and whose failure could result in the release of radioactivity causing more than 1.0 rem to the thyroid or 0.5 rem whole body dose to the average adult beyond the nearest site boundary. Class III Class III is defined as those structures and components, which are not directly related to reactor operation or containment. In Indian Point Unit 2, the only portions of the plant, which are not seismic Class I and which might carry substantial radioactivity because of required safeguards operation or requirements for safe shutdown and isolation of the reactor are portions of the chemical and volume control system and waste disposal system. The specific components in the chemical and volume control system are the volume control tank, holdup tank, and the concentrates holding tank with associated piping, valves and supports. These components are all seismic Class I. In addition, the design of the system tanks and their location were based upon the commitment that a vessel rupture would not cause doses in excess of 10 CFR 20 limits at the exclusion radius. The specific components in the waste disposal system are the gas decay tanks with their associated piping, valves and supports. These components are all seismic Class I. In addition, the gas decay tanks of the waste disposal system have been designed such that the failure of any tank will not exceed 10 CFR 20 doses at the exclusion radius. The analysis showing that the rupture of the volume control tank or a gas decay tank does not exceed the special dose limits selected for Indian Point Unit 2 is found in Section 14.2.3 of the FSAR. Those components of the chemical and volume control system that are not seismic Class I are as follows: batching tank, monitor tanks, monitor tank pumps, chemical mixing tank, and resin fill tank. In addition, the boric acid evaporator and the condensate demineralizer are not seismic Class I. Those components of the waste disposal system, which are not seismic Class I include: waste condensate tank and pumps. Failure of these components will not result in offsite doses in excess of 10 CFR 20 limits at the site exclusion radius. All components, systems, and structures classified as seismic Class I are designed in accordance with the following criteria:

1. Primary steady state stresses, when combined with the seismic stress resulting from the response to a ground acceleration of 0.05g acting in the vertical and 0.10g acting in the horizontal planes simultaneously, are maintained within the allowable stress limits accepted as good practice and, where applicable set forth Chapter 1, Page 36 of 71 Revision 22, 2010

IP2 FSAR UPDATE in the appropriate design standards, e.g., ASME Boiler and Pressure Vessel Code, USAS B31.1 Code for Pressure Piping, ACI 318 Building Code Requirements for Reinforced Concrete, and AISC Specifications for the Design and Erection of Structural Steel for Buildings.

2. Primary steady state stresses when combined with the seismic stress resulting from the response to a ground acceleration of 0.10g acting in the vertical and 0.15g acting in the horizontal planes simultaneously, are limited so that the function of the component, system or structure shall not be impaired as to prevent a safe and orderly shutdown of the plant.

All Class II structures and components are designed on the basis of a static analysis for a ground acceleration of 0.05g acting in the vertical and 0.10g acting in the horizontal directions simultaneously. The structural design of all Class III structures meets the requirements of the applicable building code, which is the "State Building Construction Code" State of New York, 1961. This code does not reference the Uniform Building Code. The Original Steam Generator Storage Facility (OSGSF) has been constructed for the storage of the original steam generators. The OSGSF is a seismic Class III structure, designed in accordance with the requirements of the State of New York Official Compilation of Codes, Rules and Regulations, Title 9, Subtitle S, 1995 edition, copyright 1999, and the American Concrete Institute (ACI) 318, Building Code Requirements for Structural Concrete, 1999. Table 1.11-1 gives the damping factors used in the design of components and structures. The design of seismic Class I structures and components utilizes the "response spectrum" approach in the analysis of the dynamic loads imparted by the earthquake. The analysis is based upon the response spectra shown on Figures 1.11-1 and 1.11-2. The following method of analysis is applied to seismic Class I structures and components, including instrumentation:

1. The natural period of vibration of the structure or component is determined.

2. The response acceleration of the component to the seismic motion is taken from the response spectrum curve at the appropriate period.

3. Stresses and deflections resulting from the combined influence of normal loads and the seismic load due to the design earthquake (0.05g acting in the vertical and 0.10g acting in the horizontal planes simultaneously) are calculated and checked against the limits imposed by the design standard.

4. Stresses and deflections resulting from the combined influence of normal loads and the seismic loads due to the maximum potential earthquake (0.10g acting in the vertical and 0.15g acting in the horizontal planes simultaneously) are calculated and checked to verify that deflections do not cause loss of function and that stresses do not produce rupture.

Chapter 1, Page 37 of 71 Revision 22, 2010

IP2 FSAR UPDATE Where the vibrator system is of a highly complex geometric shape, such as piping systems, the maximum response from the response curve with the appropriate damping factor is selected. By using this conservative value and demonstrating that the stresses are satisfactory, it becomes unnecessary to perform any further analysis to determine the natural periods of the system. For a further discussion of the models and methods used for the seismic Class I design of structures, equipment, piping, instrumentation and controls, see Section 1.11.4. 1.11.2 Classification Of Particular Structures And Equipment Examples of particular structure and equipment classifications are given below. These classifications are not intended to be all-inclusive. Item Class Buildings and Structures Containment (including all penetrations and airlocks, the concrete shield, the liner, and the interior structures) I Spent fuel pit I Control Building I Diesel Generator Building I Intake structure (to the extent that water is always available to the service water pumps) I Service water screenwell I Primary Auxiliary Building I Turbine Building III Buildings containing conventional facilities Such as the Maintenance and Outage Building III Original Steam Generator Storage Facility III Equipment, Piping, and Supports [Note - Class I components (equipment, piping, instrumentation, etc.) located in or supported on a Class II structure are protected from earthquake damage or are backed up by other Class I components located in or supported by a Class I structure.] Reactor control and protection system I Radiation monitoring system I Process instrumentation and controls I Chapter 1, Page 38 of 71 Revision 22, 2010

IP2 FSAR UPDATE Reactor Vessel and its supports I Vessel internals Fuel assemblies Rod cluster control assemblies and drive mechanisms Supporting and positioning members Incore instrumentation structure Reactor coolant system I Piping and valves (including safety and relief valves) Steam generators Pressurizer Reactor coolant pumps Supporting and positioning members Main Steam system, up to and including the isolation valve I Engineered safety features I Safety injection system (including safety injection and residual heat removal pumps, refueling water storage tank, accumulator tanks, residual heat removal heat exchangers and connecting piping and valving) Containment spray system (including spray pumps, spray headers, and connecting piping and valving) Containment air recirculation cooling system (including fans, coolers, ducts, valves, and demisters) Auxiliary building ventilation system I Condensate storage tanks I Pressurizer relief tank II Residual heat removal loop I Containment penetration and weld channel pressurization system I Component cooling loop I Instrument air system (essential sections) I Isolation valve seal water system I Sampling system II Spent fuel pit cooling loop II Fuel transfer tube I Chapter 1, Page 39 of 71 Revision 22, 2010

IP2 FSAR UPDATE Emergency power supply system I Diesel generators and fuel oil storage tank DC power supply system Power distribution lines to equipment required for transformers and switchgear supplying the engineered safety features Control panel boards Motor control centers Control Equipment, facilities and lines necessary for the above seismic Class I items I Waste disposal system I Chemical drain tank Waste holdup tanks Gas decay tanks Reactor coolant drain tank Compressors Waste holdup tank pumps Interconnecting waste gas piping Waste disposal system II or III All elements not listed as seismic Class I Containment crane I Manipulator and other cranes III Conventional equipment, tanks and piping, other than Classes I and II III Auxiliary boiler feed and service water pumps and piping I The chemical and volume control system is considered seismic Class I except for the components listed below: Batching tank II Monitor tanks III Monitor tank pumps III Chemical mixing tank II Resin fill tank III Chapter 1, Page 40 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.11.3 Design Criteria For Seismic Class I Structures And Equipment The criteria for functional adequacy of structures, equipment, piping, instrumentation, and controls follow. No loss of function implies that rotating equipment will not freeze, pressure vessels will not rupture, supports will not collapse under the load, systems required to be leaktight will remain leaktight, and components required to respond actively (such as valves and relays) will respond actively. The criteria for functional adequacy of the structures state stresses will not exceed yield when subjected to a 0.15g ground acceleration. The manner in which these criteria have been met is by limiting stresses in seismic Class I structures to meet the above criteria. For all seismic Class I piping and their supports, the criteria for functional adequacy and the manner in which the criteria are met are the following: with a ground acceleration of 0.15g horizontal, the spectral acceleration corresponding to the maximum point on the 0.5-percent critical damping response curve was used to calculate an equivalent static force imparted to the pipe at its support points. This resulted in a seismic design load approximately equal to 0.6W horizontally and 0.4W vertically taken simultaneously, where W is the weight of the pipe including static forces. The sum of the resulting additional stress plus the normal stresses was limited to 1.2 times the B31.1 code allowable. The stresses in the pipe supports and hangers were likewise limited to 1.2 times the B31.1 code allowable. Since all the buildings containing seismic Class I piping are essentially rigid structures, no amplification is expected. For seismic Class I equipment and tanks the same method was used to arrive at an equivalent static force. In each case, the total of seismic and normal stresses was limited to the applicable code allowable. The refueling water storage tank and condensate storage tank were designed in accordance with the stress limitations of American Water Works Association Standard D100.All components of the reactor coolant system and associated systems are designed to the standards of the applicable ASME code or USAS code. The loading combinations, which are employed in the design of seismic Class I components of these systems, i.e., vessels, piping, supports, vessel internals and other applicable components, are given in Table 1.11-2. Table 1.11-2 also indicates the stress limits, which are used in the design of the listed equipment for the various loading combinations. The original design criteria given above and in Table 1.11-2 have been modified in certain instances in accordance with NRC guidance given in References 3 and 4. Generic Letter 87-11 allows for the elimination of pipe whip restraints and jet impingement shields, which were installed to mitigate the effects of arbitrary intermediate pipe ruptures, provided certain criteria are met. These design criteria have also been modified in certain instances by the application of "leak before break" technology, as discussed in Section 4.1.2.4. To be able to perform their function, i.e., allow core shutdown and cooling the reactor vessel, internals must satisfy deformation limits, which are more restrictive than the stress limits shown in Table 1.11-2. For this reason the reactor vessel internals are treated separately. Chapter 1, Page 41 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.11.3.1 Piping, Vessels, and Supports The reasoning for selection of the load combinations and stress limits given in Table 1.11-2 is as follows. For the design earthquake, the nuclear steam supply system is designed to be capable of continued safe operation, i.e., for the combination of normal loads and design earthquake loading. Critical equipment and supports needed for this purpose are required to operate within normal design limits as shown in line 2 of Table 1.11-2. In the case of the maximum potential earthquake, it is only necessary to ensure that critical components do not lose their capability to perform their safety function, i.e., shut the plant down and maintain it in a safe condition. This capability is ensured by maintaining the stress limits as shown in line 3 of Table 1.11-2. No rupture of a seismic Class I pipe can be caused by the occurrence of the maximum potential earthquake. With respect to the seismic design of the piping supports, relative displacement between anchor points has been considered in the seismic analysis of the main steam lines, where largest relative displacements are expected. Analysis indicates that the stress at the highest seismically stressed point is affected by less than 10-percent when relative anchor displacements are considered. The seismic supports installed have been verified to agree with the design location and therefore the locations used in the analyses. Careful design and thorough quality control during manufacture and construction and periodic inspection during plant life, ensures that the independent occurrence of a reactor coolant pipe rupture is extremely remote. If it is assumed that a reactor coolant pipe ruptures, the stresses in the unbroken leg will be as noted in line 4 of Table 1.11-2. 1.11.3.2 Reactor Vessel Internals 1.11.3.2.1 Design Criteria for Normal Operation The internals and core are designed for normal operating conditions and subjected to loads of mechanical, hydraulic, and thermal origin. The response of the structure under the design earthquake is included in this category. The stress criteria established in Section III of the ASME Boiler and Pressure Vessel Code, Article 4, have been adopted as a guide for the design of the internals and core with exception of those fabrication techniques and materials, which are not covered by the code, such as the fuel rod cladding. Seismic stresses are combined in the most conservative way and are considered primary stresses. The members are designed under the basic principles of: (1) maintaining distortions within acceptable limits, (2) keeping the stress levels within acceptable limits, and (3) prevention of fatigue failures. 1.11.3.2.2 Design Criteria for Abnormal Operation The abnormal design condition assumes blowdown effects due to a reactor coolant pipe double-ended break. For this condition the criteria for acceptability are that the reactor be capable of safe shutdown and that the engineered safety features are able to operate as designed. Consequently, the limitations established on the internals for these types of loads are concerned principally with Chapter 1, Page 42 of 71 Revision 22, 2010

IP2 FSAR UPDATE the maximum allowable deflections. The deflection limits for critical internals structures under abnormal operation are presented in Table 14.3-14. 1.11.3.3 Reactor Vessel The criteria for movement of the reactor vessel, under the worst combination of loads, i.e., normal plus the maximum potential earthquake or normal plus reactor coolant pipe rupture loads, ensure that the radial movement of the reactor vessel will not exceed the clearance between the reactor coolant piping and the surrounding concrete. The relative motions between reactor coolant system components are controlled by the structures, which are used to support the reactor vessel, the steam generators, the pressurizer and the reactor coolant pumps. The maximum movement of the reactor vessel under the worst combination of loads, i.e., normal plus maximum potential earthquake or normal plus reactor coolant pipe rupture loads comprises an end deflection on the safety injection piping, which is small even in comparison with that resulting from thermal growth during plant heatup, and is well within the flexibility of the design of the piping system. The supports are designed to limit the stresses in the pipe to the stress limits given in Table 1.11-2. 1.11.4 Models And Methods For Seismic Class I Design The variety of design problems associated with the seismic analysis of all Class I structures, systems and equipment were approached by various methods. For the design of the reactor, recirculating pumps, and Class I piping an amplification factor of 4.0 was used with respect to ground motion of 0.15g. This amplification factor was based on the maximum for a one-half percent damping of the ground response spectrum. The fundamental frequency of the reactor building internal structure is approximately 17 cycles/sec. As can be seen from Figure 1.11-2 for this frequency level, no significant building amplification of the ground response is encountered. With the exception of the containment, primary auxiliary building, and electrical cable tunnel, no dynamic analyses were performed on Indian Point Unit 2 structures, hence no mathematical models were developed. The following methods were used in the seismic design of Class I structures. 1.11.4.1 Containment Building See Sections 2.0, 3.0, and 4.0 of the Containment Design Report for Indian Point Unit 2 containment building structures and components. 1.11.4.1.1 Steel In the design of the steel, 100-percent of the dead load and 50-percent of the live load were considered. The peak of the response curve for 0.15g ground acceleration and 1.0-percent critical damping was used to obtain the seismic forces, which were distributed by the method described in the Containment Design Report and resisted by the bracing. The 1.0-percent critical damping is conservative since the structure is shop welded and field bolted to the Chapter 1, Page 43 of 71 Revision 22, 2010

IP2 FSAR UPDATE columns. The actual critical damping value would be between 1.0-percent (welded) and 2.5-percent (bolted). A one-third increase over working stress was allowed in the design of the bracing. 1.11.4.1.2 Concrete In the design of the concrete, 100-percent of the dead load and 50-percent of the live load were considered. The Modified Rayleigh Method was used to calculate the natural period and the base shear was distributed by the same method described in the Containment Design Report. The forces determined from the response curve for a 0.15g ground acceleration with 5-percent critical damping were applied at the node points where the masses were lumped for the Rayleigh approach. These loads were resisted by the vertical walls, which acted as shear walls, and horizontal reinforcing, which resisted the moment. The Ultimate Strength Design method of ACI 318-63 was used for the design and construction of the containment building. 1.11.4.2 Control Building The dead load and equipment loads were considered. The period was determined from the formula T = 0.1 n, where n = number of stories (Design of Multistory Reinforced Concrete Building for Earthquake Motions by N. M. Newmark, et. al.). The response curve for 0.15g ground acceleration with 2.5-percent critical damping was used to determine the base shear. This base shear was distributed at the floor levels by the same method described in the Containment Design Report and resisted by a rigid frame structure with a one-third increase on allowable working stresses. The design was controlled by a deflection limitation due to the adjacent Unit 1 control building. 1.11.4.3 Diesel Generator Building Due to the light weight of the structure, the wind load controlled the design. 1.11.4.4 Fan House One hundred percent of the dead load and 50-percent of the live load were considered. The peak of the response curve for 0.15g ground acceleration with 5-percent critical damping was used for the concrete structure and the corresponding 2.5-percent was used for the steel superstructure. A one-third increase in allowable working stresses was allowed. 1.11.4.5 Boric Acid Evaporator Building One hundred percent of the dead load was considered. For method of design, see fan house. Without allowing a one-third stress increase for seismic design, the controlling factor for reinforcing design was the minimum temperature steel requirements of the ACI-318 Building Code. 1.11.4.6 Intake Structure One hundred percent of the live and dead load were considered. The peak of the response curve for 0.1g (OBE) ground acceleration with 5-percent critical damping was used to obtain the seismic loads. The effect of water sloshing was considered in the earthquake analysis (per TID-7024 "Nuclear Reactors and Earthquakes," Section 6.5). Although DBE was not explicitly considered in the calculation (the seismic forces used in the design shows that DBE is not Chapter 1, Page 44 of 71 Revision 22, 2010

IP2 FSAR UPDATE governing), the controlling factor in the design of the intake structure was the service load with the worst combination being one chamber empty and the adjacent chamber filled with water. 1.11.4.7 Waste Holdup Tank Pit One hundred percent of the dead load and 50-percent of the live load were considered (including the tank dead weight on the roof). The peak of the response curve for 0.15g ground acceleration with 5-percent damping was used to determine the base shear. Using working stress limits for the seismic design, service loads controlled the design of the top slab. The bottom slab and wall of the pit were designed for earthquake loads with stresses limited to yield multiplied by the factors recommended in Section IV-B of the ACI-318-63 "Building Code." Consideration was given to the tanks in the pit when designing the base slab. 1.11.4.8 Spent Fuel Pit The seismic loads, as determined in TID-7024 "Nuclear Reactors and Earthquakes," Section 6.5, were resisted by the reinforced concrete walls and base slab. Working stresses were used except for the moment at the base of the walls where ultimate strength design was considered with stresses limited to fy. The effects of water in the pool are accounted for in this design approach. Ground acceleration of 0.15g was used. In 1990, new high density spent fuel storage racks were installed. Prior to their installation, the spent fuel pit was reanalyzed (Ref. 6). The new racks were also analyzed (Ref. 5 and 6). 1.11.4.9 Electrical Penetration Tunnel The peak of the response curve for 0.15g ground acceleration with 5-percent critical damping was considered using working stress design limits. The load was considered to act at 2/3 L, where L = the height of the tunnel. Temperature of steel considerations controlled the design of the concrete while service loads controlled the structural steel. 1.11.4.10 Pipe Penetration Tunnel One hundred percent of the dead load, plus 50-percent of the live load were considered. The peak of the response curve for 0.15g ground acceleration with 5-percent damping was used to find the shear, which was considered as a concentrated load applied at the top slab of the tunnel. A one-third increase on working stress allowables was used in the design. 1.11.4.11 Electrical Cable Tunnel One hundred percent of the dead load, 50-percent of the surcharge, and 50-percent of live load in the tunnel were considered. The Modified Rayleigh Method was used to determine the natural period and the loads were distributed as described in the Containment Design Report. The response curve for 0.15g ground acceleration with 5-percent critical damping was used. A one-third stress increase was permitted on working stress allowables when considering the effect of seismic loads. 1.11.4.12 Shield Wall The peak of the response curve for 0.15g ground acceleration with 5-percent critical damping was used. The pipe break loads controlled the design. Chapter 1, Page 45 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.11.4.13 Retaining Wall At Equipment Entrance The wall was designed for soil pressure including a 1000 psf surcharge applied during reactor loading. The load combination that includes a seismic factor governs the design. It has been shown that the wall design was adequate. 1.11.4.14 Primary Water Storage Tank and Refueling Water Storage Tank Foundation The seismic loads on the circular wall and center pier were those supplied by the tank manufacturer. The shear force from the earthquake on the water in the tank was applied at 3/4 L above the top slab. The shear force from the earthquake on the tank was applied at L/2 above the top slab, where L = the height of the tank. The horizontal shear force from the earthquake effect on the dead weight of the foundation was determined by using the peak of the response curve for 0.15g ground acceleration with 5-percent critical damping. A triangular distribution was used. The earthquake effect of the backfill was also considered. The load was applied to the walls as the resultant of a triangular pressure distribution. The stresses were limited to working stress design limits. The temperature steel considerations controlled the design of the walls and center pier. 1.11.4.15 Condensate Water Storage Tank Foundation The seismic loads on the spread footing foundation were those supplied by the tank manufacturer. The shear forces from the earthquake on the water in the tank were applied at 3/4 L above the footing, where L = the height of the tank. The shear force from the earthquake on the tank was applied at L/2 above the top of the footing. The stresses were limited to working stress design limits. A multi degree-of-freedom modal analysis was performed on all Class I structures for Indian Point Unit 3. The results indicated that all structures except the containment structure were rigid. The only significant differences between the structural design of Units 2 and 3 seismic Class I buildings are the control building and the steel structural portion of the primary auxiliary building for Indian Point Unit 2, which are flexible steel structures. On Unit 3 they are rigid concrete structures. All seismic Class I structures on Indian Point Unit 2 except control building and containment shell are rigid and move with zero period ground acceleration. However, the design of all seismic Class I structures on Unit 2 were standardized and based on the peak acceleration of the ground response spectrum, which is extremely conservative for rigid structures. In the preceding designs, limits have been placed on stresses to ensure that all structures will respond elastically to the earthquake. If for some reason inelastic response were to occur, the period of the structure would be expected to increase. Since the majority of the structures were designed for the peak of the response curve, the effect of any change in period would be to decrease the coefficient of spectral acceleration and thus lower all shears and moments. Mathematical models were not used for seismic design of instrumentation. Ability to withstand the seismic condition is determined by actual vibration type testing of typical instrumentation equipment under simulated seismic accelerations to demonstrate its ability to perform its functions. The seismic testing is reported in Westinghouse report WCAP-7397-L (Reference 1). The locations of protection and safeguards control and electrical equipment in Indian Point Unit 2 have been identified. The most adverse location, seismically, is the control building floor at Chapter 1, Page 46 of 71 Revision 22, 2010

IP2 FSAR UPDATE elevation 53-ft, which supports the nuclear instrumentation, radiation monitoring, process instrumentation, and safeguards logic racks. Dynamic analyses of this building for the plant design basis earthquake of 0.15g show that the significant horizontal and vertical accelerations of this floor are within the specified low seismic test envelopes given in WCAP-7397-L (Reference 1). Seismic analysis of Class I equipment including heat exchangers, pumps, tanks, valves, motors, and electrical equipment components was performed using one of the following four methods:

1. Equipment, which is rigid and rigidly attached to its support structure, was analyzed for a "g" loading equal to the peak acceleration of the supporting structure at the appropriate elevation.

2. Equipment, which is not rigid and therefore a potential for response to the support motion exists, was analyzed for the peak of the floor response curve for appropriate damping values.

3. In some instances nonrigid equipment was analyzed using a multi degree-of-freedom modal analysis. All contributing modes were considered. In addition, a sufficient number of masses was included in the mathematical models to ensure that coupling effects of members within the component were properly considered.

The results of these analyses indicated that the models contained more masses than necessary, and that future analyses of comparable equipment could be considerably simplified by considering fewer masses. The method of dynamic analysis used a proprietary computer code called WESTDYN. This code used as input the inertia values, member sectional properties, elastic characteristics, support and restraint data characteristics, and the appropriate seismic response spectrum. Both horizontal and vertical components of the seismic response spectrum were applied simultaneously. The modal participation factors were combined with the mode shapes and the appropriate seismic response spectra to obtain the structural response for each mode. The internal forces and moments were computed for each mode from which the modal stresses are determined. The stresses were then summed using the root mean square method.

4. Type testing of selected electrical equipment has been conducted to demonstrate seismic design adequacy as described in WCAP-7397-L (Reference 1).

For the analysis of equipment to resist the vertical seismic component, two-thirds of the horizontal response spectrum curves were used to determine the acceleration appropriate to the vertical frequency. Engineered safeguards tanks, e.g., boric acid, accumulator and surge tanks, were analyzed using method 3 above for combined horizontal and vertical seismic excitation occurring simultaneously and in conjunction with normal loads without exceeding allowable stresses. Hydrodynamic analyses of these tanks have been performed using the methods described in Chapter 6 of the "U.S. Atomic Energy Commission - TID 7024." The stresses for these components due to the above-mentioned load combinations were found to be within allowable limits. Heat exchangers associated with the engineered safeguard systems, e.g., component cooling and residual heat removal, were analyzed using method 3 above, and the results show that stresses and deflections are within allowable limits. Chapter 1, Page 47 of 71 Revision 22, 2010

IP2 FSAR UPDATE Selected critical engineered safeguards valves were analyzed using method 3 above and the results indicated that their fundamental natural frequency was sufficiently separated from the building frequency. The results further indicated that the total stress, considering all modes, was far below the allowable stress limit. Appendages, such as motors attached to motor-operated valves, were included in the mathematical models. 1.11.4.16 Class I Piping Systems Class I piping systems were designed and analyzed as described in the succeeding paragraphs. However, in an attempt to correlate the simplified method of analysis suggested by the AEC for the H. B. Robinson Nuclear Generating Station, the following discussion is presented: If no dynamic analysis is performed on Class I piping systems, these systems for H. B. Robinson plant were to be checked to determine whether the results conform to the following formula: 1.3* K Ss + Sn 1.8 Sa [Note - *The 1.3 factor was recommended by the AEC to represent the contributions of higher modes above the fundamental mode. Detailed dynamic analyses performed on Indian Point Unit 2, and described later, indicate that where significant stresses exist in piping systems, a more realistic modal contribution factor would be 1.1. However, for the present discussion we will adhere to the 1.3 factor for additional conservatism.] where: Ss - represents seismic stress including effects of valve motors, from design calculations Sn - represents normal primary and bending stresses for loadings other than seismic, from design calculations 1.8 Sa - equals 1.8 times the allowable stress or yield stress, whichever is higher for code listed materials. K- ratio of peak acceleration of floor response spectra to acceleration used in the piping design The piping design criteria limited the deadweight and seismic stresses to 0.2 Sa. The longitudinal pressure stress is 0.5 Sa. 1.3 K (0.2 Sa) + 0.5 Sa 1.8 Sa Solving, the K-factor becomes: K=5 Chapter 1, Page 48 of 71 Revision 22, 2010

IP2 FSAR UPDATE This factor combined with the 1.3 modal contribution factor gives a combined factor of 6.5, which is more than double the original suggested multiplier of 3. Indian Point Unit 2 conservatively meets the criteria suggested for application on the H. B. Robinson Plant for seismic Class I piping. However, a different and more detailed method of analysis was actually undertaken to illustrate the conservatism of design approach used for Indian Point Unit 2. This approach is described in detail below: It is obviously necessary to use simplifying assumptions when performing initial design of piping systems, including restraints, rather than a dynamic analysis involving a trial and error procedure. Simplified design procedures are not uncommon and often suggested in codes, i.e., USAS B31.1 - Power Piping Code. A complete flexibility analysis involving detailed modeling of Class I piping systems is unnecessary if the conservatism of the simplifying assumptions used in the initial design can be demonstrated. A "third party" review was conducted to establish the adequacy and conservatism of the original design criteria for Class I piping systems as performed by the architect/engineer (United Engineers and Constructors, Inc.) and the seismic restraint supplier (Bergen-Paterson Pipe Support Corp.). The review involved the following steps:

1. Representatives from Westinghouse and United Engineers and Constructors, Inc., visited the Indian Point Unit 2 site and inspected the Class I piping systems.

2. Based upon their best engineering judgment, representative worst-case lines were selected for detailed dynamic analyses.

3. In exercising their engineering judgment, these representatives looked for the following characteristics, which would indicate possible sources of problems.

a. Amplification due to the location and elevation in building.

b. Large concentrated masses such as overhung motor-operated valves, particularly in what appear to be flexible sections of the pipe.

c. Complexity of configuration of the piping system itself such that application of the original design criteria would be difficult.

d. Manual excitation of the pipe by pushing or kicking indicated excessive flexibility either in the pipe excited or the piping attached to it.

4. The results of the dynamic analyses were compared with original design values to determine whether the design approach was conservative. Besides analysis of the reactor coolant loop, portions of the following systems were analyzed:

a. Safety injection.

b. Main steam.

c. Residual heat removal.

d. Service water.

e. Accumulator discharge.

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IP2 FSAR UPDATE

f. Containment spray.

g. Component cooling.

1.11.4.16.1 Design Approach The design and placement of seismic restraints were predicated on the principle of containing the seismic stresses without restricting the free thermal expansion of the piping system. The systems were designed to have sufficient flexibility to prevent the movements from causing failure of piping or anchors from overstress. Two fundamental principles underlie the design approach, namely:

1. The system be designed such that its fundamental natural frequency does not coincide with the exciting frequency.

2. The maximum seismic stresses in piping be less than the USAS B31.1 code allowable value. The seismic stresses were limited to 0.2 S allowable (3000 psi).

This is extremely conservative since the longitudinal pressure stress accounts for approximately 0.5 S allowable leaving a margin of safety of 0.5 S allowable, which is unused. (Note-this is based on a maximum allowable of 1.2 Sa) These fundamental principles should ensure that stresses will be within code allowable stress limits, and that the piping will not go into resonance with the exciting frequency. Tables of recommended maximum spacing of supports, for straight runs of pipe, were developed. The recommended spacing of supports was modified near bends and concentrated masses (i.e. valves) to account for additional weight and flexibility. 1.11.4.16.2 Analysis Approach In order to determine whether the design procedure resulted in an acceptable system, selected worst case Class I piping systems were modeled and a dynamic flexibility analysis performed. A detailed description of the method of analysis is given below. The analysis was performed using a proprietary computer code called WESTDYN. The code uses as input system geometry, inertia values, member sectional properties, elastic characteristics, support and restraint data characteristics, and the appropriate Indian Point seismic floor response spectrum for 0.5-percent critical damping. Both horizontal and vertical components of the seismic response spectrum are applied simultaneously. With this input data, the overall stiffness matrix of the three-dimensional piping system is generated (including translational and rotational stiffnesss). The modal participation factors are computed and combined with the mode shapes and the appropriate seismic response spectra to give the structural response for each mode. Each piping run is modeled as a three-dimensional system, which consists of straight segments, curved segments, and restraints. Straight segments are distinguished from curved segments during data output. The computer code requires that the piping be represented by a discrete mass model. Each mass includes the contribution of both the steel encasement and conveyed fluid. Where valves or other concentrated masses exist in the piping system, they were included in the model. Chapter 1, Page 50 of 71 Revision 22, 2010

IP2 FSAR UPDATE Restraints were included in the model at their proper location. The directionality of the restraints was also considered. The detailed dynamic analyses of selected worst case Class I piping indicated that the method used to design the seismic restraints was conservative. Based on this critical review of the selected worst case systems and the consistent application of the same design procedure to all completely engineered seismic Class I systems, the seismic design of other Class I systems, not analyzed, was deemed adequate. The maximum stresses imposed by the normal loads plus loads associated with the design-basis earthquake (DBE) are below 1.2S, where S is the allowable stress limit obtained from the Power Piping Code - USAS B31.1.0 - 1955. Some of the items of conservatism employed in the seismic design of Class I piping systems for Indian Point Unit 2 were:

1. The maximum longitudinal stress due to seismic excitation was limited to 0.2S rather than the usual 0.7S.

2. The maximum allowable stress was limited to 1.2S. If the combination of normal and DBE loads were considered as a faulted condition, the allowable membrane and bending stresses could be chosen as those corresponding to 20-percent to 40-percent of the material uniform strain at temperature, respectively. This would give more than a factor of 2 margin between the allowable and the maximum actual stresses.

3. A low value of the fraction of critical damping was adopted (0.5-percent). Dr. N.

M. Newmark recommended a value of 2-percent for vital piping at or just below the yield point. This would reduce the maximum amplification of the ground acceleration.

4. The maximum longitudinal stresses due to pressure, deadweight, and seismic loads were presumed to occur at the same cross-section and some point in the cross-section.

Some averaging of the response spectra was performed to smooth out the erratic response of the earthquake's random behavior. At the high frequency end of the spectra, the acceleration levels of the smoothed spectra converge to the values of the unsmoothed spectra. It is therefore concluded that the design procedure used to design seismic Class I restraints for Indian Point Unit 2 is conservative. NRC IE Bulletin (IEB) No. 79-07 was concerned with inadequacies identified in the seismic analysis of certain piping systems at several power reactors. The inadequate treatment of piping loads from earthquakes was attributed to the fact that some piping analysis codes used an algebraic summation of the loads predicted separately by computer code for both the horizontal components and the vertical component of seismic events. In accordance with the IEB, such co-directional loads should not be algebraically added unless certain more complex time-history analyses are performed. The IEB emphasized that to properly account for the effects of earthquakes on systems important to safety, such loads should be combined absolutely or by using techniques such as the sum of the squares. Chapter 1, Page 51 of 71 Revision 22, 2010

IP2 FSAR UPDATE In response to IE Bulletin No. 79-07, eight (8) Indian Point Unit No. 2 lines were reanalyzed using the UE&C-ADLPIPE-2 dynamic seismic computer code. This code utilizes the worst-case two-dimensional evaluation technique and uses the square root of the sum of the squares option for combining both intramodal and intermodal responses. The difference between the newly calculated total pipe stress and the originally calculated total pipe stress is not significant. Even after applying a 1.3 "adjustment" factor to the calculated seismic stress component, the total pipe stress remains below the allowable stress limit. Furthermore, the loads on the pipe supports and equipment nozzles were re-evaluated on the basis of the confirmatory reanalysis and found to be acceptable, as documented in Reference 9. 1.11.4.17 Reactor Coolant System Analysis for Combination Loading of Design-Basis Earthquake and Design-Basis Accident [Historical Information Only] The Indian Point Unit 2 reactor coolant system was not committed to be designed for the combination of the seismic and blowdown loads. However, an analysis for this combination of loadings was performed for the original configuration of the Indian Point Unit 3 reactor coolant system, which was identical to the original Unit 2 configuration. The analysis was performed as outlined below:

1. A lumped mass dynamic mathematical model of the primary coolant loop and support system was developed.

2. This dynamic model was subjected to multiple simultaneous time history hydraulic forcing functions for the blowdown analysis. The double-ended ruptures were located at places of large change in flexibility. Time history response of the total structure to these conditions was computed and reduced to time history stresses.

3. The dynamic model was then subjected to a floor response spectra earthquake analysis.

4. The loads as determined above were used for an evaluation of the stresses along the piping system.

5. The stresses as determined from the basis described above were lower than the allowable stresses calculated by using the approach described in WCAP-5890, Revision 1 (Reference 2) and the following parameters:

a. 20-percent of the uniform strain on the allowable membrane and average strain.

b. 23,100 psi as the at-temperature yield in the axial direction. This value was based on the minimum value of the at-temperature yield in the loop direction as measured with samples from the Unit 2 piping, and increased by 10-percent for the increase in strength in going from the loop to the axial direction. The tensile tests on the Unit 2 piping material at-temperature yielded at a minimum value of 20,900 psi, a maximum of 29,700 psi, and an average of eleven samples of 23,300 psi.

Chapter 1, Page 52 of 71 Revision 22, 2010

IP2 FSAR UPDATE Based on the above analysis, it was concluded that the Unit 2 reactor coolant system can stand the combination of blowdown and seismic loads within acceptable stress limits. In 1989, the NRC approved changes to the design bases with respect to dynamic effects of postulated primary loop pipe ruptures, as discussed in Section 4.1.2.4. In 2000, an analysis of the Unit 2 reactor coolant loop and its component supports, which incorporates the NRC approved changes, was performed with the replacement steam generators and sixteen of the original twenty-four steam generator support frame hydraulic snubbers removed (Reference 11). In line with the older analysis for Unit 3 described above, the Unit 2 analysis of 2000 included the effects of the now controlling feedwater line break at the steam generator nozzle in a similar fashion as described for the Unit 3 analysis. Based on this revised analysis, it was concluded that the Unit 2 reactor coolant system can withstand the combination of blowdown and seismic loads within acceptable stress limits. This 2000 analysis has since been updated to include a power uprate to a core power level of 3216 MWt. Combination of blowdown and seismic loads were not considered in this latest evaluation. 1.11.4.18 Service Water Lines The service water lines consist of two 24-in. diameter carbon steel pipes. They run in a common trench, which is backfilled. Assuming that the ends of a pipe are free to displace vertically but not rotate and that the maximum permissible stress is restricted to 30,000 psi, a parametric study showed that the following maximum allowable relative displacements may occur during a seismic disturbance without overstressing the pipe: Length, ft 1 10 25 50 75 100 Displacement, in. 0.002 0.20 1.25 5.01 11.27 20.04 It is therefore concluded that the service water lines could withstand, without being overstressed, relative bedrock displacements associated with the earthquakes defined for the Indian Point site. 1.11.4.19 Seismic Evaluation of the Fan Cooler and Passive Hydrogen Recombiner Systems The seismic analysis of the fan cooler system was completed in two parts.

1. Analysis of the structural steel enclosure of the fan cooler units to include the effect of supported equipment.

The structural analysis considering simultaneous incident pressures and earthquake forces was conducted on particular members, plates, and connections, which are, considered critical to the structural performance of the reactor containment fan coolers. This analysis included consideration of the mass of all components supported partially or wholly by the enclosure. The fan, fan motor, and fan motor heat exchanger, although entirely within the enclosure, are independently supported from the concrete floor that makes up the base of each unit. Chapter 1, Page 53 of 71 Revision 22, 2010

IP2 FSAR UPDATE Earthquake loadings were treated using the response spectra techniques. A horizontal force of 0.6W and a vertical force of 0.4W, where W is the weight of the member including static forces, were concentrated at the center of gravity of each member. A negative differential pressure of 1.5 psig was applied to the portion of the unit from the inlet up through the fan compartment; a negative differential pressure of 6.3 psig was applied to the charcoal filter compartment. Both of these values are consistent with unit geometry and containment environment following a loss of coolant accident or main steam line break. The charcoal filter compartment pressure is limited by a pressure equalization device installed during the 1997/1998 Maintenance Outage. All loadings were assumed to act simultaneously and comparison was made with allowable stresses consistent with specifications for installed materials. An increase in allowables was considered for loads associated with accident conditions. Where applicable, allowable concrete stresses were taken from ACI Specifications. Results of the analysis on the enclosure demonstrated that the design is adequate.

2. Evaluation of the fan motor system and its foundation.

The fan motor and its supporting structural system was evaluated using acceleration values for a maximum hypothetical earthquake. These values are 0.6g for the horizontal direction and 0.4g for the vertical direction. These accelerations were assumed to act simultaneously. The failure modes considered for the motor unit were excess deflection of the rotor shaft, which results in rubbing against the housing or by bearing failure. The failure modes for the fan are failure of the fan shaft support bearings or deflection of the fan housing and fan wheel causing binding. In addition, the potential for shear and overturning failure of the motor fan assembly at the foundation anchorage was evaluated. Based on analyses made on similar fan motor systems, it was concluded the fan cooler units in the containment are adequately designed to resist the seismic loading defined for the site and supporting building structure. The two hydrogen recombiners are located in the containment at the 95-ft elevation. The hydrogen recombiners are as shown on Figure 6.8-1. The Passive Hydrogen Recombiners (PHRs) are seismic class I, and have undergone qualification testing in accordance with IEEE 344-1987. On April 14, 2005, NRC issued IP2 License Amendment 243 which eliminated the requirement for hydrogen recombiners to provide any combustible gas control function. 1.11.4.20 Masonry Walls In response to IE Bulletin 80-11, safety related masonry walls were evaluated to demonstrate the ability to withstand the specified design load conditions without impairment of wall integrity or the performance of required safety functions. NRC acceptance of this evaluation is documented in reference 10. As a result of this evaluation, certain walls in the control building, Chapter 1, Page 54 of 71 Revision 22, 2010

IP2 FSAR UPDATE the Unit No. 1 Superheater building, the boric acid evaporator building, the fan house, and the fuel storage building have been reinforced. 1.11.5 Wind Effects The IP2 licensing basis does not include tornado protection for the design of the buildings, structures and components. Tornado protection is not a design criterion for IP2. However, the following structures were evaluated for tornado loads: containment building, primary auxiliary building, control building, fuel storage building (including the spent fuel pit), and the intake structure. Detailed information on the containment structure is found in Appendix B of the Containment Design Report. The containment structure will not be penetrated by a 4-in. x 12-in. x 12-ft wood plank traveling at 300 mph, or by a 4000-pound auto traveling at 50 mph less than 25-ft above the ground. With respect to the primary auxiliary building, control building, and fuel storage building, information from the siding manufacturer indicates that siding panels will blow out at 170 psf, which is equivalent to a 1.18 psi negative pressure. Panels fail at 60 psf external pressure, which is equivalent to a 162 mph external wind load (60 psf controls the external loading condition). The grits will fail at 90 psf, which is equivalent to a 0.62 psi negative pressure. The 3.25-in. thick siding panels are not capable of resisting any tornado-generated missiles. Spent fuel pit tornado protection is discussed in proprietary WCAP-7313-L. The intake structure is capable of resisting any wind or missile loads generated by a tornado. This is true for the structure itself, but does not necessarily include associated equipment. 1.11.6 Structural Effects The potential for damage to Class I structures due to failure of nearby Class II or Class III structures, or due to failure of Class III cranes, has been considered. The only Class I structures and components that could be endangered by failure of Class III structures are the control building, main steam piping, and feedwater piping, which could be endangered by failure of the Class III turbine building. No special provisions were provided in the original plant design except in the case of the main steam and feedwater lines up to the isolation valves, which are protected by the shield wall and the structural frame at the north end of the shield wall. Evaluations were performed and bracing was added to the turbine building during construction, as described in section 1.11.6.5, to preclude such catastrophic failures. The only Class III crane whose failure could endanger any Class I function is the fuel storage building crane. Failure of this crane will not impair a safe and orderly shutdown. The wheels of the bridge and the trolley are shaped such that sliding perpendicular to the rail would not be possible. The lateral load from an earthquake on the trolley crane rail is about 50-percent greater than the lateral loads from impact specified by the AISC Code for design within working stress limits. The stresses on the crane rail are low due to the earthquake load. For this reason no failure of the crane rail is anticipated. The Class III manipulator crane in the containment building is restrained from overturning and will not endanger Class I structures. Chapter 1, Page 55 of 71 Revision 22, 2010

IP2 FSAR UPDATE The turbine building and the fuel handling building are functionally Class III structures. However, these structures have been analyzed using a multidegree of freedom modal dynamic analysis method to ensure that there is no potential for gross structural collapse of these structures as a result of the maximum hypothetical earthquake. The results of the analyses are given below. A value of 7-percent structural damping was assumed in the analysis. Total response of the structure was determined on the basis of the square root sum of the squares basis of each mode contribution. A similar dynamic analysis was also performed to ensure that no potential gross failure of the Indian Point Unit 1 stack or superheater building could occur for the maximum hypothetical earthquake, or for the design-basis tornado for Indian Point Unit 2. The resultant dead, live, and seismic design stresses in the basic building structure is limited to 0.9 yield of the steel. The results of specific analyses are discussed in the following sections. 1.11.6.1 Seismic Analysis of the Indian Point Unit 2 Turbine Building A spectrum response analysis was performed for the turbine building considering the design-basis earthquake (DBE), which has a peak horizontal ground acceleration of 0.15g. The associated earthquake response spectrum is shown in Figure 1.11-2. The foundation was considered rigid since the footings for the structural frames of the building are underlaid by either rock or a lean concrete, which bears on rock. Also, in the analysis, interaction between the turbine and the structural frame for the building was neglected. The analysis, as performed, represents a linear elastic system. The analysis of the turbine building was performed under the assumption that the north-south motions, east-west motions and vertical motions will be uncoupled. The dynamic analysis effort was limited only to horizontal motions in the east-west and north-south directions. However, vertical components of the earthquake were considered by adding a 0.13g component to dead loads. Each of the models was simulated for the computer program called STARDYNE. A description of the modeling capabilities of STARDYNE are contained in "STARDYNE Structural Analyses Systems Users' Manual" prepared by Mechanics Research, Inc., for Control Data Corporation. The STARDYNE program was used in three ways. First, the portal frames were analyzed for a static unit force at each portal to determine their resistance to horizontal motions resulting from the turbine bay crane. This information was incorporated into the model for the analysis of the crane girder to determine the distribution of horizontal turbine bay crane loads to the various east-west portal frames. Secondly, the program was used to determine the forces induced in the frames as a result of gravity forces, and, thirdly, the STARDYNE program was used to determine the fundamental frequencies of each of the models and the characteristic shapes. In addition, the STARDYNE program is also capable of determining the modal member forces for each of the fundamental frequencies. This information for each model and mode was stored on tape along with the gravity forces for each model and later used in an earthquake analysis program to determine the maximum probable deflection, acceleration, member forces, member stresses, and the combined gravity plus earthquake member stress responses. Dynamic characteristics of the turbine building are shown in Table 1.11-4. Results of the analysis indicated that the 0.9 Fy combined load allowable stress was not violated except locally in the flange of columns where cross bracing framed in eccentric to other Chapter 1, Page 56 of 71 Revision 22, 2010

IP2 FSAR UPDATE joint members. Reduction of stresses to allowable values is accomplished by the addition of flange cover plates. While allowable stresses in the cross bracing did not exceed the 0.9 yield stress allowable, it was determined that most of the "x" cross bracing would buckle at very low compressive stress due to high l /r ratios. In order to assure the lateral stiffness of the bents and load carrying capacity as determined in the analysis, cover plates were attached to the bracing equal to the original area of the "x" crossing bracing. This assures design adequacy with only "x" cross bracing in tension assumed to be active in carrying lateral load. 1.11.6.2 Seismic Evaluation of the Fuel Storage Building Structure Above the Spent Fuel Pit The fuel storage building for Indian Point Unit 2 consists of the spent fuel pit constructed of reinforced concrete and founded on rock. The fundamental frequency of the pit is approximately 22 cps and therefore can be considered rigid. The steel superstructure above the pit encloses the pit and supports the fuel cask handling crane. This superstructure was designed as a Class III structure. The seismic loads used in the analysis of the steel superstructure were as follows:

1. Zero period ground acceleration: 0.15g horizontal, 0.10g vertical.

2. 7-percent damping.

3. Response spectrum curve as defined in Figure 1.11-2.

4. Inertial forces for each mass point are determined on the basis of the square root of the sum of the squares.

A dynamic multidegree of freedom, modal analysis of the structure was constructed as shown in Historical Figures 1.11-3 and 1.11-4. The stiffness properties of the elements were determined by the combined stiffness of the frame bents in the north-south and east-west directions taken separately. The stiffness of each bent was determined by the computer program STRUDL. The total inertial forces determined by the dynamic analysis were distributed to each individual bent and resultant member stresses were determined. The crane was assumed fully loaded. Evaluation of these seismic stresses show maximum stresses occurring in diagonal bracing. The maximum stress thus determined in the cross bracing was 18.5 ksi. The maximum combined dead and seismic column load stress determined by the analysis was 12.8 psi compression. On the basis of these results it was determined that the fuel storage building superstructure was adequately designed to carry the seismic load defined for the site. In addition to the analysis of the building structure, the fuel crane bridge was evaluated to determine the potential for the crane bridge to lift off its track support in the event of a seismic disturbance. The vertical mode fundamental frequency of the fuel storage building is approximately 9 cps. The crane bridge has also been analyzed dynamically both loaded and unloaded and for various positions of the trolley. It was determined that the crane with the trolley at the end of the span and unloaded would have a fundamental frequency of approximately 9 cps. Considering potential resonance with the fundamental vertical mode of the building at 9 cps the resulting g-loading was 1.05g. The only potential for crane lift-off will be in the unloaded condition with the trolley parked near the support. Since the unloaded crane will not be parked over the pool no potential hazard exists and vertical restraints are not required. Chapter 1, Page 57 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.11.6.3 Seismic and Wind Analysis of the Superheater Stack of Indian Point Unit 1 The Indian Point Unit 1 superheater stack has been analyzed for seismic, tornado, and vortex-shedding wind load effects. The results of this analysis are summarized below. As a result of this analysis on the existing stack it is concluded:

1. The stack can withstand a tornado wind load of approximately 300mph prior to buckling failure of the stack steel shell.

2. The maximum stress in the stack at the critical vortex-shedding frequency wind velocity is 7660 psi, which provided a 3.64 factor of safety against stack failure by this mode.

3. The maximum combined dead and seismic stress for the earthquake parameters defined for the site is 19,140 psi, which provides a 1.46 factor of safety against stack failure by this mode.

1.11.6.3.1 Load Case 1 - Tornado I. Load Criteria Wind = 300 mph L = D + W' where: L = Total load D = Dead load W'= Tornado load II. Method of Load Analysis As prescribed in ASCE Paper 3269 for uniform wind velocity with height; no gust factor. III. Allowable Stress Criteria 0.72Et a = = 27,900 psi (1 2 )r where: a = allowable stress (psi) E = modulus of elasticity (psi) t = shell thickness (in.) Chapter 1, Page 58 of 71 Revision 22, 2010

IP2 FSAR UPDATE

               = Poisson's ratio r = radius of stack (in.)

IV. Stress Determination D Wy r

             =        +          = 1.54 + 25.75 = 27.29 ksi A         I where:

y = centroidal height of stack (in.) I = moment of inertia of stack (in.4) A = cross sectional area of stack (in.2) a 27.9 Factor of Safety = = = 1.02 27.29 1.11.6.3.2 Load Case 2 - Seismic I. Load Criteria a) Zero period ground acceleration: 0.15 g horizontal; 0.10 g vertical. b) Damping 7-percent. c) Ground response curve - Figure 1.11-2. L = D + E' = E' v h where: Eh = load resulting from horizontal earthquake component EV =load resulting from vertical earthquake component II. Method of Load Analysis Multidegree of freedom modal analysis of the superheater building and stack as shown in Figure 1.11-5. The square root of the sum of the squares of seismic inertia forces at mass points is used to determine resultant shears and moments in the stack. III. Allowable Stress Criteria Chapter 1, Page 59 of 71 Revision 22, 2010

IP2 FSAR UPDATE See Load Case 1, item III. IV. Stress Determination D E' v EhXr

                 =        +       +

A A I

                = 1.54 + 0.20 + 17.4 = 19.14 a        27.9 Factor of Safety =         =

19.14

                                    = 1.46 where:

X = lever arm of node inertia force 1.11.6.3.3 Load Case 3 - Vortex-Shedding I. Expression for maximum uniformly distributed force due to vortex-shedding. P = (MF) 1 2 v2 x CL x D x L CL = Lift coefficient for a stationary circular cylinder MF = A multiplying factor applied to the lift coefficient to account for a vibrating cylinder D = Average stack diameter (ft) L = Length of stack (ft)

        = Logarithmic decrement
        = Air density (0.0023385 lb - sec2/ft4) v = F1 x Vc Vc = Critical vortex-shedding velocity (fps)

F1 = A correction factor, which accounts for the fact that stack oscillations have occurred as high as 30-percent above shedding velocity Chapter 1, Page 60 of 71 Revision 22, 2010

IP2 FSAR UPDATE fxD Vc = S S = Stronhal number f = Fundamental frequency (cps) II. Pertinent parameters CL = 0.1 MF = 4.0 D = 20-ft L = 334.5-ft

        = 0.04 (2-percent critical damping)

Vc = 42.7 fps F1 = 1.2 S = 0.27 f = 0.576 cps III. Stress criteria D Phr

         =     +       = 1.54 + 6.12 = 7.66ksi A      2I a   27.9 Factor of Safety =      =       = 3.64 7.66 In addition to the analysis performed for the existing stack it was determined that the stack with 80-ft removed from the top would have the capacity to resist a 360 mph wind for the criteria as defined in Load Case I; the seismic as defined in Load Case II; and the vortex-shedding as defined in Load Case III.

1.11.6.4 Seismic and Tornado Evaluation of the Superheater Building at Indian Point Unit 1 A spectrum response analysis was performed for the superheater building considering the design basis earthquake, which has a maximum horizontal ground motion of 0.15g. A dampening coefficient equal to seven percent was assumed for all modes. The earthquake response spectra used is shown in Figure 1.11-2 normalized to 0.15g zero period ground acceleration. In the analysis no interaction with the foundation was considered since the footings for the structural frame for the building are underlaid by rock. Also, in the analysis, the Chapter 1, Page 61 of 71 Revision 22, 2010

IP2 FSAR UPDATE stiffness interaction between the turbine building and the structural frame for the superheater building was neglected, but the mass of the turbine building was included in the dynamic analysis. The analysis, as performed, represents a linear elastic system. The analysis of the superheater building was performed under the assumption that the north-south motions, east-west motions, and vertical motions were uncoupled. The analysis effort was limited only to horizontal motions in the east-west and north-south directions, and no attempt was made to model vertical motions or to combine vertical and horizontal motions. However, vertical seismic motions have been considered in the results by increasing the dead load stress in building members by a factor equal to two thirds of the combined mode horizontal inertial g-load as determined in either the east-west or north-south direction. In each direction, north-south and east-west, the column lines were modeled in detail. These structural models were developed for elastic-static analyses obtained from the computer program STRUDL. They were used for two purposes: to develop the master stiffness matrices associated with the two directions, east-west and north-south, used in the dynamic analyses; and to determine resultant member stresses using the equivalent static seismic forces determined from the dynamic analyses. The dynamic characteristics, frequencies, and mode shapes of the superheater building were determined using the Westinghouse computer program SAND. The equivalent static forces resulting from the dynamic response were developed using a response spectrum seismic analysis performed by the Westinghouse computer program SPECTA. The equivalent static force associated with a particular mass resulting from a dynamic response is defined as the square root of the sum of the squares of the equivalent static forces associated with that mass for each mode. The equivalent static force associated with a mode and a mass point is defined as the value of the mass times the maximum acceleration associated with the mass point for that particular mode. The maximum acceleration associated with a mode and mass point is defined as follows: (Ürn ) Max= (n) Max rn (n) Max = n San M ' r r rn n = M r r 2 rn Where: n = Refers to mode n r = Refers to mass r

               Ørn = Component of Ørn in the direction of the earthquake Chapter 1, Page 62 of 71 Revision 22, 2010

IP2 FSAR UPDATE

               Ørn = Component of mode shape n for mass r Mr      = Mass lumped at point r (n)Max= Maximum modal acceleration for mode n San = Spectral acceleration for mode n from response curve for 7-percent damping

( U rn)Max = Maximum acceleration in mode n for mass point r n = Modal participation factor for mode n Sectional views in the north-south and east-west directions are shown in Figures 1.11-5 and 1.11-6. A typical column line modeled for STRUDL to determine overall column line stiffness and permit determination of resultant seismic stresses is shown in Figure 1.11-7. In Figure 1.11-8 is presented the dynamic model used to determine inertial forces. Results of the analysis showed several column lines contained diagonal bracing with stresses, which exceeded the allowable stress value of 0.9 fy. In addition several of the cross bracings showed compressive stress levels, which exceeded the expected buckling stress as determined by the l /r ratio for the member. Overstressed members can be strengthened by attaching cover plates to the angle bracing. In a few instances columns were found to be locally overstressed due to eccentric positioning of cross bracing. These areas can be reinforced by flange cover plates. Approximately 30 tons of additional plate will strengthen the structure. With respect to tornado resistance of the structure, total lateral load in the north-south direction is approximately 10-percent, and in the east-west direction 20-percent, less than the seismic-induced lateral load on the structure. Tornado loads were based on a 360-mph wind using the shape factors for a rectangular building as defined in ASCE Paper 3269. It was assumed that 20-percent of the wall area of the building was still intact as a reaction surface for the wind in addition to the total surface area of major equipment and the stack at its existing height. On the basis of this analysis, the building has approximately the same resistance capacity to a 360-mph tornado wind as it does for the 0.15g earthquake. 1.11.6.5 Evaluation of Structural Modifications In the analysis of the superheater and turbine buildings under lateral loads, the following connections were examined:

1. Gusset plates.

2. Check of connections between beams and columns to determine their adequacy to transfer horizontal shear load.

3. Check of connections at column bases in the foundation to determine their ability to transfer the given horizontal shear load. For those column base connections subjected to a net uplift load, an analysis has been performed to ensure that they are adequate for these loads.

If it was found that a connection was inadequate to support the given load, it was redesigned. Chapter 1, Page 63 of 71 Revision 22, 2010

IP2 FSAR UPDATE It is not necessary to reanalyze the turbine building after the redesign because the building stiffness characteristics are essentially the same as those assumed in the initial analysis. This is because the significant fixes involved the cross bracing system, which is made up of pairs of cross bracing members. In the initial analysis, both sets of cross bracing were assumed active. However, the bracing system was such that cross members would buckle under a very small compressive load. Therefore, lateral building load must be carried in tension by the bracing system. The fix used in the redesign was to double the area of cross bracing. The bracing in compression, due to buckling, is not active in resisting lateral building load. Therefore, only half of the cross bracing assumed in the initial analysis, which is in tension, resists this load. However, since the area of cross bracing has been doubled, the resultant effective lateral resistance is the same as that assumed in the original analysis. An initial analysis was made of the superheater building using the existing design parameters. After completion of the analysis, the overstressed members were strengthened and a dynamic reanalysis made. Tables 1.11-5, 1.11-6, and 1.11-7 give the relative comparisons in stiffness, horizontal inertial load, and frequency between the initial analysis and the reanalysis. Subsequently, retired Unit 1 superheater-associated equipment has been removed from certain areas of the superheater building and the areas refurbished to provide permanent administrative facilities. These areas do not contain any safety-related equipment. The total loading on the superheater building has been reduced from the original design loading due to the removal of superheater-associated equipment. Therefore, the administrative facilities will not adversely affect the response of the superheater building during a safe-shutdown earthquake. 1.11.7 Seismic Qualification For Safe Shutdown In response to NRC Generic Letter (GL) 87-02 and Supplement No. 1 to GL 87-02, Verification of Seismic Adequacy of Mechanical and Electrical Equipment in Operating Reactors, Unresolved Safety Issue A-46, Con Edison committed to implement Generic Implementation Procedure (GIP-2) including the clarifications, interpretations, and exceptions in the NRCs Supplemental Safety Evaluation Report (SSER-2). The NRC accepted Consolidated Edisons response and commitments regarding this issue as documented in their Safety Evaluation Report (SER) dated November 19, 1992. Consolidated Edison has verified the seismic capabilities of equipment required for safe shutdown, as documented in Reference 7. The verification utilized the Generic Implementation Procedure developed by the Seismic Qualification Utility Group and approved by the NRC. A Summary Report, including the Seismic Evaluation and Relay Evaluation Reports, was submitted to the NRC on December 31, 19967. Indian Point 2 site specific SER, dated November 8, 2000, Reference 15, provides NRC conclusions that the licensee may revise its licensing basis by incorporating the SQUG methodology. Revision 3 of the Generic Implementation Procedure (GIP-3), Reference 12, as modified and supplemented by the NRCs Supplemental Safety Evaluation Report No.2 (SSER-2), Reference 13, and No.3 (SSER-3), Reference 14, may be used as an alternative to existing methods for the seismic design and verification of modified, new, and replacement equipment. Only those portions of GIP-3 as described in Section 5 of Implementation Guidelines for Seismic Chapter 1, Page 64 of 71 Revision 22, 2010

IP2 FSAR UPDATE Qualification of New and Replacement Equipment/Parts (NARE) using the Generic Implementation Procedure, Reference 16, shall apply to the seismic design and verification of mechanical and electrical equipment, electrical relays, tanks and heat exchangers, and cable and conduit raceway systems. 1.11.8 Protection from Flooding of Equipment Important to Safety In response to NRC Guidelines for Protection from Flooding of Equipment Important to Safety, Consolidated Edison identified the potential sources of flooding outside containment that could affect safety-related equipment. The areas containing safety-related equipment that could be subject to flooding from postulated failure of water systems that are not seismic Class I were evaluated. The plant is designed so as to minimize or eliminate the vulnerability of safety-related equipment to this flooding. Modifications were made to install water level alarm switches in the adjoining Unit 1 condenser pit area, and add flap panels in doors from the primary auxiliary building and the auxiliary feed pump room. A later modification was made to install a flood control drain line and valve from the PAB to a manhole in the transformer yard. These modifications along with the implementation of an alternate safe shutdown capability, as discussed in Section 8.3, serve to mitigate the consequences of the postulated flooding. Additionally, operator action would be taken in the event of flooding from the circulating water system to prevent damage to the 480 volt switchgear in the control building. In their Safety Evaluation Report (SER) dated December 18, 1980, the NRC determined that design features and operating procedures provide assurance that the plant can be safely shut down in the event of flooding outside containment from a non-seismic component or pipe and that their guidelines (contained in Appendix A to the SER) have been satisfied.8 The Fire Protection piping added later in the PAB was analyzed for seismic loads and restrained to prevent breakage. REFERENCES FOR SECTION 1.11

1. E. L. Vogeding, Topical Report - Seismic Testing of Electrical and Control Equipment, WCAP-7397-L, Westinghouse Electric Corporation, January 1970.

2. Westinghouse Electric Corporation, Ultimate Strength Criteria to Ensure No Loss of Function of Piping and Vessels Under Earthquake Loading, WCAP-5890, Revision 1.

3. NRC Generic Letter, Relaxation in Arbitrary Intermediate Pipe Rupture Requirements, G.L. 87-11, dated June 19, 1987.

4. NRC Branch Technical Position MEB 3-1, Postulated Rupture Locations in Fluid System Piping Inside and Outside Containment.

5. Letter (Attachment B) from S. Bram, Con Edison, to NRC,

Subject:

Request for License Amendment to Technical Specification Modifying Spent Fuel Storage Requirements, dated June 20, 1989.

6. Letter (Attachment I) from S. Bram, Con Edison, to NRC,

Subject:

Indian Point Unit No. 2 Spent Fuel Storage Capacity Increase, dated January 19, 1990. Chapter 1, Page 65 of 71 Revision 22, 2010

IP2 FSAR UPDATE

7. Letter from Quinn, Con Edison, to NRC,

Subject:

Summary Report for Resolution of USI-A-46, Seismic Qualification, dated December 31, 1996.

8. Letter from S. A. Varga, NRC, to J. D. O'Toole, Con Edison,

Subject:

Safety Evaluation Report Susceptibility of Safety-Related Systems to Flooding from Failure of Non-Category I Systems, dated December 18, 1980.

9. Letter from Cahill, Con Edison, to A. Schwencer, Director of Nuclear Reactor Regulation NRC,

Subject:

Supplemental Response to IE Bulletins 79-02 and 79-07, dated November 27, 1979.

10. Letter from Steven A Varga, NRC to John D. OToole Con Edison,

Subject:

Completion of IE Bulletin 80-11, Masonry Wall Design for Indian Point Nuclear Generating Unit No. 2 (IP2), (Safety Evaluation Report included) dated October 19,1983.

11. Altran Corporation, Technical Report No. 00222-TR-001, Rev. 1, Reactor Coolant Loop Analysis for Replacement Steam Generators and Snubber Reduction

12. Generic Implementation Procedure (GIP) for Seismic Verification of Nuclear Plant Equipment, Revision 3, Updated 05/16/97 (GIP-3). Prepared by SQUG and sent to the NRC by letter dated May 16,1997.

13. Supplement No.1 to Generic Letter (GL) 87-02 that transmits Supplemental Safety Evaluation Report No.2 (SSER-2) on SQUG Generic Implementation Procedure Revision 2, as Corrected on February 14, 1992 (GIP-2), May 22, 1992.

14. NRC letter to SQUG dated December 4, 1997, Supplemental Safety Evaluation Report No.3 (SSER-3), on the Review of Revision 3 to the Generic Implementation Procedure for Seismic Verification of Nuclear Power Plant Equipment, Updated 05/16/97 (GIP-3)

15. Indian Point Nuclear Generating Station No. 2 - Plant Specific Safety Evaluation Report for Unresolved Safety Issue A-46 Program Implementation, November 8, 2000.

16. Implementation Guidelines for Seismic Qualification of New and Replacement Equipment/Parts (NARE) Using the Generic Implementation Procedure (GIP),

Revision 4, by MPR Associates Chapter 1, Page 66 of 71 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1.11-1 Damping Factors PERCENT OF COMPONENT CRITICAL DAMPING Containment structure 2.0 Concrete support structure of reactor 2.0 vessel Steel assemblies: 2.5 Bolted or riveted 1.0 welded Vital piping systems 0.5 Concrete structures above ground 5.0 Shear Wall 5.0 Rigid Frame Chapter 1, Page 67 of 71 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1.11-2 Loading Combinations and Stress Limits Loading Combinations Vessels1 Piping Supports

1. Normal Pm Sm Pm S Working stresses loads PL + PB 15. Sm PL + PB S or applicable factored load design values

2. Normal + Pm Sm Pm 12

                                                       . S                1-1/3 working design         PL + PB  15
                            . Sm                 PL + PB  12
                                                            . S           stresses or earthquake                                                              applicable loads                                                                   factored load design values

3. Normal + Pm 12. Sm Pm 12 . S Deflections and maximum . (15 PL + PB 12 . Sm) PL + PB 12. (15

                                                                  . S)    stresses of potential                                                               supports limited earthquake                                                              to maintain loads                                                                   supported equipment within their stress limits

4. Normal + Pm 12. Sm Pm 12 . S Deflections and pipe . (15 PL + PB 12 . Sm) PL + PB 12. (15

                                                                . S)      stresses of rupture                                                                 supports limited loads                                                                   to maintain supported equipment within their stress limits Where: Pm      =      primary general membrane stress; or stress intensity PL     = primary local membrane stress; or stress intensity PB     =   primary bending stress; or stress intensity Sm     = stress intensity value from ASME B and PV Code, Section III S      = allowable stress from USAS B31.1 Code for Pressure Piping notes:

1. Limited to vessels designed to ASME, Section III, Class A (or Class 1) rules. Otherwise use piping for stress limits.

Chapter 1, Page 68 of 71 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1.11-3 DELETED TABLE 1.11-4 Dynamic Characteristics of the Turbine Building MODE Frequency Values No. (cps) 1 0.5042 0.08 2 1.6141 0.12 3 2.2849 0.19 4 4.3292 0.2 5 5.2813 0.2 6 8.2814 0.18 7 12.1704 0.15 9 15.1274 0.15 10 20.754 0.15 11 22.4809 0.15 12 23.8001 0.15 13 27.3040 0.15 14 33.9678 0.15 TABLE 1.11-5 Relative Stiffness Percentages Percentage Increase In Stiffness Between First And Second Analysis (Percent) RELATIVE LOCATION IN EAST-WEST NORTH-SOUTH SUPERHEATER DIRECTION DIRECTION BUILDING BOTTOM 8 56.7 MIDDLE 18.3 41.4 TOP 19.9 10.4 Chapter 1, Page 69 of 71 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1.11-6 Inertial Loads Relative Inertial Loads for First and Second Analysis Location in (Units: Kips) Superheater East-West North-South Building Direction Direction Original Reanalysis Original Reanalysis Bottom 908 908 1091 1102 Middle 1888 1914 1687 1803 Top 1242 1271 1082 1181 TABLE 1.11-7 Frequencies Frequencies For First And Second Analysis (Units: cps) EAST-WEST NORTH-SOUTH DIRECTION DIRECTION MODE ORIGINAL REANALYSIS ORIGINAL REANALYSIS 1 0.94 1.0 0.72 0.88 2 2.07 2.15 1.58 2.13 3 4.08 4.19 3.47 4.12 1.11 FIGURES Figure No. Title Figure 1.11-1 Ten Percent of Gravity Response Spectra Figure 1.11-2 Fifteen Percent of Gravity Response Spectra Figure 1.11-3 Fuel Storage Building North-South Model [Historical] Figure 1.11-4 Fuel Storage Building East-West Model [Historical] Figure 1.11-5 Indian Point Unit 1 Superheater Building North-South Section Figure 1.11-6 Indian Point Unit 1 Superheater Building East-West Section Figure 1.11-7 Column Line G Figure 1.11-8 Representation of Lumped Mass Model of Superheater Building Used in Dynamic Analysis Chapter 1, Page 70 of 71 Revision 22, 2010

IP2 FSAR UPDATE 1.12 INSERVICE INSPECTION AND TESTING PROGRAMS 1.12.1 General The ISI Program complies with the requirements of 10 CFR 50.55a and is based upon the requirements set forth in ASME Boiler and Pressure Vessel Code, Section XI and by the applicable Code year. This program is also responsive to pertinent provisions of applicable Regulatory Guides. The Indian Point Unit 2 Inservice Inspection (ISI) and Testing (IST) Programs for the ten year interval are controlled by Program Plans and plant procedures. 1.12.2 Application The ISI program applies to Quality Groups A, B, and C systems, components (including supports), and pumps and valves as classified in accordance with Regulatory Guide 1.26, Revision 3. 1.12.3 Program Summary The ISI and IST programs identify the specific systems, components, or parts thereof to be examined and the specific pumps and valves to be tested. 1.13 CONTROL OF HEAVY LOADS In response to a December 22, 1980 Generic Letter and to NRC Staff guidelines provided in NUREG-0612, Control of Heavy Loads at Nuclear Power Plants, Con Edison performed evaluations of provisions for the handling and control of heavy loads in the vicinity of irradiated fuel or safe shutdown equipment. Control of heavy loads in the Fuel Storage Building is addressed in section 9.5.6. The NRC documented their acceptance of Con Edisons assessments in a Safety Evaluation Report dated February 19, 1985. Chapter 1, Page 71 of 71 Revision 22, 2010

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IP2 FSAR UPDATE CHAPTER 2 SITE AND ENVIRONMENT 2.1

SUMMARY

AND CONCLUSIONS This section of the FSAR sets forth the site and environmental data, which together formed the basis for the criteria for designing the facility and for evaluating the routine and accidental release of radioactive liquids and gases to the environment. These data support the conclusion that there will be no undue risk to public health and safety with the plant as designed and the environmental characteristics as described. This conclusion rests not only upon the data, but upon the scientific documentation of several independent consultants in their particular area of expertise-health physics, demography, geology, seismology, hydrology, and meteorology. Environmental characteristics of the area have been documented by field measurements and studies conducted since 1958. These studies quantified the effects on the environment of the operation of nuclear power plants. Conservative projections have been made of the probable growth of population in the area, and these projections have been taken into account in plant design both as to control of accidents and as to assumptions about operation. [Historical Information] According to 1980 population estimates, about 50 people reside within a 1100-m radius of Unit 2 (most of them to the east-southeast), and approximately 2600 live within 1-mile. Approximately 75,000 people reside within a 5-mile radius of the facility. The largest concentration of population is in the City of Peekskill, the center of which is about 2.5-miles northeast of the site. The most densely populated 15-degree sector, within 5-miles, is toward Peekskill to the northeast. The 1960 population within a 15-mile radius of the site was approximately 352,000, whereas in the year 2000 the estimated population is 1,107,195. The projections do not indicate, and there is no reason to conclude otherwise, that the land usage within this radius will shift appreciably during the intervening period. (The land is now zoned principally for residential and state park use, although there is some industrial activity and minor or isolated agricultural and grazing activity.) The outer boundary of the low-population zone has been set at 1100 m from Unit 2. Geologically, the site consists of a hard limestone in a jointed condition that provides a solid bed for the plant foundation. The bedrock is sufficiently sound to support any loads that could be expected up to 50 tons/ft2, which is far in excess of any load that may be imposed by the plant. Although it is hard, the jointed limestone formation is permeable to water. Thus, if water from the plant should enter the ground (an improbable event since the plant is designed to preclude any leakage into the ground), it would percolate to the river rather than enter any ground-water supply. About 80 million gallons of Hudson River water flow past the plant each minute during the peak tidal flow. This flow will provide additional mixing and dilution for liquid discharges from the facility. The assumption in the plant design is to treat the river water as if it were used for drinking and thus to reduce radioactive discharge, by dilution with ordinary plant effluent, to concentrations that would be tolerable for drinking water. There is a very low probability of flooding at the site. [Historical Information] Seismic activity in the Indian Point area is limited to low-level microseismicity. Detailed field investigations1-3 have been conducted in the immediate vicinity of Chapter 2, Page 1 of 119 Revision 22, 2010

IP2 FSAR UPDATE Indian Point and along the major faults in the region. To date, no evidence has been found in the rocks exposed at the surface or sediment overlying fault traces or in cores obtained in the vicinity of Indian Point that might support a conclusion that displacement has occurred along major fault systems within the New York Highlands, the Ramapo, or its associated branches during Quaternary time (the last 1.5 million years). In the vicinity of Indian Point, evidence that no displacement has occurred in the last 65 million years (since the Mesozoic) along specific major structures has been observed. The plant is designed to withstand an earthquake of Modified Mercalli Intensity VII as required by Appendix A to 10 CFR 100 "Seismic and Geologic Siting Criteria for Nuclear Power Plants." The validity of the selection of an Intensity VII earthquake was adjudicated before the Atomic Safety Licensing Appeal Board. The Appeal Board's decision (ALAB-436) verified Intensity VII as the plant's design-basis earthquake. Meteorological conditions in the area of the site were determined during a 2-year program (1955 to 1957). The validity of these conclusions has been verified by several programs, including that performed by the Atmospheric Services Department of York Services Corporation in completing a meteorological update for Consolidated Edison Company in 1981 (see Appendix 2A). These data have been used in evaluating the effects of gaseous discharges from the plant during normal operations and during the postulated loss-of-coolant accident. The evaluations indicate that the site meteorology provides adequate diffusion and dilution of any released gases. Environmental radioactivity has been measured at the site and surrounding area since 1958 in association with the operation of Indian Point Unit 1 and the construction and operation of Indian Point Units 2 and 3. Unit 3 is owned by Entergy Nuclear Indian Point 3, LLC. These measurements will be continued and reported. The radiation measurements of fallout, water samples, vegetation, marine life, etc., have shown no perceptible post-operative increase in activity. Noticeable increases in fallout have coincided with weapons-testing programs and appear to be related almost entirely to those programs. The New York State Department of Health in an independent 2-year postoperative survey4,5 found that environmental radioactivity in the vicinity of the site is no higher than anywhere else in the State of New York. [Historical Information] Consultants who have participated in the preparation of the various reports, measurements, and conclusions appearing in this chapter include Dr. Merril Eisenbud, director of Environmental Radiation Laboratory, Institute of Industrial Medicine, New York University; Dr. Benjamin Davidson (deceased), meteorologist and director, Geophysical Science Laboratory, New York University College of Engineering; Dr. James Halitsky, senior research scientist, Department of Meteorology and Oceanography, New York University, College of Engineering; Dr. Edgar M. Hoover, Regional Economic Development Institute, Inc.; Metcalf and Eddy Engineers, hydrology specialists; Quirk, Lawler, and Matusky Engineers, Environmental Science and Engineering Consultants; Mr. Karl R. Kennison, consulting civil and hydraulic engineer; and Woodward-Clyde Consultants, consulting engineers, geologists and environmental scientists. REFERENCES FOR SECTION 2.1

1. Ratcliffe 1976.

2. Ratcliffe 1980.

3. Dames & Moore 1977.

Chapter 2, Page 2 of 119 Revision 22, 2010

IP2 FSAR UPDATE

4. Hollis S. Ingraham, Consolidated Edison Indian Point Reactor Environmental and Post Operational Survey - August, 1965, Division of Environmental Health Services, New York State Department of Health.

5. Hollis S. Ingraham, Consolidated Edison Indian Point Reactor Environmental and Post Operational Survey - July, 1966, Division of Environmental Health Services, New York State Department of Health.

2.2 LOCATION 2.2.1 General [Historical Information] Indian Point is a multiunit site consisting of approximately 239 acres of land on the east bank of the Hudson River at Indian Point, village of Buchanan, in upper Westchester County, New York. Indian Point Units 2 and 3 (see Section 2.2.3) are located north and south, respectively, of Unit 1, which has been retired. The site is about 24-miles north of the New York City boundary line. The nearest city is Peekskill, located 2.5-miles northeast of Indian Point, with a population of about 20,000. An aerial photograph, Historical Figure 2.2-1, shows the site and about 58-mile2 of the surrounding area. 2.2.2 Access The site is accessible by several roads in the village of Buchanan. A paved road links the eastern boundary of the site to the existing plant. The existing wharf is used to receive heavy equipment as needed. The site is not served by rail. 2.2.3 Site Ownership And Control Entergy owns the Indian Point Units 1 and 2 Nuclear Power Plants. As shown in Figure 2.2-3, the Algonquin Gas Transmission Company has a 26 inch gas mainline and a 30 inch gas mainline on a 65 foot wide right-of-way running east to west through the property. Unit 2 is 1450-ft north of the 26-in. Algonquin gas mainline. One 30 inch main and 2-24 inch mains pass under the river to a pipeline facilities station on the easement near the river. One 24 inch main is available as a bypass alternative and ends in the pipeline facilities station while the other two continue as the 30 inch and 26 inch mains. The Georgia-Pacific Corporation has an easement, 1610-ft long and 30-ft wide, through the southerly part of the Indian Point site. The Georgia-Pacific easement is used for overhead electrical power and telephone lines and underground gas, water, and sewer lines. These easements permit Entergy to determine all activities within the right-of-way in order to ensure safe operation of the units. Units 1, 2, and 3 have a security fence surrounding the "protected" areas. Access to the protected areas is controlled via security buildings that are manned on a 24-hr basis. In addition, spaces within the protected area designated as "vital areas" are provided with additional access control. All roads within the site are continuously patrolled by security personnel. A site plot plan is shown in Historical Figure 2.2-2. Chapter 2, Page 3 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.2.4 Activities On The Site The principal activities on the site are the generation, transmission, and distribution of electrical energy; associated service activities; activities relating to the controlled conversion of the nuclear energy of fuel to heat energy by the process of nuclear fission; and the storage, use, and production of special nuclear source and byproduct materials. 2.2 FIGURES Figure No. Title Figure 2.2-1 Aerial Photo of Indian Point Site and Surrounding Area [Historical] Figure 2.2-2 Indian Point Building Identification [Historical] Figure 2.2-3 Algonquin Gas Transmission Pipeline Hudson River Crossing & Indian Point Nuclear Generation Facility 2.3 TOPOGRAPHY [Historical Information] The Indian Point Generating Station is on the east bank of the Hudson River. The river runs northeast to southwest at this point but turns sharply northwest approximately 2-miles northeast of the plant. The west bank of the Hudson is flanked by the steep, heavily-wooded slopes of the Dunderberg and West Mountains to the northwest (elevations 1086 and 1257-ft, respectively, above mean sea level) and Buckberg Mountain to the west-southwest (elevation 793-ft). These peaks extend to the west and gradually rise to slightly higher peaks. The general orientation of this high ground is northeast to southwest. One mile northwest of the site, Dunderberg bulges to the east. North of Dunderberg and the site, high grounds reaching 800-ft form the east bank of the Hudson River. At this location the Hudson River makes a sharp turn to the northwest. To the east of the site, peaks are generally lower than those to the north and west. Spitzenberg and Blue Mountains average about 600-ft in height, and there is a weak, poorly-defined series of ridges that run in a north-northeast direction. To the west of the site there are the Timp Mountains at an elevation of 846-ft. To the south of the site, elevations of 100-ft or less gradually slope towards Verplanck. The river south of the site makes another sharp bend to the southeast and then widens as it flows past Croton and Haverstraw. Historical Figure 2.3-1 shows topographic features of the site and the surrounding areas. 2.3 FIGURES Figure No. Title Figure 2.3-1 Topographical Map of Indian Point and Surrounding Area [Historical] Chapter 2, Page 4 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.4 POPULATION AND LAND USE 2.4.1 Overview The population within a 50-mile radius of the Indian Point site has been estimated for 1990. These population estimates were taken from statistics recently released by the U.S. Census Bureau. The population within the 50-mile radius of Indian Point has increased from the 1980 estimates by approximately 68,000 people, less than half of one percent. 2.4.2 Population And Land Use According to 1990 estimates, approximately 15.465 million people live within a 50-mile radius of the Indian Point site. A major part of this number live in New York City, an area 25 to 50-miles south of the plant. Approximately 1650 persons, concentrated in sectors south to southeast of the station, live within 1-mile of the plant. Approximately 74,000 persons live within 5-miles of the plant. The area surrounding the Indian Point site is generally residential with some large parks and military reservations. Some increased commercial development has occurred within a mile of the station since 1980. Most of the area to the east of the Hudson River within 15-miles of the site is zoned for residential uses. West of the Hudson within a 15-mile radius, the Palisades Interstate Park and residential areas are the dominant land uses. The only agricultural areas within 15-miles are south or northwest of the plant on the west side of the River. Several maps and tables are included to illustrate the population distribution and land use of the area. Figure 2.4-1 and Figure 2.4-2 show the sector/zone approach to the population data and the area within a 50-mile radius of the Indian Point site. Historical Figures 2.4-3 through Figure 2.4-5 illustrate the 1980 population distribution radically by sectors out to 50-miles from the plant site. Historical Figure 2.4-6 through Figure 2.4-8 show, respectively, the land uses based on official zoning maps, areas served by public utilities, and areas served by sewage systems, all as of 1970. Table 2.4-1 explains the sector/zone designations for the population maps and tables that follow. Table 2.4-2 through Table 2.4-18 give the 1990 estimated populations for all sector/zones within a 50-mile radius of the Indian Point site. The New York State Department of Commerce projects no substantial increases in population from 1986 to the year 2013 in any of the four counties in the vicinity of Indian Point. [Historical] Table 2.4-19 and Table 2.4-20 show the estimated and projected land uses by County for 1960 and 1980, respectively. These estimates were developed by the Regional Economic Development Institute, Inc., from Regional Planning Association data. 2.4.3 Low-Population Zone About 50 people reside within a 1100-m radius of Unit 2, most of them to the east-southeast. This distance was used as the outer boundary of the low population zone in the analysis of a postulated fission product release. The water boundary (Peekskill Bay) of the more densely populated area of Peekskill was used as the population center distance, which exceeds 1-1/3 times the distance from the reactor to the outer boundary of the low-population zone. A low-population zone outer boundary radius of 1100-m satisfies both 10 CFR 100.11(a)(3) and 10 Chapter 2, Page 5 of 119 Revision 22, 2010

IP2 FSAR UPDATE CFR 50.67. The low-population zone population in the year 2010 is projected to be approximately 88. 2.4.4 Exclusion Area The exclusion area for Indian Point Unit 2 includes plant property within a 520-m radius of the reactor containment. An exclusion radius of 520-m satisfies both 10 CFR 100.3(a) and 10 CFR 50.67. 2.4.5 Population Data Sources The population data used in this section were developed from the following sources:

1. 1978 Official Population Projections for New York State Counties, prepared by the Economic Development Board, New York State Department of Commerce.

2. Population by Municipality 1970-2000, prepared by the Westchester County Department of Planning, October 1979.

3. Population of Rockland County, Capacity and Forecast, 1970-2000, prepared by the Rockland Planning Board, April 1978.

4. Population Estimate and Projections, Orange County, New York, prepared by the Orange County Planning Department, March 1980.

5. Putnam County Population Projections, prepared by the Putnam County Planning Board, 1977.

6. New Jersey Revised Total and Interim Age and Sex Population Projections, 1980-2000, prepared by the New Jersey Department of Labor and Industry, Division of Planning and Research, Office of Demographic and Economic Analysis, April 1979.

7. State of Connecticut Population Projections for Connecticut Municipalities and Regions to the Year 2000, prepared by the Office of Policy and Management, Comprehensive Planning Division, February 1980.

8. Pennsylvania Projection Series, Summary Report, Employment by Labor Market Area, and Population and Labor Force by County for 1980, 1985, 1990, 1995 and 2000, Report No. 78, PPS-1, prepared by the Office of State Planning and Development, State Economic and Social Research Data Center, June 1978.

Chapter 2, Page 6 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 2.4-1 Sector and Zone Designators for Population Distribution Map1 Sector Nomenclature Zone Nomenclature Centerline of Sector in Degrees True 22.5° Sector2 Miles From Facility Zone North From Facility 0 and 360 A 0-1 1 22.5 B 1-2 2 45 C 2-3 3 67.5 D 3-4 4 90 E 4-5 5 112.5 F 5-6 6 135 G 6-7 7 157.5 H 7-8 8 180 J 8-9 9 202.5 K 9-10 10 225 L 10-15 15 247.5 M 15-20 20 270 N 20-25 25 292.5 P 25-30 30 315 Q 30-35 35 337.5 R 35-40 40 Notes:

1. An area is identified by a sector and zone alphanumeric designator (refer to Figure 2.4-1). Thus, area A1 is that area, which lies between 348.75- and 11.25-degrees true north from the facility out to a radius of 1-mile. Area G4 would be that area between 123.75-to 146.25-degrees and the 3- and 4-mile arcs from the facility.

2. The letters I and 0 have been omitted from sector designators so as to eliminate possible confusion between letters and numbers.

Chapter 2, Page 7 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-2 Population Estimates, 1990, For All Sectors Zone Population 1 1,644 2 15,130 3 18,428 4 14,225 5 24,508 6 25,922 7 28,096 8 25,967 9 36,930 10 46,488 15 342,852 20 488,652 25 920,850 30 2,171,399 35 2,276,172 40 3,451,123 45 3,416,140 50 2,199,601 Chapter 2, Page 8 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-3 Population Estimates, 1990, for Sector A (North) Sector, Zone Population A1 0 A2 70 A3 0 A4 0 A5 400 A6 390 A7 5,301 A8 5,898 A9 2,474 A10 874 A15 4,132 A20 36,987 A25 31,000 A30 57,873 A35 39,998 A40 20,100 A45 17,689 A50 40,853 Chapter 2, Page 9 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-4 Population Estimates, 1990, for Sector B (North-Northeast) Sector, Zone Population B1 0 B2 54 B3 139 B4 143 B5 1,721 B6 1,553 B7 867 B8 246 B9 2,123 B10 1,187 B15 4,343 B20 7,982 B25 20,310 B30 16,651 B35 4,800 B40 6,991 B45 8,457 B50 5,761 Chapter 2, Page 10 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-5 Population Estimates, 1990, for Sector C (Northeast) Sector, Zone Population C1 0 C2 4,879 C3 9,102 C4 4,159 C5 5,534 C6 3,895 C7 2,382 C8 1,594 C9 630 C10 1,034 C15 10,371 C20 9,685 C25 8,200 C30 12,479 C35 13,687 C40 13,067 C45 7,901 C50 6,621 Chapter 2, Page 11 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-6 Population Estimates, 1990, for Sector D (East-Northeast) Sector, Zone Population D1 49 D2 2,379 D3 2,691 D4 1,899 D5 2,324 D6 2,272 D7 4,667 D8 4,713 D9 5,982 D10 3,900 D15 32,854 D20 14,721 D25 8,961 D30 82,240 D35 21,876 D40 18,762 D45 12,991 D50 60,032 Chapter 2, Page 12 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-7 Population Estimates, 1990, for Sector E (East) Sector, Zone Population E1 59 E2 560 E3 0 E4 289 E5 279 E6 345 E7 1,769 E8 1,138 E9 3,287 E10 3,762 E15 17,702 E20 5,099 E25 22,465 E30 20,987 E35 15,730 E40 159,720 E45 162,993 E50 101,121 Chapter 2, Page 13 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-8 Population Estimates, 1990, for Sector F (East-Southeast) Sector, Zone Population F1 147 F2 305 F3 336 F4 689 F5 260 F6 987 F7 475 F8 860 F9 758 F10 1,999 F15 19,121 F20 11,728 F25 49,821 F30 120,701 F35 58,734 F40 33,691 F45 0 F50 29,199 Chapter 2, Page 14 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-9 Population Estimates, 1990, for Sector G (Southeast) Sector, Zone Population G1 575 G2 2,298 G3 1,295 G4 769 G5 420 G6 3,702 G7 3,892 G8 2,672 G9 2,159 G10 6,890 G15 27,939 G20 23,849 G25 86,999 G30 44,001 G35 17,093 G40 79,903 G45 240,102 G50 328,012 Chapter 2, Page 15 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-10 Population Estimates, 1990, for Sector H (South-Southeast) Sector, Zone Population H1 109 H2 1,782 H3 1,363 H4 741 H5 93 H6 0 H7 0 H8 78 H9 5,039 H10 5,752 H15 22,162 H20 103,969 H25 226,002 H30 252,482 H35 209,921 H40 535,969 H45 723,004 H50 469,960 Chapter 2, Page 16 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-11 Population Estimates, 1990, for Sector J (South) Sector, Zone Population J1 531 J2 650 J3 20 J4 129 J5 1,351 J6 4,012 J7 3,133 J8 4,308 J9 5,189 J10 4,321 J15 40,993 J20 55,102 J25 220,032 J30 954,691 J35 1,472,384 J40 1,907,927 J45 1,601,010 J50 702,739 Chapter 2, Page 17 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-12 Population Estimates, 1990, for Sector K (South-Southwest) Sector, Zone Population K1 174 K2 1,245 K3 1,282 K4 2,049 K5 8,093 K6 4,124 K7 2,526 K8 2,531 K9 6,291 K10 9,371 K15 86,297 K20 72,902 K25 146,895 K30 427,391 K35 321,209 K40 534,296 K45 444,572 K50 353,770 Chapter 2, Page 18 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-13 Population Estimates, 1990, for Sector L (Southwest) Sector, Zone Population L1 0 L2 63 L3 1,621 L4 2,694 L5 2,184 L6 4,059 L7 2,876 L8 902 L9 2,087 L10 4,021 L15 26,019 L20 28,753 L25 41,514 L30 94,167 L35 31,725 L40 89,824 L45 124,188 L50 54,722 Chapter 2, Page 19 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-14 Population Estimates, 1990, for Sector M (West-Southwest) Sector, Zone Population M1 0 M2 359 M3 188 M4 399 M5 169 M6 274 M7 170 M8 15 M9 96 M10 271 M15 5,139 M20 4,976 M25 14,343 M30 8,817 M35 21,625 M40 18,889 M45 47,849 M50 30,319 Chapter 2, Page 20 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-15 Population Estimates, 1990, for Sector N (West) Sector, Zone Population N1 0 N2 292 N3 214 N4 0 N5 0 N6 0 N7 0 N8 23 N9 438 N10 63 N15 3,321 N20 8,827 N25 10,234 N30 7,794 N35 14,233 N40 9,028 N45 9,007 N50 2,109 Chapter 2, Page 21 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-16 Population Estimates, 1990, for Sector P (West-Northwest) Sector, Zone Population P1 0 P2 85 P3 52 P4 0 P5 32 P6 58 P7 9 P8 626 P9 357 P10 2,004 P15 17,997 P20 9,983 P25 12,394 P30 47,277 P35 5,927 P40 9,121 P45 3,960 P50 3,917 Chapter 2, Page 22 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-17 Population Estimates, 1990, for Sector Q (Northwest) Sector, Zone Population Q1 0 Q2 0 Q3 125 Q4 189 Q5 55 Q6 0 Q7 29 Q8 321 Q9 0 Q10 1,039 Q15 7,023 Q20 9,872 Q25 10,745 Q30 12,244 Q35 10,160 Q40 7,942 Q45 5,653 Q50 6,962 Chapter 2, Page 23 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-18 Population Estimates, 1990, for Sector R (North-Northwest) Sector, Zone Population R1 0 R2 109 R3 0 R4 76 R5 1,593 R6 251 R7 0 R8 42 R9 20 R10 0 R15 17,439 R20 44,219 R25 10,935 R30 12,144 R35 17,070 R40 5,893 R45 6,764 R50 3,504 Chapter 2, Page 24 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-19 Estimated Land Use in 1960 and Projected Land Use in 19801 Within a 55-Mile Radius Intensive 1960 and 1980 Nonintensive 1960 Nonintensive 1980 1 2 3 4 5 6 7 8 9 10 11 12 Public Community Public Industrial/ Institutional Rights- Facilities Parks Rights- Grand Residential Commercial Total and Park of-Way Total Institutions Recreation of-Way Total Open Totals 1960 Square miles 1032 216 1248 696 418 1114 4062 6424 Percentage of total developed land 43 9 52 29 19 48 High 58 12 45 22 Low 32 2 15 15 1980 Square miles 2040 368 2408 876 784 682 2342 1674 6424 Percentage of total developed land 43 8 51 19 16 14 49 1960-1980 Square miles of land to be developed 1400 220 1620 1228 Percentage of total land to be developed 58 42 Notes:

1. The averages were derived from the data in Table 3. The Use of Developed Land in Selected Areas of the Regions. RPA Bulletin Number 100, Page 21, September 1962. The data for square miles excludes Monmouth County from the original Regional Plan Association (RPA) totals.

Chapter 2, Page 25 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-20 (Sheet 1 of 2) Land Use Projection by County for 1980 in Square Miles Within a 55-Mile Radius Counties in Con Ed Study Area Intensive Low Intensive Community Outside Industrial/ Facilities Parks Public Rights-State In RPA Region RPA Region Residential Commercial Community Recreation Of Way Open Institutions Conn. Fairfield 183 33 92 83 71 171 Litchfield [30]1 [6] [3] [3] [2] [5] New Haven [88] [19] [73] [65] [55] [134] N.J. Bergen 118 22 20 19 16 38 Essex 83 16 6 6 5 12 Hudson 26 5 3 3 2 6 Middlesex 126 (58)2 22 (10) 18 16 14 34 Morris 130 23 69 63 54 129 Passaic 75 14 23 21 18 43 Somerset 71 (24) 13 (4) 16 15 12 30 Sussex [34] [8] [107] [97] [83] [199] Union 63 12 6 6 5 11 Warren [3] [1] [9] [9] [7] [18] N.Y. Dutchess 106 19 152 138 117 283 Nassau 230 41 5 4 4 9 Orange 110 20 154 140 119 286 Putnam 37 6 42 38 32 79 Rockland 56 10 25 23 19 46 Chapter 2, Page 26 of 119 Revision 22, 2010

IP2 FSAR UPDATE [Historical] TABLE 2.4-20 (Sheet 2 of 2) Land Use Projection by County for 1980 in Square Miles Within a 55-Mile Radius Counties in Con Ed Intensive Low Intensive Study Area Community Outside Industrial/ Facilities Parks Public Open State In RPA RPA Region Residential Commercial Community Recreation Rights-Region Of Way N.Y. Suffolk 279 (199)2 50 (35) 92 84 72 172 Sullivan [8]1 [4] [117] [106] [90] [217] Ulster [53] [12] [207] [188] [160] [386] Westchester 162 31 53 48 42 99 Bronx 25 4 3 3 2 5 Kings 42 7 4 4 4 8 New York 14 2 1 1 1 3 Queens 65 11 7 7 6 13 Richmond 39 7 3 2 2 5 P.A. Pike [7] [1] [76] [69] [59] 142 Total RPA 2040 368 7943* 7243* 6173* 14823* Region3 Total Consolidated 2078 383 1385 1261 1073 2583 Edison Area Notes:

1. Figures in brackets are for those counties outside RPAs Region. They are added to the total for Con Eds area.

2. Figures in parentheses are those portions of the RPA Region contained in the Con Ed area.

3. Total RPA Region figures followed by

• indicate that only the portion of the counties in Con Eds area are included.

Chapter 2, Page 27 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.4 FIGURES Figure No. Title Figure 2.4-1 Schematic Sector/Zone Diagram Figure 2.4-2 Indian Point Station Ten and Fifty Mile Radius Map Figure 2.4-3 Five Mile Sector/Zone Diagram [Historical] Figure 2.4-4 Ten Mile Sector/Zone Diagram [Historical] Figure 2.4-5 Fifty Mile Sector/Zone Diagram [Historical] Figure 2.4-6 Map and Description Showing Land Usage [Historical] Figure 2.4-7 Map and Description of the Area Showing Public Utilities Figure 2.4-8 Map and Description of the Area Showing Sewage Systems Chapter 2, Page 28 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.5 HYDROLOGY The hydrologic features of the Indian Point site are relevant to the analysis of radioactive liquid discharges from the plant. These features are the Hudson River, ground water and wells, and surface-water reservoirs. During normal plant operation liquid wastes are discharged to the Hudson River through the circulating water discharge canal. Ground-water contamination from accidental ground seepage or leakage from the plant will flow to the river because of the higher elevation of the plant relative to the river. Between 2005 and 2007, GZA GeoEnvironmental (GZA), performed a comprehensive hydrogeologic investigation of the site. This investigation was initiated to understand groundwater flow and contaminant transport. During this investigation numerous borings were advanced to study the site geology, hydrology and aquifer properties. Details of the geology, hydrology and aquifer properties can be found in the GZA report5. The hydrology in the environs of the Indian Point site has been extensively studied for Con Edison by numerous consultants, augmenting the data base established through the investigations of various governmental agencies. The initial Con Edison study was conducted in 1955 by Kennison,1 who analyzed the flow characteristics of the river at the site. Metcalf and Eddy2 further examined the river flow, and also investigated local groundwater hydrology and surface-water reservoirs. The salient aspects of these and other studies3,4 are reported below. The Hudson River below the dam at Troy (immediately below the confluence of the Hudson and Mohawk Rivers) is a tide-influenced, estuarine waterway. (see Figure 2.5-1.) Fresh water from the combined Hudson and Mohawk Rivers, as well as from numerous tributaries, discharges directly into the tidal portion of the river. Seawater enters the extreme lower reaches of the river through the Narrows and the Harlem/East River. The distribution of saltwater is influenced by fresh water flow, tides, physical characteristics of the river channel, and weather. Flow in the Hudson River is controlled more by the tides than by the runoff from the tributary watershed. River width opposite the plant ranges from 4500 to 5000-ft. Water depths within 1000-ft of the shore near the site are variable with an average depth of 65-ft; at some points the depth exceeds 85-ft. River cross-sectional areas in the vicinity of the site range from 165,000 to 170,000-ft2. Tidal flow past the plant is about 80 million gpm about 80-percent of the time, and it has been estimated that this frequency flow is at least 9 million gpm in a section 500-600-ft wide immediately in front of the facility. Mean tidal flow in the vicinity of the site is over 70 million gpm. The average downstream flow (for a 17-year period preceding 1930) is in excess of:2

• 11,700,000 gpm 20-percent of the time.
• 6,800,000 gpm 40-percent of the time.
• 4,710,000 gpm 60-percent of the time.
• 3,100,000 gpm 80-percent of the time.
• 1,800,000 gpm 98-percent of the time.

The plant is designed to use the dilution characteristics of the large tidal flow and will be operated such that discharges into the river would not contravene regulatory limitations. Historical flow patterns were further examined by Quirk, Lawler, and Matusky3,4 who reported both long-term (monthly) river discharges and potential drought flows. Quirk, Lawler, and Matusky also Chapter 2, Page 29 of 119 Revision 22, 2010

IP2 FSAR UPDATE analyzed and reported on the hydraulic conveyance properties of the estuary and the effects of tide and salinity on movement in the estuary. Review of historical records indicates that flooding at the site is non-existent. Flood stages are primarily the effect of tidal influence, with the secondary influence of runoff. The highest recorded water elevation in the vicinity of the site was 7.4-ft above mean sea level (MSL), which occurred during an exceptionally severe hurricane in November 1950. Since the river water elevation would have to reach 15-ft 3-in. above MSL before it would seep into any of the Indian Point buildings, the potential for any flooding damage at the site appears to be extremely remote. Seven different flooding conditions governing the maximum water elevation at the site were investigated, including the following:

1. Flooding resulting from runoff generated by a probable maximum precipitation over the entire Hudson River drainage basin upstream of the site.

2. Flooding caused by the occurrence of any upstream dam failure concurrent with heavy runoff generated by a standard project flood.

3. Flooding due to the occurrence of a probable maximum hurricane concurrent with a spring high tide in the Hudson River.

The severest flooding condition revealed by the study results from the simultaneous occurrence of a standard project flood, a failure of the Ashokan Dam and a storm surge in New York Harbor at the mouth of the Hudson River resulting from a standard project hurricane. The water level under these conditions would reach 14-ft above MSL. Local wave action due to wind effects has been determined to add 1-ft to the river elevation producing a maximum water elevation of 15-ft above MSL at the Indian Point Site. Since this maximum water elevation is 3-in. lower than the critical elevation of 15-ft 3-in. noted earlier, it is reasonable to conclude that flooding in the Hudson River will not present a hazard to the safe operation of Indian Point. The three most severe hurricanes to hit New York Harbor (September 21, 1821; November 25, 1950 (mentioned previously); and September 12, 1960) produced tidal surges at the Battery of 11-ft, 8.2-ft and 6.3-ft, respectively. Accordingly, these surges would appear as 7.5-ft, 5.5-ft, and 4.3-ft surges at Indian Point. The 5.5-ft surge predicted for the November 25, 1950, hurricane agrees well with the actual surge that produced the 7.4-ft-high watermark recorded for Indian Point on that date. The Quirk, Lawler and Matusky report indicated that the combination of the maximum probable runoff, upstream dam failures and maximum ebb tide in the Hudson River is a less severe condition than the one postulated above. This latter scenario would cause the water level at Indian Point to be 11.7-ft above MSL, also below the critical control elevation. The report also indicates that the combination of probable maximum hurricane, spring high tide, and wave run-up will cause the water level at Indian Point to reach 14.5-ft above MSL. This is also below the critical control elevation of 15-ft-3-in. Table 2.5-1 summarizes the Indian Point water surface elevations resulting from the various combinations. Chapter 2, Page 30 of 119 Revision 22, 2010

IP2 FSAR UPDATE In view of the recorded hydrologic history of the Hudson River and New York Harbor and the predicated maximum hurricane surge at Indian Point, flooding at the site is a highly unlikely possibility. Within a 5-mile radius of the plant only one municipal water supply uses ground water. Other wells in this area are used for industrial and commercial purposes. The rock formations in the area and elevations of wells relative to the plant are such that accidental ground leakage or seepage percolating into the ground at Indian Point will not reach these sources of ground water, but will flow to the river. Only two reservoirs within a 5-mile radius are used for municipal water supplies. The first, Camp Field Reservoir, is the raw-water receiving basin for the system, which serves the city of Peekskill. This system uses the Catskill Aqueduct and Montrose Water District as alternative sources of water supply. The second reservoir, the impounding reservoir for the Stony Point water system, serves the towns of Stony Point and Haverstraw, and the villages of Haverstraw and West Haverstraw. The Stony Point system is connected to the Spring Valley Water Company to provide an alternative source of supply. A third reservoir within 5-miles of the plant, Queensboro Lake, supplies water to a state park area only. The location of these reservoirs, and others within a 15-mile radius of the site, are shown on Figure 2.5-2. The city of New York's Chelsea Pumping Station is located about 1-mile north of Chelsea, New York, on the east bank of the Hudson River, about 22-miles upstream of the site. Water will be pumped from intakes in the river at the rate of up to 100 million gal per day into the city reservoir system as required to supplement the primary supply from watersheds during severe drought conditions. This source, however, was not used during the recent 1981 drought. The discharge of any contaminant into a tidal estuary will result in its distribution throughout the estuary. Factors affecting this distribution include tide amplitude and current, river geometry, salinity distribution, and freshwater discharge. Quirk, Lawler, and Matusky investigated for Con Edison the influence of these factors and determined the effect of radioactive discharges on overall river concentrations, and specifically conditions at Chelsea Pumping station, as discussed in Section 11.1. During normal operations, the plant discharge will not exceed its maximum permissible concentration. Compliance with regulatory release limits is further discussed in Section 11.1. REFERENCES FOR SECTION 2.5

1. Letter report of Karl L. Kennison, Civil and Hydraulic Engineer, to G. R. Milne, Con Edison, November 28, 1955.

2. Metcalf and Eddy Engineers, Hydrology of Indian Point Site and Surrounding Area, October 1965.

3. Quirk, Lawler, and Matusky Engineers, Transport of Contaminants in the Hudson River above Indian Point Station, May 1966.

4. Quirk, Lawler, and Matusky Engineers, Evaluation of Flooding Conditions at Indian Point Nuclear Generating Unit No. 3, April 1970.

5. GZA, Hydrogeologic Site Investigation Report for the Indian Point Energy Center, January 7, 2008.

Chapter 2, Page 31 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 2.5-1 Water Surface Elevation at Indian Point Resulting From Stated Flow and Elevation Conditions Elevation at Indian Elevation at Elevation at Point Including the Battery - Flow at Indian Point - Local Oscillatory Component Flow Datum MSL Indian Point Datum-MSL Wave Height Datum at Indian Point (ft) (million cfs) (ft) MSL (ft)

1. Probable MSL 0.00 1.100 12.7 13.7 maximum flood

2. Probable High water 1.014 12.4 13.4 Maximum flood +/-2.20 and tidal flow

3. Probable Low water 1.165 13.0 14.0 Maximum flood 2.20 and tidal flow

4. Standard project MSL 0.00 0.705 7.2 8.2 flood and Ashokan Dam failure

5. Standard project Standard project 0.550 13.0 14.0 flood hurricane

                     +11.00

6. Standard project Standard project 0.705 14.0 15.0 flood and Ashokan hurricane Dam failure (+11.00)

7. Probable Probable maximum 13.5 14.5 maximum hurricane hurricane and +17.5 spring high tide Chapter 2, Page 32 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.5 FIGURES Figure No. Title Figure 2.5-1 Map & Description Showing Location of Sources of Potable & Industrial Water Supplies & Watershed Areas Figure 2.5-2 Hudson River Drainage Basin 2.6 METEOROLOGY 2.6.1 General [Historical Information] Meteorological parameters related to atmospheric transport and diffusion have been extensively investigated in the Indian Point area since 1955. Studies of the wind flow characteristics, induced by the topography surrounding the site, illustrated the unique valley wind system and the channeling of low level winds. Meteorological studies1-3 were conducted from 1955 to 1957 by the Research Division of New York University, under the direction of Prof. Ben Davidson, in support of Unit 1 licensing activities. Data from these studies illustrated the channeling of the air flow by the terrain into downvalley (north-northeast) and upvalley (south-southwest) regimes. Historical data collected by the U.S. Weather Bureau in 1932 also illustrated the valley wind system. Subsequent meteorological investigations were conducted from 1968 through 1972 by New York University School of Engineering and Science, Department of Meteorology and Oceanography, under the direction of Dr. James Halitsky and Mr. Edward J. Kaplin. These studies supported the earlier findings of the valley wind system by Prof. Davidson and are documented in Appendix 2A of this FSAR and in the FSAR for Indian Point Unit 3. The most recent meteorological programs and data analyses conducted in the Indian Point environs since 1972 were documented in a York Services Corporation report (Meteorological Update, September 1981). This report is included in Appendix 2A. The 10-m elevation on the 100-ft meteorological tower used for the Unit 2 siting studies is the backup tower to the 400-ft (122-m) primary tower. The 10-m tower installed at the Buchanan Service Center is also available as an additional contingency. The York Services Report summarizes the meteorological activities conducted for Indian Point from 1955 to 1981. Included are topographic effects, wind correlations, data collection, diurnal wind distribution, trajectory analyses, atmospheric stability, and wind distributions. The report substantiates previous studies conducted on the existence of the valley wind system in the environs of Indian Point. 2.6.2 Application of Site Meteorology to Safety Analysis of Loss-Of-Coolant Accident The atmospheric dispersion factors required for the safety analysis of Chapter 14 have been computed for the worst possible meteorological conditions that could prevail at the Indian Point site. A search of the records indicates that the most protracted consecutive period during which the wind direction was substantially from the same direction was 5 days. The winds in this case Chapter 2, Page 33 of 119 Revision 22, 2010

IP2 FSAR UPDATE were from the northwest and speeds ranged from 15 to 30 mph. Therefore, this case does not represent the most conservative meteorology associated with the loss-of-coolant accident. The most frequent wind flow at low heights under inversion conditions is down the axis of the valley. This direction, roughly 10- to 30-degrees, is also the direction of maximum wind frequency. Because of the relatively high frequency of inversion conditions associated with this wind direction, the safety analysis assumes that the distribution of wind speed and thermal stability during the hypothetical accident is exactly that measured at the 100-ft tower level for the 5- to 20-degree wind direction. The valley wind is diurnal in nature, that is, up-valley during unstable hours and down-valley during stable hours. In general, these local winds are most frequent under clear sky and relatively light prevailing wind conditions. The diurnal variation of the vector mean wind as measured 70-ft above river during September-October 1955 is shown in Figure 2.6-1 for conditions in which the large-scale flow was virtually zero (12 days) and in Figure 2.6-2 for conditions in which the large-scale flow (geostropic wind) was less than 16 mph (35 days). It may be seen that for these virtually stagnant prevailing wind conditions, there is a regular diurnal shift in wind direction and that the mean vector wind associated with the down-valley flow is approximately 6 mph. A measure of the magnitude of the diurnal shift in wind direction is shown in Figure 2.6-3, where the steadiness of the wind (vector) mean speed over the mean scalar speed is shown as a function of time and the strength of the prevailing flow. Where the steadiness is close to one, the persistence of a given wind direction is very high. These data indicate that a consecutive 24-hr down-valley flow with light wind speeds and inversion conditions is extremely improbable due to the diurnal variation of the steadiness. The safety analysis of the loss-of-coolant accident assumes that the accident occurs during down-valley inversion flow conditions and that this condition persists for 24 hr with average wind speeds slightly less than 2 m/sec. Figures 2.6-1 and 2.6-2 indicate that the duration of the down-valley flow is about 12 hr rather than 24 hr and that the vector mean wind speeds are approximately 2.5 m/sec. In view of the discussion above, it must be concluded that the safety analysis for the first 24 hr is conservative to within a factor of about 2. The remainder of the safety analysis assumes that for the next 30 days, 35-percent of the winds are in the 20-degree sector corresponding to the nocturnal down-valley flow and that wind speed and thermal stability are as observed over the period of 1 year as measured at the 100-ft tower location. If the observations were distributed uniformly throughout the year, slightly over 100 hr per month of 5- to 20-degree winds could be expected to occur. The analysis assumes that 276 hr of 5- to 20-degree winds occur in the first 31 days after the accident, and that about 130 of these hours are characterized by inversion conditions. Approximately 35 weak-pressure gradient days were observed in September-October 1955 or about 430 hr per month. From Figure 2.6-3, the hour during which the down-valley flow is quite persistent under weak-pressure gradient conditions are from 0 to 8 hr. Assuming that the steadiness is 1.0 during these hours (it is in fact about 0.9 or less), the number of down-valley inversion winds per month during September and October is on the order of 140 hr per month. This indicates that the meteorology assumed in the safety analysis beyond the first 24 hr is reasonable for the worst Chapter 2, Page 34 of 119 Revision 22, 2010

IP2 FSAR UPDATE months (September and October) and is undoubtedly conservative with varying degrees of conservatism for the remainder of the year. The inversion frequency assumed for the 30-day accident case is conservative because the evaluation is made from concurrent assumptions concerning the postulated meteorological conditions, namely:

1. Inversion conditions prevail for 42.4-percent of the time.

2. The wind direction is within a narrow 20-degree sector for 35-percent of the time.

This is equivalent to assuming that in the model 20-degree sector, the inversion frequency is 14.8-percent for the 30-day period. The observed annual maximum inversion frequency for a 20-degree sector is 6.2-percent (p. 29, Table 3-3, Section 1.6 of Reference 3). If we assume that the inversion frequency is spread uniformly throughout the year, almost 3 months worth of inversions in the model 20-degree sector are considered to occur in the first 31-day month after the accident. The assumption of uniform spread of inversion frequency over the year is examined above where an attempt is made to isolate those local meteorological conditions at Indian Point, which might yield the highest 30-day dose. It is concluded that the "worst" meteorological conditions are associated with the nocturnal down-valley flow, which is most frequent during September and October. REFERENCES FOR SECTION 2.6

1. New York University, Research Division, A Micrometeorological Survey of the Buchanan, N.Y., Area, NYU Technical Report 372.1, November 1955, which was Exhibit L-1, Docket 50-3, given in its entirety. The topography of the area surrounding the site is described and the effects of the topography on meteorological conditions are discussed. The types of data collected, the methods and frequencies of collection, the description and location of the equipment, and the general scope of the meteorological program are indicated in this report. Seasonal wind characteristics, including speeds, directions, and frequencies are tabulated.

2. New York University, Research Division, Evaluation of Potential Radiation Hazard Resulting From Assumed Release of Radioactive Wastes to the Atmosphere From Proposed Buchanan Nuclear Power Plant, Sections 1, 2, and 3 of NYU Technical Report 372.3, April 1957. This report was submitted to the NRC in its entirety as Exhibit L-5, Docket 50-3. These sections discuss the diffusive conditions and the climatological data of the site. The basis for evaluating the diffusion parameters selected for the safety analysis is given on pages 19 to 21. Section 3 contains tables of frequency distribution of diffusion classes and wind directions, and also wind roses.

3. New York University, Research Division, Summary of Climatological Data at Buchanan, N.Y., 1956-1957, NYU Technical Report 372.4, March 1958, was Exhibit L-6, Docket 50-3. This report summarizes the final meteorological testing at Indian Point.

Chapter 2, Page 35 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.6 FIGURES Figure No. Title Figure 2.6-1 Diurnal Variation of Mean Vector Wind for Virtually Zero Pressure Gradient Conditions Figure 2.6-2 Diurnal Variation of Mean Vector Wind for 24 Hr Periods of Weak Pressure Gradient Conditions Figure 2.6-3 Steadiness of Wind as a Function of Time of Day for Indicated Pressure Gradient Conditions 2.7 GEOLOGY AND SEISMOLOGY [Historical Information] The Indian Point site and surrounding area were studied in 1955 by Sidney Paige, consulting geologist, before the construction of Unit 1 In 1965, Thomas W. Fluhr, P.E., an engineering geologist, reviewed the geology of the site and made additional borings at the location of Unit 2. In 1982, a report by Woodward-Clyde Consultants was done to update Section 2.7 of the FSAR. The previous studies are listed in the reference list of the report. The report is included in Appendix 2B. Between 2005 and 2007, GZA GeoEnvironmental (GZA), performed a comprehensive hydrogeologic investigation of the site. This investigation was initiated to understand groundwater flow and contaminant transport. During this investigation numerous borings were advanced to study the site geology, hydrology and aquifer properties. Details of the geology, hydrology and aquifer properties can be found in the GZA report1. A seismic monitoring network exists in the vicinity of the site and data from this network is periodically evaluated. REFERENCES FOR SECTION 2.7

1. GZA, Hydrogeologic Site Investigation Report for the Indian Point Energy Center, January 7, 2008.

2.8 ENVIRONMENTAL RADIOACTIVITY Monitoring for environmental radioactivity in the vicinity of the Indian Point Station began in 1958, approximately 4 years before the operation of Unit 1. Measurements since that time have indicated that the present operation of Units 2 and 3 and the past operation of Unit 1 have had no significant effect on the environment. The monitoring program implements Section IV.B.2 of Appendix I to 10 CFR Part 50 and thereby supplements the radiological effluent monitoring program by verifying that the measurable concentrations of radioactive materials and levels of radiation are not higher than expected on the basis of the effluent measurements and the modeling of the environmental exposure pathways. Measurements of radioactivity in the environment are summarized in the Annual Radiological Environmental Operating Report, which is submitted annually as required by the plant's Technical Specifications. Chapter 2, Page 36 of 119 Revision 22, 2010

IP2 FSAR UPDATE Determinations of radioactivity in the environment are made regularly. Samples include drinking water from two nearby reservoirs and the New York City Aqueduct, river water, sediments, fish, shellfish, lake and river aquatic vegetation, land vegetation, soil from locations in the vicinity of the site, shoreline soil, air, and milk along with direct gamma radiation measurements in various locations. The overall objectives of the environmental monitoring program are as follows:

1. To establish a sampling schedule for Indian Point Units 1 and 2 that will recognize changes in radioactivity in the environs of the plant.

2. To ensure that the effluent releases are kept as low as is reasonably achievable (ALARA) and within allowable limits in accordance with 10 CFR 20.

3. To verify projected and anticipated radioactivity concentrations in the environment and related exposure from releases of radioactive material from the Indian Point site.

Results of environmental surveys conducted by Con Edison have been verified by the Bureau of Radiological Health Service of the New York State Health Department in previous years and presently, by the New York State Bureau of Environmental Radiation.1, 2 Environmental surveys have also been confirmed by Dr. Merril Eisenbud, Director of Environmental Radiation Laboratory, Institute of Environmental Medicine, New York University Medical Center, who has found that the levels of environmental radioactivity are associated with natural background and fallout of nuclear weapons testing.2 In a study of the radioactivity in the Hudson River, Mr. Sherwood Davis, Director, Bureau of Radiological Health Service, New York State Department of Health, et al., have concluded that the discharges from Indian Point Unit 1 "are a minute fraction of the federal limits."4 The above results were obtained in preoperational and operational periods of Units 1 and 2 in the late 1950s and in the 1960s. In the more recent years of the late 1970s, radiological impact evaluations have shown similar results. These evaluations of actual plant releases have been performed for inclusion in the effluent release reports and have shown that operation of the Unit 2 plant has had an insignificant impact on the environs. REFERENCES FOR SECTION 2.8

1. New York State Department of Health, Division of Environmental Health Services, Consolidated Edison Indian Point Reactor, Post Operational Survey, August 1965.

2. New York State Department of Health, Division of Environmental Health Services, Consolidated Edison Indian Point Reactor, Post Operational Survey, July 1966.

3. New York University Medical Center Institute of Environmental Medicine, Ecological Survey of the Hudson River: Progress Report No. 2, submitted to Chapter 2, Page 37 of 119 Revision 22, 2010

IP2 FSAR UPDATE Division of Radiological Health, USPHS, Contract PHS 86-95, Neg. 141, December 1966.

4. F. Cosolito, et al., Radioactivity in the Hudson River, Symposium on Hudson River Ecology, Hudson River Valley Commission of New York, October 4-5, 1966.

Chapter 2, Page 38 of 119 Revision 22, 2010

IP2 FSAR UPDATE NOTE: This information is classified as Historical Information APPENDIX 2A FACILITY SAFETY ANALYSIS REPORT (FSAR) CONSOLIDATED EDISON COMPANY OF NEW YORK, INCORPORATED INDIAN POINT NUCLEAR GENERATING UNIT NO. 2 METEOROLOGICAL UPDATE SEPTEMBER, 1981 YSC PROJECT + 01-4122 prepared by: EDWARD J. KAPLAN BRUCE R. WUEBBER Senior Scientist Staff Meteorologist ATMOSPHERIC SERVICES DEPARTMENT Chapter 2, Page 39 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE OF CONTENTS LIST OF FIGURES LIST OF TABLES 1.0 GENERAL 1.1 Historical 1.1.1 Introduction 1.1.2 Tower Siting and Instrumentation 1.1.2.1 Hudson River 1.1.2.2 General Topography 1.1.2.3 Site Configuration 2.0 METEOROLOGICAL PROGRAM AT INDIAN POINT 2.1 122m Meteorological Tower 2.1.1 Siting 2.1.2 Instrumentation 2.1.2.1 Sensor Configuration 2.1.2.2 Instrumentation Specifications 2.1.2.2.1 Climatronics F460 Wind Speed Transmitter 2.1.2.2.2 Climatronics F460 Wind Direction Transmitter 2.1.2.2.3 Climatronics TS-10 and TS-10WA Motor Aspirated Shields 2.1.2.2.4 Climatronics 100087, 100087 Temperature-Delta Temperature 2.1.2.2.5 Climatronics DP-10 Dew Probe 2.1.2.2.6 Climatronics 1000971 - Heated Rain - Snow Gauge 2.1.2.2.7 Data Collection Systems 2.1.3 Meteorological Support System 2.2 Data Log 2.2.1 Indian Point Tower IP3 2.2.2 122 Meter Meteorological Tower Chapter 2, Page 40 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.0 ANALYSES DATA 3.1 Indian Point Tower IP3 3.2 122M Meteorological Tower (IP4) 3.2.1 October 1, 1973 to August 31, 1974 3.2.2 August 1, 1978 - July 31, 1979 3.2.2.1 Surface Air Trajectories Analyses - Summary 3.2.3 March 1980 - December 1980 3.2.3.1 General 3.2.3.2 Wind Frequency Distributions 3.2.3.3 Diurnal Wind Distributions 3.2.3.4 Resultant and Concurrent Hourly Winds 3.2.3.5 Summary - Trajectory II Study 3.2.4 January 1, 1979 through December 31, 1980 3.2.4.1 Data Analyses 3.2.4.2 Wind Frequency Distributions 3.2.4.3 Diurnal Wind Direction Distributions 3.2.4.4 Wind Speed Distributions 3.2.4.5 Wind Velocities and Atmospheric Stability 3.2.4.5.1 Joint Frequency Distribution of Wind Direction and Stability 3.2.4.5.2 Frequency of Occurrence of Stability Categories 3.2.4.5.3 Average Wind Speed and Diurnal Variation as a Function of Stability Categories 4.0

SUMMARY

4.1 Meteorology 4.1.1 General 4.1.2 122 Meter Meteorological Tower System 4.1.3 Local Meteorological Characteristics 4.1.4 Conclusion

5.0 REFERENCES

Chapter 2, Page 41 of 119 Revision 22, 2010

IP2 FSAR UPDATE LIST OF FIGURES Figure No. Title Figure 1 Ground Contours at Elevation 200 Feet Figure 2 Ground Contours at Elevation 400 Feet Figure 3 Elevations in the Indian Point Region Figure 4 Water Courses in the Indian Point Region Figure 5 Existing and Historical Meteorological Towers at Indian Point Figure 6 Indian Point Meteorological Site Figure 7 Tower Configuration Figure 8 Station Configuration Figure 9 Indian Point - Meteorological Support Systems Figure 10A Two Station Wind Correlation Data Period - October 1973 Figure 10B Two Station Wind Correlation Data Period - December 1973 Figure 11 Position of One Mile Grid in Relation to Topographic Features Figure 12 Position of Wind Files on Grid Figure 13 Average March, 1980 East and West Bank Diurnal Wind Distributions Figure 14 Average June, 1980 East and West Bank Diurnal Wind Distributions Figure 15 Average December, 1980 East and West Bank Diurnal Wind Distributions Figure 16 Locations of Monitoring Sites in Relation to One Mile Grid Figure 17 Comparison of 10M Level Diurnal Wind Distributions Figure 18 Comparison of 122M Level Diurnal And Wind Distribution Figure 19 Diurnal Distribution of Wind Speeds Figure 20 Percent Probability Distribution of Wind Speeds Chapter 2, Page 42 of 119 Revision 22, 2010

IP2 FSAR UPDATE LIST OF TABLES Table 1 - Tower and Instrumentation Record Table 2 - Valid Data Log Table 3 - Comparison of Annual Percent Occurrence of Stability Categories Table 4 - Summary of Trajectory End-Points Table 5 - Summation of Trajectory End Points - August, 1978 Table 6 - Summation of Trajectory End Points - January, 1979 Table 7 - Summation Trajectory Occurrences South of Indian Point Table 8 - Locations of Stations Relative to Indian Point Table 9 - Valid Data for Trajectory Wind Sites Table 10 - Frequency Distribution of 24 Hour Resultant Wind Directions Table 11 - Summary of Two-Station Wind Correlations Piermont (Site 1) Referenced to Selected Monitoring Locations (Site 2) Table 12 - Concurrence of Two-Station Wind Directions Table 13 - Diurnal Distribution of Occurrences of Eight-Hour Trajectories With On Grid Reversals Table 14 - Summary of Trajectory End-Point Counts Table 15 - Summary of Trajectory End-Points (Percent) Table 16A - Historical Comparisons of Wind Frequency Distributions - March Table 16B - Historical Comparisons of Wind Frequency Distributions - July Table 16C - Historical Comparisons of Wind Frequency Distributions - December Table 17 - Comparison of Percent Wind Frequency Distributions - Summer Table 18 - Comparison of Percent Wind Frequency Distributions - Winter Table 19 - Comparison of Diurnal Resultant Wind Directions Table 20 - Indian Point (10M) Wind Speed (MPH) - Summer Season Table 21 - Indian Point (10M) Wind Speed (MPH) - Winter Season Table 22 - Indian Point (122M) Wind Speed (MPH) - Summer Season Table 23 - Indian Point (122M) Wind Speed (MPH) - Winter Season Table 24 - Maximum Diurnal Wind Speed (MPH) Table 25 - Annual Summary of Wind Direction Percent Frequency Distribution As A Function of Stability - 10M Level Table 26 - Summary of Wind Direction Percent Frequency Distribution As A Function of Stability - Summer Season Table 27 - Summary of Wind Direction Percent Frequency Distribution as A Function of Stability - Winter Season Table 28 - Historical Comparisons of Percent Occurrence of Stability Table 29 - Comparison of Percent Occurrence of Stability on 122 Meter Tower Table 30 - Diurnal Variation of Stability Class and Wind Speed (10M) Table 31 - Diurnal Variation of Stability Class and Wind Speed (122M) Table 32 - Diurnal Variation of Stability Class and Wind Speed (Delta-T 400-200) Table 33 - Comparisons of Average Wind Speeds (MPH) As A Function of Stability Chapter 2, Page 43 of 119 Revision 22, 2010

IP2 FSAR UPDATE FACILITY SAFETY ANALYSIS REPORT (FSAR) CONSOLIDATED EDISON COMPANY OF NEW YORK, INC. INDIAN POINT NUCLEAR GENERATING UNIT 2 METEOROLOGICAL UPDATE 1.0 GENERAL 1.1 HISTORICAL BACKGROUND 1.1.1 Introduction Meteorological data were initially collected and evaluated with respect to Indian Point, Buchanan, New York and its environment during the period from 1955 through 1957. This work was accomplished under the direction of Professor Benjamin Davidson of New York University under contract with the Consolidated Edison Company of New York, Inc. (Con Edison). The data and studies during this period were the bases for the Environmental Reports relevant to Indian Point Nuclear Generating Units 1, 2 and 3. With respect to the Facility Safety Analysis Report (FSAR) for Unit 2, the Environmental Report Supplement, Appendices Volume 1 as Appendices C, D and E contains:

• NYU Technical Report 372.1 (November, 1955), B. Davidson
• NYU Technical Report 372.3, Section 2 and 3, (April, 1957), B. Davidson and J.

Halitsky

• NYU Technical Report 372.4 (March, 1958), B. Davidson In 1968 under the direction and supervision of Dr. James Halitsky of New York University, Con Edison contracted to establish experimental meteorological monitoring, data collection and evaluation activities at the Indian Point site and at specified sites in its environment (Halitsky, Laznow and Leahy, February 1970). The original purpose of the above investigation, as noted in the reference, was modified after the studies had begun in order to provide the AEC Construction Hearings for Indian Point Unit 3 with clarification of aspects of the 1956-1957 meteorological data base for the Units 1, 2 and 3 diffusion models. This phase of data collection began in December, 1969.

A report dealing with the results of this latter phase (Halitsky, Kaplin and Laznow, NYU GSL Technical Report No. TR 7103, May 1971) appears as Appendix G in the FSAR Unit 2 Environmental Report Supplement Appendices Volume 1 and as Supplement 1 in the FSAR for Unit 3. The focus of the above report was to validate present site meteorology as representing no significant change in relation to site meteorology from the 1955-1957 period. Data collection and evaluations continued under this program and a report was submitted by Kaplin and Laznow (1972) representing the data collection period from 1 January, 1970 through 31 December 1971. A copy of this report appears in FSAR for Unit 3 as Supplement 10 (January, 1973). With respect to Indian Point on-site meteorological measurements, there were for the purpose of the reports that have been cited, three different meteorological towers at three different locations. These are specifically delineated in Figure 1 of Halitsky, Kaplin and Laznow (1971). Chapter 2, Page 44 of 119 Revision 22, 2010

IP2 FSAR UPDATE The meteorological data collected during the 1955-1957 were from the 300 Foot Tower designated as IP1. Meteorological data collected and reported by Halitsky, Kaplin and Laznow (1971) and Kaplin and Laznow (1972) were from the 100 Foot Tower designated as IP3. The base of IP3 Tower was about 200 feet from the base location of the original IP1 Tower. Using input data meteorological data from Indian Point Tower (IP3), Bowline Point and the Cape Charles along with sequences of upper air pilot balloon observations Kaplin, Laznow and Wurmbrand (1972) provided Con Edison with input information for the location of the 90-percent probability air monitoring sites, overlay patterns for the prediction of the distribution of gaseous releases and evaluation of the requirements of the AEC Safety Guide 23 (1972). All data on the IP3 Tower were collected in accordance with U.S. AEC Safety Guide No. 23, On-Site Meteorological Programs as delineated 2/17/71. The IP3 Tower was maintained from March 1972 through December 1973 by York Research Corporation, Stamford, Connecticut under contract with Con Edison. The last formal report for this tower was prepared and submitted to Con Edison by Kaplin and Kitson (1974) for the 1973 data collection period. In April of 1973, York Research Corporation under contract with Con Edison began work on the erection of an on-site 400 Foot Meteorological Tower approximately 1725 feet-S and 1750 feet-W of the IP3 Tower. The function of this tower was to develop micro-climatological data suitable for the design of cooling towers and the evaluation of their potential environmental impact on the Indian Point site and its environs. Concurrent studies were conducted to develop three dimensional aspects of the local valley flow using pibal balloons, constant level tetroons and balloon-sondes. In addition, a concurrent study was conducted to develop background levels of ambient air salt concentrations. The results of these studies were submitted in two reports: Kaplin, Kozenko and Kirshner (1974) and Kaplin, Kitson and Kozenko (1974). This latter report compared meteorological data from the IP3 Tower. At the conclusion of 1973, the primary source of reduced on-site meteorological data were from the IP4, 400 Foot Meteorological Tower. IP3 Tower systems maintenance was continued in accordance with Safety Guide 23 through October 1, 1976 and meteorological data were recorded on analog charts. Data collection was transferred to the IP4 Tower on July 1, 1976. The 400 Foot Tower (IP4) servicing, maintenance and data collection and data selective processing is on-going. Its systems have been updated to meet present requirements of NUREG-0654 Appendix 2 (1980) and proposed Revision 1 to NRC Safety Guide 1.23 (1980). The continued maintenance services, etc., of the 400 Foot Meteorological Tower by York Research Corporation/York Services Corporation was continuous under contract with Con Edison until September 30, 1978, and from October 1, 1978 through the present time under contract with the Power Authority of the State of New York (PASNY). In September, 1979, York Services Corporation under contract with Con Edison began a study of north to south surface air trajectories analyses and evaluations based on real-time wind data available from the Indian Point vicinity. This study incorporated local wind velocity data from the 400 Foot Meteorological Tower at Indian Point, the Orange and Rockland Utilities, 350 Foot Meteorological Tower in Haverstraw, New York, as well as from selected U.S. Department of Commerce, NOAA, Weather Stations. For this study, meteorological data were analyzed and evaluated for the period from August 1, 1978 through July 31, 1979. The final results of this study were presented in reports: Kaplin, E.J. and B. Wuebber, (1979) and Kaplin, E.J. (1979). As an outgrowth of these studies an expended network of surface (10M) wind velocity monitoring stations were sited at key locations along the Hudson River Valley north and south of Chapter 2, Page 45 of 119 Revision 22, 2010

IP2 FSAR UPDATE Indian Point and inland to the east. This network consisted of new anemometer stations in addition to the Indian Point 400 Foot Meteorological Tower, the Bowline 350 Foot Meteorological Tower and the U.S. Department of Commerce, NWS Station at Westchester County. These wind data were digitalized and evaluated by New York Services Corporation under contract to Con Edison for the purposes of defining surface air flow patterns within a 10-15-mile range of Indian Point with emphasis on generating refined estimated of southward movements. As completed, ten consecutive months (March 1, 1980 - December 31, 1980) were evaluated and a total of 7,264 eight-hour parcel trajectories were created objectively using appropriate local one-hour wind velocity averages on a real time basis. The results of this study were submitted to Con Edison: Kaplin, Edward J. and B. Wuebber (1981). For the purpose of this FSAR, second meteorological update, meteorological data have been analyzed and evaluated from the 400 foot (122 Meter) Indian Point Meteorological Tower for the two year period from January 1, 1979 through December 31, 1980. 1.1.2 Tower Siting and Instrumentation 1.1.2.1 Hudson River There are a number of pertinent facts about the Hudson River itself that are relevant to its ability to induce and/or influence mesoscale flow phenomena that are dominant factors in the Indian Point environs. The most important factor is that it is not a river but, rather a tidal estuary. From New York City, 154-miles north to Troy, there is no drop in the surface elevation of the river. Except for spring runoff from the Andirondacks, which can smother the tide down to Albany, there is almost imperceptible downstream current. Since there is no slope to the river surface, it will not support its own gravity flow. Any air movement within its canyons during minimal atmospheric pressure gradient periods can be strictly local cells, which may actually block continuous horizontal air movement over the water surface. Thermally induced air movement of the Atlantic sea breeze follows the natural path of the river. It has been noted, however, that Iona Island 45-miles north of the tip of Manhattan is considered the point of maximum inland intrusion. The northward movement of sea breeze does not proceed up the Hudson River Valley and Hackensack River Valley at the same speed. The inland movement along the Hackensack Valley lags the Hudson Valley movement. The Hackensack River is on the west side of the Hudson and is specifically delineated because its headwaters are just south of the South Mountains and isolated from the Hudson River by the Hook Mountains and the Palisades. The South Mountains are the east-west extension of the Hook Mountains. The South Mountains abut the Ramapo Range and form a sheer wall from the Southern boundary of the west bank community of Haverstraw. 1.1.2.2 General Topography Each of the reports cited in the Section 1.1.1 describe in some detail general topography in the Indian Point environs. The most recent was provided by Kaplin and Wuebber, 1981. Indian Point is located in the lower Hudson River Valley 27-miles due north of northern boundary of New York City (Manhattan Island). Chapter 2, Page 46 of 119 Revision 22, 2010

IP2 FSAR UPDATE The Indian Point area has been described by Halitsky, et. al., 1970, as being located roughly on the axis of a north-south valley enclosed by the Dunderberg and Buckberg Mountains to the west and Blue Mountain and Prickly Pear Hill on the east. The shape of the valley at the 200 foot and 400 foot elevation levels are given in Figures 1 and 2. At the 200 foot contour level the valley width is two miles at Dunderberg Mountain and opens southward to a width of five miles at Prickly Pear Hill. The Hudson River, flowing southward through the valley, resembles a gourd with its curved 3/4-mile thick northern neck nestling at the base of Dunderberg Mountain while the bulbous three mile thick body fills the southern part of the valley between South Mountain and Prickly Pear Hill. The Indian Point peninsula lies in the hollow of the curved neck. Beyond its northern end, the valley is split into two branches by Manitou Mountain. The Hudson River passes through the steep, narrow northwest branch between Manitou and Dunderberg Mountains. The northeast branch, between Manitou and Blue Mountains, is about 1.5-miles wide at Manitou Mountain but degenerates with distance into three tributary valleys containing Annsville Creek, Sprout Brook and Peekskill Hollow Brook with sources in the mountainous region north of Peekskill. South of Haverstraw Bay, the valley opens up rapidly to the southeast while the west bank of the Hudson River follows the blocking of the east-west orientated South Mountains to assume a southward course along the Hook Mountains to the Palisades Mountains. At elevations higher than 200 feet the solidity of the eastern wall of the valley breaks, first between Blue Mountain and Prickly Pear Hill to form at 300 feet the irregular drainage system, which supplies Furnace Brook, and then at 400 feet into an irregular array of mountain tops. The western wall is till fairly solid at 300 feet but breaks at 400 feet into two well-defined valleys containing Cedar Pond and Minisceongo Creeks. However, just to the west are Ramapo Mountains whose elevations exceed 1000 feet. Figures 3 and 4 (Kaplin and Wuebber, 1981) show the elevations of significant mountain peaks and water courses in the region of Indian Point. 1.1.2.3 Site Configuration The location of the specific meteorological towers associated with earlier studies in the Indian Point environs are available in Figure 1 of Kaplin and Laznow (1972), FSAR 3, Supplement 10, 1973 and in Figures 1a and 1d of Kaplin, Kitson and Kozenko (1974). Table 1 lists the operational periods and the instrumentation associated with each tower. Not included in this listing are those wind monitoring sites that were used for the most recent study (Kaplin and Wuebber 1981, Sec. 2.4 & 2.5). In this study, an 11 site monitoring network was established equipped with Climatronics Mark III, wind speed and direction systems. This network was operable for all or part of the period from March, 1980 through December, 1980. Chapter 2, Page 47 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1 TOWER AND INSTRUMENTATION RECORD (INCLUDES PARAMETERS NOT REQUIRED BY PROPOSED SAFETY GUIDE 1.23) Meteorological Base Elevation Operational Period Exposure Station Ft. MSL From To Parameter Instrument M. Above Grade Indian Pt (IP1) 130 1956 1957 Wind Aerovane 91 & 30 Temp. Diff. Honeywell 91-2 & 30-2 USS Jones (J) 0 1956 1957 Wind Aerovane 21 Indian Pt (IP2) 60 1968 1969 Wind Climet 30 Temp. Diff. Bristol 29 - 1.5 Montrose (MP) 60 1968 6/71 Wind Climet 30 Bowline Pt (BP) 5 9/68 11/69 Wind Climet 30 11/69 8/72 Wind Aerovane 30 9/68 11/69 Temp. Diff. Honeywell 30 - 3 11/69 2/72 Temp. Diff. Bristol 30 - 3 2/72 8/72 Temp. Diff. Climet- 30 - 3 (Rosemont) Bowline Tower 10 Note 2 Present Wind Climatronics 100, 50 & 10 Note 2 Present Temp. Diff. Climatronics 100-10 & 50-10 Trap Rock (TR) 90 1969 7/72 Wind Climet 30 USS Cape Charles 0 3/70 9/70 Wind Aerovane 30 (CC) Indian Pt (IP3) 120 11/69 9/76 Wind Aerovane 30 11/69 9/76 Wind Climet 30 6/73 9/76 Wind Climet 10 Backup Met 12/81 Present Wind Climatronics 10 System (IP3) 11/69 10/71 Temp. Diff. Honeywell 29 - 2 8/72 9/76 Temp. Diff. Climet- 30-3 & 9-3 (Rosemont) 8/72 9/76 Amb. Temp. Climet- 9 (Rosemont) 8/72 9/76 Dew Point Climet- 30,9 & 3 (Foxboro) 8/72 9/74 Net Radiation Teledyne 9 Geotech 5/70 12/70 Turbulence Bivane* 30 Note 2 - Bowline Tower is located at approximately Latitude 41° 13N and Longitude 73° 58 W. This location is about 3000 feet NW of the earlier Bowline Point Tower. It began operation in the 1972/73 time period.

• Intermittent usage.

Chapter 2, Page 48 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1 (Contd) Meteorological Base Elevation Operational Period Exposure Station Ft. MSL From To Parameter Instrument M. Above Grade Indian Pt (IP4) 117 9/73 Present Wind Climatronics 122, 38 & 10 122 M. Tower 9/73 6/80 Wind Climatronics 85 - Note 1* 6/80 Present Wind Climatronics 60 9/73 9/79 Amb. Temp. EG & G 10 9/79 Present Amb. Temp. Climatronics 10 9/73 9/79 Dew Pt. EG & G 122, 61 & 10 9/79 Present Dew Pt. Climatronics 10 9/73 9/79 Temp. Diff. EG & G 122-10 & 61-10 9/79 Present Temp. Diff. Climatronics 122-10 & 60-10 1/74 Present Net Rad. Teledyne 10 Geotech 7/80 Present Precipitation Climatronics 1 Indian Pt (IP4) ** 117 9/73 7/77 Visual Range EG & G FSM 10 10M Tower Emergency 135 7/24/80 11/81 Wind Climatronics 11.8 Control Center # Note 1* - 85 meter wind speed and wind direction moved from the 85 meter level to the 60 meter level as required by proposed Revision 1 to NRC Safety Guide 1.23.

• • Tower and System Removed 07/22/80.

1. Tower and System Removed 07/22/80.

Chapter 2, Page 49 of 119 Revision 22, 2010

IP2 FSAR UPDATE 2.0 122M METEOROLOGICAL TOWER 2.1.1 Siting The relative locations of the existing meteorological towers in historic perspective are shown in Figure 5. Specific details of site location are shown in Figure 6. 2.1.2 Instrumentation 2.1.2.1 Sensor Configuration The sensor configuration and exposure on the existing operational 122M Meteorological Tower are shown in Figures 7 and 8. 2.1.2.2 Instrumentation Specifications The following specifications apply to specific operational sensors that are a part of the total meteorological support systems at Indian Point. 2.1.2.2.1 Climatronics F460 Wind Speed Transmitter Accuracy: 0.07 M/S or 1% Range: 0-56 M/S Threshold: 0.22 M/S Distance Constant: 1.5 M 2.1.2.2.2 Climatronics F460 Wind Direction Transmitter Accuracy: +/-2° Range: 0-540° Threshold: 0.22 M/S Distance Constant: 1.5 M Damping Ratio: 0.4 at 10° initial angle of attack 2.1.2.2.3 Climatronics TS-10 and TS-10WA Motor Aspirated Shields Shield Effectiveness: Under radiation intensities of 110 W/m3 (1.6 cal/cm2/min) radiation error not exceeding 0.1°C Aspiration Rate: 3 M/S at sensor location 2.1.2.2.4 Climatronics 100087, 100087 Temperature-Delta Temperature (matched thermistor) Temperature: Range: -34 to +50°C Accuracy: +/- 0.2°C Time Constant: 10 sec. To 63% (in TS-10 Shield) Linearity: +/- 0.2% Chapter 2, Page 50 of 119 Revision 22, 2010

IP2 FSAR UPDATE Delta Temperature: Range: +/- 10°F Sensitivity: 0.02°F Accuracy: 0.1°F or +/- 5% of delta-T not to exceed 0.3°F Response Time: 10 sec. To 63% in TS-10 Shield 2.1.2.2.5 Climatronics DP-10 Dew Probe (YSI Lithium Chloride) Range: -40° to 42°C Accuracy: +/- 0.5°C Response Time: 1°C/min. 2.1.2.2.6 Climatronics 1000971 - Heated Rain - Snow Gauge (Tipping Bucket) Accuracy: +/- 1% up to 3/hr Resolution: 0.01 Size: 8 diameter x 24 height Conversion Accuracy: +/- 0.2% 2.1.2.2.7 Data Collection Systems Analog: Wind Systems: Esterline-Angus Model E1102R - Rectigraph Recorders Temperature, Dew Point, Delta Temperature: Tracer Westronics Model M11E, Multipoint Precipitation: Esterline-Angus Model MS 401C Digital: Climatronics Data Processor: 1MP/801 Tape Collection Interface: Tandeberg TD1 10-50 2.1.3 Meteorological Support System The meteorological systems at Indian Point are equipped, maintained and operated in compliance with the specification of NUREG-0654, Appendix 2 (1980); Proposed Revision 1 to NRC Regulatory Guide 1.23 (1980); and applicable regulatory requirements. The total system as presently operated is outlined in Figure 9. 2.2 DATA LOG 2.2.1 Indian Point Tower IP3 Meteorological data from the IP3 Tower were reduced and evaluated (Kaplin and Kitson, 1974) through December 1973. From the period 1974 through September 1976 the tower system was maintained, as previously noted. Analog records were provided to Con Edison for storage. Tower removed from service in September, 1976. (Reactivated as site for Backup Wind System: 12/01/81.) 2.2.2 122 Meter Meteorological Tower (IP4) Chapter 2, Page 51 of 119 Revision 22, 2010

IP2 FSAR UPDATE The data log for the 122 Meter Meteorological Tower for the period from October 1973 through August 1974 can be found in Kaplin, et. al., 1974. Subsequent to the completion of the above report and the data contained therein, the meteorological analog charts and the data collection magnetic tape were documented and transmitted to Con Edison for storage. Commencing in August 1977, wind velocity data at the 10 and 122 meter levels and the delta temperatures: 60-10M and 122-10M were reduced to hourly averages and transmitted to Con Edison in addition to the analog charts. The summary of valid data for these parameters for the period from August 1977 - July 1981 is shown in Table 2 on concurrent and total hours basis. The concurrent basis assumes that if any parameter is missing. The total basis relates an individual missing data hour to the total number of possible data hours in a month. On the concurrent basis, the average valid data collection was 92.4 +/- 10.8-percent. On a total hour basis, the average valid data collection was 98.2 +/- 2.5. Chapter 2, Page 52 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 2 VALID DATA LOG* 1977 1978 1979 1980 1981 Month Concurrent Total Concurrent Total Concurrent Total Concurrent Total Concurrent Total January N/A N/A 89.8 98.2 94.8 99.0 98.4 99.7 99.9 99.9 February N/A N/A 92.1 97.4 98.1 99.7 90.5 98.4 67.7 89.2 March N/A N/A 95.1 98.9 97.4 99.6 97.3 99.1 98.0 99.3 April N/A N/A 98.1 99.6 96.7 99.4 100.0 100.0 100.0 100.0 May N/A N/A 95.3 98.6 90.3 98.0 88.4 98.0 100.0 100.0 June N/A N/A 86.8 95.6 95.0 99.1 96.7 99.0 100.0 100.0 July N/A N/A 94.1 96.8 92.6 98.8 71.1 94.1 99.4 99.8 August 94.2 98.7 95.3 99.2 52.3 92.0 77.7 92.7 September 84.7 94.4 99.7 99.9 57.5 92.3 98.3 99.7 October 95.0 98.9 98.1 99.4 94.0 98.2 96.1 99.4 November 98.9 99.5 98.8 99.6 98.3 99.7 99.6 99.9 December 75.3 95.5 96.4 99.4 100.0 100.0 98.9 99.8 Concurrent Average: 92.4 +/- 10.8% Total Parameter Hours: 98.2 +/- 2.5%

• Based on six (6) parameters: wind data at 10 and 122M and delta-temperatures: 60-10M and 122-10M NA - Values not available.

Chapter 2, Page 53 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.0 ANALYSES DATA 3.1 INDIAN POINT TOWER IP3 A FSAR documented study with respect to the Indian Point IP3 Meteorological Tower was prepared by Kaplin and Laznow (1972) [Referenced FSAR 3 Supplement 10, 1973]. This report covered the data collection period through 1971. Additional data were subsequently provided through 1972 to provide composite joint wind velocity frequency distribution for Pasquill Stability Categories (FSAR 3 Supplement 13 and 16, 1973). Kaplin and Kitson (1974) provided an analyses of IP3 for the period March 1973 through December, 1973. This report confirmed the earlier study that wind data in the Indian Point environs based on monthly diurnal wind distributions, wind frequency distributions and joint wind stability categories are comprised of two seasons with little apparent transition. The winter season reflects little or no average diurnal variation in the hourly resultant winds, dominant winds from the west to north. The summer season is characterized by dominant north-northeast winds during the evening and early morning hours with a sharp transition to south to southwest winds during the day and another transition in late afternoon to the evening pattern. The wind frequency distributions and joint frequencies as a function of Pasquill Stability Categories were comparable in 1973 with data collected in 1970 and 1971. It is noted that the temperature gradients on the IP3 Tower were derived from delta-temperatures: 99-7 feet and wind measured at 105 feet above a grade elevation of 120 feet MSL. Kaplin, et. al., (1974) compared three months (October - December, 1973) of IP3 wind data as measured at 105 feet above grade with concurrent three months of data from the 125 foot level on the 122 Meter Meteorological Tower (IP4) (grade elevation: approximately 117 feet MSL) using a two station wind correlation program (Appendix B, Kaplin, et. al, 1974). Figures 10A and 10B show the relationships obtained for October and December, 1973. The November results were similar with the directional relationships falling between that obtained for October and December. The maximum variations between the two sites occurred with winds from E-ESE and SW to WSW. These corresponded to sectors of minimum average wind speeds. Deviations between the two sites can be attributed to local factors including terrain elevation, land use and ground cover. The wind direction displacement effects found in the two station correlations were confirmed in the monthly diurnal analyses. 3.2 122M METEOROLOGICAL TOWER (IP4) 3.2.1 October 1, 1973 to August 31, 1974 The purpose of the 122M Meteorological Tower at Indian Point (IP4) was to develop a three dimensional micro-climatological data file to be used to assess the impact of proposed cooling towers and to provide the basis for design criteria as required. The results of one year of operation of this tower were presented in a final report (with Appendices) by Kaplin, et. al., 1974. Chapter 2, Page 54 of 119 Revision 22, 2010

IP2 FSAR UPDATE Except as noted in the previous section, the meteorological data collected and evaluated were not compared at the time of this study with historical meteorological data associated with FSAR 2. This study determined that the two distinctive seasonal patterns existed at each of the four levels of wind velocity measurement: 10M, 38M, 85M and 122M. Wind directions tended to back with elevation assuming an orientation parallel to general terrain contours. In times of weak synoptic pressure gradient patterns, there were abrupt transitions in the diurnal flow patterns consistent with valley flow winds particularly during the summer season. These transitions began at the surface and progressed up to the 122M level. Chapter 2, Page 55 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3 COMPARISON OF ANNUAL PERCENT OCCURRENCE OF STABILITY CATEGORIES Stability Category Year Tower Gradient (M) A B C D E F G 1970 IP3 29 - 3 21.68 2.20 3.39 33.35 24.75 9.01 5.62 1971 IP3 29 - 3 19.17 2.75 2.97 22.74 30.87 11.69 9.75 1970-72 IP3 29 - 3 16.25 1.82 2.95 29.71 26.61 13.27 9.45 1970-72* IP3 29 - 3 6.76 2.67 2.13 32.65 40.57 11.78 3.31 1973 IP3 29 - 3 23.14 3.16 3.70 20.87 25.02 13.89 10.23 1973-74 122M-IP4 60 - 10 10.35 3.21 2.94 25.38 44.86 11.35 1.91 1979 122M-IP4 60 - 10 12.27 3.25 3.86 29.30 40.39 8.83 1.31 1980 122M-IP4 60 - 10 13.32 4.06 4.60 29.81 33.97 11.34 2.07 1979-80 122M-IP4 60 - 10 12.80 3.66 4.23 29.56 37.17 10.08 1.69

• Temperature difference corrected by a factor of 0.605; (FSAR 3, Supplement 16, April 1973)

Chapter 2, Page 56 of 119 Revision 22, 2010

IP2 FSAR UPDATE The morning transition during the summer season was sharply defined. At 0800 EST all levels had approximately the same resultant wind direction. The evening transition began just after 1800 EST and was sharply defined a the 10M and 38M levels. The 10M level reached its nocturnal northeast drainage wind by 2100 EST along with the somewhat more erradic 38M level. The 85M and 122M levels rotated systematically and did not reach their nocturnal directions (NNE-N) until 0200 EST. The systematic rotations are referenced to the average diurnal distributions on a real time basis, the upper wind levels could be disconnected from the lower wind levels with an intermediate shear zone generated by winds up to 180° out of phase. The resultant winds of the 122M Tower (IP4) associated with the diurnal variation curves for the summer season, veered after the morning transition until about noontime, then steadied out at SW-WSW before backing into the nocturnal pattern after the evening transition. The resultant summer season winds for the IP3 Tower (Kaplin and Laznow, 1972) in 1970 and 1971 veered throughout the entire day. In Kaplin and Kitson (1974), the summer IP3 diurnal resultant winds exhibited the veering - backing trait. Kaplin and Laznow (1972) indicated that the question of backing or veering was related, on any given day, not only to strength of the valley drainage flow wind but also to relative strength of local land-sea circulations. On an annual basis there were no significant differences between the percent frequency distribution of occurrences of stability categories (Pasquill) between the adjusted composite year 1970-1972 for IP3 (FSAR 3, Supplement 16, April 1973) and lower temperature gradient level on the 122M Tower (IP4). These comparisons are shown in Table 3. It is presumed that if all individual years of IP3 data were similarly adjusted prior to classification, they would also be reasonably comparable to results based on the temperature gradient 60-10M on the 122M Tower. 3.2.2 August 1, 1978 - July 31, 1979 Wind velocity data (10M level) from the 122M Meterological Tower at Indian Point and the Orange and Rockland Utilities, Inc. 100M Meteorological Tower were used to evaluate the path of air parcels in the Indian Point environs without considering stability (Kaplin and Wuebber, 1979 and Kaplin, 1979). Each hour a parcel movement was initialized from Indian Point. Each parcel was projected forward for eight consecutive hours in hourly increments. The average wind velocity at 10M level of the 122M Tower was used to determine the speed and direction of the parcel for its initial hour increment. Subsequent movement of each parcel was determined by the location of the parcel after the initial hour on a zone of influence file that assigned a wind vector to that location: Indian Point or Bowline. Prior to usage the wind velocity for selected 1978-1979 data were assessed by comparison with historical data files (1973-1974) at Indian Point. There were no variations that could not be accounted for by climatological variations of at least synoptic scale when assessed with reference to U.S. Department of Commerce, NOAA, EDS, LOCAL CLIMATOLOGICAL DATA for LA GUARDIA AIRPORT, NEW YORK and SIKORSKY AIRPORT, Bridgeport, CONNECTICUT. In considering persistent southward movement of an air parcel from Indian Point assuming that Bowline would be representative of air movement south of Grassy Point, an examination of resultant winds for August 1978 and January, 1979 (typical summer Chapter 2, Page 57 of 119 Revision 22, 2010

IP2 FSAR UPDATE and winter seasons) indicated that such movements did not occur. While concurrent hourly average north winds were found 14 times in August and 17 times in January, these occurrences represent only 13.3-percent and 16.5-percent of all north winds relative to all north winds at Bowline. These results were anticipated, particularly during periods of light winds and weak synoptic pressure fields, from the opposing patterns of the diurnal variation curves for the two monitoring sites. 3.2.2.1 Surface Air Trajectories Analyses - Summary Trajectory end points were derived on an objective basis using surface wind data from monitoring stations. The use of observed wind data appropriate to the moving air parcels location at a given time is important since these data inherently account for local wind pattern aberrations that may be of topographic and/or unique micro-meteorological origin. No individual atmospheric stability category was explicitly considered. The ability of the derived trajectories to generate realistic movement patterns is contingent on having sufficient wind monitoring sites to define the actual wind flow field in and around the area of interest on a concurrent real time basis. The area of interest was limited to ten miles south of Indian Point. For practical purposes the study area was 21 x 21 square miles subdivided in a one mile grid as shown in Figure 11. Indian Point was located near the top center of the area at grid point 10, 16. This allowed for 15-miles of due south movement. The South, Hi Tor and Hook Mountains are emphasized because of the barrier that they form for air movement due south of Indian Point. Trajectories were generated for each of the 12 months in the data file based on Indian Point and Bowline wind data. These were the only available data applicable to the study area. Trajectories were created for up to eight consecutive hours of movement. For the first hour of the trajectory, Indian Point was used as the origin of an air parcel, which would travel a distance and in a direction determined by the hourly average wind velocity at the 33 (10M) level of the Meteorological Tower. Subsequent movement depended on the location of the trajectory end point after the first hours movement. Zones for which the Bowline and Indian Point wind velocity measurements were considered as representative had been previously assigned. Trajectories for each hour of the month were computed. The end points were accumulated as summations of occurrences in their appropriate grid squares. In the process of generation, all end points were moved and accumulated whether or not they were in the 21 x 21-mile square. Only those end points within the study area boundary appear on tabular printouts. In any given period, an end point could pass out of the grid and move back in at a subsequent time interval. For August, 1978, and January, 1979, two different patterns of weather station representative areas were used as shown in Figure 12. Pattern 1, which was used for all 12 months of data, had Bowline winds dominating after passage of a line three miles south of Indian Point (through Grassy Point). Pattern 2, used for August and January only, moved this line one mile further north (through Stony Point). The influence patterns are the same for the first hours movement. Chapter 2, Page 58 of 119 Revision 22, 2010

IP2 FSAR UPDATE A summary of the August, 1978 and January, 1979, results in terms of percent of total possible trajectory end points remaining in the 21 x 21-mile area for selected trajectory time periods is shown in Table 4. TABLE 4

SUMMARY

OF TRAJECTORY END-POINTS August, 1978 January, 1979 Pattern I No. of Occur.  % Total No. of Occur.  % Total Hour 1 722 (729) 99.0 707 (721) 98.1 Hour 2 617 (728) 84.8 486 (720) 67.5 Hour 4 420 (726) 57.9 225 (718) 31.3 Hour 6 312 (724) 43.1 157 (716) 21.9 Hour 8 206 (722) 28.5 113 (714) 15.8 Pattern II Hour 1 722 (729) 99.0 707 (721) 98.1 Hour 2 595 (728) 81.7 486 (720) 67.5 Hour 4 414 (726) 57.0 226 (718) 31.5 Hour 6 308 (724) 42.5 157 (716) 21.9 Hour 8 207 (722) 28.7 115 (714) 16.1 ( ) =Total number of trajectories generated. The actual number of points within the grid network does not differentiate between those points that have never left the network and those that have recirculated. This feature takes on added importance if total distance of travel is a consideration. Summaries of occurrences within designated grid sectors are shown in Table 5 for August, 1978 and Table 6 for January, 1979. In terms of totals in the grid area, there is no significant effect of influence pattern assignment. This effect does show up in Tables 5 and 6 when the occurrences south of Indian Point are totaled. For this purpose, a SW sector is defined encompassing the area below Indian Point from the grid edge to ordinate Line 9. The S sector extends one mile south of Indian Point along ordinate Chapter 2, Page 59 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5 SUMMATION OF TRAJECTORY END POINTS AUGUST, 1978 SECTOR KEY 17-21 NW N NE 16 W I.P. E 1-15 SW S SE 1-9 10 11-21

                                                                        % TOTAL 35              164            133           46.0 HR 1                         26              8              15            6.8 232             84               25          47.2
             % TOTAL           40.6            35.4           24.0          99.0 PATTERN 1                                  PATTERN 2
                             % TOTAL                                      % TOTAL 52      93    109   41.2                   52      93     109    42.7 HR 2     20      10      7    6.0                   18      10       7     5.9 192      51     83   52.8                  148      50     108    51.4

% TOTAL: 42.8 25.0 32.2 84.8  % TOTAL 36.6 25.7 37.6 81.7

                             % TOTAL                                      % TOTAL 43      38     40   28.8                   44      38      36    28.5 HR 4      7       2      2    2.6                    7       1       2    24.4 128      27    133   68.6                  111      29     146    69.1

% TOTAL: 42.3 16.0 41.7 57.9  % TOTAL 39.1 16.4 44.4 57.0

                             % TOTAL                                      % TOTAL 33      19     18   22.4                   30      19      19    22.1 HR 6      7       0      1    2.6                    6       0       1     2.3 100      10    124   75.0                   86      17     130    75.6

% TOTAL: 44.9 9.3 45.8 43.1  % TOTAL 39.6 11.7 48.7 42.5

                             % TOTAL                                      % TOTAL 28       6      4   18.4                   28       6       4    18.4 HR 8      3       0      1    1.9                    3       1       1     2.4 67       7     90   79.6                   53      15      96    79.2

% TOTAL: 47.6 6.3 46.1 28.5  % TOTAL 40.6 10.6 48.8 28.7 Chapter 2, Page 60 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6 SUMMATION OF TRAJECTORY END POINTS JANUARY, 1979 SECTOR KEY 17-21 NW N NE 16 W I.P. E 1-15 SW S SE 1-9 10 11-21

                                                                       % TOTAL 25              52             86           23.1 HR 1                         10              10             27           6.7 154             72             271          70.3
             % TOTAL           26.7            19.0           54.3         98.1 PATTERN 1                                  PATTERN 2
                             % TOTAL                                     % TOTAL 22      21     58   20.8                   24      21     56    20.8 HR 2     14       5     16    7.2                   12       5     16     6.8 133      41    176   72.0                  100      45    207    72.4

% TOTAL: 34.8 13.8 51.4 67.5  % TOTAL 28.0 14.6 57.4 67.5

                             % TOTAL                                     % TOTAL 17       4     19   17.7                   14       4     21    17.3 HR 4      7       0      3    4.4                    7       0      3     4.4 67      16     92   77.8                   57      14    106    78.3

% TOTAL: 40.4 8.9 50.7 31.3  % TOTAL 34.5 8.0 57.4 31.5

                             % TOTAL                                     % TOTAL 16       2      7   15.9                   15       2      8    15.9 HR 6      2       1      0    3.2                    2       1      0     1.9 46       8     73   80.9                   39       8     82    82.1

% TOTAL: 40.8 7.0 51.0 21.9  % TOTAL 35.7 7.0 57.3 21.9

                             % TOTAL                                     % TOTAL 9       1      4   12.4                    6       0      4     8.7 HR 8      3       1      0    3.5                    3       1      0     3.5 34      10     51   84.1                   29       9     63    87.8

% TOTAL: 40.7 10.6 48.7 15.8  % TOTAL 33.0 8.7 58.3 16.1 Chapter 2, Page 61 of 119 Revision 22, 2010

IP2 FSAR UPDATE Line 10 to the grid base. The SE sector comprises the remaining area to the east of the S line and below Indian Point. These results are summarized below in terms of number of occurrences and percentage of total possible observations: TABLE 7 SUMMATION TRAJECTORY OCCURRENCES SOUTH OF INDIAN POINT August, 1978 Pattern 1 Pattern 2 Southwest South Southeast Southwest South Southeast Occur  % Occur  % Occur  % Occur  % Occur  % Occur  % Hour 1 232 31.8 84 11.5 25 3.4 232 31.8 84 11.5 25 3.4 Hour 2 192 26.4 51 7.0 83 11.4 148 20.3 50 6.9 108 14.8 Hour 4 128 17.6 27 3.7 133 18.3 111 15.3 29 4.0 146 20.1 Hour 6 100 13.8 10 1.4 124 17.1 86 11.9 17 2.3 130 18.0 Hour 8 67 9.3 7 1.0 90 12.5 53 7.3 15 2.1 96 13.3 JANUARY, 1979 Hour 1 154 21.4 72 10.0 271 37.6 154 21.4 72 10.0 271 37.6 Hour 2 133 18.5 41 5.7 176 24.4 100 13.9 45 6.3 207 28.8 Hour 4 67 9.3 16 2.2 92 12.8 57 7.9 14 1.9 106 14.8 Hour 6 46 6.4 8 1.1 73 10.2 39 5.4 8 1.1 82 11.5 Hour 8 34 4.8 10 1.4 51 7.1 29 4.1 9 1.3 63 8.8 The effect of the pattern change is not so much as to alter the total; rather, it is to shift the number of occurrences from the SW sector to the S and SE sectors. There are anomalies found that may be associated with recirculation. After five miles of southward movement from Indian Point, the results seem to indicate the anomaly of surface wind impaction against the South Mountain and High Tor Ridges. This anomaly occurred since there were no local wind measurements available to induce deflections. Historical studies have shown such deflections do exist. The present results cannot account for terrain unless the trajectory paths are deflected by observed surface winds. This requires a larger monitoring network, strategically placed, than was available. This need for further definition of local wind field is confirmed by the differences that appear in the results generated by Patterns 1 and 2. At the present time, based on historical studies, Pattern 2 is probably the better representation of local trajectories for the available data. Assuming a continuous 1 M/S wind speed (2.2 MPH), the number of occurrences in the south sector represent those parcels that have traveled with the effective speed (neglecting recirculation). Of the totals given, only four have traveled greater than ten miles for August (1); five for August (2); six for January (2); six of January (1); and three for January (2). These points would have to had passed through or over the South, High Tor and Hook Mountain Ridge lines. Chapter 2, Page 62 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.2.3 March 1980 - December 1980 3.2.3.1 General The results of the Trajectory II Study conducted by York Services Corporation for Con Edison have been recently submitted (Kaplin and Wuebber, 1981). It was concluded from the initial trajectory study (Kaplin and Wuebber, 1980; Kaplin, 1980) that a lack of directional persistence of low speed surface winds (10M) at Indian Point and Bowline make recirculation of local air probable. There were indications of both convergence and divergence of local air streams. Objectivity created surface air parcel trajectories generated anomalies by passing over or through abrupt terrain features. The two local monitoring sites available were unable to resolve these anomalies. In the Trajectory II Study, a supplemental network of ten surface wind monitoring stations were established for the express purpose of objectively assessing the southward movement of air parcels from Indian Point (see Figures 3 and 4). Sites were selected, specifically, in an attempt to resolve anomalous flow patterns with respect to terrain and tributary river drainage basins. A listing of sites used is shown in Table 8. A listing of valid data collected for the period is shown in Table 9. 3.2.3.2 Wind Frequency Distributions An historical evaluation of the representativeness of the data collected in 1980 was made for Indian Point and Bowline. Variations in wind frequency distributions were found to be associated with climatological variations on the synoptic-cyclonic scale. These variations can be naturally expected between any given year or set of years. For example: Over a 20 year period (1960-1979) prior to 1980 Bridgeport, WBAS, for the month of July had an average wind directional frequency for a north wind of 6.5 +/- 2.7-percent with an absolute maximum of 12.9-percent in 1974 and an absolute minimum of 3.0-percent in 1979. In 1980, the frequency set a new low of 2.8-percent. A further extreme example was found at La Guardia WBAS. Based on an eight year average (1972-1979) the northwest wind has a frequency of 12.2 +/- 4.2-percent with minimum of 6.1-percent (1973) and a maximum of 16.5 (1974). In 1980, a new maximum of 18.2-percent was observed while in 1979, the frequency was 7.9-percent, which was the second lowest value in the period. It was noted that the climatic variations of wind frequencies at Indian Point were generally minimal and less pronounced. This was attributed to topographic confinement. The wind frequency data for Indian Point and Bowline for the data collection period were adjudged to be representative (Figures 6.1-6.4, Kaplin and Wuebber, 1981). It was assumed that all concurrently collected wind velocities were representative of respective monitoring sites and relationships between sites could be evaluated. It was found that wind frequency distribution patterns in themselves were deceptive representations of the continuity of air movement in the lower Hudson River valley unless there was an understanding of the patterns of wind velocity variations on a temporal basis. Chapter 2, Page 63 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 8 LOCATIONS OF STATIONS RELATIVE TO INDIAN POINT Distance (miles) Direction Station from Indian Point (degrees) Iona Island 2.50 334 Annsville 2.20 020 Watch Hill Road 3.15 132 Jurka 6.65 122 Croton Point 6.40 155 Ossining 8.80 151 Grassy Point 3.20 191 Bowline Point 4.15 190 South Nyack 13.60 174 Piermont 15.95 173 Kingsland 13.00 163 Eastview 11.90 145 Westchester County Airport 20.00 135 Chapter 2, Page 64 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 9 YORK SERVICES CORPORATION ONE RESEARCH DRIVE, STAMFORD, CT CLIENT: CONSOLIDATED EDISON OF NEW YORK VALID DATA FOR TRAJECTORY WIND SITES PERIOD OF RECORD: 1980 SITE PERCENT VALID DATA MARCH APRIL MAY JUNE JULY AUGUST SEPT OCT NOV DEC 01-Piermont 79.23 78.75 63.71 66.67 100.00 100.00 71.25 37.10 0.00 0.00 02-Ossining 98.59 100.00 100.00 100.00 100.00 99.80 99.93 96.77 100.00 100.00 03-Iona Island 96.77 53.75 85.28 69.72 53.76 95.63 100.00 99.46 100.00 82.80 04-Jurka/Grassy 78.76 55.97 0.00 0.00 0.00 0.00 0.00 81.85 100.00 84.95 05-Kingsland 87.57 94.65 17.47 59.79 98.66 100.00 100.00 100.00 74.03 93.55 06-Watch Hill 74.46 0.00 0.00 0.00 48.79 70.56 100.00 85.48 100.00 100.00 07-South Nyack 79.17 100.00 83.87 100.00 100.00 100.00 100.00 100.00 100.00 100.00 08-Annsville 100.00 100.00 90.86 100.00 77.28 100.00 71.81 98.92 100.00 100.00 09-Eastview 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 10-Croton Point 99.73 100.00 89.85 100.00 98.12 96.77 81.39 47.04 100.00 36.96 11-West Cty Apt 99.87 100.00 100.00 99.86 100.00 100.00 100.00 100.00 100.00 100.00 12-Indian Point 99.66 100.00 94.82 99.24 95.41 94.69 99.17 99.93 100.00 99.46 13-Bowline Point 72.45 99.72 98.91 85.97 92.41 98.52 96.53 96.98 96.46 86.16 Chapter 2, Page 65 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.2.3.3 Diurnal Wind Distributions The diurnal variation curves for Indian Point and Bowline for the data collection period (March through December 1980) were found to be historically representative. For selected winter and summer months, they demonstrated all the attributes of the two season characteristics (Figure 6.21-6.27, Kaplin and Wuebber, 1981). With some variation at selected monitoring sites, it was found that the diurnal wind distributions were not only seasonally characteristic but characteristic of the monitoring site locations. They could, almost without exception, be uniquely categorized as Hudson River west bank; Hudson River east bank; or inland (Figures 6.28-6.33, Kaplin and Wuebber, 1981). The characteristics of this uniqueness were examined by combining all appropriate sites to generate average east and west bank diurnal wind distributions (Figures 6-38-6.42, Kaplin and Wuebber, 1981). The average diurnal curves for March, June and December, 1980 are shown in Figures 13, 14, and 15. A computer check revealed that individual days at any given site could be found that had observed 24 hour diurnal wind variation patterns that matched their own monthly average distributions and/or the appropriate east or west bank average diurnal distribution based on the criteria 16 or more hours of fit +/- 45° (not necessarily consecutive). While there are unique common characteristics to the diurnal wind distribution patterns in the Indian Point environs, variations in local meso-scale factors dictate that the ultimate path of an air parcel whose movement is determined by surface (10M level) wind velocities is governed by time of departure as well as point of departure. Between wind velocity monitoring sites in the region, persistent wind direction and wind speeds are not supported. This is most obvious during the summer season or at any time that the area is under the influence of a weak synoptic-cyclonic pressure gradient pattern. Between individual monitoring sites there is apparent divergence and convergence of surface air. 3.2.3.4 Resultant and Concurrent Hourly Winds A first approach at the evaluation of southward movement of air for prolonged periods of time was made for the data collection period March 1, 1980 through December 31, 1980, by examining the frequency distribution of the 24 hour resultant winds (Kaplin and Wuebber, 1981). These results are shown in Table 10 as a function of persistence category (the ratio of the resultant to arithmetic average wind speeds). At persistence levels greater than 0.9, a north wind was found in only four out of 273 possible valid cases (1.4%). The average wind speed 2.75 M/S. The high wind speeds associated with all northerly winds implies strong synoptic-cyclonic scale pressure gradient systems are the generating mechanism. Simple liner relationships between high persistent, 24-hour resultant winds between Indian Point and Piermont (bearing 173° about 16-miles from Indian Point) showed that an average 24-hour resultant wind direction of 012 +/- 22° at Indian Point was related to an average 24-hour resultant wind at Piermont of 359 +/- 24°. At the same time, for corresponding cases, the average resultant wind speed at Indian Point was about 2.5 +/- 0.9 M/S and the concurrent average resultant wind speed at Piermont was 5.6 +/- 1.2 M/S. The angular offset implies terrain tracking and high average resultant wind speeds indicate the necessity for a strong pressure gradient field. Chapter 2, Page 66 of 119 Revision 22, 2010

IP2 FSAR UPDATE Such replacements were also implied when concurrent hourly average wind data from the selected monitoring sites were correlated to Piermont. These results are shown in Table 11 for the available concurrent data collected during the period from March 1, 1980 through October 31, 1980. Out of 4,394 valid data hours, there were only 56 (1.3%) in which Indian Point and Piermont had concurrent winds from the north (350-011°). Almost half of these cases (26:0.59%) occurred in May, 1980. There was only one such hour out of 742 valid data hours in July, 1980. For July, in fact, for 3,355 concurrent data hours from five southern sites there was only one additional hour in which a site, South Nyack, had a north wind direction concurrent with Piermont. Chapter 2, Page 67 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 10 FREQUENCY DISTRIBUTION OF 24 HOUR RESULTANT WIND DIRECTIONS INDIAN POINT (10 METER LEVEL) Persistence Persistence Resultant > 0.9 > 0.8 <0.9 Wind Aver. SPDS (MPH) Aver. SPDS (MPH) Direction No. Obs. Result. Mean No. Obs. Result. Mean No. Obs. 350-011 4 5.75 6.15 6 4.87 5.62 2 012-034 20 4.97 5.22 6 4.05 4.67 4 035-056 13 4.81 4.98 8 3.56 4.04 1 057-079 5 4.06 4.24 2 2.70 3.15 2 080-101 0 0 1 102-124 0 0 0 125-146 0 0 0 147-169 0 0 1 170-191 2 2.90 3.10 1 3.20 3.70 0 192-214 10 3.38 3.53 5 2.38 2.76 5 215-238 3 4.07 4.23 7 2.69 3.11 1 237-259 0 3 4.40 5.27 2 260-281 0 1 2.40 3.00 2 282-304 0 3 3.03 3.70 1 305-326 6 4.17 4.48 4 5.15 6.00 3 327-349 13 7.82 8.22 8 4.64 5.40 7 76 54 31 Percent: 25.94 18.43 10.58 Number of Valid 24-hour Resultants ( >17 hours): 293 Percent Valid Data: 97.39 Chapter 2, Page 68 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 11

SUMMARY

OF TWO-STATION WIND CORRELATIONS PIERMONT (SITE 1) REFERENCED TO SELECTED MONITORING LOCATIONS (SITE 2) North Wind at Piermont North Wind at Site 2 Number of Observations Resultant Wind @ Site 2 Resultant Wind @ Piermont Station North Winds Concurrent Winds Speed Speed Month (Site 2) Total Piermont Site 2 All Directions North Direct. (mph) Persist. Direct. (mph) Persist. March South Nyack 588 83 33 70 10 331 3.3 0.90 025 8.5 0. Kingsland 481 71 74 133 39 004 6.7 0.96 001 9.2 0. Ossining 586 83 25 31 4 327 3.8 0.92 023 7.5 0. Iona Island 565 76 16 65 3 327 5.0 0.73 042 9.4 0. Indian Point 589 83 67 137 7 035 4.2 0.89 334 9.3 0. April South Nyack 567 50 25 63 6 346 1.9 0.65 025 6.6 0. Kingsland 490 39 68 145 13 008 3.2 0.79 018 8.1 0. Ossining 567 50 28 45 2 316 2.6 0.85 030 8.8 0. Iona Island 354 30 21 26 2 308 4.5 0.73 043 11.0 0. Indian Point 567 50 58 129 6 020 2.4 0.64 340 8.5 0. May South Nyack 359 46 20 33 2 327 3.7 0.84 032 6.8 0. Kingsland 67 2 2 9 0 256* 2.1* 0.98* 031* 10.0* 1. Ossining 474 68 37 28 4 309 6.4 0.85 036 9.9 0. Iona Island 415 66 22 20 2 301 7.3 0.89 052 6.6 0. Indian Point 422 50 58 91 26 350 3.5 0.79 356 10.1 0. June South Nyack 480 17 19 57 2 332 4.0 0.92 036 8.2 0. Kingsland 284 5 2 30 1 301* 5.6* 0.95* 031* 5.4* 0. Ossining 480 17 16 41 0 318 7.3 0.90 046 6.7 0. Iona Island 368 12 8 16 0 301 8.5 0.88 062* 9.1* 0. Indian Point 475 17 28 80 4 326 3.1 0.73 006 5.6 0. July South Nyack 744 10 25 114 1 332 2.7 0.92 041 7.9 0. Kingsland 724 10 25 98 0 337 2.3 0.55 052 5.7 0. Ossining 744 10 37 66 0 326 4.5 0.79 047 6.8 0. Iona Island 401 5 9 45 0 302* 10.0* 0.95* 054* 5.7* 0. Indian Point 742 10 35 108 1 007 2.3 0.64 358 4.7 0.

• Less than 10 valid data points Chapter 2, Page 69 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 11 (Contd)

SUMMARY

OF TWO-STATION WIND CORRELATIONS PIERMONT (SITE 1) REFERENCED TO SELECTED MONITORING LOCATIONS (SITE 2) North Wind at Piermont North Wind at Site 2 Number of Observations Resultant Wind @ Site 2 Resultant Wind @ Piermont Station North Winds Concurrent Winds Speed Speed Month (Site 2) Total Piermont Site 2 All Directions North Direct. (mph) Persist. Direct. (mph) Persist. Aug. South Nyack 744 24 44 95 1 334 3.0 0.75 058 11.9 0.95 Kingsland 744 24 52 93 1 313 4.8 0.74 054 11.7 0.92 Ossining 741 24 58 96 2 336 6.9 0.84 047 10.5 0.87 Iona Island 679 24 15 80 0 308 4.9 0.66 036 9.8 0.80 Indian Point 719 24 32 139 3 037 2.5 0.70 332 5.2 0.73 Sept. South Nyack 513 37 36 63 10 348 2.5 0.88 018 12.6 0.91 Kingsland 513 37 22 42 3 345 4.6 0.79 026 9.6 0.90 Ossining 512 36 36 87 9 354 5.0 0.86 010 9.4 0.84 Iona Island 513 37 10 34 3 322 4.8 0.76 024 11.7 0.91 Indian Point 501 37 38 102 4 031 3.4 0.83 339 7.2 0.85 Oct. South Nyack 379 24 28 83 5 335 2.8 0.91 019 10.7 0.92 Kingsland 379 24 25 64 5 018 3.0 0.73 019 11.4 0.96 Ossining 379 24 31 69 4 007 4.0 0.85 016 11.8 0.95 Iona Island 379 24 12 38 3 291 2.8 0.77 016 12.9 0.99 Indian Point 379 24 24 89 5 031 2.9 0.95 009 7.5 0.86 Table 11 can be summarized in terms of the valid wind direction data from the five designated sites concurrent with wind directions at Piermont (Note: The data from these sites should not be presumed concurrent with each other simultaneously). This summary is shown in Table 12. Chapter 2, Page 70 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 12 CONCURRENCE OF TWO-STATION WIND DIRECTIONS (Relative to Piermont) 5 Site Total  % Valid Percent (Concurrent Data Basis) Concurrent Data 6 Site Basis All Directions North Wind March 2809 62.93 15.52 2.24 April 2545 58.91 16.03 1.14 May 1737 38.91 10.42 1.96 June 2087 48.31 10.73 0.34 July 3355 75.16 12.85 0.06 August 3627 81.25 13.87 0.19 September 2552 59.07 12.85 1.14 October 1895 42.45 18.10 1.16 Since these data are derived from hourly average wind directions, it is again shown that there is little likelihood of sustaining south bound movement of air from Indian Point beyond 15-miles. 3.2.3.5 Summary-Trajectory II Study A modified version of the generic model TRAJECTORY (Kaplin and Wuebber, 1980) was used as a basis of a study, which involved the use of concurrent hourly average wind data from a network of 13-14 monitoring stations within 20-miles of Indian Point. All but four sites were located on the Hudson River shorelines. Only two sites were used to the north of Indian Point. As with the original study (Kaplin and Wuebber, 1980), emphasis was on the objective creation of the trajectories of air parcels originating hourly at Indian Point with a speed and direction equal to the average wind at the 10 meter level of Indian Point 122 Meter Meteorological Tower. In the Trajectory II Study (Kaplin and Wuebber, 1981), each parcel of air was tracked for eight consecutive hours after its movement was initiated as dictated by two factors: The movement time interval and the wind velocity at the coordinate end point of the parcel at the end of movement time interval. As in the earlier study a 21 by 21-mile grid pattern of one mile squares was used to generate tabulations of the trajectory segment end-points. While the tabulations assumed each point to be located in the center of each square, the actual coordinates within the squares were used as a starting point for the next hourly trajectory segment. Subsequent movement of a parcel from a given set of coordinates was determined by the appropriate average hourly wind velocity assigned to grid square as determined by a zone of influence file. While trajectories were objectively created, the zone of influence file required the subjective assignment of each available wind velocity monitoring site to specific grid squares. The wind velocity for any given hour within these assigned grid squares would be the same as that of the specified control monitoring site. Influence assignments were based on assessments of local wind patterns (historical and present) with some consideration for obvious meso-scale modification factors: topographic channeling, drainage flow patterns, thermally induced flow patterns, etc. Chapter 2, Page 71 of 119 Revision 22, 2010

IP2 FSAR UPDATE For the purpose of the study, all movements of parcels past the grid boundaries were assumed to continue their movements under the influence of the site whose wind was being used at the time that a boundary was crossed. When only a few wind monitoring sites are available to cover a large area, movement controlled solely by a single sites non-variant hourly average wind was not a critical factor. As the number of monitoring sites increased and zones of influence became smaller, discrete movement based on a single wind in a given hour increment would allow a parcel to move through a zone of influence without modification of its controlling wind, which could be substantially different in direction and speed than that associated with the by-passed zone. In the TRAJ3 model, as used in this study, parcel end point coordinates at the end of each hour, which had been previously determined by the non-variant wind at the parcel source at the beginning of each hour, were the resultant of 30 discrete movements (two minute intervals) within the hour interval. On this basis, a wind speed of 30 mph was required in order for a parcel to travel with a non-variant wind for more than one mile. A discrete wind velocity was reassigned to a parcel according to its coordinates (zone of influence) at the end of each two minute intervals. In effect, a parcel could, in extreme, alter its direction and speed 30 times in a given hour and not apparently move at all if it were trying to move from one zone of influence to another at a boundary line between zones and the wind in the two zones were in opposition. Such apparent anomalies were found as a matter of routine. Figure 16 shows the grid system that was used in the Trajectory 2 Study, Indian Point was located at coordinates 7,17. In the development of the trajectories recirculation was allowed. This is a parcel could leave the grid boundaries and be brought back onto the grid at a later time if dictated by a change in the wind at the monitoring site controlling its movement. In the first hour of movement for the ten months of data, only 390 parcels out of 7,344 (5.3) left the grid system and did not return. Of these 345 (4.7%) were crossings of the northern grid boundary, 40 (0.5%) were crossings of the western grid boundary and 5 (0.1%) were crossings of the southern grid boundary. There were no crossings of the eastern grid boundary. Of the five points crossing the southern boundary, all occurred in December, 1980. To cross the southern boundary a minimum speed of 7.6 M/S (17 mph) was required. During the ten months (206 days) of record there were only 35 days (11.4%) in which there are no rotation and/or recirculation (flow reversal) characteristics in at least one of the 24 trajectories created daily. This does not take into account any trajectories that may have experienced reversals wholly outside of the grid area. If flow reversals and rotational trajectories were due solely to synoptic scale meteorological patterns, then randomness would be expected in the starting time occurrence frequency as a diurnal function. The actual pattern observed for those eight hour trajectories, which contained at least one on grid or on-off-on grid flow reversal is shown in Table 13. These results include those reversals of synoptic origin and can be interpreted as consistent with the local average diurnal wind patterns induced by meso-scale phenomena when the following factors are taken into account: z Nocturnal flow patterns generally north to south with minimal with speeds z Afternoon flow patterns generally south to north with maximum wind speeds Chapter 2, Page 72 of 119 Revision 22, 2010

IP2 FSAR UPDATE Summaries of the trajectory end-point counts and percentages are shown on a monthly basis in Tables 14 and 15 for selected time increments up to the complete eight hour trajectory. The idealized valley is shown in Figure 16. It is noted that the idealized valley contains 64 out of 441 (14.5%) possible grid box end-point coordinates: ten grid boxes are north of Indian Point; and 54 are on a line with Indian Point and south. With respect to total grid, 84 boxes are to the north and 357 are to the south. The effect of dominant meso-scale factors are readily discernible in the results. Differences in north and south boundary crossings of trajectory points can be related to the normal seasonal distribution of local wind velocities as well as their diurnal distribution patterns. During the summer season it would take up to six hours of persistent north sector winds from Indian Point to Piermont to generate a south boundary crossing. Such a persistent diurnal time span is improbable. After eight hours, only 6.7 percent of all possible trajectory end-points for all time intervals in July 1980 (5,924 possible) were found to cross the southern boundary (a distance of 17-miles from Indian Point). The highest crossing percentage, 30.7 percent of 5,924 possible, occurred in December, 1980. Chapter 2, Page 73 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 13 Diurnal Distribution of Occurrences of Eight-Hour Trajectories with On Grid Reversals Starting End Number of Trajectory with Flow Reversals Hour Hour Mar Apr May June July Aug Sept Oct Nov Dec Total 0100 0900 7 9 13 13 14 9 13 7 7 10 102 0200 1000 11 9 15 9 11 10 15 10 9 6 105 0300 1100 9 11 12 15 14 14 12 12 8 8 115 0400 1200 12 11 18 14 17 13 11 12 6 10 124 0500 1300 10 11 16 17 21 13 14 11 4 9 126 0600 1400 9 11 10 13 19 13 14 11 7 11 118 0700 1500 10 8 9 10 10 12 12 7 9 9 96 0800 1600 6 7 5 10 9 10 14 7 10 7 85 0900 1700 6 10 6 8 11 6 10 4 7 8 76 1000 1800 5 7 10 11 7 8 9 4 9 7 77 1100 1900 6 8 4 4 9 7 9 6 4 5 62 1200 2000 2 8 2 4 10 8 10 5 6 4 59 1300 2100 2 8 4 5 6 5 8 4 5 6 53 1400 2200 3 4 5 5 11 5 7 6 5 6 57 1500 2300 2 7 6 4 12 11 8 6 8 6 70 1600 2400 5 5 7 6 12 11 8 6 10 7 77 1700 0100 7 6 10 9 17 12 9 6 8 8 92 1800 0200 10 7 12 8 16 10 9 9 7 9 97 1900 0300 10 3 14 11 14 8 7 6 7 10 90 2000 0400 12 9 14 7 15 8 6 6 6 8 91 2100 0500 12 8 13 5 8 8 7 7 9 8 85 2200 0600 15 9 12 10 8 6 7 8 6 8 89 2300 0700 10 12 13 8 10 9 8 9 5 9 93 2400 0800 7 5 12 10 13 11 9 8 5 11 91 TOTAL 188 193 242 216 294 227 236 177 167 190 2130 TOTAL POSSIBLE 736 712 736 712 736 736 712 736 712 736 7264 OCCURRENCE (%) 25.5 27.1 32.9 30.3 39.9 30.8 33.1 24.0 23.5 25.8 29.3 Chapter 2, Page 74 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 14

SUMMARY

OF TRAJECTORY END-POINT COUNTS March April May June July August Sept Oct Nov Dec Number of Hours/Month 744 720 744 720 744 744 720 744 720 744 Total Number of End-Points 5924 5732 5924 5732 5924 5924 5732 5924 5732 5924 Elapsed Time Number of End-Points 2 Hours - Number of Trajectories 742 718 742 718 742 742 718 742 718 742 Within Grid 490 440 577 559 593 599 498 595 538 513 In Valley (North of Indian Point) 41 29 49 65 51 57 36 39 30 44 In Valley (Indian Point and South) 95 94 150 148 305 185 178 204 139 123 Percent Grid Points in Valley 28.3 28.0 34.5 38.1 60.0 40.4 43.0 40.8 31.4 32.6 4 Hours - Number of Trajectories 740 716 740 716 740 740 716 740 716 740 Within Grid 275 244 356 355 433 361 328 371 285 297 In Valley (North of Indian Point) 20 8 17 34 28 26 15 22 13 17 In Valley (Indian Point and South) 43 59 118 106 139 146 109 156 97 78 Percent Grid Points in Valley 22.9 27.5 37.9 39.4 38.6 47.6 37.8 48.0 38.6 32.0 6 Hours - Number of Trajectories 738 714 738 714 738 738 714 738 714 738 Within Grid 173 183 257 257 342 277 250 255 175 213 In Valley (North of Indian Point) 8 5 9 13 20 9 6 16 7 11 In Valley (Indian Point and South) 18 42 93 91 105 119 80 106 63 46 Percent Grid Points in Valley 15.0 25.7 39.7 40.5 36.5 46.2 34.4 47.8 40.0 26.8 8 Hours - Number of Trajectories 736 712 736 712 736 736 712 736 712 736 Within Grid 121 135 190 194 276 226 199 188 130 160 In Valley (North of Indian Point) 9 5 6 9 9 3 5 4 6 11 In Valley (Indian Pint and South) 11 25 73 61 91 87 51 70 47 38 Percent Grid Points in Valley 16.5 22.2 41.6 36.1 36.2 39.8 28.1 39.4 40.8 30.6 Total Past South Boundary* 1643 958 1048 636 399 708 657 926 1706 1821 Percent Past South Boundary 27.7 16.7 17.7 11.1 6.7 12.0 15.0 15.6 29.8 30.7 Total Past North Boundary* 1013 1676 937 Percent Past North Boundary 17.1 28.3 15.8

• This count includes all end-points for all time intervals.

Chapter 2, Page 75 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 15

SUMMARY

OF TRAJECTORY END-POINTS (Percent) 2 Hours 4 Hours 6 Hours 8 Hours Total  % On  % In Total  % On  % In Total  % On  % In Total  % On  % In Month Traj. Grid Valley Traj. Grid Valley Traj. Grid Valley Traj. Grid Valley March 742 66.0 18.3 740 37.2 8.5 738 23.4 3.5 736 16.4 2.7 April 718 61.3 17.1 716 34.1 9.4 714 25.6 6.6 712 19.0 4.2 May 742 77.8 26.8 740 48.1 18.2 738 34.8 13.8 736 25.8 10.7 June 718 77.7 29.7 716 49.6 19.5 714 36.0 14.6 712 27.2 9.8 July 742 79.9 48.0 740 58.5 22.5 738 46.3 16.9 736 37.5 13.6 August 742 80.7 32.6 740 48.8 23.2 738 37.5 17.3 736 30.7 12.2 September 718 69.4 29.8 716 45.8 17.3 714 35.0 12.0 712 27.9 7.9 October 742 80.2 32.7 740 50.1 24.1 738 34.6 16.5 736 25.5 10.1 November 718 74.9 23.5 716 39.8 15.4 714 24.5 9.8 712 18.3 7.4 December 742 69.1 22.5 740 40.1 12.8 738 28.9 7.7 736 21.7 6.7 Average 73.7 28.1 45.2 17.1 32.7 11.9 25.0 8.5 Standard Deviation 6.5 8.4 7.0 5.3 6.7 4.6 8.1 3.3 Chapter 2, Page 76 of 119 Revision 22, 2010

IP2 FSAR UPDATE In any given month of the ten that were investigated and out of the 7,242 complete eight hour trajectories that were generated, there were a number of basic pattern types. There were those whose sequence of temporal end-points exhibited basic straight line tendencies. These were generally associated with high wind speeds. There were those whose sequence of end-points that rotated in a more or less smooth pattern but they were not usually associated with a meso-scale diurnal rotation or terrain induced deflections. In addition, there were those that exhibited characteristics of recirculation and those in which there were sharp reversals. These were induced by wind velocity changes of synoptic scale origin and/or most frequently they appeared in those trajectories, which included the morning or evening transitional periods, meso-scale induced. In the latter cases, of rotation, recirculation and reversals, the sequences may occur wholly on the grid or on-off-on on the grid. Examination of details of specific trajectories and concurrent opposing winds suggests that three block regions can be projected in the lower Hudson Valley within the grid system. These regions are: Peekskill Bay Haverstraw Bay Tappan Zee The reality of the zones of divergence and convergence on a concurrent wind basis were premised on the continuity of air flow movements locally and for air streams that were projected to cross the Hudson River. This latter feature was not uniquely demonstrated with respect to surface (10 meter) level wind velocities. If some local wind patterns were induced by thermal differentials between land and water during periods of weak geostrophic pressure gradients in accordance with sea breeze concepts: during the day air will move from cool water to warm land with return flow aloft; at night, a reverse flow pattern may develop. If this occurred on opposite shorelines of a wide river, concurrently, then there should be vertical motions induced by convergence and divergence in the mid river area or a region of air flow directionally independent of the shoreline circulations. The light wind speeds that were normally found at Croton Point during periods of weak geostrophic flow may result from its proximity to a mid river transition zone. (During periods of strong northwesterly gradients Croton Point had a high frequency of west to west northwest winds implying a cross river flow parallel to South Mountains.) In addition to the land-water effects in so far as they generate local on-shore and off-shore winds, the effect of nocturnal drainage winds should be considered. The Kingsland Park site was one example. A zone of convergence frequently develops at night between that site and Ossining. This was also apparent from the average diurnal monthly diurnal wind distributions. If only drainage winds are considered, from Figure 4, Ossining would reflect drainage from the Croton River and/or a secondary local river both of which would generate air movements from the north-northeast. Kingsland Park is, however, at the outfall to Gory Brook, while this drains from the north-northeast, it hooks in its final section and outfalls into the Hudson from the southeast. There were no measurements available upon which to alter these directions after the air streams flow into the Hudson River itself. For the objective creation of eight hour trajectories, these winds were presumed to extend into the Hudson River and generate local blocks on a concurrent hourly basis. The dominance of mesoscale flow factors on surface winds in this study have been demonstrated over and over since local meteorological data have been collected. The surface wind data sets used for this study are from the most extensive network of Chapter 2, Page 77 of 119 Revision 22, 2010

IP2 FSAR UPDATE concurrent monitoring stations that have ever been deliberately located in the region. These data have been evaluated by many of the routine methodologies common to earlier local studies. All of the data sets were found to exhibit characteristics of complex meso-scale flow fields distortions. The east shore stations were found to share some common characteristics on a daily basis and the same was true for the west shore stations. These characteristics were frequently in opposition to each other. At the same time inland stations had characteristics that were entirely different from either the east or the west shore stations. The creation and interpretation of eight hour trajectories from these data sets could not be truly separated from the concurrent flow fields on an hourly basis. The eight hour trajectories were a result of the constantly changing concurrent flow field. They were a distinct function of movement interval when based on hourly average wind velocities; and therefore, it may be presumed that in a dynamic flow field they would be equally sensitive to the wind averaging interval itself. As noted earlier, this study did not account for vertical air movement, the trajectories were therefore extremely sensitive to one crucial factor - the assumption of continuity of air movement across the Hudson River without midstream directional distortions. The results of this study indicated that continued southward movement of air parcels in the Hudson River Valley could not generally be sustained past Piermont, if, in accordance with the data evaluated, Permonts winds are assumed to be representative of the full width of Hudson River. 3.2.4 January 1, 1979 Through December 31, 1980 3.2.4.1 Data Analyses In the previous sections, with reference to the Trajectory I and Trajectory II studies, portions of 1979 and 1980 data from the Indian Point Meteorological Tower were analyzed and evaluated with reference to the studies that were in progress. Some of these analyses included references to historical data. In both of these studies it was concluded that the meteorological data being obtained at Indian Point were representative of that site and that any observed variations in wind frequency distributions and diurnal variations were assignable to transient climatological deviations from the norm on, at least, the synoptic-cyclonic scale of meteorological events. There was no indication that any changes could be attributed to local physical or dynamic modification, and/or in monitoring equipment and analyses techniques, which could introduce permanent data bias. To maximize the recent data analyses, all of 1979 and 1980 have been evaluated and compared to historical data as available. The amount of valid data for these years has been previously tabulated in Table 2. 3.2.4.2 Wind Frequency Distributions For Climatological perspective, historical comparisons have been made for selected months: March, July and December, between the wind direction, wind frequency distributions at the 10 meter level of the 122 Meter Indian Point Tower and climatological data available from Bridgeport, Connecticut and La Guardia Airport in New York City. These latter sites are within the synoptic-cyclonic scale range of Indian Point. The distributions are shown in Table 16. With respect to March and July, climatological Chapter 2, Page 78 of 119 Revision 22, 2010

IP2 FSAR UPDATE frequency extremes are noted on a nine year basis for La Guardia and a 21 year basis for Bridgeport. The fluctuations of the frequency distributions and the extremes can be associated with the frequency distribution fluctuations at Indian Point when consideration is given to the fact that winds at Indian Point are channeled by the west bank terrain. Tables 17 and 18 give the percent frequency distribution of wind direction at Indian Point at the on seasonal basis for the 10 meter and 122 meter sensor levels. These data are compared to comparable results for 1973-1974 (Kaplin, et. al.1974, Appendix D). In the summer season (Table 17), there is a frequency shift at the 10 meter level from SW and SSW in 1974 to SSW and S in the 1979-1980 period. This shift (with a directional bias) can be related to a similar shifting pattern at Bridgeport and La Guardia as found for July in Table 15. This shift tendency is also found implied at the 122 meter level. There is no reason to expect that these pattern changes are permanent. For the winter season (Table 18) there is a recent bias of wind frequencies to the NNW ad N sectors at both the 10 meter and 122 meter levels at Indian Point. These shifts have their counterpart in the March and December distributions of Table 16. It is concluded that the wind velocity data that has been collected in recent years is consistent with the data base for FSAR 2 at all measurement levels when normal climatological variations are considered. Chapter 2, Page 79 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 16A HISTORICAL COMPARISONS OF WIND FREQUENCY DISTRIBUTIONS MARCH Indian Point LaGuardia Bridgeport 1974 1979 1980 1974 1979 1980 1974 1979 1980 N .120 .078 .125 .044 .137* .085 .044 .083 .077 NNE .108 .093 .104 .032 .038 .016** .012 .020 .040 NE .115 .174 .173 .109 .105 .085 .097 .031 .052 ENE .019 .039 .057 .085 .032 .061 .069 .043 .097 E .008 .011 .030 .000** .013 .024 .020## .095 .061 ESE .004 .009 .007 .008** .013 .008** .008 .018 .031 SE .011 .004 .011 .008** .023 .032 .016 .035# .004## SSE .027 .014 .016 .040 .044* .032 .012 .031# .016 S .057 .094 .085 .085 .199* .145 .036 .082# .069 SSW .051 .133 .091 .036 .035 .057 .040 .051 .040 SW .045 .065 .046 .052 .018 .024 .085 .057 .069 WSW .020 .042 .024 .028 .024 .028 .057 .035 .032 W .036 .049 .039 .089 .069 .048 .113 .065 .113 WNW .076 .047 .024 .137 .097 .077 .145 .077 .101 NW .124 .065 .034 .165 .079 .182* .133 .097 .137 NNW .129 .085 .134 .081 .066 .081 .093 .134 .048 CALM .000 .000 .000 .000 .008 .016 .020 .047# .012

• 9 Year High (1972-1980)
  • 9 Year Low (1972-1980)

1. 1. 21 Year Low (1960-1980)
2. 21 Year High (1960-1980)

Chapter 2, Page 80 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 16B HISTORICAL COMPARISONS OF WIND FREQUENCY DISTRIBUTIONS JULY Indian Point LaGuardia Bridgeport 1974 1979 1980 1974 1979 1980 1974 1979 1980 N .034 .066 .047 .093* .036 .044 .129# .030 .028## NNE .141 .112 .111 .032 .035 .040 .057 .012## .016 NE .148 .104 .158 .057 .059 .052 .028 .015 .012## ENE .054 .085 .073 .048 .055 .040## .020 .028 .032 E .026 .044 .045 .020 .013 .016 .069 .082 .057 ESE .011 .019 .022 .012 .013 .008 .016 .054 .024 SE .020 .013 .022 .044* .031 .032 .048 .020 .016 SSE .024 .028 .026 .016** .032 .040 .020 .022 .020 S .053 .079 .085 .081** .184 .161 .073 .102 .129 SSW .110 .212 .123 .069 .085 .073 .052## .079 .113 SW .129 .097 .094 .061** .095 .133* .085 .157 .125 WSW .044 .028 .050 .081 .063 .040** .154 .114 .109 W .047 .038 .061 .149* .073 .073 .145# .105 .117 WNW .043 .024 .031 .089 .063 .081 .048 .057 .052 NW .034 .026 .024 .073 .067 .093 .004## .069 .069 NNW .015 .015 .031 .057 .043 .052 .048 .026 .069 CALM .000 .000 .000 .020 .051 .020 .004 .030 .012

• 9 Year High (1972-1980)
  • 9 Year Low (1972-1980)

1. 21 Year High (1960-1980)
   1. 21 Year Low (1960-1980)

Chapter 2, Page 81 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 16C HISTORICAL COMPARISONS OF WIND FREQUENCY DISTRIBUTIONS DECEMBER Indian Point LaGuardia Bridgeport 1973 1979 1980 1973 1979 1980 1973 1979 1980 N .068 .101 .082 .081 .044 .093 .101 .058 .048 NNE .172 .071 .148 .077 .016 .044 .048 .019 .057 NE .162 .067 .124 .081 .046 .065 .097 .030 .040 ENE .041 .023 .061 .040 .046 .044 .044 .031 .020 E .015 .018 .023 .020 .013 .008 .024 .035 .012 ESE .004 .004 .011 .016 .009 .004 .016 .013 .000 SE .004 .008 .016 .024 .011 .000 .024 .015 .004 SSE .015 .030 .023 .040 .008 .004 .024 .012 .024 S .033 .100 .070 .093 .047 .056 .044 .023 .016 SSW .048 .091 .081 .040 .058 .056 .048 .020 .024 SW .049 .042 .057 .061 .122 .081 .048 .040 .069 WSW .019 .042 .024 .044 .071 .044 .028 .093 .085 W .026 .108 .031 .089 .206 .081 .141 .233 .157 WNW .044 .094 .057 .149 .144 .121 .141 .238 .157 NW .126 .120 .070 .069 .078 .157 .101 .066 .153 NNW .074 .062 .114 .069 .059 .133 .044 .036 .097 CALM .016 .000 .000 .008 .022 .008 .024 .038 .036 Chapter 2, Page 82 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 17 COMPARISON OF PERCENT WIND FREQUENCY DISTRIBUTIONS - SUMMER Wind 10 Meter Level 122 Meter Level Direction 1974 1979 1980 1979-80 1974 1979 1980 1979-80 N 3.66 5.53 7.15 6.34 6.54 7.25 8.19 7.72 NNE 10.71 10.27 10.07 10.17 10.67 11.22 10.27 10.74 NE 15.89 12.58 13.68 13.13 6.40 4.85 5.31 5.08 ENE 6.20 7.46 7.78 7.62 2.41 1.88 2.26 2.07 E 2.85 2.68 3.27 2.97 1.83 1.88 1.60 1.74 ESE 1.59 1.59 1.97 1.78 1.66 1.00 1.28 1.14 SE 2.20 1.31 1.77 1.54 2.91 1.61 1.83 1.72 SSE 2.85 2.52 2.09 2.30 2.88 2.40 2.86 2.63 S 8.47 13.28 9.05 11.17 13.04 18.49 14.20 16.36 SSW 11.86 17.39 12.75 15.07 9.82 12.74 9.95 11.35 SW 12.23 9.14 8.17 8.65 10.64 10.22 8.32 9.28 WSW 3.63 3.26 4.08 3.67 5.56 5.14 5.21 5.18 W 3.08 4.28 4.95 4.62 4.44 4.12 5.24 4.68 WNW 2.61 2.74 3.88 3.31 4.17 3.85 4.37 4.11 NW 3.19 2.43 3.65 3.04 5.05 6.84 7.71 7.27 NNW 2.41 2.77 3.59 3.18 3.79 5.71 8.28 6.99 VAR. 6.51 0.00 0.00 0.00 1.56 0.00 0.00 0.00 CALM 0.07 0.00 0.00 0.00 0.14 0.00 0.00 0.00 MISS. 0.03 0.77 2.09 1.43 6.50 0.79 3.13 1.96 NO VALID HOURS 2951 4411 4407 8818 2952 4413 4373 8786 % HRS IN DISTR 100. 99.9 99.8 99.8 100. 99.9 99.0 99.5 Chapter 2, Page 83 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 18 COMPARISON OF PERCENT WIND FREQUENCY DISTRIBUTIONS - WINTER Wind 10 Meter Level 122 Meter Level Direction 1974 1979 1980 1979-80 1974 1979 1980 1979-80 N 7.19 10.73 11.90 11.32 7.23 11.15 10.97 11.06 NNE 13.67 10.34 10.05 10.20 13.67 13.22 14.33 13.78 NE 12.22 11.45 15.11 13.29 5.63 4.79 4.90 4.85 ENE 4.75 5.90 6.30 6.10 1.93 1.80 1.79 1.79 E 1.59 2.86 3.02 2.94 1.60 1.45 1.79 1.62 ESE 0.95 0.71 0.85 0.78 1.16 0.92 1.65 1.29 SE 0.99 0.67 1.26 0.96 1.36 1.75 1.44 1.60 SSE 1.81 1.98 1.49 1.73 1.79 1.82 1.76 1.79 S 6.44 6.50 5.63 6.06 8.67 10.57 8.26 9.41 SSW 7.39 8.11 7.65 7.88 8.25 8.64 8.04 8.34 SW 5.76 6.24 5.72 5.98 9.36 5.87 5.59 5.73 WSW 2.58 4.31 3.04 3.67 3.51 3.18 3.37 3.27 W 3.32 4.77 4.01 4.39 2.94 3.78 4.14 3.96 WNW 5.42 5.90 4.62 5.26 5.08 6.24 5.24 5.74 NW 11.88 9.63 7.10 8.36 14.01 14.56 13.46 14.01 NNW 8.80 8.80 12.07 10.44 10.42 10.27 13.28 11.78 VAR. 4.23 0.00 0.00 0.00 1.18 0.00 0.00 0.00 CALM 0.66 0.00 0.00 0.00 0.04 0.00 0.00 0.00 MISS. 2.80 1.11 0.18 0.64 5.50 0.00 0.00 0.00 NO VALID HOURS 4967 4341 4368 8709 4924 4342 4368 8710 % HRS IN DISTR 97.6 99.9 100. 100. 96.8 100.0 100. 100. Chapter 2, Page 84 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.2.4.3 Diurnal Wind Direction Distributions Seasonal diurnal distributions of the resultant wind directions for the combined 1979-1980 data period are shown in comparison to the 1973-1974 data collection period in Table 19 and in Figures 17 and 18. The diurnal patterns with the exception of the summer season at the 122 meter level from the 2300 to 0900 are nearly identical for the 1979-80 data set and the historical 1973-74 data set. The deviation of the 122 meter level during the nocturnal hours is also consistent when considered with respect to the summer wind frequency shift at the 122 meter level to a sharply defined south wind maximum. It is concluded on the basis of the diurnal wind distributions that the patterns at all levels are consistent with the data base for FSAR 2 at all measurement levels with consideration for normal climatological variations. 3.2.4.4 Wind Speed Distributions All variable valid wind speeds at the 10 meter and 122 meter levels have been evaluated on a seasonal basis to determine their diurnal characteristics and the cumulative probability distributions. The results of these analyses are shown in Tables 20 and 21 for the summer season and Tables 22 and 23 for the winter season. For visual comparison, the diurnal variability is shown in Figure 19. The probability distributions are shown in Figure 20. In this latter Figure, the annual cumulative probabilities have been included. These curves were generated by combining the cumulative points in Tables 20-23. The maximum average diurnal wind speeds from this data set can be compared with those of the 1973-1974 season as shown below in Table 24. There are no significant differences. Chapter 2, Page 85 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 19 COMPARISON OF DIURNAL RESULTANT WIND DIRECTIONS 10 Meter Level 122 Meter Level Summer Winter Summer Winter Time 1974 1979-80 1973-74 1979-80 1974 1979-80 1973-74 1979-80 0100 051 034 357 359 351 300 327 329 0200 054 036 358 002 008 313 328 333 0300 050 036 359 003 007 321 329 334 0400 042 043 360 004 355 325 334 335 0500 048 048 001 003 004 327 335 335 0600 050 041 000 007 007 331 336 337 0700 042 040 001 007 011 338 339 338 0800 017 023 357 004 012 341 340 338 0900 337 348 352 356 350 334 337 339 1000 293 302 343 351 317 312 336 336 1100 258 269 335 344 260 279 325 331 1200 246 259 329 336 241 272 320 325 1300 246 247 327 336 248 261 319 321 1400 246 249 323 333 246 260 316 316 1500 233 246 322 332 233 251 313 316 1600 228 244 316 336 221 245 306 315 1700 230 233 321 337 223 239 307 312 1800 226 237 327 340 222 239 303 313 1900 232 245 330 345 233 239 308 317 2000 290 346 340 348 248 250 313 320 2100 049 033 345 355 259 263 315 325 2200 049 018 347 358 277 276 319 329 2300 055 033 351 358 297 282 321 331 2400 045 037 354 360 317 287 321 330 Chapter 2, Page 86 of 119 Revision 22, 2010

IP2 FSAR UPDATE Job Number: 01-4122-00 TABLE 20 YORK RESEARCH CORPORATION ONE RESEARCH DRIVE, STAMFORD, CONNECTICUT 06906 CLIENT: CON EDISON CO. OF NY SITE: INDIAN POINT (10M) PARAMETER: WIND SPEED (SUMMER SEASON) UNITS: MPH DIURNAL ANALYSIS MAY 1, 79,80 - OCT. 31, 79,80 OBSERV PARAM PARAM MAX VALID

      #               AVG          STD DEV            VALUE            RANGE              PTS 1              2.775           1.643           10.500            10.000             362 2              2.748           1.634            9.000             8.500             362 3              2.780           1.643           10.000             9.500             362 4              2.805           1.713           10.000             9.400             362 5              2.748           1.615            9.500             8.900             359 6              2.813           1.790           10.000             9.500             358 7              2.936           1.867           13.000            12.500             358 8              3.103           1.806           12.000            11.400             363 9              3.323           1.899           10.500             9.900             363 10              3.481           1.663           10.000             9.400             361 11              3.809           1.712           10.500             9.900             362 12              3.969           1.618           10.000             9.400             363 13              4.133           1.801           14.000            13.400             363 14              4.104           1.825           12.000            11.000             362 15              4.019           1.833           12.500            11.500             361 16              3.825           1.804           13.000            12.400             359 17              3.562           1.759           12.000            11.400             360 18              3.076           1.641            9.000             8.500             361 19              2.669           1.609           12.000            11.500             362 20              2.593           1.634           10.000             9.500             361 21              2.706           1.763           12.000            11.500             361 22              2.702           1.752           10.000             9.500             361 23              2.710           1.744           10.000             9.500             361 24              2.694           1.684           10.000             9.500             362 TOTAL             3.171           1.802           14.000            13.500            8669 CUMULATIVE PROBABILITY DISTRIBUTION MAY 1, 79, 80 - OCT 31, 79, 80 CUMULATIVE         CUMULATIVE CATEGORY     CATEGORY        CATEGORY                            POINTS            PERCENT OP. LIMIT      POINTS         PERCENT                            < LIMIT            < LIMIT 0.500           39                0.4                            39                0.450 1.000         1177            13.6                              1216              14.025 2.000         1996            23.0                              3212              37.047 3.000         1977            22.8                              5189              59.850 5.000         2474            28.5                              7663              88.385 7.000          772            8.9                               8435              97.290 9.000          181            2.1                               8616              99.377 12.000           48            0.6                               8664              99.931 16.000           5             0.1                               8669              99.988 23.000           0             0.0                               8669              99.988 30.000           0             0.0                               8669              99.988 40.000           0             0.0                               8669              99.988 50.000           0             0.0                               8669              99.988 70.000           0             0.0                               8669              99.988 NUMBER OF VALID DATA POINTS = 8669 NUMBER OF MISSING DATA POINTS = 163 REPRESENTING 98.2-PERCENT VALID DATA Chapter 2, Page 87 of 119 Revision 22, 2010

IP2 FSAR UPDATE Job Number: 01-4122-00 TABLE 21 YORK RESEARCH CORPORATION ONE RESEARCH DRIVE, STAMFORD, CONNECTICUT 06906 CLIENT: CON EDISON CO. OF NY SITE: INDIAN POINT (122M) PARAMETER: WIND SPEED (SUMMER SEASON) UNITS: MPH DIURNAL ANALYSIS MAY 1, 79,80 - OCT. 31, 79,80 OBSERV PARAM PARAM MAX VALID

      #               AVG          STD DEV            VALUE             RANGE              PTS 1              7.821           5.161            30.000            29.500             357 2              7.595           5.070            24.000            23.500             356 3              7.581           5.153            26.000            25.400             355 4              7.413           5.097            27.000            26.500             356 5              7.233           5.043            32.000            31.400             353 6              7.436           5.383            35.000            34.500             354 7              7.499           5.616            47.000            48.400             355 8              7.332           5.442            47.000            48.400             359 9              7.717           5.320            38.000            37.400             360 10              8.086           5.399            33.000            32.000             362 11              8.881           5.312            30.000            28.500             361 12              9.595           5.217            28.000            26.000             359 13             10.232           5.394            40.000            39.000             358 14             10.409           5.450            36.000            34.500             359 15             10.762           5.485            30.000            28.000             357 16             11.087           5.397            32.000            31.000             355 17             11.262           5.184            31.000            30.400             353 18             10.764           4.780            25.000            23.500             354 19             10.266           4.919            33.000            32.000             354 20              9.879           5.015            28.000            27.000             354 21              9.256           5.022            27.000            28.400             356 22              8.781           5.299            28.000            27.400             356 23              8.383           5.140            28.000            27.400             356 24              8.070           4.944            26.000            25.500             357 TOTAL             8.888           5.388            47.000            46.500            8556 CUMULATIVE PROBABILITY DISTRIBUTION MAY 1, 79, 80 - OCT 31, 79, 80 CUMULATIVE         CUMULATIVE CATEGORY     CATEGORY        CATEGORY                             POINTS            PERCENT OP. LIMIT      POINTS         PERCENT                             < LIMIT            < LIMIT 0.500           7             0.1                                  7                0.082 1.000          169            2.0                                 176               2.057 2.000          511            6.0                                 687               8.029 3.000          638            7.5                                1325              15.484 5.000         1303            15.2                               2628              30.712 7.000         1205            14.1                               3833              44.794 9.000         1132            13.2                               4965              58.023 12.000         1565            18.3                               6530              76.312 16.000         1258            14.7                               7788              91.013 23.000          652            7.6                                8440              98.633 30.000           99            1.2                                8539              99.790 40.000           15            0.2                                8554              99.965 50.000           2             0.0                                8556              99.988 70.000           0             0.0                                8556              99.988 NUMBER OF VALID DATA POINTS = 8556 NUMBER OF MISSING DATA POINTS = 276 REPRESENTING 98.9-PERCENT VALID DATA Chapter 2, Page 88 of 119 Revision 22, 2010

IP2 FSAR UPDATE Job Number: 01-4122-00 TABLE 22 YORK RESEARCH CORPORATION ONE RESEARCH DRIVE, STAMFORD, CONNECTICUT 06906 CLIENT: CON EDISON CO. OF NY SITE: INDIAN POINT (10M) PARAMETER: WIND SPEED (WINTER SEASON) UNITS: MPH DIURNAL ANALYSIS NOV 1, 79, 80 - APR 30, 79, 80 OBSERV PARAM PARAM MAX VALID

       #             AVG           STD DEV             VALUE            RANGE              PTS 1            4.508            2.958            21.000            20.500             362 2            4.462            2.940            14.000            13.500             362 3            4.524            3.163            18.000            17.500             362 4            4.453            3.043            16.000            15.500             361 5            4.298            2.886            15.000            14.400             361 6            4.292            2.961            15.000            14.500             362 7            4.461            3.102            18.000            17.500             363 8            4.556            3.141            15.000            14.400             363 9            4.906            3.212            17.000            16.400             363 10            5.153            3.182            16.000            15.400             363 11            5.310            3.050            16.000            15.500             361 12            5.556            2.995            16.000            15.400             362 13            5.731            3.163            18.000            17.400             362 14            5.702            3.240            16.500            15.900             362 15            5.589            3.109            17.000            16.400             361 16            5.336            3.243            20.000            19.400             360 17            5.030            3.084            20.000            19.500             360 18            4.671            2.858            15.000            14.400             360 19            4.611            2.906            16.000            15.500             361 20            4.502            2.883            15.000            14.400             361 21            4.500            2.972            16.000            15.400             362 22            4.457            2.970            16.000            15.400             361 23            4.381            2.855            16.000            15.400             361 24            4.398            2.874            18.000            17.500             362 TOTAL           4.808            3.088            21.000            20.500            8678 CUMULATIVE PROBABILITY DISTRIBUTION NOV 1, 79, 80 - APR 30, 79, 80 CUMULATIVE         CUMULATIVE CATEGORY     CATEGORY        CATEGORY                             POINTS              PERCENT OP. LIMIT     POINTS         PERCENT                             < LIMIT             < LIMIT 0.500       18              0.2                                18                0.207 1.000       788             9.1                                806               9.287 2.000       1332            15.3                               2138              24.634 3.000       1178            13.6                               3316              38.207 5.000       2122            24.4                               5438              62.657 7.000       1591            18.3                               7029              80.989 9.000       841             9.7                                7870              90.679 12.000      614             7.1                                8484              97.753 16.000      181             2.1                                8665              99.839 23.000      13              0.1                                8678              99.988 30.000      0               0.0                                8678              99.988 40.000      0               0.0                                8678              99.988 50.000      0               0.0                                8678              99.988 70.000      0               0.0                                8678              99.988 NUMBER OF VALID DATA POINTS = 8678 NUMBER OF MISSING DATA POINTS = 34 REPRESENTING 99.6-PERCENT VALID DATA Chapter 2, Page 89 of 119 Revision 22, 2010

IP2 FSAR UPDATE Job Number: 01-4122-00 TABLE 23 YORK RESEARCH CORPORATION ONE RESEARCH DRIVE, STAMFORD, CONNECTICUT 06906 CLIENT: CON EDISON CO. OF NY SITE: INDIAN POINT (122M) PARAMETER: WIND SPEED (WINTER SEASON) UNITS: MPH DIURNAL ANALYSIS NOV 1, 79, 80 - APR 30, 79, 80 OBSERV PARAM PARAM MAX VALID

       #            AVG           STD DEV             VALUE            RANGE             PTS 1              10.919           6.869             49.000            48.500          359 2              10.669           5.562             38.000            37.400          359 3              10.636           6.799             40.000            39.400          358 4              10.297           6.804             37.500            36.900          358 5              10.075           6.354             34.000            33.400          358 6              9.832            6.485             32.000            31.500          358 7              9.902            6.811             44.000            43.500          356 8              10.283           7.067             38.000            37.400          356 9              10.525           6.978             36.000            35.400          358 10             10.596           7.032             37.000            36.400          360 11             11.086           6.961             36.000            35.000          362 12             11.828           7.011             34.000            33.500          362 13             12.309           7.218             48.000            47.000          362 14             12.530           7.250             45.000            44.400          362 15             12.727           7.184             43.000            42.000          362 16             12.635           6.992             40.000            38.500          362 17             12.302           6.778             42.000            41.000          361 18             11.921           6.035             32.000            31.400          361 19             11.645           6.061             32.000            31.400          360 20             11.398           6.093             37.000            36.400          360 21             11.106           6.216             40.000            39.400          360 22             10.658           6.206             39.000            38.400          358 23             10.488           6.182             36.000            35.000          358 24             10.551           6.178             41.000            40.500          359 TOTAL          11.124           6.726             49.000            48.500          4629 CUMULATIVE PROBABILITY DISTRIBUTION NOV 1, 79, 80 - APR 30, 79, 80 CUMULATIVE        CUMULATIVE CATEGORY     CATEGORY       CATEGORY                             POINTS             PERCENT OP. LIMIT     POINTS        PERCENT                             < LIMIT            < LIMIT 0.500       6              0.1                                6                0.070 1.000       105             1.2                               111              1.286 2.000       418             4.8                               529              6.130 3.000       458             5.3                               987              11.437 5.000       910             10.5                              1897             21.981 7.000       895             10.4                              2792             32.352 9.000       1009            11.7                              3801             44.044 12.000      1622            18.8                              5423             62.839 16.000      1518            17.6                              6941             80.429 23.000      1230            14.3                              8171             94.681 30.000      372             4.3                               8513             98.992 40.000      78              0.9                               8621             99.896 50.000      8               0.1                               8629             99.988 70.000      0               0.0                               8629             99.988 NUMBER OF VALID DATA POINTS = 8629 NUMBER OF MISSING DATA POINTS = 83 REPRESENTING 99.0-PERCENT VALID DATA Chapter 2, Page 90 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 24 MAXIMUM DIURNAL WIND SPEEDS (MPH) Season Level (M) 1973-1974 1979-1980 Summer 10 4.0 4.1 Summer 120 11.0 11.3 Winter 10 5.0 5.7 Winter 122 13.0 12.7 The median wind speeds extracted from Figure 20 at the 50 percent probability level are 1.1 M/S, 1.7 M/S and 1.4 M/S at the 10 meter level for the summer season, winter season, and annual basis, respectively. At the 122 meter level, these values are 3.4 M/S, 4.4 M/S and 3.9 M/S on the summer, winter and annual basis, respectively. These values bracket those presented by Kaplin and Laznow (1972) for the IP3 Tower. Corrected for exposure elevation no significant change would be expected between the two sets of values. The variation of winds during the 1979-1980 season are consistent with data obtained during the 1973-1974 operational period. There is no reason to expect any significant variations with respect to the meteorology as used in FSAR 2. 3.2.4.5 Wind Velocities and Atmospheric Stability 3.2.4.5.1 Joint Frequency Distribution of Wind Direction and Stability Stability categorizations as referenced in this study are in accordance with NRC Pasquill Tables as derived from local temperature change with elevation. Except as noted, actual temperature measured gradients have been converted to °C/100M directly from temperature difference values (°F) per difference between sensor height levels. Tables 25, 26 and 27 show the summary frequency distributions for the 10 meter level of wind direction and stability categories for the 1979-1980 data collection period. The tables show the annual, summer season and winter season summaries respectively. Chapter 2, Page 91 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 25 ANNUAL

SUMMARY

OF WIND DIRECTION PERCENT FREQUENCY DISTRIBUTION AS A FUNCTION OF STABILITY - 10M LEVEL (JANUARY 1, 1979 - DECEMBER 31, 1980) Stability Class Wind Direction A B C D E F G N 1.28 0.36 0.48 3.39 2.67 0.50 0.09 NNE 1.76 0.40 0.46 3.15 3.33 0.80 0.17 NE 0.63 0.35 0.58 4.22 4.66 2.12 0.40 ENE 0.06 0.07 0.17 1.59 2.61 1.84 0.43 E 0.01 0.03 0.03 0.64 1.49 0.59 0.11 ESE 0.01 0.01 0.01 0.27 0.73 0.21 0.04 SE 0.03 0.01 0.02 0.23 0.67 0.26 0.02 SSE 0.09 0.03 0.04 0.45 1.04 0.31 0.05 S 2.04 0.25 0.29 1.74 3.39 0.76 0.11 SSW 2.58 0.51 0.38 2.14 5.04 0.72 0.05 SW 1.16 0.33 0.35 1.89 3.03 0.51 0.03 WSW 0.49 0.17 0.16 0.96 1.44 0.39 0.02 W 0.56 0.22 0.17 1.40 1.64 0.43 0.06 WNW 0.47 0.15 0.26 1.64 1.49 0.21 0.03 NW 0.70 0.31 0.32 2.36 1.85 0.10 0.01 NNW 0.80 0.40 0.49 3.26 1.60 0.17 0.04 CALM 0.00 0.00 0.00 0.00 0.00 0.00 0.00 MISSING 0.12 0.05 0.03 0.21 0.51 0.15 0.02 TOTAL % 12.80 3.66 4.23 29.56 37.17 10.08 1.69 NO. OF HOURS 2244 641 742 5183 6519 1768 297 Chapter 2, Page 92 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 26

SUMMARY

OF WIND DIRECTION PERCENT FREQUENCY DISTRIBUTION AS A FUNCTION OF STABILITY SUMMER SEASON - 10M LEVEL (MAY 1, 1979, 80 - OCTOBER 31, 1979, 80) Stability Class Wind Direction A B C D E F G N 1.68 0.26 0.37 1.25 2.06 0.57 0.07 NNE 2.65 0.42 0.43 2.90 2.41 1.01 0.18 NE 0.58 0.31 0.46 3.46 4.44 3.17 0.35 ENE 0.11 0.10 0.24 1.38 2.66 2.62 0.39 E 0.02 0.07 0.01 0.57 1.57 0.61 0.05 ESE 0.01 0.01 0.00 0.31 1.01 0.36 0.06 SE 0.05 0.02 0.01 0.17 0.84 0.40 0.02 SSE 0.15 0.06 0.05 0.50 1.07 0.40 0.08 S 3.32 0.36 0.43 2.47 3.58 0.85 0.05 SSW 4.10 0.75 0.59 2.93 5.70 0.85 0.01 SW 1.84 0.49 0.48 2.23 3.03 0.51 0.05 WSW 0.87 0.20 0.18 0.94 1.05 0.34 0.00 W 0.88 0.28 0.19 1.38 1.42 0.34 0.07 WNW 0.80 0.09 0.25 0.94 1.03 0.15 0.05 NW 1.05 0.19 0.17 0.84 0.63 0.10 0.02 NNW 0.78 0.19 0.24 0.97 0.74 0.20 0.02 CALM 0.00 0.00 0.00 0.00 0.00 0.00 0.00 MISSING 0.22 0.06 0.01 0.31 0.68 0.22 0.03 TOTAL % 19.11 3.86 4.11 23.54 33.92 12.69 1.48 NO. OF HOURS 1687 341 363 2078 2994 1120 131 Chapter 2, Page 93 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 27

SUMMARY

OF WIND DIRECTION PERCENT FREQUENCY DISTRIBUTION AS A FUNCTION OF STABILITY WINTER SEASON - 10M LEVEL (NOVEMBER 1, 1979, 80 - APRIL 30, 1979, 80) Stability Class Wind Direction A B C D E F G N 0.87 0.46 0.60 5.56 3.28 0.44 0.11 NNE 0.85 0.38 0.49 3.41 4.26 0.59 0.16 NE 0.86 0.40 0.69 4.99 4.88 1.06 0.45 ENE 0.01 0.05 0.10 1.79 2.56 1.06 0.48 E 0.00 0.00 0.06 0.72 1.40 0.56 0.18 ESE 0.00 0.01 0.01 0.23 0.45 0.06 0.02 SE 0.02 0.00 0.02 0.29 0.49 0.11 0.01 SSE 0.02 0.01 0.03 0.40 1.01 0.23 0.02 S 0.75 0.13 0.14 1.01 3.19 0.68 0.17 SSW 1.04 0.28 0.17 1.34 4.37 0.59 0.08 SW 0.48 0.16 0.22 1.55 3.03 0.51 0.02 WSW 0.10 0.13 0.14 0.99 1.84 0.45 0.03 W 0.24 0.15 0.14 1.42 1.86 0.52 0.06 WNW 0.13 0.22 0.28 2.34 1.95 0.28 0.02 NW 0.34 0.42 0.47 3.90 3.09 0.10 0.00 NNW 0.83 0.62 0.75 5.58 2.47 0.14 0.06 CALM 0.00 0.00 0.00 0.00 0.00 0.00 0.00 MISSING 0.02 0.03 0.05 0.11 0.33 0.09 0.01 TOTAL % 6.39 3.44 4.35 35.65 40.47 7.44 1.91 NO. OF HOURS 557 300 379 3105 3525 648 161 Chapter 2, Page 94 of 119 Revision 22, 2010

IP2 FSAR UPDATE There are distinctive seasonal biases that coincide with variations in wind direction occurrence frequencies. It will be seen that these biases are consistent with variations in stability and wind speed on diurnal basis. 3.2.4.5.2 Frequency of Occurrence of Stability Categories Table 28 shows a summary of historical comparison of percent occurrence of stability categories between the various reporting periods for the IP3 Tower and of the 122M Tower (IP4) for 1973/74. (Based on concurrent wind speed and temperature gradients). The former gradients were based on temperature differences from the 30M and 2M levels while the latter were based on differential measurements between the 60M and 10M levels. On an annual basis there is generally good agreement between the results for the 122M Tower and the IP3 Tower composite year with temperature correction (FSAR 3, Supplement 13, 16). The variation in percentages at the stability extreme A and G are most probably related to the lower gradient base level of measurement on the IP3 Tower - 2 meters. One would expect higher or lower temperatures closer to the ground with less accuracy in the thermal adjustment factor in these extreme ranges. Table 29 shows a similar comparison for the 122 Meter Meteorological Tower for the 1973-1974 data collection period and the 1979-80 data collection period. These results are based on current wind speed data for the tower levels as noted. The 122M is shown based on two gradient differences: 122-10M and 122-60M. The percent occurrences of stability categories are sharply defined functions of season and, for the upper level, the vertical defined functions of season and, for the upper level, the vertical temperature gradient. This is apparent in both data sets. There are seasonal differences in the two data sets particularly in the A and G stability category extremes at the 122M level, however, is noted that the percent occurrences in the 1973-1974 data set are based on only one or less than ten observations where these percentage differentials are most extreme. Where more data points are available in each stability category as for the lower gradient level (61-10M), there is generally good agreement. Chapter 2, Page 95 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 28 HISTORICAL COMPARISONS OF PERCENT OCCURRENCE OF STABILITY Temperature Stability Class No. of Gradient (M) A B C D E F G Observ. Summer (1974) 61-10 21.44 5.46 4.37 27.92 30.21 9.28 1.31 2747 1973 (IP3)* 25.52 2.62 3.64 17.38 26.81 13.39 10.63 2935 Winter (1973/74) 61-10 4.00 1.92 2.13 23.93 53.24 12.53 2.25 4797 1973 (IP3)* 21.49 3.52 3.74 23.29 23.76 14.24 9.96 4229 Annual (1973/74) 61-10 10.35 3.21 2.94 25.38 44.86 11.35 1.91 7544 1970/72 (IP3)1 6.42 2.55 2.23 31.19 38.75 11.25 3.16 8366 1970 (IP3)* 21.68 2.20 3.39 33.35 24.75 9.01 5.62 NA 1971 (IP3)* 19.17 2.75 2.97 22.79 30.87 11.69 9.75 NA 1970/72 (IP3)1* 15.52 1.74 2.82 28.38 25.42 12.68 9.03 8366 1973 (IP3)** 23.14 3.16 3.70 20.87 25.02 13.89 10.23 NA

• Concurrent basis 1

Composite year with temperature correction - concurrent basis (FSAR 3) 1* Composite year concurrent basis (FSAR 3)

• • March-December only concurrent basis NA Not Available NOTE: Gradient for IP3 Tower: 30-2 M Chapter 2, Page 96 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 29 COMPARISON OF PERCENT OCCURRENCE OF STABILITY ON 122 METER TOWER Wind Freq Stability Class No. of Temperature Date LVL (M) A B C D E F G Observ Gradient(M) Summer 1974 10 21.44 5.47 4.37 27.92 30.21 9.28 1.31 2747 61 - 10 1979/80 10 19.34 3.93 4.22 23.96 34.35 12.73 1.48 8557 60 - 10 1974 122 0.51 0.66 2.15 47.40 41.54 7.72 0.04* 2747 122 - 10 1979/80 122 4.54 5.50 5.81 42.84 34.95 6.23 0.14** 7838 122 - 10 1974 122 0.04* 0.04* 0.11** 20.13 74.01 5.42 0.25** 2747 122 - 61 1979/80 122 0.17** 0.42 1.49 68.09 26.72 2.82 0.29 7838 122 - 60 Winter 1973/74 10 4.00 1.92 2.13 23.93 53.24 12.53 2.25 4797 61 - 10 1979/80 10 6.44 3.47 4.37 35.79 40.55 7.46 1.92 8648 60 - 10 1973/74 122 3.17 1.70 1.63 41.46 48.86 3.09 0.08** 4797 122 - 10 1979/80 122 0.59 1.18 2.78 56.07 34.75 4.29 0.31 8594 122 - 10 1973/74 122 0.02* 0.31 1.19 41.94 49.59 6.82 0.13** 4797 122 - 61 1979/80 122 0.28 0.14 0.37 71.89 25.35 1.84 0.13 8594 122 - 60 Annual 1973/74 10 10.35 3.21 2.94 25.38 44.86 11.35 1.91 7544 61 - 10 1979/80 10 12.86 3.70 4.30 29.90 37.47 10.08 1.70 17205 60 - 10 1973/74 122 2.63 1.33 1.82 43.62 46.20 4.77 0.07 7544 122 - 10 1979/80 122 2.48 3.24 4.23 49.76 34.84 5.22 0.23 16432 122 -10 1973/74 122 0.03 0.08 0.80 34.00 58.48 6.31 0.17 7544 122 - 61 1979/80 122 0.23 0.27 0.91 70.08 26.00 2.31 0.21 16432 122 - 60 It may be inferred that, except as stated for the reasons noted, the percent frequency of stability classes with the existing tower system is representative and consistent with data referenced in FSAR 2.

• Single observations
  • Less than ten observations Chapter 2, Page 97 of 119 Revision 22, 2010

IP2 FSAR UPDATE 3.2.4.5.3 Average Wind Speed and Diurnal Variation as a Function of Stability Categories Tables 30, 31, and 32 show the average wind speed and number of observations as functions of time of day and stability category for the summer and winter seasons of the combined 1979 and 1980 data collection period. The results are derived from valid wind speeds measured at 10M relative to the temperature difference 60 - 10M and 122M wind speeds based on the temperature difference between 122-10M and 122-60M. The latter gradient was generated by subtraction of gradient levels: (122-60) = (122-10) - (60-10) These tables indicate a distinctive diurnal pattern to the stability. During the summer season at the 10M level for all practical purposes G stability does not occur between 0700 to 1900 EST and F stability does not occur between 0900 and 1400. During the nocturnal hours between 1900 to 0600, for all practical purposes A, B, and C stability categories do not appear. Stability Category A appears from 0700 to 1800 EST and is the dominant gradient between 0900 and 1600 EST. Out of 2,855 observations during this time interval, the percent frequency of occurrences were 52.9 for A, 8.1 for B, 8.4 for C, 24.2 for D, 6.1 for E, 0.2 for F, and 0.04 for G. In the winter season at 10M, the A, B, and C stability categories do not appear during the nocturnal hours from 1900 to 0600 (with random singular exceptions). D is the dominant day time stability category and E is the dominant nighttime category. Somewhat similar patterns are found with the upper level gradients based on the 122-10M and 122-60M temperature gradients. Between these latter two gradients TABLE 30 Diurnal Variation of Stability Class and Wind Speed (concurrent data) TABLE 31 Diurnal Variation of Stability Class and Wind Speed (concurrent data) TABLE 32 Diurnal Variation of Stability Class and Wind Speed (concurrent data) Chapter 2, Page 98 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 33 COMPARISONS OF AVERAGE WIND SPEEDS (MPH) AS A FUNCTION OF STABILITY Year/ Anemom Stability Class Temperature Season LVL A B C D E F G Gradient (M) Summer 1974 10M 4.2 4.4 4.0 3.8 2.7 2.3 2.6 60 - 10 1979/80 10M 4.0 3.6 3.5 3.7 2.7 2.1 2.2 60 - 10 1974 122M 4.3** 14.3 12.3 9.5 7.7 4.5 4.5* 122 - 10 1979/80 122M 11.2 10.4 9.8 10.0 7.6 4.8 4.3** 122 - 10 1974 122M 3.5* 49.0* 3.8** 10.6 8.2 4.4 2.6** 122 - 60 1979/80 122M 9.3** 12.6 12.1 10.0 6.4 4.3 4.0 122 - 60 Winter 1973/74 10M 4.8 5.2 4.8 6.1 5.0 2.4 2.3 60 - 10 1979/80 10M 5.9 6.4 6.0 5.9 4.0 2.2 2.8 60 - 10 1973/74 122M 15.8 11.5 10.1 13.5 9.0 5.7 3.6** 122 - 10 1979/80 122M 13.9 12.4 13.9 12.9 8.6 5.6 8.6 122 - 10 1973/74 122M 16.0* 15.0 13.9 13.3 9.7 6.0 4.7** 122 - 60 1979/80 122M 19.9 21.0 20.4 12.3 7.8 7.2 6.3** 122 - 60 for the same data base the number of occurrences of A, B, and C stability categories in both summer and winter seasons are substantially reduced and become almost random when based on the 122-60M gradient. The obvious implication is that temperature gradient extremes are controlled by the surface level. This is consistent with Kaplin, et. al., 1974. At the upper levels, for the summer season, F stability has a distinct diurnal function. It rarely occurs during the daytime from 0800 to 2000 EST. The distribution of G stability during the summer season, while a diurnal function, is clearly biased to the early morning hours. It occurs more frequently with relation to the 122-60M gradient than the 122-10M gradient. These factors are consistent with a nocturnal cool air surface drainage flow. This routine drainage flow does not exist during the winter season, and the occurrence frequencies, while still diurnal, are clearly related to the local surface air temperatures. It is noted that during the winter seasons, drainage flow patterns are not routine occurrences. They can occur, however, during periods of weak pressure gradients. They are dependent on the horizontal and vertical temperature gradients that develop between the land (snow covered or bare ground) as well as the Hudson River (free water or ice bound). The average wind speeds as a function of stability category are shown in Table 33 for the 1973-1974 data collection period and the 1979-1980 data period. As noted, there

• Single observations

** Less than 15 observations Chapter 2, Page 99 of 119 Revision 22, 2010

IP2 FSAR UPDATE are some apparent anomalies with respect to the 122 meter level wind speeds. These are probably induced by the few data points available in 1973-1974 rather than a reality. Where the total number of occurrences were in a more reliable range, at the 10 meter and the 122 meter levels, there was good agreement between average wind speeds. Based on these data, it is concluded that within the range of normal climatological variations, there are no significant changes in the local meteorological parameters on a seasonal and annual basis. The existing meteorological data are consistent with data referenced in FSAR 2. 4.0

SUMMARY

4.1 METEROROLOGY 4.1.1 General The meteorology of the Indian Point site and its environs has been thoroughly studied over the span of years. For the past nine years the source of on-site meteorology has been the 122 Meter Meteorological Tower that became fully operative as of October 1, 1973. This tower is located at latitude: 41° 15 55 N and longitude 73° 57 08 W (N38 + 31.453 and E22 + 49.473 on the Indian Point Grid). Meteorological data from the 122 Meter Tower have, in previous studies and in this report, have been compared, in so far as possible, to those meteorological data, which were the data base for the FSAR 2 Report. 4.1.2 122 Meter Meteorological Tower System The 122 Meter Tower and support systems as presently comprised, maintained, and operated are in compliance with the meteorological measurement programs included in Regulatory Guides 1.23 and 1.97 and the criteria set forth in NUREG-0654, -0696, and -0737. The system is outlined in Figure 4. It consists of an instrumented 122 Meter Tower. The critical sensors are for winds at the 10, 60, and 122 meter levels; ambient temperature and dew point at the 10 meter level and temperature difference between the 60-10 meter and 122-10 meter levels. In addition, a precipitation gage is located within the tower complex. All sensor signals are carried to a trailer, which houses: Signal Conditioners Analog Recorders Data Acquisition System - Magnetic Tape Terminal Printer LED Satellite Displays for Control Rooms 2 and 3 Telephone Modems Dedicated Telephone Lines Air Conditioning Systems All systems are operated on primary AC voltage. A backup diesel generator within the complex provides for the automatic transfer of power if the primary source is cut off. Appropriate meteorological data are transmitted from the Meteorological Trailer to: Chapter 2, Page 100 of 119 Revision 22, 2010

IP2 FSAR UPDATE Reactor Control Rooms 2 (CON-EDISON) and 3 (PASNY) MIDAS and ARAC Computers The Emergency Control Center in addition to the computer systems and interrogation systems receives wind data from backup wind sensors. The status of the backup wind system and the MIDAS computer may be assessed by remote telephone interrogation. 4.1.3 Local Meteorological Characteristics All earlier studies and the data evaluated and included in this report indicates that the most important characteristic of the Indian Point area is the prevalence of winds from the north and south sectors. These winds are induced by meso-scale factors: terrain channeling at all times and drainage flow and land-sea circulations during periods of weak synoptic-cyclonic scale pressure gradient field. At all wind levels there are distinctive diurnal variation patterns to local winds as well as to the local winds in addition to the local atmospheric stability as determined by vertical temperature gradients related to Pasquill stability categories. Unstable A, B, and C categories are dominant daytime occurrences. Stable F and G categories are nocturnal occurrences. Because of the dominance of meso-scale factors in the Indian Point and lower Hudson Avenue Valley environs, persistent straight line flow of air from Indian Point is impossible. Paths of movement of air parcels are best generated by the use of local data on a real time basis. Recirculation of air parcels within a time frame of eight consecutive hours is a likely event. Within ten miles of Indian Point there are three zones, which indicated the probability of convergence and divergence of local surface air streams: Peekskill Bay Haverstraw Bay Tappan Zee 4.1.4 Conclusion From the evaluation of previous studies and the recent data years January 1, 1979 - December 31, 1980, it has been concluded that the meteorological data being collected at the 122 Meter Meteorological Tower are representative of the Indian Point site and are consistent with the original and expanded meteorological data basis of FSAR for Unit No. 2. All deviations of data at any given time (not otherwise specifically assigned to measurement techniques, methodologies, and, evaluation procedural changes to comply with existing Regulatory Guides) can be assigned to normal regional climatological variations in any given year on, at least, the synoptic-cyclonic meteorological scale.

5.0 REFERENCES

1) Halitsky, James, J. Laznow and D. Leahey, (1970): Wind Observations at Indian Point, Montrose and Bowline Point, 31 August 1968 to 30 June 1969. New York University, School of Engineering and Science, Department of Meteorology and Oceanography, G.S.L. Technical Report No. TR-70-3, February, 1970.

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IP2 FSAR UPDATE

2) Halitsky, James, E. J. Kaplin and J. Laznow (1971): Wind Observations at Indian Point, 26 November 1969 to October 1970. New York University, School of Engineering and Science, Department of Meteorology and Oceanography, G.S.L.

Technical Report No. TR 71-3, May, 1971.

3) Kaplin, Edward J. and J. Laznow (1972): Meteorological Observations at Indian Point, Trap Rock, Montrose Point and Cape Charles, 1 January 1970 to 31 December 1971. New York University, School of Engineering and Science, Department of Meteorology and Oceanography, G.S.L. Technical Report No. TR-73-1, December, 1972.

4) Kaplin, Edward J., J. Laznow and M. M. Wurmbrand (1972): Selection of Air Monitoring Locations, Calculations of Dispersion Patterns for Diffusion Overlays, and Recommendations for a Meteorological Program to Satisfy A.E.C. Safety Guide .23 at the Indian Point Complex. New York University, School of Engineering and Science, Department of Meteorology and Oceanography, G.S.L.

Technical Report No. TR-72-3, July, 1972.

5) Kaplin, Edward J. and C. E. Kitson (1974): Final Report - Meteorological Services and Data Evaluation Indian Point 100 Foot Tower. York Research Corporation Report NO. Y-8127, September, 1974.

6) United States Atomic Energy Commission Safety Guide 23 - Onsite Meteorological Programs, February 17, 1972.

7) Kaplin, Edward J., M. Kozenko and J. Kirschner (1974): Report No. 1, Meteorological, Ambient Salt Studies and Upper Air Studies at Indian Point, Buchanan, NY, York Research Corporation Report No. Y-8162/8166, January, 1974.

8) Kaplin, Edward J., C. E. Kitson, and M. Kozenko (1974): Final Report:

Meteorological, Upper Air and Ambient Salt Studies, York Research Corporation Report Nos. Y-8162 and Y-8166-2, October, 1974. (With Appendices A-N, I and II)

9) U.S. Nuclear Regulatory Commission (1980): NUREG-0654 (FEMA-FEP-1),

Criteria for Preparation and Evaluation of Radiological Emergency Response Plans and Preparedness in Support of Nuclear Power Plants: Appendix 2, Meteorological Criteria for Emergency Preparedness at Operating Nuclear Power Plants, January, 1980.

10) U.S. Nuclear Regulatory Commission (1980): Proposal Revision 1 to Regulatory Guide 1.23 - Meteorological Programs in Support of Nuclear Power Plants, September, 1980.

11) Kaplin, Edward J. and B. Wuebber (1979): Final Report, Surface Air Trajectory Analyses, York Services Corporation Report No. 1-1025, December, 1979.

12) Kaplin, Edward J. (1979): Executive Summary, Final Report, Surface Air Trajectory Analysis, York Services Corporation Report No. 1-1025-1, December, 1979.

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13) Kaplin, Edward J. and B. Wuebber (1981): Final Report Indian Point Trajectory II Study, York Services Corporation Report No. 01-4013, July, 1981.

14) FSAR: Final Facility Description and Safety Analysis Report, Consolidated Edison Company of New York, Inc., Nuclear Generating Station No. 2.

15) FSAR: Final Facility Description and Safety Analysis Report, Consolidated Edison Company of New York, Inc., Nuclear Generating Station No. 3.

Chapter 2, Page 103 of 119 Revision 22, 2010

IP2 FSAR UPDATE NOTE: This information is classified as Historical Information APPENDIX 2B INDIAN POINT FSAR UPDATE REVISED Prepared for: Consolidated Edison Company, Inc. of New York 4 Irving Place New York, NY 10003 Submitted by: WOODWARD-CYCLE CONSULTANTS 201 Willowbrook Boulevard Wayne, New Jersey 07470 16 April 1982 Chapter 2, Page 104 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE OF CONTENTS Page No. INTRODUCTION Physiography 2.7A-1 General Geology and Tectonics 2.7A-2 Geology of the Hudson Highlands 2.7A-3 Geology of the Manhattan Prong 2.7A-4 The Geology of the Newark Basin 2.7A-7 History of Brittle Deformation 2.7A-8 Recent Seismicity 2.7A-12 CONCLUSION 2.7A-15 REFERENCES R-1 GLOSSARY G-1 TABLES Table 1 - Geologic Time Scale Table 2 - Stratigraphic Correlation Chart Table 3 - Geologic History in the Croton Falls Area FIGURES Figure 1 - Location Map Figure 2 - Seismotectonic Map Chapter 2, Page 105 of 119 Revision 22, 2010

IP2 FSAR UPDATE INDIAN POINT FSAR UPDATE INTRODUCTION This report is intended as an update and synthesis of previous geologic reports on the area surrounding Con-Edisons Indian Point nuclear facilities. The report reflects current thinking on the geology, structure, tectonic history, neotectonics and recent seismicity in the region. Main sources of information include Ratcliffes (1976) final report on the Ramapo fault system, the Dames and Moore Geotechnical Investigation of the Ramapo Fault (1977), and the recent literature on the subject of the geology of the Manhattan and Reading Prongs. Con Edisons Indian Point power plants are located in Buchanan, New York, on the east bank of the Hudson River. The site is situated in the central portion of the Peekskill Quadrangle. Physiography The rocks in the vicinity of the Indian Point generating stations belong to three geologic provinces, the Hudson Highlands, the Manhattan Prong and the Newark Basin (Figure 1). Rocks that outcrop within the provinces range in age from Precambrian through Triassic (Jurassic?). The landscape consists of northeast trending ridges and rather broad swampy valleys. Ridges are supported by bedrock and tend to follow prominent generally northeast, structural trends. Valley walls tend to be steep, the result of modification by Pleistocene glaciation. Elevations in the area reach a maximum of 1000 feet, and range from 50 to 300 feet above sea level in low lying areas. General Geology and Tectonics The eastern third of the North American continent has been the site of episodic tectonism since Precambrian time (see Table1). Paleozoic aged tectonism has molded a broad geologically varying zone known as the Appalachian orogen. The core of the orogen is marked by intrusive rocks modified by ductile and brittle deformation and regional metamorphism. The Indian Point site lies within the Manhattan Prong of the Appalachian Mountains. The earliest recognized event in the area occurred in Precambrian time, and is known as the Grenville Orogeny. The Grenville orogeny, dated at 1.1 b.y., produced brittle and ductile deformation accompanied by regional granulite facies metamorphism and intrusive activity. The deformation and metamorphism affected the rocks of the Reading Prong and the Precambrian rocks of the Manhattan Prong. The Grenville events are not tectonically or temporarily related to the development of the Appalachian Orogen, which began in the latest Precambrian. The earliest tectonic activity in the Appalachian Orogen probably involved latest Precambrian continental rifting and associated intrusive activity. The opening of the Proto-Atlantic Ocean (Iapetus) set the stage for the development of the Appalachian geosyncline. The geosyncline received sediments from earliest Precambrian through the mid-Ordovician time. The Taconic orogeny occurred in Mid-Ordovician time, and Chapter 2, Page 106 of 119 Revision 22, 2010

IP2 FSAR UPDATE resulted in extensive thrust faulting, folding, metamorphism and intrusion in the northern Appalachians. The Taconic orogeny, generally interpreted as a continent-island arc collision was very intense in the Manhattan Prong region, and produced most of the structure evident in the current map pattern. The Acadian orogeny (Devovian), possibly a continent-continent collision, was the next pulse of orogenic activity. The Acadian orogeny caused considerable deformation, metamorphism and intrusion in New England, but was not as intense as the Taconic orogeny in the Manhattan Prong. The rocks of the Hudson Highlands (see Figure 1), an extension of the Reading Prong in New York State, consist of Precambrian gneisses and granites of Grenville (1.1 b.y.) age. The Manhattan Prong is underlain by Precambrian basement. An unconformity separates Cambro-Ordovician aged metasedimentary rocks from Precambrian rocks. The Newark Basin is filled with Triassic (Jurassic?) arkosic sediments diabase intrusives and basaltic flows. Geology of the Hudson Highlands The Hudson Highlands outcrop in a northeast (040°) trending belt, approximately 10-miles wide, north, northwest and west of the Indian Point site (Figure 1). Four major rock types are present in the vicinity of Dunderburg Mountain, across the Hudson River from Indian Point. They are quartzo-feldspathic +/- calc-silicate hornblende gneiss; migmatitic quartzo-feldspathic biotite +/- gamet gneiss; calc-silicate bearing quartzite; and gneissic hornblende granite. Granite probably intruded the gneisses during Precambrian time. Heleneck and Mose (1978) mapping in Highlands rocks near Lake Carmel, New York, recognize a mappable sequence of five rock units consisting of, gray migmatitic quartzo-feldspathic gneiss; amphibolite hornblende gneiss; leuco-granite gneiss and amphibolite; layered quartzo-feldspathic gneiss; and interlayered feldspathic quartzite and amphiboite. Highland rocks represent a sequence of Precambrian aged mio- and eugeosynclinal deposits, that have undergone a complex sequence of metamorphism and deformation. The rocks typically yield Rb/Sr ages of 1.1 billion years, the time of Grenville regional metamorphism. Mineralogic and textural evidence indicates that the rocks were metamorphosed to granulite facies, and multiply deformed during the Grenville orogeny. Recrystallization from granulite toemphibolite facies in Highland rocks near Lake Carmel accompanied folding during the Taconic orogeny (mid-Ordovician). Evidence of Taconic recrystallization in other areas of the Highlands remains equivocal. The Highlands are separated from the rocks of the Manhattan Prong and the Newark Basin by a complex fault system known as the Ramapo Fault Zone. Geology of the Manhattan Prong The Manhattan Prong is a sequence of highly deformed metamorphic rocks, trending north-northeast, from New York City through Westchester County and western Fairfield County, Connecticut. The prong is bounded on the east by Camerons Line, a complicated structure possibly representing a suture between two crustal blocks. On the west, the prong is bounded by the Newark Basin border fault and the Hudson River. The stratigraphy of the Manhattan Prong has long been the subject of controversy and frequent revision, however, most recent workers recognize five formations within the prong. In order of decreasing age they are the Fordham Gneiss, the Yonkers-Pound Chapter 2, Page 107 of 119 Revision 22, 2010

IP2 FSAR UPDATE Ridge Granite, the Lowerre Quartzite, the Inwood Marble and the Manhattan Schist (see Table 2). The Fordham and Yonkers formations are Precambrian in age, and are separated from the Cambro-Ordovician aged Lowerre, Inwood and Manhattan formations by an angular unconformity. The relative ages of the Fordham and Yonkers formations are not known with absolute certainty. The Fordham is generally considered at least Grenville in age, and the Yonkers latest Precambrian. Halls (1968) detailed subdivision of the rocks of the Manhattan Prong near White Plains, New York, served as a basis for more recent workers. Correlations of the White Plains stratigraphy to other parts of the prong are tenuous due to the complex structural and metamorphic history of the area. Difficulties in correlation are compounded by the possibility of changes in original sedimentary facies. The Fordham was divided into five members by Hall. They are: Fordham A - Brown weathering garnet biotite quartzo-feldspathic gneiss. Fordham B - Gray garnet biotite quartz feldspar gneiss interlayered with amphibolite. Fordham C - Gray biotite hornblende quartzo-feldspathic gneiss with some amphibolite. Fordham D - Rusty weathering sillimanite-garnet biotite quartzo-feldspathic gneiss. Fordham E - Silliceous biotite-quartz plagioclase gneiss. Most of the Fordham rocks probably represent metamorphosed eugeosynclinal deposits, interbedded with mafic volcanics. The Fordham formation was deformed and metamorphosed to granulite facies during the Grenville orogeny. Mineralogic and structural evidence indicates that the Fordham was recrystallized and deformed during Plaeozoic orogenesis. The Yonkers Granite is thought to represent a metamorphosed rhyolite, emplaced during the opening of the Proto-Atlantic in late Precambrian time. The assignment of formation status to the Lowerre, has been the subject of debate for nearly 100 years. It was first named and described by Merril in 1896, but many workers in the 20th century preferred to consider the Lowerre as part of the Fordham. In addition, the unconformity between the Cambro-Ordovician rocks and the Fordham, has not been recognized by all workers in the prong. The Lowerre is a relatively thin (40-ft. thickness), discontinuous unit representing an arkosic sandstone. The discontinuity of the unit is probably the result of deposition on an irregular Precambrian aged erosional surface. The Lowerre consists of quartz, with considerable amounts of potassium feldspar and minor biotite. The Lowerre is always observed in the same stratigraphic position, at the base of the Cambro-Ordovician aged cover rocks, overlying and truncating various members of the Fordham. The Inwood Marble, consisting of dolomite and calcite marble interlayered with calc-silicate schists, overlies the Lowerre. Hall has divided the Inwood into five members, that lens in and out. In map pattern, the Inwood does not appear to be continuous, the result of tectonic thinning of fairly ductile marble during deformation. The Inwood Chapter 2, Page 108 of 119 Revision 22, 2010

IP2 FSAR UPDATE represents deposition on a carbonate bank, widespread in the Appalachian orogen during the late Cambrian and early Ordovician. In the White Plains area, Hall recognizes three mappable members of the Manhattan, A, B, and C (see Table 2). Manhattan A (basal Manhattan) is a fissile sillimanite garnet biotite schist, interlayered with marble and calc-silicate schist. The Manhattan A may be a transitional facies between carbonate and clastic sedimentation. It is recognized by Ratcliffe (1976) near the Cortlandt complex, and by Brock (1977) north of White Plains. Manhattan B is a discontinuous amphibolite and Manhattan C is a brown weathering, garnet muscovite biotite schitose gneiss. The Manhattan Formation was originally deposited in a miogeosyncline, and represents pelites, mafic volcanics and greywackes. The Manhattan Prong was metamorphosed, deformed and intruded during two major orogenic episodes, the Taconic (late Ordovician) and the Acadian (Devonian). A late Acadian metamorphic and deformational event can be recognized in some locations within the prong (Brock and Mose 1979). Mose and Hall (1979) infer a mid-Ordovician unconformity within the New York City group based on structural and isotopic evidence. Brock (1977) has worked out a detailed sequence of events for the Manhattan Prong near Croton Falls, New York (see Table 3). The rocks of the prong were metamorphosed to K-feldspar sillimanite grade (upper amphibolite facies) at the peak of the Taconic orogeny. At this time the rocks underwent intense deformation, reflecting the effects of four distinct fold events during the orogeny. Taconic aged recrystallization affected the Precambrian rocks of the Manhattan Prong, destroying most Grenville aged metamorphic and structural features. Granulite facies mineralogy and textures survive as relicts within the Fordham, but are not present in the Cambro-Ordovician rocks, supporting the inferred unconformity between Fordham and younger rocks. During Silurian time, deformation eased, but the prong was intruded by the Croton Falls and Cortlandt mafic complexes. The Acadian orogeny (Devonian) produced another set of folds (F5), metamorphism of kyanite-staurolite (mid-amphibolite facies) grade, and the intrusion of the Peekskill Granite. The final Paleozoic metamorphic and deformational event occurred late in the Acadian Orogeny or during the Mississippian, causing local retrograde metamorphism (muscovite grade) and folding (F6) (Brock, 1977). The final orogenic event is seen locally as tight isoclinal folds. Late metamorphism is evidenced by recrystallization on undeformed joint faces. The ductile deformation occurred coevally with brittle deformation along the border fault and within the Manhattan Prong. The relationship of Precambrian, Paleozoic and post Paleozoic aged faulting will be discussed in a following section. One of the significant problems involving rocks of the Manhattan Prong is correlation on a regional scale (see Table 2). The Fordham Formation is often correlated with the Precambrian Highlands gneisses. Some workers have tried tracing the Highlands across the boundary fault, comparing structure and metamorphic details with Fordham rocks. Correlation is tenuous since the Grenville age yielded by the Highlands and Fordham rocks, is a metamorphic age, not a time of deposition. Correlation of the rest of the New York City Group with rocks of the surrounding region is based on similarities in lithology, structural position, radiometric age determinations and fossil evidence. The Lowerre is considered the metamorphosed equivalent of the Poughquag Sandstone. The Inwood is correlated with the Wappinger Limestone Group, and the Manhattan with the Annsville Phyllite and Hudson River Shale. Metamorphism increases from chlorite grade near the Hudson River, to K-feldspar-sillimanite grade near the Connecticut border. Chapter 2, Page 109 of 119 Revision 22, 2010

IP2 FSAR UPDATE The Geology of the Newark Basin The third geologic province in the area is the Newark-Gettysburg Basin. The basin extends 140-miles from York County, Pennsylvania, to Rockland County, New York (Figure 1). The basin, a down dropped crustal block, formed during Mesozoic time. Deposition was continuous from the late triassic through the upper Jurassic (Dames and Moore, 1977). Intrusion of the Palisades sill apparently occurred during deposition of sediments in latest Triassic-earliest Jurassic. The extrusion of the Watchung basalt flows followed later in the Jurassic. Rocks of the Newark series are in contact with the crystalline rocks of both the Manhattan and Reading Prongs, but the nature of the contact varies. At the northeastern edge of the basin, Triassic sediments unconformably lie over the Highland rocks, while the northeastern edge of the basin is in fault with the rocks of the Highlands. The Newark Group is divided into four formations, the Hammer Creek Conglomerate, the Stockton Arkose, the Lockatong Argillite, and the Brunswick Shale and Sandstone. Deposits of conglomerate and sedimentary breccia lie at the edges of the basin, reflecting proximity to the uplifted Precambrian and Paleozoic rocks that are the sources of the Triassic aged sediments. The boundary fault between the Newark series and older crystalline rocks is the Ramapo fault. Movement along the fault and subsidence of the basin, concurrent with sedimentation, produced the half-graben configuration of the basin. The rocks within the basin are not greatly deformed, displaying broad open folds of uncertain origin, gentle dips of strata, and minor faults with small offsets. History of Brittle Deformation A series of north-northeast trending faults pass through the area surrounding the Indian Point sites. The faults, some of which have been episodically active since the Precambrian time are collectively known as the Ramapo fault system. The system is composed of a number of parallel to sub-parallel branches and draws its name from the Ramapo fault, the boundary between the Reading Prong and the Newark Basin. Ratcliffe (1976) mapped the faults in the vicinity of the Indian Point site, and interpreted a chronologic sequence of fault movements. Ratcliffe classified faults utilizing radiometric ages, cross-cutting lithologic relationships, and textural evidence. Dames and Moore (1977) utilize Ratcliffes conclusions, and present evidence for timing fault movements based on geothermometry of fluid inclusions in calcite. More recently, Nelson (1980) and others have examined the stratigraphy and pollen remains in swamps and sag ponds along the Ramapo, seeking evidence of Post-Pleistocene faulting in the area. In addition, Aggarwal and Sykes (1978), Yang and Aggarwal (1981) Dames and Moore (1977), and Woodward-Cylde Consultants (Quarterly Reports, Jan 1977 - Jan 1982) have studied the recent seismicity in the region, solving for magnitude and location of earthquake epicenters. The earliest documented movement along the Ramapo fault system is Grenville age (Ratcliffe 1976). Textural evidence seen in the Canopus Pluton indicates that movement along the Canopus Hollow fault (north of Peekskill) was synchronous with emplacement of the pluton. Flow structures and mylonitization displayed within the pluton indicate crystallization during shearing. Drag folds and the overall shape of the pluton suggest right lateral strike-slip motion. The Canopus Pluton, a diorite-monzontie, Chapter 2, Page 110 of 119 Revision 22, 2010

IP2 FSAR UPDATE has been radiometrically dated at 1150 m.y. (Rb/Sr), thus providing a minimum age for movement along the fault. The Lake Peekskill fault is defined by a shear zone in the Precambrian gneisses that has not affected the Annsville Phyllite (Ordovician). The fact that the Annsville has not been deformed, places a limit on the last motion along the fault. Ratcliffe (1976) states that both the Canopus Hollow and Lake Peekskill faults were not reactivated during Paleozoic time, however, in a more recent article (Ratcliffe 1980) he suggests that movement as recent as Triassic has taken place within the Canopus Fault Zone. Dames and Moore (1977) disagree with Ratcliffes 1976 opinion, citing as evidence a sheared inlier of Poughquag Quartzite near Canopus Lake. The shear zone is considered by Dames and Moore to reflect thrust faulting along a northeast trending fault, sub-parallel to the Canopus Hollow fault. This implies post-Precambrian reactivation along the Canopus Hollow Fault. An additional strike-slip shear zone was mapped by Dames and Moore in Highland rocks, near the southwest corner of Canopus Lake, along the strike of the Canopus Hollow fault. This shear is considered additional evidence for reactivation of the Canopus Hollow fault in Paleozoic time. It is important to note that local activity occurring along part of a fault does not require movement along the entire length of a fault. Furthermore, the shear zone that displaces the Poughquag is not necessarily an extension of the Canopus shear zone, and may in fact be related to Paleozoic aged folding. Thus, Paleozoic aged movement along the Canopus Hollow fault is not required in the vicinity of Indian Point. A number of faults of Paleozoic age separate the Manhattan Prong from the Hudson Highlands. Most prominent are the Thiells fault, the Annsville fault, the Peekskill fault and the Croton Falls fault. The Peekskill and Croton Falls faults outcrop on the east bank of the Hudson River, and generally trend east-west. The Thiells fault, outcropping on the west bank of the Hudson, and the Annsville fault on the east bank trend northeast. The Ramapo fault extends northeast from Peapack, New Jersey separating the Newark Basin from the Reading Prong. At Ladentown, it splays into two branches, trending 020° and 060°, respectively. The 060° branch connects with the Thiells fault, and the more northerly trending branch extends into the Highlands, through Tomkins Cove, New York. The Peekskill, Croton Falls, Thiells and Annsville faults are primarily Paleozoic in age. The faults are marked by mylonite and ultra-mylonite, displaying retrograde chlorite grade, green schist facies mineral assemblages. Movement along the faults is generally right-lateral-strike-slip. Mid-Ordovician minimum ages of movement are inferred from cross-cutting relationships with dikes related to the Rosetown Pluton (mid-Ordovician), and radiometric ages of undeformed biotites from within shear zones (Ratcliffe, 1976). A lower Devonian K-Ar age of 396 m.y. places the most recent probable movement along the Roa Hook branch of the fault squarely within the Acadian orogeny. Similar data is available for the Thiells fault. Younger faults can be distinguished from Paleozoic aged and older faults (Ratcliffe, 1980) by their different mineralization, and cataclastic textures. Younger faults are characterized by open work breccias, clay gouge, platy fracture and deeply incised fault scarps. Older faults display healed breccia, semi-ductile mylonite shear zones, and higher temperature minerals that reflect the general pattern of Paleozic aged regional metamorphism. Reactivation and development of new faults occurred during Mesozoic time (Ratcliffe, 1980). Deep seated zones of weakness in crystalline basement were utilized in the Chapter 2, Page 111 of 119 Revision 22, 2010

IP2 FSAR UPDATE development of the Triassic basin, particularly the Ramapo-Cheesecote and Mott Farm Road faults (Ratcliffe, 1976). Structural evidence indicates that normal faulting was dominant during mesozoic time, with the latest activity along the Mott Farm Road branch of the Ramapo, dated at 163 m.y. by K-Ar methods (Ratcliffe, 1976). Late north-south vertical strike-slip faults with both left-lateral and right-lateral movements are present in the vicinity of Indian Point. The relationship of these faults to the Ramapo is uncertain, although the faults are probably Mesozoic in age (Ratcliffe, 1976; Dames & Moore, 1977). Detailed work (Ratcliffe, 1976) shows that north-south faults at Tomkins Cove across the river from Indian Point are the youngest in the area. Mineralogy and textures found in the young, Tomkins Cove faults bear a strong resemblance to faults that have been radiometrically dated as Mesozoic. In addition, the time of last movement along the faults is constrained by the lack of fault related deformation in overlying Pleistocene sediments (Ratcliffe, 1976, 1980; Dames and Moore, 1977). A group of faults located at the Indian Point site was mapped in detail by Dames and Moore (1977). Displacement along the faults is not significant, no more than a few feet. The faults are filled with undeformed euhedral calcite crystals, many of which contain fluid inclusions. Temperature equilibrium studies on the fluid inclusions indicate average formation temperatures of 160°C. Dames and Moore (1978), infer a depth of formation, by applying the geologically conservative geothermal gradient of 50°C/km. This yields a temperature of 150°C at 3 km. The amount of time necessary to expose rocks that form at a depth of 3 km is a function of denudation rates. A minimum of 45 m.y. is required to remove 3 km of material, if the rather rapid denudation rate of 15,000 yrs/meter is applied (Dames and Moore, 1978). This calculation sets another constraint on the possible minimum age of last movement on the faults. The growth of calcite in the fault zones has been attributed to circulating hydrothermal fluids related to Mesozoic igneous activity (Dames and Moore, 1978), suggesting a time of last movement in the Mesozoic. Radiometric age determination of undeformed minerals that have grown within fault zones (Ratcliffe, 1976) and the lack of fault related deformation of Pleistocene deposits and surface features (Dames and Moore, 1977); Ratcliffe 1976, 1980), provide the best evidence that the faults in the Indian Point area have not moved in the last two million years. Data from recent drill cores that intercepted the Ramapo fault plane, show that the dip of the structure is highly variable. The cores, taken at four locations along the fault, indicate that the dip is consistently to the southeast, ranging from 45° through 70° (Ratcliffe, 1980). Textural evidence observed in the cores, indicate that the dominant latest motion in the fault has been right oblique normal faulting (Ratcliffe, 1980). Recent Seismicity In the last twenty years, the catalogue of instrumentally recorded seismic events in the northeast has grown tremendously. Locally, this is the result of a dense network of seismic stations, situated in the area around Indian Point, that has been operated since 1975. Data collection by regional seismic stations has continued in the same area since 1970 with a reported detection threshold of about magnitude 2 mb (Lg). Seismic networks have provided a basis for accurately determining the location, magnitude, and in some case focal mechanism solution for many small magnitude seismic events in the area of the power plant and the Ramapo fault zone. This recent seismicity is not markedly dissimilar from the historical seismicity reported for the region. The composite data set does not define or suggest a structural association of earthquakes, however, a Chapter 2, Page 112 of 119 Revision 22, 2010

IP2 FSAR UPDATE regional overview does suggest a higher level of activity in the northern New Jersey, southeastern New York area than that of the surrounding areas. A number of hypotheses have been proposed to explain the observed pattern of seismicity. Current seismicity in the northeastern United States has been attributed to: a proposed stress system in which the maximum compressive stress trends to the northwest (Yang and Aggarwal, 1981); proximity to relatively young igneous bodies (McKeown, 1978); the response of the crust to glacial unloading (Stein and others, 1979); and, in the immediate region about the site, proposed reactivation of pre-existing fault zones based on a spatial correlation with surface traces of faults exposed within the region (Aggarwal and Sykes, 1978). Earthquakes occurring near Indian Point have been characterized as shallow focus (<10 km) and low magnitude (1.0-3.0) (Aggarwal and Sykes, 1978). Focal mechanism solutions reported for earthquakes near the Ramapo fault or the margins of the Triassic basin indicate thrust movement on faults that parallel the dominant structural grain in the exposed bedrock (Figure 2) (Aggarwal and Sykes, 1978; Yang and Aggarwal, 1981). The stress field required for this interpretation, must be compressional and oriented to the northwest (Yang and Aggarwal, 1981). Whereas some of the seismic activity in southern New York may be related to northeast trending structures, a number of other trends, transversely oriented with respect to the dominant structures are present (Ratcliffe, 1976; Pomeroy et al, 1976; Blackford and Statton, 1978; Quarterly Report for the Indian Point Seismic Monitoring Network, November 1979 through January 1980) (see Figure 2). Thompson and Bebel (1979) describe northeast and northwest trends of epicenters in the coastal plain area of New Jersey, Delaware and Pennsylvania. Many studies have attempted to quantify the stress regimes operating in the vicinity of the Ramapo fault. Dames and Moore (1977) determined the near surface stress by in-site measurements. While these results are variable, a fairly consistent northeast to eastwest trend for the horizontal component of compression was determined. This stress direction is transverse to that suggested by others (Aggarwal & Sykes, 1978) who base their interpretations on reported focal mechanism solutions. It is, however, in general agreement with regional stresses inferred from measurements and observations made throughout the northeastern United States (Sbar and Sykes, 1973). An examination of the distribution of earthquakes in the vicinity of Indian Point indicates that not all earthquakes in the region can be attributed to northeast trending faults. A sequence of earthquakes occurred near Annsville, New York, from 17 January to 23 January 1980. A composite focal mechanism solution was constructed using data recorded by the Indian Point Seismic Monitoring Network. The solution indicates thrust faulting along one of two possible planes, oriented N2°W 29°E or N16°W 62°W (Quarterly Report for the Indian Point Seismic Monitoring Network, November 1979 - January 1980). This trend is obliquely oriented to the dominant structural fabric in the region, and requires a compressive stress field oriented east-west to northeast-southwest. Low level microseismicity existing in the region is evidence that crustal adjustments are continuing in response to regional stresses. No evidence exists at the surface or in drill cores of the fault zones (in particular the Ramapo fault zone), that suggest any contemporary movement along faults exposed at the surface since the major period of activity during Mesozoic time. In fact geologic data obtained from cores of the Ramapo fault zone (Ratcliffe, 1980) show evidence of normal faulting as the last movement, which is consistent with the Mesozoic faulting regime and inconsistent with the thrust mechanism proposed by Aggarwal and Sykes (1978). To date, no satisfactory stress Chapter 2, Page 113 of 119 Revision 22, 2010

IP2 FSAR UPDATE regime has been proposed that adequately accommodates the observed pattern of low level seismicity, regional stress measurement data and those stresses inferred from recently reported focal mechanism solutions. CONCLUSION Low level microseismicity in the region is evidence that crustal adjustments are continuing in response to regional stresses. Detailed field investigations (e.g., Ratcliffe, 1976, 1980; Dames and Moore, 1977) have been conducted in the immediate vicinity of Indian Point, and along the major faults in the region. To date, no evidence has been found in the rocks exposed at the surface or sediments overlying fault traces or in cores obtained in the vicinity of Indian Point, that might support a conclusion that displacement has occurred along major fault systems within the New York Highlands, the Ramapo or its associated branches during Quaternary time (the last 1.5 m.y.). In the vicinity of Indian Point, evidence that no displacement has occurred in the last 65 m.y. (since the Mesozoic) along specific major structures has been observed. Chapter 2, Page 114 of 119 Revision 22, 2010

IP2 FSAR UPDATE Aggarwal, Y.P., and Sykes, L.R., 1978, Earthquakes, faults and nuclear power plants in southern New York and northern New Jersey: Science, v. 200, p. 425-429. American Geological Institute, 1977, Glossary of Geology, Gary, M., McAfee, R. Jr., and Wolf, C.L. eds., Am. Geol. Inst., Falls Church, Va., 805 p. Blackford, M., and Statton, C. T., 1978, The Annsville, New York earthquake of September 2, 1977. Presented at the seventy-third annual meeting of the Seismological Society of America. Brock, P.W.G., Structural and age relationships within the New York City Group in southeastern New York, Unpublished Grant Proposal. Brock, P.W.G., and Mose, D.G., 1979, Taconic and younger deformation and metamorphism in the Croton Falls area, southeastern New York: Bull. Geol. Soc. Am., Pt. II, v. 90, p. 1158-1195. Dames & Moore, 1977, Geotechnical investigation of the Ramapo Fault System in the region of the Indian Point Generating Station. Hall, L.M., 1968, Times of origin and deformation of bedrock in the Manhattan Prong, in: Zen, White, Hadley and Thompson, eds., Studies of Appalachian Geology, Northern and Maritime, Interscience Publishers, New York, p. 117-128. Heleneck, H.L. and Mose, D.G., 1978, Geology and geochronology of Precambrian rocks in the Lake Carmel region, Hudson Highlands, New York: Geol. Soc. Am., Abstracts with Programs, v. 10, no. 2, p. 47. Heleneck, H.L. and Mose D.G., 1976, Structure, petrology and geochronology of the Precambrian rocks in the Central Hudson Highlands; in: Johnson, J.H., ed., Guidebook to Field Excursions, 48th Annual Meeting of the New York State Geological Association: New York, Vassar College, p. B1 1-27. Mc Keown, F.H., 1978, Hypothesis: Many earthquakes in the central and southeastern United States are casually related to mafic intrusive bodies: Jour. Res. U.S. Geol. Surv., v. 6, p. 41-50. Merrill, F.J., 1896, The geology of the crystalline rocks of southeastern New York: New York State Museum. Ann. Rpt. No. 50, Appendix A., p. 21-31 Mose, D.G. and Hall, L.M., 1979, Rb-Sr whole-rock age determination of Member C of the Manhattan Schist and its bearing on allochthony in the Manhattan Prong, southeastern New York: Geol. Soc. Am., Abstracts with Programs, V. 11, no. 1,

p. 46.

Nelson, S., 1980, Determination of Holocene fault movement along the Ramapo Fault in southeastern New York using pollen stratigraphy: Geol. Soc. Am., Abstracts with Programs, v. 12, no. 2, p. 75. Pomeroy, P.W., Simpson, D.W., and Sbar, M.L., 1976, Earthquakes triggered by surface quarrying-the Wappingers Falls, New York sequence of June, 1974: Bull. Seis. Soc. Am. V. 66, p. 685-700. Chapter 2, Page 115 of 119 Revision 22, 2010

IP2 FSAR UPDATE Ratcliffe, N.M., 1976, Final Report on major fault systems in the vicinity of Tomkins Cove-Buchanan, New York: Report for Consolidated Edison Company, Inc. of New York, 157 p. Ratcliffe, N.M., 1980, Brittle faults (Ramapo Fault) and phyllonitic ductile shear zones in the basement rocks of the Ramapo seismic zone, New York and New Jersey, and their relationship to current seismicity, in: Manspeizer, W., ed., Field Studies of New Jersey Geology and Guide to Field Trips, 52nd Annual Meeting of the New York State Geological Association, Newark, N.J., Rutgers University, p. 278-312. Robinson, P., and Hall, L.M., 1979, Tectonic synthesis of southern New England in: The Caledonides in the USA, I.G.C.P., 1979. Blacksburg, Va. Sbar, M.L., and Sykes, L.R., 1973, Contemporary compressive stress and seismicity in eastern North America: An example of intra-plate tectonics: Bull. Geol. Soc. Am., v. 84, p. 1861-1882. Stein, S., Sleep, S., Geller, R.J., Wang, S-C, and Krueger, G.C., 1979, Earthquakes along the passive margin of eastern Canada: Geophys. Res. Letts. v.6, p. 537-540. Thompson, A.M., and Bebel, D.J., 1980, Modern seismicity in the middle Atlantic Seaboard Region, and some neotectonic Implications: Geol. Soc. Am., Abstracts with Programs, v. 12, no. 2, p. 55. Woodward-Clyde Consultants, Quarterly Reports for the Indian Point Seismic Monitoring Network, Jan., 1977-Jan., 1982. Yang, J-P, and Aggarwal, Y.P., 1981, Seismotectonics of northeastern United States and adjacent Canada. Jour. Geophys. Res., v. 86, p. 4981-4998. In addition, the following reports were used as background information: Fluhr, T.W., 1965, Memorandum on geologic features of Indian Point Nuclear Power Plant site: for Consolidated Edison Company of New York, Inc., 9 p. Paige, S., 1955, A geologic report on the Indian Point Power House Site of the Consolidated Edison Company, Inc., New York. Chapter 2, Page 116 of 119 Revision 22, 2010

IP2 FSAR UPDATE GLOSSARY amphibole - A complex chain silicate mineral rich in iron and magnesium. amphibolite - A metamorphic rock whose main components are amphibole and plagioclase feldspar amphibolite facies - Rocks formed at moderate temperature and pressure conditions during regional metamorphism. angular unconformity - An unconformity recognizable by the deposition of sediments over deformed rocks. argillite - A compact mudstone, generally not laminated and not fissile. arkose - A feldspar rich, generally coarse grained, sandstone derived from continental rocks. breccia - A generally coarse grained rock composed of angular or broken fragments of rock, which may be formed tectonically in a fault zone, or by sedimentary processes. brittle deformation - A term used to describe faulting and fracturing of rocks. calc-silicate - A descriptive term applied to minerals or rocks consisting of calcium bearing silicates, such as diopside. chlorite - A green iron-magnesium rich platy mineral. chlorite grade metamorphism - Low-grade regional metamorphism indicated by the first appearance of chlorite in rocks of appropriate composition. dike - An igneous intrusion that cuts across planar features of a rock. disconformity - A break in the time-stratigraphic record separating two sequences of rock, both of which are bedded parallel to the unconformity. ductile deformation - Occurs where rocks fold or flow when subjected to a stress field. eugeosyncline - A geosyncline or basin in which vulcanism is associated with clastic sedimentation. euhedral - A crystal bounded by well developed crystal faces. facies - A set of conditions that specify the environment of formation of rocks (metamorphic or sedimentary). geosyncline - A long linear basin, characterized by subsidence coincidental with sedimentation. geothermal gradient - The relationship of temperature to depth (pressure) in the earths crust. Chapter 2, Page 117 of 119 Revision 22, 2010

IP2 FSAR UPDATE gneiss - A metamorphic rock formed by regional metamorphism, generally high grade. granulite facies - Rocks formed at very high temperatures and pressures during regional metamorphism. intrusion - The emplacement of an igneous body in a pre-existing rock. isocinal folds - A fold in which the limbs are parallel. mafic - A term used to describe dark rocks or minerals containing large amounts of magnesium or iron. marble - A metamorphic rock consisting primarily of calcite or dolomite. metamorphic grade - Rocks of any composition that have been metamorphosed under a specific range of temperature and pressure conditions. metasediment - A metamorphosed sedimentary rock. migmatite - A mixed rock composed of metamorphic material containing segregation of igneous material formed by injection or in-situ partial melting. miogeosyncline - A geosyncline lacking volcanic deposits, commonly located adjacent to continental margins. neotectonics - Post-Miocene structural history of the earths crust. orogen - A region that has been subjected to orogeny. orogeny - The development of structures, metamorphism and igneous activity relating to mountain building. pelite - A sedimentary or metamorphic rock rich in aluminum. quartzite - A metamorphic or sedimentary rock consisting mainly of quartz. retrograde metamorphism - Recrystallization of metamorphic rocks at conditions that are lower grade than those at which the rock was originally metamorphosed. schist - A well foliated metamorphic rock that easily separates into flakes or slabs due to an abundance of platy minerals. sedimentary facies - An restricted area within a litho-stratigraphic unit representing a particular depositional environment. sill - An igneous intrusion that parallels the layering of the country rock. tectonic - Pertaining to the forces that cause crustal deformation. throw - The vertical component of fault motion. unconformity - A primary feature representing erosion or non-deposition, resulting in a break in the stratigraphic sequence. Chapter 2, Page 118 of 119 Revision 22, 2010

IP2 FSAR UPDATE TABLE 1 GEOLOGICAL TIME SCALE TABLE 2 PROPOSED CORRELATION OF THE STRATIGRAPHIC SUBDIVISIONS IN THE MANHATTAN PRONG WITH ROCKS IN ADJACENT AREAS TABLE 3 GEOLOGIC HISTORY IN THE CROTON FALLS AREA (after Brock & Mose, 1979) FIGURES Figure No. Title Figure 1 Geological Map, Southeastern New York (after Brock & Mose, 1979) Figure 2 Seismicity of Southeastern New York and New Jersey (after Dames & Moore, 1977) Chapter 2, Page 119 of 119 Revision 22, 2010

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I [, I l' GEOLOGIC HISTORY IN THE CROTON FALLS AREA (AFTER BROCK III MOSE. 19791 (F 6

• tight to isoclinal fold; deformation of Goldens Bridge granite)

(Goldens Bridge granite) Joints with muscovite and tourmaline

        ? open folds

f F5

• folded isograd. 55: retrograde IlI.Iscovite schistosity; associated

   ~
   ~

retrograde kyanite and staurolite; locally associated shearing and faulting. (Croton Falls granites intruded 387 : 37 m.y. ago) (Croton Falls ultramafic complex: includes xenoliths of refolded schists; could b synchronous with F3 folding).

        ? Open folds Quartz: garnet: sillimanite. and quartz plus two feldspar + garnet
              + sillimanite segregations.

F4

• locally developed reclined close folds and chevron folds. Faint S4:

realigned biotite and sillimanite in F3 segregations and host rocks. Sillimanite K-feldspar facies. F

• very common near-isoclinal folds; refolds Sl schistosity and F2 flasers.

3 Variable S3: slight re-alignment of sillimanite and biotite. Quartz-K-feldspar - sillimanite segregations. Sillimanite-K-feldspar facies. F

• isoclinal to near-isoclinal intrafolial folds. Strong cataclastic S2:

2 flasers of quartz-feldspar segregations and plagioclase. smeared garnets. partial realignment of biotite and sillimanite. Quartz plagioclase segregations. Fl a a few isoclinal folds. Strong Sl: alignment of biotite. sillimanite. and quartz! plagioclase segregations. Porphyroblastic garnets! plagioclase. Manhattan SChist deposition Inwood Marble deposition Lowerre Quartzite deposition unconf~nnity (Intrusion of Yonkers and Pound Ridge granite gneisses at 575-600 m.y.)

   ~    Fordham Gneiss (Hudson Highlands gneiss and related rocks metamorphosed
   ~

Lo.l a:: at 0.9-1.2 b.y. and still contain granulite facies relicts.) (:l Note: All features or events except those in parenthesis are demonstrable in one roadcut about 1 km south of the Croton Falls Complex on Route 684. TABLE 3

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                                                                                                     -~
                                                                  -~ov   1 , 79,00              APH 30,19,dU l-'lINT'ER
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                                                                   .~A ~    1,      7~,Il\l
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---

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 - - - - - - - - - - - - - G E O L O G I C TIME S C A L E - - - - - - - - - - - - - - - .

/f ABSOLUTE AGE ERA 'ERIOD EPOCH SIGNIFICANT EVENTS (I" mlllio". of yeen Befo,. P,_nt) HOLOCENE

            -         0.01-             .                  QUATERNARY                     IRECENT)

PLEISTOCENE

            -         1.8   -           -    (,)

(5 PLIOCENE WISCONSINAN GLACIATION N NEOGENE

            -        22.5   -           -    0 Z

1&.1 (,)

                                                      )-~
                                                         ~~
                                                           ~         .0 MIOCENE OLIGOCENE (J \
                                                               ~~     '1(

(C'O C'~ 1-(C' EOCENE PALEOCENE (,) CRETACEOUS

                                        *    (5 N

0I/) JURASSIC -

             -      195     -                1&.1
                                              ~                TRIASSIC OPENING OF ATLANTIC BREAK UP OF PANGEA
             -      230     -                                  PERMIAN
             -      280     "-                      r: *
                                                   .! 6

PENNSYLVANIAN ALLEGHENIAN OROGENY

             -      345     -

(,) (5 N

                                                    ~
                                                    ~~  -  I    MISSISSIPPIAN 0              DEVONIAN
             -      395     -           * ... 1&.1 c(

SILURIAN ACADIAN OROGENY

             -      435     -           *
                                              ~

ORDOVICIAN TACONIC OROGENY

             -      &02     -                                CAMBRIAN
             -      572     -
                                    .                      HADRYNIAN OPENING OF IAPETUS
             -     1000     -         Z c&:

1&.1(,) GRENVILLIAN OROGENY

             -     1800     -       . ~ i2 It   ~O HELIKIAN
             -     2600"    -         III
                                      ~

c( (,) 1&.1 APHEBIAN It "-liC,.,

                                                              ~"'
                                      ~
             -     4500

(,) (,) (,) 5 N 0 N0 0 N 1&.1 g 0 Z c( 1&.1 1&.1

                                                                                                    ~         ~      (,)  "

2600 PROTEROZOIC 572 230 650 I I I I I BAR SCALE (In million. of yUn Before p,...nt) (After: F.W.B. ven Evsinga. 19781 TABLE 1

    ~LOCATION               IIAIIIIATTAN                 PIlONG OUTCHESS      COUNTY,   WUTEIltl MASSACHUSETTS WESTERN           YERMONT AlE~

NEW YOfIK AND CONNECTICUT FlSHER,C1162a, 19UW ZEN AND HARTSHOllII, (1t661 DOLL AND OTHERS, (19'1l THiS IIEPORT KNOPF,11K21 ZE"lIMTI ZEN, Itl")

                      ~ ~       t-......:C;..*..::;II:=.E=II.::;IE::,:R.:.---i
                      == !!                I

• liE II lEIl MOOLE ORDOVICIAN ; ~ t---:A:-'~II::E"'II::I:;::E-=R--i .ALLOOIISAC FOlNATION WALLOOIISAC FOlIIIATION lIlA 'ORIIATION r--------j..../-~--r-'~. ~_ ~

                                     /         il~li~:Df:C r-.,DlIfR                           ~

L ~~;~~ COPAKE LIMESTONE

                                                                                                                .L~i~i;~:O UNlT-e
                                                                                                                                                       --J ~

CH.IIAN

i:~ IIARILE FORIlATION E""LV OfIDOYIClA" t_-----~ iii IIOCHDALE LIIIESTONE IASCOII 'ORMATION

                                                                               ..                          5                  UNIT-'                          CUTTING DOLOMITE III D-IIEIlIER                          ~       HA!.CYON LAKE       ;: 1-----;;;,UN:;,I;:,T--:E:----I---.-H..;:E;.:;L.:.IUR.:.:;;cNE"-";.:;'OR=II~A.;.;T=-IO-N--i
                        ..a'"'

II z ~ UNIT-D 1 - - - - - - - - - 1 :: I - - - - - - - - - - i ~ I----------+---C-L-A-IIE-N'-D"'""O-N-S-PR-I-N-G-'--i C-IIE_E11  : IllIlAIICLF' DOLOIIITE UlIIT-t DOLOIlITE

                                                                               *                           ...                                                DANIY 'ORIIATlON t------__i~.. P . PLA""S FQllMATlON :~t_-------+-----------__i                                                              WINOOSKI DOLOMITE
                                                                               ...                         ..                 UNIT-I g                           0                                               IIONkTON QuARTZITE CAIlIRIAN 1---------1                                    I      STlISING DOLOMITE    :; t--------i--=::..:.=..::=.:.:~--l
                                                                               ~

A-IIEIIIER UNIT-A DUIIIIAIol DOLOMITE CHtSHIRE ClUMTZIT£ CHESHIRE QUARTZITE LOWERRE QUARTZITE I'OUGHQUAG QUARTZITE DALTON FORIIATION .rDA!.TON FORIlATION t--------l--~---../- . . . . .,/".. .~_-.. . . ._. ..;'- -- ~~

                                                                                                                                                -~-                                 ---~-

PRECAli IRIAN FOIIDHAII INEISS YONKERS GNEISS INEISS IlOUNT HOLLY COMPLEX I I PROPOSED CORRELAnON OF THE STRATIGRAPHIC SUBDIViSiONS IN THE MANHATTAN PRONG WITH ROCKS IN ADJACENT AREAS (AFTER HALL. 1968) TABLE 2

(~.~. ,-,;'- _u"

• II a ..
• II' o* "

0-

                                                                                                                          \
                                                                                                                            \
                                                                                                                              \
                                                                                                                    * 'us

{//-':":

  ~

SEISMICITY OF SOUTHEASTERN NEW YORK AND NORTHERN NEW JERSEY

                                                             !AFTER DAMES 8< MOORE. 11771 I'

I

                                                                           ..l" t _   11' fI* * "

I~_. i I _ _ 0( _ , ..-'., _1' _ _ _ FIGURE 2

                                  \\        ,YANITOU "TN
                              ~,.7"O' PEAK                 ..
                              . ([~~              ~
                              ~~                        ".
                             ~-~,
                                     ~  ....- - .
                                             ~

DUNDERBERG MTN 108S*PEAK

                . BALD MTN 1120- PEAK*

,~

        ~,                                                                   BLUE MTN 830-' PEAK
\,,~

BUCKBERG MTN 793* PEAK \ L~/7 ,SPITZEN&ERG

                                                                                 ./ MTN 580' PEAK
                                       ~~TONY
                                           . POINT I'
                         ..~

PRICKL Y PEAR GRASSY POINT HILL 538' PEAK

                                                    .BOWLINE
     ,r  --

POINT HAVERSTRAW BAY

                                                                                                       /

SOUTH MTN 820* PEAK f"-...- I ~ o.q.... HUDSON RIVER

                                                                                ~ CJ~O A                          '::~~                 SHORE I I l

I U-\'~.,.~ .j LINE * , \,

                      )                                        ~~~+~.FIGURE 1                        \

f i

                    /                                         '.             AT EL. 200 FT.

{ (. ~ ~ S<rAL£: 1*-57~0*

              ;     \

BOWLINE (j :I OOM ToweR BOWLINE PT Od" -,,~, HAVERSTRAW

" .                                      BAY SOUTH. ~TN
                                                             ~o""
                                                                  . o*  ~....

HUDSON...._~ ( CI~ ';

                        ~

RIVER SHORE O~ "

                             .,,~ '<, ",

LINE'

                                +" ~ "'~. .\;; .         FIGURE, 2              \
                                    '->""A~~-f \)"~G"RO, UNO     CONTOURS:
                                              ~      ,AT EL. 400 FT.,                ,
                                                   ~'\ SCAl.t: \'" = 5700 FT. I
                                                       \   '   I              I I

FIGURE 3 ELEVATIONS IN THE INDIAN POINT REGION BEAR MTN 1284 X IONA ISLAND WEST MTN 1257 X BLUE MTN 830 X THE TIMP 8.8 X INDIAN POINT DICKERSON MTN 730 X BUCKBERG MTN 793 X SPITZENBERG MTN 580 X X PINGVP MTN 1023 SALT HILL 750 .X IRISH MTN 1174 X WATCH HILL. X T COLLABERG MTN 600 X BOWLINE POINT o CHEESCOTE MTN 960 X MOTTS HILL 580

               !     X              X LITTLe TOR MTN I

X SOUTH MTN 802 X "' HIGH TOR MTN 820 A.~

     ~
    'lit q,O                                                            .oJ W        OSSINING
 ~                                                              W
~                                                               .oJ C

l.U oI a:: W a: z oo

                                                               §

z:

KINGSLAND PARK X EASTVIEW SOUTH NYACK MT NEBO 880 WES'TCHESTER COUNTY AIRPORT X X THE NUMBER LOCATED TO THE NORTH RIGHT OF THE MTN OR HILL X CENOTES MorTORING SITE INCICATES ELEVATION ABOVE MSL IN FEET TALLMAN MTN 180 X

FIGURE 4 WATER COURSES IN THE INDIAN POINT REGION KINGSLAND PARK

    ~

cr X EASTVIEW ii

  ~

(.) WESTCHESTER COUNTY AIRPORT

 ~ NORTH          (.)

x It 5 ,",OITORING(TE

FIGURE 5 EXISTING AND HISTORICAL HETEOROLOGICAL TOWERS AT INDIAN POINT

=.. IUlUI - - -

                                                                                   ~    -flllID ,",,--
                                                                                    ~~. _.      -~~

LEGEND per~odl Tower Designation Activity IP 2 IP1 1956-1957 IP2 1968-1969

                                              ~.~~\                          IP)                    1969-1976

ir..:::n:,=,;r~.ii& J.22M( IP4 ) 197)-Present u::....__

• EMERG . BACKUP
• 1980-Present

'"""5011 OilL', lIlvlEll 122 M TOWER LOCATION: 41* 15'55"N

           'f..~"'f'l.              '1~'t'l

• Located at E.C.C.

13- 61'oe"W

                   'SCALE: 0         600      1000 Feet         prior to 12/01/~1.

FIGURE 6 \ INDIAN POINT METEOROLOGICAL SITE r--------, 122 METER :

                                                ~:
                                          'TOWER             ,

I ,. ,

                                          ,~         .,.....J N.
                                                                        ~ I          I I

AUXIUARy*1 POWER i

'PRO XIMATE FOUNDAnONI SCALE (fe.t)

I i o 5 10 GENERATOR BUILDING (TEMPORY)

                                                                                    ~.
                                                                                       \
                                                                                         \
                                                                                           \
                                                                                    ~        '.
                                                                                             ,~

f' I o PRECIPITATION GAGE I I I 1.- - - - -- .... ~ _.., __ .-. _ _.-, ~ -_~ I __ ....J FENCING

FIGURE 7 TOWER CONFIGURATION 122 METE~ LEVEL

                                                     >C WIND SPEED           )<

WIND DIRECnON TEMP. DIFFERENCE

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FIGURE 8

                                                         ~S STATION CONFIGURATION 12/01/81
                              / ...

Wq) SENSOR CROSS-'RM

                                                  /

I INSTfiUMENT BOO' ~S INSTRUMENT 8' X22' TR~ILER o . PRECIPIT~TION G~GE NOT DR~WN TO SC~Lt 122 M TOWER LOCATION: 41 9 U5'55 wN 73-57'Oa*W

FIGURE 9 INDIAN POINT-METEOROLOGICAL ~UPPORT SYSTEMS 122 MIJ'2!' TOWllti 122 Meter WS 10 we METlO"OLOCIICAL TlIALD

                                                                                                                                                     . . .CTO" CONTROL ROOMS
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FIGURE lOA TWO STATION WIND CORRELATION OA!'A PERIOD

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FIGURE 11 POSITION OF ONE MILE GRID IN RELA nON TO TOPOGRAPHIC FEATURES 1 I/V. I T I I 08 01 01 02 os 04 05 01 0'1 08 Of 10 11 12 13 1'6 15 16 11 18 19 20 21

FIGURE 12 POSITION OF WIND FILES ON GRID 1: INDIAN POINT 2: BOWLINE POINT PATTERN 1

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FIGURE 13 AVERAGE MARCH. 1980 EAST AND Wf:ST BANK DIUI(NAL WIND DISTRIBUTIONS 160 140 120 100 080 o. 060 O. West Bank 040 I.: C

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FIGURE lS AV~HAi.;E DECEMBER. 1980 ~:AST AND WES r BANK 01 URNAI, WIND DISTRIBUTIONS 160 140 O. 120 O. "'t: ::--: BanK 100

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FIGURE 16 LOCATIONS OF MONITORING SITES IN RELATION TO ONE MILE GRID 21 I ill I I _. . 'I 20 *~..dn..." __ ~~ ION ISL NO A~ 1.X\ 1'\ I-/, NNS ILLE CRm K 19 I ~- ) : 18 17 16

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FIGURE 18 COMPARISON OF 122M LEVEL .DIURNAL AND WIND DISTRIBUTION i. 160 140 120 100 080 060 040 020 360 340 1979-80 z 320 1974 0 H SU~1ER u 300 E-! r.t:l 0::: 1974 H 280 1979-80 Cl Cl zH 260 3: E-! 240 z

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FIGURE 19 DIURNAL DISTRIBUTION OF WIND SPEEDS 122M SUM..\1ER 122M WINTER 10M SUMMER 10M 1.0 o. 00 02 04 06 08 10 12 .14 16 18 20 22 HOUR OF DAY

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IP2 FSAR UPDATE CHAPTER 3 REACTOR

3.0 DESCRIPTION

The reactor core has a number of fuel regions with fuel assemblies arranged in a zone and/or checkered core pattern. The fuel rods are cold worked Zircaloy or ZIRLOTM tubes containing slightly enriched uranium-dioxide fuel. The fuel assembly is a canless type with the basic assembly consisting of the rod cluster control guide thimbles attached to the grids and the top and bottom nozzles. The fuel rods are held by the spring clip grid in this assembly, which provides support for the fuel rods. High parasitic (HIPAR) fuel was used for the initial fuel and reload fuel through Cycle 4. Low parasitic (LOPAR) fuel was loaded for Cycles 5 through 9, and optimized fuel assemblies (OFA) were loaded for Cycles 10, 11, and 12. For Cycles 13, 14 and 15, 15x15 VANTAGE+ fuel assemblies were loaded as the feed fuel. For Cycle 16, 15x15 Vantage + fuel assemblies with Performance + enhancements are loaded as feed fuel. For Cycle 17, 18, 19 and 20 15x15 Upgraded fuel design assemblies have been loaded as feed fuel. Rod cluster control assemblies and wet annular burnable absorber rods are inserted into the guide thimbles of the fuel assemblies. The absorber sections of the control rods are fabricated of silver-indium-cadmium alloy sealed in stainless steel tubes. The control rod drive mechanisms for the full-length rod cluster control assemblies are of the magnetic latch type. The latches are controlled by three magnetic coils. They are so designed that upon a loss of power to the coils, the rod cluster control assembly is released and falls by gravity to shut down the reactor. 3.1 DESIGN BASES 3.1.1 Performance Objectives The reactor thermal power analyzed is 3216 MWt (3230 MWt total NSSS power), which is the licensed power rating. Calculations and operating experience indicate that hot channel factors will be considerably less than those used for design purposes in this application. The initial reactor core fuel loading and programming were designed to yield the first cycle average burnup of 16,100 MWd/metric ton uranium (MTU). Cycles 2 through 4 reload designs yielded an average cycle burnup of 10,000 MWd/metric ton uranium, and Cycles 5 through 10 yielded an average cycle burnup of 13,500 MWd/metric ton uranium. Cycles 11, 12, and 13 achieved burnups of 18,094 MWd/MTU, 20,674 MWd/MTU, and 21,650 MWd/MTU, all of which included coastdown operation. Cycle 14 achieved a cycle burnup of 18,970 MWd/MTU. Cycle 15 achieved a cycle burnup of 22,110 MWd/MTU. Cycle 16 achieved a cycle burnup of 23,950 MWd/MTU with a coastdown. Cycle 17 achieved a cycle burnup of 18,708 MWd/MTU. Cycle 18 achieved a cycle burnup of 24,456 MWd/MTU. Cycle 20 is designed to achieve a cycle burnup of 25,193 MWd/MTU. The fuel rod cladding is designed to maintain its integrity for the anticipated core life. The effects of gas release, fuel dimensional changes, and corrosion-induced or irradiation-induced changes in the mechanical properties of cladding are considered in the design of the fuel assemblies. Chapter 3, Page 1 of 94 Revision 22, 2010

IP2 FSAR UPDATE Rod cluster control assemblies are employed to provide sufficient reactivity control to terminate any credible power transient prior to reaching the minimum departure from nucleate boiling ratio (DNBR). This is accomplished by ensuring sufficient cluster control worth to shut the reactor down by at least 1.3-percent throughout core life in the hot condition with the most reactive cluster control stuck in the fully withdrawn position. Redundant equipment is provided to add soluble poison to the reactor coolant in the form of boric acid to maintain shutdown margin when the reactor is cooled to ambient temperatures. Experimental measurements from critical experiments or operating reactors, or both, are used to validate the methods employed in the design. During design, nuclear parameters have been calculated for every phase of operation of the first core and reload cycles and, where applicable, are compared with design limits to show that an adequate margin of safety exists. In the thermal hydraulic design of the core, the maximum fuel and clad temperatures during normal reactor operation and at overpower conditions have been conservatively evaluated and found to be consistent with safe operating limitations. 3.1.2 Principal Design Criteria 3.1.2.1 Reactor Core Design Criterion: The reactor core with its related controls and protection systems shall be designed to function throughout its design lifetime without exceeding acceptable fuel damage limits, which have been stipulated and justified. The core and related auxiliary system designs shall provide this integrity under all expected conditions of normal operation with appropriate margins for uncertainties and for specified transient situations, which can be anticipated. (GDC 6) The reactor core, with its related control and protection system, is designed to function throughout its design lifetime without exceeding acceptable fuel damage limits. The core design, together with reliable process and decay heat removal systems, provides for this capability under all expected conditions of normal operation with appropriate margins for uncertainties and anticipated transient situations, including the effects of the loss of reactor coolant flow (Section 14.1.6), trip of the turbine generator (Section 14.1.8), loss of normal feedwater (Section 14.1.9), and loss of all offsite power (Section 14.1.12). The reactor protection system is designed to actuate a reactor trip for any anticipated combination of plant conditions, when necessary, to ensure a minimum departure from nucleate boiling (DNB) ratio equal to or greater than the applicable safety analysis limit DNBR. The integrity of fuel cladding is ensured by preventing excessive fuel swelling, excessive fuel densification, excessive clad heating, excessive cladding stress and strain. This is achieved by designing the fuel rods so that the following conservative limits are not exceeded during normal operation or any anticipated transient condition (Condition II events):

1. Minimum DNB ratio equal to or greater than the safety analysis limit DNBR.

2. Fuel center temperature below 4700°F.

3. Internal gas pressure (limited to value below that which could cause the diameter-to-gap to increase due to outward clad creep during steady-state operation, and which could cause excessive DNB propagation to occur).

4. Clad stresses less than the Zircaloy or ZIRLOTM yield strength.

5. Clad strain less than 1-percent.

Chapter 3, Page 2 of 94 Revision 22, 2010

IP2 FSAR UPDATE The ability of fuel designed and operated to these criteria to withstand postulated normal and abnormal service conditions is shown by analyses described in Chapter 14 to satisfy the demands of plant operation well within applicable regulatory limits. The reactor coolant pumps provided for the plant are supplied with sufficient rotational inertia to maintain an adequate flow coastdown and prevent core damage in the event of a simultaneous loss of power to all pumps as discussed in section 14.1.6.1. In the unlikely event of a turbine trip from full power without an immediate reactor trip, the subsequent reactor coolant temperature increase and volume insurge to the pressurizer result in a high pressurizer pressure trip and thereby prevent fuel damage for this transient. A loss of external electrical load of 50-percent of full power or less is normally controlled by rod cluster insertion together with a controlled steam dump to the condenser to prevent a large temperature and pressure increase in the reactor coolant system and thus prevent a reactor trip. In this case, the overpower-overtemperature protection would guard against any combination of pressure, temperature, and power, which could result in a DNB ratio less than the applicable safety analysis limit DNBR during the transient. In neither the turbine trip nor the loss-of-flow events do the changes in coolant conditions provoke a nuclear power excursion because of the large system thermal inertia and relatively small void fraction as discussed in UFSAR sections 14.1.8 and 14.1.6, respectively. Protection circuits actuated directly by the coolant conditions identified with core limits are therefore effective in preventing core damage. 3.1.2.2 Suppression of Power Oscillations Criterion: The design of the reactor core with its related controls and protection systems shall ensure that power oscillations, the magnitude of which could cause damage in excess of acceptable fuel damage limits, are not possible or can be readily suppressed. (GDC 7) The potential for possible spatial oscillations of power distribution for this core has been reviewed. It is concluded that low frequency xenon oscillations may occur in the axial dimension, and the control rods are provided to suppress these oscillations. The core is expected to be stable to xenon oscillations in the X-Y dimension. Ex-core instrumentation is provided to obtain necessary information concerning power distribution. This instrumentation is adequate to enable the operator to monitor and control xenon induced oscillations. (Incore instrumentation is used periodically to calibrate and verify the information provided by the ex-core instrumentation. The analysis, detection, and control of these oscillations is discussed in Reference 1.) 3.1.2.3 Redundancy of Reactivity Control Criterion: Two independent reactivity control systems, preferably of different principles, shall be provided. (GDC 27) Two independent reactivity control systems are provided, one involving rod cluster control assemblies and the other involving chemical shim. Chapter 3, Page 3 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.1.2.4 Reactivity Hot Shutdown Capability Criterion: The reactivity control systems provided shall be capable of making and holding the core subcritical from any hot standby or hot operating condition. (GDC 28) The reactivity control systems provided are capable of making and holding the core subcritical from any hot standby or hot operating condition, including those resulting from power changes. The maximum excess reactivity expected for the core occurs for the cold, clean condition at the beginning-of-life of the initial core. The rod cluster control assemblies are divided into control banks and shutdown banks. The control banks used in combination with chemical shim control provide control of the reactivity changes of the core throughout the life of the core during power operation. These banks of rod cluster control assemblies are used to compensate for short term reactivity changes at power that might be produced due to variations in reactor power level or in coolant temperature. The chemical shim control is used to compensate for the more slowly occurring changes in reactivity throughout core life such as those due to fuel depletion and fission product buildup. 3.1.2.5 Reactivity Shutdown Capability Criterion: One of the reactivity control systems provided shall be capable of making the core subcritical under any anticipated operating condition including anticipated operational transients sufficiently fast to prevent exceeding acceptable fuel damage limits. Shutdown margin should assure subcriticality with the most reactive control rod fully withdrawn. (GDC 29) The reactor core, together with the reactor control and protection system, is designed so that the minimum allowable DNBR is equal to or greater than the applicable safety analysis limit DNBR and there is no fuel melting during normal operation including anticipated transients. The shutdown banks are provided to supplement the control banks of rod cluster control assemblies to make the reactor at least 1.3-percent subcritical (keff = 0.987) following trip from any credible operating condition assuming the most reactive rod cluster control assembly is in the fully withdrawn position. Sufficient shutdown capability is also provided so that the minimum DNBR is equal to or greater than the applicable safety analysis limit DNBR, assuming the most reactive rod to be in the fully withdrawn position for the most severe anticipated cooldown transient associated with a single active failure, e.g., accidental opening of a steam bypass, or relief valve, or safety valve stuck open. This is achieved by the combination of control rods and automatic boric acid injection via the emergency core cooling system. The minimum design margin is 1.3-percent throughout core life, as discussed in UFSAR Section 14.2.5.2, assuming the maximum worth control rod is in the fully withdrawn position allowing 10-percent uncertainty in the control rod worth calculations. Technical Specification 3.1.1 specifies the actual minimum required shutdown margin for core design. Manually controlled boric acid addition is used to maintain the shutdown margin for the long-term conditions of xenon decay and plant cooldown. Redundant equipment is provided to guarantee the capability of adding boric acid to the reactor coolant system. Chapter 3, Page 4 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.1.2.6 Reactivity Holddown Capability Criterion: The reactivity control systems provided shall be capable of making the core subcritical under credible accident conditions with appropriate margins for contingencies and limiting any subsequent return to power such that there will be no undue risk to the health and safety of the public. (GDC 30) Normal reactivity shutdown capability is provided by control rods following a trip signal, with boric acid injection used to compensate for the long-term xenon decay transient and for plant cooldown. As discussed in response to the previous criteria, the shutdown capability maintains the minimum DNBR above the limiting value and prevents exceeding core safety limits as a result of the cooldown associated with a safety valve stuck fully open. Any time that the reactor is at power, the quantity of boric acid retained in the boric acid tanks and ready for injection always exceeds that quantity required for the normal cold shutdown. This quantity always exceeds the quantity of boric acid required to bring the reactor to hot shutdown and to compensate for subsequent xenon decay. Boric acid is pumped from the boric acid tanks by one of two boric acid transfer pumps to the suction of one of three charging pumps, which injects boric acid into the reactor coolant. Any charging pump and either boric acid transfer pump can be operated from diesel-generator power on loss of station power. On the basis of the above, the injection of boric acid is shown to afford backup reactivity shutdown capability, diverse from rod cluster controls, which normally serve this function in the short-term situation. Shutdown for long-term and reduced temperature conditions can be accomplished with boric acid injection using redundant components, thus achieving the measure of reliability implied by the criterion. Alternately, boric acid solution at lower concentration can be supplied from the refueling water storage tank. This solution can be transferred directly by the charging pumps or alternately by the safety injection pumps. The reduced boric acid concentration lengthens the time required to achieve equivalent shutdown. 3.1.2.7 Reactivity Control Systems Malfunction Criterion: The reactor protection systems shall be capable of protecting against any single malfunction of the reactivity control system, such as unplanned continuous withdrawal (not ejection or dropout of a control rod, by limiting reactivity transients to avoid exceeding acceptable fuel damage limits. (GDC 31) The reactor protection systems are capable of protecting against any single credible malfunction of the reactivity control system, by limiting reactivity transients to avoid exceeding acceptable fuel damage limits. Reactor shutdown with rods is completely independent of the normal rod control functions since the trip breakers completely interrupt the power to the rod mechanisms regardless of existing control signals. Details of the effects of continuous withdrawal of a control rod and continuous deboration are described in Section 14.1 and Section 9.2, respectively. Chapter 3, Page 5 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.1.2.8 Maximum Reactivity Worth of Control Rods Criterion: Limits, which include reasonable margin, shall be placed on the maximum reactivity worth of control rods or elements and on rates at which reactivity can be increased to ensure that the potential effects of a sudden or large change of reactivity cannot (a) rupture the reactor coolant pressure boundary or (b) disrupt the core, its support structures, or other vessel internals sufficiently to lose capability of cooling the core. (GDC 32) Limits, which include considerable margin, are placed on the maximum reactivity worth of control rods or elements and on rates at which reactivity can be increased to ensure that the potential effects of a sudden or large change of reactivity cannot (a) rupture the reactor coolant pressure boundary or (b) disrupt the core, its support structures, or other vessel internals so as to lose capability to cool the core. The reactor control system employs control rod clusters. A portion of these are designated shutdown rods and are fully withdrawn during power operation. The remaining rod clusters comprise the control groups, which are used to control load and reactor coolant temperature. The rod cluster drive mechanisms are wired into preselected groups, and are therefore prevented from being withdrawn in other than their respective groups. The rod drive mechanism is of the magnetic latch type and the coil actuation is sequenced to provide variable speed rod travel. The maximum reactivity insertion rate is analyzed in the detailed plant analysis assuming two of the highest worth groups to be accidentally withdrawn at maximum speed. This yields a reactivity insertion rate of the order of 70 pcm/sec, which is well within the capability of the overpower-overtemperature protection circuits to prevent core damage. No single credible mechanical or electrical control system malfunction can cause a rod cluster to be withdrawn at a speed greater than 72 steps per min (approximately 45-in. per min) when the Rod Control System is operating in the MANUAL mode. Single credible electrical system failures can result in withdrawal speeds greater than 72 steps per min during automatic operation. The automatic rod withdrawal feature however, has been permanently defeated at IP2. 3.1.3 Safety Limits The reactor is capable of meeting the performance objectives throughout core life under both operating and malfunction conditions without violating the integrity of the fuel elements. Thus the release of unacceptable amounts of fission products to the coolant is prevented. The limiting conditions for operation established in the Technical Specifications specify the functional capacity of performance levels permitted to assure safe operation of the facility. Design parameters, which are pertinent to safety limits are specified below for the nuclear, reactivity control, thermal and hydraulic, and mechanical aspects of the design. 3.1.3.1 Nuclear Limits The nuclear axial peaking factor Fz, and the nuclear enthalpy rise hot channel factor FN H are limited in their combined relationship so as not to exceed the FQ or DNBR limits. Chapter 3, Page 6 of 94 Revision 22, 2010

IP2 FSAR UPDATE Potential axial xenon oscillations are controlled with control rods to preclude adverse core conditions. The protection system ensures that the nuclear core limits are not exceeded. For Cycles 13 and beyond, a cycle specific Core Operating Limits Report (COLR) is generated. Provided in the COLR is the cycle specific Fq and F H limits as well as the limiting K(z) curve (normalized Fq* Power versus core height axial envelope). Also, cycle specific rod insertion limits and axial flux difference band width (delta-I) limits are provided in the COLR. Utilization of the COLR provides maximized operational and/or design flexibility, while adherence to the limits of the COLR and plant Technical Specifications assures that DNB and overpower design limits are met. The expected values for the nuclear heat flux and nuclear enthalpy rise hot channel factors for the current cycle are provided in Table 3.2-1A. 3.1.3.2 Reactivity Control Limits The control system and the operational procedures provide adequate control of the core reactivity and power distribution. The following control limits are met:

1. A minimum hot shutdown margin of 1.3% Keff is available throughout core life assuming a 10-percent uncertainty in the control rod calculation.

2. This shutdown margin is maintained with the most reactive rod cluster control assembly in the fully withdrawn position.

3. The shutdown margin is maintained at ambient temperature by the use of soluble poison.

Technical Specification 3.1.1 specifies the actual minimum required shutdown margin for core design. 3.1.3.3 Thermal and Hydraulic Limits The reactor core is designed to meet the following limiting thermal and hydraulic criteria:

1. The minimum allowable DNBR during normal operation, including anticipated transients, is equal to the applicable safety analysis limit DNBR.

2. Fuel temperature not to exceed 4700F during any anticipated operating condition or anticipated malfunction.

To maintain fuel rod integrity and prevent fission product release, it is necessary to prevent clad overheating under all operating conditions. This is accomplished by preventing a departure from nucleate boiling (DNB), which would cause a large decrease in the heat transfer coefficient between the fuel rods and the reactor coolant, resulting in high clad temperatures. The ratio of the heat flux causing DNB at a particular core location (as predicted by the WRB-1 correlation) to the existing heat flux at the same core location is the DNB ratio. A DNB ratio of 1.17 for the WRB-1 correlation corresponds to a 95-percent probability at a 95-percent confidence level that DNB does not occur. The DNB ratio for the W-3 correlation is 1.3 for pressure from 1000 to 2300 psia and 1.45 for pressure from 500 to 1000 psia. Chapter 3, Page 7 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.1.3.4 Mechanical Limits 3.1.3.4.1 Reactor Internals The reactor internal components are designed to withstand the stresses resulting from startup, steady state operation with any number of pumps running, and shutdown conditions. No damage to the reactor internals occurs as a result of loss of pumping power. Lateral deflection and torsional rotation of the lower end of the core barrel is limited to prevent excessive movements resulting from seismic disturbances and thus prevent interference with rod cluster control assemblies. Core drop in the event of failure of the normal supports is limited so that the rod cluster control assemblies do not disengage from the fuel assembly guide thimbles. The internals are further designed to maintain their functional integrity in the event of a major loss-of-coolant accident. The dynamic loading resulting from the pressure oscillations because of a loss-of-coolant accident does not cause sufficient deformation to prevent rod cluster control assembly insertion. 3.1.3.4.2 Fuel Assemblies The fuel assemblies are designed to perform satisfactorily throughout their lifetime. The loads, stresses, and strains resulting from the combined effects of flow induced vibrations, earthquakes, reactor pressure, fission gas pressure, fuel growth, thermal strain, and differential expansion during both steady state and transient reactor operating conditions have been considered in the design of the fuel rods and fuel assemblies. The assemblies are also structurally designed to withstand handling and shipping loads prior to irradiation, and to maintain sufficient integrity at the completion of design burnup to permit safe removal from the core, subsequent handling during cooldown, shipment, and fuel reprocessing. The fuel rods are supported at nine locations along their length within the fuel assemblies by grid assemblies, which are designed to maintain control of the lateral spacing between the rods through the design life of the assemblies. The magnitude of the support loads provided by the grids is established to minimize possible fretting without overstressing the cladding at the points of contact between the grids and fuel rods and without imposing restraints of sufficient magnitude to result in buckling or distortion of the rods. In addition, there are 3 Intermediate Flow Mixing (IFM) grids spaced along the fuel assembly and a protective grid (P-grid) on the bottom of the assembly. These grids do not provide any support function. The fuel rod cladding is designed to withstand operating pressure loads without rupture and to maintain encapsulation of the fuel throughout the design life. 3.1.3.4.3 Rod Cluster Control Assemblies The criteria used for the design of the cladding on the individual absorber rods in the rod cluster control assemblies are similar to those used for the fuel rod cladding. The cladding is designed to be free standing under all operating conditions and will maintain encapsulation of the absorber material throughout the absorber rod design life. Allowance for wear during operation is included for the rod cluster control assembly cladding thickness. Chapter 3, Page 8 of 94 Revision 22, 2010

IP2 FSAR UPDATE Adequate clearance is provided between the absorber rods and the guide thimbles, which position the rods within the fuel assemblies so that coolant flow along the length of the absorber rods is sufficient to remove the heat generated without overheating of the absorber cladding. The clearance is also sufficient to compensate for any misalignment between the absorber rods and guide thimbles and to prevent mechanical interference between the rods and guide thimbles under any operating conditions. 3.1.3.4.4 Control Rod Drive Assembly Each control rod drive assembly is designed as a hermetically sealed unit to prevent leakage of reactor coolant. All pressure-containing components are designed to meet the requirements of the ASME Code, Section III, Nuclear Vessels for Class A vessels. The control rod drive assemblies for the full length rods provide rod cluster control assembly insertion and withdrawal rates consistent with the required reactivity changes for reactor operational load changes. This rate is based on the worths of the various rod groups, which are established to limit power-peaking flux patterns to design values. The maximum reactivity addition rate is specified to limit the magnitude of a possible nuclear excursion resulting from a control system or operator malfunction. Also, the control rod drive assemblies for the full length rods provide a fast insertion rate during a "trip" of the rod cluster control assemblies, which results in a rapid shutdown of the reactor for conditions that cannot be handled by the reactor control system. REFERENCES FOR SECTION 3.1

1. Westinghouse Proprietary, "Power Distribution Control in Westinghouse Pressurized Water Reactors," WCAP-7208, 1968.

3.2 REACTOR DESIGN 3.2.1 Nuclear Design And Evaluation This section presents the nuclear characteristics of the core and an evaluation of the characteristics and design parameters, which are significant to design objectives. The capability of the reactor to achieve these objectives while performing safely under normal operational modes, including both transient and steady state, is demonstrated. 3.2.1.1 Nuclear Characteristics of the Design A summary of the reactor nuclear design characteristics is presented in Table 3.2-1 for Cycle 1 and Table 3.2-1A for the current cycle. Some of the presented parameters will change from cycle to cycle depending on reload core designs. 3.2.1.1.1 Reactivity Control Aspects Reactivity control is provided by neutron absorbing control rods and by a soluble chemical neutron absorber (boric acid) in the reactor coolant. The concentration of boric acid is varied as necessary during the life of the core to compensate for: Chapter 3, Page 9 of 94 Revision 22, 2010

IP2 FSAR UPDATE

1. changes in reactivity, which occur with changes in temperature of the reactor coolant from cold shutdown to the hot operating, zero power conditions,

2. changes in reactivity associated with changes in the fission product poisons xenon and samarium,

3. reactivity losses associated with the depletion of fissile inventory and buildup of long-lived fission product poisons (other than xenon and samarium), and

4. changes in reactivity due to burnable poison burnup.

The control rods provide reactivity control for:

1. fast shutdown,

2. reactivity changes associated with changes in the average coolant temperature above hot zero power (core average coolant temperature is increased with power level),

3. reactivity associated with any void formation, and

4. reactivity changes associated with the power coefficient of reactivity.

3.2.1.1.1.1 Chemical Shim Control Control to render the reactor sub-critical at temperatures below the operating range is provided by a chemical neutron absorber (boron). The boron concentration during refueling following Cycle 1 has been established as shown in Table 3.2-1, line 29. This concentration together with the control rods provides approximately 10-percent shutdown margin for these operations. The concentration is also sufficient to maintain the core shutdown without any rod cluster control rods during refueling. For cold shutdown, at the beginning of core life, a concentration (shown in Table 3.2-1, line 37) is sufficient for 1-percent shutdown with all but the highest worth rod inserted. The boron concentration (Table 3.2-1, line 29) for refueling is equivalent to less than 2-percent by weight boric acid (H3BO3) and is well within solubility limits at ambient temperature. This concentration is also maintained in the spent fuel pit since it is directly connected with the refueling canal during refueling operations. The refueling boron concentration requirement for the current cycle is shown in Table 3.2-1A. For example, the initial Cycle 1 full power boron concentration without equilibrium xenon and samarium was 1186 ppm. As the fission product poisons built up, the boron concentration was reduced to 890 ppm. This initial boron concentration was that which permitted the withdrawal of the control banks to their operational limits. The Cycle 1 xenon-free, hot zero power shutdown (k = 0.99) with all but the highest worth rod inserted, could be maintained with the boron concentration of 677 ppm. This concentration is less than the full power operating value with equilibrium xenon. 3.2.1.1.1.2 Control Rod Requirements Neutron-absorbing control rods provide reactivity control to compensate for more rapid variations in reactivity. The rods are divided into two categories according to their function. Some rods compensate for changes in reactivity due to variations in operating conditions of the reactor such as power or temperature. These rods comprise the control group of rods. The remaining rods, which provide shutdown reactivity, are termed shutdown rods. The total shutdown worth of all the rods is also specified to provide adequate shutdown with the most reactive rod stuck out of the core as discussed in Sections 14.1, 14.2.5, and 14.2.6. Chapter 3, Page 10 of 94 Revision 22, 2010

IP2 FSAR UPDATE Control rod reactivity requirements at beginning- and end-of-life for Cycle 1 are summarized in Table 3.2-2. The installed worth of the control rods for Cycle 1 is shown in Table 3.2-3. These values will vary from cycle to cycle depending on the reload core design. The difference is available for excess shutdown upon reactor trip. The control rod requirements are discussed below. 3.2.1.1.1.3 Total Power Reactivity Defect Control rods must be available to compensate for the reactivity change incurred with a change in power level due to the Doppler effect. The magnitude of this change was determined by measurement during the Cycle 1 Startup test program. The average temperature of the reactor coolant is increased with power level in the reactor. Since this change is actually a part of the power dependent reactivity change, along with the Doppler effect and void formation, the associated reactivity change must be controlled by rods. The largest amount of reactivity that must be controlled is at the end-of-life when the moderator temperature coefficient has its most negative value. The Cycle 1 moderator density coefficient range is given in Table 3.2-1, line 44, while the cumulative reactivity change for Cycle 1 is shown in the first line of Table 3.2-2. By the end of the fuel cycle, the nonuniform axial depletion causes a severe power peak at low power. The reactivity associated with this peak is part of the power defect. 3.2.1.1.1.4 Operational Maneuvering Band The control group is operated at full power within a prescribed band of travel in the core to compensate for periodic changes in boron concentration, temperature, or xenon. The band has been defined as the operational maneuvering band. When the rods reach either limit of the band, a change in boron concentration must be made to compensate for any additional change in reactivity, thus keeping the control group within the maneuvering band. 3.2.1.1.1.5 Control Rod Bite If sufficient boron is present in a chemically shimmed core, the inherent operational control afforded by the negative moderator temperature coefficient is lessened to such a degree that the major control of transients resulting from load variations must be compensated for by control rods. The ability of the plant to accept major load variations is distinct from safety considerations, since the reactor would be tripped and the plant shut down safely if the rods could not follow the imposed load variations. In order to meet required reactivity ramp rates resulting from load changes, the control rods were inserted a given distance into the core. The reactivity worth of this insertion has been defined as control rod bite. The reactivity insertion rate was sufficient to compensate for reactivity variation due to changes in power and temperature caused either by a ramp load change of 5-percent/min, or by a step load change of 10-percent. An insertion rate of 3 x 10-5 /sec is determined by the transient analysis of the core and plant to be adequate for the most adverse combinations of power and moderator coefficients. Chapter 3, Page 11 of 94 Revision 22, 2010

IP2 FSAR UPDATE Beginning with Cycle 18, plant operation with Control Bank D at bite position is no longer required. The analysis in support of the stretch power uprate eliminated the need for control rod bite. 3.2.1.1.1.6 Xenon Stability Control Control rods are capable of suppressing xenon-induced power oscillations in the axial direction, should they occur. Ex-core instrumentation is provided to obtain necessary information concerning power distribution. This instrumentation is adequate to enable the operator to monitor and control xenon-induced power oscillations. Extensive analyses, with confirmation of methods by spatial transient experiments at Haddam Neck, have shown that any induced radial or diametral xenon transients would die away naturally. A full discussion of xenon stability control can be found in Reference 1. In assessing potential power distribution instabilities arising from spatial xenon redistribution and in determining stability indices, primary reliance has been placed on time-dependent digital calculations in three dimensions representing feedback reactivity effects by means of semi-empirically fitted expressions whose coefficients were determined from other calculations using standard design analytical techniques and computer codes (e.g., LEOPARD code). To assess the level of credibility and range of uncertainty attached to xenon stability analyses, conservative values of the reactivity feedback parameters were used to arrive at a reasonable upper limit for the stability index. This technique gives reasonable assurance that the reactor will in fact be stable toward diametral xenon oscillations. Means are available to increase the moderator temperature feedback term in order to stabilize the reactor response to diametral xenon oscillation (Reference 2). The reference three-dimensional time dependent digital calculations have indicated that the Indian Point Unit 2 reactor is stable against diametral xenon transients. Cross-coupled transients are discussed in Reference 3 on the basis of full three-dimensional analyses of xenon transients. "Second overtone" xenon transients from quadrant to quadrant (X-Y transient) are also discussed in Reference 3 in the form of radial transients. That is to say that a power-xenon perturbation is introduced by moving the center control rod. "Second overtone" xenon transients from top-to-bottom (axial transient) have been analyzed with results presented in References 1 and 4. For clarity, such transients are presented here specifically for the Indian Point Unit 2 reactor. Figure 3.2-1 shows a cross-plot of the axial peaking factor and axial offset for transients at three points in core life. This plot is completely in agreement with such plots shown in Reference 1. In short, the ex-core detector based protection system is capable of detecting (by means of axial offset) transients, which result from perturbations to the "second axial overtone" of the power distribution. The separation of dimensions is a conceptual artifice, which greatly facilitates analysis of xenon transients. Full three-dimensional transient analyses have been performed, as reported in References 2 and 3. Conclusions relating to power distribution stability against spatial xenon redistribution are based on results of these analyses. Cross-coupling between axial and diametral xenon oscillations are inherently accounted for in the three- dimensional time dependent calculations. Results of these calculations do not reveal any unique problems arising from cross-coupling. See Appendix 3B for additional discussion. Chapter 3, Page 12 of 94 Revision 22, 2010

IP2 FSAR UPDATE If the core was originally operating with a symmetric quadrant-to-quadrant power distribution, the effect of excess xenon poisoning would be to flatten the power distribution because the xenon excess would be greatest where the equilibrium power has been greatest. While this additional power flattening would tend to decrease the stability of the reactor, the analytical evaluation described in Reference 3 had already assumed a power distribution flatter than expected in the actual reactors; thus, an allowance has already been made in the Reference 3 analysis to account for this effect. Furthermore, the high xenon inventory present under the postulated conditions (i.e., maximum xenon buildup) would decrease the required boron concentration and lead to a more stable reactor response from the enhanced negative moderator reactivity feedback effect. As burnup progresses, the required boron concentration further decreases resulting in an increasingly more stable reactor response to diametral xenon oscillations. This effect is greater than the effect of burnup on radial power flattening. If the equilibrium power was not balanced from quadrant to quadrant, the effect of the excess xenon poisoning might be to cause the quadrant power to reverse and perhaps to increase in magnitude. If the quadrant power tilt were to reach the limit given in the Technical Specifications, the operator would take action to maintain core thermal margins. In any event, the excess xenon, which might be present under the conditions postulated would decay naturally and cannot be regarded as a continuing source of power distribution anomalies. Similarly, a top-to-bottom power imbalance could be temporarily increased by the excess xenon poisoning. Such an imbalance cannot be a safety problem because the reactor protection system is cognizant of the axial power imbalance and if necessary will reduce trip setpoints accordingly. 3.2.1.1.1.7 Excess Reactivity Insertion Upon Reactor Trip The control requirements are nominally based on providing 1-percent shutdown at hot, zero power conditions with the highest worth rod stuck in its fully withdrawn position or to prevent return to criticality following a credible steamline break, whichever is the more limiting. The condition where excess reactivity insertion is most critical is at the end of a cycle when the steam break accident is considered. For example, the excess control available at the end of Cycle 1, hot zero power condition with the highest worth rod stuck out is 2.08-percent after allowing a 10-percent margin for uncertainty in control rod worth as shown in Table 3.2-3. 3.2.1.1.1.8 Calculated Rod Worths The complement of 53 full length control rods arranged in the pattern shown in Figure 3.2-2 meets the shutdown requirements. Table 3.2-3 lists the calculated worths of this rod configuration for beginning and end of the first cycle. The calculated reactivity worths listed are decreased in the design by 10-percent to account for any errors or uncertainties in the calculation. This worth is established for the condition that the highest worth rod is stuck in the fully withdrawn position in the core. A comparison between calculated and measured rod worths in the operating reactor shows the calculation to be well within the allowed uncertainty of 10-percent. Chapter 3, Page 13 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2.1.2 Reactor Core Power Distribution The accuracy of power distribution calculations has been confirmed through approximately 1000 flux maps during some 20 years of operation under conditions very similar to those expected for the plant described herein. Details of this confirmation are given in Reference 5. 3.2.1.2.1 Definitions Power distributions are quantified in terms of hot channel factors. These factors are a measure of the peak pellet power within the reactor core and the total energy produced in a coolant channel and are expressed in terms of quantities related to the nuclear or thermal design, namely:

• Power density is the thermal power produced per unit volume of the core (kW/liter).
• Linear power density is the thermal power produced per unit length of active fuel (kW/ft). Since fuel assembly geometry is standardized, this is the unit of power density most commonly used. For all practical purposes it differs from kW/liter by a constant factor, which includes geometry and the fraction of the total thermal power that is generated in the fuel rod.
• Average linear power density is the total thermal power produced in the fuel rods divided by the total active fuel length of all rods in the core.
• Local heat flux is the heat flux at the surface of the cladding (BTU-ft-2-hr-1). For nominal rod parameters this differs from linear power density by a constant factor.
• Rod power or rod integral power is the length integrated linear power density in one rod (kW).
• Average rod power is the total thermal power produced in the fuel rods divided by the number of fuel rods (assuming all rods have equal length).

The hot channel factors used in the discussion of power distributions in this section are defined as follows: F Q' Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods. F NQ ' Nuclear Heat Flux Hot Channel Factor, is defined as the maximum local fuel rod linear power density divided by the average fuel rod linear power density, assuming nominal fuel pellet and rod parameters. F EQ' Engineering Heat Flux Hot Channel Factor, is the allowance on heat flux required for manufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod, and eccentricity of the gap between pellet and Chapter 3, Page 14 of 94 Revision 22, 2010

IP2 FSAR UPDATE pellet and clad. Combined statistically the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux. F NH' Nuclear Enthalpy Rise Hot Channel Factor, is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power. Manufacturing tolerances, hot channel power distribution and surrounding channel power distributions are treated explicitly in the calculation of the DNBR. It is convenient for the purposes of discussion to define subfactors of FQ. However, design limits are set in terms of the total peaking factor. FQ = total peaking factor (or heat flux hot-channel factor) maximum kW / ft

                 =

average kW / ft without densification effects FQ = F NQ x F EQ

              = F NXY x F NZ x F EQ x F NU where:

F NQ and F EQ are defined above. F NU = factor for conservatism, assumed to be 1.05. F NXY = ratio of peak power density to average power density in the horizontal plane of peak local power F NZ = ratio of the power per unit core height in the horizontal plane of peak local power to the average value of power per unit core height. If the plane of peak local power coincides with the plane of maximum power per unit core height, then F NZ is the core average axial peaking factor. To include the allowance made for densification effects, which are height dependent, the following quantities are defined. S(Z) = the allowance made for densification effects at height Z in the core. P(Z) = ratio of the power per unit core height in the horizontal plane at height Z to the average value of power per unit core height. Then: Chapter 3, Page 15 of 94 Revision 22, 2010

IP2 FSAR UPDATE maximum kW/ft FQ = average kW/ft Including densification allowance FQ = max [F NXY (Z) x P(Z) x S(Z) x F NU x F EQ ] 3.2.1.2.2 Radial Power Distributions The power shape in horizontal sections of the core at full power is a function of the fuel and burnable absorber loading patterns, and the presence or absence of a single bank of control rods. Thus, at any time in the cycle, a horizontal section of the core can be characterized as (1) unrodded, or (2) with group D control rods. These two situations combined with burnup effects determine the radial power shapes, which can exist in the core at full power. The effect on radial power shapes of power level, xenon, samarium, and moderator density effects are also considered but these are quite small. The effect of nonuniform flow distribution is negligible. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater interest. As an example, Historical Figures 3.2-3 through 3.2-6 show representative radial cycle 1 power distributions for one quarter of the core for representative operating conditions. These conditions are (1) Hot Full Power (HFP) - beginning-of-life (BOL) - unrodded - no xenon, (2) HFP-BOL - Bank D in - equilibrium xenon - unrodded, (3) HFP - end-of-life (EOL) - unrodded - equilibrium xenon, (4) HFP, BOL, no xenon, part-length rods in. Figure 3.2-6 is of historical significance only, since part-length rods have been removed. Since the position of the hot channel varies from time to time a single reference radial design power distribution is selected for DNB calculations. This reference power distribution is chosen conservatively to concentrate power in one area of the core, minimizing the benefits of flow redistribution. Assembly powers are normalized to core average power. 3.2.1.2.3 Axial Power Distributions The shape of the power profile in the axial or vertical direction is largely under the control of the operator through either the manual operation of the control rods or automatic motion of rods responding to manual operation of the soluble boron system. Nuclear effects, which cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon, and burnup. Automatically controlled variations in total power output and rod motion are also important in determining the axial power shape at any time. Signals are available to the operator from the ex-core ion chambers, which are long ion chambers outside the reactor vessel running parallel to the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top and bottom signals from each of four pairs of detectors is displayed on the control panel and called the flux difference, I. Calculations of core average peaking factor for many plants and measurements from operating plants under many operating situations are associated with either I or axial offset in such a way that an upper bound can be placed on the peaking factor. For these correlations axial offset is defined as: Chapter 3, Page 16 of 94 Revision 22, 2010

IP2 FSAR UPDATE t - b axial offset = t + b and t and b are the top and bottom detector readings. 3.2.1.2.4 Local Power Peaking Fuel densification, which has been observed to occur under irradiation in several operating reactors, causes the fuel pellets to shrink both axially and radially. The pellet shrinkage combined with random hang-up of fuel pellets results in gaps in the fuel column when the pellets below the hung-up pellet settle in the fuel rod. The gaps vary in length and location in the fuel rod. Because of decreased neutron absorption in the vicinity of the gap, power peaking occurs in the adjacent fuel rods resulting in an increased power peaking factor. A quantitative measure of this local peaking is given by the power spike factor S(Z) where Z is the axial location in the core. The method used to compute the power spike factor is described in Reference 6. Results reported in Reference 14 show that fuel manufactured by Westinghouse will not densify and therefore no power spike penalty should be included in the safety analysis The power spike factor due to densification is assumed to be a local perturbation applicable to overpower transients. Thus, densification affects FQ but not FNH . The magnitude of the increased power peaking increases from no effect at the bottom of the core to a few percent at the top of the core. For fuel produced by a process other than those for which Reference 6 is applicable, specifications will be followed to ensure that the effects of densification will be no greater than has been allowed for in the design. The specifications for qualifying the extent of densification will be based on the NRC report on fuel densification (Reference 7). Results reported in a Westinghouse Topical Report concerning the spike penalty in LOCA analysis (Reference 8) show that the power spike penalty does not have to be included in the LOCA envelope. 3.2.1.2.5 Limiting Power Distributions According to the ANS classification of plant conditions, Condition I occurrences are those which are expected frequently or regularly in the course of power operation, maintenance, or maneuvering of the plant. As such, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter, which would require either automatic or manual protective action. Inasmuch as Condition I occurrences occur frequently or regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III, and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions, which can occur during Condition I operation. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating procedures for maintaining appropriate power distributions and takes any necessary remedial actions when alerted to do so by the plant instrumentation. Thus, as stated above, the worst or limiting power distribution, which can occur during normal operation is to be considered as the starting point for analysis of Condition II, III, and IV events. Chapter 3, Page 17 of 94 Revision 22, 2010

IP2 FSAR UPDATE Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (Condition II). Therefore, the limiting power shapes, which result from such Condition II events are those power shapes, which deviate from the normal operating condition at the recommended axial offset band, e.g., due to lack of proper action by the operator during a xenon transient following a change in power level brought about by control rod motion. Power shapes, which fall in this category are used for determination of the reactor protection system setpoints so as to maintain margin to overpower or DNB limits. The means for maintaining power distributions within the required hot channel factor limits are described in the Technical Specifications. A complete discussion of power distribution control in Westinghouse PWRs is included in Reference 9. Detailed background information on the design constraints on local power density in a Westinghouse PWR, on the defined operating procedures, and on the measures taken to preclude exceeding design limits is presented in the Westinghouse topical report on power distribution control and load following procedures (Reference 4). The following paragraphs summarize these reports and describe the calculations used to establish the upper bound on peaking factors. The calculations used to establish the upper bound on peaking factors, FQ and F NH , include all of the nuclear effects, which influence the radial and/or axial power distributions throughout core life for various modes of operation including load follow, reduced power operation, and axial xenon transients. Radial power distributions are calculated for the full power condition and fuel and moderator temperature feedback effects are included for the average enthalpy plane of the reactor. The steady-state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution effects neglected. The effect of flow redistribution is calculated explicitly where it is important in the DNB analysis of accidents. The effect of xenon on radial power distribution is small but is included as part of the normal design process. Radial power distributions are relatively fixed and easily bounded with upper limits. The core average axial profile, however, can experience significant changes, which can occur rapidly as a result of rod motion and load changes and more slowly due to xenon distribution. For the study of points of closest approach to axial power distribution limits, several thousand cases are examined. Since the properties of the nuclear design dictate what axial shapes can occur, boundaries in the limits of interest can be set in terms of the parameters, which are readily observed in the plant. Specifically, the nuclear design parameters, which are significant to the axial power distribution analysis are:

1. Core power level.

2. Core height.

3. Coolant temperature and flow.

4. Coolant temperature program as a function of reactor power.

5. Fuel cycle lifetimes.

6. Rod bank worths.

7. Rod bank overlaps.

Normal operation of the plant assumes compliance with the following conditions:

1. Control rods in a single bank move together with no individual rod insertion differing from the bank demand position by more than the Technical Specification limit.

Chapter 3, Page 18 of 94 Revision 22, 2010

IP2 FSAR UPDATE

2. Control banks are sequenced with overlapping banks.

3. The control bank insertion limits are not violated.

4. Axial power distribution procedures, which are given in terms of flux difference control and control bank position, are observed.

The axial power distribution procedures referred to above are part of the required operating procedures, which are followed in normal operation. Briefly they require control of the axial offset (flux difference divided by fractional power) at all power levels within a permissible operating band of a target value corresponding to the equilibrium full power value. In the first cycle, the target value changes from about +10 to percent linearly through the life of the cycle. This minimizes xenon transient effects on the axial power distribution since the procedures essentially keep the xenon distribution in phase with the power distribution. Calculations are performed for normal operation of the reactor including load following maneuvers. Beginning, middle, and end of cycle conditions are included in the calculations. Different histories of operation are assumed prior to calculating the effect of load follow transients on the axial power distribution. These different histories assume base loaded operation and extensive load following. For a given plant and fuel cycle, a finite number of maneuvers are studied to determine the general behavior of the local power density as a function of core elevation. These cases represent many possible reactor states in the life of one fuel cycle, and they have been chosen as sufficiently definitive of the cycle by comparison with much more exhaustive studies performed on some 20 or 30 different, but typical, plant and fuel cycle combinations. The cases are described in detail in Reference 4, and they are considered to be necessary and sufficient to generate a local power density limit, which, when increased by 5-percent for conservatism, will not be exceeded with a 95-percent confidence level. Many of the points do not approach the limiting envelope. However, they are part of the time histories, which lead to the hundreds of shapes, which do define the envelope. They also serve as a check that the reactor studied is typical of those more exhaustively studied. Thus, it is not possible to single out any transient or steady-state condition, which defines the most limiting case. It is not even possible to separate out a small number, which form an adequate analysis. The process of generating a myriad of shapes is essential to the philosophy that leads to the required level of confidence. A maneuver, which provides a limiting case for one reactor fuel cycle is not necessarily a limiting case for another reactor or fuel cycle with different control bank worths, enrichments, burnup, coefficients, etc. Each shape depends on the detailed history of operation up to that time and on the manner in which the operator conditioned xenon in the days immediately prior to the time at which the power distribution is calculated. The calculated points are synthesized from axial calculations combined with radial factors appropriate for rodded and unrodded planes in the first cycle. In these calculations, the effects on the unrodded radial peak of xenon redistribution that occurs following the withdrawal of a control bank (or banks) from a rodded region is obtained from two-dimensional X-Y calculations. A 1.03 factor to be applied on the unrodded radial peak was obtained from calculations in which xenon distribution was preconditioned by the presence of control rods and then allowed to redistribute for several hours. A detailed discussion of this effect may be found in Reference 4. The calculated values have been increased by a factor of 1.05 for conservatism and a factor of 1.03 for the engineering factor FEQ. Chapter 3, Page 19 of 94 Revision 22, 2010

IP2 FSAR UPDATE The envelope drawn over the calculated ([FQ x Power] max) points in Figure 3.2-7 represents an upper bound envelope on local power density versus elevation in the core for Cycles 1 through

12. For Cycles 13 and on, the anticipated normalized FQ times power versus core height limiting K(z) curve is shown in the Unit 2 COLR. It should be emphasized that this envelope is a conservative representation of the bounding values of local power density. Expected values are considerably smaller and, in fact, less conservative bounding values may be justified with additional analysis or surveillance requirements. Additionally, Figure 3.2-7 is based on a radial power distribution invariant with core elevation.

Finally, as previously discussed, this upper bound envelope is based on procedures of load follow, which require operation within an allowed deviation from a target equilibrium value of axial flux difference. [Note - Per Confirmatory Order for Indian Point Unit 2 of February 11, 1980, (letter from W. J. Cahill, Con Edison, to A. Schwencer, NRC), Indian Point Unit 2 is not operated presently in a load follow mode of operation.] These procedures are detailed in the Technical Specifications and are followed by relying only upon ex-core surveillance supplemented by the normal monthly incore core map requirement and by computer based alarms on deviation and time of deviation from the allowed flux difference band. To determine reactor protection system setpoints with respect to power distributions, three categories of events are considered, namely rod control equipment malfunctions, operator errors of commission and operator errors of omission. In evaluating these three categories of events, the core is assumed to be operating within the four constraints described above. The first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank sequence). Also included are motions of the banks below their insertion limits, which could be caused, for example, by uncontrolled dilution or primary coolant cooldown. Power distributions are calculated throughout these occurrences assuming short term corrective action, that is, no transient xenon effects are considered to result from the malfunction. The event is assumed to occur from typical normal operating situations, which include normal xenon transients. It is further assumed in determining the power distributions that total core power level will be limited by reactor trip to below 120-percent. Since the study is to determine protection limits with respect to power and axial off-set, no credit is taken for trip setpoint reduction due to flux difference. The peak power density, which can occur in such events, assuming reactor trip at or below 120-percent, is less than that required for centerline melt, including uncertainties and densification effects. The second category assumes that the operator malpositions the rod bank in violation of the insertion limits and creates short-term conditions not included in normal operating conditions. The third category assumes that the operator fails to take action to correct a flux difference violation. The resulting FQ is multiplied by 102-percent power including an allowance for calorimetric error. It should be noted that a reactor overpower accident is not assumed to occur coincident with an independent operator error. Analysis of possible operating power shapes shows that the appropriate hot channel factors FQ and F NH for peak local power density and for DNB analysis at full power are the values addressed in the Technical Specifications. FQ can be increased with decreasing power as shown in the Technical Specifications. Increasing Chapter 3, Page 20 of 94 Revision 22, 2010

IP2 FSAR UPDATE Increasing F NH with decreasing power is permitted by the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits. It has been determined that Technical Specifications are met provided that, during normal operation of the plant, there is compliance with the four conditions listed earlier in this section. When a situation is possible in normal operation, which could result in local power densities in excess of those assumed as the precondition for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms are provided for returning the core to a safe condition. 3.2.1.2.6 Power Distribution Anomalies A discussion of the means provided to monitor and control power distributions anomalies caused by misplaced control rods and xenon oscillations is given in Appendix 3B. A description of the protective function in the event of axial xenon oscillations, including calculated peaking factors and DNBRs, and the automatic trip setpoint reduction is given in References 1 and 4. Additional information on the response to ex-core ion chambers, including comparison with experimental information, is given in References 2 and 10. X-Y control rods are not required nor are they employed in the Indian Point Unit 2 reactor. A discussion of the consequences of control rod malposition is given in Appendix 3B and in Reference 2. No automatic protective function is necessary, since even the complete misalignment of a control rod in the most limiting case (see Reference 2) cannot lead to a DNBR less than the applicable safety analysis limit at operating conditions. Furthermore, (1) rod position indicators are provided, (2) the existence of an asymmetric control rod misalignment would be revealed by the ex-core instrumentation, and (3) both asymmetric and symmetric control rod misalignments can readily be detected by the incore thermocouple system as indicated in References 2 and 10. 3.2.1.2.7 Reactivity Coefficients The response of the reactor core to plant conditions or operator adjustments during normal operation, as well as the response during abnormal or accidental transients, is evaluated by means of a detailed plant simulation. In these calculations, reactivity coefficients are required to couple the response of the core neutron multiplication to the variables, which are set by conditions external to the core. Since the reactivity coefficients change during the life of the core, a range of coefficients is established to determine the response of the plant throughout life and to establish the design of the reactor control and protection system. 3.2.1.2.7.1 Moderator Temperature Coefficient The moderator temperature coefficient in a core controlled by chemical shim is less negative than the coefficient in an equivalent rodded core. One reason is that control rods contribute a negative increment to the coefficient and in a chemical shim core, the rods are only partially inserted. Also, the chemical poison density is decreased with the water density upon an increase in temperature. This gives rise to a positive component of the moderator temperature coefficient due to boron being removed from the core. This is directly proportional to the amount of reactivity controlled by the dissolved poison. Chapter 3, Page 21 of 94 Revision 22, 2010

IP2 FSAR UPDATE In order to reduce the dissolved poison requirement for control of excess reactivity, burnable absorber rods have been incorporated in the core design. The result is that changes in the coolant density will have less effect on the density of poison and the moderator temperature coefficient will be reduced. The moderator temperature coefficient is negative at the operating coolant temperature with burnable absorber rods installed. The original burnable absorber was in the form of borated Pyrex glass rods clad in stainless steel. The rods (1412 in the initial core) in the form of clusters were distributed throughout the initial core in vacant rod cluster control guide tubes as illustrated in Figures 3.2-8 and 3.2-9. Information regarding research, development, and nuclear evaluation of the burnable poison rods can be found in Reference 11. These rods initially controlled 7.2-percent of the installed excess reactivity and their addition resulted in a reduction of the initial hot zero power boron concentration in the coolant to 1318 ppm. Starting with Cycle 8, the Wet Annular Burnable Absorber (WABA) design has been used. The WABA design is described in Section 3.2.3.2.1.5. Starting with Cycle 11, the Integral Fuel Burnable Absorber (IFBA) design has also been used. The IFBA design is described in Section 3.2.3.2.1.5. The effect of burnup on the moderator temperature coefficient is calculated and the coefficient becomes more negative with increasing burnup. This is due to the buildup of fission products with burnup and dilution of the boric acid concentration with burnup. The reactivity loss due to equilibrium xenon is controlled by boron, and as xenon builds up, boron is taken out. For example, the calculated net effect and the predicted unrodded moderator temperature coefficient equilibrium xenon for Cycle 1 at full power BOL was -0.55 x 10-4/oF. With core burnup, the coefficient became more negative as boron was removed due to the buildup of plutonium and fission products. At Cycle 1 end-of-life with no boron or rods in the core, the moderator coefficient was -3.0 x 10-4/oF. Variation in moderate temperature can be seen, for example, from the Cycle 1 Figures 3.2-10 through 3.2-12. 3.2.1.2.7.2 Moderator Pressure Coefficient The moderator pressure coefficient has an opposite sign to the moderator temperature coefficient. Its effect on core reactivity and stability is small because of the small magnitude of the pressure coefficient, a change of 50 psi in pressure having no more effect on reactivity than a one-half degree change in moderator temperature. The calculated Cycle 1 beginning and end-of-life pressure coefficients are specified in Table 3.2-1, line 43. 3.2.1.2.7.3 Moderator Density Coefficient A uniform moderator density coefficient is defined as a change in the neutron multiplication per unit change in moderator density. The range of the moderator density coefficient for Cycle 1, for example, from BOL and EOL is specified in Table 3.2-1. 3.2.1.2.7.4 Doppler and Power Coefficients The Doppler coefficient is defined as the change in neutron multiplication [Note - Neutron multiplication is defined as the ratio of the average number of neutrons produced by fission in Chapter 3, Page 22 of 94 Revision 22, 2010

IP2 FSAR UPDATE each generation to the total number of corresponding neutrons absorbed.] per degree change in fuel temperature. The coefficient is obtained by calculating neutron multiplication as a function of effective fuel temperature. As an example, the Cycle 1 results, using the LEOPARD code (Reference 12), are shown in Figure 3.2-13. In order to know the change in reactivity with power, it is necessary to know the change in the effective fuel temperature with power as well as the Doppler coefficient. It is very difficult to predict the effective temperature of the fuel using a conventional heat transfer model because of uncertainties in predicting the behavior of the fuel pellets. Therefore, an empirical approach is taken to calculate the power coefficient, based on operating experience of existing Westinghouse cores. As an example, Figure 3.2-14 shows the power coefficient as a function of power for Cycle 1 obtained by this method. The results presented do not include any moderator coefficient even though the moderator temperature changes with power level. As the fuel pellet temperature increases with power, the resonance absorption in U-238 increases due to Doppler broadening of the resonances. A large temperature drop occurs across the fuel pellet-clad gap. Under certain conditions, this gap may be closed, thus resulting in lower pellet temperature. The net effect is a lower effective fuel temperature, a higher (more negative) Doppler coefficient, and a lower (less negative) power coefficient than that which exists with a pellet-clad gap. For example, the power coefficient for Cycle 1, which was determined using a closed gap model, is shown in Figure 3.2-15. Calculations indicate the stability of the reactor to xenon oscillations is relatively insensitive to the thermal model used to obtain the power coefficient. The damping factor associated with the fuel Doppler effect is: Keff T f = x T P where: T = fuel temperature P = power T The quantity is larger for the gap model than for the no gap case but since the Doppler P coefficient varies as T -1/2 the term Keff is smaller. T The net effect is that f is relatively insensitive to the thermal model in the range of power 0.5 to 1.5 of core average, which is the range of interest for stability. 3.2.1.3 Nuclear Evaluation of Current Core Three principal computer codes have been used in the nuclear design on this reload cycle. These are: PHOENIX-P (two-dimensional), ANC (two-dimensional and three-dimensional) and APOLLO (one-dimensional). Descriptions and uses for these codes are given below: PHOENIX-P is a two dimensional, multi-group transport theory code which utilizes a 70 energy-group cross section library. It provides the capability for cell lattice modeling on an assembly Chapter 3, Page 23 of 94 Revision 22, 2010

IP2 FSAR UPDATE level. In this design, PHOENIX-P is used to provide homogenized, two-group cross sections for nodal calculations and feedback models. Also, PHOENIX-P is used to generate appropriately weighted constants for the baffle/reflector regions. ANC is an advance nodal code capable of two-dimensional and three-dimensional calculations. In this design, ANC is employed as the reference model for all safety analysis calculations, power distributions, peaking factors, critical boron concentrations, control rod worth, reactivity coefficients, etc. In addition, 3D ANC is used to validate one and two-dimensional results and to provide information about radial (x-y) peaking factors as a function of axial position. It has the capability of calculating discrete pin powers from the nodal information as well. APOLLO, an advanced version of PANDA, is a two group, one-dimensional diffusion-depletion code. It uses cross sections generated by a radial averaging of the corresponding 3D model cross sections and is used as a one-dimensional axial model. Thermal feedback is included in the calculational models. The axial model is used for computing axial power distributions, differential rod worths, control rod operating limits (insertion limits, return to power limits), etc. Additional support codes are used for special calculations such as determining fuel temperatures. 3.2.2 Thermal and Hydraulic Design and Evaluation [Note - A large amount of material has been retained as historical background.] 3.2.2.1 Thermal and Hydraulic Characteristics of the Design 3.2.2.1.1 Central Temperature of the Hot Pellet The temperature distribution in the pellet is mainly a function of the uranium-dioxide thermal conductivity and the local power density. The absolute value of the temperature distribution is affected by the cladding temperature and the thermal conductance of the gap between the pellet and the cladding. The gap conductance model is selected such that when combined with the UO2 thermal conductivity model, the calculated fuel centerline temperatures reflect the inpile temperature measurements. A more detailed discussion of the gap conductance model has been provided in References 83 and 84. The temperature drop across the gap is calculated by assuming an annular gap conductance model of the following form: K gas h=

               +r 2

where: h = contact conductance, Btu/hr-ft2-oF Kgas = thermal conductivity of the gas mixture including a correction factor (Reference 25) for the accommodation coefficient for light gases, e.g., helium, Btu/hr-ft-°F

                =      diametral gap size, ft r        =      effective gap spacing due to surface roughness, ft Chapter 3, Page 24 of 94 Revision 22, 2010

IP2 FSAR UPDATE or an empirical correlation derived from thermocouple and melt radius data. The larger gap conductance value from the equation above and the empirical correlation is used to calculate the temperature drop across the gap for finite gaps. For evaluations in which the pellet-clad gap is closed, a contact conductance is calculated. The contact conductance between UO2 and zircaloy has been measured and found to be dependent on the contact pressure, composition of the gas at the interface, and the surface roughness (References 25 and 26). This information, together with the surface roughness found in Westinghouse fuel, leads to the following correlation: Kgas h = 0.6P + r r = effective gap spacing due to surface roughness, ft h = contact conductance, Btu/hr-ft2-oF P = contact pressure, psi Kgas = thermal conductivity of gas mixture at the interface including a correction factor (Reference 25) for the accommodation coefficient for light gases, e.g., helium, Btu/hr-ft-°F The thermal conductivity of uranium-dioxide was evaluated from data reported by Howard, et al.,27 Lucks, et al.,28 Daniel, et al.,29 Feith,30 Vogt, et al.,31 Nishijima, et al.,32 Ainscough, et al.,33 Godfrey, et al.,34 Stora, et al.,35 Bush,36 Asamoto, et al.,37 Kruger,38 and Gyllander.39 At the higher temperatures, thermal conductivity is best obtained by utilizing the integral conductivity to melt, which can be determined with more certainty. From an examination of the 2800°C data, it has been concluded that the best estimate for the value of Kdt is 93 watts/cm. This 0 conclusion is based on the integral values reported by Gyllander,39 Lyons, et al.,40 Duncan,41 Bain,42 and Stora.43 The design curve for the thermal conductivity is shown in Figure 3.2-38. The section of the curve at temperatures between 0oC and 1300oC is in excellent agreement with the recommendation of the IAEA panel.44 The section of the curve above 1300oC is derived for an integral value of 93 watts/cm.39,41,43 Thermal conductivity for UO2 at 95-percent theoretical density can be represented best by the following equation: 1 K = + 8.775 x 10-13 T3 11.8 + 0.0238T where: K = watt/cm°C T = °C Chapter 3, Page 25 of 94 Revision 22, 2010

IP2 FSAR UPDATE Based upon the above considerations, the maximum central temperature of the hot pellet at steady state is shown in Table 3.2-6. This temperature is well below the melting temperatures of the irradiated UO2, which is taken as 5080oF (Reference 45) unirradiated and decreasing by 58oF per 10,000 MWd/metric ton uranium. Westinghouse experience with fuel rods operating at high power ratings has been summarized in Appendix A, Indian Point No. 2 Preliminary Safety Analysis Report (Docket 50-247) and in Appendix-Section IX of the Preliminary Safeguards Report for the Saxton Reactor Operating at 35 MWt (Docket 50-146). These reports presented considerable statistical evidence of successful operation of high performance zircaloy clad fuel rods in CVTR (1368 rods) and Shipping port core I Blanket (94,920 rods). Since the date of these reports, a significant amount of additional information has been developed relating to the integrity of free standing zircaloy-clad oxide fuel rods at high power ratings. In addition, a comprehensive experimental program was initiated to extend the operating experience to higher power and to higher exposures for many of these fuel rods. This information is summarized in Figure 3.2-39. More detailed information about Westinghouse experience with high power fuel rod bars has been provided in Reference 46. 3.2.2.1.2 Heat Flux Ratio and Data Correlation Departure from nucleate boiling (DNB) is predicted upon a combination of hydrodynamic and heat transfer phenomena and is affected by the local and upstream conditions including the flux distribution. In reactor design, the heat flux associated with DNB and the location of DNB are both important. The W-3 based L-grid DNB correlation was used in design of the LOPAR fuel assemblies. The WRB-1 DNB correlation, Reference 76, is the primary DNB correlation for the analysis of the optimized and VANTAGE+ fuel assemblies. The W-3 DNB correlation47 was developed to predict the DNB flux and the location of DNB equally well for uniform and an axially nonuniform heat flux distribution. This correlation replaced the WAPD q" and H DNB correlations published in Nucleonics,48 May 1963, in order to eliminate the discontinuity of the latter at the saturation temperature and to provide a single unambiguous criterion of the design margin. The W-3 correlation, and several modifications of it, have been used in Westinghouse critical heat flux (CHF) calculations. The W-3 correlation was originally developed from single tube data,49 but was subsequently modified to apply to the "L" grid50 rod bundle data. These modifications to the W-3 correlation have been demonstrated to be adequate for reactor rod bundle design. The W-3 DNB correlation47 incorporates both local and system parameters in predicting the local DNB heat flux. This correlation includes the nonuniform flux effect and the upstream effect, which includes inlet enthalpy or length. The local DNB heat flux ratio (defined as the ratio of the DNB heat flux to the local heat flux) is indicative of the contingency available in the local heat flux without reaching DNB. The sources of the data used in developing the correlation were: WAPD-188 (1958) CU-TR-No. l (NW-208) (1964) ASME Paper 62-WA-297 (1962) CISE-R-90 (1964) CISE-R-63 (1962) DP-895 (1964) ANL-6675 (1962) AEEW-R-356 (1964) Chapter 3, Page 26 of 94 Revision 22, 2010

IP2 FSAR UPDATE GEAP-3766 (1962) BAW-3238-7 (1965) AEEW-R-213 and 309 (1963) AE-RTL-778 (1965) CISE-R-74 (1963) AEEW-355 (1965) CU-MPR-XIII (1963) EUR-2490.e (1965) The comparison of the W-3 measured to predicted DNB flux of this correlation is given in Figure 3.2-40. The local flux DNB ratio versus the probability of not reaching DNB is plotted in Figure 3.2-41. This plot indicates that with a DNBR of 1.3 the probability of not reaching DNB is 95-percent at a 95-percent confidence level. The comparison of the "L"-grid measured to predicted DNB flux is given in Figure 3.2-42. Rod bundle data without mixing vanes agree very well with the predicted DNB flux as shown in Figure 3.2-43. Rod bundle data with mixing vanes (Figure 3.2-44) show on the average an 8-percent higher value of DNB heat flux than predicted by the W-3 DNB correlation. The L-grid modified spacer factor has been formulated to reflect the improvement in DNB heat flux due to the presence of mixing vanes. It should be emphasized that the inlet subcooling effect of the W-3 correlation was obtained from both uniform and non-uniform data. The existence of an inlet subcooling effect has been demonstrated to be real and hence the actual subcooling should be used in the calculations. The W-3 correlation was developed from tests with flow in tubes and rectangular channels. Good agreement was obtained when the correlation is applied to test data for rod bundles. 3.2.2.1.3 Definition of Departure From Nucleate Boiling Ratio The DNB heat flux ratio (DNBR) as applied to the design when all flow cell walls are heated is: q" DNB,N x F 'S x 0.986 DNBR = (for LOPAR Fuel) q" loc The DNB heat flux ratio (DNBR) as applied to typical cells (flow cells with all walls heated) and thimble cells (flow cells with heated and unheated walls) is defined as: q"DNB,N DNBR= q"loc where: q" DNB,EU q" DNB,N = F and q"DNB,EU is the uniform DNB heat flux as predicted by the WRB-1 DNB Correlation and the W-3 DNB correlation (Reference 47) when all flow cells are heated. The flux shape factor to account for nonuniform axial heat flux distributions is F (Reference 47) with the "C" term modified as in Reference 49. Fs is the modified spacer factor, which uses an axial grid spacing coefficient, KS, and a thermal diffusion coefficient, TDC, based on the 20-in. grid spacing data. q"loc is the actual local heat flux. Chapter 3, Page 27 of 94 Revision 22, 2010

IP2 FSAR UPDATE The DNBR as applied to the W-3 DNB correlation when a cold wall is present is: q" DNB, N, CW DNBR = q"loc where: q"DNB,EU,Dh x CWF q"DNB,N, CW = F q"DNB,EU,Dh is the uniform DNB heat flux as predicted by the W-3 cold wall DNB correlation (Reference 49) when not all flow cell walls are heated (thimble cold wall cell). W-3 Cold Wall Factor: q" coldwall G -0.0535 CWF= =1.0- R u {13.76 1.372e 1.78

                                                                     -4.732(      )

q" w 3,Dh 10 6 P 0.14

                - 0.0619(           )      - 8.509 Dh0.107]

1000 where: Ru = (Dh - De / Dh), q"W - 3, Dh = prediction of W - 3 correlation using Dh for De The equivalent uniform DNB flux q'DNB,EU is calculated for the W-3 equivalent uniform flux DNB correlation as follows: q"DNB,EU 6 [

         = (2.022 - 0.0004302p ) + (0.1722 - 0.0000984p ) e(18.177-0.004129)          ]

10 G x 1.037 + 6 (0.1484 - 1.596 + 0.1729 ) x [ 1.157 - 0.869] 10 x[0.2664 + 0.8357e-3.151De] x [0.8258 + 0.000794 ( H sat - H in )] The heat flux is in Btu/hr-ft2 and the units of the parameters are as listed below. The ranges of parameters for applicability of the W-3 DNB correlation are: System pressure, p = 500 to 2400 psia Mass velocity, G = 1.0 x 106 to 5.0 x 106 lb/hr ft2 Equivalent diameter, De = 0.2 to 0.7-in. Quality, loc = -0.15 to +0.15 Inlet enthalpy, no limit, Btu/lb Length, L = 10 to 168-in. Heated perimeter

                    = 0.88 to 1.00 Wetted perimeter Geometries - circular tubes and rectangular channels Chapter 3, Page 28 of 94 Revision 22, 2010

IP2 FSAR UPDATE Flux = Uniform and non-uniform heat flux converted from non-uniform data by using F-factor of Reference 47. For the LOPAR fuel, which uses the "L"-grid modified W-3 DNB correlation with the modified spacer factor (Reference 50):

                                          .5 G  TDC  .35 F  S = (1.445 -.0371L)

P 225.896 ( 2

                                                                   )

e( + .2 ) -.73 + K S 10 6 .019 Where: KS = spacer factor dependent on grid type and axial spacing P = inlet pressure, psia G = inlet mass velocity, lbm/hr-ft2 TDC = thermal diffusion coefficient L = test length, ft

                         = local quality, fraction The ranges of the parameters listed above are:

1460 < P < 2430 psia 1.96 x 106 < G < 3.68 x 106 lbm/hr-ft2

       -0.15 <  < +0.15 fraction 8 < L < 14-ft Geometries ("L"-grid correlation) - circular tubes and rectangular channels Flux = Various nonuniform heat fluxes Local Nonuniform DNB Flux The local nonuniform q"DNB,N is calculated as follows:

q"DNB,EU q"DNB,N = F where: l DNB C C q(z )e l DNB Z F= Cl DNB dz at l DNB x(1 e qlocal ) 0 lDNB = distance from the inception of local boiling to the point of DNB, in inches. Z = distance from the inception of local boiling measured in the direction of flow, in inches. Chapter 3, Page 29 of 94 Revision 22, 2010

IP2 FSAR UPDATE 4.31 (1 - DNB) -1 C = 0.15 0.478 in. (Reference 49) (G/ 106) Where: G = mass velocity, lb/hr-ft2 DNB = quality of the coolant at the location where DNB flux is calculated In determining the F-factor, the value of q"local at lDNB in the above equation for the F-factor was measured as z = l DNB, the location where the DNB flux is calculated. For a uniform flux, F becomes unity so that q"DNB,N reduces to q"DNB,EU as expected. The criterion for determining the predicted location of DNB is to evaluate the ratio of the predicted DNB flux to the local heat flux along the length of the channel. The location of the minimum DNB ratio is considered to be location of DNB. 3.2.2.1.4 Procedure for Using W-3 L-grid Correlation In predicting the local DNB flux in a nonuniform heat flux channel, the following two steps are required:

1. The uniform DNB heat flux, q"DNB,EU, is computed with the W-3 L-grid correlation using the specified local reactor conditions.

2. This equivalent uniform heat flux is converted into corresponding nonuniform DNB heat flux, q"DNB,N, for the nonuniform flux distribution in the reactor. This is accomplished by dividing the uniform DNB flux by the F-factor.47 Since F is generally greater than unity, q"DNB,N will be smaller than q"DNB,EU.

To calculate the DNBR of a reactor channel, the values of q" DNB,N q" loc along the channel are evaluated and the minimum value is selected as the minimum DNBR incurred in that channel. The W-3 L-grid correlation depends on both local and inlet enthalpies of the actual system fluid, and the upstream conditions are accommodated by the F-factor. Hence, the correlation provides a realistic evaluation of the safety margin on heat flux. 3.2.2.1.5 The WRB-1 DNB Correlation The WRB-176 correlation was developed based exclusively on the large bank of mixing vane grid rod bundle critical heat flux data (in excess of 1100 points) that Westinghouse has collected. The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over a wide range of variables than the previous correlation used in design (namely the W-3 correlation). This correlation accounts directly for both typical and thimble cold wall cell effects, uniform and non-uniform heat flux profiles, and variations in rod heated length and in grid spacing. Chapter 3, Page 30 of 94 Revision 22, 2010

IP2 FSAR UPDATE The applicable range of variables is: Pressure 1440 < P < 2490 psia Local Mass Velocity 0.9 < Gloc/106 < 3.7 lb/ft2-hr Local Quality -0.2 < loc < 0.3 Heated Length, Inlet to CHF Location Lh < 14 feet Grid Spacing 13 < gsp < 32 inches Equivalent Hydraulic Diameter 0.37 < de < 0.60 inches Equivalent Heated Hydraulic Diameter 0.46 < dh < 0.68 inches Figure 3.2-44A shows measured critical heat flux plotted against predicted critical heat flux using the WRB-1 correlation. A correlation limit DNBR of 1.17 for the WRB-1 correlation has been approved by the NRC for the 15x15 fuel. 3.2.2.1.6 The W-3 DNB Correlation The W-3 DNB correlation47, 49 and 95 is used for both fuel types where the primary DNB correlation is not applicable. The WRB-1 correlation is developed based on mixing vane data and therefore is only applicable in the heated rod spans above the first mixing vane grid. The W-3 correlation, which does not take credit for mixing vane grids, is used to calculate DNBR values in the heated region below the first mixing vane grid. In addition, the W-3 correlation is applied in the analysis of accident conditions where the system pressure is below the range of the primary correlation. For system pressure in the range of 500 to 1000 psia, the W-3 correlation is 1.45 89. For system pressure greater that 1000 psia, the W-3 correlation is limited to 1.30. A cold wall factor50 is applied to the W-3 DNB correlation to account for the pressure of the unheated thermal surface. 3.2.2.1.7 Film Boiling Heat Transfer Coefficient Heat transfer after departure from nucleate boiling was conservatively assumed to be limited by film boiling immediately, and the period of transition boiling neglected. The correlation used to evaluate these film boiling heat transfer coefficients was developed by Tong, Sandberg and Bishop51 and is shown in Figure 3.2-45. 0.80 1.23 0.68 0.068 hD DG Cp g g

             = 0.0193 k f                f      k f    b        l where: b = g + l(l - ) and CP = heat capacity at constant pressure, Btu/lb-oF D = equivalent diameter of flow channel, ft h = heat transfer coefficient, Btu/hr-ft2-oF G = mass flow rate, lb/hr-ft2 k = thermal conductivity, Btu/hr-ft-oF
   = void fraction
   = density, lbs/ft3
   = viscosity, lbs/ft-hr Chapter 3, Page 31 of 94 Revision 22, 2010

IP2 FSAR UPDATE Subscripts: g = Evaluation of the property at the saturated vapor condition l = Evaluation of the property at the saturated liquid condition f = Evaluation of the property at the average film temperature w = Evaluation of the property at the wall temperature b = Evaluation of the property at the average bulk fluid condition The heat transfer correlation was developed for flow rates equal or greater than 0.8 x 106 lb/hr ft2 over a pressure range of 580 to 3190 psia, for qualities as high as 100-percent, and heat flux from 0.1 to 0.65 x 106 Btu/hr ft2. 3.2.2.2 Hot Channel Factors The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quantities. The heat flux factor considers the local maximum linear heat generation rate at a point (the "hot spot"); the enthalpy rise hot channel factor involves the maximum integrated value along a channel (the "hot channel"). 3.2.2.2.1 Definition of Engineering Hot Channel Factor Each of the total hot channel factors is composed of a nuclear hot channel factor describing the neutron flux distribution and an engineering hot channel factor, which allows for variations in flow conditions and fabrication tolerances. The engineering hot channel factors are made up of subfactors, which account for the influence of the variations of fuel pellet diameter, density, enrichment and eccentricity; fuel rod diameter; inlet flow distribution; flow redistribution; and flow mixing. 3.2.2.2.2 Heat Flux Engineering Subfactor, F EQ The heat flux engineering hot channel factor is used to evaluate the maximum linear heat generation rate in the core. This subfactor is determined by statistically combining the fabrication variations for fuel pellet diameter, density, enrichment and variation in fuel rod diameter, and has a value of 1.03 to be applied to the fuel rod surface heat flux. As shown in Reference 52, no DNB penalty need be taken for the short relatively low intensity heat flux spikes caused by variations in the above parameters or fuel pellet eccentricity 3.2.2.2.3 Enthalpy Rise Engineering Subfactor, F EH The effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise in reload analysis is directly considered in the Westinghouse version of VIPRE-01 Code 100,101 thermal subchannel analysis under any reactor operating condition (refer to Section 3.2.2.4, DNB Analysis Method for VIPRE description). The items presently considered contributing to the enthalpy rise engineering hot channel factor are discussed below:

1. Pellet diameter, density and enrichment Variation in pellet diameter, density and enrichment are considered, statistically, in establishing the limit DNBR's (see section 3.2.2.4) for the Revised Thermal Chapter 3, Page 32 of 94 Revision 22, 2010

IP2 FSAR UPDATE Design Procedure97 employed in this application. Uncertainties in the variables are determined from sampling of manufacturing data.

2. Inlet Flow Maldistribution Data have been considered from several one-seventh scale hydraulic reactor model tests (References 53, 54, and 55) in arriving at the core inlet flow maldistribution criteria to be used in the THINC56 analyses. THINC analyses made, using these data, have indicated that a conservative design basis is to consider a 5-percent reduction in the flow to the hot assembly (Reference 78).

The design basis of 5% flow reduction to the hot assembly is also used in the VIPRE analysis for the 1.4% power uprate.

3. Flow Redistribution The flow redistribution accounts for the reduction in flow in the hot channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect of the nonuniform power distribution is inherently considered in the VIPRE analysis for every operating condition, which is evaluated.

4. Flow Mixing The mixing vanes incorporated in the spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly as well as between adjacent assemblies. This mixing reduces the enthalpy rise in the hot channel resulting from local power peaking or unfavorable mechanical tolerances. The subchannel mixing model now incorporated in the VIPRE code and used in reload reactor design is based on experimental data (Reference 57).

3.2.2.3 Core Pressure Drop and Hydraulic Loads Core and vessel pressure losses are calculated by equations of the form: L V2 PL = K + F D e 2 gc (144) where: PL = unrecoverable pressure drop, lbf/in.2

               = fluid density, lbm/ft3 L        = length, ft De       = equivalent diameter, ft V        = fluid velocity, ft/sec gc       = 32.174, lbm-ft/lbf-sec2 K        = form loss coefficient, dimensionless F        = friction loss coefficient, dimensionless Fluid density is assumed to be constant at the appropriate value for each component in the core and vessel. Because of the complex core and vessel flow geometry, precise analytical values for the form and friction loss coefficients are not available. Therefore, experimental values for these coefficients are obtained from geometrically similar models.

Chapter 3, Page 33 of 94 Revision 22, 2010

IP2 FSAR UPDATE The results of full scale tests of core components and fuel assemblies are utilized in developing the core pressure loss characteristic in reload reactor design. The pressure drop for the vessel has been obtained by combining the core loss with correlation of one-seventh scale model hydraulic test data on a number of vessels (References 53 and 54) and form loss relationships (Reference 58). Moody (Reference 59) curves have been used to obtain the single phase friction factors. The fuel assembly hold-down springs are designed to keep the fuel assemblies in contact with the lower core plate under all Condition I and II events, with the exception of the turbine overspeed transient associated with a loss of external load. The hold-down springs are designed to tolerate the possibility of an over deflection associated with fuel assembly liftoff for this case and provide contact between this transient. More adverse flow conditions occur during a loss-of-coolant accident. Hydraulic loads at normal operating conditions are calculated considering the best estimate flow and accounting for the minimum core bypass flow based on manufacturing tolerances. Core hydraulic loads at cold plant startup conditions are based on the best estimate flow, but are adjusted to account for the coolant density difference. Conservative core hydraulic loads for a pump overspeed transient, which could possibly create flow rates 20-percent greater than the best estimate flow, are evaluated to be approximately twice the fuel assembly weight. The hydraulic forces are not sufficient to lift a rod control cluster during normal operation even if the rod cluster is detached from its coupling. 3.2.2.4 Thermal and Hydraulic Design Parameters The thermal and hydraulic design parameters are given in Table 3.2-6, Sheets 1-3. Sheet 3 shows parameters over a range of vessel average temperatures, giving flexibility to operate at full licensed power at various plant operating conditions. DNB Design Basis There will be at least a 95-percent probability that departure from nucleate boiling (DNB) will not occur on the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Condition I and II events), at a 95-percent confidence level. Historically, this has been conservatively met by adhering to the following thermal design basis: there must be at least a 95-percent probability that the minimum departure from nucleate boiling ratio (DNBR) of the limiting power rod during Condition I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The DNBR limit for the correlation is established based on the variance of the correlation such that there is a 95-percent probability with 95-percent confidence that DNB will not occur when the calculated DNBR is at the DNBR limit. DNB Analysis Method The THINC IV 77, 78 computer program, beginning in Cycle 10, was used to perform the thermal/hydraulic calculations for both fuel types. The THINC IV code is used to calculate coolant density, mass velocity, enthalpy, void fractions, static pressure and DNBR distributions along flow channels within a reactor core under all expected operating conditions. References 77 and 78 contain details of the THINC IV computer program, including models and correlations used. The Westinghouse version of the VIPRE-01 (VIPRE) code is used. The VIPRE code is equivalent to the THIC-VI (THINC) code and has been approved by the NRC for licensing applications to replace the THINC code. The use of VIPRE is in full compliance with the Chapter 3, Page 34 of 94 Revision 22, 2010

IP2 FSAR UPDATE conditions specified in the NRC Safety Evaluation Report (SER) on WCAP-14565-P-A (Reference 101). The design method employed for both fuel types to meet the DNB design basis is the Revised Thermal Design Procedure.97 Uncertainties in plant operating parameters, nuclear and thermal parameters, and fuel fabrication parameters are considered statistically such that there is at least a 95-percent probability that the minimum DNBR will be greater than or equal to the limit DNBR for the limiting power rod. Plant parameter uncertainties are used to determine the plant DNBR uncertainty. The DNBR uncertainty, combined with the DNBR limit, establishes a design DNBR value, which must be met in plant safety analyses. Since the parameter uncertainties are considered in determining the design DNBR value, the plant safety analyses are performed using values of input parameters without uncertainties. In addition, the limit DNBR values are increased to values designated as the safety analysis limit DNBR's. The plant allowances available between the safety analysis limit DNBR values and the design limit DNBR values is not required to meet the design basis. This allowance will be used for flexibility in the design and operation of this plant. For this design, the WRB-1 correlation is used for analysis of the Vantage+ fuel assemblies with a correlation limit of 1.17 (both typical and thimble cells). The design method employed for both fuel types to meet the DNB design basis is the Revised Thermal Design Procedure (RTDP)97. With the RTDP methodology, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes and DNB correlation predictions are considered statistically to obtain DNB uncertainty factors. Based on the DNB uncertainty factors, RTDP design limit DNBR values are determined such that there is at least a 95% probability at a 95% confidence level that DNB will not occur on the most limiting fuel rod during normal operation and operational transients and during transient conditions arising from faults of moderate frequency (Condition I and II events as defined in ANSI N18.2) To maintain DNBR margin to offset DNB penalties such as those due to fuel rod bow (see section 3.2.2.6) and potential transition core (see 3.2.2.5.1), the safety analyses were performed to DNBR limits higher than the design limit DNBR values. The difference between the design limit DNBRs and the safety analysis limit DNBRs results in available DNBR margin. The net DNBR margin, after consideration of all penalties, is available for operating and design flexibility. The Standard Thermal Design Procedure (STDP) is used for those analyses where RTDP is not applicable. In the STDP method the parameters used in analysis are treated in a conservative way from a DNBR standpoint. The parameter uncertainties are applied directly to the plant safety input values to give the lowest minimum DNBR. The DNBR limit for STDP is the appropriate DNB correlation limit increased by sufficient margin to offset the applicable DNBR penalties. For this design, the WRB-1 correlation is used for analysis of both fuel types with a correlation limit of 1.17 (both typical and thimble cells). When the core condition is outside the range of the WRB-1 correlation, the W-3 correlation is applied with a correlation limit of 1.30 (both cell types). DNB With Physical Burnout Westinghouse60 has conducted DNB tests in a 25-rod bundle where physical burnout occurred with one rod. After this occurrence, the 25-rod test section was used for several days to obtain more DNB data from the other rods in the bundle. The burnout and deformation of the rod did not affect the performance of neighboring rods in the test section during the burnout or the Chapter 3, Page 35 of 94 Revision 22, 2010

IP2 FSAR UPDATE validity of the subsequent DNB data points as predicted by the W-3 correlation. No occurrences of flow instability or other abnormal operations were observed. DNB With Return to Nucleate Boiling Additional DNB tests have been conducted by Westinghouse61 in 19 and 21 rod bundles. In these tests, DNB without physical burnout was experienced more than once on single rods in the bundles for short periods of time. Each time, a reduction in power of approximately 10-percent was sufficient to reestablish nucleate boiling on the surface of the rod. During these and subsequent tests, no adverse effects were observed on this rod or any other rod in the bundle as a consequence of operating in DNB. Hydrodynamic and Flow Power Coupled Instability Thermohydrodynamic instabilities will not occur under Condition I and II modes of operation for Westinghouse PWR reactor designs. A large power margin exists to predicted inception of such instabilities. Analysis has been performed which shows that minor plant to plant differences in Westinghouse reactor designs such as fuel assembly arrays, core power to flow ratios and fuel assembly length will not result in gross deterioration of the above power margins. 3.2.2.5 Hydraulic Compatibility 3.2.2.5.1 Transition Core Effects The entire core is now 15x15 Upgraded fuel therefore there are no transition core effects. 3.2.2.5.2 DNB Performance When Transitioning Cores The Westinghouse transition core DNB methodology is given in References 90, 91, and 92 and has been approved by the NRC via Reference 93 and 94. Using this methodology, transition cores are analyzed as if the entire core consists of one assembly type. 3.2.2.5.3 Compatibility The hydraulic resistance of the two assemblies is based on full scale hydraulic flow test data. The design hydraulic loss coefficients are verified with a confirmatory hydraulic test in the Fuel Assembly Compatibility Test System (FACTS). The results are evaluated to determine the values of the pressure drop loss coefficients. Fuel assembly vibration testing is conducted to confirm that the fuel assembly does not experience flow-induced vibration. Hydraulic vibration tests are performed in the FACTS test loop. Side-by-side VIPER tests are performed. Test results are used to demonstrate the acceptable fretting behavior design between the two designs is insignificant. A crossflow analysis using the THINC code is completed to determine the crossflow velocity profile in the transition core in the IP2 plant conditions. The crossflow is caused by the mid-grid and IFM pressure drop mismatch between adjacent fuel assemblies. Chapter 3, Page 36 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2.2.6 Effects of Rod Bow on DNBR The phenomenon of fuel rod bowing, as described in Reference 80 must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatisms in the evaluation of DNBR can be used to offset the effects of rod bow. For safety analysis of Indian Point Unit 2, sufficient margin was maintained in the design of the fuel [Note - Margin maintained between design limit DNBR and safety analysis limit DNBR +As a result of analyses performed for OFA transition, maintaining plugging devices in core is optional.] to accommodate full and low flow rod bow DNBR penalties (less than 1-percent for the worst cast, which is at a burnup of 24,000 MWd/MTU identified in Reference 81) with the incorporation of the L2/I scaling factor (I=bending moment of inertia, L=span length) to account for the 9-grid fuel span lengths. The rod bow DNBR penalties in the Intermediate Flow Mixer (IFM) grid spans are less than those in the mixing vane grid spans. The maximum rod bow penalties accounted for in the design safety analysis are based on an assembly average burnup of 24,000 MWd/MTU. At burnups greater than 24,000 MWd/MTU, credit is taken for the effect of F-delta-H burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory, and no additional rod bow penalty is required. 3.2.3 Mechanical Design And Evaluation The reactor core and reactor vessel internals are shown in cross-section in Figure 3.2-46 and in elevation in Figure 3.2-47. The core, consisting of the fuel assemblies, control rods, source rods, burnable poison rods, and plugging devices+, provides and controls the heat source for the reactor operation. The internals, consisting of the upper and lower core support structure, are designed to support, align, and guide the core components, direct the coolant flow to and from the core components, and to support and guide the incore instrumentation. A listing of the core mechanical design parameters is given in Table 3.2-7. The fuel assemblies are arranged in a checkerboard and/or roughly circular zoned pattern. The fuel assemblies contain fuel of different enrichments depending on the location of the assembly within the core. The fuel is in the form of slightly enriched uranium-dioxide ceramic pellets. The pellets are stacked to an active height of 144-in. (previously 142 in.) within ZIRLOTM (previously Zircaloy-4) tubular cladding, which is plugged and seal welded at the ends to encapsulate the fuel. The enrichments of the fuel for the first three regions in the core are given in Table 3.2-7. Reload fuel enrichment may vary up to the maximum value allowed in the Technical Specifications. Heat generated by the fuel is removed by demineralized light water, which flows upward through the fuel assemblies and acts as both moderator and coolant. The core is divided into fuel assembly regions of different enrichments. The loading arrangement for the initial cycle is indicated on Figure 3.2-48. In the past refueling took place generally in accordance with an inward loading schedule. Starting from Cycle 6 a low leakage loading pattern for core refueling design has been adopted and starting from Cycle 13, a low, low leakage loading pattern was used. This will reduce neutron fluence at the reactor vessel wall. Chapter 3, Page 37 of 94 Revision 22, 2010

IP2 FSAR UPDATE The control rods, designated as rod cluster control assemblies, consist of groups of individual absorber rods, which are held together by a spider assembly at the top end and actuated as a group. In the inserted position, the absorber rods fit within hollow guide thimbles in the fuel assemblies. The guide thimbles are an integral part of the fuel assemblies and occupy locations within the regular fuel rod pattern where fuel rods have been deleted. In the withdrawn position, the absorber rods are guided and supported laterally by guide tubes, which form an integral part of the upper core support structure. Figures 3.2-49 and 3.2-50 show a typical rod cluster control assembly in a fuel assembly. As shown in Figure 3.2-47, the fuel assemblies are positioned and supported vertically in the core between the upper and lower core plates. The core plates are provided with pins, which index into closely fitting mating holes in the fuel assembly top and bottom nozzles. The pins maintain the fuel assembly alignment, which permits free movement of the control rods from the fuel assembly into the guide tubes in the upper support structure without binding or restriction between the rods and their guide surfaces. Operational or seismic loads imposed on the fuel assemblies are transmitted through the core plates to the upper and lower support structures and ultimately to the internals support ledge at the pressure vessel flange in the case of vertical loads or to the lower radial support and internals support ledge in the case of horizontal loads. The internals also provide a form fitting baffle surrounding the fuel assemblies, which confine the upward flow of coolant in the core area to the fuel bearing region. 3.2.3.1 Reactor Internals 3.2.3.1.1 Design Description The reactor internals are designed to support and orient the reactor core fuel assemblies and control rod assemblies, absorb the control rod dynamic loads, and transmit these and other loads to the reactor vessel flange, provide a passageway for the reactor coolant, and support incore instrumentation. The reactor internals are shown in Figure 3.2-47. The internals have been designed to withstand the forces due to weight, preload of fuel assemblies, control rod dynamic loading, vibration, and earthquake acceleration. The internals were analyzed in a manner similar to Connecticut Yankee, San Onofre, Zorita, Saxton, and Yankee. Under the loading conditions, including conservative effects of design earthquake loading, the structure satisfies stress values prescribed in Section III, ASME Nuclear Vessel Code. The reactor internals are equipped with bottom-mounted incore instrumentation supports. These supports are designed to sustain the applicable loads outlined above. The components of the reactor internals are divided into three parts consisting of the lower core support structure (including the entire core barrel and thermal shield), the upper core support structure, and the incore instrumentation support structure. 3.2.3.1.1.1 Lower Core Support Structure The major component and support member of the reactor internals is the lower core support structure, shown in Figure 3.2-51. This support structure assembly consists of the core barrel, the core baffle, the lower core plate and support columns, the thermal shield, the intermediate diffuser plate, and the bottom support plate, which is welded to the core barrel. All the major Chapter 3, Page 38 of 94 Revision 22, 2010

IP2 FSAR UPDATE material for this structure is type 304 stainless steel. The core support structure is supported at its upper flange from a ledge in the reactor vessel head flange and its lower end is restrained in its transverse movement by a radial support system attached to the vessel wall. Within the core barrel are axial baffle and former plates, which are attached to the core barrel wall and form the enclosure periphery of the assembled core. The lower core plate is positioned at the bottom level of the core below the baffle plates and provides support and orientation for the fuel assemblies. The lower core plate is a 2-in. thick member through which the necessary flow distributor holes for each fuel assembly are machined. Fuel assembly locating pins (two for each assembly) are also inserted into this plate. Columns are placed between this plate and the lower core support of the core barrel in order to provide stiffness and to transmit the core load to the lower core support. Intermediate between the support plate and lower core support plate a perforated plate is positioned to diffuse uniformly the coolant flowing into the core. The one-piece thermal shield is fixed to the core barrel at the top with rigid bolted connections. The bottom of the thermal shield is connected to the core barrel by means of axial flexures. This bottom support allows for differential axial growth of the shield with respect to the core barrel but restricts radial or horizontal movement of the bottom of the shield. Rectangular tubing, in which vessel material samples can be inserted and irradiated during reactor operation, are welded to the thermal shield and extend to the top of the thermal shield. These samples are held in the rectangular tubing by a preloaded spring device at the top and bottom. The lower core support structure and principally the core barrel serve to provide passageways and control for the coolant flow. Inlet coolant flow from the vessel inlet nozzles proceeds down the annulus between the core barrel and the vessel wall, flows on both sides of the thermal shield, and then into a plenum at the bottom of the vessel. It then turns and flows up through the lower support, passes through the intermediate diffuser plate and then through the lower core plate. The flow holes in the diffuser plate and the lower core are arranged to give a very uniform entrance flow distribution to the core. After passing through the core, the coolant enters the area of the upper support structure and then flows generally radially to the core barrel outlet nozzles and directly through the vessel outlet nozzles. A small amount of water also flows between the baffle plates and core barrel to provide additional cooling of the barrel. Similarly, a small amount of the entering flow is directed into the vessel head plenum to provide cooling of the head. Both these flows eventually are directed into the upper support structure plenum and exit through the vessel outlet nozzles. Vertically downward loads from weight, fuel assembly preload, control rod dynamic loading, and earthquake acceleration are carried by the lower core plate partially into the lower core plate support flange on the core barrel shell and partially through the lower support columns to the lower core support and then through the core barrel shell to the core barrel flange supported by the vessel head flange. Transverse loads from earthquake acceleration, coolant cross flow, and vibration are carried by the core barrel shell to be distributed to the lower radial support to the vessel wall, and to the core barrel flange. Transverse acceleration of the fuel assemblies is transmitted to the core barrel shell by direct connection of the lower core plate to the barrel wall and by a radial support type connection of the upper core plate to slab-sided pins pressed into the core barrel. The main radial support system of the core barrel is accomplished by "key" and "keyway" joints to the reactor vessel wall. At equally spaced points around the circumference, an Inconel block Chapter 3, Page 39 of 94 Revision 22, 2010

IP2 FSAR UPDATE is welded to the vessel inside diameter. Another Inconel block is bolted to each of these blocks, and has a "keyway" geometry. Opposite each of these is a "key", which is attached to the internals. At assembly, as the internals are lowered into the vessel, the keys engage the keyways in the axial direction. With this design, the internals are provided with a support at the furthest extremity, and may be viewed as a beam fixed at the top and simply supported at the bottom. Radial and axial expansions of the core barrel are accommodated but transverse movement of the core barrel is restricted by this design. With this system, cycle stresses in the internal structures are within the ASME Section III limits. This eliminates any possibility of failure of the core support. In the event of downward vertical displacement of the internals, energy absorbing devices limit the displacement by contacting the vessel bottom head. The load is transferred through the energy devices of the internals. The energy absorbers, cylindrical in shape, are contoured on their bottom surface to the reactor vessel bottom head geometry. Their number and design are determined so as to limit the forces imposed to less than yield. Assuming a downward vertical displacement, the potential energy of the system is absorbed mostly by the strain energy of the energy absorbing devices. The free fall in the hot condition is on the order of 0.50-in. and there is an additional strain displacement in the energy absorbing devices of approximately 0.75-in. Alignment features in the internals prevent cocking of the internals structure during this postulated drop. The control rods are designed to provide assurance of control rod insertion capabilities under this assumed drop of internals condition. The drop distance of about 1.25-in. is not enough to cause the tips of the shutdown group of rod cluster control assemblies to come out of the guide tubes in the fuel assemblies. 3.2.3.1.1.2 Upper Core Support Assembly The upper core support assembly, shown in Figure 3.2-52, consists of the top support plate, deep beam sections, and upper core plate between which are contained 48 support columns and 61 guide tube assemblies. The support columns establish the spacing between the top support plate, deep beam sections, and the upper core plate. They are fastened at top and bottom to these plates and beams. The support columns transmit the mechanical loadings between the two plates and serve the supplementary function of supporting thermocouple guide tubes. The guide tube assemblies, shown on Figure 3.2-53, sheath and guide the control rod drive shafts and control rods and provide no other mechanical functions. They are fastened to the top support plate and are guided by pins in the upper core plate for proper orientation and support. Additional guidance for the control rod drive shafts is provided by the control rod shroud tube, which is attached to the upper support plate and guide tube. The upper core support assembly, which is removed as a unit during refueling operation, is positioned in its proper orientation with respect to the lower support structure by flat-sided pins pressed into the core barrel, which in turn engage in slots in the upper core plate. At an elevation in the core barrel where the upper core plate is positioned, the flat-sided pins are located at angular positions of 0, 90, 180, and 270 degrees. Four slots are milled into the core plate at the same positions. As the upper support structure is lowered into the main internals, the slots in the plate engage the flat-sided pins in the axial direction. Lateral displacement of the plate and of the upper support assembly is restricted by this design. Fuel assembly locating pins Chapter 3, Page 40 of 94 Revision 22, 2010

IP2 FSAR UPDATE pins protrude from the bottom of the upper core plate and engage the fuel assemblies as the upper assembly is lowered into place. Proper alignment of the lower core support structure, the upper core support assembly, the fuel assemblies and control rods is thereby assured by this system of locating pins and guidance arrangement. The upper core support assembly is restrained from any axial movements by a large circumferential spring, which rests between the upper barrel flange and the upper core support assembly and is compressed by the reactor vessel head flange. Vertical loads from weight, earthquake acceleration, hydraulic loads, and fuel assembly preload are transmitted through the upper core plate via the support columns to the deep beams and top support plate and then to the reactor vessel head. Transverse loads from coolant cross flow, earthquake acceleration, and possible vibrations are distributed by the support columns to the top support plate and upper core plate. The top support plate is particularly stiff to minimize deflection. 3.2.3.1.1.3 Incore Instrumentation Support Structures The incore instrumentation support structures consist of an upper system to convey and support thermocouples penetrating the vessel through the head and a lower system to convey and support flux thimbles penetrating the vessel through the bottom. The upper system utilizes the reactor vessel head penetrations. Instrumentation port columns are slip-connected to inline columns that are in turn fastened to the upper support plate. These port columns protrude through the head penetrations. The thermocouples are carried through these port columns and the upper support plate at positions above their readout locations. The thermocouple conduits are supported from the columns of the upper core support system. The thermocouple conduits are sealed stainless steel tubes. In addition to the upper incore instrumentation, there are reactor vessel bottom port columns, which carry the retractable, cold-worked stainless steel flux thimbles that are pushed upward into the reactor core. Conduits extend from the bottom of the reactor vessel down through the concrete shield area and up to a thimble seal line. The minimum bend radii are about 144-in. and the trailing ends of the thimbles (at the seal line) are extracted approximately 15-ft during refueling of the reactor in order to avoid interference within the core. The thimbles are closed at the leading ends and serve as the pressure barrier between the reactor pressurized water and the containment atmosphere. Mechanical seals between the retractable thimbles and the conduits are provided at the seal line. During normal operation, the retractable thimbles are stationary and move only during refueling or for maintenance, at which time a space of approximately 15-ft above the seal line is cleared for the retraction operation. Sections 7.4 and 7.6 contain more information on the layout of the incore instrumentation system. The incore instrumentation support structure is designed for adequate support of instrumentation during reactor operation and is rugged enough to resist damage or distortion under the conditions imposed by handling during the refueling sequence. 3.2.3.1.2 Evaluation of Core Barrel and Thermal Shield The internals design is based on analysis, test, and operational information. Troubles in previous Westinghouse PWRs have been evaluated and information derived has been considered in this design. For example, the Westinghouse design uses a one-piece thermal Chapter 3, Page 41 of 94 Revision 22, 2010

IP2 FSAR UPDATE shield, which is attached rigidly to the core barrel at one end and flexured at the other. The early designs that malfunctioned were multi-piece thermal shields that rested on vessel lugs and were not rigidly attached at the top. Early core barrel designs that have malfunctioned in service, now abandoned, employed threaded connections such as tie rods, joining the bottom support to the bottom of the core barrel, and a bolted connection that tied the core barrel to the upper barrel. The malfunctioning of core barrel designs in earlier service was believed to have been caused by the thermal shield, which was oscillating, thus creating forces on the core barrel. Other forces were induced by unbalanced flow in the lower plenum of the reactor. In today's rod cluster control design there are no fuel followers to necessitate a large bottom plenum in the reactor. The elimination of these fuel followers enabled Westinghouse to build a shorter core barrel. The Connecticut Yankee reactor and the Zorita reactor core barrels are of the same construction as the Indian Point Unit 2 reactor core barrel. Deflection measuring devices employed in the Connecticut Yankee reactor during the hot functional test, and deflection and strain gauges employed in the Zorita reactor during the hot-functional test provided important information that was used in the design of the present day internals, including that for Indian Point. When the Connecticut Yankee thermal shield was modified to the same design as for Southern California Edison, it, too, operated satisfactorily as was evidenced by the examination after the hot-functional test. After hot-functional tests on all of these reactors, a careful inspection of the internals was examined for any differential movement; upper core plate inside supports were examined, and the thermal shield attachments to the core barrel, including all lockwelds on the devices used to lock the bolt, were checked; no malfunctions were found. Substantial scale model testing was performed at WAPD. This included tests, which involved a complete full-scale fuel assembly, which was operated at reactor flow, temperature, and pressure conditions. Tests were run on a one-seventh scale model of the Indian Point Unit 2 reactor. Measurements taken from those tests indicated very little shield movement, on the order of a few mils when scaled up to Indian Point Unit 2. Strain gauge measurements taken on the core barrel also indicated very low stresses. Testing to determine thermal shield excitation due to inlet flow disturbances was included. Information gathered from these tests was used in the design of the thermal shield and core barrel. It was concluded, from the experience gained during the testing program and the analyses, that the design as employed on the Indian Point Unit 2 plant is adequate. In order to confirm the internals design, deflection gauges were mounted on the thermal shield top and bottom for the hot-functional test. Gauges were mounted in the top of the thermal shield equidistant from the fixed supports, and at the bottom of the thermal shield, equidistant from the six flexures, and next to the flexure supports. The internals inspection, just before the hot-functional test, included looking at mating bearing surfaces, main welds, and welds that are used on bolt locking devices. At the conclusion of the hot-functional test, measurement readings were taken from the deflectometers on the shield and the internals were re-examined at all key areas for any evidence of malfunction. Chapter 3, Page 42 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2.3.2 Core Components 3.2.3.2.1 Design Description 3.2.3.2.1.1 Fuel Assembly The 15x15 Upgraded fuel assembly, introduced in Cycle 17, is shown in Figure 3.2-61C. The assemblies are square in cross section, nominally 8.426-in. on a side, and have an overall height of approximately 159.975 inches. The fuel rods in a fuel assembly are arranged in a square array with 15 rod locations per side and a nominal centerline-to-centerline pitch of 0.563-in. between rods. Of the total possible 225 rod locations per assembly, 20 are occupied by guide thimbles for the rod cluster control rods and one for incore instrumentation. The remaining 204 locations contain fuel rods. In addition to fuel rods, a fuel assembly is composed of a top nozzle, a bottom nozzle, ten grid assemblies (plus 3 intermediate flow mixing grids starting with Cycle 13), 20 absorber rod guide thimbles, and one instrumentation thimble. The guide thimbles in conjunction with the grid assemblies and the top and bottom nozzles comprise the basic structural fuel assembly skeleton. The grid assemblies are bulge attached to the guide thimbles at each location along the height of the fuel assembly at which lateral support for the fuel rods is required. Within this skeletal framework the fuel rods are contained and supported and the rod-to-rod centerline spacing is maintained along the assembly. The original fuel design for Indian Point 2 was the Westinghouse High Parasitic (HIPAR) fuel assembly. This consisted of Zircoloy clad fuel rods, 9 Inconel grids and stainless steel instrumentation and guide thimbles. Burnable absorbers used were pyrex glass. Starting with Cycle 5, the Westinghouse Low Parasitic (LOPAR) fuel assembly was introduced. This design consisted of Zircaloy-4 clad fuel rods, 9 Inconnel grids and Zircaloy-4 instrumentation and guide thimbles. For the Cycle 8, Wet Annular Burnable Absorbers (WABA) were introduced. For Cycle 10, the Westinghouse Optimized Fuel Assembly (OFA) was introduced. This consisted of Zircaloy-4 clad fuel rods, 2 Inconel grids (top & bottom), 7 Zircaloy-4 grids and Zircaloy-4 instrumentation and guide thimbles. In addition, thimble plugs were removed from the core this cycle based on analysis performed to support removal. The assembly top nozzle design was changed to a Reconstitutable Top Nozzle (RTN) design to facilitate reconstitution of failed fuel. For Cycle 11, the OFA fuel assembly design incorporated Debris Filter Bottom Nozzles (DFBN) and Integral Fuel Burnable Adsorbers (IFBA). For Cycle 13, the Westinghouse Vantage+ fuel design was introduced (see Figures 3.2-54 and 3.2-56B). This design included ZIRLO clad fuel rods, 2 Inconel grids, 7 low pressure drop (LPD) Zircaloy-4 grids, 3 Zircaloy Integral Flow Mixing grids (IFM), ZIRLO instrumentation and guide thimbles, annular axial blankets along with the DFBN, IFBA and RTN. Use of WABAs was continued. See Reference 13. For Cycle 15, the Vantage+ fuel assembly design incorporated Performance+ features of ZIRLO grids and IFMs and a hardened coating of zirconium oxide on the bottom section of the fuel rod clad to increase debris resistance. Chapter 3, Page 43 of 94 Revision 22, 2010

IP2 FSAR UPDATE For Cycle 16, the Vantage+ fuel assembly design was further enhanced with Performance+ features that include debris mitigation features of an additional grid located at the bottom end plug of the fuel rod, a longer fuel rod end plug and a revised DFBN. Other performance+ enhancements include longer fuel rods and longer annual axial blanket (see Figure 3.2-61B). In addition to the above fuel design changes, the design burnup of the fuel assemblies has also been increasing to 62,000 MWD/MTU lead rod burnup for Cycle 16. See References 15, 16 and 17. For Cycle 17, the 15x15 Upgraded fuel assembly design was used. This design has features to address grid-to-rod fretting fuel failures. These include I-spring mid-grids, enhanced IFMs and balanced mixing vanes. In addition, the tube-in-tube thimble design was incorporated with a single-dashpot, which improves straightness. For Cycle 18, solid axial blanket pellets were introduced for the non-IFBA fuel rods. For Cycle 19, the top nozzle spring design was changed from the Vantage+ design to the standard spring design. Cycle 20 fuel is the same as Cycle 19, there are no fuel design changes. Bottom Nozzle Two types of nozzle designs were used for the HIPAR fuel assemblies. One type, which is square in cross section, is fabricated from type 304 stainless steel consisting of four side plates, 12 cross bars and four pads or feet. The side plates are welded together at the corners to form a plenum for inlet coolant to the fuel assembly. The cross bars are welded at each end to the top edges of the side plate and function as the bottom end support for the fuel rods. The bottom support surface for the fuel assembly is formed by the four pads, which are welded to the side plates in the corners. This design was used in a majority of the first core fuel assemblies. The previously used LOPAR and OFA fuel incorporate an equivalent bottom nozzle design utilizing a square perforated plate rather than the cross bars and side plate. On both designs, their respective cross bars and perforated plate prevent the fuel rods from falling through the bottom nozzles of the assembly. Coolant flow to the fuel assembly is directed from the plenum in the bottom nozzle upward to the interior of the fuel assembly and to the channel between assemblies. Axial loads imposed on the assembly, as well as the weight of the assembly are distributed through the guide thimble and the bottom nozzle to the lower core support plate. Indexing and positioning of the fuel assembly in the core is controlled through two holes in diagonally opposite pads, which mate with locating pins in the lower core plate. Lateral loads imposed on the fuel assembly are also transferred to the core support structures through the locating pins. The OFA and VANTAGE+ bottom nozzle used the reconstitutable feature found on the previously installed LOPAR fuel design, which uses a locking cup to lock the thimble screws on the guide thimble assembly, instead of the lockwire used in earlier LOPAR designs. The OFA nozzle assembly is shorter when compared to the previously installed LOPAR assembly to enhance fuel rod growth allowances. Chapter 3, Page 44 of 94 Revision 22, 2010

IP2 FSAR UPDATE The two bottom nozzle designs used in the OFAs are both square in cross section and fabricated from 304 stainless steel. The design used in earlier regions consists of a perforated plate, four angle legs, and four pads of feet. The angle legs are fastened to the plate forming a plenum space for the coolant inlet to the fuel assembly. The remaining OFA regions and the VANTAGE+ and 15x15 Upgraded fuel regions (starting with Cycle 13, Region 15) incorporate an equivalent bottom nozzle design denoted as the Debris Filter Bottom Nozzle (DFBN). This nozzle adds side plates or "skirts" to the previous design increasing structural capability for abnormal loads and providing a more defined plenum space below the nozzle. Additionally, the relatively large adapter plate flow holes of the earlier design are replaced with a new pattern of smaller flow holes. The decrease in size of the holes provides a "screen" for larger debris particles, which would otherwise cause damage if allowed to pass into the assembly. In both designs, the adaptor plates prevent accidental downward ejection of the fuel rods from the fuel assembly. The nozzles are fastened to the assembly guide tubes by stainless steel screws, which penetrate through the nozzle and mate with a threaded plug in each guide tube (Figure 3.2-57). The screw possesses a circular locking cup around the screw head, which is crimped into mating detentes (lobes) on the bottom nozzle, preventing the screw from loosening. Top Nozzle The Reconstitutable Top Nozzle (RTN) used in both the OFA, VANTAGE+ and 15x15 Upgraded fuel assemblies is a box-like structure, which functions as the fuel assembly upper structural element and forms a plenum space where the heated fuel assembly discharge coolant is mixed and directed toward the flow holes in the upper core plate. The nozzle is comprised of an adaptor plate enclosure, top plate, clamps, hold-down leaf springs and assorted hardware. Each nozzle has four sets of leaf springs. All parts, with the exception of the springs and their hold-down bolts/screws, are constructed of type 304 stainless steel. The springs are made from age hardenable Inconel 718 and the bolts/screws from Inconel 600 for Region 16 and earlier regions, and from shotpeened Inconel 718 for Regions 17 and 18. The adaptor plate portion of the nozzle is square in cross section, and is perforated by machined slots to provide for coolant flow through the plate. At assembly, the top ends of the LOPAR thimble stainless sleeves are fitted through individual bored holes in the plate and welded to the plate around the circumference of each hole. In the OFA removable top nozzle design, a groove is provided in each thimble thru-hole in the nozzle plate into which a stainless steel nozzle insert is mechanically connected by means of a preformed circumferential bulge near the top of the insert. Thus, the adaptor plate acts as the fuel assembly top end plate, and provides a means of distributing evenly among the guide thimbles any axial loads imposed on the fuel assemblies. The nozzle enclosure is actually a square tubular structure, which forms the plenum section of the top nozzle. The bottom end of the enclosure is pinned and welded to the periphery of the adaptor plate and the top end is welded to the periphery of the top plate. The top plate is square in cross section with a square central hole. The hole allows clearance for the rod cluster control absorber rods to pass through the nozzle into the guide thimbles in the fuel assembly and for coolant exit from the fuel assembly to the upper internals area. Two pads containing axial through-holes, which are located on diametrically opposite corners of the top Chapter 3, Page 45 of 94 Revision 22, 2010

IP2 FSAR UPDATE top plate provide a means of positioning and aligning the top of the fuel assembly. As with the bottom nozzle, alignment pins in the upper core plate mate with the holes in the top nozzle plate. Hold-down forces of sufficient magnitude to oppose the hydraulic lifting forces on the fuel assembly are obtained by means of the leaf spring sets, which are mounted on the top plate. The springs are fastened in pairs to the top plate at the two corners where alignment holes are not used and radiate out from the corners parallel to the sides of the plate. Fastening of each pair of springs is accomplished with a clamp, which fits over the ends of the springs and two bolts/screws (one per spring set), which pass through the clamp and spring, and thread into the top plate. At assembly, the spring mounting bolts/screws are torqued sufficiently to preload against the maximum spring load and then lockwelded to the clamp, which is counterbored to receive the bolt/screw head. The spring load is obtained through deflection of the spring pack by the upper core plate. The spring pack form is such that it projects above the fuel assembly and is depressed by the core plate when the internals are loaded into the reactor. The free end of the spring pack is bent downward and captured in a key slot in the top plate to guard against loose parts in the reactor in the event (however remote) of spring fracture. The capture of the loose end has been deleted in latter designs. Starting with Cycle 14, Region 16, the fuel has a cast top nozzle. This is a two-piece design incorporating a machined stainless steel adapter plate welded to a low-cobalt investment casting. The cast top nozzle is functionally interchangeable with the previous design and meets design criteria for the top nozzle. In addition to its plenum and structural functions, the nozzle provides a protective housing for components, which mate with the fuel assembly. In handling a fuel assembly with a control rod inserted, the control rod spider is contained within the nozzle. During operation in the reactor, the nozzle protects the absorber rods from coolant cross flows in the unsupported span between the fuel assembly adaptor plate and the end of the guide tube in the upper internals package. Plugging devices, [Note - As a result of analyses performed for OFA transition, maintaining plugging devices in the core is optional. ], which fill the ends of the fuel assembly thimble tubes at unrodded core locations and the source rods and burnable absorber rods, are all contained within the fuel top nozzle. For the RTN design, a stainless steel nozzle insert is mechanically connected to the top nozzle adaptor plate (Figure 3.2-58A) via the engagement of the preformed circumferential bulge near the top of the insert and the mating groove in the wall of the adapter plate thimble tube through-hole. The insert has four equally spaced axial slots, which allow the insert to deflect inwardly at the elevation of the bulge, thus permitting the installation and removal of the nozzle. The insert bulge is positively held in the adapter plate mating groove by placing a lock tube with a uniform OD identical to that of the thimble tube into the insert. The lock tube is secured in place by a top flare, which creates a tight fit and six non-yielding projections on the OD, which interface with the concave side of the insert to preclude escape during core component transfer. The adaptor plate acts as the fuel assembly top end plate and provides a means of evenly distributing any axial loads imposed on the fuel assemblies to the guide thimbles. Guide Thimbles The control rod guide thimbles in the fuel assemblies provide guided channels for the absorber rods during insertion and withdrawal of the control rods. Up to and including Region 18 (VANTAGE+), they are fabricated from a single piece of tubing, which is drawn to two different diameters. The OFA thimbles are Zircaloy-4 and the VANTAGE+ thimble material ZIRLOTM. The larger inside diameter at the top provides a relatively large annular area for rapid insertion Chapter 3, Page 46 of 94 Revision 22, 2010

IP2 FSAR UPDATE during a reactor trip and accommodates a small amount of upward cooling flow during normal operations. The bottom portion of the guide thimble has two sections of reduced diameter producing a "double dashpot" action when the absorber rods near the end of travel in the guide thimbles during a reactor trip. The transition zones at the dashpot sections are conical in shape so that there are no rapid changes in diameter in the tube. Starting with Region 19 (15x15 Upgraded fuel design), the guide thimbles incorporate the tube-in-tube dashpot design. The tube-in-tube design utilizes a separate dashpot tube assembly that is inserted into the guide thimble assembly pulled to a press fit over the thimble end plug and bulged into place. To maintain the same diametrical clearance between the guide thimble ID and the dashpot OD, the 15x15 upgraded nominal dashpot thickness was reduced from 0.0165 to 0.0160 inches. As the dashpot tube in the design can provide additional lateral support in that bottom thimble span, it is expected that there will be additional resistance to lateral deformation and Incomplete Rod Insertions as a result of the design modification. The 15x15 Upgraded fuel thimbles are ZIRLOTM. Flow holes are provided just above the first dashpot transition to permit the entrance of cooling water during normal operation, and to accommodate the outflow of water from the dashpot during reactor trip. The dashpot is open at the bottom by means of the drainage hole in the thimble screws that secure the bottom nozzle to the welded end plugs of the guide thimbles. This geometry is shown in Figure 3.2-57. The top ends of the thimble tubes are mechanically attached to the sleeve of the top grid. An insert is also bulge attached to the thimble and the insert upper end is in turn mechanically attached to the top nozzle as shown in Figure 3.2-58A. VANTAGE+ Grids Prior to Region 18, the VANTAGE+ assembly has twelve grids. Starting with Region 18, a thirteenth grid, the protective grid (P-grid), was added to the VANTAGE+ assembly. The top and bottom grids, as in the OFA assembly, and the P-grid are Inconel 718 non-mixing vane grids. The top and bottom grids are Inconel-718 non-mixing vane grids. Low Pressure Drop (LPD) Zircaloy grids are used for the middle grids with Zircaloy IFMs located in the three uppermost middle grid spans. The VANTAGE+ fuel assembly with PERFORMANCE+ options has ZIRLO' grids for the 3IFMs and 7 mid grids. The LPD grids have mixing vanes, diagonal springs and a reduced grid height, relative to the OFA grids. The LPD grid cells use the standard four dimples and two springs per cell for support locations. The IFMs provide mid-span flow mixing in the hottest fuel assembly spans. Each IFM cell contains four dimples, which are designed to prevent midspan channel closure and fuel rod contact with the mixing vanes. With the additional Performance+ enhancements added to the fuel starting with Region 18, a new Protective Bottom Grid (PBG) has been added. The PBG is a wider, extra grid at the very bottom of the fuel assembly that protects the fuel from debris. Its purpose is to filter out debris and hold it at an elevation below the bottom of the active core. The PBG is not a structural grid. The bottom of the PBG lies below the tops of the lower end plugs within the fuel rod. This means that any debris caught in the PBG cannot fret through the cladding and expose fuel pellets. All VANTAGE+ outside grid straps contain mixing vanes, which also act as guides during fuel handling. The grids are also attached to the thimble tubes via the bulging mechanism as shown Chapter 3, Page 47 of 94 Revision 22, 2010

IP2 FSAR UPDATE in Figure 3.2-61A. Top grid nozzle attachment is shown in Figure 3.2-58A. All grids employ the anti-snag outer strap design. A mixing vane grid is shown in Figure 3.2-59. 15x15 Upgraded Design Grids The 15x15 Upgraded fuel design still contains twelve grids with the top and bottom grids unchanged from the Vantage+ design. The thirteenth grid, the protective gird (P-grid) also remains the same as the Vantage+ design. The middle grids have changed to an I-spring design. The changes were made to improve fuel rod fretting margin. In addition to the spring change the size of the dimples was increased. The strap thickness was decreased to help offset pressure drop increase due to the I-spring and increased grid strap height. The strap height increased to create space to accommodate the increased dimples and the I-spring. The IFM grid design was enhanced to increase contact area also. Fuel Rods The fuel rods consist of uranium-dioxide ceramic pellets in slightly cold worked ZIRLOTM tubing, which is plugged and seal welded at the ends to encapsulate the fuel. Sufficient void volume and clearances are provided within the rod to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, helium released from poison burnup (IFBA rods), and fuel swelling due to accumulated fission products without overstressing of the cladding or seal welds. Shifting of the fuel within the cladding is prevented during handling or shipping prior to core loading by a stainless steel helical compression spring, which bears on the top of the fuel. At assembly, the pellets are stacked in the cladding to the required fuel height. The compression spring is then inserted into the top end of the fuel and the end plugs pressed into the ends of the tube and welded. A hold-down force of approximately four times the weight of the fuel is obtained by compression of the spring between the top end plug and the top of the fuel pellet stack. All fuel rods are internally pressurized with helium in order to minimize compressive clad stresses and creep due to coolant operating pressures. The fuel pellets are in the form of a right circular cylinder and consist of slightly enriched uranium-dioxide powder, which is compacted by cold pressing and sintering to the required density. The ends of each pellet are dished slightly to allow the greater axial expansion at the center of the pellets to be taken up within the pellets themselves and not in the overall fuel length. The 15x15 Upgrade fuel has mid-enriched annular (IFBA) and solid (non-IFBA) pellets in the axial blanket region of the fuel rod and optimized plenum spring to maximize the available plenum volume for increased burnup. The 15x15 Upgrade fuel has a longer fuel rod to allow higher fission gas release due to longer cycles. This is allowable due to the ZIRLO cladding, which has less rod growth on irradiation. For the first core, the pellets in the outer region had a density of approximately 10.3 g/cm3 (94-percent of theoretical density) while those in the two inner regions (checkerboard pattern, see Figure 3.2-48) had a density of 10.4 g/cm3 corresponding to 95-percent of theoretical density. Lower pellet densities were used to compensate for the effects of the higher burnup, which the fuel experienced in those regions. Reload cores contain 15x15 Upgraded fuel arranged in a zoned and/or checkerboard pattern. Different fuel enrichments, as listed in Table 3.2-7, are used for each of the core regions for all core loadings. Chapter 3, Page 48 of 94 Revision 22, 2010

IP2 FSAR UPDATE Each fuel rod is marked with a permanent traceability code. This aids in ensuring that rods of the proper enrichment will be loaded into each fuel assembly. The identification numbers on the fuel assembly top nozzles will then maintain the enrichment identity and ensure that the assemblies with the correct enrichment are loaded into the proper core region. Each assembly is assigned a core loading position. A record is then made of the core loading position, serial number, and enrichment. During the core loading, two independent checks are made to ensure that the actual loading position agrees with the position assigned. During initial core loading and subsequent refueling operations, detailed handling and checkoff procedures are used throughout the sequence. The initial core was loaded in accordance with the core loading diagram similar to Figure 3.2-48, which shows the location for each of the three enrichment types of fuel assemblies used in the loading. 3.2.3.2.1.2 Rod Cluster Control Assemblies The control rods or rod cluster control assemblies consist of a group of individual absorber rods fastened at the top end to a common hub or spider assembly. These assemblies, one of which is shown in Figure 3.2-49, are provided to control the reactivity of the core under operating conditions. The absorber material used in the control rods is silver-indium-cadmium alloy, which is essentially "black" to thermal neutrons and has sufficient additional resonance absorption to increase its worth significantly. The alloy is in the form of extruded single-length rods, which are sealed in stainless steel tubes to prevent the rods from coming in direct contact with the coolant. The overall control rod length is such that when the assembly has been withdrawn through its full travel, the tips of the absorber rods remain engaged in the guide thimbles so that alignment between rods and thimbles is always maintained. Since the rods are long and slender, they are relatively free to conform to any small misalignments with the guide thimble. Prototype tests have shown that the rod cluster control assemblies are very easily inserted and not subject to binding even under conditions of severe misalignment. The spider assembly is in the form of a center hub with radial vanes supporting cylindrical fingers from which the absorber rods are suspended. Handling detents and detents for connection to the drive shaft are machined into the upper end of the hub. A spring pack is assembled into a skirt integral to the bottom of the hub to stop the rod cluster control assembly and absorb the energy from the impact at the end of a trip insertion. The radial vanes are joined to the hub and the fingers are joined to the vanes by furnace brazing. A centerpost, which holds the spring pack and its retainer is threaded into the hub within the skirt and welded to prevent loosening in service. All components of the spider assembly are made from type 304 stainless steel except for the springs, which are Inconel X-750 alloy and the retainer, which is of 17-4 pH material. The absorber rods are secured to the spider so as to ensure trouble free service. The rods are first threaded into the spider fingers and then pinned to maintain joint tightness, after which the pins are welded in place. The end plug below the pin position is designed with a reduced section to permit flexing of the rods to correct for small operating or assembly misalignments. In construction, the silver-indium-cadmium rods are inserted into cold-worked stainless steel tubing, which is then sealed at the bottom and the top by welded end plugs. Sufficient diametral Chapter 3, Page 49 of 94 Revision 22, 2010

IP2 FSAR UPDATE and end clearance are provided to accommodate relative thermal expansions and to limit the internal pressure to acceptable levels. The bottom plugs are made bullet-nosed to reduce the hydraulic drag during a reactor trip and to guide smoothly into the dashpot section of the fuel assembly guide thimbles. The upper plug is threaded for assembly to the spider and has a reduced end section to make the joint more flexible. Stainless steel clad silver-indium-cadmium alloy absorber rods are resistant to radiation and thermal damage thereby ensuring their effectiveness under all operating conditions. Rods of similar design have been successfully used in a number of operating nuclear plants. 3.2.3.2.1.3 Neutron Source Assemblies Six neutron source assemblies were utilized in the first cycle core. These consisted of two assemblies with four secondary source rods each, and four assemblies with one secondary source rod and one primary source rod each. The rods in each assembly were fastened to a spider at the top end. The spider for the four secondary source rod assemblies was similar to the rod cluster control assembly spiders, while the primary source assembly spider was similar to that of the burnable poison and plugging device assemblies. Various source assembly designs are used in the reload cycles. In the first cycle core, the neutron source assemblies were inserted into the rod cluster control guide thimbles in fuel assemblies at unrodded locations. The location and orientation of each of the assemblies in the core is shown in Figure 3.2-62. The primary and secondary source rods both utilize the same type of cladding material as the absorber rods (cold-worked type 304 stainless steel tubing). The secondary source rods contain Sb-Be pellets. The primary source rods contained capsules of Pu-Be source material in the initial core loading; for reload cores, this material may vary. Design criteria similar to that for the fuel rods is used for the design of the source rods; i.e., the cladding is free standing, internal pressures are always less than reactor operating pressure, and internal gaps and clearances are provided to allow for differential expansions between the source material and cladding. 3.2.3.2.1.4 Plugging Devices In order to limit bypass flow through the rod cluster control guide thimbles in fuel assemblies, which do not contain either control rods, source assemblies or burnable absorber rods, the fuel assemblies at those locations were fitted with plugging devices. The plugging devices consist of a flat plate with short rods suspended from the bottom surface and a spring pack assembly. At installation in the core, the plugging devices fit with the fuel assembly top nozzles and rest on the adaptor plate. The short rods project into the upper ends of the thimble tubes to reduce the bypass flow area. The spring pack is compressed by the upper core plate when the upper internals package is lowered into place. Similar short rods are also used on the source assemblies to fill the ends of all vacant fuel assembly guide thimbles. All components in the plugging device, except for the springs, are constructed from type 304 stainless steel. The springs are wound from an age hardenable nickel base alloy to obtain higher strength. Coincident with implementation of the Indian Point Unit 2 OFA transition, removal of thimble plugging devices from the core was allowed. This included the removal of the thimble plugs from the OFA assemblies, previously installed LOPAR assemblies, and all new core component clusters (burnable absorbers and sources). Chapter 3, Page 50 of 94 Revision 22, 2010

IP2 FSAR UPDATE As part of the implementation of the power uprate, Cycle 17 core will contain thimble plugs for all assemblies that do not contain inserts e.g. RCCAs, WABAs or secondary sources. 3.2.3.2.1.5 Burnable Absorber Rods The burnable absorber rods are statically suspended and positioned in vacant rod cluster control thimble tubes within the fuel assemblies at nonrodded core locations. The absorber rods in each fuel assembly are grouped and attached together at the top end of the rods by a flat plate, which fits with the fuel assembly top nozzle and rests on the top adaptor plate. The plate with the absorber rods is held down and restrained against vertical motion with a spring pack, which is attached to the plate and is compressed by the upper core plate when the reactor upper internals package is lowered into the reactor. This ensures that the absorber rods cannot be lifted out of the core by flow forces. The absorber rods used during Cycles 1 through 7 consisted of borated Pyrex glass tubes contained within type 304 stainless steel tubular cladding, which was plugged and seal welded at the ends to encapsulate the glass. The glass was also supported along the length of its inside diameter by a thin-wall type 304 stainless steel tubular inner liner. Starting in Cycle 8, Wet Annular Burnable Absorber (WABA) rods were used and are described in Reference 74. As shown in Figures 3.2-69 and 3.2-70, WABA rods are composed of annular pellets containing aluminum oxide-boron carbide (Al2O3 - B4C) burnable absorber material contained within two concentric Zircaloy tubes. The Zircaloy tubes are plugged and seal welded at the ends to enclose the annular stack of absorber material. The tubes are also the inner and outer cladding of the annular burnable absorber rod. A hold-down device is placed on top of the pellet stack to hold the stack in position and to allow for pellet stack growth. The hold-down device is a C-shape Zircaloy polygonal ring clip. Within the rod is an annular plenum to allow for helium gas release from the absorber material during boron depletion. Reactor coolant flows through the inner tube and outside the outer tube of the annular rod. The annular rods are grouped and attached at the top end to a hold-down assembly and retaining plate in the same way as the borosilicate glass absorber rod. WABA rods are used in preference to standard BPRAs to provide smaller residual burnup penalty. Starting with Cycle 11, Integral Fuel Burnable Absorbers (IFBA) were used in conjunction with the WABA rods. The IFBA features a zirconium diboride coating on the fuel pellet surface on the central portion of the enriched UO2 pellets. IFBA's provide power peaking and moderator temperature coefficient control. IFBA's are described in Reference 88. 3.2.3.2.2 Evaluation of Core Components 3.2.3.2.2.1 Fuel Evaluation The fission gas release and the associated buildup of internal gas pressure in the fuel rods are calculated by overall fuel rod design models, which incorporate time-dependent fuel densification (References 68 and 69). The increase of internal pressure in the fuel rod due to this phenomena is included in the determination of the maximum cladding stresses at the end of core life when the fission product gap inventory is a maximum. Modifications to the initial core fuel design and evaluations are given in the Indian Point Unit 2 Fuel Densification Reports70 and 71. The VANTAGE+ fuel rod design bases and evaluation are given in Reference 88. The fuel rod design has not been changed as part of the 15x15 Upgraded fuel design. Chapter 3, Page 51 of 94 Revision 22, 2010

IP2 FSAR UPDATE The maximum allowable tensile strain in the cladding during steady-state operation, considering the combined effects of internal gas pressure, external coolant pressure, fuel pellet swelling and thermal expansion, and cladding creep is less than 1% from the unirradiated condition throughout core life. For Condition II transients, the total tensile strain during the transient is less than 1% from the pre-transient value. The associated stresses are below the yield strength of the material under steady-state and Condition II transient conditions. To assure that manufactured fuel rods meet a high standard of excellence from the standpoint of functional requirements, many inspections and tests are performed both on the raw material and the finished product. These tests and inspections include chemical analysis, tensile testing of fuel tubes, dimensional inspection, X-ray of both end plug welds, ultrasonic testing, and helium leak tests. In the event of cladding defects, the high resistance of uranium-dioxide fuel pellets to attack by hot water protects against fuel deterioration or decrease in fuel integrity. Thermal stress in the pellets, while causing some fracture of the bulk material during temperature cycling, does not result in pulverization or gross void formation in the fuel matrix. As shown by operating experience and extensive experimental work in the industry, the thermal design parameters conservatively account for any changes in the thermal performance of the fuel element due to pellet fracture. The consequences of a breach of cladding are greatly reduced by the ability of uranium-dioxide to retain fission products including those, which are gaseous or highly volatile. This retentiveness decreases with increasing temperature or fuel burnup, but remains a significant factor even at full power operating temperature in the maximum burnup element. Data on fuel behavior in high burnup uranium-dioxide show that it is possible to conservatively define the fuel swelling as a function of burnup and as-fabricated uranium-dioxide porosity (References 68 and 69). Actual fuel rod damage limits depend upon neutron exposure and normal variation of material properties and are greater than the design limits. For the life of the fuel rod, the actual stresses and strains are below the design limits. Thus, significant margins exist between actual operating conditions and the damage limits. The other parameters having an influence on cladding stress and strain are as follows:

1. Internal gas pressure The maximum rod internal pressure under nominal conditions will be substantially less than the calculated pressure at the design limits. The end-of-life internal gas pressure is dependent upon the fuel rod power history and will not exceed the design limit defined in Section 3.1.2.1 (item 3).

2. Cladding temperature The strength of the fuel cladding is temperature dependent. The minimum ultimate strength reduces to the design yield strength at an average cladding temperature of approximately 850°F. The maximum average cladding temperature during normal operating conditions is given in Table 3.2-6.

Chapter 3, Page 52 of 94 Revision 22, 2010

IP2 FSAR UPDATE

3. Burnup Fuel burnup results in fuel swelling, which, along with fuel thermal expansion, causes tensile cladding strain. Since rod power levels, and hence fuel temperature, decreases with burnup, the fuel pellet diameter increase with burnup is somewhat mitigated by the reduced thermal expansion. The strain design limits and stress design limits are met throughout the burnup lifetime of the fuel. These strain and stress design limits are below the cladding damage limits.

4. Fuel temperature and kW/ft The fuel is designed so that the maximum fuel temperature will not exceed 4700°F during normal operating conditions or unanticipated malfunction transients (Condition II events).

3.2.3.2.2.2 Evaluation of Burnable Absorber Rods The burnable absorber rods are positioned in the core inside rod cluster control assembly guide thimbles and held down in place by attachment to a retainer assembly compressed beneath the upper core plate and, hence, cannot be the source of any reactivity transient. Due to the low heat generation rate and the conservative design of the rods, there is no possibility for release of the poison as a result of helium pressure or clad heating during accident transients including loss-of-coolant. 3.2.3.2.2.3 Effects of Vibration and Thermal Cycling on Fuel Assemblies Analyses of the effect of cyclic deflection of the fuel rods, grid spring, rod cluster control rods, and burnable absorber rods due to hydraulically induced vibrations and thermal cycling show that the design of the components is good for an infinite number of cycles. In the case of the fuel rod grid spring support, the amplitude of a hydraulically induced motion of the fuel rod is extremely small (approximately 0.001-in.), and the stress associated with the motion is significantly small (<100 psi). Likewise, the reactions at the grid spring due to the motion is much less than the preload spring force and contact is maintained between the fuel clad and the grid spring and dimples. Fatigue of the clad and fretting between the clad and the grid support is not anticipated. The effect of thermal cycling on the grid-clad support is a slight relative movement between the grid contact surfaces and the clad. Since the number of cycles of the occurrence is small over the life of a fuel assembly (approximately 3 years), negligible wear of the mating parts is expected. Incore operation of assemblies in the Yankee Rowe and Saxton reactors using similar clad support have verified the calculated conclusions. Additional test results under simulated reactor environment in the Westinghouse Reactor Evaluation Channel also support these conclusions. The dynamic deflection of the control rods and the burnable absorber rods is limited by their fit with the inside diameter of either the upper portion of the guide thimble or the dashpot. With this limitation, the occurrence of truly cyclic motion is questionable. However, an assumed cyclic deflection through the available clearance gap results in an insignificantly low stress in either the Chapter 3, Page 53 of 94 Revision 22, 2010

IP2 FSAR UPDATE either the clad tubing or in the flexure joint at the spider or retainer plate. The above consideration assumes the rods are supported as cantilevers from the spider or the retainer plate in the case of the burnable absorber rods. A calculation assuming the rods are supported by the surface of the dashpots and at the upper end by the spider or retainer results in a similar conclusion. 3.2.3.3 Transition Cores The entire core is now 15x15 Upgraded fuel therefore there are no transition core effects. 3.2.3.4 Control Rod Drive Mechanism Design Description 3.2.3.4.1 Full Length Rods The control rod drive mechanisms are used for withdrawal and insertion of the rod cluster control assemblies into the reactor core and to provide sufficient holding power for stationary support. Fast total insertion (reactor trip) is obtained by simply removing the electrical power allowing the rods to fall by gravity. Design scram time is 2.4 seconds from gripper release to dashpot entry. The complete drive mechanism, shown in Figure 3.2-65, consists of the internal (latch) assembly, the pressure vessel, the operating coil stack, the drive shaft assembly, and the rod position indicator coil stack. Each assembly is an independent unit, which can be dismantled or assembled separately. Each mechanism pressure housing is threaded onto an adaptor on top of the reactor pressure vessel and seal welded. The operating drive assembly is connected to the control rod (directly below) by means of a grooved drive shaft. The upper section of the drive shaft is suspended from the working components of the drive mechanism. The drive shaft and control rod remain connected during reactor operation including tripping of the rods. Main coolant fills the pressure containing parts of the drive mechanism. All working components and the shaft are immersed in the main coolant and depend upon it for lubrication of sliding parts. Three magnetic coils, which form a removable electrical unit and surround the rod drive pressure housing, induce magnetic flux through the housing wall to operate the working components. They move two sets of latches, which lift, lower, and hold the grooved drive shaft. The three magnets are turned on and off in a fixed sequence by solid-state switches for the full length rod assemblies. The sequencing of the magnets produces step motion over the full length of normal control rod travel. The mechanism develops a lifting force approximately two times the static lifting load. Therefore, extra lift capacity is available for overcoming mechanical friction between the moving and the stationary parts. Gravity provides the drive force for rod insertion and the weight of the whole rod assembly is available to overcome any resistance. Chapter 3, Page 54 of 94 Revision 22, 2010

IP2 FSAR UPDATE The mechanisms are designed to operate in water at 650°F and 2485 psig. The temperature at the mechanism head adaptor will be much less than 650°F because it is located in a region where there is limited flow of water from the reactor core while the pressure is the same as in the reactor pressure vessel. A multiconductor cable connects the mechanism operating coils to the 125-V DC power supply. The power supply is described in Section 7.3.2. 3.2.3.4.1.1 Latch Assembly The latch assembly contains the working components, which withdraw and insert the drive shaft and attached control rod. It is located within the pressure housing and consists of the pole pieces for three electromagnets. They actuate two sets of latches, which engage the grooved section of the drive shaft. The upper set of latches move up or down to raise or lower the drive rod by 5/8-in. The lower set of latches have a 1/16-in. axial movement to shift the weight of the control rod from the upper to the lower latches. 3.2.3.4.1.2 Pressure Vessel The pressure vessel consists of the pressure housing and rod travel housing. The pressure housing is the lower portion of the vessel and contains the latch assembly. The rod travel housing is the upper portion of the vessel. It provides space for the drive shaft during its upward movement as the control rod is withdrawn from the core. 3.2.3.4.1.3 Operating Coil Stack The operating coil stack is an independent unit, which is installed on the drive mechanism by sliding it over the outside of the pressure housing. It rests on a pressure housing flange without any mechanical attachment and can be removed and installed while the reactor is pressurized. The three operating coils are made of round copper wire, which is insulated with a double layer of filament type glass yarn. The design operating temperature of the coils is 200°C. Average coil temperature can be determined by resistance measurement. Forced air cooling along the outside of the coil stack maintains a coil casing temperature of approximately 120°C or lower. 3.2.3.4.1.4 Drive Shaft (Drive Rod) Assembly The main function of the drive shaft is to connect the control rod to the mechanism latches. Grooves for engagement and lifting by the latches are located throughout the 144-in. of control rod travel. The grooves are spaced 5/8-in. apart to coincide with the mechanism step length and have 45-degree-angle sides. The drive shaft is attached to the control rod by the coupling. The coupling has two flexible arms, which engage the grooves in the spider assembly. A 0.25-in. diameter disconnect rod runs down the inside of the drive shaft. It uses a locking button at its lower end to lock the drive rod assembly and control rod assembly together. At its Chapter 3, Page 55 of 94 Revision 22, 2010

IP2 FSAR UPDATE upper end there is a disconnect assembly for remote disconnection of the drive rod assembly from the control rod assembly. During plant operation, the drive shaft assembly remains connected to the control rod at all times. It can be attached and removed from the control rod only when the reactor vessel head is removed. 3.2.3.4.1.5 Position Indicator Coil Stack The position indicator coil stack slides over the rod travel housing section of the pressure vessel. It detects drive rod position by means of a cylindrically wound differential transformer, which spans the normal length of the rod travel (144-in.). 3.2.3.4.1.6 Drive Mechanism Materials All parts exposed to reactor coolant, such as the pressure vessel, latch assembly, and drive rod, are made of metals, which resist the corrosive action of the water. Three types of metals are used exclusively: stainless steels, Inconel-X, and cobalt-based alloys. Wherever magnetic flux is carried by parts exposed to the main coolant, 400 series stainless steel is used. Cobalt-based alloys are used for the pins, latch tips, and bearing surfaces. Inconel-X is used for the springs of both latch assemblies and type 304 stainless steel is used for all pressure containment. Hard chrome plating provides wear surfaces on the sliding parts and prevents galling between matting parts (such as threads) during assembly. Outside of the pressure vessel, where the metals are exposed only to the reactor containment environment and cannot contaminate the main coolant, carbon and stainless steels are used. Carbon steel, because of its high permeability, is used for flux return paths around the operating coils. It is zinc-plated approximately 0.001-in. thick to prevent corrosion. 3.2.3.4.1.7 Principles of Operation The drive mechanisms, shown schematically in Figure 3.2-66, withdraw and insert their respective control rods as electrical pulses are received by the operator coils. ON and OFF sequence, repeated by switches in the power programmer, causes either withdrawal or insertion of the control rod. Position of the control rod is indicated by the transformer action of the position indicator coil stack surrounding the rod travel housing. The transformer output changes as the top of the ferromagnetic drive shaft assembly moves up the rod travel housing. Generally during plant operation the drive mechanisms hold the control rods withdrawn from the core in a static position and only one coil, the stationary gripper coil, is energized on each mechanism. Control Rod Withdrawal The control rod is withdrawn by repeating the following sequence: Chapter 3, Page 56 of 94 Revision 22, 2010

IP2 FSAR UPDATE

1. Movable Gripper Coil - ON The movable gripper armature raises and swings the movable gripper latches into the drive shaft groove.

2. Stationary Gripper Coil - OFF Gravity causes the stationary gripper latches and armature to move downward until the load of the drive shaft is transferred to the movable gripper latches.

Simultaneously, the stationary gripper latches swing out of the shaft groove.

3. Lift Coil - ON The 5/8-in. gap between the lift armature and the lift magnet pole closes and the drive rod raises one step length.

4. Stationary Gripper Coil - ON The stationary gripper armature raises and closes the gap below the stationary gripper magnetic pole, swings the stationary gripper latches into a drive shaft groove. The latches contact the shaft and lift it 1/16-in. The load is so transferred from the movable to the stationary gripper latches.

5. Movable Gripper Coil - OFF The movable gripper armature separates from the lift armature under the force of three springs and gravity. Three links, pinned to the movable gripper armature, swing the three movable gripper latches out of the groove.

6. Lift Coil - OFF The gap between the lift armature and the lift magnet pole opens. The movable gripper latches drop 5/8-in. to a position adjacent to the next groove.

Control Rod Insertion The sequence for control rod insertion is similar to that for control rod withdrawal:

1. Lift Coil - ON The movable gripper latches are raised to a position adjacent to a shaft groove.

2. Movable Gripper Coil - ON The movable gripper armature raises and swings the movable gripper latches into a groove.

Chapter 3, Page 57 of 94 Revision 22, 2010

IP2 FSAR UPDATE

3. Stationary Gripper Coil - OFF The stationary gripper armature moves downward and swings the stationary gripper latches out of the groove.

4. Lift Coil - OFF Gravity separates the lift armature from the lift magnet pole and the control rod drops down 5/8-in.

5. Stationary Gripper Coil - ON

6. Movable Gripper Coil - OFF The sequences described above are termed as one step or one cycle and the control rod moves 5/8-in. for each cycle. Each sequence can be repeated at a rate of up to 72 steps/min and the control rods can therefore be withdrawn or inserted at a rate of up to 45-in./min. The sequence timing has been modified to preclude the rod withdrawal event described in NRC Generic Letter 93-04.

Control Rod Tripping If power to the movable gripper coil is cut off, as for tripping, the combined weight of the drive shaft and the rod cluster control assembly is sufficient to move the latches out of the shaft groove. The control rod falls by gravity into the core. The tripping occurs as the magnetic field, holding the movable gripper armature against the lift magnet, collapses and the movable gripper armature is forced down by the weight acting upon the latches. 3.2.3.4.2 Part-Length Rods [Deleted] 3.2.3.5 Fuel Assembly and Rod Cluster Control Assembly Mechanical Evaluation To confirm the mechanical adequacy of the fuel assembly and full length rod cluster control assembly, functional test programs have been conducted on a full scale Indian Point Unit 2 prototype 12-ft canless fuel assembly and control rod. The prototype assembly was tested under simulated conditions of reactor temperature, pressure, and flow for approximately 1000 hr. The prototype mechanism accumulated 2,260,892 steps and 600 scrams. At the end of the test the control rod drive mechanism was still operating satisfactorily. A correlation was developed to predict the amplitude of flow excited vibration of individual fuel rods and fuel assemblies. Inspection of the fuel assembly and drive line components did not reveal significant fretting. The wear of the absorber rods, fuel assembly guide thimbles, and upper guide tubes was minimal. The control rod free fall time against 125-percent of nominal flow was less than 1.5 sec to the dashpot (10-ft of travel). Additional tests had previously been made on a full scale San Onofre mockup version of the fuel assembly and control rods (Reference 73). Chapter 3, Page 58 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2.3.5.1 One-Seventh Scale Mockup Tests A one-seventh scale model of the Indian Point Unit 2 internals was designed and built for hydraulic and mechanical testing. The tests provided information on stresses and displacements at selected locations on the structure due to static loads, flow induced loads, and electromagnetic shaker loads. Flow distribution and pressure drop information were obtained. Results of the static tests indicated that mean strains in the upper core support plate and upper support columns are below design limits. Strains and displacements measured in the model during flow tests verified that no damaging vibration levels were present. Additional information gained from the tests was the natural frequency and damping of the thermal shield and other components in air and water. Model response can be related to the full scale plant for most of the expected exciting phenomena, but across-the-board scaling is not possible. Specifically exciting phenomena, which are strongly dependent on Reynolds number cannot be scaled. In areas where Reynolds number may be important, either (1) the measured vibration amplitudes were many times lower than a level that would be damaging, or (2) full scale vibration data have been obtained. 3.2.3.5.2 Loading and Handling Tests Tests simulating the loading of the prototype fuel assembly into a core location have also been successfully conducted to determine that proper provisions had been made for guidance of the fuel assembly during refueling operation. 3.2.3.5.3 Axial and Lateral Bending Tests Axial and lateral bending tests have been performed in order to simulate mechanical loading of the assembly during refueling operation. Although the maximum column load expected to be experienced in service is approximately 1000 lb, the fuel assembly was successfully loaded to 2200 lb axially with no damage resulting. This information is also used in the design of fuel handling equipment to establish the limits for inadvertent axial loads during refueling. 3.2.4 Fixed Incore Detectors Eight fixed core neutron detectors are installed within the Unit 2 reactor as shown in Figure 3.2-

67. They provide no input to plant instrumentation nor are they needed by the operator. The detectors have been retired and are cut and capped at the seal table.

The detectors are not movable when the primary system is at operating pressure. The assemblies are extracted downward from the core during refueling. The seal table used for the movable detector assemblies is also utilized for the fixed incore detector assemblies. The installation of the fixed incore detector system is expected to cause no significant reactivity effect. If a fixed detector were to fail, the expulsion of reactor coolant would be accommodated by the charging pumps, which is common for ruptures of a very small cross section. This would enable the operator to execute an orderly shutdown. If a seal were to fail at the seal table, flow from the reactor coolant system would be through the annulus defined by the O.D. of the flux thimble and the I.D. of the conduit (approximately 0.13-in.). The fixed incore detector system is designed to Class I standards such that the likelihood of a failure causing a loss of coolant will be extremely remote. Chapter 3, Page 59 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2.4.1 Core Monitoring Verification of axial and radial power distribution during full power operation is performed using the movable incore detector system. These movable detector locations are shown in Figure 3.2-

68. The axial power distribution during operation is determined for each measured thimble location since the activity level is measured at several axial heights for each thimble.

Comparison of the measured power distribution to design predictions provides confirmation of safe operation of the reactor and confidence in design predictions. To obtain temperature maps of the core, 65 fuel assembly outlet thermocouples are located as shown in Figure 3.2-68. These thermocouples are located in the upper coolant internals package above the corresponding assemblies. Based on the average activity determined for each thimble, the measured radial power distribution can be determined. 3.2.5 Plant Computer The computer system provided for Indian Point Unit 2 is a DS&S Plant Integrated Computer System known as PICS. It is provided as an adjunct to the normal control room instrumentation to assist the operator in the operation of the reactor by monitoring reactor performance and displaying it in a consistent, well-ordered, usable form. The computer system performs functions such as scanning, signal converting, calculating, indicating, recording, and alarm annunciating. This system is not required for safety, and operation of the reactor is not in any way dependent upon the availability of the computer. Briefly, the analog scanning includes reading all inputs in a pre-established manner, checking the readings, converting values to engineering units, storing them for future use, updating information, and checking alarm conditions. Some inputs are scanned once every second and status placed in memory; other inputs are not scanned periodically but are given immediate attention. The alarm program compares the values of the inputs against the fixed or variable alarm limits and indicates when off-normal conditions exist. 3.2.6 Current Operating Cycle Indian Point Unit 2 is currently operating in the twentieth cycle. The core for this cycle uses 15x15 Upgraded fuel. In order to reduce neutron fluence to the reactor vessel shell, a low-low leakage loading pattern, shown in Figure 3.2-68A, is utilized. Integrated fuel burnable absorbers (IFBA) and wet annular burnable absorber (WABA) assemblies, described in References 74 and 88, are used. Figures 3.2-68 and 3.2-68B show the locations of core components and instrumentation. The cycle is designed for a burnup of 25,193 MWd/MTU which includes a power coastdown. The methodology and computer codes described in Reference 75, 82, 87, 96, and 97 were utilized for analysis of the current operating cycle. Fuel Temperatures were calculated using the fuel thermal models of References 69 and 84 (prior to Cycle 16 the analysis was done with Reference 83). Beginning with Cycle 11, replacement fuel had debris filter bottom nozzles and integral fuel burnable absorbers. To prevent debris from reaching the core, the nozzles have a larger number of smaller holes than previous designs. The integral fuel burnable absorber (IFBA) fuel has a thin layer of zirconium diboride or enriched zirconium diboride coated directly onto selected fuel pellets to control power peaking. A three dimensional model is used to track and predict core operating characteristics. The code predictions are compared to startup physics test results, measured core flux distributions, and Chapter 3, Page 60 of 94 Revision 22, 2010

IP2 FSAR UPDATE critical boron concentrations as a function of core burnup. It has been shown to provide a suitable and accurate means for predicting core operating conditions. REFERENCES FOR SECTION 3.2

1. Westinghouse Electric Corporation, "Power Distribution Control in Westinghouse Pressurized Water Reactors," WCAP-7208 (Proprietary), 1968.

2. R. F. Barry, et al., "Power Maldistribution Investigations," WCAP-7407-L (Proprietary), Westinghouse Electric Corporation.

3. F. B. Skogen and A. F. McFarlane, "Xenon-Induced Spatial Instabilities in Three-Dimensions," WCAP-3680-22, Westinghouse Electric Corporation, September 1969.

4. T. Morita, et al., "Power Distribution Control and Load Following Procedures,"

WCAP-8385 (Proprietary), September 1974 and WCAP-8403 (Non-Proprietary), Westinghouse Electric Corporation, September 1974.

5. F. L. Langford and R. J. Nath, "Evaluation of Nuclear Hot Channel Factor Uncertainties," WCAP-7308-L (Proprietary), April 1969 and WCAP-7810 (Non-Proprietary), Westinghouse Electric Corporation, December 1971.

6. J. M. Hellman, (Ed), "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary), Westinghouse Electric Corporation, March 1975 and WCAP-8219-A (Non-Proprietary),March 1975.

7. R. O. Meyer, "The Analysis of Fuel Densification," Division of Systems Safety, NUREG-0085, U.S. Nuclear Regulatory Commission, July 1976.

8. J. M. Hellman, C. A. Olson, and J. M. Yang, "Effects of Fuel Densification Power Spikes on Clad Thermal Transients," WCAP-8359, Westinghouse Electric Corporation, July 1974.

9. "Power Distribution Control of Westinghouse Pressurized Water Reactors,"

WCAP-7811 (Non-Proprietary), Westinghouse Electric Corporation, December 1971.

10. R. J. Johnson, "Connecticut-Yankee Tests on Detection of Power Maldistribution,"

WCAP-9010 (Proprietary), Westinghouse Electric Corporation, February 1969.

11. P. M. Wood, E. A. Bassler, et al., "Use of Burnable Poison Rods in Westinghouse Pressurized Water Resources," WCAP-7113, Westinghouse Electric Corporation, October 1967.

12. R. F. Barry, "The Revised LEOPARD Code - A Spectrum Dependent Non-Spatial Depletion Program," WCAP-2759, Westinghouse Electric Corporation, March 1965.

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13. VANTAGE+ Fuel Upgrade Reload Transition Safety Report for Indian Point Unit 2, September 1994.

14. Kersting, P.J. et al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A, March 1995.

15. Davidson, S.L. (Ed.) et al., Extended Burnup Evaluation of Westinghouse Fuel, WCAP-10125-P-A, December 1985.

16. Davidson, S.L (Editor), Westinghouse Fuel Criteria Evaluation Process, WCAP-12488-A, October 1994.

17. Letter from N.J Liparulo (Westinghouse) to R.J. Jones (NRC) Westinghouse Interpretation of Staffs Position on Extended Burnup (Proprietary), NTD-NRC-94-4275, August 29,1994.

18. Deleted

19. Deleted

20. Deleted

21. Deleted

22. Deleted

23. Deleted

24. Deleted

25. R.A. Dean, "Thermal Contact Conductance Between UO2 and Zircaloy-2,"CVNA-127, May 1962.

26. A.M. Ross and R.D. Stoute, "Heat Transfer Coefficient Between UO2 and Zircaloy-2," AECL-1552, June 1962.

27. V.C. Howard and T.G. Gulivin, "Thermal Conductivity Determinations on Uranium Dioxide by a Radial Flow Method", UKAEA IG-Report 51, November1960.

28. C.F. Lucks and H.W. Deem, Thermal conductivity and Electrical Conductivity of UO2 in Progress Reports Relating to Civilian Applications, BMI-1448 (Rev) (June 1960); BMI-1489 (Rev.) (December, 1960) and BMI-1518 (Rev.) Battelle Memorial Institute-Columbus (May, 1961).

29. J. L. Daniel, J. Matolich, Jr., and H.W. Deem, "Thermal Conductivity of UO2," HW-69945, September 1962.

30. A. D. Feith, "Thermal Conductivity of UO2 by a Radial Heat Flow Chapter 3, Page 62 of 94 Revision 22, 2010

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31. J. Vogt, L. Grandell, and U. Runfors, "Determination of the Thermal Conductivity of Unirradiated Uranium Dioxide," AB Atomenergi Report RMB-527, 1964, Quoted by IAEA Technical Report Series No. 59, "Thermal Conductivity of Uranium Dioxide."

32. T. Nishijima, T. Kawada, and A. Ishihata, "Thermal Conductivity of Sintered UO2 and AI2O3 at High Temperatures," J. American Ceramic Society, Vol. 48, pp. 31-34, 1965.

33. J. B. Ainscough and M.H. Wheeler, "Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide," in Proceedings of the Seventh Conference of Thermal Conductivity, p. 467, National Bureau of Standards, Washington, 1968.

34. T. G. Godfrey, et al., "Thermal Conductivity of Uranium Dioxide Armco Iron by and Improved Radial Heat Flow Technique," ORNL-3556, June 1964.

35. J. P. Stora, et al., "Thermal Conductivity of Sintered Uranium Oxide Under In-Pile Conditions," EURAEC-1095, August 1964.

36. A. J. Bush, "Apparatus for Measuring Thermal Conductivity to 2500°C," Report 64-1P6-401-43 (Proprietary), Westinghouse Research Laboratories, February 1965.

37. R. R. Asamoto, F.L. Anselin, and A.E. Conti, "The Effect of Density on the Thermal Conductivity of Uranium Dioxide," GEAP-5493, April 1968.

38. O. L. Kruger, "Heat Transfer Properties of Uranium and Plutonium Dioxide," Paper 11-N-68F presented at the Fall Meeting of Nuclear Division of the American Ceramic Society, Pittsburgh, Pa., September 1968

39. J. A. Gyllander, "In-Pile Determination of the Thermal Conductivity of UO2 in the Range 500-2500°C," AE-411, January 1971.

40. M. F. Lyons, et al., "UO2 Powder and Pellet Thermal Conductivity During Irradiation," GEAP-5100-1, March 1966.

41. R. N. Duncan, "Rabbitt Irradiation of UO2" CVNA-142, June 1962.

42. A. S. Bain, "The Heat Rating Required to Produce Center Melting in Various UO2 Fuels," ASTM Special Technical Publication No. 306, pp. 30-46, Philadelphia, Pa.,

1962.

43. J. P. Storal, "In-Reactor Measurements of the Integrated Thermal Conductivity of UO2 - Effect of Porosity," Trans. ANS, Vol. 13, 137-138, 1970

44. International Atomic Energy Agency, "Thermal Conductivity of Uranium Dioxide,"

Report of the Panel held in Vienna, April 1965, IAEA Technical Reports Series No. 59, Vienna, The Agency, 1966. Chapter 3, Page 63 of 94 Revision 22, 2010

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45. J. A. Christensen, R.J. Allio, and A. Biancheria, "Melting Point of Irradiated Uranium Dioxide," WCAP-6065, Westinghouse Electric Corporation, February 1965.

46. W. R. Smalley, "Survey of Experience With High Performance Fuel Rods of PWR Type, WCAP-7125, Westinghouse Electric Corporation, January 1968.

47. L. S. Tong, "Prediction of Departure From Nucleate Boiling for an Axially Non-Uniform Heat Flux Distribution," Journal of Nuclear Energy, Vol. 21, pp. 241-248, 1967.

48. L. S. Tong, H. B. Currin, and A.G. Throp II, "New Correlations Predict DNB Conditions," Nucleonics, May 1963. Also WCAP-1997, 1963.

49. L. S. Tong, "Boiling Crisis and Critical Heat Flux," AEC Critical Review Series, TID-25887, 1972.

50. F. E. Motely and F. F. Cadek, "Application of Modified Spacer Factor to L-Grid Typical and Cold Wall Cell DNB," WCAP-8030-A (Non-Proprietary),Westinghouse Electric Corporation, October, 1972.

51. L. S. Tong, R. P. Sandberg, and A.A. Bishop, "Forced Convection Heat Transfer at High Pressure After the Critical Heat Flux," ASME 65-HT-31 1965.

52. K. W. Hill, F. E. Motley, and F.F. Cadek, "Effect of Local Heat Flux Spikes on DNB in Non-Uniform Heated Rod Bundles," WCAP-8174 (Proprietary), August 1973, and WCAP-8202 (Non-Proprietary), Westinghouse Electric Corporation, August 1973.

53. G. Hetsroni, "Hydraulic Tests of the San Onofre Reactor Model, WCAP-3269-8, June 1964.

54. G. Hetsroni, "Studies of the Connecticut-Yankee Hydraulic Model," NYO-3250-2, June 1965.

55. F. D. Carter, "Inlet Orificing of Open PWR Cores," WCAP-7836, Westinghouse Electric Corporation, January 1972.

56. J. Shefcheck, "Application of the THINC Program to PWR Design," WCAP-7359-L (Proprietary), August 1969, and WCAP-7838 (Non-Proprietary), Westinghouse Electric Corporation, January 1972.

57. F. F. Cadek, "Interchannel Thermal Mixing With Mixing Vane Grids," WCAP-7667-P-A (Proprietary) and WCAP-7755-A (Non-Proprietary), Westinghouse Electric Corporation, January 1975.

58. I.E. Idel'chik, "Handbook of Hydraulic Resistance," AEC-TR-6630, 1960.

59. L.F. Moody, "Friction Factors for Pipe Flow, "Transaction of the American Society of Mechanical Engineers, Vol. 66, pp. 671-684, 1944.

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60. J. Weisman, A. H. Wenzel, L. S. Tong, D. Fitzsimmons, W. Thorne, and J. Batch, "Experimental Determination of the Departure From Nucleate Boiling In Large Rod Bundles at High Pressure," AICHE, Preprint 29, 9th National Heat Transfer Conference, Seattle, Washington, 1967.

61. L. S. Tong, H. Chelemer, J. E. Casterline, and B. Matzner, "Critical Heat Flux (DNB) in Square and Triangular Array Rod Bundles," JSME, Semi-International Symposium, Paper No. 256, Tokyo, Japan, 1967.

62. Deleted

63. Deleted

64. Deleted

65. Deleted

66. Deleted

67. Deleted

68. J. V. Miller (Ed), "Improved Analytical Models Used in Westinghouse Fuel Rod Design Computations," WCAP-8785, Westinghouse Electric Corporation, October 1976.

69. R. A. Weiner, et al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-11873-A, Westinghouse Electric Corporation, August 1988.

70. "Fuel Densification--Indian Point Nuclear Generating Station Unit No.2," WCAP-8062, Westinghouse Electric Corporation, January 1973.

71. "Addendum to Densification - Indian Point Nuclear Generating Station Unit No. 2,"

WCAP-8062-Addendum, Westinghouse Electric Corporation, dated March 22, 1973.

72. Deleted

73. Westinghouse Electric Corporation, "Large Closed Cycle Water Reactor Research and Development Program Quarterly Progress Reports for the Period January 1963 through June 1965, "WCAP-3738, 3739, 3743, 3750, 3269-2, 3269-5, 3269-6, 3269-12 and 3269-13.

74. Westinghouse Electric Corporation, "Westinghouse Wet Annular Burnable Absorber Evaluation Report," WCAP-10021-P-A, Rev.1 (Proprietary), October 1983.

75. Davidson, S. L. (Ed), et al, "Westinghouse Reload Safety Evaluation Methodology," WCAP-9273-NP-A, Westinghouse Electric Corporation, July 1985.

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76. Motley, F. E., Hill, K. W., Cadek, F. F. and Shefcheck, J., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids," WCAP-8762-P-A, July 1984 (Proprietary) and WCAP-8763, July 1976 and July 1985.

77. H. Chelemer, et al, "THINCIV-An Improved Program for Thermal-Hydraulic Analysis of Rod Bundle Cores," WCAP-7956, June 1973.

78. Hochreiter, L. E., et al, "Application of THINCIV Program to PWR Design," WCAP-8054, September 1973.

79. Chelemer, H., et al, "Improved Thermal Design Procedure," WCAP-8567, July 1975.

80. Skaritka, J., et al, "Fuel Rod Bow Evaluation," WCAP-8691, Rev. 1 (Proprietary),

WCAP-8692, Rev. 1 (non-Proprietary) July 1979.

81. Letter from C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse), "Request for Reduction in Fuel Assembly Burn-Up Limit for Calculation of Maximum Rod Bow Penalty," June 18, 1986.

82. Liu, Y. S., et al, "ANC: A Westinghouse Advanced Nodal Computer Code,"

WCAP-10965-P-A and WCAP-10966-A, September 1986.

83. Leech, W. J., et al, "Revised PAD Code Thermal Safety Model," WCAP-8720 Addendum 2 (Proprietary), October 1982.

84. Slage, W.H. (Ed.), Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Rev.1 w/errata July 2000.

84a. Gergos, B. W. and Slagle W. H., "Reload Safety Evaluation Indian Point Nuclear Plant Unit 2 Cycle 11", Attachment to THFL-91-371, Westinghouse Electric Corporation, June 1991.

85. MD2-04-25, 15x15 Upgrade Fuel Assembly - Design Closeout Package (DR02-3), 5/24/04.

86. Chelemer, H., et al, "Improved Thermal Design Procedure," WCAP-8568-A February 1989.

87. Nguyen, T. Q., et al, "Qualification of the PHOENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Core", WCAP-11596-P-A, November 1987.

87a. Davidson, S. L., "Reference Core Report VANTAGE 5 Fuel Assembly," WCAP-10444, Addendum 1, December 1985, WCAP-10445-NP-A, September 1985.

88. Davidson, S. L., Nuhfer, D. L. (Eds), "VANTAGE+ Fuel Assembly Reference Core Report", WCAP-12610-P-A and Appendices A through G, April, 1995.

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89. Letter from A. C. Thadani (NRC) to W. J. Johnson (Westinghouse), January 31, 1989,

Subject:

Acceptance for Referencing of Licensing Topical Report, WCAP-9226-P/9227-NP, "Reactor Core Response to Excessive Secondary Steam Releases."

90. Davidson, S. L., Iorii, J. A., "Reference Core Report - 17x17 Optimized Fuel Assembly," WCAP-9500-A, May 1982.

91. Letter from E. P. Rahe (Westinghouse) to J. R. Miller (NRC), "WCAP-9500 and WCAP-9401/9402 NRC SER Mixed Core Compatibility Items," NS-EPR-2573, March 19, 1982.

92. Schueren, P. and McAtee, K. R., "Extension of Methodology for Calculating Transition Core DNBR Penalties," WCAP-11837-P-A, January 1990.

93. Letter from C. O. Thomas (NRC) to E. P. Rahe (Westinghouse), "Supplemental Acceptance No. 2 for Referencing Topical Report WCAP-9500," January 1983.

94. Letter from A. C. Thadani (NRC) to W. J. Johnson (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-11837", October 1989.

95. Tong L. S., Critical Heat Flux in Rod Bundles, Two Phase Flow and Heat Transfer in Rod Bundles," Annual Winter Meeting ASME, November 1968, Page 3146.

96. Friedland, A. J. and Ray, S. "Improved THINC-IV Modeling for PWR Core Design", WCAP-12330-P, August 1989.

97. Friedland, A. J. and Ray, S. "Revised Thermal Design Procedure", WCAP 11397-P-A, April 1989.

98. Slagle, W.H., "Westinghouse Fuel Assembly Reconstitution Evaluation Methodology," WCAP-13061-NP-A, July 1993.

99. "ARMP-02 Documentations: Part II, Chapter 11 - NODE-P2 Computer Code Manuals, Vol. I and II," EPRI NP-4574-CCM, Electric Power Research Institute, Palo Alto, CA, October 1988.

100. Stewart, C.W. et al., VIPRE-01: A Thermal-Hydraulic Code for Reactor Cores. Volumes 1-3 (Revision 3, August 1989) Volume 4 (April 1987), NP-2511-CCM-A, Electric Power Research Institute. 101. WCAP-14565-P-A/WCAP-15306-NP-A, VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal-Hydraulic Safety Analysis, October 1999, Sung, Y. et al. Chapter 3, Page 67 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-1 (Sheet 1 of 3) Nuclear Design Data Cycle 1 Values STRUCTURAL CHARACTERISTICS

1. Fuel weight (UO2), lbs 217,800

2. Zircaloy weight, lbs 44,600

3. Core diameter, in. 132.75

4. Core height, in.

Reflector thickness and composition 144

5. Top water plus steel, in. ~ 10

6. Bottom water plus steel, in. ~ 10

7. Side water plus steel, in. ~ 15

8. H2O/U, (cold) core 3.91

9. Number of fuel assemblies 193

10. UO2 rods per assembly 204 PERFORMANCE CHARACTERISTICS

11. Heat output, MWt (initial rating) 2758

12. Heat output, MWt (maximum calculated turbine rating) 3216

13. Fuel burnup, MWd/metric ton uranium 16,100 First cycle enrichments, w/o

14. Region 1 2.21

15. Region 2 2.80

16. Region 3 3.20

17. Equilibrium enrichment 3.2

18. Nuclear heat flux hot channel factor, F T Q 1 2.32

19. Nuclear enthalpy rise hot channel factor, F N H

1.55 Chapter 3, Page 68 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-1 (Sheet 2 of 3) Nuclear Design Data Cycle 1 Values CONTROL CHARACTERISTICS Effective multiplication (Beginning-of-Life) with rods in, no Boron

20. Cold, no power, clean 1.113

21. Hot, no power, clean 1.057

22. Hot, full power, clean 1.031

23. Hot, full power, Xe and Sm equilibrium 1.001

24. Material 5-percent Cd; 15-percent In; 80-percent Ag

25. Full length rod cluster control assemblies, number 53

26. Part length rod cluster control assemblies (removed)

27. Number of absorber rods per rod cluster control assemblies 7,8,9,12,16, or 20

28. Total rod worth, BOL, percent (See Table 3.2-2)

Boron concentration for first core cycle loading with burnable poison rods

29. Fuel loading shutdown; rods in (k = .86) 2000 ppm (k = .90) 1615 ppm

30. Shutdown (k = .99) with rods inserted, clean, cold 849 ppm

31. Shutdown (k = .99) with rods inserted, clean, hot 572 ppm

32. Shutdown (k = .99) with no rods inserted, clean, hot 1405 ppm

33. Shutdown (k = .99) with no rods inserted, clean, cold 1370 ppm Chapter 3, Page 69 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-1 (Sheet 3 of 3) Nuclear Design Data Cycle 1 Values Boron concentration to maintain k = 1 at hot full power, No rods inserted:

34. Clean 1160 ppm

35. Xenon 860 ppm

36. Xenon and Samarium 780 ppm

37. Shutdown, all but one rod inserted, clean, cold (k = .99) 915 ppm

38. Shutdown, all but one rod inserted, clean, hot (k = .99) 677 ppm BURNABLE POISON RODS

39. Number and material 1412 Borated Pyrex Glass

40. Worth hot p 8.2-percent

41. Worth cold p 5.4-percent KINETIC CHARACTERISTICS

42. Moderator temperature coefficient at fuel power (°F-1) - 0.25 x 10-4 to

                                                                       - 3.00 x 10-4

43. Moderator pressure coefficient (psi-1) + 0.2 x 10-6 to

                                                                      + 3.00 x 10-6

44. Moderator density coefficient, delta k/gm/cm3 - 0.1 to .30

45. Doppler coefficient (°F-1) - 1.1 x 10-5 to 1.8 x 10-5

46. Delayed neutron fraction, percent 0.50 to .72

47. Prompt neutron lifetime, sec 1.50 x 10-5 to 2.0 x 10-5 Note:

1. The total flux hot channel factor (F QT ) is a generic limit.

The actual value is presented in the Technical Specifications. Chapter 3, Page 70 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-1A Nuclear Design Data Cycle 20 Values Cycle 19

1. Heat output, MWt 3216

2. Fuel loading shutdown boron >2050 concentration; rods in (k 0.95)

3. Most positive Moderator -9.77 Temperature Coefficient (pcm/°F)

4. Least Negative Doppler- Only -13.35 to -8.53 Power Coefficient, Zero to Full Power (pcm/% Power)

5. Most Negative Doppler- Only -14.74 to -8.98 Power Coefficient, Zero to Full Power (pcm/% Power)

6. Effective average delayed 0.510 to 0.621 neutron fraction Beff, percent

7. Prompt neutron lifetime, sec 11.60 to 14.17

8. Design bases minimum 1.3 shutdown (% p)

9. Nuclear Heat Flux Hot Channel 2.5 Factor (Fq) limit

10. Nuclear Enthalpy Rise Hot 1.7 Channel Factor (FdH) limit Chapter 3, Page 71 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-2 Reactivity Requirements for Control Rods for Cycle 11 Percent Requirements Beginning-of-Life End-of-Life Control Power defect 1.90 3.05 Operational maneuvering band 0.40 0.40 Control rod bite 0.10 0.10 X-Y xenon rods 0.20 0.20 Total control 2.60 3.75 Notes:

1. Design values used for performing preoperational calculations and analyses.

Chapter 3, Page 72 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-3 Calculated Rod Worths, for Cycle 1 Design Less Reactivity Shutdown Core Rod Worth 10-percent1 Requirements Margin Condition Configuration (Percent) (Percent) (Percent) (Percent) BOL, HFP 53 rods in 8.46 52 rods in 7.43 6.69 2.60 4.09 highest worth rod stuck out EOL, HFP 53 rods in 7.98 52 rods in; 6.48 5.83 3.75 2.082 highest worth rod stuck out BOL = Beginning-of-life EOL = End-of-life HFP = Hot full power Notes:

1. Calculated rod worth is reduced by 10-percent to allow for uncertainties.

2. The design basis minimum shutdown margin is 1.95-percent.

Chapter 3, Page 73 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-4 DELETED TABLE 3.2-5 DELETED Chapter 3, Page 74 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-6 (Sheet 1 of 3) Thermal and Hydraulic Design Parameters All Hipar All Lopar All OFA All Fuel Core5 Fuel Core5 Fuel Core5 VANTAGE+ Fuel Core6 Total heat Output, MWt 2758 2758 3071.4 3216 Total Heat Output, Btu/hr 9,413x106 9,413x106 10,483x106 10,973x106 Heat generated in fuel, 97.4 97.4 97.4 97.4 percent Nominal system pressure, 2250 2250 2250 2250 psia Coolant Flow Total flow rate, 129.57 128.3 121.72 123.3 x 106 lbs/hr1 Avg velocity along fuel 14.8 14.6 13.0 13.80 rods, ft/sec Avg mass flow, 2.42 2.39 2.21 2.24 x 106 lb/hr-ft2 Coolant temperature, oF Nominal inlet 541.6 541.3 547.7 538.2 Average rise in vessel 55.8 56.3 64.4 67.6 Average rise in core 58.2 58.7 67.9 71.8 Average in core 571.7 571.7 583.5 575.9 Average in vessel 569.5 569.5 579.7 572.0 Heat transfer Active heat transfer 51,400 52,100 52,100 52,100 surface area,ft2 Average heat flux, Btu/hr- 178,500 176,000 196,000 205,200 ft2 Maximum heat flux, 414,000 408,300 490,000 513,100 Btu/hr-ft2 Chapter 3, Page 75 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-6 (Sheet 2 of 3) Thermal and Hydraulic Design Parameters All Hipar All LOPAR All OFA All VANTAGE+ Fuel Core5 Fuel Core5 Fuel Core6 Fuel Core5 Heat Transfer (continued) Maximum thermal output for normal operation, KW/ft 13.42 13.42 15.863 16.63 Maximum clad surface temperature for normal operation, °F 657 657 663 NA Fuel central temperatures for nominal fuel rod dimensions, °F Maximum at 100-percent power <4700 <4700 <4700 <4700 DNB ratio Minimum DNB ratio at nominal 1.95 1.84 2.45 2.40 operating conditions (Thimble) Typical NA NA 2.33 2.50 Pressure drop, psi Across core 24.0 25.5 27.2 29.0 Across vessel, including nozzles 50.0 ~51.5 ~55.0 NA Chapter 3, Page 76 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-6 (Sheet 3 of 3) Thermal and Hydraulic Design Parameters 3230 MWt 3230 MWt (Low Temp (High Temp Extreme) 7 Extreme) 7 NSSS Power, MWt 3230 3230 Core Power, MWt 3216 3216 Thermal Design Flow, Loop gpm 80,700 80,700 Reactor Thermal Design Flow, 126.8 123.3 Total, 106 lbm/hr Reactor Coolant Pressure, psia 2250 2250 Reactor Coolant Temperature, °F Core Outlet 588.1 610.0 Vessel Outlet 583.7 605.8 Core Average 552.6 575.9 Vessel Average 549.0 572.0 Vessel/Core Inlet 514.3 538.2 Zero Load Temperature 547 547 Percent Tube Plugging 108 108 Core Bypass Percent 6.54 6.54 Notes:

1. The thermal design flow rate for all the all HIPAR core reflects a 5-percent flow reduction (to account for postulated 25-percent steam generator tube plugging). The thermal design flow rate for the all LOPAR core reflects a 6-percent flow reduction (5-percent reduction to account for postulated 25-percent steam generator tube plugging and an additional 1-percent reduction to account for an all LOPAR fuel core). For all OFA and all VANTAGE+ Fuel Cores, the thermal design flow rate reflects a 5-percent flow reduction (to account for postulated 10-percent steam generator tube plugging).

2. This power level is based upon a peaking factor (Fq) of 2.32.

3. This power level is based upon a peaking factor (Fq) of 2.50

4. Increased bypass flow is due to thimble plug deletion, and IFMs.

5. This data is historic only.

6. This data reflects the current core with uprated power of 3216 MWt.

7. This data is for analysis extremes covering a range of vessel average temperatures.

8. The tube plugging level is supported by the Thermal-hydraulic safety analyses.

Chapter 3, Page 77 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-7 Core Mechanical Design Parameters1 Fuel assemblies Number 193 Rod array 15 x 15 Rods per assembly 2042 Rod pitch, in. 0.563 Overall dimensions 8.426 x 8.426 HIPAR6/LOPAR6 8.424 x 8.424 OFA 8.426 x 8.426 VANTAGE+/15x15 Upgraded Number of grids per assembly (HIPAR6/LOPAR6/OFA) 9 (VANTAGE+/15x15 Upgraded) 13 Number of instrumentation thimbles 1 Number of guide thimbles 20 Diameter of guide thimbles, upper part, in., HIPAR6 0.545 O.D. x 0.515 I.D. Diameter of guide thimbles, lower part, in., HIPAR6 0.484 O.D. x 0.454 I.D. Diameter of guide thimbles, upper part, in., LOPAR6 0.546 O.D. x 0.512 I.D. Diameter of guide thimbles, lower part, in., LOPAR6 0.489 O.D. x 0.455 I.D. Diameter of guide thimbles, upper part, in., OFA/V+ 0.533 O.D. x 0.499 I.D. Diameter of guide thimbles, lower part, in., OFA/V+ 0.489 O.D. x 0.455 I.D. Diameter of guide thimbles, upper part, in., 15x15 Upgraded 0.533 O.D. x 0.499 I.D. Diameter of guide thimbles, lower part, in., 15x15 Upgraded 0.487 O.D. x 0.455 I.D. Fuel rods Number 39,369 (+3 stainless steel rods) Outside diameter, in. 0.422 Diametral gap, in. 0.0075 Clad thickness, in. 0.0243 Clad material Zircaloy (HIPAR6/LOPAR6/OFA) ZIRLOTM (VANTAGE+/15x15 Upgraded) Overall length 148.6, HIPAR6 151.9, LOPAR6 152.17, OFA 152.55, VANTAGE+ 152.88,V+ w/P+ Enhancements/15x15 Upgraded Length of end cap, overall, in. 0.688, HIPAR6 0.265, LOPAR6 0.357 OFA/v+ (TOP) 0.430 OFA/V+ (BOTTOM) 0.350,V+w/P+/15x15 Upgraded (TOP) 0.810,V+w/P+/15x15 Upgraded (BOTTOM) Length of end cap, inserted in rod 0.250, HIPAR6 0.200, LOPAR6 0.130, OFA/V+/15x15 Upgraded Active fuel length, in. 142, HIPAR6 144, LOPAR6 144, OFA/V+/15x15 Upgraded Chapter 3, Page 78 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-7 (Cont.) Core Mechanical Design Parameters1 Fuel pellets Material UO2 sintered Density (percent of theoretical) Region 20B 94.49 15x15 Upgraded Region 21A 95.85 15x15 Upgraded Region 21B 95.74 15x15 Upgraded Region 22A 95.50 15x15 Upgraded Region 22B 95.50 15x15 Upgraded Feed enrichments w/o3 Region 20B 4.94 15x15 Upgraded Region 21A 4.60 15x15 Upgraded Region 21B 4.96 15x15 Upgraded Region 22A 4.80 15x15 Upgraded Region 22B 4.95 15x15 Upgraded Diameter, in. 0.3659 Length, in. 0.4390 0.500 (Blanket) Rod cluster control assemblies Neutron absorber 5-percent Cd, 15-percent In, 80-percent Ag Cladding material Type 304 SS - cold worked Clad thickness, in. 0.019 Number of clusters 53 Number of control rods per cluster 20 Length of rod control, in. 156.436 (overall) Length of absorber section, in. 142.00 Chapter 3, Page 79 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3.2-7 (Cont.) Core Mechanical Design Parameters1 Core structure Core barrel I.D., in. 148.0 O.D., in. 152.5 Thermal shield I.D., in. 158.5 O.D., in. 164.0 Wet Annular Burnable Absorber (WABA) Rods Number 640 Pellet Stack Length 128 Pellet Material Al2O3-B4C Boron Loading (Natural) .0243 g/cm (B-10) .0060 g/cm Pellet O.D. /I.D. .318"/.278" Tube material ZIRLOTM Outer tube O.D. /I.D. .3810"/.3290" Inner tube O.D. /I.D. .2670"/.2250" Integral Fuel Burnable Absorber (IFBA) Rods Number 5396 Absorber Zirconium Diboride B-10 Loading (mg/inch)4 2.65 (1.5X) 2.21 (1.25X) IFBA Coating Length 120/128inches Notes:

1. All dimensions are for cold conditions. Data is for all fuel types unless otherwise stated.

2. Twenty-one rods are omitted: Twenty provide passage for control rods and one contains incore instrumentation.

3. Reload fuel regions have variable enrichments depending on energy requirements, the number of assemblies being fed, and the degree of low leakage (i.e. number of feed assemblies on the periphery).

4. Nominal values.

5. Deleted

6. Symbol representing old and removed fuel assemblies.

Chapter 3, Page 80 of 94 Revision 22, 2010

IP2 FSAR UPDATE 3.2 FIGURES Figure No. Title Figure 3.2-1 Typical Power Peaking Factor Versus Axial Offset Figure 3.2-2 Rod Cluster Groups - Cycle 1 [Historical] Figure 3.2-3 Assembly Average Power & Burnup, Cycle 1 Calculations, BOL, Unrodded Core [Historical] Figure 3.2-4 Assembly Average Power & Burnup, Cycle 1 Calculations, EOL, Unrodded Core [Historical] Figure 3.2-5 Assembly Average Power Distribution Cycle 1 Calculations, BOL, Group C4 Inserted [Historical] Figure 3.2-6 Assembly Average Power Distribution Cycle 1 Calculations, BOL Part-Length Rods In [Historical] Figure 3.2-7 Cycle 1 Maximum FQ X Power Versus Axial Height During Normal Operation [Historical] Figure 3.2-7A Deleted - See Unit 2 COLR For Normalized K (Z) - Fq Vs. Axial Height For Cycle 17 Figure 3.2-8 Burnable Poison & Source Assembly Locations - Cycle 1 Figure 3.2-9 Burnable Poison Rod Locations - Cycle 1 [Historical] Figure 3.2-10 Moderator Temperature Coefficient Vs Moderator Temperature - EOL, Cycle 1 [Historical] Figure 3.2-11 Moderator Temperature Coefficient Vs Moderator Temperature - BOL, Cycle 1 Full Power [Historical] Figure 3.2-12 Moderator Temperature Coefficient Vs Moderator Temperature - BOL, Cycle 1 Zero Power [Historical] Figure 3.2-13 Doppler Coefficient Vs Effective Fuel Temperature - Cycle 1 [Historical] Figure 3.2-14 Power Coefficient Vs Percent Power - Cycle 1 [Historical] Figure 3.2-15 Power Coefficient - Closed Gap Model Figure 3.2-16 Deleted Figure 3.2-17 Deleted Figure 3.2-18 Deleted Figure 3.2-19 Deleted Figure 3.2-20 Deleted Figure 3.2-21 Deleted Figure 3.2-22 Deleted Figure 3.2-23 Deleted Figure 3.2-24 Deleted Figure 3.2-25 Deleted Figure 3.2-26 Deleted Figure 3.2-27 Deleted Figure 3.2-28 Deleted Figure 3.2-29 Deleted Figure 3.2-30 Deleted Figure 3.2-31 Deleted Figure 3.2-32 Deleted Figure 3.2-33 Deleted Figure 3.2-34 Deleted Figure 3.2-35 Deleted Chapter 3, Page 81 of 94 Revision 22, 2010

IP2 FSAR UPDATE Figure 3.2-36 Deleted Figure 3.2-37 Deleted Figure 3.2-38 Typical Thermal Conductivity Of UO2 Figure 3.2-39 High Power Fuel Rod Experimental Program Figure 3.2-40 Typical Comparison Of W-3 Prediction And Uniform Flux Data Figure 3.2-41 Typical W-3 Correlation Probability Distribution Curve Figure 3.2-42 Comparison Of "L" Grid Typical And Thimble Cold Wall Cell Rod Bundle DNB Data For Non-Uniform Axial Heat Flux With Predictions Of W-3 X F'SL Figure 3.2-43 Typical Comparison Of W-3 Correlation With Rod Bundle DNB Data (Simple Grid Without Mixing Vane) Figure 3.2-44 Typical Comparison Of W-3 Correlation With Rod Bundle DNB Data (Simple Grid With Mixing Vane) Figure 3.2-44A Typical Measured Versus Predicted Critical Heat Flux-WRB-1 Correlation Figure 3.2-45 Typical Stable Film Boiling Heat Transfer Data And Correlation Figure 3.2-46 Core Cross Section Figure 3.2-47 Reactor Vessel Internals Figure 3.2-48 Core Loading Arrangement - Cycle 1 [Historical] Figure 3.2-49 Typical Rod Cluster Control Assembly Figure 3.2-50 Rod Control Cluster Assembly Outline Figure 3.2-51 Core Barrel Assembly Figure 3.2-52 Upper Core Support Structure Figure 3.2-53 Guide Tube Assembly Figure 3.2-54 Fuel Assembly And Control Cluster Cross Section - HIPAR, LOPAR, And OFA And VANTAGE+ Figure 3.2-55 HIPAR Fuel Assembly Figure 3.2-56 LOPAR Fuel Assembly Figure 3.2-56A OFA Fuel Assembly Figure 3.2-56B VANTAGE+ Fuel Assembly Figure 3.2-57 Guide Thimble To Bottom Nozzle Joint Figure 3.2-58 LOPAR Top Grid To Nozzle Attachment Figure 3.2-58A OFA And VANTAGE+ Top Grid To Nozzle Attachment Figure 3.2-59 Spring Clip Grid Assembly Figure 3.2-60 Mid-Grid Expansion Joint Design Plan View Figure 3.2-61 Elevation View - LOPAR Grid To Thimble Attachment Figure 3.2-61A Elevation View-VANTAGE+ Grid To Thimble Attachment Figure 3.2-61B Vantage+ Fuel Assembly With Performance+ Enhancements Figure 3.2-61C 15x15 Upgraded Fuel Assembly Figure 3.2-62 Cycle 1 - Neutron Source Locations [Historical] Figure 3.2-63 HIPAR Burnable Poison Rod Figure 3.2-64 LOPAR Burnable Poison Rod Figure 3.2-65 Control Rod Drive Mechanism Assembly Figure 3.2-66 Control Rod Drive Mechanism Schematic Figure 3.2-67 Thimble Location - Fixed Incore Detectors Chapter 3, Page 82 of 94 Revision 22, 2010

IP2 FSAR UPDATE Figure 3.2-68 Cycle 14 Incore Detector, Thermocouple And Flow Mixing Device Locations Figure 3.2-68A Cycle 20 Region And Fuel Assembly Locations Figure 3.2-68B Cycle 20 Core Components And Fresh IFBA Locations Figure 3.2-69 Comparison Of Borosilicate Glass Absorber Rod With WABA Rod Figure 3.2-70 Wet Annular Burnable Absorber Rod 3.3 TESTS AND INSPECTIONS 3.3.1 Reactivity Anomalies To eliminate possible errors in the calculations of the initial reactivity of the core and the reactivity depletion rate, the predicted relation between fuel burnup and the boron concentration, necessary to maintain adequate control characteristics, must be adjusted (normalized) to reflect actual core conditions. When full power is reached initially, and with the control rod groups in the desired positions, the boron concentration is measured and the predicted curve adjusted to this point. As power operation proceeds, the measured boron concentration is compared with the predicted concentration and the slope of the curve relating burnup and reactivity was compared with that predicted. This process of normalization is completed before a cycle burnup of 60 Effective Full Power Days (EFPDs) is reached. Thereafter, actual boron concentration is compared with prediction, and the reactivity status of the core can be continuously evaluated. Any reactivity anomaly greater than 1-percent would be unexpected, and its occurrence would be thoroughly investigated and evaluated. The methods employed in calculating the reactivity of the core versus burnup and the reactivity worth of boron versus burnup are given in Section 3.2.1. 3.3.2 Thermal And Hydraulic Tests And Inspections General hydraulic tests on models were used to confirm the design flow distributions and pressure drops.1,2 Fuel assemblies and control and drive mechanisms were also tested. Onsite measurements were made to confirm the design flow rates. Vessel and internals inspections were also reviewed to check such thermal and hydraulic design values as bypass flow. As part of startup physics testing, a series of core power distribution measurements were made over the entire range of operation in terms of design control rod configuration by means of the core movable detector system. These measurements were analyzed and the results compared with the analytical predictions upon which the safety analysis was based with regard to both radial and axial power distribution. The design hot-channel factors were used as criteria for acceptable results. 3.3.3 Core Component Tests And Inspections To ensure conformance of all materials, components, and assemblies to the design requirements, a release point program is established with the assembly manufacturer, which requires upgrading of all raw materials, special processes (i.e., welding, heat treating, nondestructive testing, etc.) and those characteristics of detail parts, which directly affect the assembly and alignment of the reactor internals. The upgrading is accomplished by the issuance of an inspection release by quality control after conformance has been verified. Chapter 3, Page 83 of 94 Revision 22, 2010

IP2 FSAR UPDATE A resident quality control representative performs a surveillance/audit program at the manufacturer's facility and witnesses the required tests and inspections and issues the inspection releases. An example is the radiographic examination of the welds joining core barrel shell courses. Components and materials supplied by Westinghouse to the assembly manufacturer are subjected to a similar program. Quality control engineers develop inspection plans for all raw materials, components, and assemblies. Each level of manufacturing is evaluated by a qualified inspector for conformance, i.e., witnessing the ultrasonic testing of core plant raw material. Upon completion of specified events, all documentation is audited prior to releasing the material or component for further manufacturing. All documentation and inspection releases are maintained in the quality control central records section. All materials are traceable to the mill heat number. In conclusion, a set of "as-built" dimensions are taken to verify conformance to the design requirements and assure proper fit-up between the reactor internals and the reactor pressure vessel. 3.3.3.1 Quality Assurance Program The quality assurance program plan of the Westinghouse Nuclear Fuel Division is summarized in Reference 3. The program provides for control over all activities affecting product quality, commencing with design and development and continuing through procurement, materials handling, fabrication, testing and inspection, storage, and transportation. The program also provides for the indoctrination and training of personnel and for the auditing of activities affecting product quality through a formal auditing program. Westinghouse drawings and product, process, and materials specifications identify the inspections to be performed. 3.3.3.2 Quality Control Quality control philosophy is generally based on the following inspections being performed to a 95-percent confidence that at least 95-percent of the product meets specification, unless otherwise noted.

1. Fuel system components and parts The characteristics inspected depend upon the component parts; the quality control program includes dimensional and visual examinations, check audits of test reports, material certification, and nondestructive examination, such as X-ray and ultrasonic.

All material used in this core is accepted and released by quality control. Chapter 3, Page 84 of 94 Revision 22, 2010

IP2 FSAR UPDATE

2. Pellets Inspection is performed for dimensional characteristics such as diameter, density, length, and squareness of ends. Additional visual inspections are performed for cracks, chips, and surface conditions according to approved standards.

Density is determined in terms of weight per unit length and is plotted on zone charts used in controlling the process. Chemical analyses are performed on a specified sample basis throughout pellet production.

3. Rod inspection The fuel rod inspection consists of the following nondestructive examination techniques and methods, as applicable:

a. Each rod is leak tested using a calibrated mass spectrometer, with helium being the detectable gas.

b. Rod welds are inspected by ultrasonic test or X-ray in accordance with a qualified technique and Westinghouse specification.

c. All rods are dimensionally inspected prior to final release. The requirements include such items as length, camber, and visual appearance.

d. All fuel rods are inspected by gamma scanning or other approved methods to ensure proper plenum dimensions.

e. All fuel rods are inspected by gamma scanning, or other approved methods to ensure that no significant gaps exist between pellets.

f. All fuel rods are active gamma scanned to verify enrichment control prior to acceptance for assembly loading.

g. Traceability of rods and associated rod components is established by quality control.

4. Assemblies Each fuel assembly is inspected for compliance with drawing and/or specification requirements. Other incore control component inspection and specification requirements are given in paragraph 4.2.3.4 of Reference 3.

5. Other inspections The following inspections are performed as part of the routine inspection operation:

a. Tool and gauge inspection and control, including standardization to primary and/or secondary working standards. Tool inspection is performed at Chapter 3, Page 85 of 94 Revision 22, 2010

IP2 FSAR UPDATE prescribed intervals on all serialized tools. Complete records are kept of calibration and conditions of tools.

b. Audits are performed of inspection activities and records to ensure that prescribed methods are followed and that records are correct and properly maintained.

c. Surveillance inspection, where appropriate, and audits of outside contractors are performed to ensure conformance with specified requirements.

6. Process control

a. To prevent the possibility of mixing enrichments during fuel manufacture and assembly, strict enrichment segregation and other process controls are exercised.

The uranium-dioxide powder is kept in sealed containers. The contents are fully identified both by descriptive tagging and preselected color coding. A Westinghouse identification tag completely describing the contents is affixed to the containers before transfer to powder storage. Isotopic content is confirmed by analysis. Powder withdrawal from storage can be made by only one authorized group, which directs the powder to the correct pellet production line. All pellet production lines are physically separated from each other and pellets of only a single nominal enrichment and density are produced in a given production line at any given time. Finished pellets are placed on trays identified with the same color code as the powder containers and transferred to segregated storage racks within the confines of the pelleting area. Samples from each pellet lot are tested for isotopic content and impurity levels prior to acceptance by quality control. Physical barriers prevent mixing of pellets of different nominal densities and enrichments in this storage area. Unused powder and substandard pellets are returned to storage in the original color-coded containers. Loading of pellets into the clad is performed in isolated production lines, and again only one enrichment and density loaded on a line at a time. A serialized traceability code is placed on each fuel tube to provide unique identification. The end plugs are inserted and then inert-welded to seal the tube. The fuel tube remains coded and traceability identified until just prior to installation in the fuel assembly. At the time of installation into an assembly, the traceability codes are removed and a matrix is generated to identify each rod in its position within a given assembly. The top nozzle is inscribed with a permanent identification number providing traceability to the fuel contained in the assembly. Similar traceability is provided for burnable poison, source rods, and control rodlets, as required. Chapter 3, Page 86 of 94 Revision 22, 2010

IP2 FSAR UPDATE REFERENCES FOR SECTION 3.3

1. G. Hetsroni, "Hydraulic Tests of the San Onofre Reactor Model, "WCAP-3269-8, Westinghouse Electric Corporation, 1964.

2. G. Hetsroni, "Studies of the Connecticut-Yankee Hydraulic Model, "WCAP-2761, Westinghouse Electric Corporation, 1965.

3. J. Moore, "Nuclear Fuel Division Quality Assurance Program Plan;" WCAP-7800, Revision 5, Westinghouse Electric Corporation, November 1979.

Appendix 3A EXPERIMENTAL VERIFICATION OF CALCULATIONS FOR BORON BURNABLE POISON RODS A number of experiments were performed at the Westinghouse Reactor Evaluation Center to investigate the reactivity worth of Pyrex glass tubing similar to that employed in the Indian Point Unit 2 core as burnable poison rods. Several configurations with and without glass burnable poison rods and with fuel loadings representative of power reactors were tested. The reactor used was a rectangular core 4-ft high with 29 or 30 fuel rods on a side. In each case the water height was adjusted until the reactor was just critical. Analyses were performed for each of the configurations measured to determine the adequacy of the methods used to calculate burnable poison rod worths in the design of the Indian Point Unit 2 core. The results of the calculations for the different experimental configurations are listed in Table 3A-1. In each case the eigenvalue should be compared to the appropriate reference eigenvalue (core with fuel only) to eliminate the systematic bias, which appears in the clean core calculation. The discrepancy between the eigenvalue calculated for the unpoisoned and poisoned cases has been related to the fractional error in the neutron current into the boron. This error is also given in Table 3A-1. The burnable poison rods used in Indian Point Unit 2 correspond to the thick-walled tubes and in these cases the agreement is generally better than 5-percent. Chapter 3, Page 87 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3A-1 (Sheet 1 of 2) Calculations and Burnable Poison Rod Worths1 Just Critical Calculated Keff Error in Poison Case Configuration Loading Water Height, cm For Critical Core Absorption, Percent 1 No inserts - clean 252 = 625 64.10 2 One solid glass rod 252-1 = 624 70.00 3 No inserts clean core 292 = 841 43.25 .9969152 4 Uniform thin wall glass 292-25 = 816 69.61 .994633 4.6 5 Uniform thick wall glass 292-25 = 816 83.81 .994800 3.5 6 Central assembly pattern - 292-72 = 769 39.52 1.000085 72 water holes 7 Central assembly pattern - 292-72 = 769 83.76 .996162 1.4 36 thin wall glass 36 water holes 8 Central assembly pattern - 292-72 = 769 114.95 .996392 0.8 36 thick wall glass 36 water holes 9 4 water holes array 292-16 = 825 42.25 10 No inserts - clean 302 = 900 40.69 .9965962 Chapter 3, Page 88 of 94 Revision 22, 2010

IP2 FSAR UPDATE TABLE 3A-1 (Sheet 2 of 2) Calculations and Burnable Poison Rod Worths1 Just Critical Calculated Keff Error in Poison Case Configuration Loading Water Height, cm For Critical Core Absorption, Percent 11 Uniform water holes 302-36 = 864 38.91 .99777 12 Uniform voids (SS cladding) 302-36 = 864 42.62 13 Uniform thin wall glass tubes 302-36 = 864 67.18 .993915 4.8 14 Uniform thick wall glass tubes 302-36 = 864 82.56 .994875 2.5 15 Uniform - .260-in. Ag-In-Cd 302-32 = 868 92.34 16 Alternate thin wall glass tubes 302-18 = 882 49.42 .994963 6.2 17 Alternate thick wall glass tubes 302-18 = 882 51.90 .994965 5.1 18 Alternate solid glass rods - bare 302-18 = 882 54.91 .995925 1.8 19 Alternate pattern - Ag-In-Cd (.260) 302-18 = 882 53.89 20 Alternate pattern - Ag-In-Cd (.330) 302-18 = 882 59.94 21 Four assembly pattern - water holes 302-80 = 820 37.13 .999352 22 Four assembly pattern - 302-80 = 820 68.73 .997573 1.8 40 thick wall glass 40 water holes Notes:

1. Reactivity data of glass rods, encased in 0.020-in. SS 0.395-in. I.D., 0.435-in. O.D.

Nominal glass rod data: solid rod, 0.396-in. O.D., L = 48-in.; thick wall I.D. = 0.228-in., O.D. = 0.375-in., L = 48-in. thin wall I.D. = 0.316-in., O.D. = 0.394-in., L = 48-in.

2. Reference case.

Chapter 3, Page 89 of 94 Revision 22, 2010

IP2 FSAR UPDATE Appendix 3B POWER DISTRIBUTION CONTROL 3B.1 GENERAL The spatial stability of the xenon distribution in large PWRs has been the subject of extensive investigation by Westinghouse. General studies (performed in part under the Euratom Xenon Program) are reported in WCAP-3680-20, 21, 22, and 231-4 and specific studies related to the Indian Point Unit 2 reactor are reported in WCAP-7407-L5 (Westinghouse Proprietary). Confidence that the reactor can be maintained within thermal limits (design nuclear hot channel factors) is provided by the following:

1. Results of the extensive analytical investigation of potential spatial instability arising from redistribution of xenon in the Indian Point Unit 2 reactor lead to the conclusions that (a) the reactor may be unstable toward axial spatial oscillations and (b) is stable toward radial or diametral (quadrant to quadrant) xenon spatial oscillations.

2. Stability towards diametral (X-Y) xenon oscillations was demonstrated during Cycle 1 startup tests per a report submitted by Con Edison to the NRC (Reference 6).

3. Continuous monitoring and appropriate alarm functions of both axial and diametral power tilts, using signals from the eight ex-core ion chambers, with additional information provided by the core exit thermocouples and moveable incore flux detectors.

4. Since the core is expected to be X-Y stable, automatic protection against diametral transients is not required. However, an alarm function is provided to alert the operator to the existence of such tilts before a limiting value on diametral power tilt is reached.

Stability toward diametral oscillations was verified at startup. As burnup progresses, the reactor becomes increasingly stable toward diametral oscillations due to the decreasing soluble boron concentration and hence the continuously increasing moderator temperature coefficient feedback effect.

5. Control rod cluster malpositioning even under the most limiting case will not lead to a DNBR = 1.30 at operating conditions. Means for detecting such a misalignment are also provided.

3B.2 SPATIAL XENON STABILITY 3B.2.1 AXIAL XENON STABILITY The potential existence of axial power distribution anomalies due to xenon redistribution have been reported in WCAP-7208.7 Results of these studies have shown that the reactor will be unstable toward xenon oscillations in this dimension; consequently, power shaping devices (i.e., control rods) and automatic protection (i.e., trip setpoint reduction with excessive axial power imbalance) are provided. Operating philosophy and procedures for monitoring and controlling axial power anomalies have been described in References 7 and 8. The primary Chapter 3, Page 90 of 94 Revision 22, 2010

IP2 FSAR UPDATE

8. The primary means of detecting axial power distortions will be by means of the ex-core ion chambers with appropriate operator display signals. Tests in the Connecticut Yankee reactor (References 5 and 9) have verified the capability of these ex-core ion chambers to detect significant axial power imbalances.

3B.2.2 DIAMETRAL XENON STABILITY Results of the analytical investigations (primarily three-dimensional transient analyses reported in References 3 and 5) indicate that the Indian Point Unit 2 reactor will be stable toward diametral xenon oscillations; consequently, X-Y control rods are not required. Comparison with experimental results in the Connecticut Yankee reactor tend to confirm the validity of the less conservative calculations (see Reference 5, Figure 3-1). A test was performed at startup to demonstrate that artificially induced diametral oscillations decrease in amplitude as a function of time. Furthermore, extensive monitoring with appropriate display and alarm function is provided to alert the operators in the event a diametral power tilt should develop in the course of reactor operation. Consequently, no automatic safety protection against diametral xenon instability is required. 3B.2.3 ANALYTICAL TECHNIQUES In assessing potential power distribution anomalies arising from spatial xenon redistribution, primary reliance has been placed on time-dependent two-group diffusion calculations in three-dimensions including pointwise feedback effects due to coolant density and fuel pellet temperature changes. Means of incorporating the reactivity feedback effects are described in References 2 and 3 using semi-empirically fitted expressions whose coefficients were determined by other calculations (e.g., LEOPARD). In some cases, survey calculations were performed in one or two dimensions using both digital and modal techniques (see Reference 1), to indicate trends and to identify the significance and relative importance of the various contributing parameters. In performing three-dimensional time dependent stability analyses, standard design techniques (i.e., the LEOPARD Code) were used to compute the effect of the various feedback parameters on local reactivity. These results were fitted by a semi-empirical expression as described in Sections 2.2 and 3.3 of Reference 2. These analytical fits, with appropriate coefficients as determined from LEOPARD type calculations, were then used in the three-dimensional spatial power calculations, which included coupled thermal hydraulic effects. 3B.2.4 INSTRUMENTATION AND CONTROL Instrumentation and appropriate display is provided to ensure that the reactor will be maintained within thermal limits (design hot nuclear channel factors) in the presence of power distribution anomalies caused by time-dependent xenon redistribution. Primary reliance is placed on the eight ex-core ion chambers supplemented by information derived from the core exit thermocouples and from the movable incore fission chambers. The operator will have the ex-core detector information available, backed up by the core exit thermocouples and the movable incore detector readouts. The following ex-core detector information is provided for the operator to alert him to the existence of any core instabilities, axial or diametral: Chapter 3, Page 91 of 94 Revision 22, 2010

IP2 FSAR UPDATE

1. Four indicators, which indicate the difference between the top and bottom detectors. These signals will initiate alarms.

2. Eight indicators, which read out the individual currents of the four top and four bottom detectors.

3. One alarm for the four top detectors when the maximum to average flux is exceeded.

4. One alarm for the four bottom detectors when the maximum to average flux is exceeded.

5. Four 2-pen recorders; two detectors at 180 degrees are on the same recorder.

6. Two 2-pen recorders; power level proportional to total current, i.e., combined top and bottom detector outputs.

7. One total current deviation alarm, i.e., when any one top and bottom total current deviates by a pre-set amount from the other three total current outputs, the operator is alerted to this condition.

With these indications and alarms, the operator has many cross-checks and comparisons available to him. Failure of one top or bottom detector will provide the operator with instant indication and alarm. The ex-core detectors, backed by the movable incore detectors, provide more than adequate information. Operation with one ex-core ion chamber out of service does not compromise the safety of the plant. 3B.3 CONTROL ROD POSITIONING Normal control rod operations have been described in Section 3.0. A deviation in the position of one or more control clusters relative to the position of the control bank can potentially lead to:

1. Asymmetric fuel depletion.

2. Reduction in shutdown margin.

3. Reduction in DNB margin.

Rod misalignment is not a safety problem, which requires automatic protection because (1) asymmetric fuel depletion could possibly lead to unacceptable power distributions, but only if the condition were to persist for many hundreds of hours, (2) misalignment of sufficient magnitude to consume the standard 1-percent k shutdown is not possible, because it would require an entire control bank to be several feet below the desired position; the complete misalignment of a single control cluster will reduce trip reactivity by not more than 0.2-percent k; and (3) misalignment of a single control cluster by as much as the entire height of the core with the most pessimistic xenon spatial distribution will not result in a DNBR less than 1.30 at operating conditions. Deviation of 15-in. will not result in a power distribution worse than design. Chapter 3, Page 92 of 94 Revision 22, 2010

IP2 FSAR UPDATE Misalignment of a rod is most limiting when the last control group (which may be partly inserted at full power) is fully inserted but one cluster is full-out. It has been shown for Indian Point Unit 2 (Reference 5) that this case cannot lead to DNBR less than 1.30 at operating conditions even with the worst possible xenon distribution and the control bank (less one cluster) fully inserted. Each control cluster has its own position indicator channel. The rod position indicator channel is sufficiently accurate to detect a rod +/-7.5-in. away from its demand position for indicated control rod position less than or equal to 210 steps withdrawn. An indicated misalignment 12 steps does not exceed the power peaking factor limits. A misaligned rod of +17 steps allows for greater instrumentation error when indicated control rod position is greater than or equal to 211 steps withdrawn. The reactivity worth of a rod at this core height (211 + steps) is not sufficient to perturb power shapes to the extent that peaking factors are affected. The rod position indication system is the primary source of rod position information, but additional means, namely, ex-core ion chambers and movable incore fission chambers, are available. Except for the central control rod cluster, a power tilt will result from any significant control rod misalignment and such a power tilt would be detected by the ex-core ion chambers. Also, the movable incore fission chamber system can also be used to detect and/or investigate a suspected control rod malpositioning. REFERENCES FOR APPENDIX 3B

1. C. C. Poncelet and A. M. Christie, "Xenon-Induced Spatial Instabilities in Large Pressurized Water Reactors," WCAP-3680-20, Westinghouse Electric Corporation, March 1968.

2. F. B. Skogen and A. F. McFarlane, "Control Procedures for Xenon-Induced X-Y Instabilities in Large Pressurized Water Reactors," WCAP-3680-21, Westinghouse Electric Corporation, February 1969.

3. F. B. Skogen and A. F. McFarlane, "Xenon-Induced Spatial Instabilities in Three-Dimensions," WCAP-3680-22, Westinghouse Electric Corporation, September 1969.

4. A. M. Christie, et al., "Control of Xenon Instabilities in Large Pressurized Water Reactors," WCAP-3680-23, Westinghouse Electric Corporation, September 1969.

5. R. F. Barry, et al., "Power Maldistribution Investigations," WCAP-7407-L (Proprietary Class 2), Westinghouse Electric Corporation.

6. Letter from C. Newman, Con Ed, to AEC,

Subject:

Indian Point Unit No. 2 Results of the X-Y Xenon Stability Tests, dated October 17, 1974.

7. Westinghouse Electric Corporation, "Power Distribution Control of Westinghouse Pressurized Water Reactors," WCAP-7208 (APD Proprietary Class 2), September 1968.

8. T. Morita, et al., "Power Distribution Control and Load Following Procedures," WCAP-Chapter 3, Page 93 of 94 Revision 22, 2010

IP2 FSAR UPDATE WCAP-8385 (Proprietary), September 1974 and WCAP-8403 (Non-Proprietary), September 1974.

9. R. J. Johnson, "Connecticut-Yankee Tests on Detection of Power Maldistribution,"

WCAP-9010 (NES Proprietary Class 2), Westinghouse Electric Corporation, February 1969. Chapter 3, Page 94 of 94 Revision 22, 2010

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W-3 CORRELATION PROBABILITY DISTRIBUTION CURVE INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-41 TYPICAL W-3 CORRELATION PROBABILITY DISTRIBUTION CURVE MIC. No. 1999MC3678 REV. No. 17B

1.2...-------------------------------. LEGEND: L GRID DATA ba 8' HEATED LENGTH COS J.t 26" 1.0 o 14' HEATED LENGTH /ol SINE /ol26" o 14' HEATED LENGTH /ol SINE IJ. 26" COLDWALL

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().2 DNB QUALITY: -15.12 TO 13.341 O - - - - - . l l o - . ._ _--.llo-.. L....., ..I....- ..I....- -'--_ _- - - I I o 0.2 0.4 0.6 0.8 1.0 1.2 1.4 q" DNB (PREDICTED), 106 BTU/HR-FT2 INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-42 COMPARISON OF "L" GRID TYPICAL AND THIMBLE COLD WALL CELL ROD BUNDLE DNB DATA FOR NON-UNIFORM AXIAL HEAT FLUX WITH PREDICTIONS OF W-3 x F'SL MIC. No. 1999MC3679 REV. No. 17A

2 23 2004 11 :52:01 AM vnd 1999MC3680-0-17B.DWG

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MIC. No. 1999MC3681 REV. No. 17B

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CORE BAFFLE CORE BARREL THERMAL SHIELD REACTOR VESSEL FUEL ASSEMBLY 193 FUEL ASSEMBLIES INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-46 CORE CROSS SECTION MIC. No. 1999MC3684 REV. No. 17A

CONTROL ROO i-IJSI-tlii'j------- DRIVE MECHANISM INSTRUMENTAnON PORTS CLOSURE HEAD LIFTING LUG ASSEMBLY jf=H~::"'=J~t-J.-+:iu_THERMAL SLEEVE UPPER SUPPORT ~~---F~Q-l__CONTROlROO PLATE =-~+-t:~~ SHROUD TUBE INTERNALS ffi~~~~~~t"HOL[)"DOWNSPRING SUPPORT LEDGE ~~~~~\fr- ALIGNMENT PIN CORE BARREL 'r-,:DN~_ CONTROL ROO GUIDE TUBE SUPPORT COLUMN I1H;1--~ll_CONTROL ROO DRIVE SHAFT

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EN RICHMENTS 2.2 W!O 2.8 W!O 3.3 W!O INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-48 CORE LOADING ARRANGEMENT

                           - CYCLE 1 MIC. No. 1999MC3686      REV. No. 17A

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FUEL FlCOS/AUY .2£)4 Hote: (1) All dim. corrected to 68~F ~ 2° (2) Dimensions Ire expressed in inches INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-54 FUEL ASSEMBLY AND CONTROL CLUSTER CROSS SECTION - HIPAR, LOPAR, OFA AND VANTAGE+ MIC. No. 1999MC3708 REV. No. 17A

1.280 REf. I 2.250 DIA. IREfl 22.66 41.16 REf. L 59.66 REF. L. 76.54 REF. 133.66 93.41 REf.

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169._ 169.280 Note: Dimensions are expressed in inches. INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-55 HIPAR FUEL ASSEMBLY MIC. No. 1999MC3709 REV. No. 17A

159.71 -------------1 Note: Dimensions are expressed in inches. INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-56 LOPAR FUEL ASSEMBLY MIC. No. 1999MC3710 REV. No. 17A

OPTIMIZED (OFA) FUEL ASSEMBLY OFA RECONSTITUTABLE TOP NOZZLE: II 159.975 INCONEL GRID OFA ZIRC 134.07 GRIDS (7) 1/3.26 FUEL ROD LENGTH = 152.17 MID GRID HEIGHT = 2.25 93.70 76.77 59.84 41.28 23.47 INCONEL GRID 6.383 DEBRIS FILTER BOTTOM NOZZLE INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-56A OFA FUEL ASSEMBLY MIC. No. 1999MC3711 REV. No. 17A

INGONEL LPD ZIRe. GRIDS Ci) 133.70 t 1 122.78 i 112. BS J---+-- 1 I FIA (3) 10250 t 33."33 1 B~.35 t FUEL ROD LENGTH 152.55 lAID GRIQ HEIGHT 1.50

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• VANTAGE+ FUEL WITH PERFORMANCE + UFSAR FIGURE 3.2-568 FEATURES WILL HAVE ZIRLO' MIDGRIDS.

VANTAGE+ FUEL ASSEM8LY MIC. No. 1999MC3712 REV. No. 17A

_____ ZIRCALOY THIMBLE

                       ~        tZIRLOr\l forVANTAGE+l SPOT WELD
                        ~SSINSERT THIMBLE/END PLUG INERT*GAS FUSION WELD                               ZIRCALOY THIMBLE END PLUG INSERT/END PLUG WELD INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-57 GUIDE THIMBLE TO BOTTOM NOZZLE JOINT MIC. No. 1999MC3713            REV. No. 17A

BUTT WE LD ALL AROUND TOP NOZZLE ADAPTER PLATE STAINLESS STEEL SLEEVE EXPANSION LOBE TOP GRID INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-58 LOPAR TOP GRID TO NOZZLE ATTACHMENT MIC. No. 1999MC3714 REV. No. 17A

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TOP GRID TO NOZZLE ATTACHMENT MIC. No. 1999MC3715 REV. No. 17A

INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-59 SPRING CLIP GRID ASSEMBLY MIC. No. 1999MC3716 REV. No. 17A

zr;CALCY SLEEVE THI!'18LE:

                        ~UEL  ROO EXPANSION LOaE INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-60 MID-GRID EXPANSION JOINT DESIGN PLAN VIEW MIC. No. 1999MC3717           REV. No. 17A

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                                      ~IO  G~IO EXPANSION LOBE ~~~~~~-ZIRC SLEEVE THIMBLE INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-61 A ELEVATION VIEW - VANTAGE+

GRID TO THIMBLE ATTACHMENT MIC. No. 1999MC3719 REV. No. 17A

3 2 2004 6:50:50 AM vnd1999MC3720-0-17B.DWG o INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-61 B VANTAGE+ FUEL ASSEMBLY WITH PERFORMANCE+ ENHANCEMENTS MIC. No. 1999MC3720 REV. No. 17B

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1

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INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-63 HIPAR BURNABLE POISON ROD MIC. No. 1999MC3722 REV. No. 17A

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CABLE CONN'eCTION 10.71B" X 10.718" OPERATING COIL STACK ASSEMBLY PRESSURE HOUSING

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BARcd Cladding Zone Number Pnmous Design SA WABA Desiqn 0*1 AU' Wamr t(WdJ 1 - 2 St4liniess steef Zin:aIOV 2*3 Air Helium !. n **** , 3*4 Borosilicate glass .41203 - S4 C 4-5 Ai,. HeJium V?ZYa 5-6 Stainless steej Zin:aJoy INDIAN POINT UNIT No. 2 UFSAR FIGURE 3.2-69 COMPARISON OF BOROSILICATE GLASS ABSORBER ROD WITH WABA ROD MIC. No. 1999MC3730 REV. No. 17A

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IP2 FSAR UPDATE CHAPTER 4 REACTOR COOLANT SYSTEM 4.0 GENERAL DESCRIPTION The reactor coolant system, shown in Plant Drawing 9321-2738 [Formerly UFSAR Figure 4.2-1], consists of four similar heat transfer loops connected in parallel to the reactor vessel. Each loop contains a reactor coolant circulation pump and a steam generator. The system also includes a pressurizer, pressurizer relief tank, connecting piping, and instrumentation necessary for operational control. 4.1 DESIGN BASES 4.1.1 Performance Objectives The reactor coolant system transfers the heat generated in the core to the steam generators where steam is generated to drive the turbine generator. Demineralized light water is circulated at the flow rate and temperature consistent with achieving the reactor core thermal-hydraulic performance presented in Section 3.2. The water also acts as a neutron moderator and reflector and as a solvent for the neutron absorber used in chemical shim control. The reactor coolant system provides a boundary for containing the coolant under operating temperature and pressure conditions. It serves to confine radioactive material and limits any uncontrolled release to the secondary system and to other parts of the plant to acceptable values under conditions of either normal or abnormal reactor behavior. During transient operation, the system heat capacity attenuates thermal transients generated by the core or extracted by the steam generators. The reactor coolant system accommodates coolant volume changes within the protection system criteria. By appropriate selection of the inertia of the reactor coolant pumps, the thermal-hydraulic effects are reduced to a safe level during the pump coastdown that would result from a loss-of-offsite power situation. The layout of the system ensures the natural circulation capability following a loss of offsite power, to permit decay heat removal without overheating the core. Portions of the system piping are used by the safety injection system to deliver cooling water to the core during a loss-of-coolant accident. 4.1.2 General Design Criteria General design criteria (GDC) that apply to the reactor coolant system are given below. 4.1.2.1 Quality Standards Criterion: Those systems and components of reactor facilities, which are essential to the prevention, or the mitigation of the consequences, of nuclear accidents, which could cause undue risk to the health and safety of the public shall be identified and then designed, fabricated, and erected to quality standards that reflect the importance of the safety function to be performed. Where generally recognized codes and standards pertaining to design, materials, fabrication, and inspection are used, they shall be identified. Where adherence to such codes or standards does not suffice to assure a quality product in keeping with the safety function, they shall be supplemented or modified as necessary. Quality assurance Chapter 4, Page 1 of 85 Revision 22, 2010

IP2 FSAR UPDATE programs, test procedures, and inspection acceptance criteria to be used shall be identified. An indication of the applicability of codes, standards, quality assurance programs, test procedures, and inspection acceptance criteria used is required. Where such items are not covered by applicable codes and standards, a showing of adequacy is required. (GDC 1) The reactor coolant system is of primary importance with respect to its safety function in protecting the health and safety of the public. Quality standards of material selection, design, fabrication, and inspection conform to the applicable provisions of recognized codes and good nuclear practice (Section 4.1.7). Details of the quality assurance programs, test procedures, and inspection acceptance levels are given in Sections 4.3.1 and 4.5. Particular emphasis is placed on the assurance of quality of the reactor vessel to obtain material whose properties are uniformly within tolerances appropriate to the application of the design methods of the code. 4.1.2.2 Performance Standards Criterion: Those systems and components of reactor facilities, which are essential to the prevention or to the mitigation of the consequences of nuclear accidents, which could cause undue risk to the health and safety of the public shall be designed, fabricated, and erected to performance standards that will enable such systems and components to withstand, without undue risk to the health and safety of the public, the forces that might reasonably be imposed by the occurrence of an extraordinary natural phenomenon such as earthquake, tornado, flooding condition, high wind, or heavy ice. The design bases so established shall reflect: (a) appropriate consideration of the most severe of these natural phenomena that have been officially recorded for the site and the surrounding area and (b) an appropriate margin for withstanding forces greater than those recorded to reflect uncertainties about the historical data and their suitability as a basis for design. (GDC 2) All piping, components, and supporting structures of the reactor coolant system are designed to Class I requirements, as discussed in Sections 1.11 and 4.1.4.3. The reactor coolant system is located in the containment whose design, in addition to being a Class I structure, also considers accidents or other applicable natural phenomena. Details of the containment design are given in Chapter 5. 4.1.2.3 Records Requirements Criterion: The reactor licensee shall be responsible for assuring the maintenance throughout the life of the reactor of records of the design, fabrication, and construction of major components of the plant essential to avoid undue risk to the health and safety of the public. (GDC 5) Records of the design of the major reactor coolant system components and the related engineered safety features components are maintained for the life of the plant. Records of fabrication are maintained in the manufacturers' plants as required by the appropriate code or other requirements pending submittal to Westinghouse or ENIP2. They are Chapter 4, Page 2 of 85 Revision 22, 2010

IP2 FSAR UPDATE available at any time throughout the life of the plant. Records of changes made to the plant as described in the FSAR are maintained for the life of the plant. 4.1.2.4 Missile Protection Criterion: Adequate protection for those engineered safety features, the failures of which could cause an undue risk to the health and safety of the public, shall be provided against dynamic effects and missiles that might result from plant equipment failures. (GDC 40) 4.1.2.4.1 Original Design Basis The dynamic effects during blowdown following a loss-of-coolant accident are evaluated in the detailed layout and design of the high-pressure equipment and barriers that afford missile protection. Support structures are designed with consideration given to fluid and mechanical thrust loadings. Original plant design basis required that the steam generators be supported, guided, and restrained in a manner that prevents rupture of the steam side of a generator, the steam lines, and the feedwater piping as a result of forces created by a reactor coolant system pipe rupture. These supports, guides, and restraints also prevent rupture of the primary side of a steam generator as a result of forces created by a steam or feedwater line rupture. Original plant design basis also required that the mechanical consequences of a pipe rupture as a result of forces created by a reactor coolant system pipe rupture be restricted by design such that the functional capability of the engineered safety features would not be impaired. 4.1.2.4.2 Revised Design Basis In 1989, the NRC approved elimination of the necessity for considering and protecting against dynamic effects of postulated primary loop pipe ruptures from the design basis of Indian Point Unit 2. "Leak before break" technology was applied as permitted by revised General Design Criterion 4 of 10CFR50, Appendix A. References 1, 2, 3 and 4 contain further information. With the elimination of the necessity for considering and protecting against the dynamic effects of postulated primary loop ruptures from the design basis of Indian Point 2, breaks have been postulated in the following branch lines: the Accumulator branch line in the cold leg, the Pressurizer Surge line and the Residual Heat Removal (RHR) line in the hot leg. This is discussed in Reference 5. These breaks are the new design basis breaks with respect to considering and protecting against the dynamic effects of postulated ruptures. In general, these new breaks are significantly less severe than the original design basis breaks. The following is a description of how the dynamic effects of these new breaks have been considered and protected against. The dynamic effects have been grouped into three categories as follows:

1. Containment subcompartment pressurization

2. Break reaction forces (i.e. forcing function analysis, asymmetric blowdown loading) used in structural support design and for confirming the structural integrity of systems and components

3. Missile Protection (pipe whip, jet impingement and missiles)

Chapter 4, Page 3 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.1.2.4.2.1 Containment Subcompartment Pressurization The dynamic effects of these new breaks with respect to containment subcompartment pressurization is discussed in Section 14.3.5.4.3.2. 4.1.2.4.2.2 Break Reaction Forces As discussed in Reference 5, the new break locations were used to develop hydraulic forcing functions for revised structural analyses. These analyses demonstrate that for the new break locations the structural integrity of the reactor vessel internals, core components including fuel assemblies, and the reactor coolant loop will be maintained and will preserve the ability to maintain a coolable geometry for the rated stretch power conditions. The dynamic effects of postulated breaks with respect to break reaction forces is discussed further in Sections 1.3.7, 1.11.3, 4.2.4, 4.3.1.3, Appendix 4B.7, 5.1.1.1.5, 5.1.5, 6.1.1.5, and 6.2.2.5. 4.1.2.4.2.3 Missile Protection Since the new break locations remain inside the missile barrier, engineered safety features and associated systems remain protected from loss of function due to dynamic effects and missiles, which might result from these breaks. This protection is discussed further in Sections 1.3.7, 4.2.4, 5.1.2.5, 6.1.1.3, 6.2.2.5, and 8.2.2.6. 4.1.3 Principal Design Criteria The criteria that apply solely to the reactor coolant system are given below. 4.1.3.1 Reactor Coolant Pressure Boundary Criterion: The reactor coolant pressure boundary shall be designed, fabricated and constructed so as to have an exceedingly low probability of gross rupture of significant uncontrolled leakage throughout its design lifetime. (GDC 9) The reactor coolant system in conjunction with its control and protective provisions is designed to accommodate the system pressures and temperatures attained under all expected modes of plant operation or anticipated system interactions and to maintain the stresses within applicable code stress limits. Fabrication of the components that constitute the pressure retaining boundary of the reactor coolant system is carried out in strict accordance with the applicable codes. In addition, there are areas where equipment specifications for reactor coolant system components go beyond the applicable codes. Details are given in Section 4.5.1. The materials of construction of the pressure retaining boundary of the reactor coolant system are protected by control of coolant chemistry from corrosion phenomena that might otherwise reduce the system's structural integrity during its service lifetime. System conditions resulting from anticipated transients or malfunctions are monitored, and appropriate action is automatically initiated to maintain the required cooling capability and to limit system conditions so that continued safe operation is possible. Chapter 4, Page 4 of 85 Revision 22, 2010

IP2 FSAR UPDATE The system is protected from overpressure by means of pressure-relieving devices as required by Section III of the ASME Boiler and Pressure Vessel Code. Isolated sections of the system are provided with overpressure-relieving devices discharging to closed systems such that the system code allowable relief pressure within the protected section is not exceeded. 4.1.3.2 Monitoring Reactor Coolant Leakage Criterion: Means shall be provided to detect significant uncontrolled leakage from the reactor coolant pressure boundary. (GDC 16) Positive indications in the control room of leakage of coolant from the reactor coolant system to the containment are provided by equipment that permits continuous monitoring of containment air activity and humidity and of runoff from the condensate-collecting pans under the cooling coils of the containment air recirculation units. This equipment provides indication of normal environmental conditions within the containment. Any increase in the observed parameters could be an indication of change within the containment, and the equipment provided is capable of monitoring this change. The basic design criterion is the detection of deviations from normal containment environmental conditions including air particulate activity, radiogas activity, humidity, condensate runoff, and in the case of significant leakage, the liquid inventory in the process systems and containment sump. Further details are supplied in Sections 4.2.7 and 6.7. 4.1.3.3 Reactor Coolant Pressure Boundary Capability Criterion: The reactor coolant pressure boundary shall be capable of accommodating without rupture the static and dynamic loads imposed on any boundary component as a result of an inadvertent and sudden release of energy to the coolant. As a design reference, this sudden release shall be taken as that which would result from a sudden reactivity insertion such as rod ejection (unless prevented by positive mechanical means), rod dropout, or cold water addition. (GDC 33) The reactor coolant boundary is shown to be capable of accommodating without further rupture the static and dynamic loads imposed as a result of a sudden reactivity insertion such as a rod ejection. Details of this analysis are provided in Section 14.2.6.10. The operation of the reactor is such that the severity of an ejection accident is inherently limited. Since control rod clusters are used to control load variations only and core depletion is followed with boron dilution, only the rod cluster control assemblies in the controlling groups are inserted in the core at power, and at full power these rods are only partially inserted. A rod insertion limit monitor is provided as an administrative aid to the operator to ensure that this condition is met. By using the flexibility in the selection of control rod groupings, radial locations and positions as a function of load, the design limits the maximum fuel temperature for the highest worth ejected rod to a value that precludes any resultant damage to the primary system pressure boundary, i.e., gross fuel dispersion in the coolant and possible excessive pressure surges. The failure of a rod mechanism housing causing a rod cluster to be rapidly ejected from the core is evaluated as a theoretical, though not a credible, accident. While limited fuel damage could result from this hypothetical event, the fission products are confined to the reactor coolant Chapter 4, Page 5 of 85 Revision 22, 2010

IP2 FSAR UPDATE system and the reactor containment. The environmental consequences of rod ejection are less severe than from the hypothetical loss of coolant, for which public health and safety are shown to be adequately protected. Refer to Section 14.2.6. 4.1.3.4 Reactor Coolant Pressure Boundary Rapid Propagation Failure Prevention Criterion: The reactor coolant pressure boundary shall be designed and operated to reduce to an acceptable level the probability of rapidly propagating type failure. Consideration is given (a) to the provisions for control over service temperature and irradiation effects, which may require operational restrictions, (b) to the design and construction of the reactor pressure vessel in accordance with applicable codes, including those, which establish requirements for absorption of energy within the elastic strain energy range and for absorption of energy by plastic deformation and (c) to the design and construction of reactor coolant pressure boundary piping and equipment in accordance with applicable codes. (GDC 34) The reactor coolant pressure boundary is designed to reduce to an acceptable level the probability of a rapidly propagating type failure. In the core region of the reactor vessel it is expected that the notch toughness of the material will change as a result of fast neutron exposure. This change is evidenced as a shift in the nil-ductility transition temperature (NDTT), which is factored into the operating procedures in such a manner that full operating pressure is not obtained until the affected vessel material is above the design transition temperature (DTT) in the ductile material region. The pressure during startup and shutdown at the temperature below NDTT is maintained below the threshold of concern for safe operation. The DTT is a minimum of NDTT plus 60oF and dictates the procedures to be followed in the hydrostatic test and in station operations to avoid excessive cold stress. The value of the DTT is increased during the life of the plant as required by the expected shift in NDTT and as confirmed by the experimental data obtained from irradiated specimens of reactor vessel materials during the plant lifetime. Further details are given in Section 4.1.6. All pressure-containing components of the reactor coolant system are designed, fabricated, inspected, and tested in conformance with the applicable codes. Further details are given in Section 4.1.7. 4.1.3.5 Reactor Coolant Pressure Boundary Surveillance Criterion: Reactor coolant pressure boundary components shall have provisions for inspection, testing, and surveillance of critical areas by appropriate means to assess the structural and leaktight integrity of the boundary components during their service lifetime. For the reactor vessel, a material surveillance program conforming with current applicable codes shall be provided. (GDC 36) The design of the reactor vessel and its arrangement in the system provides the capability for accessibility during service life to the entire internal surfaces of the vessel including the nozzle to reactor coolant piping welds and the top and bottom heads. The reactor arrangement within the containment provides sufficient space for inspection of the external surfaces of the reactor coolant piping, except for the area of pipe within the primary shielding concrete. Monitoring of the nil-ductility transition temperature properties of the core region plates, forgings, weldments, and associated heat-treated zones is performed in accordance with ASTM E185 Chapter 4, Page 6 of 85 Revision 22, 2010

IP2 FSAR UPDATE (Recommended Practice for Surveillance Tests on Structural Materials in Nuclear Reactors). Samples of reactor vessel plate materials are retained and cataloged in case further engineering development shows the need for further testing. The material properties surveillance program includes not only the conventional tensile and impact tests, but also fracture mechanics specimens. The fracture mechanics specimens are the wedge opening loading type specimens. The observed shifts in nil ductility transition temperature of the core region materials with irradiation are used to confirm the calculated limits to startup and shutdown transients. To define permissible operating conditions below the design transition temperature, a pressure range is established, which is bounded by a lower limit for pump operation and an upper limit, which satisfies reactor vessel stress criteria. To allow for thermal stresses during heatup or cooldown of the reactor vessel, an equivalent pressure limit is defined to compensate for thermal stress as a function of rate of change of coolant temperature. Since the normal operating temperature of the reactor vessel is well above the maximum expected design transition temperature, brittle fracture during normal operation is not considered to be a credible mode of failure. 4.1.4 Design Characteristics 4.1.4.1 Design Pressure The reactor coolant system design and operating pressure together with the safety, power relief, and pressurizer spray valves setpoints and the protection system setpoint pressures are listed in Table 4.1-1. The design pressure allows for operating transient pressure changes. The selected design margin considers core thermal lag, coolant transport times and pressure drops, instrumentation and control response characteristics, and system relief valve characteristics. The design pressures and data for the respective system components are listed in Tables 4.1-2 through 4.1-6. Table 4.1-7 gives the design pressure drop of the system components. 4.1.4.2 Design Temperature The design temperature for each component is selected to be above the maximum coolant temperature in that component under all normal and anticipated transient load conditions. The design and operating temperatures of the respective system components are listed in Tables 4.1-2 through 4.1-6. 4.1.4.3 Seismic Loads The seismic loading conditions are established by the "design earthquake" and "maximum potential earthquake." The former is selected to be typical of the largest probable ground motion based on the site seismic history. The latter is selected to be the largest potential ground motion at the site according to seismic and geological factors and their uncertainties. For the design earthquake loading condition, the nuclear steam supply system is designed to be capable of continued safe operation. Therefore, for this loading condition critical structures and equipment needed for this purpose are required to operate within normal design limits. The seismic design for the maximum potential earthquake is intended to provide a margin in design that ensures capability to shut down and maintain the nuclear facility in a safe condition. In this case, it is only necessary to ensure that the reactor coolant system components do not lose Chapter 4, Page 7 of 85 Revision 22, 2010

IP2 FSAR UPDATE their capability to perform their safety function. This has come to be referred to as the "no-loss-of function" criteria and the loading condition as the "no-loss-of-function earthquake" loading condition. The criteria adopted for allowable stresses and stress intensities in vessels and piping subjected to normal loads plus seismic loads are defined in Section 1.11. These criteria ensure the integrity of the reactor coolant system under seismic loading. For the combination of normal and design earthquake loadings, the stresses in the support structures are kept within the limits of the applicable codes. For the combination of normal and no-loss-of-function earthquake loadings, the deflections and stresses in the support structures are limited to values as necessary to ensure their integrity and to maintain supported equipment within their stress limits as stated in Table 1.11-2. 4.1.5 Cyclic Loads All components in the reactor coolant system are designed to withstand the effects of cyclic loads due to reactor system temperature and pressure changes. These cyclic loads are introduced by normal unit load transients, reactor trip, and startup and shutdown operations. The number of thermal and loading cycles used for design purposes and the bases thereof are given in Table 4.1-8. There is a station program, which tracks these thermal and loading cycles. During unit startup and shutdown, the rates of temperature and pressure changes are limited as indicated in Section 4.4.1. The cycles are estimated for equipment design purposes (40-year life) and are not intended to be an accurate representation of actual transients or actual operating experience. For example, the number of cycles for plant heatup and cooldown at 100oF per hr was selected as a conservative estimate based on an evaluation of the expected requirements. The resulting number, which averages five heatup and cooldown cycles per year, could be increased significantly; however, it is the intent to represent a conservative realistic number rather than the maximum allowed by the design. Although loss-of-flow and loss-of-load transients are not included in the tabulation because the tabulation is only intended to represent normal design transients, the effects of these transients have been analytically evaluated and are included in the fatigue analysis for primary system components. Over the range from 15-percent full power up to and including but not exceeding 100-percent of full power, for the purpose of cyclic load definition, the reactor coolant system and its components are designed to accommodate 10-percent of full power step changes in plant load and 5-percent of full power per minute ramp changes without reactor trip. The reactor coolant system will accept a complete loss of load from full power with reactor trip. In addition, the turbine bypass and steam dump system makes it possible to accept a step load decrease of 50-percent of full power without reactor trip. These transient capability definitions bracket the transient design bases used for the Regulating Systems as discussed in Section 7.3. 4.1.6 Service Life The service life of reactor coolant system pressure components depends upon the end-of-life material radiation damage, unit operational thermal cycles, quality manufacturing standards, environmental protection, and adherence to established operating procedures. Chapter 4, Page 8 of 85 Revision 22, 2010

IP2 FSAR UPDATE The reactor vessel is the only component of the reactor coolant system that is exposed to a significant level of neutron irradiation and it is therefore the only component that is subject to material radiation damage effects. The nil-ductility transition temperature shift of the vessel material and welds, due to radiation damage effects, is monitored by a radiation damage surveillance program, which conforms with ASTM E185 standards. Reactor vessel design is based on the transition temperature method of evaluating the possibility of brittle fracture of the vessel material, as a result of operations such as leak testing and plant heatup and cooldown. To establish the service life of the reactor coolant system components as required by the ASME (Section III) Boiler and Pressure Vessel Code for Class "A" vessels, the unit operating conditions have been established for the 40-year design life. These operating conditions include the cyclic application of pressure loadings and thermal transients. The number of thermal and loading cycles used for design purposes are listed in Table 4.1-8. 4.1.7 Codes And Classifications The quality assurance criteria specified below apply to all nuclear Class I piping and fittings. All pressure-containing components of the reactor coolant system are designed, fabricated, inspected, and tested in conformance with the applicable codes listed in Table 4.1-9. Shop and field fabrication requirements, documentation, and quality assurance examinations all comply with those found in USAS B31.7 for Class I nuclear piping. Quality control techniques used in the fabrication of the reactor coolant system are equivalent to those used in the manufacture of the reactor vessel, which conforms to Section III of the ASME Boiler and Pressure Vessel Code. The piping is designed to the USAS B31.1 (1955 and Summer 1973) Code for Power Piping using the allowable stresses found in the Nuclear Code Cases N-7 and N-10 for pipe and fittings, respectively. The quality assurance requirements required by Westinghouse in the purchase and examination of the reactor coolant piping ensures that the quality level of a Westinghouse plant is comparable to that delineated by USAS B31.7, Class I, Code for Nuclear Piping.

1. All materials conform to ASTM specifications listed for B31.7 Class I, Nuclear Piping. In addition, all materials are certified, identified, and marked to facilitate traceability thus complying with the requirements of USAS B31.7, Class I, Code for Nuclear Piping.

2. Piping base materials are examined by quality assurance methods having acceptance criteria that meet the requirements set forth in USAS B31.7, Class I, Code for Nuclear Piping.

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IP2 FSAR UPDATE

3. All welding procedures, welders, and welding operators are qualified to the requirements of ASME Section IX, Welding Qualifications, which is in compliance with the requirements of USAS B31.7, Class I, Code for Nuclear Piping.

4. All welds are examined by nondestructive testing methods and to the extent prescribed in USAS B31.7 for Class I nuclear piping.

5. All branch connection nozzle welds of nominal sizes of 3-in. and larger are 100-percent radiographed. This exceeds the requirements of USAS B31.7 for Class I piping, since it includes nominal sizes of 6-in. and larger for 100-percent radiography.

6. All finished welds are liquid penetrant examined on both the outside and inside (if accessible) surfaces as required by USAS B31.7, Class I. In addition, nozzle welds in nominal sizes 2-in. and smaller are progressively examined after each 0.25-in. increment of weld deposit in lieu of radiography.

7. Hydrostatic testing is performed on the erected and installed piping. This requirement is the same as in USAS B31.7, Class I.

Hence, the Westinghouse quality assurance requirements implemented in the procurement of Indian Point Unit 2 piping and fittings are equal to and in some instances exceed the requirements of USAS B31.7. The reactor coolant system is classified as Class I for seismic design, requiring that there will be no loss of function of such equipment in the event of the assumed maximum potential ground acceleration acting in the horizontal and vertical directions simultaneously, when combined with the primary steady state stresses. The design and stress criteria specified in USAS B31.7 are not directly comparable to that of USAS B31.1 (1955 and Summer 1973). The following describes how USAS B31.1 (1955 and Summer 1973) was used in the design of the primary coolant piping and the ASME B&PV Code Section III, Subsection NB, 1986 Edition for the presurizer surge line including the effects of Thermal Stratification on Indian Point Unit 2. A thermal expansion flexibility stress analysis was performed on the main primary coolant piping and pressurizer surge line (including the effects of Thermal Stratification) in accordance with the criteria set forth in USAS B31.1 (1955 and Summer 1973) for the reactor coolant piping and the ASME B&PV Code Section III 1986 Edition for the pressurizer surge line including the effects of Thermal Stratification. For the reactor coolant piping the analysis was performed to ensure that the stress range is within the limits prescribed in B31.1. As per the requirements of USAS B31.1, no fatigue analysis is required and hence, no fatigue analysis of the reactor coolant loop piping is performed. For the pressurizer surge line including the effects of Thermal Stratification, the analysis was performed to ensure that the stress range and number of thermal cycles (usage factor) are safely within the limits prescribed in ASME B&PV Code Section III, Subsection NB, 1986 Edition. In addition, seismic analysis were performed on the composite piping, which included the combined stress effects of all the sustained (pressure and weight) loadings plus seismic vertical / horizontal loading components. The resultant reactions of the piping due to the separate and combined effects of thermal, sustained, and seismic loadings were factored into the piping as interconnected. In turn, the equipment supporting structures were checked for adequate design including the added effects of these same loadings. Thus the total design analyses including Chapter 4, Page 10 of 85 Revision 22, 2010

IP2 FSAR UPDATE pipe, equipment, and structures considered the effects of thermal expansion, sustained, and seismic loadings. Thermally induced stresses arising from temperature gradients are limited to a safe and low order of magnitude in assigning a maximum permissible time rate of temperature change on plant heatup, cooldown, and incremental loadings in the plant operation procedure. An added margin of conservatism is obtained through the use of thermal sleeves in nozzles wherein a cold fluid is introduced into a pipe conveying a significantly hotter fluid or vice versa. Typical examples are the charging line, pressurizer surge line, and residual heat return nozzle connections to the primary coolant loop piping. The thermal sleeve is no longer in place on the 10" SI line to the 23 Cold Leg. A detailed analysis demonstrated that the fatigue usage factor and stresses for the nozzle in line 353 still meet the requirements of ASME Section III of the Boiler and Pressure Vessel Code for continued operation through the life of the plant with the thermal sleeve not in place. REFERENCES FOR SECTION 4.1

1. Letter from Stephen B. Bram, Con Edison, to Document Control Desk, NRC,

Subject:

Leak-Before-Break, dated May 23, 1988.

2. Letter from Stephen B. Bram, Con Edison, to Document Control Desk, NRC,

Subject:

Leak-Before-Break (LBB) Submittal (TAC 68318), dated November 18, 1988.

3. Letter from Stephen B. Bram, Con Edison, to Document Control Desk,

Subject:

Leak-Before-Break (LBB) Submittal (TAC 68318), dated January 12, 1989.

4. Letter from Donald S. Brinkman, NRC, to Stephen B. Bram, Con Edison,

Subject:

Safety Evaluation Report on Elimination of Dynamic Effect of Postulated Primary Loop Pipe Ruptures from Design Basis for Indian Point Unit 2 (TAC No. 68318), dated February 23, 1989.

5. WCAP-12187 - Consolidated Edison Company of New York, Inc., Indian Point Unit 2, NSSS Stretch Rating - 3083.4 MWT, Engineering Report, Volume II. -

Westinghouse Proprietary Class 2 Chapter 4, Page 11 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.1-1 Reactor Coolant System Pressure Settings Category Pressure (psig) Design pressure 2485 Operating pressure (at pressurizer) 22351 Safety valves 24851 Power relief valves 23351 Pressurizer spray valves (open) 22601 High pressure trip 23631 High pressure alarm 2300/23351,2 Low pressure trip 19281 Low pressure alarm 21851 Hydrostatic test pressure 3110 Notes:

1. Nominal values

2. The fixed high alarm PC-456F is a redundant alarm, and will not annunciate unless there is a failure of the 2300 psig alarm/control circuit.

TABLE 4.1-2 Reactor Vessel Design Data Design/operating pressure, psig 2485/2235 Hydrostatic test pressure, psig 3110 Design temperature, °F 650 Overall height of vessel and closure head, ft-in. (bottom head OD to top of control rod mechanism housing) 43-9 11/16 Water volume, (with core and internals in place), ft3 4647 Thickness of insulation, min., in. 3 Number of reactor closure head studs 54 Diameter of reactor closure head studs, in. 7 ID of flange, in. 167 1/16 OD of flange, in. 205 ID at shell, in. 173 Inlet nozzle ID, in. 27 1/2 Outlet nozzle ID, in. 29 Clad thickness, min., in. 5/32 Lower head thickness, min., in. 5 5/16 Vessel belt-line thickness, min., in. 8 5/8 Closure head thickness, in. 7 Reactor coolant inlet temperature, °F 514.31 Reactor coolant outlet temperature, °F 605.81 Reactor coolant flow, lb/hr 1.268 x 108 Notes:

1. Reactor Coolant inlet temperature is for the low Tavg case and Reactor Coolant outlet temperature is for the high Tavg case.

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IP2 FSAR UPDATE TABLE 4.1-3 Pressurizer and Pressurizer Relief Tank Design Data Pressurizer Design/operating pressure, psig 2485/2235 Hydrostatic test pressure (cold), psig 3110 Design/operating temperature, °F 680/653 Water volume, full power, ft3 10801 Steam volume, full power, ft3 7201 Surge line nozzle diameter, in./pipe schedule 14/Sch 140 Shell ID, in./calculated minimum shell thickness, in. 84/4.1 Minimum clad thickness, in. 0.188 Electric heaters capacity, kW 1800 Heatup rate of pressurizer using heaters only, °F/hr 55 (approximately) Power relief valves Number 2 Set pressure (open), psig 2335 Capacity, lb/hr saturated steam per valve 179,000 Safety valves Number 3 Set pressure, psig2 2485 Capacity, lb/hr saturated steam per valve 408,000 Pressurizer Relief Tank Design pressure, psig 100 Rupture disc release pressure, psig 100 Design temperature, °F 340 Normal water temperature, °F Containment ambient Total volume, ft3 1800 Rupture disc relief capacity, lb/hr 1.224 x 106 Notes:

1. Present operation is at a Tavg of 562°F. In the safety analysis discussed in section 14.1, a reduced flow is assumed to account for a postulated 25% steam generator tube plugging.

Actual values will depend on Tavg and actual percentage of tube plugging.

2. Allowance for error is specified in the Technical Specifications.

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IP2 FSAR UPDATE TABLE 4.1-4 Steam Generator Design Data Number of steam generators 4 Design pressure, reactor coolant/steam, psig 2485/1085 Reactor coolant hydrostatic test pressure (tube side-cold), psig 3107 Design temperature, reactor coolant/steam, °F 650/556 Reactor coolant flow, Thermal Design, gpm/loop 80,700 Total heat transfer surface area, ft2 43,467 Heat transferred, Btu/hour 2755 x 106 Steam conditions at full load, outlet nozzle: Low Tavg

• High Tavg
• Steam flow, lb/hr 3.50 x 106 3.51 x106 Steam temperature, °F 488.0 513.3 Steam pressure, psia 610.1 766.3 Feedwater temperature, °F 436.2 436.2 Overall height, ft-in. 63-1.625 Shell OD, upper/lower, in. 166/127.0 Shell thickness, upper/lower, in. 3.5/2.63 Number of U-tubes 3214 U-tube diameter, in. 0.875 Tube wall thickness, (average), in. 0.050 Number of manways/ID, in. 4/16 Number of handholes/ID, in. 6/6 Number of Inspection Openings/ID, in. 1/3 3230 MWt Zero Power Low Tavg/High Tavg*

Reactor coolant water volume (unplugged), ft3 924 924 Primary side fluid heat content, Btu 23.67 x 106 / 23.630 x 106 24.11 x 106 Secondary side water volume, ft3 1493/1599 2778.5 Secondary side steam volume, ft3 32434/3128 1949 Secondary side fluid heat content, Btu 39.96 x 106 / 75.31 x 106 Note:

• Refers to low (548.9 deg F) and high (571.9 deg F) Tavg and 0% tube plugging cases for design.

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IP2 FSAR UPDATE TABLE 4.1-5 Reactor Coolant Pumps Design Data Number of pumps 4 Design pressure/operating pressure, psig 2485/2235 Hydrostatic test pressure (cold), psig 3110 Design temperature (casing), °F 650 RPM at nameplate rating 1189 Suction temperature, °F 555 1 Net positive suction head, ft 170 1 Developed head, ft 272 1 Capacity, gpm 89,700 1 Seal water injection, gpm 8 Seal water return, gpm 3 Pump discharge nozzle ID, in. 27 1/2 Pump suction nozzle ID, in. 31 Overall unit height, ft 28.38 Water volume, ft3 192 Pump-motor moment of inertia, lb/ft2 82,000 Motor Data: Type AC induction, single speed, air cooled Voltage 6600 Insulation class B thermoplastic epoxy Phase 3 Frequency, cps 60 Starting Current, amp 2950 1 Input (hot reactor coolant), kW 4221 1 Input (cold reactor coolant), kW 5673 Power, HP (nameplate) 6000 Note:

1. These values represent the pump hydraulic design point. Actual heads, flows, temperatures, currents, and powers are dependant upon system parameters such as reactor internals changes, percentage of steam generator tube plugging, and plant operating Tavg. For use in analyses or evaluations, values reflecting the current conditions should be obtained.

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IP2 FSAR UPDATE TABLE 4.1-6 Reactor Coolant Piping Design Data Reactor inlet piping ID, in. 27 1/2 Reactor inlet piping nominal thickness, in. 2.375 Reactor outlet piping ID, in. 29 Reactor outlet piping nominal thickness, in. 2.50 Coolant pump suction piping ID, in. 31 Coolant pump suction piping nominal thickness, in. 2.656 Pressurizer surge line piping ID, in. 11.5 Pressurizer surge line piping nominal thickness, in. 1.25 Design/operating pressure, psig 2485/2235 Hydrostatic test pressure, (cold) psig 3110 Design temperature, °F 650 Design temperature, (pressurizer surge line) °F 680 3 Water volume, (all 4 loops including surge line) ft 1156 TABLE 4.1-7 Reactor Coolant System Design Pressure Drop Pressure Drop (psi) Across pump discharge leg 1.2 Across vessel, including nozzles 51.5 Across hot leg 1.1 Across steam generator 31.8 Across pump suction leg 2.8 Total pressure drop 88.4 Notes:

1) DPs based on best estimate flow conditions.

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IP2 FSAR UPDATE TABLE 4.1-8 Thermal and Loading Cycles Transient Condition Design Cycles1

1. Plant heatup at 100°F per hr 200 (5/yr2)

2. Plant cooldown at 100°F per hr 200 (5/yr)

3. Plant loading at 5-percent of full power per min 14,500 (1/day)

4. Plant unloading at 5-percent of full power per min 14,500 (1/day)

5. Step load increase of 10-percent of full power (but not to exceed full power) 2000 (1/wk)

6. Step load decrease of 10-percent of full power 2000 (1/wk)

7. Step load decrease of 50-percent of full power 200 (5/yr)

8. Reactor trip 400 (10/yr)

9. Hydrostatic test at 3110 psig pressure 5 (preoperational)

10. Hydrostatic test at 2485 psig pressure and 400°F temperature 5 (postoperational)

11. Steady state fluctuations the reactor coolant average temperature for purposes of design is assumed to increase and decrease a maximum of 6°F in one minute. The corresponding reactor coolant pressure variation is less than 100 psig. It is assumed that an infinite number of such fluctuations will occur.

Notes:

1. Estimated for equipment design purposes (40-yr life) and not intended to be an accurate representation of actual transients, or to reflect actual operating experience.

2. This transient includes pressurizing to 2235 psig.

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IP2 FSAR UPDATE TABLE 4.1-9 Reactor Coolant System - Design Code Requirements CODE APPLICABLE COMPONENT CODE EDITION ADDENDA Reactor vessel ASME III1 Class A 1965 Summer 1965 and Code Cases 1332, 1335, 1339, 1359 Control rod drive mechanism ASME III1 Class A 1965 Summer 1966 Steam generators Tube side ASME III1 Class A 1965 Summer 1966 Shell side4 ASME III1 Class C 1965 Summer 1966 Reactor coolant pump volute5 ASME III1 Class A 1965 Winter 19652 Pressurizer ASME III1 Class A 1965 Summer 1966 Pressurizer relief tank ASME III1 Class C 1964 Winter 1965 Pressurizer safety valves: Old Buy ASME III1 1971 Winter 1972 New Buy 1974 Summer 1975 Reactor coolant piping USAS B31.13 1955 System valves, fittings, piping USAS B31.13 1955 Notes:

1. ASME Boiler and Pressure Vessel Code, Section III, Nuclear Vessels.

2. Not stamped, but built in accordance with this edition and addenda.

3. USAS B31.1 Code for pressure piping.

4. The shell side of the generator conforms to requirements for Class A vessels and is so stamped as permitted under the rules of Section III.

5. The reactor pump, though not a coded vessel, was designed to Section III of the ASME Boiler and Pressure Vessel Code.

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IP2 FSAR UPDATE 4.2 SYSTEM DESIGN AND OPERATION 4.2.1 General Description The reactor coolant system consists of four similar heat transfer loops connected in parallel to the reactor vessel. Each loop contains a steam generator, a pump, loop piping, and instrumentation. The pressurizer surge line is connected to one of the loops. Auxiliary system piping connections into the reactor coolant piping are provided as necessary. A flow diagram of the system is shown in Plant Drawing 9321-2738 [Formerly UFSAR Figure 4.2-1], and a schematic flow diagram in Figure 4.2-2. The total design volume of the reactor coolant system, at rated operating conditions, is approximately 12,250-ft3. The nominal liquid volume of the reactor coolant system, at rated operating conditions and with 0% Steam Generator tube plugging, is 11,350 cubic feet. The containment boundary shown on the flow diagram indicates those major components, which are to be located inside the containment. The intersection of a process line with this boundary indicates a functional penetration. Reactor coolant system design data are listed in Tables 4.1-2 through 4.1-6. Pressure in the system is controlled by the pressurizer, where water and steam pressure is maintained through the use of electrical heaters and sprays. Steam can either be formed by the heaters, or condensed by a pressurizer spray to minimize pressure variations due to contraction and expansion of the coolant. Instrumentation used in the pressure control system is described in Chapter 7. Spring-loaded steam safety valves and power-operated relief valves are connected to the pressurizer and discharge to the pressurizer relief tank, where the discharged steam is condensed and cooled by mixing with water. 4.2.2 Components 4.2.2.1 Reactor Vessel The reactor vessel is cylindrical with a hemispherical bottom and a flanged and gasketed removable upper head. The vessel is designed in accordance with Section III (Nuclear Vessels) of ASME Boiler and Pressure Vessel Code. Figure 4.2-3 is a schematic of the reactor vessel. The materials of construction of the reactor vessel are given in Table 4.2-1. Coolant enters the reactor vessel through inlet nozzles in a plane just below the vessel flange and above the core. The coolant flows downward through the annular space between the vessel wall and the core barrel into a plenum at the bottom of the vessel where it reverses direction. Approximately 95-percent of the total coolant flow is effective for heat removal from the core. The remainder of the flow includes the flow through the rod cluster control guide thimbles, the leakage across the outlet nozzles, and the flow deflected into the head of the vessel for cooling the upper flange. All the coolant is united and mixed in the upper plenum and the mixed coolant stream then flows out of the vessel through exit nozzles located on the same plane as the inlet nozzles. A one-piece thermal shield, concentric with the reactor core, is located between the core barrel and the reactor vessel. It is attached to the core barrel. The shield, which is cooled by the coolant on its downward pass, protects the vessel by attenuating much of the gamma radiation Chapter 4, Page 19 of 85 Revision 22, 2010

IP2 FSAR UPDATE and some of the fast neutrons, which escape from the core. This shield minimizes thermal stresses in the vessel that result from heat generated by the absorption of gamma energy. This protection is further described in Section 3.2.3. Fifty-eight core instrumentation nozzles are located on the lower head. The reactor closure head and the reactor vessel flange are joined by fifty-four 7-in. diameter studs. Two metallic 0-rings seal the reactor vessel when the reactor closure head is bolted in place. A leakoff connection is provided between the two 0-rings to monitor leakage across the inner 0-ring. A leakoff connection is also provided beyond the outer 0-ring seal. The vessel is insulated with metallic reflective-type insulation supported from the nozzles. Insulation panels are provided for the reactor closure head, which are supported on the refueling seal ledge and vent shroud support rings. The reactor vessel internals are designed to direct the coolant flow, support the reactor core, and guide the control rods in the withdrawn position. The reactor vessel contains the core support assembly, upper plenum assembly, fuel assemblies, control cluster assemblies, surveillance specimens, and incore instrumentation. Surveillance specimens made from reactor vessel steel are located between the reactor vessel wall and the thermal shield. These specimens will be examined at selected intervals to evaluate reactor vessel material nil-ductility transition temperature changes as described in Section 4.5.2. The factor by which the maximum specimen exposure exceeds that at the vessel wall (at the location of maximum vessel wall exposure) has a maximum value of 3.5. Four of the eight irradiation specimens will lead the vessel wall maximum exposure by this factor. Ring forgings have been used for closure flanges; no other forgings have been used in the reactor vessel shell sections. The eight primary inlet and outlet nozzles have been provided with nozzle safe ends (forgings). These safe ends have been overlaid in the field with stainless steel weld metal. The Charpy V-notch and drop weight tests for the reactor vessel plates and forgings are discussed in Section 4.2.5. The reactor internals are described in detail in Section 3.2.3 and the general arrangement of the reactor vessel and internals is shown in Figure 3.2-47. Reactor vessel design data are listed in Table 4.1-2. 4.2.2.2 Pressurizer The general arrangement of the pressurizer is shown in Figure 4.2-4 and the design characteristics are listed in Table 4.1-3. The pressurizer maintains the required reactor coolant pressure during steady-state operation, limits the pressure changes caused by coolant thermal expansion and contraction during normal load transients, and prevents the pressure in the reactor coolant system from exceeding the design pressure. The pressurizer contains replaceable direct immersion heaters, multiple safety and relief valves, a spray nozzle and interconnecting piping, valves, and instrumentation. The electric heaters located in the lower section of the vessel regulate the reactor coolant system pressure by Chapter 4, Page 20 of 85 Revision 22, 2010

IP2 FSAR UPDATE keeping the water and steam in the pressurizer at saturation temperature. The heaters are capable of raising the temperature of the pressurizer and contents at approximately 55oF/hr during startup of the reactor. The pressurizer is designed to accommodate positive and negative surges caused by load transients. The surge line, which is attached to the bottom of the pressurizer, connects the pressurizer to a hot leg of a reactor coolant loop. During a positive surge caused by a decrease in plant load, the spray system, which is fed from the cold leg of a coolant loop, condenses steam in the vessel to prevent the pressurizer pressure from reaching the setpoint of the power operated relief valves. The spray valves on the pressurizer are power operated. In addition, the spray valves can be operated manually by a valve controller in the control room. A small continuous spray flow is provided to ensure that the pressurizer liquid is homogeneous with the coolant and to prevent excess cooling of the spray piping. During a negative pressure surge caused by an increase in plant load, flashing of water to steam and generation of steam by automatic actuation of the heaters keep the pressure above the minimum allowable limit. Heaters are also energized on high water level during positive surges to heat the subcooled surge water entering the pressurizer from the reactor coolant loop. The pressurizer is constructed of low alloy steel with internal surfaces clad with austenitic stainless steel. The heaters are sheathed in austenitic stainless steel. The pressurizer vessel surge nozzle is protected from thermal shock by a thermal sleeve. A thermal sleeve also protects the pressurizer spray nozzle connection. 4.2.2.3 Steam Generators Each loop contains a vertical shell and U-tube steam generator. A steam generator of this type is shown in Figure 4.2-5. Principal design parameters are listed in Table 4.1-4. The steam generators are designed and manufactured in accordance with Section III (Nuclear Vessels) of the ASME Boiler and Pressure Vessel Code. The secondary side of the steam generator shall not be pressurized above 200 psig if the temperature of the steam generator is below 70°F. Reactor coolant enters the inlet side of the channel head at the bottom of the steam generator through the inlet nozzle, flows through the U-tubes to an outlet channel, and leaves the generator through another bottom nozzle. The inlet and outlet channels are separated by a partition. Manways are provided to permit access to the U-tubes and the moisture-separating equipment. Feedwater to the steam generator enters just above the top of the U-tubes through a feedwater ring. The water flows downward through an annulus between the tube wrapper and the shell and then upward through the tube bundle where part of it is converted to steam. Certain plant operating conditions affecting the steam generator can result in the steam generator water level dropping below the feedwater sparging ring. As a result of these conditions and the waterhammer that can occur in the feedwater system, "J" tubes are installed on the feedwater sparging rings inside the steam generators. These "J" tubes preclude the rapid draining of the feedwater sparging rings and prevent steam from entering these rings even if they are uncovered. In the very remote event that the feedwater system would experience another large pressure wave, additional pipe restraints were installed in 1974 along the feedwater pipe. This modification is intended to preclude the rebound-type failure of the feedwater line at the containment penetration supports. All modifications that were made and the test program that Chapter 4, Page 21 of 85 Revision 22, 2010

IP2 FSAR UPDATE was performed as a result of these conditions were accomplished in accordance with the Quality Assurance Program for operating nuclear plants that was currently in effect. The steam-water mixture from the tube bundle passes through a steam swirl vane assembly, which imparts a centrifugal motion to the mixture and separates the water particles from the steam. The water spills over the edge of the swirl vane housing and combines with the feedwater for another pass through the tube bundle. The steam rises through additional separators, which further reduce the moisture content of the steam. A steam-generator blowdown system exists to perform several functions. Primarily it is used in maintaining the secondary side water chemistry of the steam generators within specifications. It also provides water samples from the secondary side of the steam generator as well as a means of draining the shell sides for inspection and/or maintenance. The steam generator is constructed primarily of low alloy steel. The heat transfer tubes are Inconel. The tubes undergo thermal treatment following tube-forming operations. The interior surfaces of the channel heads and nozzles are clad with austenitic stainless steel and the side of the tube sheet in contact with the reactor coolant is clad with Inconel. The tube-to-tube sheet joint is welded. The primary nozzles are provided with safe ends with weld metal overlay. Tubes are examined and defective tubes are repaired, plugged or sleeved as required by the technical specifications. The upper limit for tube plugging is 20-percent. Nozzle dam retention rings are permanently welded to the channel head cladding and provide a means of attachment for the temporary installation of nozzle dams during refueling and/or maintenance outages. The inspection ports and handholes are on the secondary side of the steam generators and any possible leakage (the only possible failure) would be inside containment. The possibility of secondary loss of water has been evaluated in Section 14.1.9, Loss of Normal Feedwater, and Section 14.2.5, Rupture of a Steam Pipe. These sections show that the types of failure possible due to secondary leakage or loss of water have already been analyzed. 4.2.2.4 Reactor Coolant Pumps Each reactor coolant loop contains a vertical, single-stage centrifugal pump that employs a controlled leakage seal assembly. A view of a controlled leakage pump is shown in Figure 4.2-6 and the principal design parameters for the pumps are listed in Table 4.1-5. The reactor coolant pump estimated performance and net positive suction head characteristics are shown in Figure 4.2-7. The performance characteristic is common to all of the higher specific speed centrifugal pumps and the "knee" at about 45-percent design flow introduces no operational restrictions, because the pumps operate at full speed. During normal operation, the reactor coolant pumps are supplied from the unit auxiliary bus and are therefore tied to the turbine-generator frequency (speed). On occurrence of unit turbine trip, the pump electrical buses are transferred from the unit auxiliary transformer to the station auxiliary transformer with an intentional delay of 30 sec. Further details are given in Section 14.1.8.3. On most electrical events, which cause the turbine to be tripped, the reactor coolant pump buses are transferred to offsite power and the unit is tripped simultaneously and the pumps will Chapter 4, Page 22 of 85 Revision 22, 2010

IP2 FSAR UPDATE therefore not exceed their normal running speed. If for some unlikely reason the only plant trip is a turbine over-speed trip an over-frequency trip relay circuit is provided that will trip the turbine-generator. This trip circuit first locks out the 6.9 kV dead bus transfer at 62.2 +/-0.1 Hz (1866 +/- 3 rpm) and then trips the main generator at 62.5 +/-0.1 Hz (1875 +/- 3 rpm). Termination of power to the in-house 6.9 kV buses 1-4 limits the reactor coolant pumps overspeed to maintain the RCS flow condition below the design limit per section 4.2.2.5.4. Reactor coolant is pumped by the impeller attached to the bottom of the rotor shaft. The coolant is drawn up through the impeller, discharged through passages in the diffuser, and exits through a discharge nozzle in the side of the casing. The motor-impeller can be removed from the casing of the piping. All parts of the pumps in contact with the reactor coolant are austenitic stainless steel or equivalent corrosion resistant materials. The pump employs a controlled leakage seal assembly to restrict leakage along the pump shaft, a second seal that directs the controlled leakage out of the pump, and a third seal that minimizes the leakage of water and vapor from the pump into the containment atmosphere. A portion of the high-pressure water flow from the charging pumps is injected into the reactor coolant pump between the impeller and the controlled leakage seal. Part of the flow enters the reactor coolant system through a labyrinth seal in the lower pump shaft to serve as a buffer to keep reactor coolant from entering the upper portion of the pump. The remainder of the injection water flows along the drive shaft, through the controlled leakage seal, and finally out of the pump. A very small amount that leaks through the second seal is also collected and removed from the pump. Component cooling water is supplied to the motor bearing oil coolers of the reactor coolant pumps. The component cooling water system also provides cooling flow to the thermal barrier heat exchanger of the reactor coolant pumps to minimize heat transfer from the high-temperature primary coolant to the seal area environment, to cool primary system water that could leak through the thermal barrier labyrinth seals, and to provide adequate seal cooling in the event that seal injection flow was lost. In the event of loss of offsite power, the reactor coolant pump motor is deenergized and both cooling water supplies (seal injection and component cooling flow) are terminated; however, the plant diesel generators are immediately started and the component cooling water pumps are automatically loaded (in sequence) onto the emergency buses and started (no operator action required). Once the automatic loading of the emergency buses has been completed, the operator has the option of manually loading a charging pump onto one of the diesels and reestablishing normal seal injection flow. The squirrel cage induction motor driving the pump is air cooled and has oil lubricated thrust and radial bearings. A water lubricated bearing provides radial support for the pump shaft. An extensive test program was conducted for several years to develop the controlled leakage shaft seal for pressurized-water reactor applications. Long-term tests were conducted on less than full-scale prototype seals as well as on full-size seals. Operating experience with other large size, controlled leakage shaft seal pumps has also been available. The reactor coolant pump motor bearings are of conventional design, the radial bearings are the segmented pad type, and the thrust bearings are tilting pad Kingsbury bearings. All are oil lubricated - the lower radial bearing and thrust bearings are submerged in oil, and the upper radial bearing is oil fed from an impeller integral with the thrust runner. Both high and low oil Chapter 4, Page 23 of 85 Revision 22, 2010

IP2 FSAR UPDATE levels would signal an alarm in the control room to alert the operator to possible pending bearing problems. Each motor bearing contains embedded temperature detectors, and so initiation of failure, separate from or in combination with a loss of oil, would be indicated and alarmed in the control room as a high bearing temperature. This would require a reactor trip followed by pump shutdown. Even if these indications were ignored, and the bearing proceeded to failure, the low melting point Babbitt metal on the pad surfaces would ensure that no sudden seizure of the bearing would occur. In this event the motor would continue to drive, as it has sufficient reserve capacity to operate even under such conditions. However, it would demand excessive currents and at some stage would be shut down because of high current demand. It may be hypothesized that the pump impeller might severely rub on a stationary member and then seize. Analysis has shown that under such conditions, assuming instantaneous seizure of the impeller, the pump shaft would fail in torsion just below the coupling to the motor. This would constitute a loss of coolant flow in the one loop, the effect of which is analyzed in Section 14.1.6. Following the seizure, the motor would continue to run without any overspeed and the flywheel would maintain its integrity, as it would still be supported on a shaft with two bearings. There are no other credible sources of shaft seizure other than impeller rubs. Any seizure of the pump bearing would be precluded by shearing of the graphitar in the bearing. Any seizure in the seals would result in a shearing of the anti-rotation pin in the seal ring. The motor has adequate power to continue pump operation even after the above occurrences. Indications of pump malfunction in these conditions would be initially by high temperature signals from the bearing water temperature detector, and excessive No. 1 seal leakoff indications, respectively. Following these signals, pump vibration levels would be checked. These would show excessive levels, indicating some mechanical trouble. Again, the pump would be shut down for investigation. The design specifications for the reactor coolant pumps include as a design condition that the pumps are designed to withstand seismic load equivalent to 0.28 g in the vertical direction and 0.40 g in the horizontal direction and the seismic loads shall be considered acting simultaneously. Besides examining the externally produced loads from the nozzles and support lugs, an analysis was made of the effect of gyroscopic reaction on the flywheel and bearings and in the shaft due to rotational movements of the pump about a horizontal axis during the maximum seismic disturbance. The pump would continue to run unaffected by such conditions. In no case does any bearing stress in the pump or motor exceed or even approach a value, which the bearing could not carry. The design requirements of the bearings are primarily aimed at ensuring a long life with negligible wear so as to give accurate alignment and smooth operation over long periods of time. To this end, the surface bearing stresses are held at a very low value, and even under the most severe seismic transients or other accidents, do not begin to approach loads, which cannot be adequately carried for short periods of time. Because there are no established criteria for short-time stress-related failures in such bearings, it is not possible to make a meaningful quantification of such parameters as margins-to-failure, safety factors, etc. A qualitative analysis of the bearing design, embodying such considerations, gives assurance of the adequacy of the bearing to operate without failure. Chapter 4, Page 24 of 85 Revision 22, 2010

IP2 FSAR UPDATE As is generally the case with machines of this size, the shaft dimensions are predicated on avoidance of shaft critical speed conditions, rather than actual levels of stress. There are many machines as large as, and larger than these, that are designed to run at speeds in excess of first shaft critical. However, it is considered desirable in a superior product to operate below first critical speed; the reactor coolant pumps are designed in accordance with this philosophy. This results in a shaft design, which even under the severest postulated transient, gives very low values of actual stress. While it would be possible to present quantitative data of imposed operational stress relative to maximum tolerable levels, if the mode of postulated failure were clearly defined, such figures would have little significance in a meaningful assessment of the adequacy of the shaft to maintain its integrity under operational transients. However, a qualitative assessment of such factors gives assurance of the conservative stress levels experienced during these transients. So in each of these cases, where it is the functional requirements of the component that control its dimensions, it can be seen that if these are met, the stress-related failure cases are more than adequately satisfied. It is thus considered to be out of the bounds of reasonable credibility that any bearing or shaft failure could occur that would endanger the integrity of the pump. 4.2.2.5 Reactor Coolant Pump Flywheel Integrity The reactor coolant pump flywheels were fabricated from two rolled, vacuum-degassed, ASTM A-533 Grade B Class 1 steel plates. The plates are bolted together with bolts aligned perpendicular to the plane of the plates. Thus the bolts carry no stress during operation. The flywheel blanks were flame-cut from the plates, with allowance for exclusion of flame-affected metal. They were then machined to the specified dimensions and the bolt holes were drilled. Two plates were then bolted together, the finished flywheel attached to the motor shaft, and the whole unit balanced to yield vibration levels at operating speed less than 0.001-in. double amplitude. The reactor coolant pump flywheel is shown in Figure 4.2-8. A nil-ductility transition temperature less than +10oF was specified. A minimum of three Charpy tests, parallel and normal to the rolling direction, were made from each plate to determine that each blank satisfied design requirements. The finished flywheels were subjected to 100-percent volumetric ultrasonic inspection. The finished machined bores were also subjected to magnetic particle or liquid penetrant examination. These design-fabrication techniques yield flywheels with primary stress at operating speed (shown in Figure 4.2-9) less than 50-percent of the minimum specified material yield strength at room temperature (100 to 150oF). A fracture mechanics evaluation was made on the reactor coolant pump flywheel. This evaluation considered the following assumptions: Chapter 4, Page 25 of 85 Revision 22, 2010

IP2 FSAR UPDATE

1. Maximum tangential stress at an assumed overspeed of 125-percent compared with a maximum expected overspeed of 109-percent.

2. A through crack through the thickness of the flywheel at the bore.

3. 400 cycles of startup operation in 40 years.

Using critical stress intensity factors and crack growth data attained on flywheel material, the critical crack size for failure was greater than 17-in. radially and the crack growth data was 0.030-in. to 0.060-in. per 1000 cycles. Periodic nondestructive examinations of the reactor coolant pump flywheels are included in the inservice inspection and testing program discussed in Chapter 1.

4. Normal operating speed of the flywheel (pump).

The primary coolant pumps run at 1189 rpm and may operate briefly at overspeeds up to 109-percent (1295 rpm) during loss of offsite load. For conservatism, however, 125-percent of operating speed was selected as the design speed for the primary coolant pumps. For the overspeed condition, which would not persist for more than 30 sec, pump operating temperature would remain at about the design value. However, the limiting condition for the RCS system components is the effect the excess RCS flow has on the reactor internals. This limit is a statistically derived RCS flow value of 115.8% at a confidence level of 95%.

5. Bursting speed of the flywheel.

Bursting speed of the flywheels was calculated on the basis of Robinson's results (Reference 1) to be 3900 rpm, more than three times the operating speed. This is confirmed using Griffith-Irwin theory as detailed in Reference 2. 4.2.2.6 Pressurizer Relief Tank Principal design parameters of the pressurizer relief tank are given in Table 4.1-3. The tank is shown on Figure 4.2-10. Steam and water discharge from the power relief and safety valves pass to the pressurizer relief tank, which is partially filled with water. The tank normally contains water in a predominantly nitrogen atmosphere. Steam is discharged under the water level to condense and cool by mixing with the water. The tank is equipped with a spray and a drain to the waste disposal system, which are operated to cool the tank following a discharge. The tank size is based on the requirement to condense and cool a discharge equivalent to 110-percent of full power pressurizer steam volume. The tank is protected against a discharge exceeding the design value by two rupture disks that discharge into the reactor containment. The rupture disks on the relief tank have a combined relief capacity equal to the combined capacity of the pressurizer safety valves. The tank design pressure (and the rupture disk's bursting pressure) is twice the calculated pressure resulting from the maximum safety valve discharge described above. This margin is to prevent Chapter 4, Page 26 of 85 Revision 22, 2010

IP2 FSAR UPDATE deformation of the disk. The tank and rupture disk holder are also designed for full vacuum to prevent tank collapse if the tank contents cool without nitrogen being added. The discharge piping from the safety and relief valves to the relief tank is sufficiently large to prevent backpressure at the safety valves from exceeding 20-percent of the setpoint pressure at full flow. The pressurizer relief tank, by means of its connection to the waste disposal system, provides a means for removing any noncondensable gasses from the reactor coolant system, which might collect in the pressurizer vessel. The tank is constructed of carbon steel with a corrosion-resistant coating on the internal surface. 4.2.2.7 Piping A schematic of the reactor coolant piping is shown on Figure 4.2-2. The general arrangement of the loop piping is shown on Plant Drawings 9321-2502, 9321-2506, 9321-2508 [Formerly UFSAR Figures 5.1-3, 5.1-5, and 5.1-7]. Piping design data are presented in Table 4.1-6. The austenitic stainless steel reactor coolant piping and fittings that make up the loops are 29-in. ID in the hot legs, 27.5-in. ID in the cold legs, and 31-in. ID between the steam generator outlet and reactor coolant pump suction. The pressurizer relief line, which connects the outlets of the pressurizer safety and relief valves to the inlet nozzle flange on the pressurizer relief tank, is constructed of carbon steel. Smaller piping, including the pressurizer surge and spray lines, drains, and connections to other systems are austenitic stainless steel. All piping connections are welded except for flanged connections at the pressurizer relief tank and at the safety valves, and the vacuum fill connection closure. In response to NRC Bulletin 88-11, thermal stratification effects on the pressurizer surge line have been evaluated for the design life of the plant18. The stress and fatigue analyses results are within the ASME Code allowables for the surge line. Thermal sleeves are installed at the following locations where high thermal stresses could otherwise develop due to rapid changes in fluid temperature during normal operational transients:

1. Return lines from the residual heat removal loop (safety injection lines) See Section 4.1.7.

2. Both ends of the pressurizer surge line.

3. Pressurizer spray line connection to the pressurizer.

4. Charging lines and auxiliary charging line connections.

4.2.2.8 Valves All valve surfaces in contact with reactor coolant are austenitic stainless steel or equivalent corrosion-resistant materials. Connections to stainless steel piping are welded. Valves that Chapter 4, Page 27 of 85 Revision 22, 2010

IP2 FSAR UPDATE perform a modulating function are equipped with either two sets of packing and an intermediate leakoff connection or have been designed with live-loaded packing which will either control of mitigate the potential for valve stem leakage due to modulating service. 4.2.2.9 Component Supports The support structures for the reactor coolant components are described in Appendix 4B and Chapter 5. 4.2.3 Pressure-Relieving Devices The reactor coolant system is protected against overpressure by control and protective circuits such as the high-pressure trip and by code relief valves connected to the top head of the pressurizer. The relief valves discharge into the pressurizer relief tank, which condenses and collects the valve effluent. The schematic arrangement of the relief devices is shown in Figure 4.2-1; the valve design parameters are given in Table 4.1-3. Valve sizes are determined as indicated in Section 4.3.4. Power-operated relief valves and code safety valves are provided to protect against pressure that is beyond the pressure limiting capacity of the pressurizer spray. Acoustic sensors installed on the code safety valve discharge lines provide indication in the control room of the "flow" or "no flow" condition of the safety valves. Direct valve position indication is also provided for the power-operated relief valves. The pressurizer relief tank is protected against a steam discharge exceeding the design pressure value by two rupture discs, which discharge into the reactor containment. The rupture disc relief conditions are given in Table 4.1-3. 4.2.4 Protection Against Proliferation Of Dynamic Effects Engineered safety features and associated systems are protected from loss of function due to dynamic effects and missiles, which might result from a loss-of coolant accident. Protection is provided by missile shielding and/or segregation of redundant components. This is discussed in detail in Chapter 6. The reactor coolant system is surrounded by a concrete shield wall. This wall provides shielding to permit access into the containment annular region during full-power operation for inspection and maintenance of miscellaneous equipment. This shielding wall also provides missile protection for the containment liner plate. The concrete deck over the reactor coolant system also provides for shielding and missile damage protection. Lateral bracing is provided near the steam-generator upper tube sheet elevation to resist lateral loads, including those resulting from seismic forces and pipe rupture forces. Additional bracing is provided at a lower elevation to resist pipe rupture loads. Missile protection afforded by the arrangement of the reactor coolant system is illustrated in the containment structure drawings, which are given in Chapter 5. [Historical Information Only] This paragraph is retained for historical purposes only. The integrity of the reactor coolant system as may be affected by asymmetric loss-of-coolant accident loads due to postulated pipe breaks in the primary loop coolant piping was considered Chapter 4, Page 28 of 85 Revision 22, 2010

IP2 FSAR UPDATE in WCAP-9117 (Reference 3) and in a subsequent submittal to the NRC on June 15, 1978 (Reference 4). The combination of LOCA and safe shutdown earthquake loads applied to the results of WCAP-9117 was described in the September 3, 1980 follow-up submittal (Reference 5), and is applicable to both Indian Point Units 2 and 3, based on the similarity between the two units. The safe shutdown earthquake results for Unit 3 would be similar to those for Unit 2, and the total strains also apply to Unit 2. The NRC Safety Evaluation Report (Reference 6) concluded that the assessment of asymmetric loss-of-coolant accident and safe shutdown earthquake loads for the Indian Point Unit 2 was acceptable. This conclusion was based upon the installation of pipe motion limiters in the primary shield wall and was contingent upon the verification of shield plug assumptions, which included the determination of the effects of the plugs as missiles, and the determination that structural components do not inhibit plug displacement. This verification was provided in a subsequent submittal to the NRC on June 10, 1986 (Reference 22). In 1989, the NRC approved changes to the design bases with respect to dynamic affects of postulated primary loop pipe ruptures, as discussed in Section 4.1.2.4. 4.2.5 Materials Of Construction Each of the materials used in the reactor coolant system is selected for the expected environment and service conditions. The major component materials are listed in Table 4.2-1. All reactor coolant system materials that are exposed to the coolant are corrosion resistant. They consist of stainless steels and Inconel, and they are chosen for specific purposes at various locations within the system for their superior compatibility with the reactor coolant. Reactor coolant chemistry is further discussed in Section 4.2.8. It is characteristic of stress corrosion that combinations of alloy and environment, which result in cracking are usually quite specific. Environments that have been shown to cause stress-corrosion cracking of stainless steels are free alkalinity in the presence of a concentrating mechanism, and the presence of chlorides and free oxygen. With regard to the former, experience has shown that deposition of chemicals on the surface of tubes can occur in a steam blanketed area within a steam generator. In the presence of this environment, stress-corrosion cracking can occur in stainless steels having the nominal residual stresses resulting from normal manufacturing procedures. However, the steam generator contains Inconel tubes. Testing to investigate the susceptibility of heat exchanger construction materials to stress corrosion in caustic and chloride aqueous solutions has indicated that Inconel alloy has excellent resistance to general and pitting-type corrosion. Considerable experience with Inconel in steam generator and heat exchanger applications has been accumulated in the industry. Since 1962, widespread adoption of Inconel for steam generator tubes in nuclear stations is evident: as for example, Connecticut-Yankee; San Onofre; PM-1, Sundance; PM-3A, McMurdo Sound; CVTR; NPD, and Hanford N-Reactor. Materials with lead traces in the overall composition were present in the secondary side of the referenced plants. The use of lead in the materials of the secondary side of the Indian Point plant has been minimized to the practical limit of that occurring as trace elements in metallurgical alloys and as such is insignificant. All external insulation of reactor coolant system components is compatible with the component materials. The cylindrical shell exterior and closure flanges to the reactor vessel are insulated with metallic reflective insulation. The closure head is also insulated with metallic reflective Chapter 4, Page 29 of 85 Revision 22, 2010

IP2 FSAR UPDATE insulation. All other external corrosion-resistant surfaces in the reactor coolant system are insulated with low or halide-free insulating material as required. The reactor vessel was fabricated by Combustion Engineering, Inc. A sketch of the reactor vessel showing all materials in the beltline region is shown in Figure 4.2-11. Information on each of the welds and plates in the beltline region is shown in Tables 4.2-2 through 4.2-5, and Tables 4.2-5 through 4.2-8, respectively. Information relative to weld and plate material included in the material surveillance program is shown in Tables 4.2-2 and 4.2-6 through 4.2-8. Details concerning the reactor vessel radiation surveillance program are provided in WCAP-7323 (Reference 7) and in the Technical Specifications. The reactor vessel plate or forging material opposite the core is purchased to a specified Charpy V-notch test result of 30-ft-lb or greater at a corresponding nil-ductility transition temperature (NDTT) of 40oF or less, and the material is tested to verify conformity to specified requirements and to determine the actual NDTT value (see Table 4.2-9). In addition, this plate is 100-percent volumetrically inspected by ultrasonic test using both longitudinal and shear wave methods. The remaining material in the reactor vessel and other reactor coolant system components meets the appropriate design code requirements and specific component function. The reactor vessel material is heat-treated specifically to obtain good notch-ductility, which ensures a low NDTT, and thereby gives assurance that the finished vessel can be initially hydrostatically tested and operated near room temperature within the restrictions of NDTT + 60°F. The stress limits established for the reactor vessel are dependent upon the temperature at which the stresses are applied. As a result of fast neutron irradiation in the region of the core, the material properties will change, including an increase in the NDTT. An initial maximum value of NDTT of 40°F has been established during fabrication in this region. The techniques used to measure and predict the integrated fast neutron (E > 1 MeV) fluxes at the sample locations are described in Appendix 4A. The calculation method used to obtain the maximum neutron (E > 1 MeV) exposure of the reactor vessel is identical to that described for the irradiation samples. Since the neutron spectra at the samples and vessel inner surface are identical, the measured transition shift for a sample can be applied with confidence to the adjacent section of reactor vessel for some later stage in plant life. The maximum exposure of the vessel will be obtained from the measured sample exposure by appropriate application of the calculated azimuthal neutron flux variation. The evaluation of the second surveillance capsule is discussed in detail in Reference 8. The analysis for the third surveillance capsule, removed during the 1984 refueling outage, is documented in Reference 9. The analysis of the fourth surveillance Capsule V, removed during the 1987-88 refueling outage (end of Cycle 8), is documented in Reference 11. The Capsule V received a fast neutron (E > 1 MeV) fluence of 5.3 x 1018 n/cm2 in 8.6EFPY at the end of Cycle 8. See Reference 11. The maximum integrated fast neutron (E > 1 MeV) exposure of the vessel for 32 EFPYs is calculated to be 1.39 x 1019 n/cm2 based on the measurements from the fourth surveillance Capsule V. Fast neutron fluences corresponding to 32 EFPYs at various reactor vessel thicknesses are given in Table 4.2-10. Chapter 4, Page 30 of 85 Revision 22, 2010

IP2 FSAR UPDATE The calculated neutron exposure exceeds the value of 0.85 x 1019 n/cm2 (E > 1MeV) reported in the First Supplement to the Preliminary Facility Description Safety Analysis Report. The reasons for the increase are:

1. Anticipated increase in reactor power from 2758 MWt to 3071.4 MWt in Cycle 10 and then to 3114.4 MWt in Cycle 16 and subsequently to 3216 MWt in Cycle 17.

2. Revision of analysis methodology including upgrading of neutron cross sections and codes.

3. Core design considerations involving changes in loading patterns.

The above projected exposure to reactor vessel from Capsule V measurements is based upon using the standard loading pattern (only fresh fuel assemblies at core periphery) in Cycles 1 thru 5 and the low leakage loading pattern (mixture of fresh and spent fuel assemblies) or L3P starting from Cycle 6. Furthermore, it assumes Indian Point Unit 2 operation at stretch power operation at 3071.4 MWt core power level starting from Cycle 10, then at the Appendix K uprated power level of 3114.4 MWt starting in Cycle 16 and subsequently at 3216 MWt starting in Cycle 17. The maximum reference temperature, RTNDT for the Indian Point Unit 2 vessel core beltline materials at the 1/4 thickness and the 3/4 thickness after 32 effective full power years of operation are projected to be 240°F and 194°F, respectively, based on calculations performed per Regulatory Guide 1.99, Revision 2, using data obtained from evaluation of Surveillance Capsule V. (Ref.11). This data provides the basis for subsequent calculation of Adjusted Reference Temperature values for determination of allowable pressure/temperature limits for operation to 25 EFPY, as described in Reference 19. To evaluate the NDTT shift of welds, heat affected zones, and base material for the vessel, test coupons of these material types have been included in the reactor vessel surveillance program described in Section 4.5.2. The methods used to measure the initial NDTT of the reactor vessel baseplate material are given in Appendix 4A. The reference nil ductility transition temperatures for pressurized thermal shock evaluation (RTPTS) have been estimated11,12,13 in accordance with 10 CFR 50.61(b)(2). The values at 15 EFPY and also at the end of the license term are well below the screening criteria of 270°F (for plates and axial weld materials) and 300°F (for circumferential weld materials), based on a low-leakage core design. The NRC has accepted this analysis.14 Additional information in response to Generic Letter 92-01, Revision 1, is given in reference 21. With regard to electroslag welding of Class I components, the Indian Point Unit 2 90-degree elbows were electroslag welded. The following efforts were performed for quality assurance of these components.

1. The electroslag welding procedure employing one-wire technique was qualified in accordance with the requirements of ASME Boiler and Pressure Vessel Code, Section IX and Code Case 1355 plus supplementary evaluations as requested by Westinghouse. The following test specimens were removed from a 5-in. thick weldment and successfully tested:

a. Six transverse tensile bars - as welded.

b. Six transverse tensile bars - 2050oF, H2O quench.

Chapter 4, Page 31 of 85 Revision 22, 2010

IP2 FSAR UPDATE

c. Six transverse tensile bars - 2050oF, H2O quench + 750oF stress relief heat treatment.

d. Six transverse tensile bars - 2050oF, H2O quench, tested at 650oF.

e. Twelve guided side bend test bars.

2. The casting segments were surface conditioned for 100-percent radiographic and penetrant inspections. The acceptance standards were ASTM E-186 severity level 2 (except no category D or E defectiveness was permitted) and ASME Section III, Paragraph N-627, respectively.

3. The edges of the electroslag weld preparations were machined. These surfaces were penetrant inspected prior to welding. The acceptance standards were ASME Section III, Paragraph N-627.

4. The completed electroslag weld surfaces were ground flush with the casting surface. Then the electroslag weld and adjacent base material were 100-percent radiographed in accordance with ASME Code Case 1355. Also, the electroslag weld surfaces and adjacent base material were penetrant inspected in accordance with ASME Boiler and Pressure Vessel Code Section III, Paragraph N-627.

5. Weld metal and base metal chemical and physical analyses were determined and certified.

6. Heat treatment furnace charts were recorded and certified.

Two of the Indian Point Unit 2 reactor coolant pump casings were electroslag welded. The following efforts were performed for quality assurance of these two components.

1. The electroslag welding procedure employing two- and three-wire technique was qualified in accordance with the requirements of the ASME Boiler and Pressure Vessel Code, Section IX and Code Case 1355 plus supplementary evaluations as requested by Westinghouse. The following test specimens were removed from an 8-in.-thick and from a 12-in.-thick weldment and successfully tested for both the two-wire and the three-wire techniques, respectfully.

a. Two-wire electroslag process in.-thick weldment.

(1) 6 transverse tensile bars - 750oF postweld stress relief. (2) 12 guided side bend test bars.

b. Three-wire electroslag process in.-thick weldment.

(1) 6 transverse tensile bars - 750oF postweld stress relief. (2) 17 guided side bend test bars. (3) 21 Charpy V-notch specimens. (4) Full-section macroexamination of weld and heat affected zone. Chapter 4, Page 32 of 85 Revision 22, 2010

IP2 FSAR UPDATE (5) Numerous microscopic examinations of specimens removed from the weld and heat affected zone regions. (6) Hardness survey across weld and heat affected zone.

c. A separate weld test was made using the two-wire electroslag technique to evaluate the effects of a stop and restart of welding by this process. This evaluation was performed to establish proper procedures and techniques as such an occurrence was anticipated during production applications due to equipment malfunction, power outages, etc. The following test specimens were removed from an 8-in.-thick weldment in the stop-restart-repaired region and successfully tested.

(1) 2 transverse tensile bars - as welded. (2) 4 guided side bend test bars. (3) Full section macroexamination of weld and heat affected zone.

d. All of the weld test blocks in (a), (b), and (c) above were radiographed using a 24-MeV Betatron. The radiographic quality level obtained was between 0.05 to 1-percent. There were no discontinuities evident in any of the electroslag welds.

(1) The casting segments were surface conditioned for 100-percent radiographic and penetrant inspections. The radiographic acceptance standards were ASTM E-186 severity level 2 (except no category D or E defectiveness was permitted) for section thickness up to 4.5-in. and ASTM E-280 severity level 2 for section thicknesses greater than 4.5-in. The penetrant acceptance standards were ASME Boiler and Pressure Vessel Code, Section III, Paragraph N-627. (2) The edges of the electroslag weld preparations were machined. These surfaces were penetrant inspected prior to welding. The acceptance standards were ASME Boiler and Pressure Vessel Code, Section III, Paragraph N-627. (3) The completed electroslag weld surfaces were ground flush with the casting surface. Then the electroslag weld and adjacent base material were 100-percent radiographed in accordance with ASME Code Case 1355. Also, the electroslag weld surfaces and adjacent base material were penetrant inspected in accordance with ASME Boiler and Pressure Vessel Code, Section III, Paragraph N-627. (4) Weld metal and base metal chemical and physical analyses were determined and certified. (5) Heat treatment furnace charts were recorded and certified. Chapter 4, Page 33 of 85 Revision 22, 2010

IP2 FSAR UPDATE The two remaining Indian Point Unit 2 reactor coolant pump casings were submerged arc welded. Quality Assurance procedures and Quality Assurance inspections equivalent to the above were also exercised on these casings. 4.2.6 Maximum Heating And Cooling Rates The reactor system operating cycles used for design purposes are given in Table 4.1-8 and described in Section 4.1.5. The reactor coolant system heatup, cooldown, and leak test limitations curves are included in the Technical Specifications. Starting with a minimum water level, sufficient electrical heaters are installed in the pressurizer to permit a heatup rate of 55°F/hr. This rate takes into account the small continuous spray flow provided to maintain the pressurizer liquid homogeneous with the coolant. The fastest cooldown rates, which result from the hypothetical case of a break of a main steam line are discussed in Section 14.2.5. 4.2.7 Leakage The existence of leakage from the reactor coolant system to the containment regardless of the source of leakage is detected by one or more of the following conditions:

1. Radiation sensitive instruments provide the capability for detection of leakage from the reactor coolant system. The containment air particulate monitors are quite sensitive to low leak rates. The containment radiogas monitors are less sensitive but are used in addition to the air particulate monitor.

2. A third mechanism used in leak detection is the humidity detectors. These provide a means of measuring overall leakage from all water and steam systems within the containment, which can affect containment humidity. The humidity monitoring method is considered supplemental to the radiation monitoring methods.

3. A leakage detection system collects and measures moisture condensed from the containment atmosphere by cooling coils of the main air recirculation units. The condenser moisture includes, of course, any leaks from the cooling coils themselves. This system provides a dependable and accurate means of measuring the total leakage from these sources. Condensate flows of approximately 1.0 gpm to 15 gpm per detector can be measured by this system.

Condensate flows can be determined using weir calibration curves in conjunction with the weir water head displayed by the weir water meter, or by direct reading of the weir integrated condensate flow on the weir meter.

4. An increase in the amount of coolant makeup water, which is required to maintain normal level in the pressurizer or an increase in containment sump level provide additional means of detecting leakage.

The Technical Specifications provide the requirements and bases for leakage detection. In considering potential leakage from the reactor coolant system containing primary coolant at high pressure, four categories are described and evaluated in Section 6.7.1. These include leakage paths to the reactor coolant drain tank, leakage paths to the pressurizer relief tank, leakage paths to the containment environment, and leakage paths to the interconnecting systems. Chapter 4, Page 34 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.2.7.1 Maximum Leak Rates The maximum leak rate from an unidentified source that will be permitted during normal operation is specified in the Technical Specifications. Leakage from the reactor coolant system is collected in the containment or by the other closed systems. These closed systems are: the steam and feedwater system, the waste disposal system, and the component cooling system. Assuming the existence of the maximum allowable activity in the reactor coolant, the rate of unidentified leakage is a conservative limit on what is allowable before the guidelines of 10 CFR 20 would be exceeded. With the limiting reactor coolant activity and assuming initiation of a leak from the reactor coolant system to the component cooling system, the radiation monitor that samples the component cooling pump discharge downstream of the component cooling heat exchangers would annunciate in the control room and initiate closure of the surge tank vent line in the component cooling system. In the case of failure of the closure of the vent line and resulting continuous discharge in the atmosphere via the component cooling surge tank vent, the resultant dose at the site boundary would be within the limit allowed by 10 CFR 20. Leakage directly into the containment indicates the possibility of a breach in the coolant envelope. The limitation specified by the Technical Specifications for a source of leakage not identified is sufficiently above the minimum detectable leakage rate to provide a reliable indication of leakage. The leakage limit is well within the capacity of one coolant charging pump. The conservative approach that is used in the design and fabrication of the components that constitute the primary system pressure boundary together with the operating restrictions, which are imposed for system heatup and cooldown give adequate assurance that the integrity of the primary system pressure boundary is maintained throughout plant life. The periodic examination of the primary pressure boundary via the inservice inspection program (specified in the Technical Specifications) will physically demonstrate that the operating environment will have no deleterious effect on the primary pressure boundary integrity. The maximum unidentified leak rate that is permitted during normal operation is well within the sensitivity of the leak detection systems incorporated within the containment, and it reflects good operating practice based on operating experience gained at other PWR plants. Detection of leakage from the primary system directs the operator's attention to potential sources of leakage, such as valves, and permits timely evaluation to ensure that any associated activity release does not constitute a public hazard, that the reactor coolant inventory is not significantly affected, and that the leakage is well within the capability of the containment drainage system. See also Section 6.7 for a further discussion of leakage detection. 4.2.7.2 Leakage Prevention Reactor coolant system components are manufactured to exacting specifications, which exceed normal code requirements (as outlined in Section 4.1.7). In addition, because of the welded construction of the reactor coolant system and the extensive nondestructive testing to which it is subjected (as outlined in Section 4.5), it is considered that leakage through metal surfaces or welded joints is very unlikely. However, some leakage from the reactor coolant system is permitted by design from the reactor coolant pump seals. Also, all sealed joints are potential sources of leakage even though the Chapter 4, Page 35 of 85 Revision 22, 2010

IP2 FSAR UPDATE most appropriate sealing device is selected in each case. Thus, because of the large number of joints and the difficulty of ensuring complete freedom from leakage in each case, a small integrated leakage is considered acceptable. Leakage from the reactor through its head flange will leak-off between the double O-ring seal and actuate an alarm in the control room. 4.2.7.3 Locating Leaks Experience has shown that hydrostatic testing is successful in locating leaks in a pressure containing system. The Reactor Coolant System shall be tested for leakage at normal operating pressure prior to plant startup following each refueling outage, in accordance with the requirements of the applicable edition and addenda of the ASME Section XI Code. Leak test of the Reactor Coolant System is required by the ASME Boiler and Pressure Vessel Code, Section XI, to ensure leak tightness of the system during operation. The test frequency and conditions are specified in the Code. Testing of repairs, replacements or modifications for the Reactor Coolant System shall meet the requirements of the applicable edition and addenda of the ASME Section XI Code. For repairs on components, the thorough non-destructive testing gives a very high degree of confidence in the integrity of the system, and will detect any significant defects in and near the new welds. In all cases, the leak test will assure leak-tightness during normal operation. Methods of leak location, which can be used during plant shutdown include visual observation for escaping steam or water, or for the presence of boric acid crystals near the leak. The boric acid crystals are transported outside the reactor coolant system in the leaking fluid and deposited by the evaporation process. 4.2.8 Water Chemistry The water chemistry is selected to provide the necessary boron content for reactivity control and to minimize corrosion of reactor coolant system surfaces. All materials exposed to reactor coolant are corrosion resistant. Periodic analyses of the coolant chemical composition are performed to monitor the adherence of the system to the required reactor coolant water quality. Maintenance of the water quality to minimize corrosion is accomplished using the chemical and volume system and sampling system that is described in Chapter 9. 4.2.9 Reactor Coolant Flow Measurement Elbow taps are used in the primary coolant system as an instrument device that indicates the status of the reactor coolant flow. The basic function of this device is to provide information as to whether or not a reduction in flow rate has occurred. The correlation between flow reduction and elbow tap read out has been well established by the following equation: P 2

             =( )

Po o where Po is the reference pressure differential with the corresponding referenced flow rate o, and P is the pressure differential with the corresponding referenced flow rate. The full flow reference point was established during initial plant startup. The low flow trip point was then Chapter 4, Page 36 of 85 Revision 22, 2010

IP2 FSAR UPDATE established by extrapolating along the correlation curve. The technique has been well established in providing core protection against low coolant flow in Westinghouse PWR plants. The expected absolute accuracy of the channel is within 10-percent and field results have shown the repeatability of the trip point to be within 1-percent. The analysis of the loss-of-flow transient is presented in Section 14.1.6. 4.2.10 Reactor Coolant Vent System 4.2.10.1 Design Basis The remote reactor coolant vent system has been designed and installed in accordance with NUREG-0737 Item II.B.1 to allow for remote manual venting of gases from the reactor vessel head should they accumulate there. The power-operated relief valve system acts as the remote operated vent system for the pressurizer (see Section 4.2.3) and as a redundant backup to the vessel head vent system. 4.2.10.2 System Description 4.2.10.2.1 Power-operated Relief Valve System The power-operated relief valve system is discussed in Sections 4.2.3 and 4.3.4. 4.2.10.2.2 Remote Reactor Head Vent System The original manual reactor vessel head vent line has been extended and two motor operated valves have been installed in series to facilitate venting of the reactor vessel head from the control room. The release point is located above the operating floor at an elevation of approximately 105-ft and is situated so that the discharge of the system will not impinge on any structures, systems, or components essential to the reactor safe shutdown or mitigation of a design basis accident. The power-operated relief valve system relieves to the pressurizer relief tank. The remote reactor head vent, and the power-operated relief valves and their associated block valves are in three separate lines and are supplied with three independent emergency power sources so that at least one vent path will remain functional after the single failure of an emergency power train. Potential seat leakage through both valves is vented directly to the containment atmosphere and is detected and monitored as part of the reactor coolant system leakage requirements specified in the Technical Specifications. The two series motor operated head vent valves are closed and deenergized during normal plant operation. The circuit breakers will be locked open to prevent inadvertent operation. If the need should arise for venting the reactor, the two breakers of the remote head vent valves will be reenergized and the valves opened as necessary from the central control room accident assessment panel. 4.2.10.3 Design Criteria The reactor coolant vent system piping, valves, components, and supports are classified seismic Class I and Class A. They have been designed and installed in accordance with the original requirements for reactor coolant pressure boundary installations, and ASME and ANSI Chapter 4, Page 37 of 85 Revision 22, 2010

IP2 FSAR UPDATE codes applicable to Indian Point Unit 2. The piping, valves, and fittings were fabricated from stainless steel and are compatible with reactor coolant chemistry. To alleviate the potential hazard of missiles, the remote reactor head vent system was installed such that it does not come close to and have the ability to damage safety-related systems required for safe reactor shutdown or mitigation of a design basis accident. 4.2.10.4 Design Evaluation Consistent with NUREG-0737 Item II.B.1 Clarification A(4), the new remote reactor head vent system was designed with sufficient flow restriction that in the event of inadvertent opening or line breaks, normal makeup charging flow from the chemical and volume control system is capable of precluding actuation of the safety injection system. The original reactor vessel head vent consisted of a 3/4-in. line with a manual (locally operated) shutoff valve and bolted blind flange and was used only for routine operations when the reactor was shut down. When the remotely operated head vent system was installed, the blind flange was removed and additional nominally 3/4-in. tubing (9/16-in. ID) was run from the existing 3/4-in. NPS line to the new motor-operated head vent valves. Those portions of the system, which were revised were designed and constructed to the same criteria as the original Indian Point Unit 2 pressure boundary components. A specific calculation has been performed for the worst case break location for the revised vent system (i.e., the interface between the 9/16-in. tubing and the original 3/4-in. head vent piping). This calculation determined that even at this worst case location, the break flow would be well within the capacity of two chemical and volume control system charging pumps without actuating safeguards equipment. Thus, failure of the vent system would not result in a break size corresponding to the definition of a loss-of-coolant accident. 4.2.11 Reactor Vessel Level Indication System The reactor vessel level indication system (RVLIS) has been installed in accordance with the requirements of NUREG-0737. The system is mainly part of the "Inadequate Core Cooling (ICC) Instrumentation" in improving the reliability of the plant operator to diagnose the approach of inadequate core cooling and to assess the adequacy of responses taken to restore cooling. The system also provides assistance to the operator in determining the presence of voids in the vessel. Additional information is given in Section 7.5.2. 4.2.11.1 Design Basis The system has been designed to provide continuous indication of coolant level inside the reactor and to assist the operator in determining the presence of voids in the vessel. The system was designed by Westinghouse as a Class A-Class 1E system. 4.2.11.2 System Description The RVLIS, shown in Plant Drawing 208798 [Formerly UFSAR Figure 4.2-12], is based on the differential pressure principle as sensed by taps located at the top and bottom of the reactor vessel. The top tap is installed on an unused control rod penetration and the bottom tap uses an unused incore instrument thimble to the seal table. The differential pressure is transmitted through filled capillary systems to transmitters outside containment. The temperature sensors are mounted on each capillary inside the containment. The signals from the temperature Chapter 4, Page 38 of 85 Revision 22, 2010

IP2 FSAR UPDATE sensors and transmitters are routed to a Class 1E panel in the cable spreading room. The temperature compensated signals of level indication are indicated on the accident assessment panel in the Unit 1/Unit 2 central control room. The reactor vessel level indication system, which was installed in response to NUREG-0737, Inadequate Core Cooling Instrumentation, has been approved by the NRC. REFERENCES FOR SECTION 4.2

1. E. L. Robinson, "Bursting Tests of Steam-Turbine Disk Wheels," Transactions of the ASME, July 1944.

2. D. H. Winne and B. M. Wundt, "Application of the Griffity-Irwin Theory of Crack Propagation to the Bursting Behavior of Discs, Including Analytical and Experimental Studies," ASME Paper Number 57-A-249, December 1, 1957.

3. Westinghouse Electric Corporation, "Analysis of Reactor Coolant System for Postulated Loss-of-Coolant Accident: Indian Point Unit 3 Nuclear Power Plant,"

WCAP-9117 (Proprietary) and WCAP-9130 (Nonproprietary).

4. Letter from W. J. Cahill, Con Edison, to V. Stello, NRC,

Subject:

Reactor Vessel Supports, dated June 15, 1978.

5. Letter from P. Zarakas, Con Edison, to T. Novak, NRC,

Subject:

Effects of Combining Seismic (SSE) and LOCA Responses on WCAP-9117 Analysis, dated September 3, 1980.

6. Letter from S. A. Varga, NRC, to J. D. O'Toole, Con Edison,

Subject:

Safety Evaluation Report on Asymmetric LOCA Loads, dated March 28, 1984.

7. S. Yanichko, "Indian Point 2 Reactor Vessel Radiation Surveillance Program,"

WCAP-7323, Westinghouse Electric Corporation, May 1969.

8. Letter from J. D. O'Toole, Con Edison, to S. A. Varga, NRC,

Subject:

Application for Amendment to Operating License, dated May 5, 1982.

9. Letter from J. D. O'Toole, Con Edison, to S. A. Varga, NRC,

Subject:

Reactor Vessel Material Surveillance Program, Analysis of Capsule Z, dated May 7, 1984.

10. Deleted (superceded by Reference 11).

11. "Reactor Vessel Material Surveillance Program for Indian Point Unit No. 2 Analysis of Capsule V", SWRI Final Report Project No. 17-2108, March 1990.

12. Letter from J.D. O'Toole, Con Edison, to S.A. Varga, NRC,

Subject:

Reference Transition Temperature For Pressurized Thermal Shock Evaluations, dated January 22, 1986.

13. Letter from Murray Selman, Con Edison, to Document Control Desk, NRC, transmitting supplemental information, dated January 12, 1987.

Chapter 4, Page 39 of 85 Revision 22, 2010

IP2 FSAR UPDATE

14. Letter from Marylee M. Slosson, NRC, to Murray Selman, Con Edison,

Subject:

Projected Values of Material Properties for Fracture Toughness Requirements for Protection Against Pressurized Thermal Shock Events, dated February 27, 1987.

15. Deleted

16. Deleted

17. Deleted

18. Letter from Stephen B. Bram, Con Edison, to Document Control Desk, NRC,

Subject:

Close-out for NRC Bulletin 88-11 Pressurizer Surge Line Stratification, dated October 1, 1991.

19. Letter from Jefferey F. Harold, NRC, to Paul H. Kinkel, Con Edison,

Subject:

Issuance of Amendment for Indian Point Nuclear Generating Unit No. 2 (TAC No. M96944), dated February 27, 1998.

20. Deleted

21. Letter from Stephen B. Bram, Con Edison, to Document Control Desk, NRC.

Subject:

Reactor Vessel Structural Integrity, 10 CFR 50.54(f), (Generic Letter 92-01, Revision 1), dated July 6, 1992.

22. Letter from J. O'Toole, Con Edison, to M Slosson, NRC, Dated June 10, 1986.

23. Letter from R. Capra, NRC, to S. Bram,, Con Edison, dated December 24, 1987,

Subject:

Indian Point Nuclear Generating Unit No. 2 Steam Generator Girth Weld Repair (TAC 66684).

24. Letter from R. Capra, NRC, to S. Bram,, Con Edison, dated October 28, 1988,

Subject:

Steam Generator Girth Weld Repair Safety Evaluation for Indian Point Nuclear Generating Unit No. 2 (TAC 66684).

25. Letter from R. Capra, NRC, to S. Bram,, Con Edison, dated January 4, 1991,

Subject:

Review of Mid-Cycle Steam Generator Girth Weld and Feedwater Inlet Nozzle Work (TAC 72961).

26. WCAP-15629, Rev. 1, Indian Point Unit 2 Heatup and Cooldown Curves for Normal Operation and PTLR Support Documentation, Westinghouse Electric Co., December 2001.

Chapter 4, Page 40 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-1 Materials of Construction of the Reactor Coolant System Components Component Section Materials Reactor vessel Pressure plate SA-302, Gr. B Shell and nozzle forgings SA-302, Gr. B / SA-336 Cladding, stainless weld rod Type 304 equivalent Thermal shield and internals A-240 Type 304 Stainless steel, Insulation Aluminum Steam generator Pressure plate SA-533, Grade A Class 2 Cladding, stainless weld rod Type 304 equivalent Cladding for tube sheets Inconel Tubes SB-163, Thermally Treated (Code Case N-20) Channel head castings SA-216 WCC Pressurizer Shell SA-302 Gr. B Heads SA-216 WCC External plate (support skirt) SA-516, Gr. 70 Cladding, stainless Type 304 equivalent Internal plate SA-240 Type 304 Spray Nozzle SA-376 Type 316 Pressurizer relief Shell A-285 Gr. C tank Heads A-285 Gr. C Internal surface coating Amercoat 55 system Piping Pipes A-376 Types 304 and 316 Fittings A-351 CF8M Nozzles A-182 Type F316 Pump Shaft Type 304 Impeller A-351 CF8 Casing A-351 CF8M Valves Pressure containing parts A-351 CF8 and CF8M; A-182 Type F316, and ASME SA182 Type F316 ASTM A479 Type 316 Chapter 4, Page 41 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-2 Identification of Indian Point Unit 2 Reactor Vessel Beltline Region Weld Metal Weld Welding Weld Weld Wire Flux Post-Weld Location Process Control Type Heat No. Type Lot Heat No. No. Treatment Nozzle shell Submerged - RACO W5214 Linde 3600 1125 +/- 25°F Arc 3 1092 vertical seam +Ni 200 N7048A 25 hr-FC 1-042 A, B, and C Inter shell Submerged - RACO W5214 Linde 3600 1125 +/- 25°F vertical Arc 3 1092 seam circle +Ni 200 N7048A 25 hr-FC seam 8-042 Nozzle shell to Submerged - RACO W5214 Linde 3600 1125 +/- 25°F Arc 3 1092 inter seam 2- +Ni 200 N7048A 25 hr-FC 042 A, B, and C Inter shell to Submerged M1.03 RACO 34B009 Linde 3708 1150 +/- 25°F lower Arc 3 1092 shell circle +Ni 200 N9867A 40 hr-FC seam 9-042 Lower shell Submerged - RACO W5214 Linde 3576 1150 +/- 25°F vertical Arc 3 1092 seams 3-042 A +Ni 200 - 40 hr-FC and B Chapter 4, Page 42 of 85 Revision 22, 2010

IP2 FSAR UPDATE Surveillance Submerged - RACO W5214 Linde 3600 1150 +/- 25°F weld Arc 3 1092

                                                               +Ni 200         N7048A                                   19 3/4 hr-FC TABLE 4.2-3 Chemical Composition of Reactor Vessel Beltline Region Weld Metal Weld Wire                          Flux                                           Weight Percent Type            Heat            Type              Lot      C       Mn        P         S      Si    Mo    Cr Ni Cu No.                               No.

RACO 3 W5214 Linde 1092 3600 .11 1.20 .021 .012 .19 .52 -- -- -- RACO 3 34B00 Linde 1092 3708 .14 2.01 .010 .017 .04 .51 -- -- --1 9 RACO 3 W5214 Linde 1092 3576 .12 1.15 .021 .012 .21 .56 -- -- -- Surveillance Weld - Not Performed Notes:

1. Chemical analysis of bare wire - No as-deposited analysis available.

TABLE 4.2-4 Mechanical Properties of Reactor Vessel Beltline Region Weld Metal Energy Shelf T RT NDT1 at 10°F ND Energ YS UTS Elong RA Weld Wire Flux (°F) (ft-lbs) T1 y (ksi) (ksi) Percent Percen (°F) (ft-lbs) t Type Heat Type Lot No. No. RACO 3 W5214 Linde 3600 0 103,93,95 0 -- 65.5 80.0 31.0 71.5 1092 RACO 3 34B00 Linde 3708 0 84,71,90 0 -- 67.9 84.2 31.0 69.8 9 1092 RACO 3 W5214 Linde 3576 0 57,51,69 0 -- 68.5 85.0 27.5 68.5 1092 Surveillance Weld 0 78,74,81 0 121 64.75 80.85 27.7 72.7 NOTES:

1. Estimated per NRC Standard Review Plan Section 5.3.2.

Chapter 4, Page 43 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-5 Maximum End-of-Life Fluence at Vessel Inner Wall Locations Seam or Fluence Plate or Weld Location Plate No. (n/cm2) Nozzle Shell Vertical Seam 1-042A 6.6 x 1017 Nozzle Shell Vertical Seam 1-042B 4.4 x 1017 Nozzle Shell Vertical Seam 1-042C 1.1 x 1018 Nozzle Shell to Inter. Shell Circle Seam 8-042 1.3 x 1018 Intermediate Shell Vertical Seam 2-042A 8.8 x 1018 Intermediate Shell Vertical Seam 2-042B 8.8 x 1018 Intermediate Shell Vertical Seam 2-042C 5.0 x 1018 Intermediate Shell to Lower Shell Circle Seam 9-042 1.6 x 1019 Lower Shell Vertical Seam 3-042A 7.0 x 1018 Lower Shell Vertical Seam 3-042B 7.0 x 1018 Nozzle Shell Plate B2001-1 1.3 x 1018 Nozzle Shell Plate B2001-2 1.3 x 1018 Nozzle Shell Plate B2001-3 1.3 x 1018 Intermediate Shell Plate B2002-1 1.6 x 1019 Intermediate Shell Plate B2002-2 1.6 x 1019 Intermediate Shell Plate B2002-3 1.6 x 1019 Lower Shell Plate B2003-1 1.6 x 1019 Lower Shell Plate B2003-2 1.6 x 1019 TABLE 4.2-6 Identification of Reactor Vessel Beltline Region Plate Material Mat'l Heat Treatment Compon Plate No. Heat Spec No. Supplier Austenitize Temper Stress Relief ent No. Nozzle B2001-1 B4679 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 60hr-FC Nozzle B2001-2 B4701 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 60hr-FC Nozzle B2001-3 A9870 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 50hr-FC Inter B2002-11 B4688 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 50hr-FC Inter B2002-21 B4701 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 50hr-FC Inter B2002-31 B4922 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 40hr-FC Lower B2003-1 B4791 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 40hr-FC Lower B2003-2 B4782 A302B Lukens 1550-1650°F 4 hr- 1200-1250°F 1125-1175°F shell Mod. WQ 4hr-AC 40hr-FC Notes:

1. Surveillance Material.

Chapter 4, Page 44 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-7 Chemical Composition of Reactor Vessel Beltline Region Plate Material, Weight Percent Plate No. C Mn P S Si Ni Mo Cu B2001-11 0.22 1.35 0.010 0.022 0.24 0.50 0.46 0.20 B2001-21 0.23 1.27 0.011 0.021 0.23 0.43 0.47 0.14 B2001-31 0.23 1.35 0.012 0.025 0.26 0.50 0.48 0.19 B2002-12 0.20 1.28 0.010 0.019 0.25 0.65 0.46 0.19 B2002-22 0.22 1.30 0.014 0.020 0.22 0.46 0.50 0.17 B2002-32 0.22 1.29 0.011 0.018 0.25 0.60 0.46 0.25 B2003-11 0.23 1.33 0.011 0.025 0.23 0.66 0.48 0.20 B2003-21 0.21 1.30 0.010 0.021 0.23 0.48 0.45 0.19 Notes:

1. Surveillance Material - No analysis performed other than reported by supplier.

2. Best estimate Cu and Ni weight percent 26.

Chapter 4, Page 45 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-8 Mechanical Properties of Reactor Vessel Beltline Region Plate Material Shelf T RT NDT NDT1 Energy1 YS UTS Elongation RA Plate No. (°F ) (°F ) (ft-lb) (ksi) (ksi) (percent) (percent) B2001-1 -10 24 69 67.25 87.75 26.00 64.45 B2001-2 -10 18 63.5 63.25 85.25 27.25 65.75 B2001-3 -10 25 69 65.25 86.75 25.00 63.75 B2002-1 -20 34 70 70.75 91.50 25.00 64.75 B2002-2 -30 21 73 65.00 85.25 26.50 67.00 B2002-3 -10 21 73.5 68.95 90.50 26.75 67.75 B2003-1 -20 20 71 65.75 87.25 27.75 65.50 B2003-2 -20 -20 88 61.25 81.60 30.75 70.50 B2002-1 - 34 76 67.17 88.40 25.20 67.6 B2002-2 - 34 75 64.55 87.15 27.65 69.8 Surveillance B2002-3 - 39 72.5 65.32 87.32 26.30 67.0 Test Data Notes:

1. Estimated from longitudinal data per NRC Standard Review Plan Section 5.3.2.

Chapter 4, Page 46 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-9 Summary of Charpy V-notch and Drop Weight Tests 30-ft-lb Drop Weight Component Grade Fix (°F ) NDT (°F ) Head dome A533B CL1 -2 10 Head peel segment A533B CL1 -10 10 Head peel segment A533B CL1 12 0 Upper shell plate A533B CL1 33 -10 Upper shell plate A533B CL1 31 -10 Upper shell plate A533B CL1 9 -10 Intermediate shell plate A533B CL1 14 -20 Intermediate shell plate A533B CL1 -11 -30 Intermediate shell plate A533B CL1 18 -10 Lower shell plate A533B CL1 -5 -20 Lower shell plate A533B CL1 -32 -20 Bottom peel segment A533B CL1 -12 -20 Bottom peel segment A533B CL1 -9 -10 Bottom dome A533B CL1 8 -30 Head flange A508 CL2 10 - Vessel flange A508 CL2 -18 - Inlet Nozzle A508 CL2 -102 - Inlet Nozzle A508 CL2 -84 - Inlet Nozzle A508 CL2 -95 - Inlet Nozzle A508 CL2 -51 - Outlet Nozzle A508 CL2 -32 - Outlet Nozzle A508 CL2 <10 - Outlet Nozzle A508 CL2 -45 - Outlet Nozzle A508 CL2 <10 - Chapter 4, Page 47 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.2-10 Reactor Vessel Beltline Fluence Fast Neutron Fluence (>1 MeV) 32 Effective Full Power Years (n/cm2)1 Reactor vessel Interior surface 1.39 X 1019 1/4 vessel thickness (1/4 T) 9.04 X 1018 3/4 vessel thickness (3/4 T) 3.48 X 1018 Notes:

1. These values are calculated based upon experimental results from the measurements on the fourth surveillance capsule V. See Reference 11.

4.2 FIGURES Figure No. Title Figure 4.2-1 Reactor Coolant System Flow Diagram - Replaced with Plant Drawing 9321-2738 Figure 4.2-2 Reactor Coolant System Schematic Flow Diagram Figure 4.2-3 Reactor Vessel Figure 4.2-4 Pressurizer Figure 4.2-5 Steam Generator Assembly Figure 4.2-6 Reactor Coolant Pump Figure 4.2-7 Reactor Coolant Pump Estimated Performance Characteristics Figure 4.2-8 Flywheel Figure 4.2-9 Reactor Coolant Pump Flywheel Tangential Stress vs Radius Figure 4.2-10 Pressurizer Relief Tank Figure 4.2-11 Identification & Location of Beltline Region Material for the Indian Point Unit 2 Reactor Vessel Figure 4.2-12 Reactor Vessel Level Instrumentation System Flow Diagram - Replaced with Plant Drawing 208798 4.3 SYSTEM DESIGN EVALUATION 4.3.1 Safety Factors The safety of the reactor vessel and all other reactor coolant system pressure-containing components and piping is dependent on several major factors including design and stress analysis, material selection and fabrication, quality control, and operations control. Chapter 4, Page 48 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.3.1.1 Reactor Vessel A stress evaluation of the reactor vessel has been carried out in accordance with the rules of Section III of the ASME Nuclear Pressure Vessel Code. The evaluation demonstrates that stress levels are within the stress limits of the code. Table 4.3-1 presents a summary of the results of the stress evaluation. The most significant transients with regard to cumulative fatigue of the reactor vessel are loss of load transient and loss-of-flow transients. A summary of fatigue usage factors for components of the reactor vessel is given in Table 4.3-2. The effect of gamma-ray heating on the cumulative usage factor is negligible. The cycles specified for the fatigue analysis are the results of an evaluation of the expected plant operation coupled with experience from nuclear power plants now in service. These cycles include five heatup and cooldown cycles per year, a conservative selection considering that the vessel may not complete more than one cycle per year during normal operation. The vessel design pressure is 2485 psig, while the normal operating pressure will be 2235 psig. The resulting operating membrane stress is therefore amply below the code-allowable membrane stress to account for operating pressure transients. To preclude the possibility of brittle failure, a reactor vessel material surveillance program that meets the requirements of 10CFR50 App. H, is implemented to monitor the change in reactor vessel materials due to neutron radiation. The radiation induced shift in Reference Temperature nil-ductility transition (RTNDT) is periodically assessed during the life of the plant by testing of vessel material samples that are irradiated cumulatively by securing them near the inside wall of the vessel in the core area. Regulatory Guide 1.99, Rev.2, Radiation Embrittlement of Reactor Vessel Materials, is utilized to predict the radiation induced change in the (RTNDT) and calculate a new Adjusted Reference Temperature (ART). To compensate for any increase in the (RTNDT) caused by irradiation, the heatup and cooldown pressure temperature limits given in the Technical Specifications are periodically changed to comply with 10CFR50, Appendix G, Fracture Toughness Requirements. The vessel closure contains fifty-four 7-in. studs. The stud material has a minimum yield strength of 104,100 psi at design temperature. The membrane stress in the studs when they are at the steady state operational condition is approximately 40,000 psi. This means that 21 of the 54 studs have the capability of withstanding the hydrostatic end load on the vessel head without the membrane stress exceeding yield strength of the stud material at design temperature. The normal operating temperature always exceeds even the highest anticipated DTT during the life of the plant. Thus the emphasis of conservative operation is placed on heatup and cooldown because long term irradiation of the vessel raises the DTT and thereby limits the heatup or cooldown rates. The conservatism in setting up the temperature-pressure relationship limits stated above are:

1. Use of a stress concentration factor of 4 on assumed flaws in calculating the stresses.

2. Use of nominal yield of material instead of actual yield.

Chapter 4, Page 49 of 85 Revision 22, 2010

IP2 FSAR UPDATE

3. Neglecting the increase in yield strength resulting from radiation effects.

4.3.1.2 Steam Generators Calculations confirm that the steam generator tube sheet will withstand the loading (which is a quasi-static rather than a shock loading) by loss of reactor coolant. The rupture of primary or secondary piping has been assumed to impose a maximum pressure differential of 2485 psi across the tubes and tube sheet from the primary side or a maximum pressure differential of 1035 psi across the tubes and tube sheet from the secondary side, respectively. Under these conditions there is no rupture of the primary to secondary boundary (tubes and tube sheet). This criterion prevents any violation of the containment boundary. An examination of stresses under these conditions shows that for the case of a 2485 psi maximum tube sheet pressure differential, the stresses are within acceptable limits. The tubes were designed to the requirements (including stress limitations) of Section III for normal operation, assuming 1700 psi as the normal operating pressure differential. Hence, the secondary pressure loss accident condition imposes no extraordinary stress on the tubes beyond that normally expected and considered in Section III requirements. An evaluation determined the extent of tube wall thinning that could be tolerated under accident conditions. The worst-case loading conditions are assumed to be imposed upon uniformly thinned tubes at the most critical location in the steam generator. Under such a postulated design basis accident, vibration is short enough duration that there is no endurance issue to be considered. The steam generator tubes, existing originally at their minimum wall thickness and reduced by a conservative general corrosion and erosion loss, provide an adequate safety margin (sufficient wall thickness) in addition to the minimum required for a maximum stress less than the allowable stress limit, as defined by the ASME Code. Studies have been made on tubing of the size in the replacement steam generators under accident loadings. The results show that the maximum Level D Service condition stress due to combined pipe rupture and safe shutdown earthquake loads is less than the allowable limit. The tube thickness required to achieve the acceptable stress is less than the minimum steam generator tube wall thickness, which is reduced to account for assumed general corrosion and erosion rate. Thus, an adequate safety margin is exhibited. The general corrosion rate is based on a conservative weight-loss rate for Alloy 600 tubing in flowing, 650°F primary-side reactor coolant fluid. The estimated weight loss, based on testing when equated to a thinning rate and projected over a 60-year design objective, is much less than the assumed corrosion allowance of 3 mils. This leaves the remainder of the general corrosion allowance for thinning on the secondary side. Potential sources of tube excitation are considered, including primary fluid flow within the U-tubes, mechanically induced vibration, and secondary fluid flow on the outside of the U-tubes. The effects of primary fluid flow and mechanically induced vibration, including those developed by the canned-motor pump, are acceptable during normal operation. The primary source of potential tube degradation due to vibration is the hydrodynamic excitation of the tubes by the secondary fluid. This area has been emphasized in both analyses and tests, including evaluation of steam generator operating experience. Chapter 4, Page 50 of 85 Revision 22, 2010

IP2 FSAR UPDATE Three potential tube vibration mechanisms related to hydrodynamic excitation of the tubes have been identified and evaluated. These include potential flow-induced vibrations resulting from vortex shedding, turbulence, and fluid-elastic vibration mechanisms. Nonuniform, two-phase turbulent flow exists throughout most of the tube bundle. Therefore, vortex shedding is possible only for the outer few rows of the inlet region. Moderate tube response caused by vortex shedding is observed in some carefully controlled laboratory tests on idealized tube arrays. However, no evidence of tube response caused by vortex shedding is observed in steam generator scale model tests simulating the inlet region. Bounding calculations consistent with laboratory test parameters confirmed that vibration amplitudes would be acceptably small, even if the carefully controlled laboratory conditions were unexpectedly reproduced in the steam generator. Flow-induced vibrations due to flow turbulence are also small. Root mean square amplitudes are less than allowances used in tube sizing. These vibrations cause stresses that are two orders of magnitude below fatigue limits for the tubing material. Therefore, neither unacceptable tube wear nor fatigue degradation due to secondary flow turbulence is anticipated. Fluid elastic tube vibration is potentially more severe than either vortex shedding or turbulence because it is a self-excited mechanism. Relatively large tube amplitudes can feed back proportionally large tube driving forces if an instability threshold is exceeded. Tube support spacing in both the tube support plates and the anti-vibration bars in the U-bend region provides tube response frequencies such that the instability threshold is not exceeded for secondary fluid flow conditions for tubes effectively supported. This approach provides large margins against initiation of fluid elastic vibration for tubes effectively supported by the tube support system. Small clearances between the tubes and the supporting structure are required for steam generator fabrication. These clearances introduce the potential that any given tube support location may not be totally effective in restraining tube motion if there is a finite gap around the tube at that location. Fluid-elastic tube response within available support clearances is therefore theoretically possible if secondary flow conditions exceed the instability threshold when no support is assumed at the location with a gap around the tube. This potential has been investigated both with tests and analyses for the U-bend region where secondary flow conditions have the potential to exceed the instability threshold if a tube does not contact provided supports as a result of fabrication tolerances. Tube vibration response is shown to have wear potential within available design margins even for limiting tube fit-up conditions, based on previous experience in fabricating steam generators with fit-up control typical of the replacement steam generator. The replacement steam generator includes a number of features that minimize the potential for tube wear at tube supports. Provisions to minimize the potential for wear include the spacing between the tube supports, the configuration of the broached hole through the support plate, the surface finish of the broached hole in the tube support plate, the clearance between the tube and the hole in the tube support plate, and the tube support plate material selection. Tube bending stresses corresponding to tube vibration response remain more than two orders of magnitude below fatigue limits as a consequence of vibration amplitudes constrained by available clearances. The analyses and tests for limiting postulated fit-up conditions include simultaneous contributions from flow turbulence. Chapter 4, Page 51 of 85 Revision 22, 2010

IP2 FSAR UPDATE As outlined, analyses and tests demonstrate that unacceptable tube degradation resulting from tube vibration is not expected for the replacement steam generators. Operating experience with steam generators having the same size tubes and similar flow conditions supports this conclusion. The U-bend fatigue (discussed in NRC Bulletin 88-02) is not a consideration in the replacement steam generators. The mechanism considered in Bulletin 88-02 requires denting of the top tube support plate. But this is not expected with the stainless steel tube support plates in the replacement steam generator. The stress limits for Service Level D that allow inelastic deformation are supplemented with the requirements of "Rules for Evaluation of Service Loadings with Level D Service Limits," Appendix F of ASME Code, Section III. The limits and rules of Appendix F confirm that pressure boundary integrity and core support structural integrity are maintained but do not confirm operability. The limits and rules of Appendix F do not apply to the portion of the component or support in which the failure has been postulated. The structural stress analyses performed on the replacement steam generators consider the loadings specified. These loads result from thermal expansion, pressure, weight, earthquake, pipe rupture, and plant operational thermal and pressure transients. Dynamic effects of pipe rupture, including the loss of coolant accident, are not included in loading combinations when the leak-before-break criteria are satisfied. The combination of safe shutdown earthquake plus pipe rupture loads by square-root-sum-of-the-squares is considered. The dynamic effects of pipe rupture that are combined with safe shutdown earthquake in loading combinations are combined using the square-root-sum-of-the-squares method. The integrity of the pressure boundary of safety-related components is provided by the use of the ASME Code. The replacement steam generators, including the transition cone, lower shell, tubesheet, and channel head are constructed to the requirements of the ASME Code, Section III, 1980 Edition, plus winter 1981 addenda, which is reconciled to the design code of record, the 1965 ASME Boiler and Pressure Vessel Code, Section III, plus Addenda thru summer 1966. Using the methods and equations in the ASME Code, stress levels in the components and supports are calculated for various load combinations. These load combinations may include the effects of internal pressure, dead weight of the component and insulation, and fluid, thermal expansion, dynamic loads due to seismic motion, and other loads. The evaluation of the stress levels and fatigue usage for the steam generator pressure boundary is calculated for the specified loading conditions and demonstrates that the values are less than the allowable limits. These calculations are documented in a Stress Report as required by the ASME Code. Evaluation of the secondary shell in contact with secondary water assumes a 0.050 inch corrosion allowance. The analysis of the support plates assumes no corrosion allowance. The ASME Code, Section III requires that a design specification be prepared for ASME components. The specification conforms to and is certified to the requirements of ASME Code, Section III. The Code also requires a design report for safety-related components, to demonstrate that the as-built component meets the requirements of the relevant ASME Design Specification and the applicable ASME Code. The design specifications and design reports will be completed by the Combined License applicant or his agent. Design specifications for ASME components and piping are prepared utilizing procedures that meet the ASME Code. The design report includes as-built reconciliation. Chapter 4, Page 52 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.3.1.3 Piping The reactor coolant system piping has been designed for normal and emergency conditions. For the emergency condition, the piping has been designed and analyzed for seismic loads and blowdown forces due to a loss-of-coolant accident. By design, the main piping of the reactor coolant loop is not subjected to induced pressure pulse vibrations from the reactor coolant pump impeller or from the pistons of the charging pump. In 1989, the NRC approved changes to the design bases with respect to dynamic affects of postulated primary loop pipe ruptures, as discussed in Section 4.1.2.4. 4.3.2 Reliance On Interconnected Systems The principal heat removal systems, which are interconnected with the reactor coolant system are the steam and power conversion, the safety injection, and residual heat removal systems. The reactor coolant system is dependent upon the steam generators, and the steam, feedwater, and condensate systems for decay heat removal from normal operating conditions to a reactor coolant temperature of approximately 350oF. The layout of the system ensures the natural circulation capability to permit adequate core cooling following a loss of power to all main reactor coolant pumps. Further details are given in Section 14.1.6.1. The NRC reviewed the Indian Point 2 response to issues concerning natural circulation cooldown in their safety evaluation report dated August 1, 1983 (reference 20), and determined that Con Edison met the requirements of Generic Letter 81-21. The flow diagram of the steam and power conversion system is shown on Plant Drawings 227780, 9321-2017, 235308 [Formerly UFSAR Figure 10.2-1 sheets 1 to 3], 9321-2025 [Formerly UFSAR Figure 10.2-4], 9321-2018 and 235307 [Formerly UFSAR Figure 10.2-5 sheets 1 and 2], and 9321-2019 [Formerly UFSAR Figure 10.2-7]. In the event that the condensers are not available to receive the steam generated by residual heat, the water stored in the feedwater system may be pumped into the steam generators and the resultant steam vented to the atmosphere. The auxiliary feedwater system will supply water to the steam generators in the event that the main feedwater system is unavailable. The safety injection system is described in Section 6.2. The residual heat removal system is described in Section 9.3. 4.3.3 System Integrity A complete stress analysis that reflects consideration of all design loadings detailed in the design specification has been prepared by the manufacturer. The analysis shows that the reactor vessel, steam generator, reactor coolant pump casing, and pressurizer comply with the stress limits of Section III of the ASME Code. A similar analysis of the piping shows that it complies with the stress limits of the applicable USAS Code. As part of the design control on materials, Charpy V-notch toughness test curves were run on all ferritic material used in fabricating pressure parts of the reactor vessel, steam generator, and pressurizer to provide assurance for hydrotesting and operation in the ductile region at all times. In addition, drop weight tests were performed on the reactor vessel plate material. As an Chapter 4, Page 53 of 85 Revision 22, 2010

IP2 FSAR UPDATE assurance of system integrity, all components in the system were hydrotested at 3110 psig prior to initial operation. 4.3.4 Overpressure Protection The reactor coolant system is protected against overpressure by safety valves located on the top of the pressurizer. The safety valves on the pressurizer are sized to prevent system pressure from exceeding the design pressure by more than 10-percent, in accordance with Section III of the ASME Boiler and Pressure Vessel Code. The capacity of the pressurizer safety valves is determined from considerations of: (1) the reactor protective system, and (2) accident or transient conditions, which may potentially cause overpressure. The combined capacity of the safety valves is equal to or greater than the maximum surge rate resulting from complete loss of load without a direct reactor trip or any other control, except that the safety valves on the secondary plant are assumed to open when the steam pressure reaches the secondary plant safety valve setting. Details of the analysis are reported in Section 14.1.8. Experience has shown that the safety valve capacity so determined is adequate for all the other transients as the results of Section 14.1 show. The report "Summary Report of Safety and Relief Valve Installation and Re-Analysis for ASME Class 1 and Class 2 Systems in Indian Point Unit No. 2" (Reference 5) describes the general scope, design and installation criteria, significant assumptions, methods of analysis, and maximum combined stresses for those applicable safety and relief valves in the reactor coolant system, main steam system, chemical and volume control system, safety injection system, component cooling water system, and service water system. In response to NUREG-0737 Section II.D.1, a test program for the pressurizer safety and relief valves was formulated by the Electric Power Research Institute to provide full-scale test data confirming the functional ability of the reactor coolant system power-operated relief valves and safety valves for expected operating and accident conditions, and to obtain sufficient piping thermal-hydraulic load data to permit confirmation of models that may be used for plant-unique analysis of safety and relief discharge piping systems. The Indian Point 2 plant-specific evaluations regarding this generic issue are contained in Con Edison submittals to NRC dated July 1, 1982, September 15, 1982, June 15, 1984, June 14, 1985 and October 18, 1985. This program satisfied the requirements of NUREG-0737, as documented in NRC's Safety Evaluation Report (SER) dated August 5, 1987 (Reference 16). Item II.K.3.2 of NUREG-0737 required licensees of pressurized water reactors to submit a report to the NRC staff documenting the various actions taken to decrease the probability of a small break LOCA caused by a stuck-open PORV and show how these actions constitute sufficient improvements to safety. Based upon the results of the report submitted in response to item II.K.3.2, licensees were to assess whether an automatic PORV isolation system was required. If required, licensees were to submit a system design that uses the PORV block valve to automatically protect against a small break LOCA caused by a stuck open PORV. The Westinghouse Owners Group submitted a generic report to the NRC staff in response to Item II.K.3.2 (Reference 17). Con Edison's response to the NRC on this matter (Reference 18) adopted the conclusions reached in the aforementioned report as applicable to IP2, namely that the concept of an automatic PORV block valve closure system cannot be warranted on the Chapter 4, Page 54 of 85 Revision 22, 2010

IP2 FSAR UPDATE basis of providing additional protection against a PORV LOCA. On this basis, Con Edison proposed no modifications to provide automatic isolation of the PORVs. The NRC reviewed Con Edison's submittal and found that the requirements of NUREG-0737, Item II.K.3.2 were met with the existing PORV, safety valve and reactor high pressure trip setpoints and that an automatic PORV isolation system was not required for IP2 (Reference 19). 4.3.4.1 Reactor Coolant System Overpressure Protection System An overpressure protection system to prevent reactor coolant system pressure exceeding the 10 CFR 50 Appendix G curves has been installed. It is a three-channel, analog, curve-tracking arrangement, which would initiate an appropriate chain of coincidence logic for the purpose of automatically preventing a violation of the operating Technical Specifications temperature/pressure curves for the reactor vessel. In order to develop the overpressure protection system setpoint limit curve for the Technical Specifications, heatup and cooldown limit curves are calculated (Ref.21), using the adjusted RTNDT (reference nil-ductility temperature) corresponding to the limiting beltline region material of the reactor vessel. The adjusted RTNDT (ART) of the limiting material in the core region of the reactor vessel is determined by using unirradiated reactor vessel material fracture toughness properties, estimating the radiation induced change in RTNDT, and adding margin for uncertainty. The unirradiated RTNDT is designed as the higher of either the drop weight nil ductility transition temperature (NDTT) or the temperature at which the material exhibits at least 50 ft-lb of impact energy and 35 mils lateral expansion (transverse to the primary working direction), less 60F degrees. The method used to calculate the ART values at 1/4T and 3/4T locations, (where T is the thickness of the reactor vessel at the beltline region not including the cladding), complies with Nuclear Regulatory Commission Regulatory Guide 1.99, Revision 2, Radiation Embrittlement of Reactor Vessel Materials. The heatup and cooldown curves are generated using the most limiting ART values and the methodology documented in Westinghouse Report WCAP-14040-NP-A, Rev.2, Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown limit curves, with the following exceptions:

1) The fluence values used are calculated fluence values, rather than best estimate fluence values.

2) The Kic critical stress intensities are used in place of Kia critical stress intensities, in compliance with ASME Code Case N-640,

3) The 1996 version of Appendix G to ASME Section XI is used instead of the 1989 version, and

4) Pressure-temperature limit curves were generated with the most limiting circumferential weld ART in conjunction with Code Case N-588. These curves are bounded by curves using the standard axial flaw methodology of the ASME Code 1996, App. G with the ART from the limiting plate material.

The heatup and cooldown pressure-temperature limit curves (10CFR50 App. G limits) so obtained (Ref.21) are valid for 29.2 EFPY. Thermal-hydraulic analysis (Ref.22) accounting for the effects of pressure bias and pressure overshoot during mass and heat input transients, is utilized to develop setpoint curves to actuate the Power Operated Relief Valves (PORVs) and prevent the reactor coolant system from exceeding the 10 CFR50 App. G limits. The analysis demonstrates that a single power operated relief valve is capable of mitigating the worst possible mass or heat input transient, Chapter 4, Page 55 of 85 Revision 22, 2010

IP2 FSAR UPDATE thereby ensuring that the peak reactor coolant system pressure for the Indian Point Unit No. 2 remains below the 10 CFR50 Appendix G limits were such a transient to occur. Overpressure Protection System setpoint curves are further adjusted for instrument error, and these latter curves are utilized as heatup and cooldown limits in plant operating procedures. Also, additional administrative controls are utilized to protect the Residual Heat Removal System from reactor coolant overpressurization events. The overpressure protection system does not change the primary system operation or relief system operation during normal plant operation. The system allows for more close control of system heatup and cooldown through more accurate instrumentation and monitoring. Spurious opening and/or closing of the power operated relief valves is essentially eliminated by the new two-out-of-three logic (if one channel were to fail the valve would not malfunction). Thus, the probability of occurrence or the consequences of an accident or malfunction of equipment important to safety previously evaluated is not increased. The overpressure protection system maintains the existing operational function of all existing plant components. There is an expansion of certain component functions to enhance the controllability of primary system pressure during heatup and cooldown. Inasmuch as there is better control of existing plant components and no change to their operation, the possibility for an accident or malfunction of a different type than any evaluated previously is not increased. The overpressure protection system allows heatup and cooldown guidelines to be strictly followed both by automatic and manual means, thereby reducing the possibility of violating significant parameters and maintain an orderly heatup and cooldown. Thus the margin of safety as defined in the bases for the facility Technical Specifications is not reduced. NRC acceptance of the Indian Point 2 low temperature overpressure protection system and the relevant Consolidated Edison submittals are contained in References 12, 13, 14 and 15. 4.3.4.2 Nitrogen System A nitrogen system actuates the power operated relief valves (PORVs) PCV-455C and PCV-456. The PORV nitrogen system is tapped from the existing nitrogen supply header to Safety Injection (SI) system accumulators at a location downstream of pressure regulator valve PCV-942 and relief valve RV-1816, which is set at 1100 psig. The nitrogen pressure to the PORVs is reduced to 100 psig by pressure regulator valves PRV-3100 and PRV-3101. Containment isolation of the PORV nitrogen system is provided by valves 4312 and 863. The instrument nitrogen system includes two accumulators, each holding approximately 13-ft3 of nitrogen. In case the nitrogen supply is lost, these accumulators, with a minimum initial pressure of 600 psig, can support cycling (full open/close) the power operated relief valves for a minimum of 10 minutes. The nitrogen system is provided with pressure indicating alarms located on the SKF panel in the control room to provide information to the operator in case of low pressure in the nitrogen accumulators. Pressure alarms also on the SKF panel provide indication for nitrogen supply regulator malfunction. The PORV nitrogen system is designated Class A, Seismic Class I. The accumulators are designed to ASME Section VIII, Div.1 and piping is designed to ANSI B31.1. The PORV nitrogen system piping can withstand 1100 psi, the relief setting of valve RV-1816. The design Chapter 4, Page 56 of 85 Revision 22, 2010

IP2 FSAR UPDATE of the PORV nitrogen system further considered the potential to generate missiles. To ensure that none of the components of the nitrogen system would become a source of missiles, the valves are forged, have a bolted clamseal bonnet and have stems which back seat. This rules out the possibility of ejecting valve stems as, even if it were assumed that the stem threads fail, the back seat or the upset end cannot penetrate the bonnet and thereby become a missile. Also, the valves have been designed against bonnet-body connection failure and subsequent bonnet ejection by means of (1) using the design practice of ASME Section VIII, which limits the allowable stress of bolting material to less than 20-percent of its yield strength, (2) using the design practice of ASME Section VIII for flange design, and (3) by controlling the load during the bonnet-body connection stud tightening process. The pressure-containing parts except the flange and studs are designed per criteria established by USAS B16.5. Flanges and studs, where used, are designed in accordance with ASME Section VIII. 4.3.4.3 Evaluation of the Overpressure Protection System With the overpressure protection system enabled, the power-operated relief valves will open automatically to prevent the reactor pressure vessel pressure from exceeding the Appendix G limits during a temperature range and for the effective full power years as defined in the Indian Point Unit 2 Technical Specifications, and there is a pressure excursion over the setpoint. With the Overpressure Protection System enabled, the power-operated relief valve isolation motor-operated valves are in the open position. Existing wide-range cold leg reactor coolant system temperature signals (TE-413, 433, and 443) are designed to perform two primary functions in this system: (1) provide the arming and disarming function and (2) serve as the independent variable in computing the reference Appendix G limit to which the system pressure limit must be adhered. The arming function is initiated when the reactor coolant system temperature falls below a temperature defined by the Technical Specification. At the OPS enable temperature, the motor-operated valves (MOV-535 and 536) on the pressurizer will either be manually or automatically opened and the overpressure protection system logic system will be armed to prevent a possible overpressurization condition. Also, one half of a two-out-of-two coincidence logic will be satisfied to allow the relief valves to open in the event of an impending overpressure condition. These same temperature signals are also fed into three respective function generators whose task is to output values of pressure as a function of the input temperature, which are the maximum reactor coolant system pressures (Appendix G limit pressures) allowed at those temperatures. The difference between these maximum permissible reactor coolant system pressures and the actual reactor coolant system pressure, transmitted by 0 to 1500-psig transmitters (PT-413, 433, 443), is computed in each of the three channels, and if any two-out-of-three of these differences is smaller than a preset minimum, a trip open condition will be initiated for each pressurizer power operated relief valve designated as train "A" (MOV-536 and PCV-456) and train "B" (MOV-535 and PCV-455C). Various alarms and lights related to reactor coolant system overpressurization arming, actuation or nonavailability of train "A" or "B" are located in the SGF and FB panels. The alarms to indicate arming of the reactor coolant system overpressurization trains and actuation of the reactor coolant system overpressurization train "A" and train "B" are located on the SG panel. The motor-operated valves can be closed in the armed region by putting the motor-operated valves selector switch into the full locked position. White lights (one for each Chapter 4, Page 57 of 85 Revision 22, 2010

IP2 FSAR UPDATE train), which indicate that the reactor coolant system overpressurization train is not available are located on the FB panel above the control switches. As a protection against a common air supply failure causing inoperability of both power operated relief valves, the air system has been replaced by a nitrogen system with accumulators to supply each valve. (See description of nitrogen supply system, Section 4.3.4.2.) The electrical supply is from the 125VDC power panels, which are supplied by 480VAC through 125VDC battery chargers with backup by the station emergency batteries. The electrical activation uses two-out-of-three logic for valve actuation. Manual disconnect switches provide a means to interrupt the power to a SOV, which will then result in the closure of the associated PORV. Operation with the switches closed permits the PORVs to open or close automatically or be manually operated to perform their pressure relief function. In the event of a fire in most of the fire zones in Fire Area A, these switches can be manually opened to prevent or mitigate the spurious opening of the PORVs due to a hot short in the control circuitry. To ensure that the PORVs can perform their pressure relief function, the block valves are interlocked to open automatically when the pressurizer pressure reaches a preset limit below the pressure at which the PORVs open. In addition, the PORV actuation and reclosure setpoint calibration is checked each 24 months. Operation with the PORVs and block valves closed will prevent the spurious opening of both a PORV and its associated block valve in the event of a fire in certain fire zones. 4.3.5 Incident Potential The potential of the reactor coolant system as a cause of accidents is evaluated by investigating the consequences of certain credible types of components and control failures as discussed in Sections 14.1 and 14.2. Reactor coolant pipe rupture is evaluated in Section 14.3. 4.3.6 Redundancy Each loop of the reactor coolant system contains a steam generator and a reactor coolant pump. Operation at reduced reactor power is possible with one loop out of service as limited by the facility Technical Specifications. For added reliability, power to the reactor coolant pumps is normally supplied by electrically separated buses as shown in Plant Drawing 231592 [Formerly UFSAR Figure 8.2-5]. The remote reactor head vent valves and the power operated relief valves and block valves are supplied with diverse and independent emergency power sources as described in Section 4.2.10. REFERENCES FOR SECTION 4.3

1. H. Kihara and K. Masubuchi, "Effect of Residual Stress on Brittle Fracture,"

Welding Journal, Volume 38, April 1959.

2. Robertson, "Propagation From Brittle Fracture in Steel," Journal of the Iron and Steel Institute, 1953.

3. L. Porse, "Reactor Vessel Design Considering Radiation Effects," ASME Paper Number 63-WA-100.

4. Deleted Chapter 4, Page 58 of 85 Revision 22, 2010

IP2 FSAR UPDATE

5. Letter from W. J. Cahill, Con Edison, to J. F. O'Leary, NRC,

Subject:

"Summary Report of Safety and Relief Valve Installation and Re-Analysis of ASME Class 1 and Class 2 Systems in Indian Point Unit 2," dated July 13, 1972.

6. Deleted

7. Deleted

8. Deleted

9. Westinghouse Electric Corporation, "Consolidated Edison Company of New York, Inc., Indian Point Unit 2 NSSS Stretch Rating - 3083.4 MWT Engineering Report," WCAP 12187 (Proprietary).

10. Deleted

11. Deleted

12. Letter (with attachments) from A. Schwencer, NRC to W. Butler, NRC,

Subject:

Safety Evaluation of the Electrical, Instrumentation and Control (EI&C) Aspects of the Low Temperature Overpressure Protection System, dated June 14, 1978.

13. Letter (with attachments) S. Varga, NRC to J. OToole, Con Edison,

Subject:

Indian Point 2- Low Temperature Overpressure Protection System, dated April 24, 1984.

14. Letter (with attachments) S. Varga, NRC to J. OToole, Con Edison,

Subject:

Low Temperature Overpressure Protection System at Indian Point Nuclear Generating Plant, Unit 2, dated June 28, 1984.

15. Letter (with attachments) J. OToole, Con Edison to S. Varga, NRC,

Subject:

Response to NRC Safety Evaluation Report, dated July 17, 1984.

16. Letter (with attachments) from M. Slosson, NRC to M. Selman, Con Edison,

Subject:

Safety Evaluation Report, NUREG-0737, Item II.D.1, Performance Testing of Relief and Safety Valves for Indian Point Nuclear Generating Unit No. 2, dated August 5, 1987.

17. WCAP-9804, "Probabilistic Analysis and Operational Data is Response to NUREG-0737, Item II.K.3.2, for Westinghouse NSSS Plants", Westinghouse Electric Corporation, February 1981

18. Letter (with attachments) from J. D. O'Toole, Con Edison, to D. G. Eisenhut, NRC, Response to NUREG-0737, Clarification of TMI Action Plan Requirements, Dated February 26, 1981

19. Letter (with attachments) from S. A. Varga, NRC, to J. D. O'Toole, Con Edison,

Subject:

NUREG-0737 Items II.K.3.1 - Automatic PORV Isolation and II.K.3.2 - Report on PORVs the Indian Point Nuclear Generating Plant, Dated September 13, 1983 Chapter 4, Page 59 of 85 Revision 22, 2010

IP2 FSAR UPDATE

20. Letter (with attachment) from S.A. Varga, NRC, to J.D. OToole, Con Edison,

Subject:

Natural Circulation Cooldown For The Indian Point Nuclear Generating Plant, Unit No. 2 (IP-2), Dated August 1, 1983.

21. Westinghouse Electric Company Report: WCAP16752-NP, Indian Point Unit 2 Heatup and Cooldown Limit Curves for Normal Operation, January 2008.

22. Con Edison Calculation No. FMX-00270-01, Indian Point Unit 2 Overpressure Protection System (OPS) Thermal Hydraulic Analysis, Setpoint Development and Technical Specification Revision February 26, 2009.

Chapter 4, Page 60 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.3-1 Summary of Primary Plus Secondary Stress Intensity for Components of the Reactor Vessel Allowable Stress 3Sm(psi) Area Stress Intensity (psi) (Operating Temperature) Control rod housing 77,700 (1) 69,900 Head flange 45,370 80,100 Vessel flange 52,140 80,100 Closure studs 109,400 110,400 Primary nozzles - inlet 45,500 80,100 outlet 49,390 80,100 Core support pad 55,280 69,900 Bottom head to shell 34,100 80,100 Bottom instrumentation 55,500 69,900 Nozzle belt to shell 37,900 80,100 Head Adapter Plugs 27,630 48,600 Note:

1. A simplified elastic plastic analysis was performed to justify exceeding the 3Sm limit.

TABLE 4.3-2 Summary of Cumulative Fatigue Usage Factors for Components of the Reactor Vessel Item Usage Factor1 Control rod housing 0.01 Head flange 0.0107 Vessel flange 0.0229 Stud bolts 0.944 Primary nozzles - inlet 0.050 outlet 0.281 Core support pad (lateral) 0.904 Bottom head to shell 0.004 Bottom instrumentation 0.201 Nozzle belt to shell 0.0029 Head Adapter Plugs 0.0036 Notes:

1. As defined in Section III of the 1965 ASME Boiler and Pressure Vessel Code, Nuclear Vessels.

TABLE 4.3-3 DELETED TABLE 4.3-4 DELETED Chapter 4, Page 61 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.4 SAFETY LIMITS AND CONDITIONS 4.4.1 System Heatup And Cooldown Rates Operating limits for the reactor coolant system with respect to heatup and cooldown rates are defined in the Technical Specifications. The stress level of material in the reactor vessel, or in other reactor coolant system components, is a combination of stresses caused by internal pressures and by thermal gradients. The latter are significant as they may result from a rate of change of reactor coolant temperature. Operating restrictions are imposed to limit the combined stresses to 20-percent of minimum yield stress when at the design transition temperature (DTT). The DTT is defined as the initial nil-ductility transition temperature (NDTT) plus the increase in NDTT due to irradiation experienced, plus 60oF. This stress limit (20-percent of yield strength) is reduced linearly to a value of 10-percent of yield at a temperature 200oF below DTT. Curves are incorporated in the plant operating procedures, which define the operating limits for initial operation and for end of life operation. To establish the latter, an adjustment is made for the maximum expected NDTT shift (240oF), which the reactor vessel material will experience because of the fast neutron dose it will receive. The predicted shift will be verified by the surveillance program testing. The limits for initial operation are used to define operational limitations, and these curves are periodically updated to reflect irradiation exposure of the vessel and the results of the surveillance program. 4.4.2 Reactor Coolant Activity Limits The plant systems are designed for operation with activity in the reactor coolant systems corresponding to 1-percent fuel defects. The accident analyses presented in Chapter 14 include the calculation of doses resulting from the release of activity initially contained in the primary system. The reactor coolant system operational activity limit is defined in the Technical Specifications. 4.4.3 Maximum Pressure The reactor coolant system serves as a barrier preventing radionuclides contained in the reactor coolant from reaching the atmosphere. In the event of a fuel cladding failure, the reactor coolant system is the primary barrier against the uncontrolled release of fission products. By establishing a system pressure limit, the continued integrity of the reactor coolant system is assured. Thus, the safety limit of 2735 psig (110-percent of design pressure) has been established. This represents the maximum transient pressure allowable in the reactor coolant system under the ASME Code, Section III. Reactor coolant system pressure settings are given in Table 4.1-1. 4.4.4 System Minimum Operating Conditions Minimum operating conditions for the reactor coolant system for all phases of operation are given in the Technical Specifications. Chapter 4, Page 62 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.5 INSPECTIONS AND TESTS 4.5.1 Inspection Of Materials And Components Prior To Operation Table 4.5-1 summarizes the quality assurance program for all reactor coolant system components. In this table all of the nondestructive tests and inspections, which were required by Westinghouse specifications on reactor coolant system components and materials are specified for each component. All tests required by the applicable codes are included in this table. Westinghouse requirements, which are more stringent in some areas than those requirements specified in the applicable codes, are also included. The fabrication and quality control techniques used in the fabrication of the reactor coolant system are equivalent to those used for the reactor vessel. Westinghouse required, as part of its reactor vessel specification, that certain special tests that were not specified by the applicable codes be performed. These tests are listed below:

1. Ultrasonic testing - Westinghouse required that a 100-percent volumetric ultrasonic test of reactor vessel plate for both shear wave and longitudinal wave be performed. Section III Class A vessel plates were required by code to receive only a longitudinal wave ultrasonic test on a 9-in. x 9-in. grid. The 100-percent volumetric ultrasonic test is a severe requirement, but it ensured that the plate is of the highest quality.

2. Radiation surveillance program - In the surveillance programs, the evaluation of the radiation damage is based upon pre-irradiation and post-irradiation testing of Charpy V-notch, tensile, and wedge opening loading fracture mechanism type.

4.5.2 Reactor Vessel Surveillance Program This program is directed toward evaluation of the effects of radiation on the fracture toughness of reactor vessel steels based on the transition temperature and fracture mechanics approaches, and is in accordance with ASTM E-185, "Recommended Practice for Surveillance Tests on Structural Materials in Nuclear Reactors." The reactor vessel surveillance program uses eight specimen capsules, which are located about 3-in. from the vessel wall directly opposite the center portion of the core. The capsules can be removed when the vessel head is removed. The capsules contain reactor vessel steel specimens from the shell plates and forgings located in the core region of the reactor, associated weld metal, and heat affected zone metal. In addition, correlation monitors made from fully documented specimens of SA302 Grade B material obtained through Subcommittee II of ASTM Committee E10, Radioisotopes and Radiation Effects, are inserted in the capsules. The 8 capsules contain at least 27 tensile specimens, 256 Charpy V-notch specimens (which will include weld metal and heat-affected zone material), and 42 wedge opening loading specimens. Dosimeters including pure Ni, Al-Co (0.15-percent Co), Cd shielded Al-Co, Cd0 shielded Np-237, and Cd0 shielded U-238 are placed in the impact specimens, tensile specimens, or filler blocks drilled to contain the dosimeters. The dosimeters permit evaluation of the flux seen by the specimens and vessel wall. In addition, thermal monitors made of low melting alloys are included to monitor temperature of the specimens. The specimens are enclosed in a tight fitting stainless steel sheath to prevent corrosion. Chapter 4, Page 63 of 85 Revision 22, 2010

IP2 FSAR UPDATE Irradiation of the specimens will be higher than the irradiation of the vessel because the specimens are located in the vicinity of the core corners and are closer to the core than the vessel itself. Since these specimens will experience higher irradiation and are actual samples from the materials used in the vessel, the nil-ductility transition temperature (NDTT) measurements will be representative of the vessel at a later time in life. Data from fracture toughness samples (wedge opening loading specimens) are expected to provide additional information for use in determining allowable stresses for irradiated material. The Indian Point Unit 2 reactor vessel surveillance program was developed on the requirements provided in ASTM E-185 in effect at the time of construction. The details of the program are provided in WCAP-7323, Consolidated Edison Co., Indian Point Unit No. 2 Reactor Vessel Surveillance Program, Dated May 1969. The requirements of this program, currently form the basis for the reactor vessel surveillance program, as modified by the requirements of 10CFR50, Appendix H which state that the test procedures and reporting requirements must meet the requirements of ASTM E-185-82 to the extent practicable for the configuration of the specimens in the capsule. The following is a list of the surveillance program capsules along with the actual (past) and anticipated (future) withdrawal schedule based on the latest fluence and embrittlement calculations performed in accordance with the requirements of Regulatory Guide 1.99, Revision 2 (WCAP-15629). Capsule Location Lead Factor Withdrawal Date T 320° 3.42 End of Cycle 1 Y 220° 3.48 End of Cycle 2 Z 40° 3.53 End of Cycle 5 V 4° 1.18 End of Cycle 8 S 140° 3.5 Retired in Place** U* 176° 1.2 Spare W* 184° 1.2 End of Life*** X* 356° 1.2 Spare

• The withdrawal schedule of these capsules is interchangeable due to common materials and lead factors.
  • Capsule S may be withdrawn during the RFO 19 if modified tooling capable of removing the capsule is available. If not withdrawn, no capsule is required. If withdrawn, testing will be coordinated with industry to optimize the usefulness of the test data.
    • At the end of life as currently licensed, Capsule W (or U or X) will be withdrawn.

Results of Surveillance Capsule analyses are discussed in Section 4.2.5. 4.5.3 Primary System Quality Assurance Program Table 4.5-1 summarizes the quality assurance program with regard to inspections performed on primary system components. In addition to the inspections shown in Table 4.5-1, there are those that the equipment supplier performed to confirm the adequacy of material he received and those performed by the material manufacturer in producing the basic material. The inspections of reactor vessel, pressurizer, and steam generator are governed by ASME Code requirements. The inspection procedures and acceptance standards required on pipe materials and piping fabrication are governed by USAS B31.1 and Westinghouse requirements, and are equivalent to those performed on ASME coded vessels. Procedures for performing the Chapter 4, Page 64 of 85 Revision 22, 2010

IP2 FSAR UPDATE examinations are consistent with those established in the ASME Code Section III and are reviewed by qualified Westinghouse engineers. These procedures were developed to provide the highest assurance of quality material and fabrication. They considered not only the size of the flaws, but equally as important, how the material was fabricated, the orientation and type of possible flaws, and the areas of most severe service conditions. In addition, the surfaces most subject to damage as a result of the heat treating, rolling, forging, forming, and fabricating processes received a 100-percent surface inspection by magnetic particle or liquid penetrant testing after all these operations were completed, although flaws in plates are inherently laminations in the center. All reactor coolant plate material was subject to shear as well as longitudinal ultrasonic testing to give maximum assurance of quality. All forgings received the same inspection. In addition, 100-percent of the material volume was covered in these tests as an added assurance over the grid basis required in the code. Westinghouse quality control engineers monitored the supplier's work, witnessing key inspections not only in the supplier shop but in the shops of subvendors of the major forgings and plate material. Normal surveillance included verification of records of material, physical and chemical properties, required tests, and qualification of supplier personnel. An independent surveillance of the conformance to the fabrication and installation specifications and the quality control requirements of, among other things, the reactor coolant system components, was carried out by the United States Testing Company for Con Edison. Equipment specifications for fabrication required that suppliers submit the manufacturing procedures (welding, heat treating, etc.) to Westinghouse where they were reviewed by qualified Westinghouse engineers. This also was done on the field fabrication procedures to ensure that installation welds were of equal quality. Con Edison engineers witnessed the hydrostatic test of the reactor vessel. Cleaning of reactor coolant system piping and equipment was accomplished before and/or during erection of various equipment. Stainless steel piping was cleaned in sections as specific portions of the systems were erected. Pipe and units large enough to permit entry by personnel were cleaned by locally applying approved solvents (Stoddard solvent, acetone, and alcohol) and demineralized water, and by using a rotary disc sander or 18-8 wire brush to remove all trapped foreign particles. Section III of the ASME Boiler and Pressure Vessel Code requires that nozzles carrying significant external loads shall be attached to the shell by full penetration welds. This requirement has been carried out in the reactor coolant piping where all auxiliary pipe connections to the reactor coolant loop were made using full penetration welds. The reactor coolant system components were welded under procedures, which required the use of both preheat and postheat. Preheat requirements, not mandatory under code rules, were performed on all weldments, including P1 and P3 materials, which are the materials of construction in the reactor vessel, pressurizer, and steam generators. Preheat and postheat of weldments both served a common purpose: the production of tough, ductile metallurgical structures in the completed weldment. Preheating produces tough ductile welds by minimizing the formation of hard zones, whereas postheating achieves this by tempering any hard zones, which may have formed due to rapid cooling. Chapter 4, Page 65 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.5.4 Inservice Inspection Considerations The inservice inspection and testing program is discussed in Chapter 1. 4.5.5 Reactor Coolant System Surveillance A preoperational and inservice structural surveillance program for the reactor vessel and reactor coolant system boundary was originally established as part of the Indian Point Unit 2 initial plant conditions. This program was designed to ensure the continued integrity of the reactor coolant system boundary and included specifications, as follows:

1. Prior to initial plant operation, an ultrasonic survey was made of reactor vessel shell welds, vessel nozzles, vessel flange welds, piping system butt welds, and major welds on the pressurizer, steam generator, coolant piping and components to establish preoperational system integrity, and establish baseline data.

2. An inspection interval of 10 years was established.

3. Postoperational nondestructive inspections were provided for. The results obtained from compliance with this specification were to be evaluated after 5 years, and the conclusions of this evaluation reviewed with the NRC.

4. The structural integrity of the reactor coolant system boundary was to be maintained throughout the life of the plant at the level required by the original acceptance standards. Any evidence as a result of the inspections that defects have initiated or grown, were to be investigated, including evaluation of comparable areas of the reactor coolant system.

5. The following definitions apply to the nondestructive inspection methods.

a. UT - Volumetric examination using ultrasonic techniques.

b. RT - Volumetric examination using radiography.

c. PT - Surface examination using liquid penetrant methods.

d. V - Visual examination by direct vision or by means of remote viewing devices.

e. IV - Indirect visual examination performed during periods when the reactor coolant system is subjected to hydrostatic test pressure.

6. Detailed records of each inspection shall be maintained to allow comparison and evaluation of future inspections.

Current requirements for the primary system surveillance program are discussed in Section 5.5.6 of the facility Technical Specifications and in the Inservice Inspection and Testing Program, Chapter 1. During the first ten year inspection of the reactor vessel, an indication was discovered in a longitudinal weld in the lower shell course. While the NRC in their October 16, 1984 safety Chapter 4, Page 66 of 85 Revision 22, 2010

IP2 FSAR UPDATE evaluation concurred that the size of the indication was acceptable for plant operation, they required an augmented inspection program for the reactor vessel, which was incorporated into the Technical Specifications. By safety evaluation dated July 12, 1988, the NRC concluded that the required augmented inspection could be discontinued. In addition, inservice surveillance of the steam generator tubes that are part of the primary coolant pressure boundary is detailed in Section 5.5.7 of the Technical Specifications. This surveillance program is to ensure their continued integrity and includes inspection requirements, corrective measures, reports, and NRC approval as a condition for plant operability. This program for inservice inspection of steam generator tubes exceeds the requirements of Regulatory Guide 1.83, Revision 1, July 1975. 4.5.6 Reactor Coolant Vent System Testing The testing of the remote reactor head vent and power operated relief valves system valves is performed in accordance with ASME Code Section XI requirements for Category B valves. Chapter 4, Page 67 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 1 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5

1. Steam generator 1.1 Tube sheet 1.1.1 Forging Yes Yes 1.1.2 Cladding Yes6 Yes7 1.2 Channel head 1.2.1 Casting Yes Yes 1.2.2 Cladding Yes 1.3 Secondary shell and head 1.3.1 plates Yes 1.4 Tubes Yes Yes 1.5 Nozzles (forgings) Yes Yes 1.6 Weldments 1.6.1 Shell, longitudinal Yes Yes 1.6.2 Shell, circumferential Yes Yes 1.6.3 Cladding (channel head-tube sheet joint cladding restoration) Yes 1.6.4 Steam and feedwater nozzles to shell Yes Yes 1.6.5 Support brackets Yes 1.6.6 Tube to tube sheet Yes Chapter 4, Page 68 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 2 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5 1.6.7 Instrument connections (primary and secondary) Yes 1.6.8 Temporary attachments after removal Yes 1.6.9 After hydrostatic test (all welds and complete channel head) Yes 1.6.10 Nozzle safe ends Yes Yes (if forgings) 1.6.11 Nozzle safe ends Yes (if weld deposit)

2. Pressurizer 2.1 Heads 2.1.1 Casting Yes Yes 2.2.2 Cladding Yes 2.2 Shell 2.2.1 Plates Yes Yes 2.2.2 Cladding Yes 2.3 Heaters 2.3.1 Tubing8 Yes Yes 2.3.2 Centering of element Yes Chapter 4, Page 69 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 3 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5 2.4 Nozzle Yes Yes 2.5 Weldments 2.5.1 Shell, longitudinal Yes Yes 2.5.2 Shell, circumferential Yes Yes 2.5.3 Cladding Yes 2.5.4 Nozzle safe end Yes Yes (if forging) 2.5.5 Nozzle safe end Yes (if weld deposit) 2.5.6 Instrument connections Yes 2.5.7 Support skirt Yes 2.5.8 Temporary attachments Yes after removal 2.5.9 All welds and cast heads Yes after hydrostatic test 2.6 Final Assembly 2.6.1 All accessible weld surfaces after hydrostatic test Yes

3. Primary Coolant Piping 3.1 Fittings (castings) Yes Yes 3.2 Fittings (forgings) Yes Yes Chapter 4, Page 70 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 4 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5 3.3 Pipe9 Yes Yes 3.4 Weldments 3.4.1 Circumferential Yes Yes 3.4.2 Nozzle to run pipe Yes Yes (no RT for nozzles less than 3-in.) 3.4.3 Instrument connections Yes

4. Pumps 4.1 Casting Yes Yes 4.2 Forgings Yes Yes 4.2.1 Main shaft Yes Yes 4.2.2 Main studs Yes Yes 4.2.3 Flywheel (rolled plate) Yes 4.3 Weldments 4.3.1 Circumferential Yes Yes 4.3.2 Instrument connections Yes

5. Reactor Vessel 5.1 Forgings 5.1.1 Flanges Yes Yes 5.1.2 Studs Yes Yes Chapter 4, Page 71 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 5 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5 5.1.3 Head adapters Yes Yes 5.1.4 Head adapter tube Yes Yes 5.1.5 Instrumentation tube Yes Yes 5.1.6 Main nozzles Yes Yes 5.1.7 Nozzle safe ends Yes Yes (if forging is employed) 5.2 Plates Yes Yes 5.3 Weldments 5.3.1 Main seam Yes Yes 5.3.2 CRD head adapter Yes connection 5.3.3 Instrumentation tube Yes connection 5.3.4 Main nozzles Yes Yes 5.3.5 Cladding Yes10 Yes 5.3.6 Nozzle-safe ends Yes Yes (if forging) 5.3.7 Nozzle safe ends Yes Yes (if weld deposits) 5.3.8 Head adaptor forging Yes Yes to head adaptor tube 5.3.9 All welds after hydrotest Yes Chapter 4, Page 72 of 85 Revision 22, 2010

IP2 FSAR UPDATE TABLE 4.5-1 (Sheet 6 of 6) Reactor Coolant System Quality Assurance Program Component RT1 UT2 PT3 MT4 ET5

6. Valves 6.1 Castings Yes Yes 6.2 Forgings Yes Yes (No UT for valves two inches and smaller)

Notes:

1. RT - Radiographic.

2. UT - Ultrasonic.

3. PT - Dye Penetrant.

4. MT - Magnetic Particle.

5. ET - Eddy Current.

6. Flat Surfaces Only.

7. Weld Deposit Areas Only.

8. Or a UT and ET.

9. Except pressurizer surge line - UT only.

10. UT of Clad Bond-to-Base Metal.

Chapter 4, Page 73 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4.6 METAL IMPACT MONITORING SYSTEM 4.6.1 General The metal impact monitoring system is designed to enable early detection of any debris, detached internal structural items, and hardware present in the reactor coolant system. A metal impact monitoring system for Indian Point Unit 2 was installed during the 1976 refueling outage and was operational when the plant returned to service in September 1976. At that time, component "signature acquisition" of the nuclear steam supply system components (baseline data) was obtained at selected plant operating conditions for future reference. The metal impact monitoring system was modified during the 1982 refueling outage. 4.6.2 Description This system involves the use of a metal impact monitoring system capable of detecting changes in reactor coolant system vibrations and converting that input into an electronic signal thereby providing an indication to operating personnel that an undesireable level of foreign material may be present in the reactor coolant. While the installed system has no control capability, it is nevertheless quite valuable as an advisory system. Metal impact monitoring is accomplished by the installation of specially developed transducers (accelerometers) mounted on the exterior of the reactor coolant system and steam generators. When the interior of the reactor coolant system is struck by bouncing debris, the structure is shock excited producing local wall accelerations that are detected by the transducers, amplified, conditioned, and fed to the metal impact monitoring system. The metal impact monitoring system further conditions the signals for recording and display in the control room. The transducers are located on the following equipment:

1. Reactor vessel head.

2. Incore instrumentation penetration (below reactor vessel).

3. Steam generators.

Chapter 4, Page 74 of 85 Revision 22, 2010

IP2 FSAR UPDATE APPENDIX 4A DETERMINATION OF REACTOR PRESSURE VESSEL NIL-DUCTILITY TRANSITION TEMPERATURE (NDTT) 4A.1 MEASUREMENT OF INTEGRATED FAST NEUTRON (E > 1.0 MEV) FLUX AT THE IRRADIATION SAMPLES The energy dependent neutron fluxes at the irradiation samples are obtained from the DOT(1), a two-dimensional discrete ordinates transport theory code. Dosimeters in the surveillance program include CdO shielded U-238, Np-237, Co-Al, Cu, Ni, Cd shielding Co-Al, and Fe from specimens, which will be contained in the capsule assemblies. The specific activities of the dosimeters are to be determined by the multichannel analyzer and NAI scintillation detector. The equipment calibration shall be accomplished with 54Mn and 60Co radioactivity standards obtained from the U.S National Bureau of Standards or the equivalent. All activities will be corrected to the time-of-removal (TOR) at reactor shutdown. Infinite dilute saturated activities (ASAT) will be calculated for each of the dosimeters because ASAT is directly related to the product of the energy dependent microscopic activation cross-section and the neutron flux density. The relationship between ATOR and ASAT is given by: m=n A TOR = Pm (1 - e - T m ) (e - t m ) A SAT m =1 Where: = decay constant for the activation product, 1/day tm = decay time after operating period m, days Tm = operating days Pm = average fraction of full power during operating period Pm = average fraction of full power during operating period The primary result desired from the dosimeter analysis in the total neutron fluence (E > 1 MeV) that the surveillance specimens and pressure vessel have received. The average flux density at full power is given by: o/ = ASAT/No Where: o/ = energy dependent neutron flux density, n/cm2-sec

                  = spectrum averaged activation cross-section, cm2 No = number of target atoms per mg The total neutron flux fluence is then equal to the product of the averaged neutron flux and the equivalent reactor operating time at full power.

4A.2 CALCULATION OF INTEGRATED FAST NEUTRON (E > 1.0 MEV) FLUX AT THE IRRADIATION SAMPLES In the analysis of the neutron environment within a pressurized water reactor geometry, predictions of the spatial neutron flux magnitude, and energy spectra are made with the DOT (two-dimensional discrete ordinates transport theory code). First, the radial and azimuthal distributions are obtained from an R, computation normalized to the reactor core power Chapter 4, Page 75 of 85 Revision 22, 2010

IP2 FSAR UPDATE density representative of the axial midplane. A second calculation in R, Z geometry is used to provide relative axial variations of neutron flux in the pertinent regions of the pressure vessel. A three-dimensional description of the neutron environment is then constructed by assuming separability and using the relation: (R, , Z, E) = (R, , E) x F(Z, E) Where o/ (R, , E) represents the absolute neutron flux magnitude at the core midplane as determined from the R, computation and F(Z, E) is the relative axial distribution obtained from the R, Z analysis and normalized to unity at the core midplane. From a neutronic standpoint, the inclusion of the surveillance capsule structures in the R, analytical model is significant. Neutron dosimetry from these capsules provides a means for evaluating the analytical model by direct comparison with measurement. Since the presence of the capsules has a marked impact on both the neutron flux magnitude and energy spectrum, a meaningful comparison of measurement and calculation can be made only if these perturbation effects are properly accounted for in the analysis. Two distinct sets of transport calculations are carried out. The first, a single computation in the conventional forward mode, is used primarily to obtain relative neutron energy distributions throughout the reactor geometry as well as to establish relative radial distributions of exposure parameters ( (E > 1.0 MeV), (E > 0.1 MeV), and dpa) through the vessel wall. The neutron spectral information is required for the interpretation of neutron dosimetry withdrawn from surveillance capsules as well as for the determination of exposure parameter ratios: i.e., dpa/(E > 1.0 MeV), within the pressure vessel geometry. The relative radial gradient information is required to permit the projection of measured exposure parameters to locations interior to the pressure vessel wall; i.e., the 1/4T, 1/2T, and 3/4T locations. The second set of calculations consists of a series of adjoint analyses relating the fast neutron flux (E > 1.0 MeV) at surveillance capsule positions, and several azimuthal locations on the pressure vessel inner radius to neutron source distributions within the reactor core. The importance functions generated from these adjoint analyses provide the basis for all absolute exposure projections and comparison with measurement. These importance functions, when combined with cycle specific neutron source distributions, yield absolute predictions of neutron exposure at the locations of interest for each of the operating fuel cycles; and establish the means to perform similar predictions and dosimetry evaluations for all subsequent fuel cycles. It is important to note that the cycle specific neutron source distributions utilized in these analyses include not only spatial variations of fission rates within the reactor core; but, also account for the effects of varying neutron yield per fission and fission spectrum introduced by the build-in of plutonium as the burnup of individual fuel assemblies increased. The absolute cycle specific data from the adjoint evaluations together with relative neutron energy spectra and radial distribution information from the forward calculation provide the means to:

1. Evaluate neutron dosimetry obtained from the surveillance capsule program.

2. Extrapolate dosimetry results to key locations at the inner radius and through the thickness of the pressure vessel wall.

3. Enable a direct comparison of analytical prediction with measurement.

Chapter 4, Page 76 of 85 Revision 22, 2010

IP2 FSAR UPDATE

4. Establish a mechanism for projection of pressure vessel exposure as the design of each new fuel cycle evolves.

The forward transport calculation is carried out in R, geometry using the DOT two-dimensional discrete ordinates code1 and the SAILOR cross-section library2. The SAILOR library is a 47 group ENDFB-IV based data set produced specifically for light water reactor applications. In these analyses anisotropic scattering was treated with a P3 expansion of the cross-sections and the angular discretization was modeled with an S8 order of angular quadrature. The reference forward calculation is normalized to a core power. All adjoint analyses are also carried out using an S8 order of angular quadrature and the P3 cross-section approximation from the SAILOR library. Adjoint source locations are chosen at several azimuthal locations at the pressure vessel inner radius and at the geometric center of surveillance capsules positioned at 4o and 40o relative to the core cardinal axes. Again these calculations are run in R, geometry to provide neutron source distribution importance functions for the exposure parameter of interest; in this case, (E > 1.0 MeV). Having the importance functions and appropriate core source distributions, the response of interest could be calculated as: R(r, ) = r E I(r, ,E)S(r, ,E)r dr d dE where: R(r, ) = (E > 1.0 MeV) at radius r and azimuthal angle I (r, , E) = Adjoint importance function at radius r, azimuthal angle , and neutron source energy E. S (r, , E) = Neutron source strength at core location r, . and energy E. Forward transport as well as the adjoint analyses for Indian Point Unit 2 were carried out and summarized in Reference 3. In the R, analysis, the discretization of the angular flux is represented by a symmetric S8 quadrature. However, in the R, Z case the use of this relatively low order quadrature set can often prove to be inadequate. At large depths within the pressure vessel, the axial distribution of neutron flux is dominated by neutron streaming in the annulus between the pressure vessel wall and the primary biological shield. To account for this effect a high resolution angular quadrature is required. Therefore, in this analysis a 124 angle asymmetric quadrature is employed. For regions of the reactor, which are above the core midplane, this quadrature is constructed with 109 angles biased in the upward directions, i.e., the direction of prime interest, and 15 angles biased downward. For analysis below the core midplane, the quadrature is reversed with 109 angles biased in the downward direction. Complete descriptions of both the symmetric S8 and the asymmetric 124 angle quadratures are given in Reference 1. The calculated fast neutron flux distributions may be used in conjunction with damage trend curves to predict the degree of embrittlement of the reactor vessel steel over its service life. The accuracy of these neutron flux profiles depends on the analyst's ability to define an appropriate core power distribution, the adequacy of the cross-sections used in the transport analysis, and the applicability of the geometric modeling of the reactor. Taken as a whole, these factors combine to yield an overall uncertainty of 20-percent in the prediction of neutron flux and fluence within the pressure vessel wall. Chapter 4, Page 77 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4A.3 MEASUREMENT OF THE INITIAL NIL-DUCTILITY TRANSITION TEMPERATURE OF THE REACTOR PRESSURE VESSEL BASE PLATE AND FORGINGS MATERIAL The unirradiated or initial NDTT of pressure vessel reactor materials was measured by two methods. These methods were the drop weight test per ASTM E208 and the Charpy V-notch impact test (Type A) per ASTM E23. The NDTT is defined in ASTM E208 as the temperature at which a drop weight test specimen is broken in a series of tests in which duplicate no-break performance occurs at a temperature of 10oF higher. The NDTT temperature, as determined by drop weight tests is the RTNDT if, at 60oF above the NDTT, at least 50-ft-lbs of energy and 35 mils lateral expansion are obtained in Charpy V tests on specimens oriented in the weak direction (traverse to the direction of maximum working). The NDTT has been correlated with Charpy V-notch impact tests results. For SA 302B and A508 Class 2 steels the Charpy V-notch "fix" temperature, which corresponds to NDTT is the temperature at 30-ft-lbs in accordance with Section III Table N-421 of the ASME Code for Nuclear Vessels. The curve of the temperature versus energy observed in breaking the specimen was plotted. To obtain this curve 15 tests were performed, which include three tests at five different temperatures. The intersection of the energy versus temperature curve with the 30-ft-lbs ordinate is designated as NDTT. As part of the Westinghouse surveillance program referred to above, Charpy V impact tests, tensile tests, and fracture mechanics specimens are taken from the plate of forging material. To assess any possible uncertainties in the consideration of NDTT shift for welds, heat affected zone and base metal, test specimens of these three "material types" have been also included in the reactor vessel surveillance program. Encapsulated specimens are located on the outside diameter surface of the thermal shield where the fast neutron flux density is about three times that at the adjacent vessel wall surface. The capsules also contain several dosimeter materials for experimentally determining the average neutron flux density at each capsule location during the exposure period. REFERENCES FOR APPENDIX 4A

1. R. G. Soltesz, et al, "Nuclear Rocket Shielding Methods, Modification, Updating and Input Data Preparation - Volume 5, Two-Dimensional Discrete Ordinates Transport Techinque," WANL-PR-(LL)-304, August, 1970.

2. "ORNL RSIC Data Library Collection DLC-76, SAILOR Coupled Self Shielded, 47 Neutron, 20 Gamma-Ray, P3, Cross Section Library for Light Water Reactors".

3. S.L. Anderson "Plant Specific Fast Neutron Exposure Evaluation of the Indian Point Unit 2 Reactor Pressure Vessel and Surveillance Capsules Fuel Cycles 1 through 9". Westinghouse report PSE-REA-88/127, July 1988.

Chapter 4, Page 78 of 85 Revision 22, 2010

IP2 FSAR UPDATE APPENDIX 4B SUPPORT STRUCTURES FOR REACTOR COOLANT SYSTEM COMPONENTS The reactor coolant system components and their supports are designed as seismic Class I components as discussed in Section 1.11. In 2003, the reactor coolant loop and its component supports were re-analyzed due to a power uprate. This latest analysis does not consider the coincident combination of blowdown and seismic loads. 4B.1 REACTOR VESSEL The reactor vessel support structure consists of a circular box section ring girder fabricated of carbon steel plates. The bottom flange of the girder is in continuous contact (except for openings for neutron detectors) with a non-yielding concrete foundation. The reactor vessel has four supports located at alternate nozzles and cooled by the component cooling system. Each support bears on a support shoe, which is fastened to the support structure. The support shoe is a structural member that transmits the support loads to the supporting structure. The support shoe is designed to restrain vertical, lateral, and rotational movement of the reactor vessel, but allows for thermal growth by permitting radial sliding at each support on bearing plates. 4B.2 STEAM GENERATORS The steam generators are supported within a caged structural system consisting of four connected columns welded together, fabricated of carbon steel members, with provisions for limited movement of the structure in a horizontal direction with a system of "Lubrite" plates, hydraulic snubbers, guides, and stops to accommodate piping expansion. The "Lubrite" plates, hydraulic snubbers, guides, and stops were originally designed as a rigid support to resist the action of seismic and pipe break loads. In 2000, the number of hydraulic snubbers supporting the steam generator frame in the direction of the hot leg, has been reduced from the original six down to two per steam generator. The two remaining snubbers are located at the upper support point of the frame at Elevation 92'-0". The analysis of the reactor coolant loop and of the steam generator support structure accounts for the replacement steam generator and for the reduced number of hydraulic snubbers. 4B.3 REACTOR COOLANT PUMP Each reactor coolant pump is supported on a three-legged structural system consisting of three connected columns fabricated of carbon steel members, structural sections, and pipe. Provisions for limited movement of the structure in any horizontal direction to accommodate piping expansion is accomplished with a sliding "Lubrite" base plate arrangement, and a system of tie rods and anchor bolts, which restrain the structure from movement beyond the calculated limits. Chapter 4, Page 79 of 85 Revision 22, 2010

IP2 FSAR UPDATE 4B.4 PRESSURIZER The pressurizer is supported on a free-standing structural system consisting of six connected columns fabricated of carbon steel members, all welded together and secured at the base by anchor bolts. 4B.5 PIPING The reactor coolant piping layout is designed on the basis of providing "floating" supports for the steam generator and reactor coolant pump in order to absorb the thermal expansion from the fixed or anchored reactor vessel. A comprehensive thermal analysis has been performed to ensure that stresses induced by linear thermal expansion are within code limits. 4B.6 APPLICABILITY OF UNIT 3 PIPE BREAK ANALYSES TO UNIT 2 A report (Reference 1) entitled, "Analysis of Reactor Coolant System for Postulated Loss-of-Coolant Accident: Indian Point Unit 3 Nuclear Power Plant," has been submitted to the NRC. This report postulates pipe breaks at the locations in the primary loop, which induce the most severe asymmetric loads on the reactor vessel. The analyses performed included the effects of the addition of pipe motion limiters and demonstrate the adequacy of the entire system. Reference 4 of Section 4.2 addresses the applicability of this report to Unit 2. Because of the similarity of the plants the nature of the system response, and the installation in Unit 2 of the modifications discussed in that report, the conclusions stated for Unit 3 in that report are found to be applicable to Unit 2. 4B.7 LEAK BEFORE BREAK In 1989, the NRC approved elimination of the necessity for considering and protecting against dynamic effects of postulated primary loop pipe ruptures from the design basis of Indian Point Unit 2 as discussed in Section 4.1.2.4. "Leak before break" technology was applied as permitted by revised General Design Criterion 4 of 10CFR50, Appendix A. References 2, 3, 4 and 5 contain further information. REFERENCES FOR APPENDIX 4B

1. "Analysis of Reactor Coolant System for Postulated Loss-Of-Coolant Accident:

Indian Point Unit 3 Nuclear Power Plant." WCAP-9117 (Proprietary) and WCAP-9130 (Non-Proprietary), Westinghouse Electric Corporation.

2. May 23, 1988 letter, Bram to Document Control Desk, subject: Leak-Before-Break (LBB).

3. November 18, 1988 letter, Bram to Document Control Desk, subject: Leak-Before-Break (LBB) Submittal (TAC 68318).

4. January 12, 1989 letter, Bram to Document Control Desk, subject:

Leak-Before-Break (LBB) Submittal (TAC 68318). Chapter 4, Page 80 of 85 Revision 22, 2010

IP2 FSAR UPDATE

5. Letter from Donald Brinkman, NRC, to Stephen B. Bram, Con Edison,

Subject:

Safety Evaluation Report on Elimination of Dynamic Effect of Postulated Primary Loop Pipe Ruptures from Design Basis for Indian Point Unit 2 (TAC No. 68318), dated February 23, 1989. APPENDIX 4C SENSITIZED STAINLESS STEEL 4C.1 INTRODUCTION Westinghouse has evaluated the use of sensitized stainless steel for reactor components in pressurized water reactors. The results of this evaluation are summarized in WCAP 7477-L (Reference 1), which cover the nature of sensitization conditions leading to stress corrosion and associated problems with both sensitized and non-sensitized stainless steel. The results of extensive testing and service experience that justify the use of stainless steel in the sensitized condition for components in Westinghouse systems is presented in the report. Sensitized stainless steel is subject to stress corrosion and must not be exposed to certain environments that will cause cracking. Chlorides and fluorides are the most important contaminants, although oxygen, low pH, elevated temperature, and high stress generally must also be present to cause cracking. When subjected to environments that cause cracking, the cracks are usually intergranular in sensitized stainless steel. The stainless steel safe-ends on the reactor vessel, pressurizer, and steam generator nozzles may become somewhat sensitized during stress relief of the vessel. The post weld heat treatment (PWHT) temperatures and minimum time are consistent with ASME Section III requirements. The degree of sensitization of the safe-ends varies from plant to plant, depending on the materials used and the detailed processing performed by the various vendors. For Indian Point Unit 2, the specific design and construction practices are discussed in the following sections. The outer diameter and inner diameter safe-ends of the reactor vessel were overlaid with type 308L and Inconel weld metal to eliminate any question of intergranular attack in areas where there is limited accessibility for inservice inspection and plant maintenance. There is complete accessibility to the remaining reactor coolant system components. The pre-operational inspection of the reactor coolant system components provides assurance that there is no stress corrosion cracking of sensitized stainless steel. 4C.2 REACTOR COOLANT SYSTEM NOZZLE SAFE-ENDS 4C.2.1 Reactor Vessel Primary Nozzle Safe-Ends

1. Method of Fabrication (See Figure 4C-1)

a. Wrought stainless steel - Type 316 Forging welded to SA-336 nozzle with Inconel weld metal. Attached prior to final post weld heat treatment.

b. Forging was overlaid on ID and OD with type 308L stainless and Inconel weld metal. This was performed in the field after the primary coolant piping was attached to the nozzles.

Chapter 4, Page 81 of 85 Revision 22, 2010

IP2 FSAR UPDATE

2. Inspection

a. Forging safe-ends were examined by ultrasonic testing and penetrant testing at Combustion Engineering using Section III acceptance standards.

b. Weld overlay of the ID and OD surfaces was examined by ultrasonic testing and penetrant testing. The acceptance standards are shown below:

(1) Ultrasonic Acceptance Standards Each discontinuity that produced a response equal to or exceeding the calibration reference line and was 0.5-in. or greater in length was considered rejectable and removed. Discontinuities that produced a response equal to or greater than the calibration reference line and exceed 0.25-in., but were less than 0.5-in. in length were considered acceptable if separated by a minimum distance of 2-in. from similar discontinuities. Each discontinuity that produced a response between 50 and 100-percent of the calibration reference line and exceeded one inch but was not more than 1.5-in. in length, were acceptable if separated by a minimum distance of 2-in. from similar indications. (2) Penetrant Inspection Acceptance Standards (a) Examination of welds by liquid penetrant methods were made over an area including the welds and base metal extending for at least 0.5-in. on each side of weld. (b) Surfaces examined by fluid penetrant methods were free of laps, fissures, cracks, other linear indications. (c) Weld area and adjacent wrought type base metal(s) - In any 6-in. length of weld and adjacent base metal examined, there were no indications greater than 0.62-in. in maximum dimension, nor were there more than six indications with sum of maximum dimensions specified herein. Any 6-in. length of weld was interpreted to denote the 6-in. length selected in the least favorable location with respect to the discontinuities disclosed by the inspection test. All surfaces examined were free of linearly disposed indications of four or more indications in a line and each separated by 1/16-in. or less, edge to edge. (d) Weld area and adjacent cast type base metal(s) - In any 6-in. length of weld examined, there were no indications greater than those defined in 3, above. The adjacent cast base metal was free of random indications in excess of those shown in the following table for a distance of not less than 0.5-in. from toe(s) of weld: Chapter 4, Page 82 of 85 Revision 22, 2010

IP2 FSAR UPDATE Size of Indications, In.Number per Square In.

                                > 1/8                                     None
                                >1/16 < 2/8                               2
                                < 1/16                                    10 (e) All surfaces examined were free of linearly disposed indications of four or more indications in a line and each separated by 1/16-in. or less, edge to edge. Rounded indications were those which were circular or elliptical with the length less than twice the width.

4C.2.2 STEAM GENERATOR PRIMARY NOZZLE SAFE-ENDS (See Figure 4C-2)

1. Method of Fabrication Weld metal buttering applied to carbon steel (A-216 Casting) nozzles prior to final post weld heat treatment. Stainless weld metal for the first layer was type 309L, and for the balance was type 308L.

2. Inspection Buttered safe-ends were examined by penetrant testing and radiography testing using ASME Boiler and Pressure Vessel Code Section III acceptance standards.

4C.2.3 PRESSURIZER (See Figure 4C-3)

1. Method of Fabrication Wrought stainless steel pipe or forgings welded to carbon steel (A-216 Casting) nozzles with type 309 weld metal before post weld heat treatment. The surge nozzle safe-end was fabricated from SA-312 pipe, type 316, and the spray, relief, and safety nozzle safe-ends from SA-182 forgings, type 316.

2. Inspection Wrought material was examined by ultrasonic testing and penetrant testing using Section III acceptance standards.

4C.3 REACTOR COOLANT SYSTEM CONSTRUCTION All primary piping and fittings were given a solution annealing treatment consisting of heating to 1900 - 1950oF, holding 1 hr/in. of thickness, and water quenching. This ensured that the material would not be sensitized. Main coolant pipe welds are of type 308 or type 316 stainless steels. Welding was performed by the manual metal arc process after the root pass was completed using an insert followed by three layers using the manual gas shielded tungsten arc process. The maximum energy input possible with the manual metal arc process is on the order of 20,000 joules per linear inch of weld. With the large heat sink available in this thick-walled pipe (2.375 to 3.00-in.) and the interpass temperature control of 350oF maximum, there will be no sensitization of the solution-treated pipe during welding. Chapter 4, Page 83 of 85 Revision 22, 2010

IP2 FSAR UPDATE Venting provisions have been made at high points throughout the reactor coolant system to relieve entrapped air when the system is filled and pressurized. Principally, vents are installed on the reactor coolant pumps with additional vents available on the control rod drive mechanisms, on instruments, and on a number of connecting pipes. For normal venting of the reactor coolant system, only the principal venting points are used. The amount of oxygen, which could be trapped in the remaining small volumes becomes negligible as the system is pressurized and the oxygen is scavenged by the hydrazine specifically added for this purpose prior to operation. During operation, the oxygen levels are kept low consistent with water chemistry requirements as described in the Technical Specifications. In addition to the high point vents, a connection is installed downstream of the Power-operated Relief Valves to permit pulling the air out of the system under vacuum during system refilling. 4C.4 REACTOR COOLANT SYSTEM OPERATIONAL STRESSES To avoid unusual stresses in areas where nozzle safe-ends are joined to the piping, precautions were taken to eliminate unnecessary stresses due to erection of the various components of the reactor coolant system. The primary coolant system piping closure pieces were two pipe fitting subassemblies located between the steam generator and the primary coolant pump. The 40-degree elbow of the loop piping was first installed on the steam generator outlet nozzles. Then the gap to be closed by the closure pieces was physically measured between the 40-degree elbow outlet and the inlet nozzle of the pump. These measured dimensions for each individual loop were compensated and adjusted for the expected field weld shrinkage. The resulting net true dimensions were then transmitted to the pipe shop fabricator who prepared the final closure pipe subassemblies for each primary coolant loop. Upon welding these specially dimensioned pipe subassemblies in place, the primary coolant system closure was accomplished for each loop in a condition, which is free from cold spring. As a precaution that the behavior of the reactor coolant system during operating conditions would be as predicted, measurements were made at incremental temperature increases during the hot functional test. The measurements were made to check the movement of the components at temperature and pressure to ensure interferences were not present. Data taken during the test were compared with the flexibility analysis predictions and evaluated. 4C.5 INSERVICE INSPECTION CAPABILITY As a final check on the adequacy of the precautions taken to avoid any reactor coolant system failure as a result of severely sensitized stainless steel, a postoperational inspection plan was developed for the nozzle safe-ends within the reactor coolant system boundary. The pressurizer and steam generator stainless steel safe-ends that were subjected to the furnace atmosphere during final stress relief are accessible for visual, surface, and volumetric inspection upon removal of the insulation at each safe-end. The reactor vessel safe-ends, which were subjected to the furnace atmosphere, are accessible for limited inspection by removal of the special access plugs provided in the primary concrete just above each nozzle. Upon removal of these plugs and the insulation on the safe-end, approximately 120-degrees of the top segment of the safe-ends are accessible for direct visual, surface, and remote volumetric inspection. As specially designed devices for remote ultrasonic inspection and applicable procedures become available, and when metallurgical considerations indicate that this type of inspection is appropriate and necessary, such inspections will be accomplished utilizing the internal access Chapter 4, Page 84 of 85 Revision 22, 2010

IP2 FSAR UPDATE to the reactor vessel safe-ends. Requirements for inspection of the reactor coolant system are detailed in the facility Technical Specifications. REFERENCES FOR APPENDIX 4C

1. WCAP-7477L (Proprietary), Westinghouse Electric Corporation.

APPENDIX 4C FIGURES Figure No. Title Figure 4c-1 Primary Nozzle Combustion Engineering Reactor Vessel Figure 4c-2 Primary Nozzle Tampa Steam Generators Figure 4c-3 Spray or Surge Nozzle Tampa Pressurizer Chapter 4, Page 85 of 85 Revision 22, 2010

10e 55 THE RMAl 5lEE VE"** PIPE OR FORGING TYPE {309 55" I N5TALLEO AFTER 55 WELD TYPE 316 55 OR INCONEl FINAL P.WH:l OVERLAY I CR FO...l..ID-IN LII\ER **

             -~-----t                  - -----.-..----+-~---
                                                                                  //J,.~cs      NOZZL E

. Butt weld. ~ Type 309 0.08 ( max; TiG root. balance TyPe 309 0.09 C max; SHAW Attachment weld of thermal sleeve and rolled-in liner - Type 308 l 0.04 ( max. TIG (made afrer final PWHT)

• .. Rolled-in liner welded top and l!ottorn for spray. safety. *and, ,

relief nozzles - Type 109 followed by Type 308 l weld o~erlay for surge nozzle INDIAN POINT UNIT No. 2 UFSAR FIGURE 4C-3 SPRAY OR SURGE NOZZLE TAMPA PRESSURIZER MIC. No. 1999MC3745 REV. No. 17A

---

l

                      'I'
             ~
               ~.\:                :,1 ZERP=:LZ=:FTAr,j //',jl INDIAN POINT UNIT No. 2 UFSAR FIGURE 4.2-1 REACTOR COOLANT SYSTEM FLOW DIAGRAM MIC. No. 1999MC3430             REV. No. 178 VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

PRESSURIZER REACTOR VESSEL VIEW LOOKING WEST INDIAN POINT UNIT No. 2 UFSAR FIGURE 4.2-2 REACTOR COOLANT SYSTEM SCHEMATIC FLOW DIAGRAM MIC. No. 1999MC3733 REV. No. 17A

                  ~,.:..:.:

INDIAN POINT UNIT No. 2 UFSAR FIGURE 4.2-3 REACTOR VESSEL MIC. No. 1999MC3734 I REV. No. 17A

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TYPE 316 SS TYPE 308L INCONEl CSNOZZLE FORGING OVERLAY OVERLAY INCONEL TYPE 308L INCONEL 10 OVERLAY OVERLAY INDIAN POINT UNIT No. 2 UFSAR FIGURE 4C-1 PRIMARY NOZZLE COMBUSTION ENGINEERING REACTOR VESSEL MIC. No. 1999MC3743 REV. No. 17A

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IP2 FSAR UPDATE CHAPTER 5 CONTAINMENT SYSTEMS 5.1 CONTAINMENT STRUCTURES 5.1.1 Design Basis The reactor containment completely encloses the entire reactor and reactor coolant system and ensures that essentially no leakage of radioactive materials to the environment would result even if gross failure of the reactor coolant system were to occur. The liner and penetrations are designed to prevent any leakage through the containment. The structure provides biological shielding for both normal and accident situations. The reactor containment is designed to safely withstand several conditions of loading and their credible combinations. The major loading conditions are:

1. Occurrence of a gross failure of the reactor coolant system, which creates a high-pressure and temperature condition within the containment.

2. Coincident failure of the reactor coolant system with an earthquake or wind.

5.1.1.1 Principal Design Criteria 5.1.1.1.1 Quality Standards Criterion: Those systems and components of reactor facilities, which are essential to the prevention, or the mitigation of the consequences, of nuclear accidents, which could cause undue risk to the health and safety of the public shall be identified and then designed, fabricated, and erected to quality standards that reflect the importance of the safety function to be performed. Where generally recognized codes and standards pertaining to design, materials, fabrication, and inspection are used, they shall be identified. Where adherence to such codes or standards does not suffice to assure a quality product in keeping with the safety function, they shall be supplemented or modified as necessary. Quality assurance programs, test procedures, and inspection acceptance criteria to be used shall be identified. An indication of the applicability of codes, standards, quality assurance programs, test procedures and inspection acceptance criteria used is required. Where such items are not covered by applicable codes and standards, a showing of adequacy is required. (GDC 1) The containment system structure is of primary importance with respect to its safety function in protecting the health and safety of the public. Quality standards of material selection, design, fabrication, and inspection governing the above features conforms to the applicable provisions of recognized codes and good nuclear practice. The concrete structure of the reactor containment conforms to the applicable portions of ACI18-63. Further elaboration on quality standards of the reactor containment is given is Section 5.1.1.5. Chapter 5, Page 1 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.1.1.2 Performance Standards Criterion: Those systems and components of reactor facilities, which are essential to the prevention or to the mitigation of the consequences of nuclear accidents, which cause undue risk to the health and safety of the public shall be designed, fabricated, and erected to performance standards that enable such systems and components to withstand, without undue risk to the health and safety of the public, the forces that might reasonably be imposed by the occurrence of an extraordinary natural phenomenon such as earthquake, tornado, flooding condition, high wind or heavy ice. The design bases so established shall reflect: (a) appropriate consideration of the most severe of these natural phenomena that have been officially recorded for the site and the surrounding area and (b) an appropriate margin for withstanding forces greater than those recorded to reflect uncertainties about the historical data and their suitability as a basis for design. (GDC 2) All components and supporting structures of the reactor containment are designed so that there is no loss of function of such equipment in the event of maximum potential ground acceleration acting in the horizontal and vertical directions simultaneously. The dynamic response of the structure to ground acceleration, based on the site characteristics and on the structural damping, is included in the design analysis. The reactor containment is defined as a Class I structure for purposes of seismic design (Section 1.11). Its structural members have sufficient capacity to accept, without exceeding specified stress limits, a combination of normal operating loads, functional loads due to a loss-of-coolant accident, and the loadings imposed by the maximum potential earthquake. 5.1.1.1.3 Fire Protection Criterion: A reactor facility shall be designed to ensure that the probability of events such as fires and explosions and the potential consequences of such events will not result in undue risk to the health and safety of the public. Noncombustible and fire resistant materials shall be used throughout the facility wherever necessary to preclude such risk, particularly in areas containing critical portions of the facility such as containment, control room, and components of engineered safety features.(GDC 3) Fire protection in all areas of the nuclear electric plant is provided by structure and component design that optimizes the containment of combustible materials and maintains exposed combustible material below the ignition temperature. The station is designed on the basis of limiting the use of combustible materials in construction by using fire-resistant materials to the greatest extent practical. Containment liner thermal insulation does not support combustion. The bearing oil systems for the reactor coolant pump motors are self-contained. All oil-containing equipment associated with the reactor coolant pump motors is also completely enclosed by an oil-collecting system, which in the event of an oil leak, will contain and channel away the oil to remote storage containers. 5.1.1.1.4 Records Requirement Criterion: The reactor licensee shall be responsible for assuring the maintenance throughout the life of the reactor of records of the design, fabrication, and Chapter 5, Page 2 of 89 Revision 22, 2010

IP2 FSAR UPDATE construction of major components of the plant essential to avoid undue risk to the health and safety of the public. (GDC 5) Records of the design, fabrication, construction, and testing of the reactor containment are maintained throughout the life of the reactor. 5.1.1.1.5 Reactor Containment Criterion: The containment structure shall be designed (a) to sustain, without undue risk to the health and safety of the public, the initial effects of gross equipment failures, such as a large reactor coolant pipe break, without loss of required integrity, and (b) together with other engineered safety features as may be necessary, to retain for as long as the situation requires, the functional capability of the containment to the extent necessary to avoid undue risk to the health and safety of the public. (GDC 10). The design pressure and temperature of the containment exceeds the peak pressure and temperature occurring as the result of the complete blowdown of the reactor coolant through any rupture of the reactor coolant system up to and including the hypothetical double-ended severance of a reactor coolant pipe. Energy contribution from the steam system is included in the calculation of the containment pressure transient due to reverse heat transfer through the steam generator tubes. The supports for the reactor coolant system are designed to withstand the blowdown forces associated with the sudden severance of the reactor coolant piping so that the coincidental rupture of the steam system is not considered credible. In 1989, the NRC approved changes to the design bases with respect to dynamic affects of postulated primary loop pipe ruptures, as discussed in Section 4.1.2.4. The containment structure and all penetrations are designed to withstand, within design limits, the combined loadings of the design-basis accident and design and maximum potential seismic conditions. All piping systems that penetrate the vapor barrier are anchored at the liner. The penetrations for the blowdown, and sample lines are designed so that the penetration is stronger than the piping system and so that the vapor barrier is not breached due to a hypothesized pipe rupture. The pipe rupture loads for the main steam and feedwater lines are resisted by the supports located away from their penetrations and do not affect the integrity of the penetrations for these lines. The pipe capacity in flexure is assumed to be limited to the plastic moment, based upon the yield strength of the pipe material. All lines (with the exception of sample tubing) connected to the primary coolant system that penetrate the vapor barrier are also restrained near the secondary shield walls (i.e., walls surrounding the steam generators and reactor coolant pumps) and are each provided with at least one valve between the shield wall and the reactor coolant system. These restraints are designed to withstand the thrust, moment, and torque resulting from a hypothesized rupture of the attached pipe. All isolation valves are supported to withstand, without impairment of valve operability, the combined loadings of the design basis accident and design and maximum potential seismic conditions. Section 5.1.5 includes a discussion of the details of the design of primary system supports. In addition, the design pressure will not be exceeded during any subsequent long-term pressure Chapter 5, Page 3 of 89 Revision 22, 2010

IP2 FSAR UPDATE transient determined by the combined effects of heat sources, such as residual heat and limited metal-water reactions, structural heat sinks, and the operation of the engineered safeguards, which uses only the emergency electric power supply. 5.1.1.1.6 Reactor Containment Design Basis Criterion: The reactor containment structure, including openings and penetrations, and any necessary containment heat removal systems, shall be designed so that the leakage of radioactive materials from the containment structure under conditions of pressure and temperature resulting from the largest credible energy release following a loss-of-coolant accident, including the calculated energy from metal-water or other chemical reactions that could occur as a consequence of failure of any single active component in the emergency core cooling system, will not result in undue risk to the health and safety of the public. (GDC 49) The following general criteria are followed to ensure conservatism in computing the required structural load capacity:

1. In calculating the containment pressure, rupture sizes up to and including a double-ended severance of reactor coolant pipe are considered.

2. In considering postaccident pressure effects, various malfunctions of the emergency systems are evaluated. Contingent mechanical or electrical failures are assumed to disable one of the diesel generators, two of the five fan-cooler units, and one of the two containment spray units. Equipment, which can be run from diesel power is described in Chapter 8.

3. The pressure and temperature loadings obtained by analyzing various loss-of-coolant accidents, when combined with operating loads and maximum wind or seismic forces, do not exceed the load-carrying capacity of the structure, its access opening or penetrations.

The most stringent case of these analyses is summarized in Section 14.3.5.3.7. 5.1.1.1.7 Nil-ductility Transition Temperature Requirement for Containment Material Criterion: The selection and use of containment materials shall be in accordance with applicable engineering codes. (GDC 50). The selection and use of containment materials comply with the applicable codes and standards tabulated in Section 5.1.1.5. The concrete containment is not susceptible to a low-temperature brittle fracture. The containment liner is enclosed within the containment and thus is not exposed to the temperature extremes of the environs. The containment ambient temperature during operation is between 50 and 130°F. This includes both hot operating and cold shutdown conditions. The Containment liner was specified for impact testing at a temperature of 30°F below the minimum service temperature of 50°F. The large Containment steel penetrations, equipment hatch and personnel lock, were specified for impact testing at -50°F which is more than 30°F below the Chapter 5, Page 4 of 89 Revision 22, 2010

IP2 FSAR UPDATE outside containment temperature of -5°F. These tests assure that the Nil Ductility Transition Criterion of GDC 50 is met. 5.1.1.2 Supplementary Accident Criteria Systems relied upon to operate under postaccident conditions, which are located external to the containment and communicate directly with the containment, are considered to be extensions of the leakage boundary. The pressure retaining components of the containment structure are designed for the maximum potential earthquake ground motion of the site combined with the simultaneous loads of the design basis accident as follows:

1. The liner is designed to ensure that no average strains greater than the strain at the guaranteed yield point occur at the factored loads. In regions of local stress concentrations or stresses due to localized secondary load effects, the liner is permitted to yield but the maximum liner strain is limited to 0.5-percent.

2. The mild steel reinforcement is designed to ensure that no strains greater than the strain at the guaranteed yield point occur at a cross section under the factored loads. The local yielding of reinforcing bars are permitted around the large openings for load combinations that include seismic loads.

The pressure-retaining components of containment subject to deterioration or corrosion in service are provided with appropriate protective means or devices (e.g., protective coatings). 5.1.1.3 Energy and Material Release The design pressure is not exceeded during any subsequent long-term pressure transient determined by the combined effects of heat sources such as residual heat and metal-water reactions, structural heat sinks, and the operation of other engineered safety features utilizing only the emergency onsite electric power supply. The mass and energy releases to and the accident pressure and temperature effects on the containment structures, are those created by the hypothetical large break loss-of-coolant accident as presented in Section 14.3.5. The following loadings are considered in the design of the containment in addition to the pressure and temperature conditions described above:

1. Structure dead load.

2. Live loads.

3. Equipment loads.

4. Internal test pressure.

5. Earthquake.

6. Wind.

5.1.1.4 Engineered Safety Features Contribution Five types of engineered safety features are included in the design of this facility to ensure containment integrity. These systems are discussed in Chapter 6 and their effectiveness is analyzed in Chapter 14. Chapter 5, Page 5 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.1.5 Codes and Standards The design, materials, fabrication, inspection, and proof testing of the containment vessel complies with the applicable parts of the following codes and standards. Code Title

1. ASTM A-333, Gr. 1 Specification for Seamless and Welded Steel Pipe for Low Temperature Service

2. ASTM A-181 Forged or Rolled Steel Pipe Flanges, Forged Fittings, and Valves and Parts for General Service

3. ASTM A-300, Cl. 1, Firebox Specification for Notch Toughness Requirements for Normalized Steel Plates for Pressure Vessels

4. ASTM A-201, Gr. B Specification for Carbon Silicon Steel Plates of Intermediate Tensile Ranges for Fusion Welded Boilers and other Pressure Vessels

5. ASTM A-36 Specification for Structural Steel

6. ASTM A-131, Gr. C Specification for Structural Steel for Ships

7. ASTM A-240 Specification for Heat-Resisting Chromium and Chromium-Nickel Stainless Steel Plate, Sheet, and Strip for Fusion-Welded Unfired Pressure Vessels

8. ASTM A-312 Specification for Seamless and Welded Austenitic Stainless Steel Pipe

9. ASTM A442, Grade 60 Specifications for Pressure Vessel Plates, Carbon Steel, Improved Transition Properties

10. ASME Boiler and Pressure Nuclear Nuclear Vessels Vessels Vessel Code-Section III

11. ASME Boiler and Pressure Unfired Unfired Pressure Vessels Pressure Vessels Vessel Code-Section VIII

12. ASME Boiler and Pressure Welding Welding Qualifications Qualifications Vessel Code-Section IX

13. ASTM C-33 Standard Specifications for Concrete Aggregates

14. ASTM C-150 Standard Specifications for Portland Cement

15. ASTM C-172 Standard Method of Sampling Fresh Concrete

16. ASTM C-31 Standard Method of Making and Curing Concrete Compression and Flexure Test Specimens in the Field

17. ASTM C-39 Standard Method of Test for Compressive Strength of Molded Concrete Cylinders

18. ASTM-C-350 Specifications For Fly Ash For Use As AN Admixture in Portland Cement Concrete

19. ASTM C-94 Specifications for Ready Mixed Concrete Chapter 5, Page 6 of 89 Revision 22, 2010

IP2 FSAR UPDATE

20. ASTM C-42 Standard Methods of Securing, Preparing, and Testing Specimens from Hardened Concrete for Compressive and Flexural Strengths

21. ASTM C-494 Specifications for Chemical Admixtures for Concrete

22. ASTM A-305 Specifications for Minimum Requirements for Deformations of Deformed Steel Bars for Concrete Reinforcement

23. ASTM A-408 Specifications for Special Large Size Deformed Billet-Steel Bars for Concrete Reinforcement

24. ASTM A-432 Specification for Deformed Billet Steel Bars for Concrete Reinforcement with 60,000 psi Minimum Yield Strength

25. Research Council of Riveted and Specification For Structural Joints Using Bolted Structural Joints of the ASTM A-325 Bolts Engineering Foundation

26. ACI-613 Recommended Practice for Selecting Proportions for Concrete

27. ACI-306 Recommended Practice for Winter Concreting

28. ACI-318, Part IV-B Structural Analysis and Proportioning of Members-Ultimate Strength Design

29. ACI-318 Building Code Requirements for Reinforced Concrete

30. ACI- 505 Specification for the Design and Construction of Reinforced Concrete Chimneys

31. ACI-315 Manual of Standard Practice for Detailing Reinforced Concrete Structures

32. ASA N6.2 Safety Standards for the Design, Fabrication and Maintenance of Steel Containment Structures for Stationary Nuclear Power Reactors

33. ASA A58.1 American Standard Code Requirements for Minimum Design Loads in Buildings and Other Structures

34. State Building and Construction Code for the State of New York

35. SSPC-SP-6 Commercial Blast Cleaning 5.1.2 Containment Structure Design 5.1.2.1 General Description The reactor containment structure is a reinforced concrete vertical right cylinder with a flat base and hemispherical dome. A welded steel liner with a minimum thickness of 0.25-in. is attached to the inside face of the concrete shell to ensure a high degree of leaktightness. The design Chapter 5, Page 7 of 89 Revision 22, 2010

IP2 FSAR UPDATE objective of the containment structure is to contain all radioactive material, which might be released from the core following a loss-of-coolant accident. The structure serves as both a biological shield and a pressure container. The structure, as shown on Plant Drawings 9321-2501, 9321-2502, 9321-2503, 9321-2506, 9321-2507, 9321-2508, [Formerly UFSAR Figures 5.1-2 through 5.1-7] and Figures 5.1-1 consists of side walls measuring 148-ft from the liner on the base to the springline of the dome, and has an inside diameter of 135-ft. The side walls for the cylinder and the dome are 4-ft 6-in. and 3-ft 6-in. thick respectively. The inside radius of the dome is equal to the inside radius of the cylinder so that the discontinuity at the springline due to the change in thickness is on the outer surface. The cylindrical part of the liner is substantially round. The difference between the minimum and maximum inside diameters at any selected cross section does not generally exceed 0.25-percent of the nominal diameter at that elevation. Between elevations 43-ft and 95-ft, the maximum diameter of any cross section is 135-ft 2-in., and the minimum diameter is 134-ft 10-in. except at the liner closing the temporary opening in the northwest quadrant where a minimum diameter of 134-ft 8-5/8-in. was measured. This portion of the liner was erected after all exterior concrete work was completed and is within the local buckle allowance of the liner plates. Above elevation 95 ft the tolerance on inside diameter does not exceed 0.50-percent of the nominal diameter of the selected cross section. The liner is erected true and plumb so that the deviation does not exceed 1/500 of the height at the selected cross section (allowing for 2-in. local buckling of the liner plates). Particular care is taken in matching edges of cylindrical and hemispherical sections to ensure that all joints are properly aligned. Maximum permissible offset of completed joints is 25 percent of nominal plate thickness. Plates buckled beyond acceptable limits are cut out and replaced with new plates. The flat concrete base mat is 9-ft thick with the bottom liner plate located on top of this mat. The bottom liner plate is covered with 3-ft of concrete, the top of which forms the floor of the containment. Where uplift from pressure occurs at the outer areas of the mat, the 9-ft thick mat has sufficient flexural capacity to resist the uplift. No hydraulic uplift exists since the bottom elevation of the mat is considerably higher than that of the high water level. The large mass of the containment including interior concrete and equipment makes the structure inherently stable from overturning due to seismic motion. In addition, keying action from the reactor pit and sumps, plus friction between the concrete and rock, prevents a sliding of the structure from horizontal ground motion. The basic structural elements considered in the design of the containment structure are the base slab, side walls, and dome acting as one structure under all possible loading conditions. The liner is anchored to the concrete shell by means of stud anchors. The lower portions of the cylindrical liner are insulated to avoid thermal deformation of the liner under accident conditions. The containment structure is inherently safe with regard to common hazards such as fire, flood, and electrical storm. The thick concrete walls are invulnerable to fire and only an insignificant Chapter 5, Page 8 of 89 Revision 22, 2010

IP2 FSAR UPDATE amount of combustible material, such as lubricating oil in pump and motor bearings, is present in the containment. Internal structures consist of equipment supports, shielding, reactor cavity and canal for fuel transfer, and miscellaneous concrete and steel for floors and stairs. All internal structures are supported on the mat with the exception of equipment supports secured to the intermediate floors. A 3-ft thick concrete ring wall serving as a missile and partial radiation shield surrounds the reactor coolant system components and supports the polar-type reactor containment crane. A 2-ft thick reinforced concrete floor covers the reactor coolant system with removable gratings in the floor provided for crane access to the reactor coolant pumps. The four steam generators, pressurizer, and various piping penetrate the floor. Spiral stairs provide access to the areas below the floor. The refueling canal connects the reactor cavity with the fuel transport tube to the spent fuel pool. The floor and walls of the canal are concrete, with wall and shielding water providing the equivalent of 6-ft of concrete. The refueling canal floor is 5-ft thick. The concrete walls and floor are lined with 0.25-in. thick stainless steel plate. The linings provide a leakproof membrane that is resistant to abrasion and damage during fuel handling operation. Waterproofing is provided in the areas of the containment in contact with backfill to prevent ground-water seepage. This consists of a coat of bitumastic No. 50, a 0.625-in.-thick layer of hardboard insulation, and a second coat of bitumastic No. 50. Fill for innermost 5-ft from containment walls is crushed rock of maximum size of 6-in. and minimum amount of fines. All fill is free of vegetable matter. 5.1.2.2 Design Load Criteria The following loads are considered to act upon the containment structure creating stresses within the component parts.

1. Dead load consists of the weight of the concrete wall, dome, liner, insulation, base slab, and the internal concrete. Weights used for dead load calculations are as follows:

a. Concrete 150 lb/ft3

b. Reinforcing steel 490 lb/ft3 using nominal cross-sectional areas of reinforcing as defined in ASTM for bar sizes.

c. Steel lining 490 lb/ft3 using nominal cross-sectional area.

d. Insulation 6 lb/ft3 including stainless steel jacket.

2. Live load consists of snow and construction loads on the dome and major components of equipment in the containment. Snow and ice loads are assumed to be applied uniformly to the top surface of the dome at an estimated value of 20 lb/ft2 of horizontal projection of the dome. This loading represents approximately Chapter 5, Page 9 of 89 Revision 22, 2010

IP2 FSAR UPDATE 2-ft of snow, which is considered to be a conservative amount since the slope of the dome will tend to cause much of the snow that falls on it to slide off. A construction live load of 50 lb/ft2 has been used on the dome, but will not be considered to act concurrently with the snow load. Equipment loads are considered as specified on the drawings supplied by the manufacturers of the various pieces of equipment. Design live loads inside the containment building are as follows:

a. Elevation 68-ft-0-in. 10-ft strip adjacent to crane wall = 600 psf Remaining strip = 100 psf

b. Elevation 95-ft-0-in. Concrete slab = 500 psf Grating areas = 100 psf

3. The internal pressure transient used for the containment design and its variation with time is based on a postulated large break LOCA of 47 psig and liner temperature of 247oF. For the free volume of 2,610,000-ft3 within the containment, the design pressure is 47 psig. This pressure transient is more severe than those calculated for various loss-of-coolant accidents, which are presented in Section 14.3.

4. Thermal expansion stresses due to an internal temperature increase caused by a loss-of-coolant accident is considered. The maximum temperature at the uninsulated section of the liner under accident conditions is 247°F. For the 1.25 times and 1.50 times design pressure loading conditions given in Section 5.1.2.4, the corresponding liner temperatures will be 285°F and 306°F respectively. The minimum external ambient design temperature, averaged over a 24 hour period, is 0°F. The liner maximum temperature following a loss-of-coolant accident with an outside air temperature of 0°F was calculated to be less than 247oF at the Stretch Power Uprate (SPU) power rating of 3216 MWt for the core. The initial containment air temperature in the SPU analysis for the liner temperature was set to 110oF, which is the maximum expected operating temperature at 100%

power with an outside temperature of 0°F.

5. The ground acceleration for the design earthquake has been determined to be 0.1g applied horizontally and 0.05g applied vertically. These values have been resolved as conservative numbers based upon recommendations from Dr.

Lynch, Director of Seismic Observatory, Fordham University. A dynamic analysis is used to arrive at equivalent design loads. Additionally, a hypothetical ground acceleration of 0.15 g horizontal and 0.10 g vertical is used to analyze for the no-loss-of-function. This is discussed in Section 5.1.3.11, Seismic Design. Due to symmetry of the containment structure, torsional loads generated by an earthquake are insignificant and have not been considered. Tornado loads have not been considered in the design of the Unit 2 containment; however, the seismic bars provide a more than adequate mechanism to Chapter 5, Page 10 of 89 Revision 22, 2010

IP2 FSAR UPDATE withstand the torsional effect if it were to occur. An evaluation of the effect of tornado loads on the containment structure is presented in Appendix B of the Containment Design Report.

6. The American Standards Association "American Standard Code Requirements for Minimum Design Loads in Buildings and Other Structures" (A58.1-1955) designates the site as being in a 25 psf zone for wind loads. In this code, for height zones between 100 and 499-ft, the recommended wind pressure on a flat surface is 40 psf. Correcting for the shape of the containment by using a shape factor of 0.60, the recommended pressure becomes 24 psf. The state building and construction code for the State of New York stipulates a wind pressure up to 30 psf on a flat surface for heights up to 300 feet. For design, a 30 psf basic wind load has been used from ground level up.

7. Internal pressure was applied to test the structural integrity of the containment shell up to 115-percent of the design pressure. For this structure, the test pressure is 54 psig. The containment is also structurally designed to withstand an external pressure 2.5 psig higher than the internal pressure.

5.1.2.3 Material Specifications Basically five materials are used for the construction of the containment structure. These are:

1. Concrete.

2. Reinforcing steel.

3. Plate steel liner.

4. Insulation.

5. Protective Coating.

Basic specifications for these materials are as follows:

1. Concrete is a dense, durable mixture of sound coarse aggregate, fine aggregate, cement, and water. Cement conforms to ASTM, Specification C-150-65 "Standard Specification for Portland Cement," Type I (Normal), or Type II (moderate heat of hydration) requirements. Whenever high early strength is required, Type III Cement is used. Water is free from any injurious amounts of acid, alkali, salts, oil, sediment, or organic matter. The concrete has a minimum density of 150 lb/ft3. The 28-day standard compressive strength of the concrete is 3000 psi. Adequate means of control are used in the manufacture of the concrete. To ensure the values of compressive strength are attained as a minimum, concrete samples are tested in accordance with the following ASTM Standards:

ASTM C-172 - Standard Method of Sampling Fresh Concrete ASTM C Standard Method of Making and Curing Concrete Compression and Flexure Test Specimens in Field ASTM C Standard Method of Test for Compressive Strength of Molded Concrete Cylinders Chapter 5, Page 11 of 89 Revision 22, 2010

IP2 FSAR UPDATE All making and testing of concrete samples have been performed by Vacca Testing Laboratory and Research Company, Inc. At certain specifically evaluated locations, non-structural surface type cracks and delaminations in the containment concrete have been repaired by injection of engineering approved epoxy grout. Although non-structural in nature, these repairs were performed in accordance with the requirements of IWL-4210 of the 1992 ASME Boiler and Pressure Vessel Code, Section XI, as applicable.

2. Reinforcing steel for the dome, cylindrical walls and base mat is high-strength, deformed billet steel bars conforming to ASTM Designation A432-65 "Specification for Deformed Billet Steel Bars for Concrete Reinforcement with 60,000 psi Minimum Yield Strength." This steel has a minimum yield strength of 60,000 psi, a minimum tensile strength of 90,000 psi, and a minimum elongation of 7-percent in an 8-in. specimen. Reinforcing bars No. 11 and smaller in diameter are lapped spliced in the mat for flexural loadings and spliced by the Cadweld process in the walls and dome for tension loading. Bars No. 14S and 18S are spliced by the Cadweld process only. A certification of physical properties and chemical content of each heat of reinforcing steel delivered to the job site has been issued from the steel supplier. The splices used to join reinforcing bars have been tested to ensure that they will develop at least 125-percent of the minimum yield point stress of the bar. The test program required cutting out, at random, approximately 3-percent, completed splices and testing to determine their breaking strength.

3. The plate steel liner is carbon steel conforming to ASTM Designation A442-65 "Standard Specification for Carbon Steel Plates with Improved Transition Properties," Grade 60. This steel has a minimum yield strength of 32,000 psi and a minimum tensile strength of 60,000 psi with an elongation of 22-percent in an 8-in. gauge length at failure.

The liner is 0.25-in. thick at the bottom, 0.50-in. thick in the first three courses, except 0.75-in. thick at penetrations, a minimum of 0.34-in. in the general area at elevation 46-ft. due to past corrosion, and 0.375-in. thick for remaining portion of the cylindrical walls and 0.50-in. thick in the dome. The 0.34-in. minimum thickness affects the calculated stress levels presented in the Containment Design Report and the Containment Liner Stress Analysis Report. However, evaluation of the reduced minimum thickness has concluded that no design criteria are exceeded. The liner material has been tested to ensure an NDTT more than 30°F lower than the minimum operating temperature of the liner material. Impact testing has been done in accordance with Section N331 of Section III of the ASME Boiler and Pressure Vessel Code. A 100-percent visual inspection of liner anchors was made prior to pouring concrete.

4. The material used for the original insulation of the liner plate was polyvinylchloride with stainless steel jacket. This insulation has been selected to withstand the calculated temperature and pressure conditions associated with a postulated large break LOCA of 47 psig and liner temperature of 247oF. The carbon steel liner with an inorganic zinc protective coating makes contact with Chapter 5, Page 12 of 89 Revision 22, 2010

IP2 FSAR UPDATE the polyvinylchloride insulation, the stainless steel, and the sealant. However, these materials do not react with each other. Because the insulation panels are jacketed with stainless steel and sealed at the joints, the insulation will not be subjected to the moisture and high humidity atmosphere of the containment during an accident. Manufacturer's tests on the polyvinylchloride insulation indicated that the insulation was capable of withstanding periodic compression at 60 psig at temperatures from 40°F to 120°F and a single compression under accident conditions without any detriment or change to the insulation properties. The manufacturer's analog transient analysis indicated only a 5°F rise in liner temperature 1000 sec after an exposure to 310°F for the entire duration of the analysis. This provides a factor of safety of approximately 15 on specified tolerable temperature rise in the liner. A factor of safety of 2 is provided on specified insulation performance versus tolerable temperature rise in liner. The maximum normal operating temperature of the containment was changed from 120°F to 130°F by Amendment 149 to the Facility Operating License DPR-26 for IP-2 dated March 27,1990. Evaluations performed show the insulation material used on the containment liner is adequate for use at the higher operating temperature. For additional information on the liner insulation and the modifications made to it in 1973, see Section 5.1.7.

5. One 3 mil shop coat of Carbozinc No. 11 primer and one 4 mil minimum finish coat of Phenoline No. 305 as manufactured by the Carboline Company have been applied to the liner, as well as essentially all painted surfaces in containment, in accordance with the manufacturer's recommendations.

The effect of the postaccident environment on protective coatings was conservatively evaluated for Indian Point Unit 2. The coatings showed no deterioration after a number of cycles. A more thorough discussion on the qualifications of the protective coatings applied during construction is presented in WCAP-7198-L.1 In addition, various areas inside containment have been repaired and recoated with other DBA qualified coatings approved for use at Indian Point 2. Protective coatings used inside the containment are procured, applied, and maintained in compliance with Regulatory Guide 1.54 (June 1973), Quality Assurance Requirements for Protective Coatings Applied to Water Cooled Nuclear Power Plants. New quality requirements will be developed based on its provisions, but specific requirements, such as documented site meetings, field demonstrations, substrate priming, applicator reporting, inspection reporting and report forms will be considered on a job-by-job basis. Quality of both materials and construction of the containment structure was ensured by a continuous program of quality control and inspection by Con Edison, and/or its field representatives, and Westinghouse Atomic Power Division, and United Engineers and Constructors Inc., as described in Section 5.1.2.6. Chapter 5, Page 13 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.2.4 Design Stress Criteria The design is based upon limiting load factors that are used as the ratio by which loads will be multiplied for design purposes to ensure that the loading deformation behavior of the structure is one of elastic, tolerable strain behavior. The load factor approach is being used in this design as a means of making a rational evaluation of the isolated factors, which must be considered in ensuring an adequate safety margin for the structure. This approach permits the designer to place the greatest conservatism on those loads most subject to variation and which most directly control the overall safety of the structure. In the case of the containment structure, therefore, this approach places minimum emphasis on the fixed gravity loads and maximum emphasis on accident and earthquake or wind loads. The loads utilized to determine the required limiting capacity of any structural element on the containment structure are computed as follows:

1. C = 1.0D +/- 0.05D + 1.5 P + 1.0 (T + TL)

2. C = 1.0D +/- 0.05D + 1.25 P + 1.0 (T' + TL') + 1.25E

3. C = 1.0D +/- 0.05D + 1.0P + 1.0 (T" + TL") + 1.0E' Symbols used in these formulae are defined as follows:

C = Required load capacity of section. D = Dead load of structure and equipment loads. P = Accident pressure load as shown on pressure-temperature transient curves. T = Load due to maximum temperature gradient through the concrete shell and mat based upon temperature associated with 1.5 times accident pressure. TL = Load exerted by the liner based upon temperatures associated with 1.5 times accident pressure. T' = Load due to maximum temperature gradient through the concrete shell and mat based upon temperatures associated with 1.25 times accident pressure. TL' = Load exerted by the liner based upon temperatures associated with 1.25 times accident pressure. E = Load resulting from either design earthquake or wind, whichever is greater. T" = Load due to maximum temperature gradient through the concrete shell and mat based upon temperatures associated with the accident pressure. TL" = Load exerted by the liner based upon temperatures associated with the accident pressure. E' = Load resulting from assumed hypothetical earthquake. Chapter 5, Page 14 of 89 Revision 22, 2010

IP2 FSAR UPDATE A chart for allowable versus actual stresses has been included in the Containment Design Report. Load condition (1) indicates that the containment will have the capacity to withstand loadings at least 50-percent greater than those calculated for the postulated loss-of-coolant accident alone. Results of analysis using load condition (1) are shown in Figure 5.1-11. Load condition (2) indicates that the containment will have the capacity to withstand loadings at least 25-percent greater than those calculated for the postulated loss-of-coolant accident with a coincident design earthquake. Results of analysis using load condition (2) are shown in Figure 5.1-12. Load condition (3) indicates that the containment will have the capacity to withstand loads at least equal to those calculated for the postulated loss-of-coolant accident with a coincident hypothetical earthquake defined in Section 5.1.2.2. Results of analysis using load condition (3) are shown in Figure 5.1-13. The mat has been analyzed using load conditions (1), (2) and (3) as shown in Figures 5.1-14 through 5.1-16 and also for loads occurring only at operating and test pressure conditions. For loads, see Table 5.1-1, Flooded Weights-Containment Building. The loads resulting from wind on any portion of the structure do not exceed those resulting from earthquake. The capacity of all structural components, with the minor exceptions of outer rebar at large containment openings addressed in Section 3.4.4 of the Containment Design Report, exceeds or is equal to the capacity required by the most severe loading combination. The loads resulting from the use of these equations will hereafter be termed "factored loads. The load factors used in these equations are based upon the load factor concept employed in Part IV-B, "Structural Analysis and Proportioning of Members Ultimate Strength Design" of ACI 318-63. Because of the refinement of the analysis and the restrictions on construction procedure, the load factors in the design primarily provide for a safety margin on the load assumptions. The design includes the consideration of both primary and secondary stresses. The design limit for tension member (i.e., the capacity required for the design load) is based upon the yield stress of the reinforcing steel. The theoretical load carrying capacity of steel reinforced concrete cross-sections are reduced by a capacity reduction factor "", which provides for the possibility that small adverse variations in material strengths, workmanship, dimensions, and control, while individually within required tolerances and the limits of good practice, occasionally may combine to result in under-capacity. For tension members, the factor "" has been established as 0.95. The factor "" is 0.90 for flexure and 0.85 for diagonal tension, bond, and anchorage. For principle compression and tension, the liner stresses are maintained below 0.95 specified minimum yield at normal operating temperature (i.e., =0.95). For shear, the liner stresses are maintained below 0.6 yield. Chapter 5, Page 15 of 89 Revision 22, 2010

IP2 FSAR UPDATE The liner is designed to assure that no strains greater than the strain at the guaranteed yield point will occur at the factored loads. In regions of local stress concentrations or stresses due to localized secondary load effects, the liner is permitted to yield but the maximum liner strain is limited to 0.5-percent. Sufficient anchorage is provided to ensure elastic stability of the liner. The basic design concept for the liner stud anchorage is the ductility of the anchorage that assures stud failure due to shear, tension or bending stress without the stud connection causing failure or tear of the liner plate. References 2 and 3 provide information on design of stud connection. The studs in the 0.50-in. plate are installed on 24-in. horizontal and 28-in. vertical grid and in the 0.375-in. plate on a 24-in. horizontal and 14-in. vertical grid. Studs are centered between vertical bars. In the dome, 5-ft by 5-ft panels are anchored in the center by studs and by Tbars at the edges. The 0.50-in. diameter bent welding studs are 9-in. long minimum and 9.50-in. long maximum with a 2-in. 90 degree hook at the end. An arc stud welding process was used on all bent welding studs. The arc stud welding process produces a circular weld around the 0.50-in. diameter stud with a diameter (outside to outside of weld) equal to 0.678-in. and a height equal to 0.157-in. The design considers the possibility of daily stress reversals due to ambient temperature changes for the life of the plant, and fatigue limit of the studs exceeds the design requirements. However, to accommodate possible fatigue failure in the plate-to-stud weldment, the depth of penetration to the liner plate is controlled to avoid impairment of liner integrity. The boundary conditions in the cylinder are determined by assuming a buckling model (shown in Figures 5.1-17 through 5.1-19) in which the studs form the low points and the center of the panels form the high points of a series of peaks and valleys thus forming a set of panels whose edges represent points of inflection. The analytical procedure used is a simply supported plate under biaxial compression. A Mohr's circle analysis is used to find the normal and shear stresses on this simply-supported plate. The critical buckling stress is derived considering a plate whose length is equal to one-half of the diagonal distance between studs. This critical buckling load is 38.1 ksi for the 0.375-in. liner and 38.4 ksi for the 0.50-in. liner, which is higher than the yield strength of the liner, 32 ksi; therefore, the liner plate will begin to yield before the critical buckling stress is reached, and buckling failure does not control the design. Since shear reduces the stability of a plate subjected to compressive stresses, critical shear is considered and it was found that critical buckling is controlled by normal stresses rather than shear stresses. This is determined by considering the magnitude of both the normal and the shear stresses on the panel. The magnitude of the shear is so low that it shows no effect on the previously stated critical buckling stresses. In the dome the liner will be considered clamped at the stiffeners forming a 5-ft by 5-ft grid panel pattern. The center of each panel is fixed by a stud. Assuming points of inflection at the one-quarter point a distance of 1-ft 3-in. occurs between points of simple support. The critical buckling load is 58.1 ksi, which is also higher than the yield strength of the liner. At maximum strain in the liner, the studs will not fail. This maximum strain due to an unbalanced load would occur in a panel adjacent to a buckled panel. Since this adjacent stud will not fail, no zipper effect will occur and massive buckling of the liner and mass failure of anchors is not credible. The anchorages can fail by failure of the studs in shear or tension, by studs pulling out from the concrete, or by studs separating from the liner plate. The most likely mode of failure is by tensile failure of the stud. The anchors are designed so that failure occurs in the anchor rather Chapter 5, Page 16 of 89 Revision 22, 2010

IP2 FSAR UPDATE than the plate, thereby ensuring that the leaktight integrity of the containment liner will be maintained. If failure should develop, it would be a random stud failure due to poor workmanship during stud attachment. This failure would not impair the liner integrity nor would it cause progressive failure. The anchor must resist tensile and shearing loads. Tests have indicated that the lateral load needed to prevent column buckling is 1-percent of the axial yield load. Conservatively doubling this value to account for uncertain field conditions, a value of 2-percent is used.4 The total load per plate would be 24-in. x 0.50-in. x 32,000 psi = 384,000 lb. Therefore, the tensile load per anchor is 384,000 lb x 0.02 = 7680 lb, which yields a stress of 7680/0.2 = 38,400 psi. This compares with a yield value of 50,000 psi and a tensile strength of 60,000 psi in the studs. This does not consider the internal pressure, which provides further stability against buckling. The shear load on the anchor is due to the strain in the liner. Assuming the liner approaches its yield strain of 0.1-percent, the anchor deflection would be 28-in. x .001 = .028-in. Tests on the stud anchor have shown a maximum deflection of about 0.1-in. can be tolerated before failure of the stud. 5.1.2.5 Missile Protection Except for the upper portions of the steam generators and the Pressurizer, the high pressure reactor coolant system equipment is surrounded by the 3-ft concrete shield wall enclosing the reactor coolant loop and by the 2-ft concrete operating floor. In 1989, the NRC approved changes to the design basis with respect to dynamic effects of postulated primary loop ruptures, as discussed in Section 4.1.2.4. A structure is provided over the control rod drive mechanism to block any missiles generated from fracture of the mechanisms. Systems containing hot pressurized fluids and that might affect the engineered safeguards components have been carefully checked against the possibility of being sources of missiles. The general criterion adopted has been to take provision, when necessary, against the generation of missiles rather than allow missile formation and try to contain their effects. Once the design requirement that the above systems are not to be sources of missiles has been set forth, identification of potential deficiencies and generation of adequate fixes took place through the quality assurance program. The following examples illustrate how this approach has been implemented. 5.1.2.5.1 Valves Valves installed in the nuclear steam supply system have stems with back seat. This rules out the probability of ejecting valve stems; even if it were assumed that the stem threads fail, analysis shows that the back seat or the upset end cannot penetrate the bonnet and thereby become a missile. Additional interference is encountered with air- and motor-operated valves. Chapter 5, Page 17 of 89 Revision 22, 2010

IP2 FSAR UPDATE Valves with nominal diameter larger than 2-in. have been designed against bonnet-body connection failure and subsequent bonnet ejection by means of: (1) following the EPRI recommendations12 regarding bolting practices; (2) using the design practice of ASME Section VIII for flange design; and (3) by controlling the pre-load during the bonnet body connection stud tightening process. The pressure-containing parts except the flange and studs are designed per criteria established by USAS B16.5. Flanges and studs are designed in accordance with ASME Section VIII. Piping and flange materials of construction are procured per ASTM A182, F316, or A351, GR CF8M. Stud and nut material is ASTM A193-B7 and A194-2H. The proper stud torquing procedures and the use of a torque wrench, with indication of the applied torque, limit the stress of the studs to the allowable limits established in the EPRI Good Bolting Practices Reference Manual (NP-5067).12 This stress level is far below the material yield, i.e., about 105,000 psi. The complete valves are hydro-tested per USAS B16.5 (1500 lb USAS valves are hydro-tested to 5400 psi). The cast stainless steel bodies and bonnets are radiographed and dye penetrant tested to verify soundness. Valves with nominal diameter of 2-in. or smaller are forged and generally have screwed bonnet with canopy seal. The canopy seal is the pressure boundary while the bonnet threads are designed to withstand the hydrostatic end force. The pressure containing parts are designed per criteria established by the USAS B16.5 specification. 5.1.2.5.2 Reactor Coolant Pump Flywheel The reactor coolant pump flywheel is not considered to be a credible source of missiles because of conservative design and care in manufacture and inspection. The flywheel material is ASTM A-533 having an nil-ductility transition temperature less than 10oF. The design results in a primary stress less than 50-percent of the material yield strength at operating speed. The flywheel was subjected to 100-percent volumetric ultrasonic inspection, which is repeated at intervals during plant life. The finished machined bore is subjected to either magnetic particle or liquid penetrant examination. The design overspeed of the pump is 125-percent. The maximum pump overspeed on loss of external load is 112-percent. For an additional discussion of integrity of the reactor coolant pump flywheel, see Section 4.2.2. 5.1.2.6 Quality Control To ensure a high degree of confidence in plant design, construction, workmanship, materials, and performance, a quality control program has been in effect for this project in which the following principal organizations have their respective responsibilities:

1. Consolidated Edison Company of New York, Inc. as initial owner and operator of the plant.

2. Westinghouse Electric Corporation as the turnkey plant contractor and supplier of major equipment.

3. United Engineers and Constructors Inc. as architect-engineer, construction manager, and constructor.

Chapter 5, Page 18 of 89 Revision 22, 2010

IP2 FSAR UPDATE The function and responsibility in the quality control program of each of the above organizations is as follows: 5.1.2.6.1 Consolidated Edison Company of New York, Inc. (Con Edison) - Initial Licensee A qualified field representative was assigned to the field during the construction period. His responsibilities included continuous inspection of the construction of the containment building to ensure that all materials used and work performed was strictly in accordance with the plans and specifications. The Con Edison representative, through instructions received from the home office, had the power to stop the construction until any discrepancies were corrected and the work once more was in compliance with the specifications and plans. The Con Edison representative was in constant communication and consultation with the construction superintendent in matters regarding quality control. In addition, personnel from U.S. Testing Laboratories were assigned to this project to monitor the inspection of the construction and obtain samples of the materials for testing. 5.1.2.6.2 Westinghouse Electric Corporation For the assurance of plant integrity and quality, Westinghouse performed the following functions regarding the containment building:

1. Reviewed and approved the containment design criteria, material specifications and detail design concepts before they were released for construction. This work was done by qualified structural engineers at the company's home office.

2. Reviewed the construction and inspection methods employed by United Engineers and Constructors Inc.

Westinghouse Pressurized Water Reactor Division, Nuclear Power Services Group had a field quality assurance representative in residence during the construction period. His function was the same as the Con Edison representative mentioned above. He reported discrepancies to the Westinghouse Construction and Services resident engineer who had the authority to stop the work until the discrepancy was resolved. In addition to this, he audited the construction files, and verified that records were complete, accurate, and adequate for quality assurance. Nuclear Power Service Headquarters quality assurance engineers also made trips to the site to audit, monitor, and review the project with regard to site quality assurance. Construction practices were observed for conformance to codes, specifications, and approved procedures. 5.1.2.6.3 United Engineers and Constructors Inc. The responsibilities of United Engineers and Constructors Inc. in the quality control of the containment building were as follows:

1. They inspected all materials delivered to the job site, and examined the suppliers' certified test reports of physical and chemical properties for those components furnished by them.

Chapter 5, Page 19 of 89 Revision 22, 2010

IP2 FSAR UPDATE

2. They inspected fabrication of major components of the containment structure in the shop. Trip reports are available at the site.

3. They maintained an adequate force of qualified supervisory personnel at all times.

4. They supervised and were fully responsible for the quality of work performed by their subcontractors and for the craft labor employed and supervised by them.

5. They maintained as part of their field engineering force, qualified personnel who performed a thorough inspection of each construction operation.

No changes in design or specifications were allowed without the approval of the engineer in charge of design. 5.1.3 Containment Stress Analysis 5.1.3.1 General The structural design of the containment meets the requirements established by 1961 edition of "The State Building and Construction Code for the State of New York" so far as these provisions are applicable. All concrete structures have been designed, detailed, and constructed in accordance with the provisions of "Building Code Requirements for Reinforced Concrete" (ACI 318-63) so far as these provisions are applicable. 5.1.3.2 Method of Analysis Basically three separate structural components have been analyzed, each in equilibrium with loads applied to it and with constraints occurring at the juncture of the structures. The three components are:

1. The 135-ft ID hemispherical dome.

2. The 135-ft ID cylinder.

3. The base slab.

Mathematically, the dome and cylinder have been treated as thin-walled shell structures, which results in a membrane analysis. Since the thickness of the dome and cylinder is small in comparison with the radius of curvature (1/20 and 1/15) and there are no discontinuities such as sharp bends in the meridonal curves, the stresses due to pressure and wind or earthquake are calculated by assuming that they are uniformly distributed across the thickness. Since the concrete is not assumed to resist any tensile or shear forces, radial shear reinforcing has been introduced in the lower portion of the wall in the form of hooked diagonal stirrups and diagonally bent bars as shown in Figure 5.1-1. Diagonal shear reinforcing, at 45° and 135° to the circumferential direction, are placed in the center of the cylinder wall for the full height of the wall and a distance above the springline into the dome to resist earthquake shears. The diagonal bars are discontinued in the upper area of the dome (beyond about 30 degrees above the springline), where the seismic shears are small and are carried by the dome reinforcing steel lying in the plane of principal tension. Chapter 5, Page 20 of 89 Revision 22, 2010

IP2 FSAR UPDATE The base slab has been treated as a flat circular plate supported on a rigid nonyielding foundation. The limiting cases in the design of the wall for discontinuity moments and shears were considered. One case considered an uncracked wall and the other considered a cracked wall with the steel acting as a spring constant. The value of c varied from zero in the cracked case to .14 in the uncracked case. In the uncracked case, variations in Ec will have no effect on the answer since Ec appears in both the numerator and the denominator of the stiffness formulation. For the above variation in Ec and c, the values of discontinuity moment and shear vary by 14-percent and 7-percent respectively at the base. These are the maximum deviations of the wall forces since the wall will actually vary from uncracked to cracked with an increase in containment height rather than be cracked or uncracked for the total height. In the area of thermal stress, the entire wall section will be cracked and no variation in Ec or c need be considered. The liner stresses depend on the strains of the reinforcing steel and are not related to the concrete properties. Shrinkage and creep effects will be relieved by cracking during the pressure test and will not be included in accident design considerations. The finite element computer program has the capabilities of taking into account variations in c and Ec and axisymmetric loads. However, it is not necessary to take into account the variations in c and Ec for the reasons stated above. The computer program used to study the general behavior of the structure and to generate boundary conditions was the axisymmetric shell structure program. This computer program, developed by Franklin Institute Research Laboratories, is designed to handle arbitrarily shaped shells of revolution subjected to axisymmetric as well as nonaxisymmetric loadings. The method of analysis consists of subdividing the shell into elements having continuous meridians with continuous first and second derivatives so that the first and second fundamental forms of the resulting shell elements are continuous throughout the element. By expanding the dependent variables in Fourier series in the circumferential direction, and assigning unspecified functions for the meridional variation, the independent variables are separated and a system of ordinary differential equations results for the dependent variables in terms of the meridional independent variable. Particular and complementary solutions of these ordinary differential equations are then found for each of the elements and each of the circumferential harmonics individually. The matching of the elements is achieved by writing the required boundary conditions. The idealized section used with the axisymmetric shell structure program consists of five layers whose moment of inertia is equal to that of the actual section. The wall section is considered as cracked with the reinforcing carrying all loads. A finite element program, with the capability to incorporate thermal loads, was used to analyze the containment shell considering the effect of the equipment hatch opening. The shell was idealized into 10 layers with alternate layers of steel and concrete. Section 5.1.3.10 provides more information on the finite element analysis. Chapter 5, Page 21 of 89 Revision 22, 2010

IP2 FSAR UPDATE The computer program can handle the loads in the form of either surface traction or edge loads or both. Analysis of the liner is presented in the Containment Liner Stress Analysis Report. The report also contains a description of analytical procedures arriving at forces, shears, and moments in the structural shell. 5.1.3.3 Dome Analysis The analysis of the hemispherical dome has been performed by the super-position of membrane forces resulting from gravity, accident pressure, and accident thermal loads. In addition, earthquake or wind loading create both direct and shear stresses in the dome, and the operating temperature of the liner creates tension and compression. All of the combined direct stresses are developed in the reinforcing steel encased in the concrete. In the upper area of the dome (about 30 degrees above the springline), where the seismic shears are small, seismic shears are carried by dome reinforcing steel lying in the plane of the principal tension. The dome reinforcing is spliced to the vertical steel in the cylindrical concrete wall, so that a continuity between the dome and the cylinder is realized. See Figure 5.1-20 for a section of wall, dome and for reinforcing in the dome. 5.1.3.4 Cylinder Analysis The analysis of the cylinder is by superposition of membrane forces resulting from gravity, pressure and thermal loads, overturning due to earthquake or wind and shears due to earthquake or wind. The concrete has been reinforced circumferentially using steel hoops and vertically by straight bars. Diagonal bars have been placed to resist the horizontal and vertical shears due to earthquake or wind. The required capacity of the diagonal bars has been designed so that the horizontal component per foot of the diagonals is equal to the maximum value of shear flow. A check was made to ensure that no net compressive force results in the diagonal bars because of the combination of seismic shear load and internal pressure load. Although, in the cylinder, the liner has some capacity available to resist the seismic shears, no credit is taken for this capacity. For all of the cylinder and the lower areas of the dome, the diagonal reinforcing has been designed to accommodate all seismic shears. No credit has been taken for the dowel action of the vertical and horizontal bars in resisting seismic shear. Only in the upper area of the dome (beyond about 30 degrees above the spring line) where the seismic shears are small is the liner counted on to resist shear. For all of the cylinder and the lower areas of the dome, the diagonal reinforcing has been designed to accommodate all seismic shears. No credit has been taken for the dowel action of the vertical and horizontal bars in resisting seismic shear. 5.1.3.5 Base Mat Analysis The base slab was treated as a flat circular plate supported on a rigid non-yielding foundation. For loads applied uniformly around the slab, the analysis considers a 1-ft wide beam fixed at a point where the vertical shear is equal to zero. This is the point where the downward pressure on the mat and the dead weight overcome the uplift at the containment wall base mat juncture from pressure and earthquake loadings. Radial and circumferential reinforcing is provided at the top and bottom of the mat to resist moments in the areas where uplift occurs. Temperature Chapter 5, Page 22 of 89 Revision 22, 2010

IP2 FSAR UPDATE steel was added in other areas to meet requirements of the (ACI-318) Code. Diagonal tension reinforcement was added to meet requirements of ACI-318 Code. See Figure 5.1-23 for base slab reinforcing detail. Moments and shears were calculated by writing equations for moment and shear in terms of X using the containment wall-base slab juncture as the origin with X increasing toward the center of the containment building. The point along the circumference of the containment wall chosen as the end of the beam is a point where the maximum tension from the earthquake will exist. Since the containment structure is considered a beam in all earthquake analyses, the maximum uplift for which the mat is designed will occur at only one point on the circumference and will represent the worst possible uplift on the mat. All stresses were calculated using Part IV-B Structural Analysis and Proportioning of Members - Ultimate Strength Design of the Building Codes Requirements for Reinforced Concrete (ACI318-63). No rebar stresses exceed 0.90 fy. A gradient with an operating temperature of 120°F inside the containment and a 50°F temperature at the mat-rock interface was considered and stresses were negligible. Ambient accident temperatures have no appreciable effect on the base slab. The maximum operating temperature of the containment is 130°F. The effect of elevated operating temperature on the structural elements was evaluated in 1987 and was found acceptable. It is not possible to show that the design on nonyielding rock is more conservative than assuming the rock to be elastic. However, due to the installation of temperature reinforcing, the design is conservative. Reinforcing and concrete stresses are very low when considering the rock to be elastic. To substantiate the above statement, the following studies were performed:

1. The foundation modulus were determined using the expression:15 4Gr kZ = o 1

where: kZ = The vertical spring constant of a circular base supported on an elastic foundation E G= 2 (1 + ) rO = Radius of Foundation

                 = Poisson's Ratio To obtain the foundation modulus, kZ is divided by the area of the circular base to yield Chapter 5, Page 23 of 89 Revision 22, 2010

IP2 FSAR UPDATE k 4G k = z x o A ro (1 ) Substituting for G 2E ko = r o (1 - 2 )

2. The first case examined was that of a rectangular strip loaded with 1.5 times design accident pressure plus dead load using conservative properties for the Dolomitic limestone:7,14 E = 6.0 x 106 psi

        =0 Applying these values ko = 4370 lbs/in.3 The "characteristic"  is defined as:6 k     1/4
        =      4 EI Where:

E is the modulus of elasticity of the structural base (concrete), I is the moment of inertia of the structural base, k = ko b , (b = width of base) using base properties

        = 7.56 x 10-3-in.-1 Where l> beams may be considered as infinite in length.6 Taking the length of beam as being the base diameter l = 13.1 >

The beam was then analyzed as a beam of unlimited length loaded over an area equal to the base diameter with an 80 psi uniform load. The solution to this problem gives Chapter 5, Page 24 of 89 Revision 22, 2010

IP2 FSAR UPDATE q y c

                    =      (2  Da  Db )

2k q M c

                    =         (B a + B b )

2 4 q Qc = (C a C b ) 4 where yC is deflection of point being considered MC is the moment at point being considered QC is shear at point being considered q is the uniform load a is the distance from point under consideration to end of load b is distance from point under consideration to other end of load. Bx = e-x sin x Cx = e-x(cos x - sin x) D = e-x cos x Maximum moment occurs at mid-point of load and is equal to 352-in.-lbs/in. For the area of the mat where there is only temperature reinforcing, the maximum moment would cause a stress of 30 psi in the reinforcing. The maximum shear would occur at the ends and is equal to 2.64 kips/in. This shear would cause a shear stress in an unreinforced concrete section of 26.4 psi.

3. A second case examined was for the foundation material being less rigid than the concrete base. The model was the same for the first case:

Assumed Erock = 2.6 x 106 psi

               =0 For this case, the following were determined:

ko = 1890 lb/in.3

                = 6.2 x 10-3-in.-1 Chapter 5, Page 25 of 89 Revision 22, 2010

IP2 FSAR UPDATE Mmax = 3.66-in.-kips/in. Qmax = 3.23 kips/in. Srebar = 312 psi conc = 32.3 psi As a final study, the maximum deflection as calculated in the first case was imposed as a settlement of the base mat for the outer portion and a section of the mat was analyzed for this settlement. A 30-ft section was used with fixity at the reactor pit, the remainder cantilevered from the pit. The resulting moment and shear are as follows: M = 142-in.-kips/in. q = 396 lbs resulting in a rebar stress of 12.2 ksi and a shear stress of 4.0 psi. From the above, it can be seen that the assumption that a foundation on rock is a rigid unyielding foundation is a valid assumption and that temperature reinforcing provides much greater resistance than required to accommodate the effects of any elastic deformation of the subgrade. 5.1.3.6 Analysis of Liner and Reinforcing Steel Approximately 67-percent of the inclined bars, provided to resist radial shear at the base of the containment wall, are secondary vertical bars, which are inside the primary vertical bars on the outside face and inside face of the wall. These bars are continuous and are bent across the wall where reinforcing is required to resist the radial shear. The remaining 33-percent of the required steel area is provided by stirrups that are hooked around the vertical bars by means of a 90 degree hook. Only one-third of the shear reinforcing at a particular elevation is made up of these hooked bars, which occur at four elevations up the wall. See Figure 4.16 of the Containment Design Report. Since the stud anchors are hooked around reinforcing bars, concrete stresses for pull out loads are negligible. For high shear loads, which would be caused if a stud anchor should fail or be missing, local crushing of the concrete occurs; however, integrity of the anchor and liner plate is not impaired. See Figures 5.1-21 and 5.1-22. The lowest elevation at which these hooked bars are used is at a point where only 65-percent of the maximum shear at the base is present. The remaining three levels are in regions where the shear is less than 25-percent of maximum base shear. Since the large majority of the shear is resisted by continuous vertical bars, a minimal amount of load must be transmitted to the vertical bars. The hooked stirrups will mechanically transmit the small amount of shear, which they carry. The main function of the stirrups is to contain the formation of the diagonal tension crack. The mechanical anchorage of the stirrups is sufficient for this purpose. Chapter 5, Page 26 of 89 Revision 22, 2010

IP2 FSAR UPDATE There are no significant structural loadings, which must be transferred through the liner such as those required for crane brackets or machinery equipment mounts. Miscellaneous spray system piping, instrumentation, conduit, and insulation, which are attached to the liner can be supported by the free-standing liner without inducing significant stresses in the liner or liner anchorage. Liner stress is imposed on the cylindrical penetration as a circular uniform load acting around the circumference of the penetration. The liner plate is locally thickened at the penetrations to take care of additional stresses. The liner can accommodate any shear it will see due to thermal expansion or earthquake. An investigation was made on the thermal effects, based on the conservative assumptions that the base mat was fully fixed against any thermal movement thereby restraining the liner from movement. The 3-ft fill slab was then subjected to thermal growth. No excessive forces were introduced into the liner and the welds on the test channels were found to be sufficient to prevent any shear failure of the test channels from the liner due to movement of the 3-ft fill mat. Seismic shear of the interior concrete is resisted by the keying action of the reactor pit and the sump for the recirculation pumps in addition to the weld channels. Considerable resistance is also provided by friction between the liner and the 3-ft slab. Jet forces cannot remove the liner panels since the forces will be compressing the insulation panels against the liner and exterior wall. The panels are anchored to the liner with 3/16-in. diameter stainless steel studs. The consequence of an insulation panel being displaced from the liner during or as a consequence of an accident is that the exposed liner would tend to expand. The unequal strain between the exposed and unexposed portions of the liner causes a shear load on the liner anchor, and a local yielding in compression of the exposed portion of the liner. The liner anchor stud has the capacity to accommodate much greater strains than would be experienced at yield strain in the liner. 5.1.3.7 Containment Interior Structure The interior structure may be separated into five main structural components. They are:

1. 3-ft thick fill slab.

2. 3-ft thick crane wall.

3. 4-ft to 6-ft thick refueling canal.

4. 2-ft thick operating floor slab.

5. Primary shield wall.

The method of design, stress analysis, critical stresses and locations are as follows:

1. 3-ft thick fill slab - The controlling loads on the 3-ft slab are the reactions are from the primary equipment supports due to various postulated pipe breaks. The slab was designed as a series of radial beams running under the equipment supports and spanning between the reactor support wall and the crane wall. Stresses in reinforcing were limited to 0.9 fy. Maximum stresses occur immediately below the primary equipment supports.

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IP2 FSAR UPDATE

2. 3-ft thick crane wall - The crane wall was designed for a 7 psi differential pressure occurring immediately after a primary pipe break and prior to pressure equalization.

Although the stress levels associated with this pressure differential were sufficiently low to establish that the concrete could resist the pressure loading, sufficient reinforcing was provided to resist all membrane forces without any contribution from the concrete. Stresses were limited to 0.9 fy. The membrane hoop stress was 33 ksi and the axial vertical rebar stress was 14.3 ksi. A two dimensional finite element analysis was performed to determine the effect of the jet forces associated with the pipe break on the crane wall. The jet force associated with a pipe break has been based on the static force PA where P is the primary system operating pressure and A is the cross sectional area of the coolant pipe. The analysis indicated that in local areas (at the application of the force) yielding of the crane wall rebar will occur. The load was assumed to act at the mid-height of the wall, thus causing maximum bending moment. The ability of the wall to support the dead load of the crane was checked, considering the yielded area indicated by the computer analysis as unable to carry load. A beam 12-ft long and 5-ft deep (the underside of the operating floor to the top of the potential yield portion of the crane wall) was found to provide more than twice the ultimate capacity required. This analysis was very conservative for three reasons:

1. A jet force load at this location would cause little yielding since it is not located at mid span.

2. The haunch at the underside of the operating floor was not considered.

3. The membrane effect of the circular crane wall was not taken into account.

Further stability of the crane wall was demonstrated by determining the ultimate failure load by means of a yield line analysis. This analysis indicated that the structure has the capacity, through strain energy of structural response, to resist the uniform jet force load of 1500 kips or 975 kips with the 7 psi pressure differential without failure. The containment internal concrete is essentially rigid; (fundamental frequency 18.6 cps) therefore, seismic loads were calculated using the maximum ground acceleration (0.15g). The crane wall was initially considered as a cantilever beam with a frequency of approximately 13 cps and the base shear was determined by the response spectrum approach. The base shear was distributed to the individual nodes by the formula: W h V Fx = x x Wh Where V = base shear Wx = weight of node under consideration hx = distance from base to section under consideration. 3 Wh = Summation of the product of weights and heights of all nodes Chapter 5, Page 28 of 89 Revision 22, 2010

IP2 FSAR UPDATE The moment at the base was determined and the uplift calculated by considering a circular ring of thickness equal to the area of steel per in. This maximum uplift, which occurs at one point at the base of the structure stresses the rebar to 5.2 ksi. The crane wall was also designed to resist steam and feed water pipe break reactions of 340 kips and 200 kips where supports are connected to the wall. The extra steel provided for pipe break loads is available in the form of steel buttresses to resist pressure, jet force, and seismic loads; however, it was not considered in the analysis.

3. 4-ft to 6-ft thick refueling canal - The refueling canal was designed for the 7 psi pressure differential. The wall resists the pressure by spanning vertically between the refueling floor and the operating floor. Stresses were limited to 0.9 fy.

An analysis was performed to check the effects of the jet force load the cross section was found to be sufficient to provide stability. A yield line analysis was performed and provided the basis for the above. The seismic load was determined by the same procedure used for the crane wall. The average load in kips/ft was distributed over the wall and the vertical span was conservatively assumed to carry the entire load. The resulting bending moment was found to be well within the capacity of the wall.

4. 2-ft thick operating floor slab - Because of the many openings in the floor for equipment, the floor was designed as a series of beams. Principal loadings were D.L. + 500 psf live load and 7 psi upward pressure differential + D.L. The first loading (D.L. + 500 psf live load) was designed in accordance with Part IV-B of ACI 318. Stresses for the pressure differential case were limited to 0.9 fy.

The operating floor was investigated. There appears to be very little area of the operating floor, which could be reached by the expanding jet of water from a break in the reactor coolant system. The jet will be greatly dispersed in the distance between the primary coolant piping and the underside of the operating floor. The only area of the floor, which could be struck by a jet spans between areas of the floor heavily reinforced as beams. The span cross section consists of a T-beam with the 2-ft thick floor acting as the flange and the 7-ft high biological shielding wall as the web. This section can resist the jet force load within 0.9 fy stress limit on the rebar.

5. Primary Shield Wall - This was designed for two loading conditions due to a split in the reactor. The stress in the reinforcing was limited to the tensile strength of the bars. The first load considered was a 1-ft wide longitudinal split along the length of the reactor. The vessel is assumed accelerated through a 6-in.

distance against the support wall by the jet force caused by a 2200 psi pressure acting through a 26.4-ft long by 1-ft wide longitudinal vessel rupture, which results in an impact load of 650 k/ft. This load is imposed by considering an impact factor of two. The maximum rebar stress is 69.5 ksi. The second load considered a pressure buildup of 1000 psi inside the pit due to release of reactor contents. This produces a rebar stress of 86 ksi. The rebar used is ASTM A 432 with specified yield of 60 ksi and ultimate tensile strength of 90 ksi. Chapter 5, Page 29 of 89 Revision 22, 2010

IP2 FSAR UPDATE To protect the containment base liner, an average of 2-ft of concrete above the containment liner plus a 1-in. liner plate embedded on top of the concrete was provided at the bottom of the containment reactor cavity pit. Below the containment liner plate is 4.5-ft of structural concrete poured on rock. Temperature differential conditions as a result of a LOCA are considered to be of such short duration that the effects were not used in the design of interior structures for stress analysis. A sketch of the design conditions is given in Figure 5.1-24. During normal operations, the only significant transient temperature gradients occur during startup. The minimum containment internal temperature is limited to 50°F. The maximum operating containment internal temperature is 130°F. Forced movement of containment air is used to limit the concrete temperature surrounding the reactor vessel. This forced air movement of the containment air as well as normal convection and radiation is expected to limit the concrete temperature differentials in the range of 5°F to 10°F. To demonstrate the large margin available in the concrete crane wall and the primary shield wall, a conservative assumption of a 30°F temperature gradient has been evaluated. The evaluation included the gradient effect through the crane wall, the 6-ft thick portion of the primary shield wall below the reactor coolant pipe nozzle, the 5-ft thick portion of the primary shield wall where the nozzles penetrate the wall, and the 4-ft thick wall above the shield wall. The maximum rebar stress was found to be 4500 psi and occurs in the vertical rebar in the crane wall. The maximum compressive concrete stress was found to be 226 psi and occurs in the hoop direction in the 5-ft portion of the primary shield wall. These stresses are approximately 20-percent of the allowable working stress values and will have no significant effect on the design adequacy of the structures analyzed. 5.1.3.8 Pressure Stresses 5.1.3.8.1 Accident Pressure Pressure effects on the containment structure may be divided into two types: (1) membrane stresses and (2) discontinuity stresses.

1. For membrane stress analysis, the dome and cylinder are treated as thin-walled shell structures. (The thickness to radius ratio for the dome is 1/20 and the cylinder 1/15. These ratios are smaller than the 1/10 criterion for thin-walled shell analysis.8 Membrane forces are resisted by steel reinforcing.

2. Discontinuity stresses occur at the juncture of the cylinder and the mat and the juncture of the cylinder and dome. Discontinuity effects are determined as follows:

a. The radial growth of the shell is computed based on membrane stress in the reinforcing and liner.

Chapter 5, Page 30 of 89 Revision 22, 2010

IP2 FSAR UPDATE

b. The flexural rigidity of the meridional wall section is determined based on a cracked section analysis in accordance with conventional reinforced concrete design techniques.

c. Moments and shears are calculated based on having consistent deformation for the two elements at the point of discontinuity.

Discontinuity effects at the spring line are very slight due to the small difference in radial growth between the dome and cylinder. Since the circumferential reinforcing in the dome and cylinder vary, stresses and therefore deformations are essentially equal. The mat is considered as offering complete fixity; no credit is taken for the liner at the base in resisting moments since at the point of maximum shear the bond between the liner and concrete is insufficient to transmit complementary beam shear. A slip surface between the concrete and liner is formed and the liner is subjected to membrane forces only. The 9-ft thick mat is subjected to the following due to pressure inside the containment building:

1. Uplift at the juncture with the wall.

2. Moment and shear due to discontinuity effects with the wall.

3. Downward pressure loading due to internal pressure.

The 9-ft mat is designed to accommodate the flexural effects of these loads. At the crane wall, the mat is founded on the unyielding rock and further pressure loads are transmitted through bearing directly into the rock. Resistance to these loads is based on a cracked concrete section. No credit is taken for the liner for the same reasons given for the wall. Discontinuity shears in both the cylinder and mat are resisted by either bent bars or stirrups. In the outer portions of the base mat, the slab is raised off of the rigid foundation under accident loadings; thus no frictional resistance can be offered by the rigid foundation. Where the uplift is overcome, the only load of any consequence, which must be resisted by the mat is the radial tension. The restraint, which is imposed by the rigid foundation on the bottom portion of the base mat, effectively eliminates all radial tension in the mat. However, for conservatism this restraint has been neglected in the analysis of the mat for radial tension. The hoop and radial reinforcing supplied as temperature reinforcing is more than adequate for this purpose. 5.1.3.8.2 Soil Pressure Portions of the containment structure are subjected to the effects of backfill bearing against the containment wall. The effects on the structure are:

1. Shear and overturning effects due to seismic response and interaction between the soil and structure.

2. Discontinuity effects caused by the soil restraining deformation of the structure under accident pressures.

Chapter 5, Page 31 of 89 Revision 22, 2010

IP2 FSAR UPDATE To determine the shear and overturning effects two limiting cases were investigated. The first was the case where the structure and soil move out of phase. It was assumed that the structure was subjected to the passive pressure of the soil with the mass of soil, within the shear failure envelope, accelerated against the structure with ground acceleration. In the second case the soil and structure move in phase. For this case it was assumed that the structure was subjected to the active pressure of the soil with the mass of soil, within the shear failure envelope, accelerated with the structure at ground acceleration. These loads were then treated as external loads on the structure. See Section 3.1.5 of the Containment Design Report for additional information. To determine the discontinuity effects caused by soil restraint, the structure was analyzed for the passive pressure case. The restraint of the deformation of the structure due to the soil was calculated. Vertical and circumferential bending moments due to this restraint were then determined. Reinforcing bar stresses were calculated and found to be minor. This analysis was then verified by a finite element analysis. In this analysis, full contribution of the backfill was assumed. During the course of construction it became necessary to build a retaining wall in a substantial area of the backfill, to facilitate construction. The retaining wall extends over 50-ft in plan and includes all of the high fill points assumed in the analysis and design. It can therefore be concluded that the analysis was conservative in that the backfill effects on the completed structure would be only a fraction of that assumed in the original design. 5.1.3.9 Thermal Stresses Temperature effects on the containment structure may be divided into two separate considerations: one effect is due to a thermal gradient through the wall, the other is caused by the rapid temperature rise of the liner under accident conditions. The reinforced-concrete wall restrains the liner from growing, resulting in compression in the liner and additional tension in the reinforcing.

1. Calculation of gradient stresses is based on method of analysis outlined in ACI 505-54, "Specification for the Design and Construction of Reinforced Concrete Chimneys."9 The gradient used is linear with 120°F on the inside and 0°F exterior concrete temperature (-5°F ambient). The maximum operating temperature of the containment is 130°F. The effect of elevated operating temperature (up to 150°F) on the structural elements was evaluated in 1987 and was found to be acceptable.

The ACI method assumes a cracked section in which the concrete carries no tension. The neutral surface (surface at which no thermal stress exists) is determined. Stresses in the liner and reinforcing are calculated based on the assumption that there is no distortion of the wall; i.e., variation of strain through the wall thickness is linear.

2. To determine the effects due to rapid rise in liner temperature, there are two basic assumptions made. The first is that the effects are internal in nature; i.e.,

the compressive force in the liner is balanced by a tensile force in the reinforcing. The second is that there is no distortion of the wall. Chapter 5, Page 32 of 89 Revision 22, 2010

IP2 FSAR UPDATE Because temperature effects are internal in nature and do not affect the overall tensile load carrying capability of the structure, local yielding of reinforcing under accident conditions is acceptable. The temperature gradient through the wall is essentially linear on both the insulated and uninsulated portions and is a function of the operating temperature internally and the average ambient temperature externally. Accident temperatures mainly affect the liner, rather than the concrete and reinforcing bars, due to the insulating properties of the concrete. By the time the temperature of the concrete adjacent to the liner begins to rise significantly, the internal pressure and temperature in the containment shell due to maximum thermal gradient will not influence the capacity of the structure to resist the other forces. Temperature effects induce stresses in the structure, which are internal in nature; tension outside and compression in the inside of the shell such that the resultant force is zero. Loading combinations concurrent with these temperature effects may cause local stresses in the outside horizontal and vertical bars to reach yield; however, as local yielding is reached, any further load is transferred to the unyielded elements. At the full yield condition, the magnitude of final load resisted across a horizontal and vertical section remains identical to that which would be carried if the temperature effects were not considered. Thus, the overall carrying capacity of the structure and the factor of safety of the structural elements are not affected. 5.1.3.10 Analysis of Openings The methods followed in design of large openings are described in Section 3.4 of the Containment Design Report (CDR). Included are descriptions of the safety factors used in design. Sample calculations are provided, listing all the criteria and analyzing the effects of all pertinent factors, such as cracking. Also addressed in the CDR is how the existence of biaxial tension in concrete (cracking) has been taken care of in the design, and how the normal and shear stresses due to axial load, two-directional bending, two-directional shear, and torsion are combined. Additionally, the criteria for the design of the thickened part of the wall around the openings is stated. The methods used to check the design of the thickened stiff part of the shell around large openings and its effect on the shell, torsional stresses, and shrinkage considerations are also addressed in Section 3.4 of the Containment Design Report. This section also describes how deformations and forces are handled around the large openings and in the transition zones into the main portion of the structure. In the cylindrical section of the containment, where there are large openings for access hatchways and penetrations, the reinforcing bars (hoop, vertical and diagonal) are continued without interruption around the openings. No bar terminates at any openings as illustrated around the penetration in Figure 5.1-1. Also additional bars have been furnished locally to take the stresses developed around large openings. Concrete is locally thickened at the equipment access hatchway area to accommodate all the reinforcing bars required in this area. A finite element analysis is performed on the large openings. Representation of the structure is by rectangular elements; each element consists of ten layers of orthotropic, elastic material to Chapter 5, Page 33 of 89 Revision 22, 2010

IP2 FSAR UPDATE represent the reinforcement, concrete and the liner. About 1000 degrees of freedom are considered in the model. This analysis is used as a check on the adequacy of the large openings. Results appear in the Containment Design Report. A finite element analysis of the equipment hatch area indicated local liner plastic deformations during the pressure test. For the order of magnitude and location of these stresses, see Section 3.4 of the Containment Design Report. These deformations have no influence on the structure during the pressure test due to the ductility of the studs and liner plate. The limiting elastic liner deformations during test pressure will be from tensile stresses. During an accident loading they will be from compressive stresses. Therefore, a relationship between the pressure and accident loads cannot be determined directly. However, the test pressure demonstrates the ductile behavior of the liner. Since the containment is not subject to accident temperatures during the testing, no direct correlation between test and accident conditions can be made in evaluating thermal stresses at large openings. The liner is stressed beyond the yield point in very local areas adjacent to the transition from the thickened equipment hatch boss to the cylinder wall. The maximum stress is equal to 39.28 ksi for the 1.5P loading condition. The strain corresponding to this stress (0.17-percent) is below the limits (0.5-percent) stated in Section 2.2.4 of the Containment Design Report. The average liner stress in the cylinder for the 1.5P load combination is approximately -15 ksi in the vertical direction and -2.0 ksi in the horizontal direction. The maximum rebar stress associated with the 1.5P load combination is approximately 66 ksi in the 4'-6" portion of the containment wall cylinder. For a complete discussion of liner stresses, see the Containment Liner Stress Analysis Report. For a detailed discussion of liner stresses in the equipment hatch area and further justification of the stresses noted above, see Section 3.4.4 of the Containment Design Report. All reinforcing is continuous around penetrations. Steps have been taken to ensure that no local crushing of concrete will occur. From Reference 16, it has been determined that in order to prevent local crushing of the concrete, a minimum bend diameter of 31 times the bar diameter is required when the reinforcing is stressed to yield. The angle of bend in the rebar determines the force that will be transmitted to the concrete in the event the bar tries to straighten out due to tension. For this reason most bars are bent at 10 degrees except at large penetrations including the equipment hatch, personnel lock, main steam and feedwater, and air purge penetrations, where the deviation of the bar from its centerline is too large to permit a 10 degree bend. In these cases the bars have been bent at 30 degrees but a tie-back system is used, which prevents a buildup of forces. To prevent this buildup, (in all cases except the equipment hatch penetration), the line of force makes an angle of one-half of the angle of bend, from a horizontal line from the vertical bars and from a vertical line for the horizontal bars and is tangent to the outside of the penetration. At the personnel and equipment hatches a large void will be carried since, due to the large offset of the bars from their centerline, it will take the bars longer to return to their centerline after passing the penetration. To prevent any cracking and spalling of concrete and to add lost strength to the cross-section, these voids have been filled with added rebar, which achieves bond by means of mechanical anchorage. Chapter 5, Page 34 of 89 Revision 22, 2010

IP2 FSAR UPDATE The same precautions mentioned above have been taken with the seismic bars. See Figure 5.1-25. For penetrations between 9-in. and 18-in. in diameter, all the reinforcing bars including primary and secondary vertical bars and diagonal bars have been grouped around the penetrations. Due to the continuity of the bars and the relatively small opening size, no special provisions need be made to resist normal, shear, and bending stresses. The penetrations are keyed into the concrete, thus creating an edge loading, which will put torsion into the wall. The loads are small and the rebar will feel little effects from this torsional loading. For penetrations greater than 18-in. up to 48-in. in diameter, the bars are continuous. Due to the large angle of bend of these bars, a tie-back system is used, which offers additional resisting strength to shear, bending, and torsional stresses. 5.1.3.11 Seismic and Wind Design The design of the containment, which is a Class I structure (see Section 1.11), is based on a "response spectrum" approach in the analysis of the dynamic loads imparted by earthquake. The seismic design takes into account the acceleration response spectrum curves as developed by G. Housner. Seismic accelerations have been computed as outlined in TID-702410 and Portland Cement Association Publication.11 The following damping factors have been used: Percent Component Critical Damping

1. Containment structure 2.0

2. Concrete support structure of reactor vessel 2.0

3. Steel assemblies:

a. Bolted or riveted 2.5

b. Welded 1.0

4. Vital piping systems 0.5

5. Concrete structures above ground:

a. Shear wall 5.0

b. Rigid frame 5.0 As indicated in Section 5.1.2.2, ground accelerations used for design purposes are 0.1g applied horizontally and 0.05g applied vertically. The natural period of vibration is computed by the Rayleigh method; in this method, the containment structure is analyzed as a simple cantilever intimately associated with the rock base and with broad base sections of adequate strength to assure full and continued elastic response during seismic motions. Further, both bending and shear deformations are considered.

The structure is divided into sections of equal length and loaded laterally by dead weight of the section and any equipment and live load occurring at the section. Deflections caused by shear and moments are then determined, and the end deflection is given the value ' = 1.0 with corresponding values determined for other sections. The natural period of vibration for the structure is then determined by setting potential energy equal to kinetic energy and solving for the period. Chapter 5, Page 35 of 89 Revision 22, 2010

IP2 FSAR UPDATE 2 T = 2 Y0 dm 1/2 g dm where Y0 = maximum actual deflection deflection of section under consideration

                =

maximum actual deflection g = acceleration due to gravity dm = weight of section under consideration T = period in sec. Based on an uncracked concrete section, the period is determined to be 0.241 sec. A more realistic calculation for a cracked section, using reinforcing steel and liner as the resisting elements, yields a period 0.936 sec. Using the derived period and entering the acceleration spectral curves, Figures 1.11-1 and 1.11-2 of Section 1.11, and applying a 2-percent critical damping, a spectral acceleration for the containment was selected. This value was derived to determine the base shear. The distribution of base shear is a triangular loading assumption. This assumption yields a load distribution pattern with zero loading at the base to a maximum loading at the spring line of the dome. Above this line, the loading decreases due to a change in section and consequently change in weight. This load distribution allows the determination of shears and moments at any critical section through the containment from which the appropriate unit stresses are obtained. Seismic shears are resisted by diagonal reinforcing except in the upper areas of the dome. No credit is taken for the reinforcing in compression. From 30 degrees above the springline, where the seismic shears are small, the shears are carried by dome reinforcing steel lying in the plane of principal tension A finite element analysis was performed on the basemat using loads determined for the three basic loading conditions specified in the Containment Design Report. Maximum hoop moment caused by lack of symmetry of the seismic loading was found to be 454 in.kips/in. This compares with a capacity of 690-in.-kips/in. for the in-place hoop reinforcing. Tornado loads have not been considered in the design of the Indian Point Unit 2 Containment Building; however, similarity in design of Indian Point Unit 3 (where such loads are considered) indicates that seismic reinforcement bars provide a more than adequate mechanism to withstand the torsional effect of Tornado loads. Chapter 5, Page 36 of 89 Revision 22, 2010

IP2 FSAR UPDATE The torsional effect results from wind striking the containment building at an angle from the normal, as shown in Figure 5.1-26. The torsional force is due to the component of the wind tangential to the surface of the containment building and is equal to: Ft = ACD (q) (sin ) Where A = surface area of the containment CD = 0.5 from A.S.C.E. Paper 3269 - "Transactions of the A.S.C.E.," Vol. 126 Part II 1961, p. 1165 (coefficient of drag) q = 0.002558 V2 (wind pressure)

                = 45 degrees This assumption is conservative in that the actual tangential force would be the result of skin friction and the effects would be negligible.

This component of torsional force is computed from a direct wind loading as based on A.S.C.E. Paper 3269. Torsional shear is a maximum at the juncture of the walls and base slab and varies to zero at the top of the dome. The torsional effect can be converted to a shear per lineal foot around the circumference of the containment by distributing the shear over the circumference of the seismic reinforcing. The seismic bars provide a more than adequate mechanism to withstand this torsional effect. The maximum stress in the bars under this loading is 17 ksi. See Figure 5.1-26. 5.1.3.12 Cathodic Protection During the initial Licensing process, a complete survey and tests to determine the need for cathodic protection on Indian Point Unit 2 was made by the A. V. Smith Engineering Company of Narberth, Pennsylvania. Electrical resistivity measurements and a visual inspection of the area away from the river, where the turbine generator building, reactor building, primary auxiliary building and associated facilities are located indicated that the environment is mostly rock with areas of dry sandy clay. The electrical resistivity of the soil ranged from 3,500 to 30,000 ohm-cm with the majority of the readings being above 10,000 ohm-cm. On this basis, it was determined that cathodic protection was not required on underground facilities in areas away from the river or the containment building liner, although a protective coating on pipes was recommended to eliminate any random localized corrosion attack. An analysis of Hudson River water data, obtained from the Con Edison plant chemist, showed the electrical resistivity of the water to vary over an extremely wide range due to salt intrusion from the ocean. The range of resistivity has been from 59 to 10,000 ohm-cm with a large number reading in the 300 ohmcm area. This value was considered to be extremely corrosive and the following structures in the area near the river were placed under cathodic protection: Chapter 5, Page 37 of 89 Revision 22, 2010

IP2 FSAR UPDATE

1. Circulating water lines.

2. Service water lines.

3. Bearing piles.

4. Sheet piling (earth and water side) and wing wall anchorage system.

5. Metallic structures inside intake structure (traveling screens, bar racks, circulating water pump suction, service water pump suction).

In 2008, a cathodic protection field survey and assessment of underground structures at Indian Point Unit 2 was performed by PCA Engineering of Pompton Lakes, New Jersey. A positive shift in pipe potential was found where the City Water supply piping from the City Water Tank crosses the Algonquin Gas pipes. As a result the City Water supply piping in the vicinity of the gas pipes was placed under cathodic protection. The cathodic protection system for the Circulating Water lines and the Service Water lines were found not to be functional and the rectifiers were removed. In order to assure the lines will perform their functions, the buried pipes are inspected as part of the Buried Piping Program. Inspections of buried piping are initially performed using Guided Wave (GW) ultrasonic inspection techniques to locate potential areas of degradation. If significant degradation is detected during the GW inspections, excavation is performed to uncover the affected sections of piping and a direct visual inspection and UT thickness measurements are performed. Repairs and / or replacements are implemented as required to restore degraded piping sections to within the required structural margins of safety. In addition to the inspections performed as part of the Buried Piping Program, the nuclear safety related portion of the service water piping is further subjected to pressure and / or flow testing as required by ASME XI, Subsection IWA-5244. Visual inspections on the inside surface of the SW piping are also performed under the GL 89-13, Service Water program. Based on the results of the inspections and testing, the Service Water system is structurally adequate to perform its required safety function. The cathodic protection system for the Traveling Water Screens and Bar Racks were found not to be effective and the installed cathodic protective systems were retired. The original Traveling Water Screens which were carbon steel were upgraded to stainless steel frames, baskets, and chains. The splash housings are also stainless steel. The Bar Racks, replaced in the mid-1990s are of carbon steel construction and are epoxy coated with a tar epoxy expected to provide corrosion protection over a 40-year period. The guides for the Screens and Racks are carbon steel channels mounted in a concrete through. The rate of corrosion is slow and the Screens and Racks are on a regular PM cycle that checks for degraded conditions. The Service Water and Circulating Water pumps suction are not cathodically protected. Rather, the Service Water Pumps suction is inspected and refurbished as part of the Service Water Pumps Preventing Maintenance (PM) activities. The Circulating Water Pumps are inspected and refurbished according to Preventive Maintenance program requirements. 5.1.3.13 Containment - Shear Crack The arrangement of reinforcing bars in the containment shell is such that a reinforcing bar crosses any potential crack plane. Any cracks resulting from diagonal tension caused by shearing forces will be carried by reinforcing bars, which span across the crack. Thus all shears will be carried by the reinforcing bars and none by the concrete. Chapter 5, Page 38 of 89 Revision 22, 2010

IP2 FSAR UPDATE The reinforcing bars are almost all continuous throughout the containment structure; however, where a bar terminates this is accomplished by means of a 180 degree hooked bar. In no case are bars simply terminated without providing means for additional anchorage. Throughout the cylinder, the meridional reinforcing is continuous. Beyond the springline, the bars extend radially toward the center of dome. As the bars reach a 6-in. spacing, which is one-half the required spacing, alternate bars have been dropped off by means of reinforcing splice plates. The splice piece consists of a plate with two Cadweld sleeves welded on the incoming side and one sleeve welded on the outgoing side. Thus, the number of bars present is halved and the spacing is increased to the required 12-in. This is repeated to the top of the dome where a three layered grid pattern has been used to maintain the continuity of the rebars. The bars in the grid pattern have been Cadwelded to the same type reinforcing splice plates described above, but the Cadweld is beveled to obtain the desired direction of the grid. At the base in the area of high discontinuity stresses, additional No. 18S bars have been provided. At the point where they were no longer needed, they have been Cadwelded to a No. 11 bar, which is terminated with a 180 degree hook. All seismic bars have been terminated in a 180 degree hook. In no case was a No. 18S bar terminated in this way since the minimum 180 degree hook could not be provided in a 4-ft 6-in. thick wall. Radial shear reinforcing stirrups were terminated by hooking around vertical bars. 5.1.4 Containment Penetrations 5.1.4.1 General In general, a penetration consists of a sleeve embedded in the concrete and welded to the containment liner. The weld to the liner is shrouded by a continuously pressurized channel, which is used to demonstrate the integrity of the penetration-to-liner weld joint. The pipe, electrical conductor cartridge, duct or equipment access hatch passes through the embedded sleeve and the ends of the resulting annulus are closed off, either by welded end plates, bolted flanges or a combination of these. (See Figures 5.1-27 through 5.1-31.) Differential expansion between a sleeve and one or more hot pipes passing through it is accommodated by using a bellows type expansion joint between the outer end of the sleeve and the outer end plate, as shown on Figure 5.1-30. Pressurizing connections are provided to continuously demonstrate the integrity of the penetration assemblies. 5.1.4.2 Types of Penetrations 5.1.4.2.1 Electrical Penetrations The electrical penetration system consists of 60 electrical penetrations including the following: 48 Crouse-Hinds, 1 Westinghouse, 10 Conax and 1 spare sleeve (below flood-up level.) The Crouse-Hinds and Westinghouse types are identical in design (see Fig 5.1-27). This is because Westinghouse took over the Crouse-Hinds manufacturing facility and design after the Chapter 5, Page 39 of 89 Revision 22, 2010

IP2 FSAR UPDATE original plant penetrations were purchased. The design of this type of electrical penetration utilizes a single canister that is sealed at both ends by a combination of metal and ceramic seals. Epoxy layers on both ends provide a physical support for the conductors within the penetration canister. All of the Westinghouse and Crouse-Hinds penetrations are welded to the sleeve inside containment. The entire canister assembly is constantly pressurized by the weld channel pressurization system and monitored for any leakage. The Conax penetrations are of a modular design consisting of a stainless steel header and 18 independently mounted conductor feedthrough modules (Figure 5.1-28 and 5.1-29), which can be individually removed and relocated. The header plate and the individual feedthrough modules are the pressure-retaining boundary. This type penetration does not have a sealed canister. The conductor modules are threaded into the header plate and the header plate itself is welded to the sleeve, which goes though the containment wall. Leakage monitoring of the Conax penetrations is accomplished by interconnecting ports machined in the header plate to each conductor feedthrough module. A small hole is provided on each conductor feedthrough module stainless steel tubular housing allowing the feedthrough module to be pressurized when the header plate's parts are pressurized. Metal compression fittings (swaging type) are used for mounting the conductor feedthrough modules to the header in a double seal manner. The individual conductors passing through the feedthrough module are surrounded by polysulfone and are sealed (swaged) at each end of the feedthrough housing. The length of the housing (feedthrough tube) is roughly 2-ft longer than the sleeve within, which the penetration is installed. Six of the Conax penetrations are welded to their sleeve outside containment and four are welded to their sleeve inside containment to accommodate differences in the sleeves into which they are welded. Weld channel rings are used to create a double weld seal between the header plate and the containment sleeve. All of the weld joints necessary maintain containment integrity are monitored for leaks with the weld channel pressurization system. If a minor leak should develop at any of the plants electrical penetrations, a release from inside containment to outside should not occur since each penetration is double sealed and pressurized to maintain a positive pressure (between 49 and 55 psi), which is higher than anticipated containment accident pressures. 5.1.4.2.2 Piping Penetrations Double barrier piping penetrations are provided for all piping passing through the containment. The pipe is centered in the embedded sleeve, which is welded to the liner. End plates are welded to the pipe at both ends of the sleeve. Several pipes may pass through the same embedded sleeve to minimize the number of penetrations required. In this case, each pipe is welded to both end plates. A connection to the penetration sleeve is provided to allow continuous pressurization of the compartment formed between the piping and the embedded sleeve. . These penetrations are listed as Hot in Table 5.2-1. In the case of piping carrying hot fluid, the pipe is insulated and cooling is provided to reduce the concrete temperature adjoining the embedded sleeve. Local areas are allowed to have increased temperatures not to exceed 250ºF. Cooling is provided for hot penetrations through the use of air-to-air heat exchangers. These are made in accordance with the ASME Boiler and Pressure Vessel Code, Section VIII, by welding together one flat sheet and one embossed sheet of 10 gauge carbon steel material, the embossment forming coolant passages. The unit is rolled into the form of a cylinder with an outside diameter slightly smaller than the respective inside diameter of the penetration sleeve. The exchanger is placed inside the sleeve and outside the pipe insulation, Chapter 5, Page 40 of 89 Revision 22, 2010

IP2 FSAR UPDATE with the inlet and outlet coolant connections penetrating the sleeve between the outside concrete wall surface and the bellows expansion joint. The coolant to be used is ambient air fed by a rotary blower, which is backed up with a full sized spare. The isolation features and criteria for piping penetrations are given in Chapter 6. Figure 5.1-30 shows typical hot and cold pipe penetrations. A total of 107 pipes pass through 53 penetration sleeves, 23 of which are considered thermally hot. In addition, two spare penetration sleeves (capped and pressurized) are available for the possible future addition of piping. All piping penetrations are designed for normal loads within the stress limits of the ASME Code, Section VIII. All piping penetrations except main steam and feedwater are designed as anchors for the pipes passing through them and will transmit piping loads to the reinforced concrete wall. The anchorage strength exceeds the maximum combined forces imposed by the effects on the piping penetration of dead load, loads induced from a loss of coolant accident, thermal expansion of the pipe, penetration air pressure, and earthquake loads. The piping penetrations are designed to transmit the above combined loadings to the concrete structure without exceeding the yield strength of penetration steel. In addition, each piping penetration is designed to withstand, within emergency load criteria, the effect of the rupture of a pipe passing through that penetration at or near the penetration. The main steam and feedwater penetrations are designed so that the pipes themselves are effectively enclosed for blowdown just inside and just outside the wall. These anchors are designed to prevent a main steam or feedwater pipe rupture from causing a breach of containment at the penetrations. The anchors are designed to 90-percent of yield strength. All piping penetrating the containment is designed to meet the requirements of USAS B31.1 (1955) Power Piping Code. Pipes that penetrate the containment building wall and that are subject to machinery-originated vibratory loadings, such as from the reactor coolant pumps, have their supports spaced in such a manner that the natural frequency of the piping system immediately adjacent to the penetrations is greater than the dominant frequencies of the pump. Pipeline vibration was checked during preliminary plant operation and where necessary, vibration dampers were fitted. This checking and fitting effectively eliminates vibrating loads as a design consideration. 5.1.4.2.3 Equipment and Personnel Access Hatches An equipment hatch has been provided. It is fabricated from welded steel and furnished with a double-gasketed flange and a bolted, dished door. The hatch barrel is embedded in the containment wall and welded to the liner. Provision is made to continuously pressurize the space between the double gaskets of the door flanges, and the weld seam channels at the liner joint, hatch flanges and dished door. Pressure is relieved from the double gasket spaces prior to opening the joints. The personnel hatch is a double door, mechanically latched, welded steel assembly. A quick-acting type equalizing valve connects the personnel hatch with the interior of the containment vessel for the purposes of equalizing pressure in the two systems when entering or leaving the containment. Two spring-loaded check valves in series are installed to Chapter 5, Page 41 of 89 Revision 22, 2010

IP2 FSAR UPDATE allow pressure relief inside the air locks to the containment interior. The 16-ft diameter equipment hatch opening and the 8-ft 6-in. diameter personnel hatch are the only openings, which require special design consideration. The personnel hatch doors are interlocked to prevent both being opened simultaneously and to ensure that one door is completely closed before the opposite door can be opened. Remote indicating lights and annunciators situated in the control room indicate the door operational status. An emergency lighting and communications system operating from an external emergency supply is provided in the lock interior. Emergency access to either the inner door, from the containment interior, or the outer door, from outside, is possible by the use special door unlatching tools. The design is in accordance with Section VIII of the ASME Code. [Deleted] The design basis Fuel Handling Accident (FHA) analysis does not credit accident mitigation via Containment isolation subsequent to a postulated fuel assembly drop. However, Containment closure after a FHA event is an option that can reduce total exposure and is a good ALARA practice. The roll-up door serves as a mechanism that will support rapid closure of Containment in the event a radiation release occurs during fuel handling. Containment closure subsequent to a total loss of Residual Heat Removal (RHR) cooling in MODE 6 is accomplished through installation of the equipment hatch or temporary closure plate. The roll-up door is currently not considered a suitable device that can be credited for Containment closure subsequent to a loss of RHR, pending NRC approval of an outstanding license Amendment request. 5.1.4.2.4 Special Penetrations

1. Fuel Transfer Penetration - A fuel transfer penetration is provided for fuel movement between the refueling transfer canal in the reactor containment and the spent fuel pit. The penetration consists of a 20-in. stainless steel pipe installed inside a 24-in. pipe. The inner pipe acts as the transfer tube and is fitted with a pressurized double-gasketed blind flange in the refueling canal and a standard gate valve in the spent fuel pit. This arrangement prevents leakage through the transfer tube in the event of an accident. The outer pipe is welded to the containment liner and provision is made by use of a special seal ring for pressurizing all welds essential to the integrity of the penetration during plant operation. Bellows expansion joints are provided on the pipes to compensate for any differential movement between the two pipes or other structures. Figure 5.1-31 shows a sketch of the fuel transfer tube.

2. Containment Supply and Exhaust Purge Ducts - The ventilation system purge ducts are each equipped with two quick-acting, tight-sealing valves (one inside and one outside of the containment) to be used for isolation purposes. The valves are manually opened for containment purging, but are automatically closed upon receipt of a safety injection signal or high-containment radiation signal. The space between the valves is pressurized above design pressure while the valves are normally closed during plant operation. See Section 5.3, Containment Ventilation System, and Section 6.4, Containment Air Recirculation Cooling System.

Seismic Class I debris screens inside the primary containment protect the primary containment isolation valves in the containment purge and pressure relief Chapter 5, Page 42 of 89 Revision 22, 2010

IP2 FSAR UPDATE exhaust ducts from debris that may inhibit their correct operation. The screens are stainless-steel wire mesh and are mounted over the exhaust ducts. Two solenoid-controlled, pneumatically operated butterfly valves are provided for each purge penetration, one on each side of the containment building wall. Two penetrations, one supply and one exhaust, are required. Valves are spring-loaded to fail closed. The space between the valves is pressurized from the pressurization system through an electrically operated three-way solenoid valve. This pressure is maintained only when valves are closed and must be relieved before butterfly valves can be opened. Failure to release this pressure will prevent the inside containment valves from opening. By procedure the outside containment valves are opened after the inside containment valves are open. Failure of any of the valves to open will prevent the containment building purge supply fan from running. Tripping of the containment building purge supply fan will automatically close the inside containment butterfly valves. By procedure the outside containment butterfly valves must then be closed. When these valves are closed the space between the valves is automatically pressurized. Failure of any valves to close will prevent the adjacent space from being pressurized and will sound the loss of pressurization alarm. Loss of pressure for either zone will be displayed by individual indicating lights at the main control board. The valve control solenoids for the inside containment isolation valves FCV-1170 and FCV-1172 and pressurization solenoids are controlled from a single control switch on the fan room control panel. The valve control solenoids for outside containment isolation valves FCV-1171 and FCV-1173 are controlled from a switch in the control room. The cycle is initiated by setting the fan room control switch to the "open" position. This will energize the pressurization alarm. When the pressure between the valves has been relieved, the valve control solenoids for the inside containment isolation valves are energized and these two valves are opened. If for any reason, either of the two inside containment isolation valves fail to open within a given time after the cycle is initiated, both of these valves will close and pressure will be restored. The circuit is interlocked to prevent inadvertent opening of the valves during a safety injection condition. Once the inside containment purge valves have been opened, the operator has a predetermined time to place the control switch for the outside containment purge valves to the "open" position and once opened to start the purge supply fan. Failure to do so will cause the inside containment purge valves to close. Position indicating lights for each of the four valves are provided on the fan room control panel and the main control board.

3. Sump Penetrations - The piping penetration in the containment sump area is not of the typical sleeve-to-liner design. In this case, the pipe is welded directly to the base liner. The weld to the liner is shrouded by a test channel, which is used to demonstrate the integrity of the liner.

Chapter 5, Page 43 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.4.3 Design of Containment Penetrations 5.1.4.3.1 Criteria The liner is basically not a load-carrying member. Because it is subjected to strains imposed by the reinforced concrete, the liner has been reinforced at each penetration in accordance with the ASME Code Section VIII. The weldments of liner to penetration sleeve are of sufficient strength to accommodate stress concentrations and adhere strictly to ASME Code Section VIII requirements for both type and strength. The penetration sleeves and plates are designed to accommodate all loads imposed on them under operating conditions (thermal effects and internal penetrations and test pressures) and accident conditions (loads resulting from all strains, internal pressures, and seismic movements). All reinforcing bars except stirrups and facing bars that are not counted on to carry any load are continuous around the openings. Liner stress is imposed on the cylindrical penetration as a circular uniform load acting around the circumference of the penetration. The liner plate is locally thickened at the penetrations to take care of additional stresses. 5.1.4.3.2 Materials The materials for penetrations, including the personnel and equipment access hatches, together with the mechanical and electrical penetrations, are carbon steel, conforming with the requirements of the ASME Nuclear Vessels Code and exhibiting ductility and welding characteristics compatible with the main liner material. As required by the Nuclear Vessels Code, the penetration materials meet the necessary Charpy V-notch impact values at a temperature 30°F below lowest service metal temperature, which is 50°F within containment and -5°F outside the containment. The stainless steel bellows of the hot penetration expansion joints were protected from damage in transit and during construction by sheet metal covers fastened in place at the fabricator's shop.

1. Piping Penetrations: Materials Piping Penetration Material Specification Penetration Sleeve in. dia. and under ASTM-A333, Gr. 1 Over 12-in. dia. ASTM-A201, Gr. B (see exception below) normalized to A300 CL. 1, Firebox in. dia. containment ASTM-A53, Gr. B sump suction

                                  - Rolled shapes             ASTM-A36, A131, Gr. C

2. Electrical Penetrations: Materials The penetration sleeves to accommodate the electrical penetration assembly cartridges are schedule 80 carbon steel in accordance with ASTM-A333, Gr. 1, except where otherwise noted. The electrical cartridges have been secured to the penetration sleeve so that all possible leak paths between the cartridge and sleeve will be blocked by a pressurized zone.

Chapter 5, Page 44 of 89 Revision 22, 2010

IP2 FSAR UPDATE

3. Access Penetrations: Materials The equipment and personnel access hatch material is as follows:

Item Material Specification Equipment hatch insert ASTM A516, Gr. 60 normalized to ASTM A300, CL. 1, Firebox Equipment hatch flanges ASTM A516, Gr. 60 normalized to ASTM A300, CL. 1, Firebox Equipment hatch head ASTM A516, Gr. 60 normalized to ASTM A300, CL. 1, Firebox Personnel hatch ASTM A516, Gr. 60 normalized to ASTM A300, CL. 1, Firebox 5.1.4.4 Leak Testing of Penetration Assemblies A preoperational proof test was applied to each penetration by pressurizing the necessary areas to 54 psig. This pressure was maintained for a sufficient time to allow soap bubble and Freon sniff tests of all welds and mating surfaces. Any leaks found were repaired and retested; this procedure was repeated until no leaks existed. 5.1.4.5 Construction The qualification of welding procedures and welders has been in accordance with Section IX, "Welding Qualifications," of the ASME Boiler and Pressure Vessel Code. The repair of defective welds has been in accordance with Paragraph UW-38 of Section VIII, "Unfired Pressure Vessels." 5.1.4.6 Testability of Penetrations and Weld Seams All penetrations, the personnel air lock, and the equipment hatches are designed with double seals, which will be normally pressurized at or above the containment design pressure. Individual testing at 115-percent of containment design pressure is also possible. The containment ventilation purge ducts are equipped with double isolation valves and the space between the valves is permanently piped into the penetration pressurization system. The space can be pressurized to 115-percent of design pressure when the isolation valves are closed. The purge valves fail in the closed position upon loss of power (electric or air). All welded joints in the liner have steel channels welded over them on the inside of the vessel. During construction, the channel welds were tested by means of pressurizing sections with Freon gas and checking for leaks by means of a Freon sniffer. These welds were also then continuously pressurized at 50 psig. Chapter 5, Page 45 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.4.7 Accessibility Criteria Limited access to the containment through personnel air locks is possible with the reactor at power or with the primary system at design pressure and temperature at hot shutdown. After shutdown, the containment vessel is purged to reduce the concentration of radioactive gases and airborne particulates. This purge system has been designed to reduce the radioactivity level to doses defined by 10 CFR Part 20 for a 40-hr occupational work week within 2 to 6 hr after plant shutdown. Since negligible fuel defects are expected for this reactor, much less than the 1-percent fuel rod defects used for design and purging of the containment is normally accomplished in less than 2 hr. To ensure removal of particulate matter and radioactive gases, the purge air is passed through a high efficiency and charcoal filters before being released to the atmosphere through the purge vent. The primary reactor shield has been designed so that access to the primary equipment is limited by the activity of the primary system equipment and not the reactor. 5.1.4.8 Penetration Design - Computations The penetration sleeves and end plates are designed to accommodate all loads imposed on them. The sleeve and end plate loads include the effects of internal pressure; concentrated loads imposed by the sleeve anchors on the concrete as the anchors strain in conjunction with wall movement under both operating and accident conditions; thermal effects due to both gradient and thermal reactions of the particular item passing through the sleeve; shear, bending, and compression due to accident end pressures; and shear and bending due to seismic movements of the particular item passing through the penetration. The sleeve and expansion joint are designed to remain within ASME Code Section VIII stress limitations with small strains under all or any combinations of loadings mentioned above. For design computations of penetrations and the shell adjacent to them, see Figures 5.1-32 and 5.1-33. In Section 5.1.4.8.1, the formula for radial deformation of a hole in a plate subjected to biaxial stresses is determined by performing an integration of the tangential strains around the periphery of the hole. In Section 5.1.4.8.2, the relationship between the deflection determined from above to the final plate and penetration sleeve deformations is developed and the formulas for stress in the liner and the stress in the penetration sleeve are developed. Section 5.1.4.8.3 shows a summary of the liner and penetration stresses and states the assumptions made in the analysis. In addition, thermal loads have been investigated for their effect on the shell adjacent to the penetration sleeve and found to be insignificant (38 psi bearing stress on the concrete is the maximum stress on the concrete shell). 5.1.4.8.1 Radial Deformation of a Hole in a Plate From Reference 17, page 81 o = S - 2 S cos 2 + [S'-2S' cos(2-)] where Chapter 5, Page 46 of 89 Revision 22, 2010

IP2 FSAR UPDATE S = Horizontal stress in liner S' = Vertical stress in liner 1 D = (S 2S cos 2 + [S'2S' cos (2 )]) r sin d E r D = S sin d 2S cos 2 sin d + S' sin d 2S' cos (2 ) sin d E cos (2 ) sin d = cos 2 sin d

           =   (1  2 sin  ) (sin )d 2
           =   (sin   2 sin ) d 3
                                                +  sin d) 2 sin  cos  2
           =  ( cos )  2 3        3 4        
           =    cos  + 2 3 sin2  cos  + cos 3

therefore cos(2 ) sin d = cos 3 2

                                               - sin 3

2 cos r cos 2 cos 2

=     S cos   2S              + sin2  cos    S' cos   2S'           sin2  cos E                       3         3                                 3          3 r          2           2              2          2 
=     S + S + S' S'     S  S  S'+ S' E          3           3              3          3 r            4            4 
=

E 2S + 3 S + 2S' 3 S' r

=     [ 10 S + 2 S' ]

E 3 3

 =   2   r    [ 5S + S']                      (for stresses in the same direction) 3E
                     = 2 r [ 5S - S'] (for stresses in the opposite direction) 3E Chapter 5, Page 47 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.4.8.2 Plate and Sleeve Deformation UN= Pl Res. + Sleeve S Sl tpl R2 UN = 1 (1 ) R + E 2Etsleeve t R2 UN = S1 R (1 ) + pl E 2tsleeve UNE S1 = tpl R R (1 ) + 2tsleeve S1 tpl R Ssleeve = 2tsleeve UNE t R p1 Ssleeve = tp1 R R (1 ) + 2t 2tsleeve sleeve where: 1 2

        =                     4 3 (1 -   )

R2t2 sleeve

       *S1 = Stress in Liner tpl = plate thickness, in.

Chapter 5, Page 48 of 89 Revision 22, 2010

IP2 FSAR UPDATE R = radius, in.

         = poisson ratio E = modulus of elasticity 5.1.4.8.3       Summary Penetration                       Stress in Sleeve                   Stress in Liner (ksi)                            (ksi)

Air purge -23.8 -19.5 Main steam -33.4 -27.94 Typical mech. Penetration -31.0 -31.1 Electrical penetration A) C and T1 -22.5 -29.5 B) T and T1 +18.2 +19.7 Fuel transfer A) C and T1 -25.7 -25.6 B) T and T1 +20.8 +16.6 [Note - 1. First letter represents direction of vertical liner stress; second letter represents direction of horizontal liner stress, C signifies compression and T signifies tension.] A) Ignores effects of insulation in the vertical direction. B) Considers effects of insulation. Conservative Assumptions

1. The weld pressurization channel stiffens the area.

2. The liner alone was designed for stress concentration effects while the cracked concrete was ignored.

3. The unrestrained growth is based on maximum growth from a stress concentration consideration.

4. The main steam and mechanical penetrations have been considered in a noninsulated zone when they are just inside the insulated zone. The compression in the hoop direction will be greatly reduced or perhaps go into tension, thus reducing the stresses.

5. The allowable stress in the sleeve is 56,700 psi except for the stainless steel fuel transfer penetration, which is 49,500 psi. These values come from Table N-421 and Figure N-414 of the ASME Nuclear Vessel Code Section III.

5.1.5 Primary System Supports In 1989, the NRC approved changes to the design bases with respect to dynamic effects of postulated primary loop pipe ruptures, as discussed in Section 4.1.2.4. Chapter 5, Page 49 of 89 Revision 22, 2010

IP2 FSAR UPDATE In 2000, an analysis (Reference 19) of the reactor coolant loop and its component supports, which incorporates the NRC approved changes, was performed to reflect the replacement steam generator and removal of sixteen of the original twenty-four steam generator support frame hydraulic snubbers. The analysis also reflected the de-activation of the original horizontal and vertical pipe rupture restraints, located on the cross over legs at the steam generator end. Based on this revised analysis, it was concluded that the Unit 2 reactor coolant system can withstand the combination of blowdown and seismic loads within acceptable stress limits. By reducing the number of snubber and de-activating the rupture restraints the extent of maintenance, inspection and testing requirements is reduced and the reliability of the Reactor Coolant System is enhanced by reducing the possibility of equipment malfunction. In 2003, the reactor coolant loop and its component supports were re-analzed due to a power uprate. This latest analysis does not consider the coincident combination of blowdown and seismic loads. The original design basis is described in the following paragraphs. The primary system supports, steam generator, reactor coolant pump, pressurizer, and reactor vessel were designed to withstand pipe break or seismic acceleration based on the following:

1. The break is either a circumferential or longitudinal pipe rupture of area equivalent to the pipe cross section occurring anywhere in the system piping.

The longitudinal rupture occurs at any point 360 degrees around the pipe. The support system is designed to withstand the steady thrust equivalent to the product of system operating pressure and pipe rupture area without exceeding yield stress in the support members. The stress limits on the vessels and piping are tabulated in Section 1.11. The component supports prevent rupture of reactor coolant piping in the remaining intact loops which could result from an assumed rupture in any one loop, thereby ensuring that the path for safety injection flow to the core is available. Additionally, the supports are designed to prevent secondary piping rupture as a result of rupture in the primary loop and vice versa.

2. The nuclear steam supply system and its support system are designed such that the nuclear steam supply system is capable of continued safe operation for the combination of normal loads and the design earthquake loading. The equipment and supports operate within normal design limits for the design earthquake. The system and its supports are also designed to withstand the maximum potential earthquake without loss of function. The seismic response curves for both the design and maximum potential earthquake and the stress limits are presented in Section 1.11. Component loads are obtained from the curve using the appropriate period and damping.

3. The primary system supports were not originally designed to resist combined seismic and accident loads. They were designed as statically uncoupled component supports.

A complete reactor coolant system loop, including the steam generator and the reactor coolant pump supports, has been analyzed for combined dead, seismic, and blowdown loads. Stresses were determined by means of the three-dimensional frame computer program, STRUDL. The dead load assumed is the flooded weight of the component. The seismic load considered is 0.6g horizontal acceleration times the flooded mass of the component at the center of gravity of the component acting in the N-S, E-W, NW-SE and SW-NE directions analyzed Chapter 5, Page 50 of 89 Revision 22, 2010

IP2 FSAR UPDATE separately. The horizontal earthquake component acting on the steam generator is assumed to be carried by the upper steam generator. The vertical component of earthquake is assumed equal to 0.4g acting simultaneously with the horizontal load at the center of gravity of both the pump and steam generator. The system was analyzed for each separate accident or pipe rupture resulting in a jet load equal to 1500 kips as shown in Figure 5.1-34. The combined dead plus seismic plus accident maximum resultant member axial stress and axial plus bending stress (in parentheses) for the steam generator and pump supports is shown in Figures 5.1-35 through 5.1-42 (stresses are expressed in ksi.) The section views of the support shown can be identified by the isometric views of the pump and steam generator supports shown in Figures 5.1-43 and 5.1-44. Negative values indicate compression and positive values indicate tensile stress. Since response of the primary systems is elastic, deformations are very small and were not considered as design parameters required to verify the design adequacy of the supports. It should be noted the stresses shown are not for a particular combined blowdown or seismic load case but rather the worst combination for a given member; hence, the values shown are upper limits for each member and could not in fact actually occur in the combination shown. It should also be noted that the primary support structures are designed as trusses rather than frames hence the bending stresses indicated are secondary in nature. 5.1.5.1 Steam Generator The steam generators are supported within a caged structural system, consisting of four connected trusses, all welded together, fabricated of carbon steel members, with provisions for limited movement of the structure in a horizontal direction to accommodate piping expansion with a system of "Lubrite" plates, hydraulic snubbers, guides, and stops. The "Lubrite" plates, hydraulic snubbers, guides, and stops were originally designed as a rigid support to resist the action of seismic and pipe break loads. In 2000, the number of hydraulic snubbers supporting the steam generator frame in the direction of the hot leg, has been reduced from the original six down to two per steam generator. The two remaining snubbers are located at the upper support point of the frame at Elevation 92'-0". The analysis of the reactor coolant loop and of the steam generator support structure accounts for the replacement steam generator and for the reduced number of hydraulic snubbers. The following are loading conditions that the structure was originally designed to resist:

1. Vertical dead weight of pipe and vessel, flooded = 1,000 kips

2. Seismic loads:

a. Horizontal load of 474 kips acting at the centroid of the steam generator vessel, located near the top of the support structure, which is directly transferred to the hydraulic snubbers, guides, and stops, and in turn to the bottom of the 2-ft thick concrete operating floor slab at elevation 93-ft.

b. Vertical load of 320 kips transferred as axial load to the base plates and anchor bolts at elevation 46-ft.

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3. Primary system - longitudinal pipe rupture:

a. Reaction at the nozzle of the steam generator from the pipe between the reactor and the steam generator elbow, produces a force of 1090 kips in any direction and an overturning moment or torsional moment of (1090 kips x 4.25-ft) 4632-ft-kips. Overturning and torsional moments are resisted by the support system at elevation 46-ft and horizontal forces are distributed, through the truss action, to elevations 46-ft and 93-ft.

b. Reactions at the nozzle of the steam generator from the pipe between the steam generator elbow and reactor coolant pump elbow, produces a force of 850 kips in any direction and a torsional moment or overturning moment of (850 kips x 5.0-ft) 4250-ft-kips. Overturning and torsional moments are resisted by the support system at elevation 46-ft, and horizontal forces are distributed, through the truss action, to elevations 46-ft and 93-ft.

4 Primary system - circumferential break:

a. Reactions at the nozzle of the steam generator from the pipe between the reactor and steam generator produces a horizontal force of 1490 kips. This force is transferred through the vessel support to the two vertical trusses of the structural system, which in turn, transmits it as horizontal reactions at the slabs at elevations 46-ft and 93-ft. The moment produced by this force is (1490 kips x 2-ft) 2980-ft-kips and is less than the dead load resisting moment (500 kips x 10-ft) 5000-ft-kips, and the vertical forces at elevation 46-ft are all compressive, no uplift.

b. Reactions at the nozzle of the steam generator from a pipe between the steam generator and the reactor coolant pump produces a horizontal force of 1700 kips plus an overturning moment of (1700 kips x 4.25-ft) 7225-ft-kips, or a vertical force of 1700 kips and an overturning moment of (1700 kips x 5.33-ft) 9061-ft-kips. The horizontal force and moments are transferred to the structural system and the reactions are resisted at the slabs at elevations 46-ft and 93-ft, or the vertical force and moment are resisted at elevation 46 ft.

5. Secondary system - longitudinal rupture in steam pipe:

Reactions at the nozzle of the steam generator from the steam pipe longitudinal rupture at the top of the vessel produce:

a. Horizontal force of 600 kips and a torsional moment of 2400-ft-kips.

Horizontal force is transferred through the vessel to the structural support system, which in turn transmits it as horizontal reactions to the slabs at elevations 46-ft and 93-ft. The torsional moment is transferred through the vessel to the structural system, which in turn, transmits it to the base at elevation 46-ft.

b. Vertical upward or downward force of 600 kips and an overturning moment of 2400-ft-kips. Upward forces are overcome by the operating weight of the steam generator. Downward force is added to the operating weight and transferred to the base at elevation 46-ft. Overturning moment is transferred Chapter 5, Page 52 of 89 Revision 22, 2010

IP2 FSAR UPDATE through the vessel supports to the structural system, which in turn, transmits it as vertical reactions at the base, elevation 46-ft.

6. Secondary system - circumferential break in steam pipe:

Reaction at the nozzle of the steam generator from the steam pipe guillotine break at the top of the vessel produces a horizontal force of 600 kips. This force is transferred through the vessel to the structural system, which in turn transmits it as horizontal reactions of 1085 kips at elevation 93-ft and 485 kips at elevation 46-ft.

7. Secondary system - feedwater pipe breaks:

The reactions from circumferential and longitudinal pipe breaks in the feedwater system are resisted in a manner similar to steam pipe breaks listed under preceding sections (5) and (6), but are much smaller in magnitude. Maximum longitudinal 1600-ft-kips, maximum circumferential 200 kips. 5.1.5.2 Reactor Coolant Pump The reactor coolant pump is supported on a three-legged structural system consisting of three connected trusses fabricated of carbon steel members, structural sections and pipe, supported from elevation 48-ft-6-in. Provisions for limited movement of the structure in any horizontal direction to accommodate piping expansion is accomplished with a sliding "Lubrite" base plate arrangement and a system of tie rods and anchor bolts that restrain the structure from movement beyond the calculated limits. To improve the ability of the reactor coolant pumps to meet combined LOCA and seismic loads, two of the reactor coolant pump holddown bolts have been replaced with higher strength ASTM A540 steel bolts. The following are loading conditions that the structure was originally designed to resist:

1. Vertical dead weight of pipe and pump flooded = 206 kips.

2. Seismic:

a. Horizontal load of approximately 117 kips acting at the centroid of the pump assembly, which is transferred by the structural system and piping to the tie rods and base of the supporting structure at elevation 48-ft-6-in. This load includes the seismic effect of the support self-weight.

b. Vertical seismic load of approximately 78 kips transferred directly as axial load to the base plates and anchor bolts. This load includes the seismic effect of the support self-weight.

3. Primary system - longitudinal rupture:

a. Reaction at the nozzle of the pump from a pipe break in the pipe between the steam generator elbow and pump elbow produces a torsional moment of 3825-ft-kips, together with a horizontal force of 850 kips or an overturning moment of 3825-ft-kips, together with a vertical up or down force of 850 kips.

Torsional forces are resisted by the structural stability of the primary piping connected to the pump. Chapter 5, Page 53 of 89 Revision 22, 2010

IP2 FSAR UPDATE Reactions from horizontal forces are resisted by the tie rods connected to the steam generator and reactor support structures. Forces caused by an overturning moment are resolved into horizontal and vertical components, which are resisted by tension in the anchor bolts, axial load on the foundations, and tension in the tie rods.

b. Reaction at the nozzle of the pump from a pipe break in the pipe between the pump and the reactor, produces a torsional moment of 6880-ft-kips, together with a horizontal force of 1165 kips, or an overturning moment of 6880-ft-kips, together with a vertical up or down force of 1165 kips.

Torsional forces are resisted by the structural stability of the primary piping connected to the pump. Reactions from the horizontal forces are resisted by the tie rods connected to the walls. Forces caused by an overturning moment are resolved into horizontal and vertical components, which are resisted by: (1) Tension in the anchor bolts. (2) Axial load on the foundations. (3) Tension in the tie rods.

4. Primary system - circumferential break:

a. Reactions at the nozzle of the pump from a pipe break in the pipe between the steam generator and pump produces a horizontal force on the structure of 1700 kips. This force is resisted directly by the bumper located against the elbow of the pipe. Components of the force are then transferred to the base of the structure and the tie rods connecting the pump support to the steam generator support system.

b. Reactions at the nozzle of the pump from a pipe break in the pipe between the pump and the reactor produces a torsional moment of 3240-ft-kips and a horizontal force of 1340 kips on the structure.

Torsional forces are resisted by the structural stability of the remaining primary piping connected to the pump. Reactions from the horizontal forces are resisted by tie rods connected to the walls. 5.1.5.3 Pressurizer Pressurizer is supported on a free-standing structural system, consisting of six connected trusses fabricated of carbon steel members, all welded together and secured at the base by anchor bolts at elevation 46-ft. The following are loading conditions that the structure has been designed to resist:

1. Vertical dead weight of pipe and vessel flooded is 360 kips. The self-weight of the support is 21 kips.

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2. Seismic:

a. Horizontal seismic load of 174 kips acting at the centroid of the pressurizer vessel, which coincides in elevation with the slab at elevation 95-ft, is directly transferred through the concrete embedded guides to the slab. This load excludes the seismic effect of the support self-weight.

b. Vertical seismic load of 123 kips transferred through the structural system as axial forces to the base plates and anchor bolts at elevation 46-ft. This load excludes the seismic effect of the support self-weight.

3. Longitudinal pipe rupture

a. Reaction at the surge pipe nozzle of the pressurizer produces either a torsional moment of 734-ft-kips and a horizontal force of 234 kips or an overturning moment of 734-ft-kips and a horizontal or vertical force of 234 kips.

These moments and forces are resisted by the structural system and transferred to the base at elevation 46-ft.

4. Circumferential pipe break:

a. Reaction at the surge pipe nozzle of the pressurizer produces a horizontal force of 234 kips and an overturning moment of 734-ft-kips.

These moments and forces are resisted by the structural system and transferred to the base at elevation 46-ft. 5.1.5.4 Reactor Vessel Support Girder The reactor vessel is supported on four cooling plates that are fastened to the top flange of a circular box section ring girder, fabricated of carbon steel plates. The bottom flange of the girder is in continuous contact with a nonyielding concrete foundation. In addition to the reactor vessel weight and piping reactions of the girder has been designed to support the conditions of loading for pipe break and seismic forces as outlined in Figure 5.1-45. 5.1.5.5 Reactor Vessel Rupture The reactor pressure vessel is enclosed by a 6-ft thick circular reinforced concrete shield wall that is designed to sustain the internal pressure and provide missile protection for the containment liner in the highly unlikely failure of the reactor vessel due to a longitudinal split. All stresses will be maintained within specified minimum ultimate rebar tensile stress. In the event of a circumferential reactor break, the 0.25-in. basemat liner plate at the bottom of the containment reactor cavity pit directly under the reactor vessel is protected by 2-ft of concrete with a 1-in. steel liner plate embedded on top of the concrete. Directly below the reactor cavity pit containment basemat liner plate, 4.5-ft of concrete is poured on rock. Refer to Figures 5.1-46 through 5.1-51. Chapter 5, Page 55 of 89 Revision 22, 2010

IP2 FSAR UPDATE As discussed in Section 5.1.3.7, in the event of reactor vessel failure, a pressure build up of 1000 psi and rebar stresses of 86 ksi (assuming all concrete is cracked) inside the pit due to release of reactor contents is assumed. Since the integrity of the wall is not jeopardized, the integrity of the vessel support that is supported on the wall will not be jeopardized. Deflection of the shield wall will not cause large stresses in the vessel support since a lubricated surface is provided on the shoes, allowing the vessel support to slide. 5.1.5.6 Circumferential Cracking The worst circumferential crack location from the standpoint of downward missiles is just below the reactor coolant system piping nozzles. As the following calculations show, this missile will not violate the containment structure and liner integrity. As a consequence of this circumferential crack, the downward missile represented by bottom vessel head has the following characteristics at the time of impact on the cavity floor:

1. Weight: 381,000 lb

2. Cross sectional area of crater: 63-ft2

3. Downward velocity: 213-ft/sec

4. Concrete crushing strength: 4000 psi The depth of penetration has been calculated by using the Petri formula for penetration into an infinite, thick concrete slab, as reported in Nav. Docket P-51:

W V2 D = K log10 (1 + ) A 215,000 where: D = depth of penetration, ft K = penetration coefficient for 4000 psi concrete W = missile weight, lb A = missile area, ft2 V = missile velocity, ft/sec The following parameters have been used: K = 2.8 x 10-3 W = 381,000 lb A = 63-ft2 V = 213 ft/sec The result is a depth of penetration of 1.4-ft. As mentioned above, the 0.25-in. basemat liner is covered by 2-ft of concrete with a 1-in. steel plate on top. As it can be readily seen, even neglecting the 1-in. steel plate in the penetration calculations, the containment liner will not be reached. Chapter 5, Page 56 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.5.7 Longitudinal Splitting The cavity wall is designed to withstand the forces and internal pressurization associated with a longitudinal split without gross damage. See Section 5.1.3.7 for a discussion of the analysis of this assumed accident condition. 5.1.6 Containment Structure Design Evaluation 5.1.6.1 Reliance On Interconnected Systems The containment leakage limiting boundary is provided in the form of a single, carbon steel liner on the vessel having double barrier weld channels and penetrations. Each system whose piping penetrates this boundary is designed to maintain isolation of the containment from the outside environment. Provision is made to continuously pressurize penetrations and weld channels and to monitor leakage from this pressurization. 5.1.6.2 System Integrity and Safety Factors Pipe Rupture - Penetration Integrity - The penetrations for the blowdown and sample lines are designed so that the penetrations are stronger than the piping system and so that the vapor barrier will not be breached due to a hypothesized pipe rupture. The pipe rupture loads for the main steam and feedwater lines are resisted by the supports located away from their penetrations and do not affect the integrity of the penetrations for these lines. Major Component Support Structures - The support structures for the major components are designed to resist all thrust forces, moments and torques associated with either a reactor coolant system or main steam pipe break. All primary structural steel elements are designed for stresses not exceeding yield stress due to these forces. Containment Structure Components Analyses - The details of radial, longitudinal, and horizontal shear analyses for the containment reinforced concrete are given in Section 5.1.3. 5.1.6.3 Performance Capability Margin The containment structure is designed based upon limiting load factors, which are used as the ratio by which accident and earthquake loads are multiplied for design purposes to ensure that the load/deformation behavior of the structure is one of elastic, low strain behavior. This approach places minimum emphasis on fixed gravity loads and maximum emphasis on accident and earthquake loads. Because of the refinement of the analysis and the restrictions on construction procedures, the load factors primarily provide for a safety margin on the load assumptions. Load combinations and load factors used in the design, which provide an estimate of the margin with respect to all loads, are tabulated in Section 5.1.2. 5.1.7 Liner Insulation Insulation is provided on approximately the first 43-ft of the containment liner to limit the temperature rise in the liner under accident conditions to 80°F above ambient and thereby avoid excessive liner compressive stress during the accident. The first 18-ft (elev. 46-ft to 64-ft, except in the piping penetration area in the southeast quadrant where the insulation rose only 16-ft to the 62-ft elevation) was installed as part of the original containment design. In 1973 an additional 25-ft (elev. 64-ft to 89-ft) was added. The first 18-ft (elev. 64-ft to 82-ft) covers the Chapter 5, Page 57 of 89 Revision 22, 2010

IP2 FSAR UPDATE entire circumference of the liner while the upper 7-ft (elev. 82-ft to 89-ft) only covers part of the circumferential area in the north and south-southwest quadrants where the main steam and feedwater lines extend up along the crane wall. The insulation panels are attached to the steel containment liner by means of 3/16-in. diameter stainless steel studs welded to the liner on the basis of six per panel. The insulation panels are protected by stainless steel jacketing on the exposed faces and sealed at the joints. Details of the insulation installation are given in Table 5.1-2. The insulation has been designed to meet the following operational requirements:

1. Normal operating temperature of 120°F. (The maximum normal operating temperature of the containment was changed from 120°F to 130°F by Amendment 149 to the Facility Operating License DPR-26 for IP-2 dated March 27, 1990. Evaluations performed show the insulation material used on the containment liner is adequate for use at the higher operating temperature.)

2. Under accident conditions, the rise in liner temperature not to exceed 80°F. The analyses performed to support the Stretch Power Uprate (SPU) also performed analyses of the containment liner under the most limiting conditions for liner stress and showed a temperature rise well under allowed 80oF.

3. Insulation panels to be rated nonburning in accordance with ASTM procedure D1692.

4. To be removable by sections for inspection of the containment liner.

5.1.8 Minimum Operating Conditions (For Containment Integrity) Containment integrity internal pressure limitations and leakage rate requirements are established in the facility Technical Specifications. 5.1.9 Containment Structure-Inspection And Testing 5.1.9.1 Initial Containment Leakage Rate Testing Criterion: Containment shall be designed so that integrated leakage rate testing can be conducted at the peak pressure calculated to result from the design basis accident after completion and installation of all penetrations and the leakage rate shall be measured over a sufficient period of time to verify its conformance with required performance. (GDC 54) After completion of the containment structure and installation of all penetrations and weld channels, an initial integrated leakage rate test was conducted at the containment design pressure (47 psig), maintained for a minimum of 24 hr, verifying that the leakage rate is no greater than 0.1-percent by weight of the containment volume per day at design basis accident conditions. This leakage rate test was performed using the absolute method. In addition, a reduced pressure integrated leakage rate test was conducted at a pressure not less than 50 percent of the containment design pressure and maintained for a minimum of 24 hr. Chapter 5, Page 58 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.9.2 Periodic Containment Leakage Rate Testing Criterion: The containment shall be designed so that an integrated leakage rate can be periodically determined by test during plant lifetime. (GDC 55) The containment is tested in accordance with 10 CFR 50 Appendix J as discussed in section 5.1.12. A leak rate test at the containment design pressure using the same method as the initial leak rate test can be performed at any time during the operational life of the plant, provided the plant is not in operation and precautions are taken to protect instruments and equipment from damage. 5.1.9.3 Provisions for Testing of Penetrations Criterion: Provisions shall be made to the extent practical for periodically testing penetrations, which have resilient seals or expansion bellows to permit leak tightness to be demonstrated at the peak pressure calculated to result from occurrence of the design basis accident. (GDC 56) Penetrations are designed with double seals, which are continuously pressurized above accident pressure. The large access openings such as the equipment hatch and personnel air locks are equipped with double gasketed doors and flanges with the space between the gaskets connected to the pressurization system. The system uses a supply of clean, dry, compressed air that will place the penetrations under an internal pressure above the peak calculated accident pressure. A permanently piped monitoring system is provided to continuously measure leakage from all penetrations. Leakage from the monitoring system is checked by continuous measurement of the integrated makeup air flow. In the event excessive leakage is discovered, each penetration can then be checked separately at any time. 5.1.9.4 Provisions for Testing of Isolation Valves Criterion: Capability shall be provided to the extent practical for testing functional operability of valves and associated apparatus essential to the containment function for establishing that no failure has occurred and for determining that valve leakage does not exceed acceptable limits. (GDC 57) Capability is provided to the extent practical for testing the functional operability of valves and associated apparatus during periods of reactor shutdown. Initiation of containment isolation employs coincidence circuit, which allow checking of the operability and calibration of one channel at a time. Removal or bypass of one signal channel place that circuit in the half-tripped mode. Hydrostatic tests of isolation valves in series are performed by first testing the upstream valve with the second valve open, then opening the upstream valve and closing the second valve, so that each valve will have an independent test. Chapter 5, Page 59 of 89 Revision 22, 2010

IP2 FSAR UPDATE The main steam and feedwater barriers and isolation valves in systems that connect to the reactor coolant system are hydrostatically tested to measure leakage. Valves in the residual heat removal system are not considered to be isolating valves in the usual sense inasmuch as the system would be in operation under accident conditions. Field and operational inspection and testing have been divided into three phases:

1. Construction tests; those taking place during erection of the containment building liner.

2. Preoperational tests; those taking place after the containment structure was erected and all penetrations were complete and installed.

3. Postoperational tests; monitoring during reactor operation.

5.1.10 Construction Tests During erection of the liner, the following inspection and tests were performed. 5.1.10.1 Bottom Liner Plates All liner plate welds are tested for leaktightness by vacuum box. The box is evacuated to at least a 5 psi pressure differential with the atmospheric pressure. After completion of a successful leak test, the welds were covered by channels. A strength test was performed by applying 54 psig air pressure to the channels in the zone for a period of 15 min. The zone of channel covered welds was pressurized to 47 psig with a 20-percent by weight of freon-air mixture. The entire run of the channel-to-plate welds was then traversed with a halogen leak detector. The sensitivity of the leak detector was 1 x 10-9 standard cc per second. The sniffer was held approximately 0.5-in. from the weld and traversed at a rate of about 0.5-in./sec. The detection of any amount of halogen indicated a leak requiring weld repairs and retesting. After the halogen test was completed, all liner welds not accessible for radiography were pressurized with air to 47 psig and soap-tested. Any leaks indicated by bubbles were repaired and retested. Where leaks occurred, welds were removed by arc gouging, grinding, chipping, and/or machining before rewelding. In addition, the zone of channels was held at the 47 psig air pressure for a period of at least 2 hr. The drop in pressure did not exceed the equivalent of a leakage of 0.05-percent of the containment building volume per day. Compensation for change in ambient air temperature was made. 5.1.10.2 Vertical Cylindrical Walls and Dome For the liner, a complete radiograph was made of the first 10-ft of full penetration weld made by each welder or welding operation. A minimum of a 12-in. film "spot" radiograph was made every 50-ft of weld thereafter on the side walls and dome, except where backup plates were Chapter 5, Page 60 of 89 Revision 22, 2010

IP2 FSAR UPDATE used. The radiograph films were reviewed by United Engineers and Constructors. When a spot radiograph showed defects that required repair, two adjacent spots were radiographed. If defects requiring repair were shown in either of these, all of the welding performed by the responsible operator or welder was 100-percent radiographed to determine the area of defect. The performance and acceptance standards for all radiography were ASME Section VIII, Paragraph UW51. The liner plate-to-plate welds were tested for leaktightness by vacuum box techniques. After successful completion of the spot radiography and vacuum box tests and subsequent repair of all defects, the channels were welded in place over all seam welds in a predetermined zone. A strength test was performed on the liner plate weld and the channel weld by pressurizing the channel with air at 54 psig for 15 min. In addition, each zone of channel covered weld was leak-tested using the freon-air mixture at 47 psig. In locations where radiography was not possible, such as the lower courses of shell plates where backup plates were used, and where liner bottom welds and floor plate welds were made to angles and tees, the liner fabricator welded on a 2-in. long overrun coupon. The overrun coupon was chipped off, marked for location and given to United Engineers and Constructors for testing. These welds were also vacuum box tested. Welded studs were visually inspected, and at least one at the beginning of each day's work and another at approximately mid-day were bend-tested to 45 degrees for each welder. Studs failing visual or bend-testing were removed. While the liner is not a pressure vessel, industry experience has shown that leaks in pressure vessels normally occur at joints. For this reason, and following current liner fabrication practice, there was no radiographic or other nondestructive examination of liner plate. 5.1.10.3 Penetrations Strength and leak tests of individual penetration internals and closures and sleeve weld channels were performed in a similar manner to the above and all leaks repaired and the penetration or weld channel retested until no further leaks were found. See Figures 5.1-53 through 5.1-56 for the areas of the containment and liner, which were instrumented for the strength test. 5.1.11 Preoperational Tests All penetrations and the welds joining these penetrations to the containment liner and the liner seam welds were designed to provide a double barrier, which can be continuously pressurized at a pressure higher than the design pressure of the containment. This blocks all of these potential sources of leakage with a pressurized zone and at the same time provides a means of monitoring the leakage status of the containment, which is more sensitive to changes in the leakage characteristics of these potential leakage sources. After the containment building was complete with liner, concrete structures, and all electrical and piping penetrations, equipment hatch and personnel locks were in place, the following tests were performed. Chapter 5, Page 61 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.1.11.1 Strength Test A pressure test was made on the completed building using air at 54 psig. This pressure was maintained on the building for a period of at least 1 hr. During this test, measurements and observations were made to verify the adequacy of the structural design. For a description of observations, cracks, strain gauges, etc., refer to Reference 18. 5.1.11.2 Integrated Leakage Rate Test: (Type A) The integrated leakage rate tests were performed on the containment building at 47 psig using the absolute method. This leakage test was performed with the double penetration and weld channel zones open to the containment atmosphere. The leakage rate demonstrated by this test was equal to or less than 0.1-percent of the containment free volume per day at design basis accident conditions. After it was assured that there were no defects remaining from construction, a sensitive leak rate test was conducted. 5.1.11.3 Sensitive Leak Rate Test: (Type B) The sensitive leak rate test included only the volume of the weld channels and double penetrations. This test was considered more sensitive than the integrated leakage rate test, as the instrumentation used permitted a direct measurement of leakage from the pressurized zones. The sensitive leak rate test was conducted with the penetrations and weld channels at 50 psig and with the containment building at atmospheric pressure. The leak rate for the double penetrations and weld channel zones was equal to or less than 0.2-percent of the containment free volume per day. 5.1.11.4 Containment Isolation Valve Test: (Type C) These tests were conducted to detect leaks through certain containment isolation valves. 5.1.12 Postoperational Tests Containment testing is conducted in accordance with the Technical Specifications and 10 CFR Appendix J, including integrated leakage rate tests at the containment design pressure, In 1997, the Technical Specifications were amended to allow the use of 10 CFR 50 Appendix J, Option B (as modified by approved exemptions) and NRC Regulatory Guide 1.163 dated September 1995 for integrated leakage rate tests, air lock tests, and containment isolation valve operability tests. REFERENCES FOR SECTION 5.1

1. "Evaluation of Protection Coatings For Use in Reactor Containment," WCAP-7198-L, April 1968.

2. J.B. Stellmeyer, W. H. Munse, and E. A. Selby, Fatigue Tests of Plates and Beams with Stud Shear Connections, Highway Research Record, No. 76.

3. Robert C. Singleton, "The Growth of Stud Welding," Welding Engineer, July 1963.

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4. L. S. Beedle, "Plastic Design of Steel Frames," P. 131, Copyright 1958, Second print, 1961.

5. Design Data, Nelson Concrete Anchor, Printed August 1, 1961.

6. M. Hetenyi, "Beams on Elastic Foundation," Ann Arbor Press, 1955.

7. S. Timoshenko, and D. H. Young, "Elements of Strength of Materials," D. Van Nostrand Company, Inc., 1962.

8. S. Timoshenko, and S. Woinowsky-Krieger, Theory of Plates and Shells, Second Edition, McGraw-Hill, 1959.

9. American Concrete Institute, Specification for the Design and Construction of Reinforced Concrete Chimneys ACI-505-54, ACI Manual of Concrete Practices, Part 2, 1967.

10. United States Atomic Energy Commission, "Nuclear Reactors and Earthquakes,"

TID-7024, 1963.

11. J. Blume, N. Newmark, and L. Corning, "Design of Multistory Reinforced Concrete Buildings for Earthquake Motions," Portland Cement Association, 1961.

12. Electric Power Research Institute, Good Bolting Practices, A Reference Manual for Nuclear Power plant Maintenance Personnel, Volume 1: Large Bolt Manual, NP5067, 1987.

13. Letter from Donald Brinkman, NRC, to Stephen Bram, Con Edison,

Subject:

Emergency Amendment to Increase the Service Water Temperature to 90°F (TAC 73764), dated August 7, 1989.

14. Raymond J. Roark, "Formulas for Stress and Strain," McGraw-Hill Book Company, Fourth Edition.

15. R. V. Whitman, and F. E. Richart, Jr. "Design Procedures for Dynamic Loaded Foundations," A.S.C.E., Journal of the Soil Mechanics and Foundations Division, Vol. 93, No. SM6, 1967.

16. Paul F. Rice, Detailing and Placing of Reinforcing Bars, Concrete Construction, January 1965.

17. S. Timoshenko and Goodier, Theory of Elasticity, Second Edition, McGraw-Hill Book Co., 1961.

18. Wiss, Janney, Elstner & Associates, Inc., Consulting & Research Engineers, Structural Response of Secondary Containment Vessel During Structural Integrity Test at Indian Point Power Generating Station Unit No. 2, prepared for Wedco Corporation, April 22, 1971.

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19. Altran Corporation, Technical Report No. 00222-TR-001, Rev. 1, Reactor Coolant Loop Analysis for Replacement Steam Generators and Snubber Reduction" TABLE 5.1-1 Flooded Weights - Containment Building Item Flooded Operating Weight, lbs Pressurizer -1 346,000 Steam generators - 4 3,746,000 Reactor - 1 (a) Vessel 868,000 (b) Internals 420,000 (c) Piping 1,000,000 Reactor pumps - 4 824,000 Accumulator tanks - 4 529,000 175-ton polar crane - 1 650,000 Ventilation fans - 5 656,000 Reactor coolant drain tank - 1 20,000 Pressure relief tank - 1 100,000 Other miscellaneous equipment 100,000 Total 9,259,000 Chapter 5, Page 64 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.1-2 Containment Liner Insulation Properties

1. Elevation 46-ft to 64-ft liner insulation:

• 1-1/4-in. polyvinylchloride insulation, Vinylcel, as manufactured by Johns-Mansville and 1-1/2 in. Pittsburgh Corning Foamglass Insulation.
• 0.019-in. thick stainless steel jacket (exposed side) except for areas using Pittsburgh Corning Foamglass Insulation in which a jacket thickness of 0.024 is used.
• Insulation adhesive is Johns-Manville Dutch Brand FN12 or an approved equal.

2. Elevation 64-ft to 89-ft liner insulation:1

• 1.5-in. thick FOAMGLASR with density of 8.5 to 9 lb/ft3, as manufactured by Pittsburgh Corning Corporation. This insulation has a thermal conductivity of 0.5 - 0.525 BTU-in/hr-ft2-°F and a specific heat (Cp) of 0.18 BTU/lb-°F.
• 1/16-in. commercial grade pure asbestos paper backing adjacent to the liner plate on the unexposed face.
• The adhesive bonding the FOAMGLASR to the asbestos paper is Cadoprene No. 434 and bonding the stainless steel jacket to the FOAMGLASR is Cadoseal No. 700 by Epolux Manufacturing Corporation.

Note: 1 Insulation from Elevation 82-ft to 89-ft only covers part of the circumferential area in the north and south-southwest quadrants. 5.1 FIGURES Figure No. Title Figure 5.1-1 Containment Structure Figure 5.1-2 Containment Building General Arrangement Plans, Sheet 1 - Replaced with Plant Drawing 9321-2501 Figure 5.1-3 Containment Building General Arrangement Plans, Sheet 2 - Replaced with Plant Drawing 9321-2502 Figure 5.1-4 Containment Building General Arrangement Plans, Sheet 3 - Replaced with Plant Drawing 9321-2503 Figure 5.1-5 Containment Building General Arrangement Elevation - Sheet 1 Replaced with Plant Drawing 9321-2506 Figure 5.1-6 Containment Building General Arrangement Elevation - Sheet 2 Replaced with Plant Drawing 9321-2507 Figure 5.1-7 Containment Building General Arrangement Elevation - Sheet 3 Replaced with Plant Drawing 9321-2508 Figure 5.1-8 Deleted Figure 5.1-9 Deleted Chapter 5, Page 65 of 89 Revision 22, 2010

IP2 FSAR UPDATE Figure 5.1-10 Deleted Figure 5.1-11 Cylinder and Dome-Load Condition (A) - 1.5P Figure 5.1-12 Cylinder and Dome-Load Condition (B) - 1.25P Figure 5.1-13 Cylinder and Dome-Load Condition (C) - 1.0P Figure 5.1-14 Loading Diagram in Mat-Load Condition (A) - 1.5P Figure 5.1-15 Loading Diagram in Mat-Load Condition (B) - 1.25P Figure 5.1-16 Loading Diagram in Mat-Load Condition (C) - 1.0P Figure 5.1-17 Weld Stud Connection at Panel Low Point Figure 5.1-18 Weld Stud Connection At Panel Low Point Figure 5.1-19 Weld Stud Connection at Panel Center Figure 5.1-20 Wall Section Figure 5.1-21 Cylinder Base Slab Liner Juncture Figure 5.1-22 Typical Base Mat Liner Detail Figure 5.1-23 Base Slab Reinforcing Detail Figure 5.1-24 Reactor Cavity Pit Figure 5.1-25 Equipment Hatch Personnel Lock, Main Steam and Feedwater, Air Purge - Rebar Figure 5.1-26 Torsional Effects Figure 5.1-27 Typical Electrical Penetration Figure 5.1-28 CONAX Penetrations - Outside Containment Weld Figure 5.1-29 CONAX Penetrations - Inside Containment Weld Figure 5.1-30 Typical Piping Penetration Figure 5.1-31 Fuel Transfer Tube Penetration (Conceptual Drawing) Figure 5.1-32 Containment-Stresses on Penetrations and Liner - Sheet 6 Figure 5.1-33 Containment-Stresses on Penetrations and Liner - Sheet 7 Figure 5.1-34 Assumed Pipe Rupture Accident Break Locations Figure 5.1-35 Steam Generator Support-Section 1-1 Figure 5.1-36 Steam Generator Support-Section 2-2 Figure 5.1-37 Steam Generator Support-Section 3-3 Figure 5.1-38 Steam Generator Support-Section 4-4 Figure 5.1-39 Steam Generator Support-Plan Location Elevation 60 and 63 Figure 5.1-40 Steam Generator Support-Plan Location Elevation 60 and 63 Figure 5.1-41 Pump Support-Section 2-2 and 3-3 Figure 5.1-42 Pump Support-Section 3-3 Figure 5.1-43 Isometric View-Steam Generator Support Figure 5.1-44 Isometric View-Reactor Coolant Pump Support Figure 5.1-45 Maximum Forces Acting on a Reactor Vessel Support Figure 5.1-46 Plan View 60 Ft-0 In. Figure 5.1-47 Typical Layer-Reactor Ring Figure 5.1-48 Section 5-5 Figure 5.1-49 Section 18-18 Figure 5.1-50 Plan View at Elevation 19 Ft-7 In. Figure 5.1-51 Section A-A and Section B-B Figure 5.1-52 Deleted Figure 5.1-53 Containment Equipment Hatch Strain Gauge Test Locations Figure 5.1-54 Containment Temporary Opening in NW Quadrant Strain Gauge Test Locations Figure 5.1-55 Containment Strain Gauge Test Locations Figure 5.1-56 Containment Proof Test Gross Deformation Measurements Chapter 5, Page 66 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.2 CONTAINMENT ISOLATION SYSTEM 5.2.1 Design Basis Each system whose piping penetrates the containment leakage limiting boundary is designed to maintain or establish isolation of the containment from the outside environment under the following postulated conditions:

1. Any accident for which isolation is required (severely faulted conditions) with

2. A coincident independent single failure or malfunction (expected fault condition) occurring in any active system component within the isolated bounds.

Piping penetrating the containment is designed for pressures at least equal to the containment design pressure. Containment isolation valves are provided as necessary in lines penetrating the containment to ensure that no unrestricted release of radioactivity can occur. Such releases might be due to rupture of a line within the containment concurrent with a loss-of-coolant accident, or due to rupture of a line outside the containment that connects to a source of radioactive fluid within the containment. In general, isolation of a line outside the containment protects against rupture of the line inside concurrent with a loss-of-coolant accident, or closes off a line, which communicates with the containment atmosphere in the event of a loss-of-coolant accident. Isolation of a line inside the containment prevents flow from the reactor coolant system or any other large source of radioactive fluid in the event that a piping rupture outside the containment occurs. A piping rupture outside the containment at the same time as a loss-of-coolant accident is not considered credible, as the penetrating lines are seismic Class I design up to and including the second isolation barrier and are assumed to be an extension of containment. The isolation valve arrangement provides two barriers between the reactor coolant system or containment atmosphere and the environment. System design is such that failure of one valve to close will not prevent isolation, and no manual operation is required for immediate isolation. Automatic isolation is initiated by a containment ventilation isolation signal, a Phase A isolation signal ("T" signal), a Phase B isolation signal ("P" signal), or manually. See Section 5.2.4 or Chapter 7.0 for further details. The containment isolation valves have been examined to ensure that they are capable of withstanding the maximum potential seismic loads. To ensure their adequacy in this respect:

1. Valves are located in a manner to reduce the accelerations on the valves.

Valves suspended on piping spans are reviewed for adequacy for the loads to which the span would be subjected. Valves are mounted in the position recommended by the manufacturer.

2. Valve yokes are reviewed for adequacy and strengthened as required for the response of the valve operator to seismic loads.

Chapter 5, Page 67 of 89 Revision 22, 2010

IP2 FSAR UPDATE

3. Where valves are required to operate during seismic loading, the operator forces are reviewed to ensure that system function is preserved. Seismic forces on the operating parts of the valve are small compared to the other forces present.

4. Control wires and piping to the valve operators are designed and installed to ensure that the flexure of the line does not endanger the control system.

Appendages to the valve, such as position indicators and operators, are checked for structural adequacy. Isolation valves are provided as necessary for all fluid system lines penetrating the containment to ensure at least two barriers for redundancy against leakage of radioactive fluids to the environment in the event of a loss-of-coolant accident. These barriers, in the form of isolation valves or closed systems, are defined on an individual line basis. In addition to satisfying containment isolation criteria, the valving is designed to facilitate normal operation and maintenance of the systems and to ensure reliable operation of other engineered safeguards systems. With respect to numbers and locations of isolation valves, the criteria applied are generally those outlined by the seven classes described in Section 5.2.2 below. 5.2.2 System Design The seven classes listed below are general categories into which lines penetrating containment may be classified. The seal water referred to in the listing of categories is provided by the isolation valve seal water system described in Section 6.5. The following notes apply to these classifications.

1. The "not-missile-protected" designation refers to lines that are not protected throughout their length inside containment against missiles generated as the result of a loss of coolant accident. These lines, therefore, are not assumed invulnerable to rupture as a result of a loss of coolant.

2. In order to qualify for containment isolation, valves inside the containment must be located behind the missile barrier for protection against loss of function following an accident.

3. Manual isolation valves that are locked closed or otherwise closed and under administrative control during power operation qualify as automatic trip valves.

4. A check valve qualifies as an automatic trip valve in certain incoming lines not requiring seal water injection.

5. The double disc type of gate valve is used to isolate certain lines. When sealed by water injection, this valve provides two barriers against leakage of radioactive liquids or containment atmosphere.

6. In lines isolated by globe valves and provided with seal water injection, the valves are generally installed so that the seal water wets the stem packing.

Chapter 5, Page 68 of 89 Revision 22, 2010

IP2 FSAR UPDATE

7. Loss of seal water through those isolation valves closed only by a containment isolation phase B signal is prevented by solenoid operated valves in the seal water injection lines. Excessive loss of seal water through motor operated isolation valves that could fail to close in response to a containment isolation signal is limited by flow restrictive orifices installed in the seal water lines. A water seal at the failed valve is ensured by proper slope of the protected line, or a loop seal, or by additional valves on the side of the isolation valves away from the containment.

8. Isolated lines between the containment and the second outside isolation valve are designed to the same seismic criteria as the containment vessel and are considered to be an extension of containment.

5.2.2.1 Class 1, Outgoing Lines, Reactor Coolant System Outgoing lines connected to the reactor coolant system that are normally or intermittently open during reactor operation are provided with at least two automatic trip valves in series located outside the containment. Automatic seal water injection is provided for lines in this classification. An exception to the general classification is the residual heat removal loop's reactor coolant system suction line, which has two barriers established by normally closed valves located outside containment. 5.2.2.2 Class 2, Outgoing Lines Outgoing lines not connected to the reactor coolant system that are normally or intermittently open during operation and not missile protected or that can otherwise communicate with the containment atmosphere following an accident are provided at a minimum with two automatic trip valves in series or a single automatic double-disc gate valve outside containment. Automatic seal water injection is provided for lines in this classification. Most of these lines are not vital to plant operation following an accident. 5.2.2.3 Class 3, Incoming Lines Incoming lines connected to open systems (i.e., systems that are in some way connected to the containment environment) outside containment, and not missile protected or that can otherwise communicate with the containment atmosphere following an accident are provided with one of the following arrangements outside containment:

1. Two automatic trip valves in series, with automatic seal water injection. This arrangement is provided for lines that are not necessary to plant operation after an accident.

2. Two manual isolation valves in series, with manual seal water injection. This arrangement is provided for lines that remain in service for a time, or are used periodically, subsequent to an accident.

Incoming lines connected to closed systems outside containment, and not missile protected or that can otherwise communicate with the containment atmosphere, are provided either with two isolation valves in series outside containment with seal water injection, or at a minimum, with Chapter 5, Page 69 of 89 Revision 22, 2010

IP2 FSAR UPDATE one check valve or normally closed isolation valve located either inside or outside containment. The closed piping system outside containment provides the necessary isolation redundancy for lines that contain only one isolation valve. Exception is the containment spray headers, for which valving is based on safeguards requirements. 5.2.2.4 Class 4, Missile Protected Lines Incoming and outgoing lines that penetrate the containment and that are normally or intermittently open during reactor operation and are connected to closed systems inside the containment and protected from missiles throughout their length, are provided with at least one manual isolation valve located outside the containment. Seal water injection is not required for this class of penetration. An exception is the residual heat exchanger cooling water lines for which design is based on safeguards requirements. 5.2.2.5 Class 5, Normally Closed Lines Penetrating the Containment Lines that penetrate the containment and that can be opened to the containment atmosphere but that are normally closed during reactor operation are provided with two isolation valves in series or one isolation valve and one blind flange. 5.2.2.6 Class 6, Special Service Lines There are a number of special groups of penetrating lines and containment access openings. These are discussed below. Each ventilation purge duct penetration is provided with two tight-closing butterfly valves, which are normally closed during reactor power operation and are actuated to the closed position automatically upon a containment isolation or a containment high radiation signal. One valve is located inside and one valve is located outside the containment at each penetration. The space between valves is pressurized by air from the weld channel and penetration pressurization system whenever they are closed. Blind flanges can also be used for containment isolation in place of automatic purge isolation valves, provided they meet the same design criteria as the isolation valves. The containment pressure relief line is similarly protected. However, because the line is periodically opened during reactor power operation, three tight closing butterfly valves in series are provided, one inside and two outside the containment. These valves also are actuated to the closed position upon a containment isolation or containment high radiation signal. The two intravalve spaces are pressurized by air from the weld channel and penetration pressurization system whenever they are closed. The equipment access closure is a bolted, gasketed closure that is air sealed during reactor operation. The personnel air locks consist of two doors in series with mechanical interlocks to ensure that one door is closed at all times. Each air lock door and the equipment closure are provided with double gaskets to permit pressurization between the gaskets by the weld channel and penetration pressurization system. (See Section 6.6.) The fuel transfer tube penetration inside the containment is designed to present a missile-protected and pressurized double barrier between the containment atmosphere and the atmosphere outside the containment. The penetration closure is treated in a manner similar to Chapter 5, Page 70 of 89 Revision 22, 2010

IP2 FSAR UPDATE the equipment access hatch. A positive pressure is maintained between the double gaskets of the tube cover flange to establish the double barrier between the containment atmosphere and the inside of the fuel transfer tube. The interior of the fuel transfer tube is not pressurized. Seal water injection is not required for this penetration. The following lines would be subjected to pressure in excess of the isolation valve seal water system pressure (~52 psig) in the event of an accident, due to operation of the safety injection system recirculation pumps:

1. Residual heat removal loop inlet line.

2. Residual heat removal loop outlet line.

3. Bypass line from residual heat exchanger outlet to safety injection pumps suction.

4. Residual heat removal pumps mini-flow line.

5. Residual heat removal loop sample line.

6. Recirculation pump discharge sample line.

These lines are isolated by double disk gate valves or double valves, which can be sealed by nitrogen gas from the high pressure nitrogen supply of the isolation valve seal water system. A self-contained pressure regulator operates to maintain the nitrogen injection pressure slightly higher than the maximum expected line pressure. These valves are closed or intermittently operated during reactor operation, and the nitrogen gas injection is manually initiated. Lines, which are capable of communicating with the containment atmosphere (normally filled with air or vapor) include:

1. Steam jet air ejector return line to containment.

2. Containment radiation monitor inlet and outlet lines.

In an accident condition the space between the two containment isolation valves in each line are sealed by pressurizing with air from the weld channel and penetration pressurization system. The air is introduced into each space at approximately 2 psi above the containment design pressure through a separate line from the weld channel and penetration pressurization system. Parallel, redundant, fail-open valves in each injection line open on the appropriate containment isolation signal to provide a reliable supply of pressurizing air. A flow-limiting orifice in each injection line prevents excessive air consumption if one of these valves spuriously fails to open, or if one of the containment isolation valves fails to respond to the "trip" signal. 5.2.2.7 Class 7, Steam and Feedwater Lines These lines and the shell side of the steam generator are considered basically as an extension of the containment boundary and as such must not be damaged as a consequence of reactor coolant system damage. This requires that the steam generator shell, feed and steam lines within the containment are to be classified and designed for the reactor coolant system missile-protected category. The reverse is also true in that a steam break is not to cause damage to the reactor coolant system. 5.2.3 Isolation Valves And Instrumentation Diagrams Figures 5.2-1 through 5.2-28 show all valves in lines leading to the atmosphere or to closed systems on both sides of the containment barrier, valve actuation and preferential failure Chapter 5, Page 71 of 89 Revision 22, 2010

IP2 FSAR UPDATE modes, the application of "trip" (containment isolation) signals, relative location of the valves with respect to missile barriers, and the boundaries of seismic Class I designed lines. The item numbers in these figures align with the item numbers in Table 5.2-1. Figure 5.2-29 defines the nomenclature and symbols used. 5.2.4 Valve Parameters Tabulation A summary of the fluid systems lines penetrating containment and the valves and closed systems employed for containment isolation is presented in Table 5.2-1. Each valve is described as to type, operator, position indication and open or closed status during normal operation, shutdown and accident conditions. Information is also presented on valve preferential failure mode, automatic trip by the containment isolation signal, and the fluid carried by the line. Containment isolation valves are provided with actuation and control equipment appropriate to the valve type. For example, air-operated globe and diaphragm (Saunders Patent) valves are generally equipped with air diaphragm operators, with fail-safe operation provided by the control devices in the instrument air supply to the valve. Motor-operated gate valves are capable of being supplied from reliable onsite emergency power as well as their normal power source. Manual and check valves, of course, do not require actuation or control systems. The containment isolation system is brought into service by one of three conditions: phase A isolation signal, phase B isolation signal, or containment ventilation isolation signal. The automatically tripped isolation valves are actuated to the closed position by any of these isolation signals. The first of these signals is derived in conjunction with safety injection actuation, and trips the majority of the automatic isolation valves. These are valves in the so-called "nonessential" [Note - "Nonessential" process lines are defined as those, which do not increase the potential for damage to in-containment equipment when isolated. "Essential" process lines are those providing cooling water and seal water flow for the reactor coolant pumps. These services should not be interrupted unless absolutely necessary while the reactor coolant pumps are operating.] process lines penetrating the containment. This is defined as "phase A" isolation and the trip valves are designated by the letter "T" in the isolation diagrams, Figures 5.2-1 through 5.2-29. This signal also initiates automatic seal water injection (See Section 6.5). The second, or "phase B", containment isolation signal is derived upon actuation of the containment spray system, and trips the automatic isolation valves in the so-called "essential"* process lines penetrating the containment. These trip valves are designated by the letter "P" in the isolation diagrams. Containment ventilation isolation represents closing of the three ventilation lines to the containment and will be automatically activated by high containment radioactivity, a phase A isolation signal, or automatic containment spray (and associated phase B) actuation; see Section 5.3 for further information on the containment, heating, cooling and ventilation system. A manual containment isolation signal can be generated from the control room for either phase A or phase B isolation. These signals perform the same functions as the automatically derived signals. The containment ventilation isolation signal can be manually activated by a manual safety injection signal, a manual phase A containment isolation signal, or a manual containment spray signal. Chapter 5, Page 72 of 89 Revision 22, 2010

IP2 FSAR UPDATE Nonautomatic isolation valves, i.e., remote stop valves and manual valves, are used in lines, which must remain in service, at least for a time, following an accident. These are closed manually if and when the lines are taken out of service. Standard closing times available with commercial valve models are adequate for the sizes of containment isolation valves used. Valves equipped with air-diaphragm operators generally close in approximately 2 sec. The typical closing time available for large motor-operated gate valves is 10 sec. Closing times of greater than 10 seconds are permitted on a case by case basis if properly justified by an individual valve evaluation. The large butterfly valves used to isolate the containment ventilation purge ducts are equipped with air piston operators and spring returns capable of closing the valves. The butterfly valves used to isolate the 10-in. pressure relief line are equipped with air piston operators each with a separate accumulator air supply on each valve capable of closing the valves. These valves all fail to the closed position on loss of control signal or instrument air. Allowable closure time for these valves is less than or equal to 3 seconds. 5.2.5 Valve Operability All containment isolation valves, actuators, and controls are located so as to be protected against missiles that could be generated as the result of a loss-of-coolant accident. Only valves so protected are considered to qualify as containment isolation valves. Only isolation valves located inside containment are subject to the high-pressure, high-temperature, steam-laden atmosphere resulting from an accident. Operability of these valves in the accident environment is ensured by proper design, construction, and installation, as reflected by the following considerations:

1. All components in the valve installation, including valve bodies, trim and moving parts, actuators, instrument air and control and power wiring, are constructed of materials sufficiently temperature resistant to be unaffected by the accident environment. Special attention is given to electrical insulation, air operator diaphragms and stem packing material.

2. In addition to normal pressures, the valves are designed to with-stand maximum pressure differentials in the reverse direction imposed by the accident conditions.

This criterion is particularly applicable to the butterfly-type isolation valves used in the containment purge lines. Valve actuators are installed on these butterfly valves and travel is limited to a maximum of 60 degrees to ensure that the valves will be able to close against the maximum calculated design-basis accident pressure of 47 psig. An adjustable position setting on the actuators allows the valves to be opened to a full 90-degree position when containment integrity is not required. Chapter 5, Page 73 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 1 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks 1 Pressurizer 5.2-1 549 Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) No G Hot relief tank to 548 Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) gas analyzer RCS

2. Pressurizer 5.2-1 518 Check - No Closed Closed Closed - - - Yes/No G Cold relief tank N2 3418 Globe Sole. No Open Open Op/Cl FC - -

supply tank 3419 Globe Sole. No Open Open Op/Cl FC - - RCS 4136 Dia. Manual No Closed/OI Closed/OI Op/Cl - - -

3. Pressurizer 5.2-1 552 Dia Air Yes Closed Closed Closed FC T Water (A) No W Cold relief tank 519 Dia Air Yes Closed Closed Closed FC T Water (A) makeup -

RCS

4. Residual heat 5.2-2 741A1A Check - No Closed Open Open - - RHR Yes W Hot removal 7441A DDV Motor Yes Open Open Op/CL FAI - Nit (M) May be closed return depending on

     - ACS/SIS                                                                                                                             accident condition

5. Resid. Heat 5.2-2 888A DDV Motor Yes Closed/OI Closed Op/Cl FAI - Nit (M) Yes W Hot**

removal loop 888B DDV Motor Yes Closed/OI Closed Op/Cl FAI - Nit (M) to - S.I.pumps - ACS/SIS To sampling 5.2-2 958 Globe Motor No Closed/OI Closed Closed FAI - Nit (M) Yes W Hot May be used during system 959 Globe Motor No Closed/OI Closed Closed FAI - Nit (M) Yes W shutdown and after ACS/SS 990D Globe Manual No LC/OI LC LC - - Nit (M) No - - accident RHR pump 5.2-2 1870 Globe Motor Yes Open Open Op/Cl FAI - Nit (M) Yes/No W Hot mini-flow line 743 Globe Motor Yes Open Open Op/Cl FAI - Nit (M) Yes/No W Chapter 5, Page 74 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 2 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

6. Residual heat 5.2-2 7321A DDV Manual No LC/OI Open Closed - - Nit (M) No W Hot removal loop-out - ACS

7. Containment 5.2-2 885A DDV2A Motor Yes Closed/OI Closed Closed2 FAI - RHR No2 W Cold 2. Normally closed but sump 885B DDV2A Motor Yes Closed/OI Closed Closed2 FAI - RHR may be opened after recirculation - accident if normal ACS/SIS recirculation path from recirculation pump not available 2A. The upstream disc (nearest containment) of 885A and the downstream disc (RHR Loop side) of 885B have a 3/16" hole to prevent pressure locking

8. Letdown line 5.2-3 201 Globe Air Yes Open Open Closed FC T Water (A) No W Hot

    - CVCS                202      Globe    Air    Yes     Open        Open      Closed   FC        T     Water (A)

9. Charging line 5.2-3 205 Gate Motor No Open Open Op/Cl FAI - Water (M) Yes* W Cold

• May be used

    - CVCS                226      Globe  Motor    No      Open        Open      Op/Cl    FAI       -     Water (M) Yes*                     depending on accident.

227 Globe Motor No Closed/OI Closed Op/Cl FAI - Water (M) Yes*

10. Reactor 5.2-4 250ABCD Globe Motor No Open Open Op/Cl FAI - Water (M) Yes3 W Cold 3. Manual isolation if coolant pump 4925, Globe Motor No Open Open Op/Cl FAI - Water (M) Yes3 W Cold and when pumps are seal-water 4926, stopped.

supply lines 4927, (4) - CVCS 4928

11. Reactor 5.2-4 222 DDV Motor Yes Open Open Closed FAI P Water (A) No W Cold coolant pump seal water return - CVCS

12. Reactor 5.2-5 956E Globe Motor Yes Op/Cl Op/Cl Closed FAI T Water (A) Yes4 W Hot 4. Used to take coolant 956F Globe Motor Yes Op/Cl Op/Cl Closed FAI T Water (A) postaccident RCS sample line - samples SS

13. Fuel transfer 5.2-5 A Blind - No Closed - - - - Air5 No W Cold Flange is double tube - FHS flange gasketed in refuel-ing canal (missile protected).

5. Normally seal with air (WCPPS)

Chapter 5, Page 75 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 3 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Posit. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

14. Containment 5.2-6 869A,B DDV Motor No Open Op/Cl Op/Cl FAI - Water (M) Yes W Cold spray 867A,B Check - No Closed Closed Op/Cl - - -

headers (2) - 878A Globe Manual No LC/OI Closed Closed - - - SIS

15. Safety 5.2-7 850A DDV Motor No Open Open Op/Cl FAI - Water(M) Yes W Hot**

injection 851A DDV Motor Yes Open Open Op/Cl FAI - Water(M) Yes W Hot** headers (2) - 851B DDV Motor Yes Open Open Op/Cl FAI - Water(M) Yes W Hot** SIS 850B DDV Motor No Open Open Op/Cl FAI - Water(M) Yes W Hot**

16. Safety 5.2-7 859A Globe Manual No. LC/OI Closed Closed - - Water (A) No W Cold injection test 859C Globe Manual No LC/OI Closed Closed - - Water (A) line - SIS

17. Accumulator/ 5.2-8 4312 Check - No Closed Closed Closed6 - - - No G Cold 6. Could be OPS N2 863 Globe Air Yes Op/Cl Op/Cl Closed6 FC T - No opened depending supply - SIS on type of accident

18. Accumulator 5.2-8 956G Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) No W Cold Valves A and B sample - SS 956H Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) opened intermittently to take sample

19. Primary 5.2-9 1786 Dia Air Yes Open Closed Closed FC T Water (A) No G Hot system vent 1787 Dia Air Yes Open Closed Closed FC T Water (A) header and 3416 Globe Sole. No Open Open Op/Cl FC - - Yes/No G Hot N2 supply line 3417 Globe Sole. No Open Open Op/Cl FC - -

    - WDS                 5459     Dia  Manual    No   Closed/OI    Closed     Op/Cl     -        -          -

1616 Check - No Closed Closed Op/Cl - - -

20. Reactor 5.2-9 1788 Dia Air Yes Op/Cl Op/Cl Closed FC T Water (A) No G Hot Valves opened coolant drain 1789 Dia Air Yes Op/Cl Op/Cl Closed FC T Water (A) intermittently tank to gas analyzer -

WDS

21. RCDT pump 5.2-9 1702 Dia Air Yes Open Op/Cl Closed FC T Water (A) No W Cold Valves open discharge - 1705 Dia Air Yes Open Op/Cl Closed FC T Water (A) intermittently WDS Chapter 5, Page 76 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 4 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

22. Reactor 5.2-10 797 DDV Motor Yes Open Op/Cl Closed FAI P Water (A) No7 W Cold 7. Could be used coolant pump depending on the type cooling water of accident in - ACS

23. Reactor 5.2-10 784 DDV Motor Yes Open Op/Cl Closed FAI P Water (A) No8 W Cold 8. Could be used coolant pump depending on the type water out of accident (6) - ACS

24. Reactor 5.2-10 FCV-625 DDV Motor Yes Open Op/Cl Closed FAI P Water (A) No9 W Cold 9. Could be used coolant pump depending on the type water out of accident (3) - ACS

25. Resid. Heat 5.2-11 CS - - - - - - - - - Yes W Hot Residual heat exch. Cooling exchanger and water in - associated component ACS cooling lines are a missile protected closed system CS - - - - - - - - - Yes W Hot Component cooling system closed

26. Resid. Heat 5.2-11 822A1B Gate Motor Yes Closed Open Open FAI - - Yes W Cold exch. Cooling 822B1B Gate Motor Yes Closed Open Open FAI - - Yes W Cold water return CS - - - - - - - - - Yes W Cold Component cooling

     - ACS                                                                                                                                  system closed

27. Recir. Pump 5.2-12 753H1B Gate Manual No Open Open Op/Cl - - - Yes W Cold May be closed cooling water depending on accident supply - ACS condition CS - - - - - - - - - Yes W Cold Component cooling system closed

28. Recir. Pump 5.2-12 753G1B Gate Manual No Open Open Op/Cl - - - Yes W Cold May be closed cooling depending on accident heater return condition

     - ACS CS        -      -       -       -         -            -      -        -          -      Yes     W      Cold     Component cooling system closed

29. Excess 5.2-13 791 Dia Air Yes Closed Closed Closed FC T Water (A) No W Cold letdown heat 798 Dia Air Yes Closed Closed Closed FC T Water (A) exchanger cooling water in - ACS Chapter 5, Page 77 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 5 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

30. Excess letdown 5.2-13 796 Globe Air Yes Closed Closed Closed FC T Water (A) No W Cold heat exchanger 793 Dia Air Yes Closed Closed Closed FC T Water (A) cooling water out - ACS

31. Containment 5.2-13 1728 Dia Air Yes Open Open Closed FC T Water (A) No W Cold sump pump 1723 Dia Air Yes Open Open Closed FC T Water (A) discharge - WDS 31a. Sampling system 5.2-13 5132 Globe Motor Yes Closed Closed Cl/Op FAI T Water (A) No/Yes10 W Cold 10. Can be used to return return - WDS 4399 Globe Motor Yes Closed Closed Cl/Op FAI T Water (A) No/Yes10 W Cold highly radioactive water to containment after post-accident analysis of sampling system.

32. Containment air 5.2-14 PCV-1234 Dia Air Yes Open Open Closed FC T Air (A) No11 G Cold 11. May be opened for air sample in - rad. PCV-1235 Dia Air Yes Open Open Closed FC T Air (A) No11 G Cold sampling following accident mon. when the containment pressure is below 5 psig

33. Containment air 5.2-14 PCV-1236 Dia. Air Yes Open Open Closed FC T Air (A) No12 G Cold 12. May be opened for air sample out - PCV-1237 Dia. Air Yes Open Open Closed FC T Air (A) No12 G Cold sampling following accident rad. mon. when the containment pressure is below 5 psig

34. Air ejector 5.2-14 PCV-1229 Globe Air Yes Closed Closed Closed FC T Air (A) No G Cold discharge to PCV-1230 Globe Air Yes Closed Closed Closed FC T Air (A) No G Cold containment sec sys

35. Main steam - CS - - - - - - - - - - - Hot Steam generators headers13 13. (four penetrations)

Chapter 5, Page 78 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 6 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Acid. Water Hot/Cold Remarks

36. Main - CS - - - - - - - - - - - Hot Steam generators (four feedwater penetrations) headers

37. Steam 5.2-15 PCV1214 Globe Air Yes Open Op/Cl Closed FC T Water (A) No W Hot *(four penetrations) generator PCV1215 blowdown/sa PCV1216 mple sec. sys. PCV1217 PCV1214A Globe Air Yes Open Op/Cl Closed FC T Water (A) No W Hot PCV1215A PCV1216A PCV1217A

38. S.G. System deleted blowdown sample

39. Ventilation 5.2-16 SWN-41 BV Motor No Open Open Op/Cl FAI - SWS Yes W Cold Fan cooler units -

system water SWN-42 Relief - No Closed Closed Closed - - SWS - - - missile protected, cooling water SWN-43 Gate(2) Manual No LC/OI Closed Op/Cl - - SWS Yes - - closed system in - SWS14 Globe(3) 14. (five penetrations)

40. Ventilation 5.2-16 SWN-44 BV Motor No Open Open Op/Cl FAI - SWS Yes W Cold Fan cooler units -

system motor missile protected, cooling water closed system out - SWS15 5216 SWN-51 Globe Motor No Open Open Op/Cl FAI - SWS Yes W Cold 15. (Five penetrations) 40a. Ventilation 5216 SWN-71 Globe Motor No Open Open Op/Cl FAI - SWS Yes W Cold 16. Five penetrations system motor cooling water out - SWS16

41. Service air 5.2-17 SA-24 Dia Manual No LC/OI LC LC - - Water (A) No G Cold SA-24-1 Dia Manual No LC/OI LC LC - - Water (A)

42. Not assigned Chapter 5, Page 79 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 7 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

43. Weld channel 5.2-17 PCV1111-11B Ball Manual No Open Open Op/Cl - - - Yes G Cold 17. Two pressurization PCV1111-21B Ball - - penetrations air supply CS - - - - - - - - - penetration press (PPS)17 system

44. Spare 5.2-17 580A Needle Manual No LC/OI LC LC - - - No G Cold Penetration 580B Needle Manual No LC/OI LC LC - - - capped inside containment and outside containment downstream of valve 580B

45. Auxiliary steam 5.2-18 UH-43 DDV Manual No LC/OI Closed18 LC - - Water (A) No G Hot 18. May be supply opened during shutdown for cont. heating

46. Auxiliary steam 5.2-18 UH-44 DDV Manual No LC/OI Closed19 LC - - Water (A) No W Hot 19. May be supply opened during condensate shutdown for return cont. heating

47. City water 5.2-18 MW-17 Gate Manual No LC/OI Closed20 LC - - Water (A) No W Cold 20. May be MW-17-1 Gate Manual No LC/OI Closed20 LC - - Water (A) opened during shutdown for maintenance or fire protection purposes

48. Purge supply 5.2-19 FCV-1170 BV Air Yes Closed21 Open Closed FC CVI Air (A) No G Cold 21. May be open duct in - vent. FCV-1171 BV Air Yes Closed21 Open Closed FC CVI Air (A) for safety related sys. purging, or to facilitate safety related surveillance or maintenance.

49. Purge exhaust 5.2-19 FCV-1172 BV Air Yes Closed22 Open Closed FC CVI Air (A) No G Cold 22. May be open duct out - FCV-1173 BV Air Yes Closed22 Open Closed FC CVI Air (A) for safety related vent. sys. purging, or to facilitate safety related surveillance or maintenance.

50. Containment 5.2-19 PCV-1190 BV Air Yes Closed23 Closed Closed FC CVI Air (A) No G Cold 23. Opened pressure relief PCV-1191 BV Air Yes Closed23 Closed Closed FC CVI Air (A) intermittently for

    - vent                 PCV-1192      BV     Air   Yes     Closed23     Closed   Closed   FC       CVI    Air (A)                              pressure relief.

51. Recirculation 5.2-20 990A Globe Motor Yes Closed Closed Op/Cl FAI T Nit (M) No/Yes24 W Cold 24. Used pump 990B Globe Motor Yes Closed Closed Op/Cl FAI T Nit (M) periodically after discharge accident to sample line sample recirculation fluid.

Chapter 5, Page 80 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 8 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

52. Pressurizer 5.2-20 956A Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) No/Yes25 W Hot 25. Could be used for steam space 956B Globe Air Yes OP/Cl OP/Cl Closed FC T Water (A) No/Yes25 W Hot taking postaccident sample line samples.

53. Pressurizer 5.2-20 956C Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) No/Yes26 W Hot 26. Could be used for liquid space 956D Globe Air Yes Op/Cl Op/Cl Closed FC T Water (A) No/Yes26 W Hot taking postaccident sample line samples.

54. Containment 5.2-21 1814A Globe Manual No LO Open Op/Cl - - - Yes G Cold

55. Pressure 1814B

56. Instrumentation 1814C CS - - - - - - - - - - - -

57. Postaccident 5.2-22 SOV-5018 Globe Sole. Yes Closed Closed Both26a FC - - Yes G Cold 26a. Isolation valves containment SOV-5020 Globe Sole. are opened sampling SOV-5022 Globe Sole. intermittently after an system supply SOV-5024 Globe Sole. accident.

and return lines SOV-5019 Globe Sole. Yes Closed Closed Both26a FC - - Yes G Cold (7) SOV-5021 Globe Sole. SOV-5023 Globe Sole. SOV-5025 Globe Sole.

58. Spare 5.2-23

59. Spare 5.2-24

60. Spare

61. Spare

62. Spare

63. Not assigned Chapter 5, Page 81 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 9 of 10) Posit. Valve Indic. Posit. Fluid No. or In Posit. Posit. On Cont. Testing/ Used Gas Item Penetration Dia- Closed Valve Oper. Cont. Normal During After Power Isolation Sealing After or Penetration No. and System gram System Type Type Room Posit. Shutdown Accident Fail Trip Method1 Accid. Water Hot/Cold Remarks

64. Instrument/ai 5.2-25 IA-39 Check - No Open Open Both - - - Yes/No27 G Cold 27. Could be used to r postaccident PCV1228 Dia Air Yes Open Open Both FC T - Yes/No27 G Cold resupply instrument venting air to containment supply post-accident.

65. Postaccident 5.2-25 E-2 Dia Air No Closed/OI Closed Both FC - Air (A) Yes/No28 G Cold 28. Could be used venting E-1 Dia Air No Closed/OI Closed Both FC - Air (A) Yes/No28 G Cold after accident if exhaust line E-3 Dia Air No Closed/OI Closed Both FC - Air (A) Yes/No28 G Cold containment venting E-5 Dia Air No Closed/OI Closed Both FC - Air (A) Yes/No28 G Cold were deemed necessary.

66. Deleted

67. Containment 5.2-26 A Blind - No Closed Closed Closed - - - No. Gas Cold 30. Two penetrations leak test air Flange line30 B Blind - No Closed Closed Closed - - - No Gas Cold Flange W/Test Conn.

68. Equipment - CS - - - - - - - - Air (A) No - -

access

69. Personnel air 5.2-27 85A, 95A31 Ball Interlk Yes Closed Closed Closed - - Air (A) No Gas Cold 31. 85A & 95A may be lock (2) 85B, 95B31 Ball w/door Air (A) open when 85B & 95B 85C, 95C Spring - No Closed Closed Closed - - - No Gas Cold are closed. 85B & 95B check may be open when 85D, 95D Spring - No Closed Closed Closed - - - No Gas Cold 85A & 95A are closed.

check

70. Steam 5.2-28 IIP-500 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold 32. Depending on generator IIP-501 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold accident type level, IIP-502 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold pressurizer IIP-503 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold level, and IIP-504 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold pressure IIP-505 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold pneumatic IIP-506 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold indication IIP-507 Globe Manual No Closed/OI Closed Both - - - Yes32 - Cold lines (4)

Chapter 5, Page 82 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.2-1 Containment Piping Penetrations and Valving (Sheet 10 of 10) Key: FAI - fail as is RCS - reactor coolant system Hot - insulated & cooled penetration FC - fail closed ACS - auxiliary coolant system Cold - standard piping or equipment penetration FO - fail open WDS - waste disposal system Hot** - insulated non cooled penetration LC - locked closed SIS - safety injection system IVSWS -isolation valve seal water system LO - locked open SS - sampling system WCPPS -weld channel pressurization system BV - butterfly valve CVCS - chemical and volume control system RHR -residual heat removal system DDV - double disk gate valve Vent - ventilation system Dia - diaphragm valve SWS - service water system T - containment isolation signal - phase A FH - fuel handling P - containment isolation signal - phase B PPS - penetration pressurization system A - automatic CVI - containment ventilation isolation signal M - manual CS - closed system Op/Cl - open/closed Nit - nitrogen OI - may be opened intermittently to support plant operations Notes:

1. Sealing Methods and Type C Testing:

• For valves sealed by IVSWS water (designated Water (A) or Water (M)), minimum Type C test pressure is 52 psig.
• For valves sealed by IVSWS nitrogen (designated Nit (M)), minimum Type C test pressure is 47 psig.
• For valves sealed by WCPPS (designated Air (A)), minimum Type C test pressure is 47 psig.
• For valves sealed by RHR system fluid (designated RHR), minimum Type C test pressure is 52 psig (valves 741A,885A,885B).
• For valves sealed by service water system (designated SWS), minimum Type C test pressure is 52 psig (valve series SWN-41, SWN-42, SWN-43, SWN-44, SWN-51, SWN-71). Either the A or B valve(s) may serve as the required containment isolation valve(s) for the SWN-41, SWN-44 and SWN-71 series. Designation of the B valve(s) in the SWN-44 series requires the codesignation of the SWN-51 valves associated with the penetration(s) as an additional required containment isolation valve(s) (see Figure 5.2-16).
• For all other isolation valves not sealed by a system, gas (ie. Nitrogen or air) is the test medium at a minimum Type C pressure of 47 psig.

1A. These valves testable only when at cold shutdown (741A, 744, 732). 1B. These valves are excluded from Type C testing per License Amendment N0. 63, dated August 28, 1980 (822A, 822B, 753G, 753H, PCV-1111-1, PCV-1111-2). Chapter 5, Page 83 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.2 FIGURES Figure No. Title Figure 5.2-1 Containment Isolation System Penetration Schematics Figure 5.2-2 Containment Isolation System Penetration Schematics Figure 5.2-3 Containment Isolation System Penetration Schematics Figure 5.2-4 Containment Isolation System Penetration Schematics Figure 5.2-5 Containment Isolation System Penetration Schematics Figure 5.2-6 Containment Isolation System Penetration Schematics Figure 5.2-7 Containment Isolation System Penetration Schematics Figure 5.2-8 Containment Isolation System Penetration Schematics Figure 5.2-9 Containment Isolation System Penetration Schematics Figure 5.2-10 Containment Isolation System Penetration Schematics Figure 5.2-11 Containment Isolation System Penetration Schematics Figure 5.2-12 Containment Isolation System Penetration Schematics Figure 5.2-13 Containment Isolation System Penetration Schematics Figure 5.2-14 Containment Isolation System Penetration Schematics Figure 5.2-15 Containment Isolation System Penetration Schematics Figure 5.2-16 Containment Isolation System Penetration Schematics Figure 5.2-17 Containment Isolation System Penetration Schematics Figure 5.2-18 Containment Isolation System Penetration Schematics Figure 5.2-19 Containment Isolation System Penetration Schematics Figure 5.2-20 Containment Isolation System Penetration Schematics Figure 5.2-21 Containment Isolation System Penetration Schematics [Replaced with Plant Drawing 235296] Figure 5.2-22 Containment Isolation System Penetration Schematics Figure 5.2-23 Containment Isolation System Penetration Schematics Figure 5.2-24 Containment Isolation System Penetration Schematics Figure 5.2-25 Containment Isolation System Penetration Schematics Figure 5.2-26 Containment Isolation System Penetration Schematics Figure 5.2-27 Containment Isolation System Penetration Schematics Figure 5.2-28 Containment Isolation System Penetration Schematics Figure 5.2-29 Containment Isolation System Penetration Schematics Chapter 5, Page 84 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.3 CONTAINMENT HEATING, COOLING AND VENTILATION SYSTEM 5.3.1 Design Basis 5.3.1.1 Performance Objectives The containment heating, cooling and ventilation systems are designed to accomplish the following:

1. Remove the normal heat loss from all equipment and piping in the reactor containment during plant operation and to maintain a normal ambient temperature of 130oF or less.

2. Provide sufficient air circulation and filtering of iodine throughout all containment areas to permit safe and continuous access to the reactor containment within two hours after reactor shutdown assuming defects exist in 1-percent of the fuel rods.

3. Provide for positive circulation of air across the refueling water surface to assure personnel access and safety during shutdown.

4. Provide containment heating, if required, to assure a minimum containment ambient temperature of 50°F before the reactor is taken above the cold shutdown condition.

5. Provide for purging of the containment vessel to the plant vent for dispersion to the environment. The rate of release does not permit offsite dose to exceed Offsite Dose Calculation Manual (ODCM).

6. Provide for depressurization of the containment vessel following an accident.

The postaccident design and operating criteria are detailed in Section 6.4.

7. Provide for continuous pressure relief via an exhaust system.

In order to accomplish these objectives the following systems are provided:

1. Containment recirculation cooling system

2. Control rod drive mechanism cooling system

3. Containment purge and pressure relief system

4. Containment auxiliary charcoal filter system

5. Steam heating system 5.3.1.2 Design Characteristics-Sizing The design characteristics of the equipment required in the containment for cooling, filtration and heating to handle the normal thermal and air cleaning loads during normal plant operation are presented in Table 5.3-1. In certain cases where engineered safeguards functions also are served by the equipment, component sizing is determined from the heavier duty specifications associated with the design basis accident detailed further in Section 6.4.

Chapter 5, Page 85 of 89 Revision 22, 2010

IP2 FSAR UPDATE 5.3.2 System Design 5.3.2.1 Piping and Instrumentation Diagram The containment ventilation, purging, and recirculation cooling and filtration systems flow diagram is shown in Plant Drawing 9321-4022 [Formerly UFSAR Figure 5.3-1]. The containment ventilation systems and main plant vent are designed as Class I structures. 5.3.2.2 Containment Cooling and Ventilation System Air recirculation cooling during normal operation is accomplished using air handling units discharged into a common header ductwork distribution system to ensure adequate flow of cooled air throughout the containment. The cooling coils in each air handling unit transfer up to 61.7 x 106 Btu/hr in the event of an accident when supplied with approximately 1600 gpm cooling water at 95°F inlet temperature and steam-air flowrate of 64,500 cfm. Each air-handling unit consists of the following equipment arranged so that, during normal and accident operation, air flows through the unit in the following sequence: cooling coils, moisture separators (demisters), centrifugal fan with direct-drive motor, and distribution header. The fans and motors of these units are equipped with vibration sensors to detect abnormal operating conditions in the early stages of the disturbance. The normal air flow rate per air-handling unit is approximately 70,000 cfm. Section 6.4.2 provides additional information on the operation of this system. The following additional systems supplement the main containment recirculation system:

1. Control rod drive cooling system consisting of fans and ductwork to circulate air through the control rod drive mechanism shroud and discharge it to the main containment volume. Four direct driven axial flow fans are provided for use.

There are two power supplies for each fan.

2. Two unit steam heaters are located in containment to provide additional area heating as required. The containment purge supply is also provided with steam pre-heating to supplement containment heating as required.

5.3.2.3 Containment Purge System The containment purge system is independent of the primary auxiliary building exhaust system, (except for the common exhaust fans) and includes provisions for both supply and exhaust air. The supply system includes roughing filters, heating coils, fan, supply penetration with two butterfly valves for bubble tight shutoff, and a purge supply distribution header inside containment. The exhaust system includes exhaust penetration with two butterfly valves identical to those above, exhaust ductwork, filter bank with roughing, HEPA and charcoal filters, fans and exhaust vent. The purge supply and exhaust flow rates are nominally 23,000 cfm and 25,000 cfm respectively. The quick closing purge isolation valves close upon receipt of an accident signal. Allowable closure times for these valves are specified in Section 5.2.4. During power operation, the purge system is routinely not operated. Prior to purging the containment air, particulate and gas monitor indications of the closed containment activity levels will be used to guide routine releases from the containment. During power operation, the containment air particulate and gas monitor indications will help determine desirability of using Chapter 5, Page 86 of 89 Revision 22, 2010

IP2 FSAR UPDATE either one or both of two auxiliary particulate and charcoal filter units installed in the containment primarily for preaccess cleanup. When containment purging is in progress for access following reactor shutdown, releases from the plant vent are continuously monitored with a gas monitor, as described in Section 11.2. 5.3.2.4 Purge System Isolation Valves The purge supply and exhaust duct butterfly valves, both inside and outside the containment, are normally closed during power operation. They may be opened for safety related reasons, i.e., pressure control or to facilitate safety related surveillance or maintenance. The opening angle is limited during operation so that the valves can close against a differential pressure (see Section 5.2.5). The spaces between the closed valves are pressurized with air by the weld channel and penetration pressurization system. The valves are designed for rapid automatic closing by the containment isolation signal (derived from any safety injection signal), upon a signal of high activity level within the containment in the event of a radioactivity release when the purge line is open, or upon a manually initiated signal. To ensure optimum sealing of the resilient valve seats, the two valves located outside containment are enclosed and a minimum ambient temperature is maintained. 5.3.2.5 Containment Pressure Relief Line The normal pressure changes in the containment during reactor power operation, and during plant cooldown if the containment purge system is not operating, will be handled by the containment pressure relief line. This line is equipped with three quick-closing butterfly type isolation valves, one inside and two outside the containment. The valves will be automatically actuated to the closed position by the containment isolation signal, by a containment high radioactivity signal, or upon a manually initiated signal. Each of these air operated valves is equipped with an accumulator to assure each can close even if the air supply is lost. The two intra-valve spaces are pressurized with air by the Weld Channel and Penetration Pressurization System when the valves are closed. The pressure relief line discharges to the plant vent. The opening angle of the pressure relief valves is limited during operation so the valves can close against a differential pressure. 5.3.2.6 Containment Purge and Pressure Relief Isolation Reset Opening of the purge and pressure relief isolation valves following an isolation signal requires deliberate operator action by resetting all isolation signals and depressing both Containment Ventilation Isolation reset push buttons. Further, in order to reset the Containment Ventilation Isolation signal for Train B, the control switches for the purge and pressure relief isolation valves in Train B must first be placed in the closed position. In addition, guard plates are placed over the reset buttons. These three features preclude the possibility of inadvertently opening these valves. REFERENCES FOR SECTION 5.3

1. Deleted Chapter 5, Page 87 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.3-1 (Sheet 1 of 2) Containment Cooling and Ventilation System - Principal Component Data Summary Units Required for System Units Installed Unit Capacity Normal Operation Containment cooling and recirculation Demister 5 72,500 cfm 52 Cooling coils - normal 5 2.20 x 106 Btu/hr1 52 Cooling coils - DBA 5 61.7 x 106 Btu/hr3 Fans 5 72,500 cfm 52 Fan pressure - 7.21-in. H2O (Note 7) Fan motors 5 350 hp 52 (440 V, 3-phase) Control rod drive mechanism cooling Fans, standard 4 15,000 cfm 3 conditions Fan pressure - 5.5-in. H2O Fan motors 4 30 hp 3 Reactor compartment cooling Part of containment

                                  -                    12,000 cfm recirculation system Refueling canal air sweep Part of containment
                                  -                    17,500 cfm recirculation system Purge supply                  1                   23,000 cfm6              Optional Fan pressure                  -                    2.5-in. H20 Fan motors                  1                       40 hp 25 psig steam preheat 1 set                3 x 106 Btu/hr            Optional coils Air filters, roughing            -                          -                    1 Chapter 5, Page 88 of 89 Revision 22, 2010

IP2 FSAR UPDATE TABLE 5.3-1 (Sheet 2 of 2) Containment Cooling and Ventilation System - Principal Component Data Summary Units Required for System Units Installed Unit Capacity Normal Operation Purge exhaust Fans, standard 24 55,500 cfm6 Optional conditions Fan pressure - 10.3-in. H2O Optional Fan motors 2 125 hp Optional Plenums 1 -5 Optional HEPA filters - - Optional HECA filters 1 - Optional (charcoal adsorbers) Containment auxiliary charcoal filter Fans, standard 2 8,000 cfm Optional conditions Fan pressure - 5.0-in. H20 Optional Fan motors 2 10 hp Optional Filters and charcoal 2 8,000 cfm Optional filters; roughing, HEPA Steam heating Heaters, 2 400,000 Btu/hr each Optional 25 psig steam Notes:

1. This value reflects the increase in air side flow rate due to removal of the original plant HEPA Filters.

2. Depends on time of year and containment atmospheric temperature.

3. Based on minimum assumed performance at 271°F containment temp 95°F Service Water temp, 1600 gpm Service Water flow and 64,500 cfm air flow rate.

4. The two exhaust fans are used interchangeably or as backup for:

1. Ventilation of primary auxiliary building.

2. Containment building purge system.

5. Purge (25,000 cfm) and primary auxiliary building exhaust (55,500 cfm) are fed into a common plenum.

6. Purge supply (23,000 cfm) and purge exhaust (25,000 cfm) are the nominal, as built, flow rates for the purge system (+/- 10%).

7. At 72,500 cfm flow rate.

5.3 FIGURES Figure No. Title Figure 5.3-1 Containment Cooling and Ventilation System [Replaced with Plant Drawing 9321-4022] Chapter 5, Page 89 of 89 Revision 22, 2010

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(-10.9) For Section location, see Figure 5.1-44 INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1 -42 (THIS FIGURE DEPICTS ORIGINAL HISTORICAL DESIGN DATA AND DOES NOT REFLECT THE CURRENT CONFIGURATION OR ANALYSIS.) PUMP SUPPORT-SECTION 3-3 MIC. No. 1999MC3781 REV. No. 17A

ISOMETRIC Of iYPIC.J\L 5TEftJ-.,\ qENERt'X"rO~ { 5uPPORT ~ reP 6'-2 Or H.P. EL. L.P. EL. 15'-1 0" r--------------------, INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1 -43 ISOMETRIC VIEW-STEAM GENERATOR SUPPORT MIC. No. 1999MC3782 REV. No. 17A

ISOME.iR\c OF TYPIC 1\L REA.C.TORCOOl ANT PVMP

• SUPPoRT'

                ,2 1 TOP oF" suppoRT STRUCTURe:.

61'-11 3/4" INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1-44 ISOMETRIC VIEW-REACTOR COOLANT PUMP SUPPORT MIC. No. 1999MC3783 REV. No. 17A

MAXIMUM FORCES ACTING ON A REACTOR VESSEL SUPPORT A B C D I REACTOR PIPE PIPE ~RTH::. VESSEL RUAKE Z WEIGHT BREAK BREAK IfJ VERTICA A+B A+C A+D

                           & PIPING REACTION CASE II      CASE II pI RECT ION P (1b) 934,000      -        525.000      395.000  934,000  1,459,OOC 1.329.00(

R (1b) 322,000 - - - 322.000 322.000 322.00( T (lb) 140.000 1.187,OOC 710,000 969,000 ,327,000 850.00(J 1I-.109.00( PIPE BREAK CASE II

                                       /1'/,...

IS BETWEEN PUMP AND ELL ON REACTOR NOZZLE I I

                      ..... ~---+--

PIPE BREAK CASE III IS AT STEAM GENERATOR b INLET EARTHQUAKE IN Z T1 DIRECTION ACTS NORMAL TO REACTOR COOLANT OUTLET LINE R-RADIAL T-TANGENTIAL P-VERTICAL PLAN P INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1-45 (THIS FIGURE DEPICTS ORIGINAL HISTORICAL DESIGN DATA AND DOES NOT REFLECT THE CURRENT CONFIGURATION OR ANALYSIS.) MAXIMUM FORCES ACTING ON A REACTOR VESSEL SUPPORT MIC. No. 1999MC3784 REV. No. 17A

2A4MJ(5~, -~

                    , flir INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1 -46 PLAN VIEW 60 FT -  0 IN.

MIC. No. 1999MC3785 REV. No. 17A

         \    ZII aMK.c. 801 2.. IJMJ<C 801 2 Ilt;MJ:.C.~O 2 2J""MJ<(..8DZ TYP LAYER (i) REACTOR RING SCALE 1,(:    I~O 5ECTION                \-\

INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1-47 TYPICAL LAYER-REACTOR RING MIC. No. 1999MC3786 REV. No. 17A

N07E

• CAOWELD SPLJCf ,-iLL HOOPS 8EfORE. THE'! ARE SET IN PLACE INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1 -48 SECTION 5-5 MIC. No. 1999MC3787 I REV. No. 17 A

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                                           - -                                         INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.1-49 SECTION 18-1 8 MIC. No. 1999MC3788        I REV. No. 17A

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SECTIOIoJ C-C DEFORMATION MEASUREMENTS

                                   ;; ----         ------ri                                                                                                                                                 MIC, No,  1999MC3795  I REV. No. 171<

I I

Item 1 Pressurizer Relief Tank to Gas Analyzer Item 2 Pressurizer Relief Tank N2 Supply Item 3 Pressurizer Relief Tank Makeup TO VENT HEADER (ITEM 19) ASWI I S-1 I I'-I T I I 548 I 1549 /T I I GAS J-I--........ AN AL YZER cr: W N cr:

J (f)

(f) w cr: 0... f-r---411-- N2 SUPPLY I (100 PSIG) I S-1 I 4136 .--1 Res 1--r-9>L.....j~_ _-r--II~......I---<IIII__ MAK EU P I WATER I I I-I 519 I Z w I .-.-1 I~ I S-1 I- I z ASWI o u PRESSURIZER RELIEF TANK STANDPIPE MAKEUP ....- + - - - - - - - - - - - - ' Although the Pressurizer Relief Tank is missile protected, these penetrating lines can become exposed to containment atmosphere if the pressurizer discharge header is breached during the accident. INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-1 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3411 REV. No. 17 A VALVE POSITIONS ARE NOT CONTROUED BY THIS FIGURE

Item 4 Residual Heat Removal Return Item 5 Residual Heat Removal Loop To RWST Safety Injection Pumps Item 6 Residual Heat Removal Loop Out Item 7 Containment Sump Recirculation Line RCS ~---1.-,---l1 r---------l I RESIDUAL HE A T L~X_~Ii~~Q~~J I I l/1 741A 741 I 744 r------------l I RESIDUAL I HEAT REMOVALI~---..l L ~~_~~ I J N (IVSWS) 2 RECIRCULA TION PUMP N

                                                                              , 2 I

I I _ _ ..., _ _ ...J (IVSWS) I S.I.S. I I (CLOSED I I SYSTEM) I NOTE: 14" DRN Line No. 10 I has 6" thick insulation between valves 730 and 731 RCS --I~D-1--- __-----_at4l_--~____,_L=_'~ __+_-----'---_____1 II{I l----H-_ _ ---------+'~"=-=-:--~_=_==___-----.J CONTAINMENT/ SUMP NOTE: A 3/16" hole is drilled through CONTAINMENT the upstream disc (nearest MISSILE BARRIER containment) of 885A and the downstreom disc (RHR Loop Side) of 8858 to prevent pressure locking. INDIAN POINT UNIT No. 2 N2 - Manual N2 Pressurization UFSAR FIGURE 5.2-2 SS - Sampling System CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS ENTIRE SYSTEM SHOWN IS SEISMIC CLASS I DESIGN MIC. No. 1999MC3382 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 8 LETDOWN LINE PRT ASWI I T

                     )

(

                   \
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2010) I

                         \
                          )              200B                               I

( S i CVCS RCS T (CLOSED SYSTEM) REGEN. t MISSILE BARRIERi HX I 200C FROM RHR LOOP j I CONTAINMENT PRT - PRESSURIZER RELIEF TANK ITEM 9 CHARGING LINE REGEN. HEAT RCS 210B 204B AUX. SPRAY 211 1 1-- 9 RCS 374 210A 204A 1 MISSILE BARRIER CONTAINMENT --1 SLOPE INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-3 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3383 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 10 REACTOR COOLANT PUMP SEAL-WATER SUPPLY LINES

                                ---  f--_---+1~O~
                                                           ~

250,A,B,C,D MSWI I

        ~                    251,E,F,G,H      I    L                                241,A,B,C,D CVCS MISSILE BARRIERI I                      I    L SLOPE I---CONTAINMENT 4925,4926 4927,4928 ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DESIGN ITEM 11  REACTOR COOLANT PUMP SEAL-WATER RETURN RCS
                             -I EXCESS    I L1ETDOWN HX        I (j) 0-L                              218

J 0- I PR 0

0::: L_ 215 JDT I P

                     ,-                                                    223 r-t-~CVCS I

I I (CLOSED SYSTEM) MISSILE BARRIER

--1 I

I CONTAINMENT _I I SLOPE HCV-3003 1 I I S-1 .......- - - I DT - REACTOR COOLANT DRAIN TANK P - TRIPPED CLOSED BY CONTAINMENT ISOLATION SIGNAL PHASE B PRT - PRESSURIZER RELIEF TANK INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-4 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3385 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 12 REACTOR COOLANT SYSTEM SAMPLE LINES MISSILE BARRIERJ LCONTAINMENT I I ASWI T I I (f) U 12 I 0:: I SAMPLING (f) L0 956E 956F I I I SYSTEM I U I 0:: 954A 955A SLOPE S-1--- --I ITEM 13 FUEL TRANSFER TUBE AA J 7 DOUBLE GASKETED BLIND F ANGE SPENT FUEL I PIT 11-----.....+1~3 ___1~ I o I I I I S-1--- - - I I MISSILE BARRIER---I l--- CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-5 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3386 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 14 CONTAINMENT SPRAY HEADERS RWST 878A SEE NOTE 1 I 1 14 D.O LC 867A if) Q... L

)

Q... I MSWI >- 878B a:::: Q... if) LC f-Z w L Z f-Z 0 a:::: w SEE NOTE 1 I I 14 D.O. 867B I u a:::: f-L ___ a:::: z

         <<                         w 1m                         L            I W                         z         MSWI
         -..l if) f-if)                       Z L                         0 U

RWST - REFUELING WATER STORAGE TANK ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DESIGN NOTE 1: INDIAN POINT UNIT No. 2 FLANGED ELBOW USED DURING REFUELING OPERATIONS TO UFSAR FIGURE 5.2-6 FILL REACTOR REFUELING CAVITY AND CANAL CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3387 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 15 SAFETY INJECTION HEADERS ITEM 16 SAFETY INJECTION TEST LINE 856A

      ~

8568 15 850A I D.O. (/) I (L

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                          ~I                            I 0::::

w z ASWI 4: 0:::: I- ~I Z o CD u ~ ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DESIGN (/) (/) RWST - REFUELING WATER STORAGE TANK

2 INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-7 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3388 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 17 ACCUMULATOR N2 SUPPLY OPS t')

                                  ~                                                                     TRUCK 0)

I FILL ()

r: o OC)

D OC)

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2 o

0::: 1.L I I rCONTAINMENT I I----MISSILE BARRIER INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-8 CONTAINMENT ISOLATION SYSTEM PEN ETRATION SCH EMATICS MIC. No. 1999MC3389 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 19 PRIMARY SYSTEM VENT HEADER AND N2 SUPPLY LINE ITEM 20 REACTOR COOLANT DRAIN TANK TO GAS ANALYZER ITEM 21 RCDT PUMP DISCHARGE I (SLOPE 19 liT A~WI I T

               ,------------t!

I+-...... VENT HEADER f-- Z w e 1786

                                                 @I 17871

2 z (S-1)- --1

(/) CL f-- o z o o ---.J o (/) o ~ t-----TI~ N2 SUPPLY 5460 1616 1 (100 PSIG)

                                                                             @)            :

(S-1)-, T t--t------tII~ GAS ANAL YZER 1 1 ASWI 1 (S-1)-1 I t--t------tII~ HOLDU PTANK I 1 ASWI 1 1 (S-1)-1 INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-9 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3390 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 22 REACTOR COOLANT PUMP COOLING WATER IN (j) P P CL L CL

          ...--+----1                                                   COMPONENT COOLING

() 0::: 770 WATER LOOP o I-o L_ ASWI (CLOSED SYSTEM) SLOPE ITEM 23 REACTOR COOLANT PUMP WATER OUT (6 in.) (j) CL L P P CL I

     ~ ~+-...l--_D_R_A_IN.....2'=-'(3                                   COMPONENT COOLING L                                                                  WATER LOOP o

0::: 1.L I o (CLOSED SYSTEM) I L -ASWI MISSILE BARRIER-----j I I LSLOPE r-CONTAINMENT ITEM 24 REACTOR COOLANT PUMP COOLING WATER OUT (3 in.) (j) CL L CL () 0::: ....+ - - - - . . . . - . : COMPONENT COOLING L WATER LOOP o 0::: 1.L o L_ ASWI (CLOSED SYSTEM) SLOPE ENTIRE COMPONENT COOLING WATER SYSTEM IS SEISMIC CLASS 1 DESIGN INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-10 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3391 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 25 RESIDUAL HEAT EXCHANGER COOLING WATER IN ITEM 26 RESIDUAL HEAT EXCHANGER COOLING WATER RETURN I I s-- 11,1: 820A

                                             /~'e I~                                    I--

I ~~22A I I 2

                        .....                                                            I W
                                    ~

( S f-

                      )  .....                                                           I (J)

I (J)~ 2 I OW Zf-I  ! 125

                                                                             .........     -(J)
                                                                                           ---l>-

I I I O(J) u 0 X W I 26 ...... I UO W IIIl" I f-(J)

                        ~~

f- ZO ( 'e

                                             ,~ ~~22B
       <<~                                                                                  W ---l W---l                                                                             I I    I<<                                                        I                       I ZU 0'----"

u \ 0..

       ---lii

Jf-0'----"

i7i W

                    ;      ........ 11,1:

820B S I I I 2 0 U L __ I 0::: I I

                   ~

I I MISSILE CONTAINMENT BARRIER ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DESIGN S - OPEN S.I. SIGNAL, PHASE A INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-11 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3392 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 27 RECIRCULATION PUMP COOLING WATER SUPPLY ITEM 28 RECIRCULATION PUMP COOLING WATER RETURN r--- 753H

2 W

(f) I-(f) 0... >-

2 (f) ,.--..

) :2 0... OW

               '\  z
                 ) ~-i-
                         /       '\
                                   /

zl- _(f)

                                                                              -.J>-

O(f) o 00

                   -.J  '\                                                         W
             /     :::)    '-- /                                              I-(f) o                                                          ZO 0:::                                                       W -.J o                                                          ZO w                                                          0"--"

0::: 0...

2 o

o 753G 753A 28 MISSILE CONTAINMENT BARRIER ENTIRE COMPONENT COOLING SYSTEM IS SEISMIC CLASS 1 DESIGN INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-12 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3393 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 29 EXCESS LETDOWN HEAT EXCHANGER COOLING WATER IN ITEM 30 EXCESS LETDOWN HEAT EXCHANGER COOLING WATER OUT ITEM 31 CONTAINMENT SUMP PUMP DISCHARGE ITEM 31 A SAMPLING SYSTEM RETURN TO CONTAINMENT SUMP r--- I 2 29 w f-(f) 790 )- (f) ,.-.... 2 C)W Zf-

                                                                               -(/)
                                                                               -l)-

O(f) \ o 00 W f-(f) T ZO W -l zo 0'--" 0-30 2 o o 31 f---J-----1.- WA S TE I HOLDUP TANK L -t-S-1 I ASWI I I I I T I T I

     .--_+-               ~31'-'-'A'__II_II__-----J'----     -___<IIl      -  FROM SAMPLING SYSTEM 5132                   4399 I       0                      
            ---- MISSILE   ' - - - CON TAINMEN T BARRIER ENTIRE COMPONENT COOLING SYSTEM IS SEISMIC CLASS 1 DESIGN INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-13 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3394                       REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 32 CONTAINMENT AIR SAMPLE IN ITEM 33 CONTAINMENT AIR SAMPLE OUT ITEM 34 AIR EJECTOR DISCHARGE TO CONTAINMENT T I 32 GAS AND PARTICULATE MONITORS T T

                 --+='--------     -----l--      --+-4I- FROM AI R EJE CTOR PCV-1229        PCV- 123 0I o             I S-1 ------j I               I BARRIER    I MISSILE----I               j-CONTAINMENT AA -     AUTOMATIC PRESSURIZATION WITH AIR FROM PENETRA TION PRESSURIZATION SYSTEM INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-14 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3395             REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 37 STEAM GENERA TOR BLOWDOWN/SAMPLE (4 PENETRATIONS) 851 0:: f-0:: Z m w

2 Z

w  :::{

        -l (7)1 (f)        ~I

2 o

() PCV-1214 PCV-1214A PCV-1215 PCV-1215A PCV-1216 PCV-1216A PCV-1217 PCV-1217A I BLOWDOWN TANK SAMPLING SYSTEM I I I I ASWI (S-1)---,

• THIS DRAIN LINE HAS BEEN REMOVED ON STEAM GENERA TOR 21 INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-15 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3396 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

2 23 2004 11 :47:03 AM vnd1999MC3397-1-17B.DWG ITEM 39 10" VENTILATION SYSTEM COOLING WATER IN (5) ITEM 40 10" VENTILATION SYSTEM COOLING WATER OUT (5) ITEM 40A 2" VENTILATION SYSTEM MOTOR COOLING WATER OUT (5) SWN 42

                ,---_...----_ _                        --""3~9~(A_2_)---l...---L-----I~ H ~

SWN SWN 41A 41B (B) ('.J  ::2: (A1) w f- f-fJJ Z

) >-

fJJ

                                                                                                ~SWN-40 COOLING WATER 0::::            0 W
                          ---l,--....,     W                                                    ~SWN                         SUPPLY o      ---l       fJJ                                                  ~ 40-1 0<<                0 uo               d zo.... "---/
                          <<~

l.L "---/

                                                                             ~r----

[M] l--4=0'-t-------1 ' " ........,.----'--~ COOLING WATER

                                          ,--....,                   SWN    SWN                                               RETURN TO r----Lo::::-~                                     44A   44B                                               DISCHARGE
                      ..--- C)

W f-fJJ (A) (B1 ) CHANNEL NZ >- 0:::: ~ fJJ 0u,--...., 0 f- X ---l W Ow << fJJ

                      ~            U       0 f- -         ---l Z       <<     0.... u
                      <<W            ~     "---/

l.L I "---/

        ~MISSILE                                            I BARRIER                                                                                    IRM I=     RADIATION MONITORS CONTAINMENT--I ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DESIGN ALL MOv'S ARE DEENERGIZED OPEN PLANT MANAGEMENT MAY DESIGNATE THE "A" OR THE "B" VALVE(S) IN THE 41, 44 AND 71 SERIES AS THE REQUIRED CONTAINMENT ISOLATION                                         INDIAN POINT UNIT No. 2 VALVE(S). DESIGNATION OF THE "B" VALVE(S)

IN THE SWN-44 SERIES REQUIRES THE CODESIGNA TION OF THE RESPECTIVE 51 SERIES UFSAR FIGURE 5.2-16 VALVE(S) AS ADDITIONAL REQUIRED CONTAINMENT CONTAINMENT ISOLATION SYSTEM ISOLATION VALVE(S). CONTAINMENT ISOLATION VALVE PENETRATION SCHEMATICS SURVEILLANCE REQUIREMENTS MUST HAVE BEEN SATISFIED FOR THE MIC. No. 1999MC3397 I REV. No. 178 DESIGNATED VALVE(S). VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

3 2 2004 6:40:53 AM vnd 1999MC3398-1 -17B.DWG ITEM 41 SERVICE AIR f-----+---+-'--'---~~--"---- ~f--__+----<III_- SE RVI CE AI R

                     ~                                                      L.C.I TYPICAL                                              1 (S-1)~ ~

1 INSIDE OUTSIDE CONTAINMENT CONTAINMENT ITEM 43 WELD CHANNEL PRESSURIZATION AIR SUPPLY 43 o f----- PCV-1111-1 o PENETRATION PRESSURIZATION SYSTEM 43 o f----- (CLOSED SYSTEM) PCV-1111-2 o ALL PIPING IN PRESSURIZATION SYSTEM IS SEISMIC CLASS 1 INSIDE OUTSIDE CONTAINMENT CONTAINMENT ITEM 44 SPARE

                       ~I w                          I 0:::1                      1 C=~f-'-44-'-------1          --'----------'-----=-~
                       ~

IT) ] w L.C. 580BI

                       ---l Vi                                                ~           I
                       ~I                         1                      ~   (S-1)~---,

I INSIDE CONTAINMENT I OUTSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-17 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3398 REV. No. 178 VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 45 AUXILIARY STEAM SUPPLY ITEM 46 AUXILIARY STEAM CONDENSATE RETURN ITEM 47 CITY WATER I (S-1)---. -, I L.C.' 5 o~', 5 ~ ~ AUX. STEAM SUPPLY

          \4   UH-43         ,

I ,' , ASWI' I , I , , 5

• 46 .,.' I
• CONDENSATE RETURN 1

I o UH-44 L.C. I I I 5\--~----- 47 ~ L.C. MW-17-1 o , t ~ L.C.I MW-17 ,

                                            ,   ~     CITY WATER
                           , (S-1)---..J I

ASWI CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-18 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3399 TREV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 48 PURGE SUPPLY DUCT ITEM 49 PURGE EXHAUST DUCT CVI CVI I I I ~1:48 FCV-1170

                             .....   ~II FCV-1171 I

I I

                                                         .....      FROM PURGE SUPPLY FANS I             ([)    I                         I I

r-MISSILE BARRIE~ AA (S-l).......- .J I CONTAINMENT I I CVI I CVI I

               ~1:49 ~ ~II I

I I FCV-1172 FCV-1173 I .... TO PURGE EXHAUST FANS I ([) ITEM 50 CONTAINMENT PRESSURE RELIEF CVI CVI CVI I

  ~H~I:50~ ~I :I-------c--~ I 1--1--:- -l"~ l~6 PCV-1190              PCV-1191             PCV-1192       I AN PLANT VENT I                  ([)                                            ©     I I

!---MISSILE BARRIER (S-1).......- J AA AA I I I I CONTAINMENT---j INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-19 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3400 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 51 RECIRCULATION PUMP DISCHARGE SAMPLE LINE SIS ~1-_11~51 __~_T_I,------_~_T_----II"~ 993 990A T 9908 SAMPLING SYSTEM ([) , @

                                            ,I N2 (IVSWS)

INSIDE OUTSIDE CONTAINMENT ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 ITEM 52 PRESSURIZER STEAM SPACE SAMPLE ITEM 53 PRESSURIZER LIQUID SPACE SAMPLE PRESS. STEAM 52 I----l-.... SAMPLING SPACE (RCS) 9568' SYSTEM I (S-1)---,I ASWI T PRESS. LIQUID I----'-~ SAMPLING SPACE (RCS) 953 956D' SYSTEM

                                                           , (S-1)---,

ASWI INSIDE OUTSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-20 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC340 1 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEMS 54, 55, AND 56 CONTAINMENT PRESSURE INSTRUMENTATION LINES (CLOSED SYSTEM) 1--_--+55 (CLOSED SYSTEM) 1--_--+56

                                  .J \ 1814C (CLOSED SYSTEM)     ~        ~

CONDENSATE POT INSIDE OUTSIDE CONTAINMENT ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-21 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3402 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 57 POSTACCIDENT CONTAINMENT SAMPLING LINES (SUPPLY AND RETURN) 5 7 L.O. FAN UN IT ---t1lilll"-~F-'------1 23 18750 I - -__~I-----II~I------~ TO H2 / 02 FAN UN IT ---I1IiIII"-~F-57'---_L,_0-;, SOV-5018 SOV-5019 ANAL YZER 24 1875E ([) FAN UNIT 57 L.O. 21 I 1875A ro 0 ro

                                                              ~
                                                     ---l U1 FAN UNIT 22                             1875C                                      TO H2 / 02 SOV-5020    SOV-5021   ANAL YZER

([) FAN UNIT 157 L.O. 25 1875F INSIDE I OUTSIDE CONTAINMENT CONTAINMENT ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 157 L.O. RETU RN A ~----i~-----l 1875G SOV-5022 SOV-5023 ([) FROM H2 / 02 ANALYZERS I L.O. RE TU RN B 57

                ~------i~-----l 1875H     SOV-5024 SOV-5025

([) INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-22 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3403 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 58 SPARE I fAA I: 58/ :1 OUTSIDE INSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-23 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3404 IREV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEMS 59, 60, 61, and 62 SPARE I fAA II 59V II I I AA Ir I~I I I I r AA I~I I I I r AA I~I I INSIDE OUTSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-24 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3405 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

POSTACCIDENT (P.A.) CONTAINMENT VENTING SYSTEM ITEM 64 INSTRUMENT AIR/P.A. VENTING SUPPLY LINE ITEM 65 P.A. VENTING EXHAUST LINE T INSTRUMENT AIR PCV-1228 65 E-1 EA-1 ENTIRE SYSTEM SHOWN E-2 IS SEISMIC CLASS 1 o E-3 TO VENT STACK E-5

                               ~

141I--- AA INDIAN POINT UNIT No. 2 INSIDE OUTSIDE CONTAINMENT UFSAR FIGURE 5.2-25 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3406 REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 66 DELETED ITEM 67 CONTAINMENT LEAK TEST AIR LINE (2) l II-r >--67--- 1H><I-J ([) I SEISMIC CLASS 1 INSIDE OU TSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-26 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3407 IREV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 69 PERSONNEL AIR LOCK (2) CONTAINMENT I I I INSIDE OUTSIDE 85D 85C 95D 95C

                         .~, 5800*

SEE NOTE 2A~ 5801* AA AA e~ e~ 1 85B SEE NOTE SEE NOTE 1 85A 95B - 95A _

                                                -                                                  I--

INNER DOOR OUTER DOOR I-- I-- AA

   -.4-        EQUALIZING BALL VALVE INTERLOCKED WITH DOOR SPRING LOADED CHECK VALVE ENTIRE SYSTEM SHOWN IS SEISMIC CLASS 1 DEISGN NOTE1:    85A & 95A MAY BE OPEN WHEN 85B & 95B ARE CLOSED.

85B & 95B MAY BE OPEN WHEN 85A & 95A ARE CLOSED. 2 TEST VALVES 5800 & 5801 ARE LOCATED AT 83 FT. & 95 FT. AIRLOCKS RESPECTIVELY INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-27 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3408 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

ITEM 70 STEAM GENERA TOR LEVEL INDICATION LINES (2) PRESURIZER LEVEL INDICATION LINES (1) PRESSURIZER PRESSURE INDICATION LINES (1) I (£) _~1~70~ ~1-- ~1- -I~~ TO INDICATIORS IIP500 IIP501 IIP502 IIP503 I IIP504 IIP505 IIP506 IIP507 INSIDE OUTSIDE CONTAINMENT INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-28 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3409 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

LEGEND FOR SYMBOLS, CONTAINMENT ISOLATION SYSTEM VALVES OPERATORS Globe Air diaphragm Diaphragm (DIA) Air cylinder Gate Motor Double Disc Gate (DDV) Solenoid

              --           Check 11"11          Butterfly (BV)
                                                      ~

Relief ~J T Test Connection Self contained VALVE POSITION (NORMAL) pressure regulator [:><J open Needle

                                                          ~         closed Non-Return (Piston Type)

NOTATION ASWI - AUTOMATIC SEAL WATER INJECTION MSWI - MANUAL SEAL WATER INJECTION AA - AUTOMATIC PRESSURIZATION WITH AIR N2 - MANUAL PRESSURIZATION WITH NITROGEN LO - LOCKED OPEN LC - LOCKED CLOSED P - TRIPPED CLOSED BY CONTAINMENT ISOLATION SIGNAL, PHASE A T - TRIPPED CLOSED BY CONTAINMENT ISOLATION SIGNAL, PHASE B S-1 - SEISMIC CLASS I S - OPEN S.1. Signal, Phase A D.O. - DEENERGIZED OPEN CVI - CONTAINMENT VENTILATION ISOLATION (AUTOMATIC SIGNAL) IVSWS - ISOLATION VALVE SEAL WATER SYSTEM INDIAN POINT UNIT No. 2 UFSAR FIGURE 5.2-29 CONTAINMENT ISOLATION SYSTEM PENETRATION SCHEMATICS MIC. No. 1999MC3410 I REV. No. 17A VALVE POSITIONS ARE NOT CONTROLLED BY THIS FIGURE

Figures 5.1-2, 5.1-3, 5.1-4, 5.1-5, 5.1-6, and 5.1-7 Supplied as plant drawings 9321-F-2501, 2502, 2503, 2506, 2507, and 2508. Redacted by NRC staff as sensitive information.

IP2 FSAR UPDATE CHAPTER 6 ENGINEERED SAFETY FEATURES

6.0 INTRODUCTION

The central safety objective in reactor design and operation is the control of reactor fission products. The following methods are used to ensure this objective:

1. Core design to preclude release of fission products from the fuel (Chapter 3).

2. Retention of fission products in the reactor coolant for whatever leakage occurs (Chapters 4 and 6).

3. Retention of fission products by the containment for operational and accidental releases beyond the reactor coolant boundary (Chapters 5 and 6).

4. Optimizing fission product dispersal to minimize population exposure (Chapters 2 and 11).

The engineered safety features are the provisions in the plant that embody methods 2 and 3 above to prevent the occurrence or to ameliorate the effects of serious accidents. The engineered safety features systems in this plant are the containment system, detailed in Chapter 5, the safety injection system, detailed in Section 6.2, the containment spray system, detailed in Section 6.3, the containment air recirculation cooling system, detailed in Section 6.4, the isolation valve seal-water system, detailed in Section 6.5, and the containment penetration and weld channel pressurization system, detailed in Section 6.6. Evaluations of techniques and equipment used to accomplish the central objective including accident cases are detailed in Chapters 5, 6, and 14. 6.1 GENERAL DESIGN CRITERIA Criteria applying in common to all engineered safety features are given in Section 6.1.1. Thereafter, criteria that are related to engineered safety features, but which are more specific to other plant features or systems, are listed and cross referenced in Section 6.1.2. 6.1.1 Engineered Safety Features Criteria 6.1.1.1 Engineered Safety Features Basis for Design Criterion: Engineered safety features shall be provided in the facility to back up the safety provided by the core design, the reactor coolant pressure boundary, and their protection systems. Such engineered safety features shall be designed to cope with any size reactor coolant piping break up to and including the equivalent of a circumferential rupture of any pipe in that boundary, assuming unobstructed discharge from both ends. (GDC 37) The design, fabrication, testing, and inspection of the core, reactor coolant pressure boundary, and their protection systems give assurance of safe and reliable operation under all anticipated normal, transient, and accident conditions. However, engineered safety features are provided in Chapter 6, Page 1 of 174 Revision 22, 2010

IP2 FSAR UPDATE the facility to back up the safety provided by these components. These engineered safety features have been designed to cope with any size reactor coolant pipe break up to and including the circumferential rupture of any pipe assuming unobstructed discharge from both ends as discussed in Section 14.3.3.3. They are also designed to cope with any steam or feedwater line break up to and including the main steam or feedwater lines as discussed in Section 14.2.5. Limiting the release of fission products from the reactor fuel is accomplished by the safety injection system, which by cooling the core, keeps the fuel in place and substantially intact and limits the metal water reaction to an insignificant amount. The safety injection system consists of high- and low-head centrifugal pumps driven by electric motors and of passive accumulator tanks that are self-energized and which act independently of any actuation signal or power source. The release of fission products from the containment is limited in three ways:

1. Blocking the potential leakage paths from the containment. This is accomplished by:

a. A steel-lined, reinforced-concrete reactor containment with testable, double-sealed penetrations, and liner weld channels, the spaces of which are continuously pressurized above accident pressure and which form a virtually leaktight barrier to the escape of fission products should a loss-of-coolant accident occur. (Section 6.6.2 lists those portions of the Weld Channel Pressurization System that have been disconnected because repairs have been determined not to be practical.)

b. Isolation of process lines by the containment isolation system, which imposes double barriers in each line that penetrates the containment except for lines used during the accident. Pipes penetrating the containment are sealed as shown in Table 5.2-1. This table presents the sealing method for all containment piping penetrations and valving.

2. Reducing the fission product concentration in the containment atmosphere. This is accomplished by containment spray, which removes elemental iodine vapor and particulates from the containment atmosphere by washing action. The spray is chemically treated during the recirculation phase to enhance iodine retention.

3. Reducing the containment pressure and thereby limiting the driving potential for fission product leakage. This is accomplished by cooling the containment atmosphere by the following independent systems of approximately equal heat removal capacity that together also function to ensure the containment design criteria is maintained even with an assumed single failure:

a. Containment spray system.

b. Containment air recirculation cooling system.

Chapter 6, Page 2 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.1.1.2 Reliability and Testability of Engineered Safety Features Criterion: All engineered safety features shall be designed to provide such functional reliability and ready testability as is necessary to avoid undue risk to the health and safety of the public. (GDC 38) A comprehensive program of plant testing was formulated for all equipment, systems, and system control vital to the functioning of engineered safety features. The program consisted of performance tests of individual pieces of equipment in the manufacturer's shop, and integrated tests of the systems as a whole. Periodic tests of the actuation circuitry and mechanical components ensure reliable performance, upon demand, throughout the life of the plant. The initial tests of individual components and the integrated test of the system as a whole complemented each other to ensure the performance of the system as designed and to prove proper operation of the actuation circuitry. Routine periodic testing of the engineered safety features components is performed, in accordance with the Technical Specifications. 6.1.1.3 Missile Protection Criterion: Adequate protection for those engineered safety features, the failure of which could cause an undue risk to the health and safety of the public, shall be provided against dynamic effects and missiles that might result from plant equipment failures. (GDC 40) A loss-of-coolant accident or other plant equipment failure might result in dynamic effects or missiles. For engineered safety features that are required to ensure safety in the event of such an accident or equipment failure, protection is provided primarily by the provisions that are taken in the design to prevent the generation of missiles. In addition, protection is also provided by the layout of plant equipment or by missile barriers in certain cases. (Refer to Section 5.1.2 for a discussion of missile protection.) Injection paths leading to unbroken reactor coolant loops are protected against damage resulting from the maximum reactor coolant pipe rupture by layout and structural design considerations. Injection lines penetrate the main missile barrier, which is the crane wall, and the injection headers are located in the missile protected area between the crane wall and the containment wall. Individual injection lines, connected to the injection header, pass through the barrier and then connect to the loops. The separation of the individual injection lines is provided to the maximum extent practicable. The movement of the injection line, associated with a rupture of a reactor coolant loop, is accommodated by line flexibility and by the design of the pipe supports such that no damage outside the missile barrier is possible. The containment structure is capable of withstanding the effects of missiles originating outside the containment and which might be directed toward it so that no loss-of-coolant accident can result. All hangers and anchors are designed in accordance with USAS B31.1, Code for Pressure Piping. This code provides minimum requirements on material, design, and fabrication with ample safety margins for both dead and dynamic loads over the life of the equipment. Concrete Chapter 6, Page 3 of 174 Revision 22, 2010

IP2 FSAR UPDATE missile barriers, bumpers, walls and other concrete structures are designed in accordance with ACI 318-63, Building Code Requirements for Reinforced Concrete. In 1989, the NRC approved changes to the design basis with respect to dynamic effects of postulated primary loop ruptures, as discussed in Section 4.1.2.4 6.1.1.4 Engineered Safety Features Performance Capability Criterion: Engineered safety features, such as the emergency core cooling system and the containment heat removal system, shall provide sufficient performance capability to accommodate the failure of any single active component without resulting in undue risk to the health and safety of the public. (GDC 41) Each engineered safety feature provides sufficient performance capability to accommodate any single failure of an active component and still function in a manner to avoid undue risk to the health and safety of the public. The extreme upper limit of public exposure is taken as the levels and time periods presently outlined in 10 CFR 50.67 (i.e., 25 rem total effective dose equivalent (TEDE) at the exclusion radius for the worst two hour interval, 25 rem TEDE over the duration of the accident at the low-population-zone distance and 5 rem TEDE for the operators in the control room for the duration of the accident). The accident condition considered is the hypothetical case of a release of fission products per NUREG-1465. Also, the total loss of all outside power is assumed concurrently with this accident. With minimum engineered safety features systems functioning, the offsite exposure would be within 10 CFR 50.67 limits as discussed in Section 14.3.6. Under these accident conditions, the containment air recirculation cooling system and the containment spray system are designed and sized so that both systems, each operating with partial effectiveness, are able to supply the necessary postaccident cooling capacity to ensure the maintenance of containment integrity, that is, keeping the pressure below design pressure at all times assuming that the core residual heat is released to the containment as steam. Partial effectiveness is defined as the operation of a system with at least one active component failure. Containment spray relies on a sufficient amount of passive sodium tetraborate stored in containment to raise the pH of the recirculating solution for continued iodine removal following an accident. The containment spray system alone is able to supply the post accident iodine removal required to restrict the offsite exposure to within 10 CFR 50.67 limits. 6.1.1.5 Engineered Safety Features Components Capability Criterion: Engineered safety features shall be designed so that the capability of these features to perform their required function is not impaired by the effects of a loss-of-coolant accident to the extent of causing undue risk to the health and safety of the public. (GDC 42) Instrumentation, pumps, fans, cooling units, valves, motors, cables, and penetrations located inside the containment are selected to meet the most adverse accident conditions to which they may be subjected. These items are either protected from containment accident conditions or are designed to withstand, without failure, exposure to the worst combination of temperature, pressure, and humidity expected during the required operational period. Chapter 6, Page 4 of 174 Revision 22, 2010

IP2 FSAR UPDATE In response to NRC Generic Letter 95-07, safety-related power-operated gate valves have been evaluated for susceptibility to pressure locking and thermal binding. The results of this evaluation identified that those potential conditions will not prevent the plant from achieving a safe shutdown, as all valves evaluated remain operable. This conclusion is based upon valve design, plant configuration during normal, accident, and post accident operating modes and sufficient actuator thrust to open the valve. The details of the system and valve evaluations are documented in Reference 1. In response to NRC Generic Letter 96-06, isolated pipe line segments that penetrate containment have been analyzed to evaluate their susceptibility to overpressurization caused by thermal expansion of the contained fluid in the event of a design basis accident. The results show that potential overpressurization will not cause lines to fail; all remain operable. Those that are protected by safety relief devices were further evaluated for the effects of stuck-open relief valves under accident conditions. No failure modes were identified that would adversely affect the ability for safety-related systems to perform their intended functions during accidents. The safety injection system pipes serving each loop are restrained in such a manner as to restrict potential accident damage to the portion of piping downstream of the crane wall that constitutes the missile barrier in each loop area. The restraints are designed to withstand, without failure, the thrust force of any branch line severed from the broken reactor coolant pipe and discharging fluid to the atmosphere and to withstand a bending moment equivalent to that which produces a failure of the piping under the action of free-end discharge to the atmosphere or a motion of the broken reactor coolant pipe to which the injection pipes are connected. This prevents possible failure at any point upstream from the support point, including the branch line connection into the piping header. In 1989, the NRC approved changes to the design basis with respect to dynamic effects of postulated primary loop ruptures, as discussed in Section 4.1.2.4. Designated valves that are located in areas that would have excessive radiation levels in the event of a release of fission products from the core are provided with capability for remote operation. 6.1.1.6 Accident Aggravation Prevention Criterion: Protection against any action of the engineered safety features, which would accentuate significantly the adverse after effects of a loss of normal cooling shall be provided. (GDC 43) The introduction of borated cooling water into the core results in a negative reactivity addition. The control rods insert and remain inserted. The supply of water by the safety injection system to cool the core cladding does not produce significant metal-water reaction (<1.0-percent). The delivery of cold safety injection water to the reactor vessel following accidental expulsion of reactor coolant does not cause further loss of integrity of the reactor coolant system boundary. Chapter 6, Page 5 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.1.1.7 Sharing of Systems Criterion: Reactor facilities may share systems or components if it can be shown that such sharing will not result in undue risk to the health and safety of the public. (GDC 4) The engineered safety features at Indian Point 2 do not share systems or components with other units. 6.1.2 Related Criteria The following are criteria, which although related to all engineered safety features, are more specific to other plant features or systems, and, therefore, are discussed in other chapters, as listed. Name Discussion Quality Standards (GDC 1) Chapter 4 Performance Standards (GDC 2) Chapter 4 Records Requirements (GDC 5) Chapter 4 Instrumentation and Control Systems (GDC 12) Chapter 7 Engineered Safety Features Protection Systems (GDC 15) Chapter 7 Emergency Power (GDC 39) Chapter 8 REFERENCES FOR SECTION 6.1

1. NRC Letter dated May 20, 1999, Jeffrey F. Harold to A. Alan Blind,

Subject:

Safety Evaluation of Generic Letter 95-07, "Pressure Locking and Thermal Binding of Safety-Related Power-Operated Gate Valves," for Indian Point Nuclear Generating Unit No.2 (TAC M93473). 6.2 SAFETY INJECTION SYSTEM At Indian Point Unit 2 the emergency core cooling function is performed by the safety injection system. Therefore, whenever the term "emergency core cooling system" or ECCS is referenced in the document it is synonymous with the safety injection system. 6.2.1 Design Basis 6.2.1.1 Emergency Core Cooling System Capability Criterion: An emergency core cooling system with the capability for accomplishing adequate emergency core cooling shall be provided. This core cooling system and the core shall be designed to prevent fuel and clad damage that would interfere with the emergency core cooling function and to limit the clad metal-water reaction to acceptable amounts for all sizes of breaks in the reactor coolant piping up to the equivalent of a double-ended rupture of the largest pipe. The performance of such emergency core cooling system shall be evaluated conservatively in each area of uncertainty. (GDC 44) Chapter 6, Page 6 of 174 Revision 22, 2010

IP2 FSAR UPDATE Adequate emergency core cooling is provided by the safety injection system (which constitutes the emergency core cooling system) whose components operate in three modes. These modes are delineated as passive accumulator injection, active safety injection, and residual heat removal recirculation. The primary purpose of the safety injection system is the automatic delivery of cooling water to the reactor core in the event of a loss-of-coolant accident. This limits the fuel clad temperature and thereby ensures that the core will remain intact and in place, with its essential heat transfer geometry preserved. This protection is afforded for:

1. All pipe break sizes up to and including the hypothetical instantaneous circumferential rupture of a reactor coolant loop, assuming unobstructed discharge from both ends.

2. A loss of coolant associated with the rod ejection accident.

3. A steam-generator tube rupture.

The basic design criteria for loss-of-coolant accident evaluations prior to codification under 10 CFR 50.46 were as follows:

1. The cladding temperature is to be less than:

a. The melting temperature of Zircaloy-4.

b. The temperature at which gross core geometry distortion, including clad fragmentation, may be expected.

2. The total core metal-water reaction will be limited to less than 1-percent.

These criteria ensure that the core geometry remains in place and substantially intact to such an extent that effective cooling of the core is not impaired. Subsequently, the basic design criteria for loss-of-coolant accident calculations have been revised to those required under 10 CFR 50.46. For any rupture of a steam pipe and the associated uncontrolled heat removal from the core, the safety injection system adds shutdown reactivity so that with a stuck rod, no offsite power, and minimum engineered safety features, there is no consequential damage to the reactor coolant system and the core remains in place and coolable as discussed in Section 14.2.5. Redundancy and segregation of instrumentation and components are incorporated to ensure that postulated malfunctions will not impair the ability of the system to meet the design objectives. The system is effective in the event of loss of normal station auxiliary power coincident with the loss of coolant, and is tolerant of failures of any single component or instrument channel to respond actively in the system. During the recirculation phase, the system is tolerant of a loss of any part of the flow path since backup alternative flow path capability is provided as described in Section 6.2.3.3. The ability of the safety injection system to meet its capability objectives is presented in Section 6.2.3. The analysis of the accidents is presented in Chapter 14. Chapter 6, Page 7 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.2.1.2 Inspection of Emergency Core Cooling System Criterion: Design provisions shall, where practical, be made to facilitate inspection of physical parts of the emergency core cooling system, including reactor vessel internals and water injection nozzles. (GDC 45) Design provisions are made to the extent practical to facilitate access to the critical parts of the reactor vessel internals, pipes, valves, and pumps for visual or boroscopic inspection for erosion, corrosion, and vibration-wear evidence and for nondestructive test inspection where such techniques are desirable and appropriate. 6.2.1.3 Testing of Emergency Core Cooling System Component Criterion: Design provisions shall be made so that components of the emergency core cooling system can be tested periodically for operability and functional performance. (GDC 46) The design provides for periodic testing of active components of the safety injection system for operability and functional performance. Power sources are arranged to permit individual actuation of each active component of the safety injection system. The safety injection pumps and residual heat removal pumps can be tested periodically during plant operation using the minimum flow recirculation lines provided. The residual heat removal pumps are used every time the residual heat removal loop is put into operation. All remote-operated valves can be exercised, and actuation circuits can be tested either during normal operation or routine plant maintenance. 6.2.1.4 Testing of Emergency Core Cooling System Criterion: Capability shall be provided to test periodically the operability of the emergency core cooling system up to a location as close to the core as is practical. (GDC 47) An integrated system test can be performed when the plant is cooled down and the residual heat removal loop is in operation. This test would not introduce flow into the reactor coolant system, but would demonstrate the operation of the valves, pump circuit breakers, and automatic circuitry upon the initiation of safety injection. Level and pressure instrumentation is provided for each accumulator tank, and accumulator tank pressure and level are continuously monitored during plant operation. Flow from the tanks can be checked at any time using test lines as described in Section 6.2.5.3.1. 6.2.1.5 Testing of Operational Sequence of Emergency Core Cooling System Criterion: Capability shall be provided to test initially, under conditions as close as practical to design, the full operational sequence that would bring the emergency core cooling system into action, including the transfer to alternate power sources. (GDC 48) Chapter 6, Page 8 of 174 Revision 22, 2010

IP2 FSAR UPDATE The design provides for the capability to test initially, to the extent practical, the full operational sequence up to the design conditions for the safety injection system to demonstrate the state of readiness and capability of the system. Details of the operational sequence testing are presented in Section 6.2.5. 6.2.1.6 Codes and Classifications Table 6.2-1 lists the codes and standards to which the safety injection system components are designed. 6.2.1.7 Service Life All portions of the system located within the containment are designed to operate without benefit of maintenance and without loss of functional performance for the duration of time the component is required. Per the 12/06/04 NRC generic SER on NEI 04-07 (Reference 5), evaluations of PWR post accident emergency recirculation performance to address the potential impact of debris blockage per GSI-191 (Reference 4) and Generic Letter 2004-02 (Reference 3) will use a mission time of 30 days. 6.2.2 System Design And Operation 6.2.2.1 System Description Adequate emergency core cooling following a loss-of-coolant accident is provided by the safety injection system shown in Plant Drawing 9321-2735 [Formerly UFSAR Figure 6.2-1]. Plant Drawing 235296 [Formerly UFSAR Figures 6.2-2] and Figures 6.2-2 through 6.2-5 depict how this system concept is translated into plant layout design. The system components operate in the following possible modes:

1. Injection of borated water by the passive accumulators.

2. Injection by the safety injection pumps drawing borated water from the refueling water storage tank.

3. Injection by the residual heat removal pumps also drawing borated water from the refueling water storage tank.

4. Recirculation of spilled reactor coolant, injected water, and containment spray system drainage back to the reactor from the recirculation sump by the recirculation pumps. The residual heat removal pumps provide backup recirculation capability through the independent containment sump as described in Section 6.2.3.3.

The initiation signal for core cooling by the safety injection pumps and the residual heat removal pumps is the safety injection signal, which is described in Section 7.2.3.2.3. 6.2.2.1.1 Injection Phase The principal components of the safety injection system, which provide emergency core cooling immediately following a loss of coolant are the accumulators (one for each loop), the three safety injection (high-head) pumps, and the two residual heat removal (low-head) pumps. The safety Chapter 6, Page 9 of 174 Revision 22, 2010

IP2 FSAR UPDATE safety injection and residual heat removal pumps are located in the primary auxiliary building. The accumulators, which are passive components, discharge into the cold legs of the reactor coolant piping when pressure decreases to the minimum Technical Specification value, thus rapidly ensuring core cooling for large breaks. They are located inside the containment, but outside the crane wall, therefore each is protected against possible missiles. The safety injection signal opens certain of the safety injection system isolation valves, provides confirmatory open signals to system isolation valves that are normally open, and starts the safety injection pumps and residual heat removal pumps. The three safety injection pumps (high-head) deliver borated water to two separate discharge headers. The flow from the discharge headers can be injected into the four cold legs and two hot legs of the reactor coolant system. The motor-operated isolation valves in the four cold-leg injection lines are open during normal plant operation. The motor-operated isolation valves in the two hot-leg injection lines are closed during normal plant operation. The hot-leg injection lines are provided for later use during hot-leg recirculation following a reactor coolant pressure boundary break. The high-head safety injection system is configured with two cold leg injection lines physically connected to the reactor coolant pressure boundary and the other two lines connected to the accumulator discharge lines upstream of the pressure boundary. Since a small break in the reactor coolant pressure boundary can include a cold leg injection line, safety injection flow capability can be limited by the resulting flow from only three intact cold leg injection lines. Depending on the assumed single failure, either two or three safety injection pumps can be operating. To maximize the fraction of safety injection flow delivered to the reactor coolant system with a broken cold leg injection line, the four cold leg injection lines are flow balanced to within an allowable range. The resulting system flow capability is sufficient for the makeup of coolant following a small break that does not immediately depressurize the reactor coolant system to the accumulator discharge pressure. Credit is not taken for operator action to isolate a broken cold leg injection line. For large breaks, the reactor coolant system would be depressurized and voided of coolant rapidly (about 26 sec for the largest break as shown in Figure 14.3-12) and a high flow rate is required to recover quickly the exposed fuel rods and limit possible core damage as discussed in Section 14.3.3.3.1. To achieve this objective, one residual heat removal pump and two safety injection pumps are required to deliver borated water to the cold legs of the reactor coolant loops. Two residual heat removal and three safety injection pumps are available to provide for an active component failure. Delivery from these pumps supplements the accumulator discharge. Since the reactor coolant system backpressure is relatively low (rapid depressurization for large breaks), a broken injection line would not appreciably change the flows in the other injection line's delivery to the core. The residual heat removal pumps take suction from the refueling water storage tank. In addition, the charging pumps of the chemical and volume control system are available but are not required to augment the flow of the safety injection system. Because the injection phase of the accident is terminated before the refueling water storage tank is completely emptied, all pipes are kept sufficiently filled with water before recirculation is initiated to ensure the systems remain operable and perform properly. Water level indication and alarms on the refueling water storage tank give the operator ample warning to terminate the injection phase. Additional level indicators and alarms are provided in the recirculation and Chapter 6, Page 10 of 174 Revision 22, 2010

IP2 FSAR UPDATE containment sumps, which also give backup indication when injection can be terminated and recirculation initiated. 6.2.2.1.2 Recirculation Phase After the injection operation, coolant spilled from the break and water collected from the containment spray are cooled and returned to the reactor coolant system by the recirculation system. When the break is large, depressurization occurs due to the large rate of mass and energy loss through the break to containment. In the event of a large break, the recirculation flow path is within the containment. The system is arranged so that the recirculation pumps take suction from the recirculation sump in the containment floor and deliver spilled reactor coolant and borated refueling water back to the core through the residual heat exchangers. The system is also arranged to allow either of the residual heat removal pumps to take over the recirculation function. The residual heat removal pumps would only be used if backup capacity to the internal recirculation loop is required as described in Section 6.2.3.3. Water is delivered from the containment to the residual heat removal pumps from the separate containment sump inside the containment. For small breaks, the depressurization of the reactor coolant system is augmented by steam dump from and auxiliary feedwater addition to the steam generators. For the smaller breaks in the reactor coolant system where recirculated water must be injected against higher pressures for long-term cooling, the system is arranged to deliver the water from residual heat removal heat exchanger 21 to the high-head safety injection pump suction and by this external recirculation route to the reactor coolant loops. If this flow path is unavailable, an alternate flow path is provided as indicated in Table 6.2-11. Thus, if depressurization of the reactor coolant system proceeds slowly, the safety injection pumps may be used to augment the flow-pressure capacity of the recirculation pumps in returning the spilled coolant to the reactor. In this system configuration, the recirculation pump (or residual heat removal pump) provides flow and net positive suction head to the operating safety injection pumps. To prevent safety injection pump flow in excess of its maximum allowable (i.e., runout) limit, variable flow orifices are installed at the discharge of the safety injection pumps and the hot and cold leg motor-operated isolation valves are preset with mechanical stops based on data from operational flow testing to limit system maximum flow capability. The recirculation pumps, the residual heat removal heat exchangers, piping, and valves vital to the function of the recirculation loop are located in a missile-shielded space inside the polar crane support wall on the west side of the reactor primary shield. There are two sumps within the containment, the recirculation sump and the containment sump. Both sumps collect liquids discharged into the containment during the injection phase of the design-basis accident. As part of the resolution of GSI-191 and Generic Letter 2004-02, various flow channeling barriers are installed in the Vapor Containment, EL 46-0 to force the recirculation flow into the Reactor Cavity Sump area, up and out the Incore Instrumentation Tunnel, through the Crane Wall via the three nominal 20 inch square openings and into the annulus area outside the Crane Wall. The recirculation flow will migrate towards the Recirculation Sump Strainer or the Containment Sump Strainer depending on which pump(s) are operating. Flow channeling barriers are installed on the reactor Cavity Platform, EL 29-4, around the Incore Instrumentation Tunnel, on the Chapter 6, Page 11 of 174 Revision 22, 2010

IP2 FSAR UPDATE Instrumentation Tunnel, on the Recirculation Sump Trenches, at the Containment Sump, and on Crane Wall penetrations up to the flood level. Flow channeling barrier doors are installed in the Northeast and Northwest quadrant openings of the Crane Wall. In addition, flow channeling barrier doors are installed in the North and South entrances to the Recirculation Sump area. Perforated plate is installed on the RHR Heat Exchanger Platform, EL 66-0 to preclude debris from washing through the existing grating and into the Recirculation Sump area. Forcing the recirculation flow path into the Reactor Cavity Sump area (a low velocity zone) allows the larger debris an opportunity to settle. The Recirculation Sump and Containment Sump strainers consist of a matrix of multi-tube top-hat modules, which are fabricated from perforated stainless steel plate and mounted in the horizontal position. The perforated plate has 3/32 diameter holes sized to limit downstream affects. The top-hat modules have four (4) layers of perforated surfaces for straining debris from the sump fluid. Typical Recirculation Sump and Containment Sump strainer top-hat modules consist of a 12-1/2 diameter outer perforated tube with a respective 10-1/2 diameter inner perforated tube and a second set of tubes, which consist of a 7-1/2 diameter outer perforated tube with a respective 5-1/2 diameter inner perforated tube. The top-hat modules feature an internal vortex suppressor, which prevents air ingestion into the piping system. Stainless steel mesh has been installed between each pair of perforated plate tubes to minimize fiber bypass through the strainers. The top-hat modules are attached to strainer water boxes. Frame structures supporting sections of grating are installed above the Internal Recirculation and Containment Sump strainers including the sump strainer extension in the containment annulus providing for additional vortex suppression function. The Containment Sump Level Detection System is discussed in Section 6.7.2.13. The Recirculation Sump relies on two, connected water boxes with 249 top-hat modules in the sump pit for the purpose of preventing particles greater than 3/32 in diameter from entering the suction of the recirculation pumps. The recirculation sump strainer has effective surface area of ~3,156 square feet and an effective interstitial volume of ~476 cubic feet. Water will enter the top-hat modules through the perforated plates and flow through the stainless steel mesh inside either of the two (2) annuli flow paths within each top-hat module. Upon exiting the top-hat modules, water will flow into either of the two connected strainer water boxes, flow over the Recirculation Sump weir wall and into the Recirculation Pump Bay towards the pumps. The Containment Sump relies on a water box with 23 top-hat modules in the Containment Sump pit and a plenum extension out into the annulus with a water box with an additional 40 top-hat modules for the purpose of preventing particles greater than 3/32 in diameter from entering the Containment Sump suction line to the RHR Pumps. The Containment Sump strainer has a combined effective surface area of ~1182 square feet (~412 sq. ft. for sump pit and ~770 sq. ft. for the annulus extension) and an effective combined interstitial volume of ~161 cubic feet. Water will enter the top-hat modules through the perforated plates and flow through the stainless steel mesh inside either of the two (2) annuli flow paths within each top-hat module. Upon exiting the top-hat modules, water will flow into the strainer water box, which is connected to the Containment Sump suction line and the RHR System. The containment sump level detection system is discussed in Section 6.7.1.2.13. Each sump strainer is qualified to handle the post LOCA design basis accident debris loads predicted by the mechanistic evaluations required by GL 2004-02. There are two classifications for the debris generated by an RCS break: 1) conventional debris (e.g., insulation, tags, coatings, dust and dirt), 2) chemical debris (principally the precipitation of Aluminum based compounds Sodium Aluminum Silicate and Aluminum Oxy Hydroxide) which are conservatively Chapter 6, Page 12 of 174 Revision 22, 2010

IP2 FSAR UPDATE predicted by use of a model detailed in WCAP-16530-NP-A (Reference 6). An Argonne National Laboratory (ANL) formula was used to predict the post-LOCA chemical precipitation temperature. The precipitation temperature is determined from the post-accident containment sump pool conditions (Aluminum concentration, temperature and pH). Chemical precipitants are not predicted to develop prior to the required switchover to hot-leg recirculation. Consequently, the internal recirculation sump strainer qualification uses predicted head losses associated with conventional debris loads up to the switchover to hot-leg recirculation and then conventional and chemical debris loads after the transfer to hot-leg recirculation occurs when reduced sump flow rates are expected to be less than two HHSI pumps at runout (2 x 675 GPM). Per the 12/06/04 NRC generic SER on NEI 04-07 (Reference 5), evaluations of PWR post accident emergency recirculation performance to address the potential impact of debris blockage per GSI-191 (Reference 4) and Generic Letter 2004-02 (Reference 3) will use a mission time of 30 days. The internal recirculation sump strainers are qualified for a GL 2004-02 defined 30-day mission time. The containment sump strainers were qualified using the same methodology but are qualified from 24 hours post large break LOCA until the end of the 30 day mission time. The containment sump is not required to handle the same full debris loads at the start of the recirculation phase as the recirculation sump since the only postulated failure that would require its use is a passive failure, which is only postulated after 24 hours into the event (reference Technical Specification Amendment #257). However, to maintain redundancy for the more probable small break LOCAs, the containment sump strainers have been qualified for 6 inch diameter breaks and smaller from the start of the recirculation phase. A condition of Amendment

1. 257 was that the Emergency Operating Procedures continue to utilize the containment sump as an alternative path should both the recirculation sump trains become unavailable.

As identified above in the strainer descriptions, both sump strainers are constructed of concentric cylindrical tubes perforated with 3/32 inch diameter holes and have stainless steel mesh behind the perforations to reduce the quantity of fine fibers able to pass through the strainer perforations, although some fine fibers and particulates may still pass through the strainer. The passive sump strainers provide adequate protection to the downstream components from the majority of the accident generated debris. As part of the resolution of GL 2004-02, an analysis of the components downstream of the strainers required for accident mitigation was performed to ensure satisfactory operation for the defined 30-day mission time. Pumps, isolation and throttle valves, orifices, instrument connections, and piping were examined using the guidance provided by revision 1 of WCAP-16406 (Reference 7) with justification provided for any methodology deviations. All equipment was found to have sufficient clearance: to allow passage of debris, to limit blockage to an acceptable level, and / or to have sufficient resistance to wear as to not affect there function for the defined 30-day mission time. Chemical effects on the fuel elements were also examined and not predicted to interfere with heat transfer per analysis based on WCAP-16793-NP (Reference 8). The low-head external recirculation loop via the containment sump line and the residual heat removal pumps provides backup recirculation capability to the low-head internal recirculation loop. The containment sump line has two remote motor-operated normally closed valves located outside the containment and a remote motor-operated butterfly valve inside containment. The high-head external recirculation flow path via the high-head safety injection pumps is required for the range of small-break sizes for which the reactor coolant system pressure remains in excess of the shutoff head of the recirculation pumps at the end of the injection phase. The recirculation pumps, or residual heat removal pumps if backup capability is required, are also used to provide Chapter 6, Page 13 of 174 Revision 22, 2010

IP2 FSAR UPDATE required, are also used to provide flow to the high-head safety injection pumps during hot leg recirculation. The external recirculation flow paths within the primary auxiliary building are designed so that external recirculation can be initiated immediately after the accident. Those portions of the safety injection system outside of the containment, which are designed to circulate, under postaccident conditions, radioactivity contaminated water collected in the containment meet the following requirements:

1. Shielding to limit radiation levels.

2. Collection of discharges from pressure-relieving devices into closed systems.

3. Means to detect and control radioactivity leakage into the environs.

These criteria are met by minimizing leakage from the system. External recirculation loop leakage is discussed in Section 6.2.3.8. The radiological consequences of external recirculation loop leakage following a design basis accident are presented in Section 14.3.6.6. Detection and control of leakage from external recirculation loop components is also discussed in Section 6.7. One recirculation pump and one residual heat exchanger of the recirculation system provide sufficient cooled recirculated water to keep the core flooded with water by injection through the cold-leg connections while simultaneously providing, sufficient containment recirculation spray flow to reduce containment airborne activity. These systems are kept sufficiently filled with water to ensure the systems remain operable and perform properly. Three of the five fan cooler units prevent the containment pressure from rising above design limit. Analysis demonstrates that flow will be determined by system resistance provided by the physical configuration of the recirculation piping and components, and will be hydraulically balanced such that sufficient flow is established to the core and the spray header. Only one pump and one heat exchanger are required to operate for this capability at the earliest time recirculation spray is initiated. With both recirculation pumps in operation and both spray header valves open, a recirculation spray flow rate can be established such that no containment cooling fans (Section 6.4) are required. Likewise with five containment cooling units in operation, no containment spray is required to maintain containment pressure below its design limit. The system is also arranged to allow either of the residual heat removal pumps to take over the recirculation function following a passive failure as defined in Section 6.2.3.3. This design ensures that heat removal from the core and containment is effective in the event of a pipe or valve body rupture. 6.2.2.1.3 Cooling Water The service water system (Section 9.6) provides cooling water to the component cooling loop, which in turn cools the residual heat exchangers, both of which are part of the auxiliary coolant systems (Section 9.3). Three non-essential service water pumps are available to take suction from the river and discharge to the two component cooling heat exchangers. Three component cooling pumps are available to discharge through their heat exchangers and deliver to the two residual heat exchangers. During the recirculation phase following a loss-of-coolant-accident, only one residual heat removal heat exchanger, one recirculation or residual heat removal pump, one non-essential service water pump, one component cooling water pump, one auxiliary component cooling water pump, and one component cooling water heat exchanger are required to meet the core-cooling function. The auxiliary component cooling water pump is required only to support the function of the recirculation pump. With the exception of the residual heat removal heat exchangers and the recirculation pumps, all of the cited equipment is located outside of the Chapter 6, Page 14 of 174 Revision 22, 2010

IP2 FSAR UPDATE outside of the containment. 6.2.2.1.4 Changeover From Injection Phase to Recirculation Phase Assuming that the three high-head safety injection pumps, the two residual heat removal pumps, and the two containment spray pumps (Section 6.3) are running at their maximum capacity, the time sequence, from the time of the safety injection signal, for the changeover from injection to recirculation in the core of a large rupture is as follows:

1. In approximately 15 min, sufficient water has been delivered to provide the required net positive suction head to start the recirculation pumps.

2. In approximately 20 min, (a) low-level alarms on the refueling water storage tank sound, and (b) the redundant containment and recirculation sump level indicators show the sump water level. The alarm(s) serve to alert the operator to start the switchover to the recirculation mode. The redundant containment and recirculation sump level indicators provide verification that the refueling water storage tank water has been delivered during the injection phase, as well as giving consideration to the case of a spurious (i.e., early) refueling water storage tank low-level alarm. The operator would see on the control board that the redundant sump level indications are at the appropriate points; switchover via the eight-switch sequence is performed at that time.

3. With the initiation of the eight-switch sequence (i.e., switch No. 1), only one spray pump will continue in operation. This spray pump will continue to draw from the refueling water storage tank until the level drops below 2 feet.

Recirculation pump motors are 2-ft 2-in. above the highest water level after the addition of the injected water to the spilled coolant. The changeover from injection to recirculation takes place when the level indicator or level alarms on the refueling water storage tank indicate that the fluid has been injected. The level indicators in the containment sump will also verify that the level is sufficient within the containment. The sequence is followed regardless of which power supply is available. All switches are grouped together on the safeguard control panel. The component position lights verify when the function of a given switch has been completed. Should an individual component fail to respond, the operator can take corrective action to secure appropriate response from the backup component. The manual switchover by the operator:

1. Terminates safety injection signal in order that the control logic permits manipulation of the system (at any time following completion of the auto-start sequence).

2. Closes switches one and three (removes and isolates unnecessary loads from the diesels).

Switch One:

a. Trips one (i.e., pump 22) of three high-head safety injection pumps if all three are operating (no action if two are operating), and isolates the pump suction to Chapter 6, Page 15 of 174 Revision 22, 2010

IP2 FSAR UPDATE to the refueling water storage tank if the tripped pump is the middle safety injection pump (i.e. pump 22).

b. Trips one containment spray pump if both are operating (no action if one is operating).

c. Closes isolation valves at the inoperative spray pump discharge.

Switch Three:

a. Trips both residual heat removal pumps.

b. Sends close signals to isolation valves in the residual heat removal pump suction and discharge headers, which are administratively reenergized later in the sequence. (Technical Specifications require the motor operators for these valves to be deenergized.)

3. Closes switch two (establishes cooling flow for residual heat removal heat exchangers)

a. Starts one service water pump, non-essential header (the second or third pump is given a start signal if the first or second pump fails to start).

b. Starts one component cooling pump (the second or third pump is given a start signal if the first or second pump fails to start).

4. Isolates one RHR heat exchanger flow path. (if both are open)

5. Closes switch four (initiates internal recirculation flow).

a. Opens valves on discharge of recirculation pumps.

b. Starts recirculation pump 21 (if pump 21 fails to start, uses manual start on pump 22). (Pump 22 control switch is adjacent to switch four).

6. Checks flow to reactor coolant system via the low-head injection lines to ensure minimum flow requirements are established. If minimum flow requirements are established, the closes switch seven and switch eight to establish low-head recirculation.

If minimum flow requirements are not established, then closes switch six and switch eight to establish high-head recirculation.

7. Close switch six (supplies recirculation for reactor coolant system pressures greater than 150 psig, which impedes flow via the low-head injection lines).

a. Opens valves 888A and 888B to provide a flow path from the recirculation pump discharge to the high-head safety injection pump suction.

b. Activates the low-pressure alarm circuit off of PT-947. If for some reason PT-947 alarm is not activated by this switch (RS-6), the operator can switch "HI Chapter 6, Page 16 of 174 Revision 22, 2010

IP2 FSAR UPDATE HEAD PUMP LO SUCTION PRESS ALARM" to activate this alarm. This latter switch is on the safeguards panel.

c. Closes valves 842 and 843 (high-head pump test line) (if their control feed interlock switches were first placed to the "OFF" position), and sends a close signal to valves 746 and 747 (residual heat removal heat exchanger discharge).

8. Closes switch seven (removes the two running safety injection pumps from service since they are no longer needed).

9. Closes switch eight (completes the isolation of the safety injection system and containment spray system lines from the refueling water storage tank).

a. Closes the valve on the spray test line.

b. Sends a close signal to the valve in the safety injection pumps suction line from the refueling water storage tank, which is administratively reenergized later in the sequence. (Control power for this valve is deenergized in accordance with Technical Specifications requirements).

10. Close switch five (Establishes additional cooling capability if adequate power is available i.e. all diesel breakers are either open or racked out, or at least one breaker from each of the three diesels is racked in and closed).

a. Starts second service water pump, non-essential header (the third pump is given the start signal if the second pump fails to start).

b. If (a) completed, starts second component cooling pump (the third pump is given the start signal if the second pump fails to start).

c. If (b) completed, starts recirculation pump 22 (unless already running). [Note:

running two (2) recirculation pumps is restricted to Low Head Recirculation. If High Head Recirculation is required, operator action is taken to prevent two recirculation pumps from operating simultaneously.] Although the listed switches are manual, each automatically causes the operations listed. An indicating lamp is provided to show the operator when the operations of a given switch have been performed. In addition, lamps indicating completion of the individual functions for a given switch are provided. These lamps are adjacent to the switches. Should an individual component fail to respond, the operator can take corrective action to secure appropriate response from controls within the control room. Remote-operated valves for the injection phase of the safety injection system (Plant Drawings 9321-2735 and 235296 [Formerly Figure 6.2-1]), which are under manual control (i.e., valves, which normally are in their ready position and do not receive a safety injection signal) have their positions indicated on a common portion of the control board. At any time during operation when one of these valves is not in the ready position for injection, it is shown visually on the board. Table 6.2-2 lists the instrumentation readouts on the control board and assessment panel, which the operator can monitor during recirculation. In addition, an audible annunciation alerts the operator to the condition. Chapter 6, Page 17 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.2.2.1.5 Location of the Major Components Required for Recirculation The residual heat removal pumps are located in the residual heat removal pump room, which is on the basement floor of the primary auxiliary building (elevation 15-ft). The residual heat exchangers are on a platform above the basement floor in the containment building (elevation 66-ft). The recirculation pumps are directly above the recirculation sump in the containment building (elevation 46-ft). The component cooling pumps and heat exchangers are located in the primary auxiliary building (elevations 68-ft and 80-ft, respectively). The auxiliary component cooling water pumps are located in the Mezzanine area of the piping penetration area at elevation 68-ft. The service water pumps are located at the river water intake structure, and the redundant piping to the component cooling heat exchangers is run underground, until it surfaces just prior to its penetrating the Primary Auxiliary Building exterior wall. 6.2.2.2 Steam Line Break Protection A large break of a steam system pipe rapidly cools the reactor coolant causing insertion of reactivity into the core and the depressurization of the system. Compensation is provided by the injection of boric acid from the refueling water storage tank. The analysis of the steam line rupture accident is presented in section 14.2.5. 6.2.2.3 Components All associated components, piping, structures, and power supplies of the safety injection system are designed to seismic Class I criteria. All components inside the containment are capable of withstanding or are protected from differential pressure that may occur during the rapid pressure rise to 47 psig in 10 sec. Electrical equipment that has been determined to be important to safety and located in potentially harsh environments are environmentally qualified to ensure performance of their safety function under postaccident temperature, pressure, and humidity conditions. Emergency core cooling components are either austenitic or an equivalent corrosion-resistant stainless steel, and hence, are compatible with the spray solution over the full range of exposure in the postaccident regime. Corrosion tests performed with simulated spray indicated negligible attack, both generally and locally, in stressed and unstressed stainless steel at containment and emergency core cooling system conditions. These tests are discussed in Reference 1. The quality standards of all safety injection system components are given in summary form in Table 6.2-3. Chapter 6, Page 18 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.2.2.3.1 Accumulators The accumulators are pressure vessels filled with borated water and pressurized with nitrogen gas. During normal plant operation, each of the four accumulators is isolated from the reactor coolant system by two check valves in series. Should the reactor coolant system pressure fall below the accumulator pressure, the check valves open and borated water is forced into the cold legs of the reactor coolant system. Mechanical operation of the swing-disk check valves is the only action required to open the injection path from the accumulators to the core via the cold leg. The level of borated water in each accumulator tank is adjusted remotely as required during normal plant operations. Refueling water is added using the accumulator topping pump (or safety injection pump 22 or 23). Water level is reduced by draining to the reactor coolant drain tank or through the chemistry sampling panel. Samples of the solution in the tanks are taken at the sampling station for periodic checks of boron concentration. Pressure is adjusted by adding nitrogen as required. The accumulators are passive engineered safety features since the gas forces injection and no external source of power or signal transmission is needed to obtain fast-acting, high-flow capability when injection is required. One accumulator is attached to each of the four cold legs of the reactor coolant system. The design capacity of the accumulators is based on the assumption that flow from one of the accumulators spills onto the containment floor through the ruptured loop. The flow from the three remaining accumulators provides sufficient water to fill the volume outside of the core barrel below the nozzles, the bottom plenum, and approximately one-half the core. To assure the independence of the accumulators from each other, operating procedures require that only one liquid fill valve and only one nitrogen stop valve can be open at a time when reactor temperature is equal to or greater than 350°F. The accumulators are carbon steel, internally clad with stainless steel and designed to ASME Section III, Class C. Connections for remotely draining or filling the fluid space, during normal plant operation, are provided. Redundant level and pressure indicators are provided with readouts on the control board. Each indicator is equipped with high- and low-level alarms. The accumulator design parameters are given in Table 6.2-4. 6.2.2.3.2 Boron Injection Tank The boron injection tank has been removed. 6.2.2.3.3 Refueling Water Storage Tank In addition to its normal duty to supply borated water to the refueling canal for refueling operations, this tank provides borated water to the safety injection pumps, the residual heat removal pumps, and the containment spray pumps for the Loss-of-Coolant Accident. These systems are kept sufficiently filled with water to ensure the systems remain operable and perform properly. During plant operation, this tank is aligned to these pumps. Chapter 6, Page 19 of 174 Revision 22, 2010

IP2 FSAR UPDATE The capacity of the refueling water storage tank is based on the requirement for filling the refueling canal; a minimum of 345,000 gal is required by the Technical Specifications to be maintained in the refueling water storage tank. This capacity provides an amount of borated water to assure:

1. A sufficient volume of water on the floor to permit the initiation of recirculation (246,000 gal).

2. A volume water sufficient to allow time for completing the switchover to recirculation and securing High Head Injection and Containment Spray Flow from the RWST (60,000 gal).

3. A sufficient volume of water to allow for instrument inaccuracies, additional margin, and for water that is physically unavailable from the bottom of the tank (39,000 gal).

4. The RWST water volume injected into containment, when added to accumulator discharge to the reactor coolant system, assures no return to criticality with the reactor at cold shutdown and no control rods inserted into the core.

The water in the tank is borated to ensure a minimum shutdown margin as discussed in Section 14.1.5.2.1. The maximum boric acid concentration is approximately 1.4 wt percent boric acid. At 32oF, the solubility limit of boric acid is 2.2-percent. Therefore, the concentration of boric acid in the refueling water storage tank is well below the solubility limit at 32oF. Steam heating is provided for the tank, and the outside lines are heat traced to maintain the temperature above freezing. Each of two redundant channels of refueling water storage tank level instrumentation provide level indication and low-level alarms in the central control room. In addition, a third instrument provides local level indication. The design parameters are presented in Table 6.2-6. 6.2.2.3.4 Pumps The three high-head safety injection pumps for supplying borated water to the reactor coolant system are horizontal centrifugal pumps driven by electrical motors. Parts of the pump in contact with borated water are stainless steel or an equivalent corrosion-resistant material. Each safety injection pump is sized at 50-percent of the capacity required to meet the design criteria outlined in Section 6.2.1. The design parameters are presented in Table 6.2-7; Figure 6.2-6 gives the performance characteristics of these pumps. A minimum flow bypass line is provided on each pump discharge to recirculate flow to the refueling water storage tank in the event the pumps are started with the normal flow paths blocked. Valves in the minimum flow bypass line (which are normally open) are equipped with motor operators. If either valve closes, an alarm annunciates in the control room. Power is de-energized to prevent spurious valve closure. The safety injection pump bearing oil is cooled by CCW circulating water pumps using component cooling water as a heat sink. The CCW circulating water pumps are directly Chapter 6, Page 20 of 174 Revision 22, 2010

IP2 FSAR UPDATE connected to the injection pump motor shaft. The pump seals are designed to operate during the injection phase without forced component cooling water flow. During the recirculation phase, cooling water is supplied by the component cooling system or alternately from the primary water system. Emergency backup is available via connections to the city water system. The two residual heat removal (low-head) pumps of the auxiliary coolant system are used to inject borated water at low pressure into the reactor coolant system. The two recirculation pumps are used to recirculate fluid from the recirculation sump back to the reactor, to the spray headers, or to suction of the safety injection pumps. The recirculation pumps will only be required to operate during the recirculation phase. In addition, the recirculation pumps are required to be operable for a period of one year. Per the 12/06/04 NRC generic SER on NEI 04-07 (Reference 5), evaluations of PWR post accident emergency recirculation performance to address the potential impact of debris blockage per GSI-191 (Reference 4) and Generic Letter 2004-02 (Reference 3) will use a mission time of 30 days. All four of these pumps are of the vertical centrifugal type driven by electric motors. The recirculation pumps are open suction, well-type pumps. Parts of the pumps, which contact the borated water solution during recirculation are stainless steel or an equivalent corrosion-resistant material. A minimum flow bypass line is provided on the discharge of the residual heat removal heat exchangers to recirculate cooled fluid to the suction of the residual heat removal pumps should these pumps be started with their normal flow paths blocked. There are two normally open motor-operated valves in this line. The control power to the two normally open motor operated valves is locked open. The emergency procedures ensure that the RHR pumps are not run in parallel for extended time periods with RCS pressure at or above their shutoff head. A minimum flow bypass, discharging back into the recirculation sump, is provided to protect the recirculation pumps should these flow paths be blocked. Valves in these lines are manually operated and are in the open position during normal plant operation. Figures 6.2-7 and 6.2-8 give the performance characteristics of these pumps. The design parameters are presented in Table 6.2-7. The recirculation pump motors are air-to-water cooled in a similar manner as the containment cooling fan motors described in section 6.4.2.2.5, item 2. The motor fans are integral to the recirculation pump motor shafts. Cooling water to the motor heat exchanger is component cooling water. The sump water cools the pump bearings. The two auxiliary component cooling water pumps are started during the injection phase. However, their function during this phase is not required to protect the recirculation pump motors from the containment atmosphere. Since the recirculation pumps do not operate during injection, their motors do not experience any self-heating. Without this self-induced heat up, the motors functional capabilities and EQ characteristics have been shown in motor qualification testing to be unaffected by the post-LOCA environment. Even with an auxiliary component cooling water pump running, effectively no motor cooling occurs during the injection phase because the air circulating fans integral to the motor that drive cooling air through the heat exchanger and motor are not operating. Details of the component cooling pumps and service water pumps, which serve the safety injection system, are presented in Section 9.3 and 9.6, respectively. The pressure-retaining parts of the high-head safety injection pumps are castings conforming to ASTM A-296, Grade CA-15 or ASME SA-487, Grade CA-6NM. The pressure-retaining parts of the residual heat removal pumps and the recirculation pumps are castings conforming to ASTM A-351, Grade CF-8A (chromium content 21.0 to 22.5%) and ASTM A-351, Grade CF-8, respectively. Stainless steel forgings are procured per ASTM A-182, Grade F304 or F316, or ASTM A-336, Class F8 or F8M, and stainless steel plate is constructed to ASTM A-240, Type 304 or 316. All bolting material conforms to ASTM A-193. Material such as Monel is used at Chapter 6, Page 21 of 174 Revision 22, 2010

IP2 FSAR UPDATE points of close running clearances in the pumps to prevent galling and to ensure continued performance ability in high-velocity areas subject to erosion. All pressure-retaining parts of the pumps were chemically and physically analyzed, and the results were checked to ensure conformance with the applicable ASTM specification. In addition, all pressure-retaining parts of the pump were liquid-penetrant inspected in accordance with Appendix VIII of Section VIII of the ASME Boiler and Pressure Vessel Code. The acceptance standard for the liquid-penetrant test is USAS B31.1, Code for Pressure Piping, Case N-10. The pump design was reviewed with special attention to the reliability and maintenance aspects of the working components. Specific areas include the evaluation of the shaft seal and bearing design to determine whether adequate allowances had been made for shaft deflection and clearances between stationary parts. Where welding of pressure-containing parts was necessary, a welding procedure including joint detail was submitted for review and approval by Westinghouse. The procedure included evidence of qualification necessary for compliance with Section IX of the ASME Boiler and Pressure Vessel Code Welding Qualifications. This requirement also applied to any repair welding performed on pressure-containing parts. The pressure-containing parts of the pump were assembled and hydrostatically tested to 1.5 times the design pressure for 30 min. Each pump was given a complete shop performance test in accordance with Hydraulic Institute Standards. The pumps were run at design flow and head, shutoff head, and three additional points to verify performance characteristics. Where net positive suction head is critical, this value was established at design flow by means of adjusting suction pressure. An accumulator topping pump is provided to fill the accumulators rather than using safety injection pump 22 or 23. The pump is a double-diaphragm type with a capacity of approximately 5 gpm (293 gph). It is located in the safety injection system pump area and is operated from a local key-locked push button switch. The topping pump is capable of withstanding the safe shutdown earthquake but does not operate following safety injection actuation. 6.2.2.3.5 Heat Exchangers The two residual heat exchangers of the auxiliary coolant system are sized for the cooldown of the reactor coolant system. Table 6.2-8 gives the design parameters of the heat exchangers. During the recirculation phase following a loss-of-coolant-accident, only one residual heat removal heat exchanger is required to ensure that heat removal requirements from the core and containment are met. The ASME Boiler and Pressure Vessel Code has strict rules regarding the wall thicknesses of all pressure-containing parts, material quality assurance provisions, weld joint design, radiographic and liquid-penetrant examination of materials and joints, and hydrostatic testing of the unit as well as requiring final inspection and stamping of the vessel by an ASME Code inspector. The designs of the heat exchangers also conform to the requirements of the Tubular Exchanger Manufacturers Association (TEMA) for Class R heat exchangers. Class R is the most rugged Chapter 6, Page 22 of 174 Revision 22, 2010

IP2 FSAR UPDATE class of TEMA heat exchangers and is intended for units where safety and durability are required under severe service conditions. Items such as tube spacing, flange design, nozzle location, baffle thickness and spacing, and impingement plate requirements are set forth by TEMA standards. In addition to the above, additional design and inspection requirements were imposed to ensure rugged, high-quality heat exchangers such as the following:

1. Confined-type gaskets, general construction and mounting brackets suitable for the plant seismic design requirements.

2. Tubes and tube sheet capable of withstanding full shell-side pressure and temperature with atmospheric pressure on the tube side, ultrasonic inspection in accordance with Paragraph N-324.3 of Section III of the ASME Code of all tubes before bending, penetrant inspection in accordance with Paragraph N-627 of Section III of the ASME Code of all welds and hot- or cold-formed parts.

3. A hydrostatic test duration of not less than 30 min, the witnessing of hydro and penetrant tests by a qualified inspector, a thorough final inspection of the unit for good workmanship and the absence of any gouge marks or other scars that could act as stress concentration points, and a review of the radiographs and of the certified chemical and physical test reports for all materials used in the unit.

The residual heat exchangers are conventional vertical shell and U-tube type units. The tubes are seal welded to the tube sheet. The shell connections are flanged to facilitate shell removal for inspection and cleaning of the tube bundle. Each unit has an SA-285, Grade C, carbon steel shell; an SA-234 carbon steel shell end cap; SA-213, Type 304, stainless steel tubes; an SA-240, Type 304, stainless steel channel; an SA-240, Type 304, stainless steel channel cover; and an SA-240, Type 304, stainless steel tube sheet. 6.2.2.3.6 Valves All parts of valves used in the safety injection system in contact with borated water are austenitic stainless steel or an equivalent corrosion-resistant material. The motor operators on the injection line isolation valves are capable of rapid operation. All valves required for the initiation of safety injection or isolation of the system have remote position indication in the control room. Valving is specified for exceptional tightness. All valves, except those which perform a control function, are provided with backseats that are capable of limiting leakage. The estimated leakage of backseated valves outside containment is provided in Table 6.2-9. Those valves, which are normally open are backseated, except when operational considerations do not allow. [Note - The following valves may not be backseated based on operational requirements: 744, 850A, 850B, 851A, 851B, 883, 885A, 885B, 887A, 887B, 888A, 888B and 958.] Normally closed globe valves are installed with recirculation flow under the seat to prevent the leakage of recirculated water through the valve stem packing. Relief valves are totally enclosed. Control and motor-operated valves, 2.5-in. and above, which are exposed to recirculation flow, are generally provided with double-packed stuffing boxes and stem leakoff-connections that are piped to the waste disposal system. Chapter 6, Page 23 of 174 Revision 22, 2010

IP2 FSAR UPDATE The check valves that isolate the safety injection system from the reactor coolant system are installed immediately adjacent to the reactor coolant piping to reduce the probability of a safety injection line rupture causing a loss-of-coolant accident. A relief valve is installed in the safety injection pump discharge header discharging to the pressurizer relief tank to prevent overpressure in the lines that have a lower design pressure than the reactor coolant system. RV-855 is a thermal relief valve which protects the Safety Injection System piping and components from overpressurization due to thermal expansion of fluid in the system or from in-leakage of reactor coolant. The setpoint of RV-855 was changed to 1670 psig to ensure that the valve does not lift when operating the SI system at a pressure near the shutoff head of the SI pumps. The gas relief valves on the accumulators protect them from pressures in excess of the design value. 6.2.2.3.7 Motor-Operated Butterfly Valve (Containment Sump Valve) The pressure-containing parts (body, disks) of the valves employed in the safety injection system are designed per criteria established by USAS B16.5 or MSS SP-67 specifications. The materials of construction for these parts are procured per ASTM A182, F316 or A351, Grade CF8M or CF8. All material in contact with the primary fluid, except the packing and the liner, is austenitic stainless steel or an equivalent corrosion-resistant material. The liner is EPT-NORDEL (Du Pont). The pressure-containing cast components are radiographically inspected as outlined in ASTM E-71, Class 1 or Class 2. The body and disk are liquid-penetrant inspected in accordance with ASME Boiler and Pressure Vessel Code, Section VIII, Appendix VIII. The liquid-penetrant acceptable standard is as outlined in USAS B31.1, Case N-10. The entire assembled unit is hydrotested as outlined in MSS SP-67, with the exception that the test is maintained for a minimum period of 30 min. The motor operator is evaluated in accordance with the GL 89-10 Motor Operated Valve Program to assure its capability to meet the required stem torque for opening and closing. The shaft material is ASTM A276, Type 316, condition B, or precipitation hardened 17-4 pH stainless steel procured and heat treated to Westinghouse specifications. These materials are selected because of their corrosion-resistant, high-tensile properties, and their resistance to surface scoring by the packing. The motor operator is located above the maximum sump fluid level and therefore is never submerged. The motor operator is extremely rugged and is noted throughout the power industry for its reliability. The unit incorporates a hammer-blow feature that allows the motor to impact the disks away from the fore or backseat upon opening or closing. This hammer-blow feature not only impacts the disk but allows the motor to attain its operational speed. The valve is assembled, hydrostatically tested, seat-leakage tested (fore and back), operationally tested, cleaned, and packaged per specifications. All manufacturing procedures employed by the valve supplier such as welding, repair welding, and testing are submitted to Westinghouse for approval. The valve operator completes its cycle from one position to the other in approximately 120 sec. Chapter 6, Page 24 of 174 Revision 22, 2010

IP2 FSAR UPDATE Valves that must function against system pressure are designed such that they function with a pressure drop equal to full system pressure across the valve disk. 6.2.2.3.8 Motor-Operated Gate Valves The pressure-containing parts (body, bonnet, and disks) of the valves employed in the safety injection system are designed per criteria established by USAS B16.5 or MSS SP-66 specifications. The materials of construction for these parts are procured per ASTM A182, F316 or A351, Grade CF8M or CF8. All material in contact with the primary fluid, except the packing, is austenitic stainless steel or an equivalent corrosion-resistant material. The pressure-containing cast components are radiographically inspected as outlined in ASTM E-71, Class 1 or Class 2. The body, bonnet, and disks are liquid-penetrant inspected in accordance with ASME Boiler and Pressure Vessel Code, Section VIII, Appendix VIII. The liquid-penetrant acceptable standard is as outlined in USAS B31.1, Case N-10. When a gasket is employed, the body-to-bonnet joint is designed per ASME Boiler and Pressure Vessel Code, Section VIII, or USAS B16.5 with a fully trapped, controlled compression, spiral wound gasket with provisions for seal welding, or of the pressure seal design with provisions for seal welding. The body-to-bonnet bolting and nut materials are procured per ASTM A193 and A194, respectively. The entire assembled unit is hydrotested as outlined in MSS SP-61, with the exception that the test is maintained for a minimum period of 30 min. The seating design is of the Darling parallel disk design, the Crane flexible wedge design, or the equivalent. These designs have the feature of releasing the mechanical holding force during the first increment of travel. Thus, the motor operator is evaluated in accordance with the GL 89-10 Motor Operated Valve Program to assure its capability to meet the required stem thrust for opening and closing. The disks are guided throughout the full disk travel to prevent chattering and to provide ease of gate movement. The seating surfaces are hard faced (Stellite No. 6 or equivalent) to prevent galling and reduce wear. The stem material is ASTM A276, Type 316, condition B, or precipitation hardened 17-4 pH stainless steel procured and heat treated to Westinghouse specifications. These materials are selected because of their corrosion-resistant, high-tensile properties, and their resistance to surface scoring by the packing. The valve stuffing box was originally designed with a lantern ring leakoff connection with a minimum of a full set of packing below the lantern ring and a maximum of one-half of a set of packing above the lantern ring; a full set of packing is defined as a depth of packing equal to 1-1/2 times the stem diameter. An alternate packing arrangement may be installed in these valves upon approval for substitution. Experience with designs utilizing live load and graphite packing has been favorable. The motor operator is extremely rugged and is noted throughout the power industry for its reliability. The unit incorporates a hammer-blow feature that allows the motor to impact the disks away from the fore or backseat upon opening or closing. This hammer-blow feature not only impacts the disk but allows the motor to attain its operational speed. The valve is assembled, hydrostatically tested, seat-leakage tested (fore and back), operationally tested, cleaned, and packaged per specifications. All manufacturing procedures employed by the valve supplier such as hard facing, welding, repair welding, and testing are submitted to Westinghouse for approval. Chapter 6, Page 25 of 174 Revision 22, 2010

IP2 FSAR UPDATE For those valves that must function on the safety injection signal, approximately 10-sec operation is required. For all other valves in the system, the valve operator completes its cycle from one position to the other in approximately 120 sec. Operating times greater than these values are permitted on a case by case basis if properly justified by an individual safety evaluation. Valves that must function against system pressure are designed such that they function with a pressure drop equal to full system pressure across the valve disk. 6.2.2.3.9 Manual Valves The stainless steel manual globe, gate, and check valves are designed and built in accordance with the requirements outlined in the motor-operated valve description above with the following exceptions:

1. Alternate materials, evaluated to be equivalent, have been used in some replacement valves.

2. Liquid-penetrant inspection of the body, bonnet, and disks to ASME V Article 6 with acceptance per ASME III has been used on some replacement valves.

The carbon steel valves are built to conform with USAS B16.5. The materials of construction of the body, bonnet, and disk conform to the requirements of ASTM A105, Grade II; A181, Grade II; or A216, Grade WCB or WCC. Alternate materials, evaluated to be equivalent, have been used in some replacement valves. The carbon steel valves pass only nonradioactive fluids and are subjected to hydrostatic test as outlined in MSS SP-61, except that the test pressure is maintained for at least 30 min/in. of wall thickness. Since the fluid controlled by the carbon steel valves is not radioactive, the double packing and seal weld provisions are not provided. 6.2.2.3.10 Accumulator Check Valves The pressure-containing parts of this valve assembly are designed in accordance with MSS SP-

66. All parts in contact with the operating fluid are of austenitic stainless steel or of equivalent corrosion-resistant materials procured to applicable ASTM or WAPD specifications. The cast pressure-containing parts are radiographed in accordance with ASTM E-94 and the acceptance standard as outlined in ASTM E-71, Class 1 or Class 2. The cast pressure-containing parts, machined surfaces finished hard facings, and gasket bearing surfaces are liquid-penetrant inspected per the ASME Boiler and Pressure Vessel Code, Section VIII, and the acceptance standard is as outlined in USAS B31.1, Code Case N-10. The final valve is hydrotested per MSS SP-61, except that the test pressure is maintained for at least 30 min. The seat leakage is conducted in accordance with the manner prescribed in MSS SP-61, except that the acceptable leakage is 2 cm3/hr-in. nominal pipe diameter.

The valve is designed with a low-pressure drop configuration with all operating parts contained within the body, which eliminates those problems associated with packing glands exposed to boric acid. The clapper arm shaft is manufactured from 17-4 pH stainless steel heat treated to Westinghouse specifications. The clapper arm shaft bushings are manufactured from Stellite No. 6 material. The various working parts are selected for their corrosion-resistant, tensile, and bearing properties. Chapter 6, Page 26 of 174 Revision 22, 2010

IP2 FSAR UPDATE The disk and seat rings are manufactured from forgings. The mating surfaces are hard faced with Stellite No. 6 to improve the valve seating life. The flexible disc-hinge connection permits the disc to completely contact the seat even if there is minor seat movement. The valves are intended to be operated in the closed position with a normal differential pressure across the disk of approximately 1650 psi. The valves remain in the closed position except for testing and safety injection. Since the valve will normally not be required to operate in the open condition, hence be subjected to impact loads caused by sudden flow reversal, it is expected that this equipment will not have difficulties performing its required functions. When the valve is required to function, a differential pressure of less than 25 psig will shear any particles that may attempt to prevent the valve from functioning. Although the working parts are exposed to the boric acid solution contained within the reactor coolant loop, a boric acid "freeze up is not expected with this low a concentration. The experience derived from the check valves employed in the safety injection system of the Carolina-Virginia Tube Reactor (CVTR) in a similar system has indicated that the system is reliable and workable. The CVTR emergency injection system, maintained at atmospheric conditions, was separated from the main coolant piping by one 6-in. check valve. A leak detection pit was provided in the CVTR to accumulate any leakage coming back through the check valve. A level alarm provided a signal on excessive leakage. There was a gas volume in the upper space of the loop. The pressure differential was 1500 psi and the system was stagnant. The valve was located 2 to 3-ft from the main coolant piping, which resulted in some heatup and cooldown cycling. The CVTR went critical late in 1963. Since that time and up to initial operation of Indian Point Unit 2, the level alarm in the detection pit had never gone off due to check valve leakage. 6.2.2.3.11 Relief Valves The accumulator relief valves are sized to pass nitrogen gas at a rate in excess of the accumulator gas fill line delivery rate. The relief valves will also pass water in excess of the maximum expected leak rate of 1.0 gpm identified in Technical Specifications with leakage being from the reactor coolant system into an accumulator through an accumulator discharge line. The accumulators are provided with level and pressure alarms. Operator response to inleakage causing these alarms to actuate would preclude the need for the relief valves to perform in a water relief capacity. The safety injection test line relief valve is provided to relieve any overpressure, that might build up in the high-head safety injection piping due to thermal expansion of fluid in the system or from leakage from the reactor coolant system past the SI header check valves. The valve will pass a nominal 15 gpm (2.25 x 105cm3/hr), which is far in excess of the manufacturing design in leakage rate from the reactor coolant system of 24 cm3/hr. 6.2.2.3.12 Leakage Limitations of Valves Valving is specified for exceptional tightness. Normally open valves have backseats that limit leakage as shown in Table 6.2-9. Normally closed globe valves are installed with recirculation flow under the seat to prevent stem leakage from the more radioactive fluid side of the seat. Chapter 6, Page 27 of 174 Revision 22, 2010

IP2 FSAR UPDATE Motor-operated valves, which are exposed to recirculation flow, are generally provided with double-packed stuffing boxes and stem leakoff connections that are piped to the waste disposal system. The specified leakage across the valve disk required to meet the equipment specification and hydrotest requirements is as follows:

1. Conventional globe - 3 cm3/hr-in. of nominal pipe size.

2. Gate valves - 3 cm3/hr-in. of nominal pipe size; 10 cm3/hr-in. for 300- and 150-lb USA Standard.

3. Motor-operated gate valves - 3 cm3/hr-in. of nominal pipe sizes; 10 cm3/hr-in. for 300- and 150-lb USA Standard.

4. Check valves - 3 cm3/hr-in. of nominal pipe size; 10 cm3/hr-in. for 300- and 150-lb USA Standard.

5. Accumulator check valves - 2 cm3/hr-in. of nominal pipe size; relief valves are totally enclosed.

Leakage from components of the recirculation loop including valves, is given in Table 6.2-9. 6.2.2.3.13 Piping All safety injection system piping in contact with borated water is austenitic stainless steel. Piping joints are welded except for the flanged connections at the safety injection pumps, the recirculation pumps, and valve 741A. The piping beyond the accumulator stop valves is designed for reactor coolant system conditions (2485 psig, 650oF). All other piping connected to the accumulator tanks is designed for 700 psig and 400oF. The safety injection pump and residual heat removal pump suction piping (210 psig at 300oF) from the refueling water storage meets net positive suction head requirements of the pumps. The safety injection high-pressure branch lines (1500 psig at 300oF) are designed for high-pressure losses to limit the flow rate out of the branch line, which may have ruptured at the connection to the reactor coolant loop. The system design incorporates the ability to isolate the safety injection pumps on separate headers such that full flow from at least one pump is ensured should a branch line break. The piping is designed to meet the minimum requirements set forth in (1) the USAS B31.1 Code (1955) for the Pressure Piping, (2) Nuclear Code Case N-7, (3) USAS Standards B36.10 and B36.19, and (4) ASTM Standards with supplementary standards plus additional quality control measures. Minimum wall thicknesses are determined by the USAS Code (1955) formula found in the power piping Section 1 of the USAS Code (1955) for Pressure Piping. This minimum thickness has been increased to account for the manufacturer's permissible tolerance of 12.5-percent on the Chapter 6, Page 28 of 174 Revision 22, 2010

IP2 FSAR UPDATE nominal wall. Purchased pipe and fittings have a specified nominal wall thickness that is no less than the sum of that required for pressure containment, mechanical strength, and manufacturing tolerance. Thermal and seismic piping flexibility analyses have been performed. Special attention is directed to the piping configuration at the pumps with the objective of minimizing pipe imposed loads at the suction and discharge nozzles. Piping is supported to accommodate expansion due to temperature changes during the accident. Piping between valves 730 and 731 (Line 10) has 6" thick insulation to assure operability of these valves during design basis accident conditions. Pipe and fittings materials are procured in conformance with all requirements of ASTM and USAS specifications. All materials are verified for conformance to specification and documented by certification of compliance to ASTM material requirements. Specifications impose additional quality control upon the suppliers of pipes and fittings as listed below.

1. Check analyses are performed on both the purchased pipe and fittings.

2. Pipe branch lines 2.5-in. and larger between the reactor coolant pipes and the isolation stop valves conform to ASTM A376 and meet the supplementary requirement S6 for ultrasonic testing. Fittings conform to the requirements of ASTM A403. Fittings 2.5-in. and above have requirements for ultrasonic testing inspection similar to S6 of A376.

Shop fabrication of piping subassemblies is performed by reputable suppliers in accordance with specifications that define and govern material procurement, detailed design, shop fabrication, cleaning, inspection, identification, packaging, and shipment. Welds for pipes sized 2.5-in. and larger are butt welded. Reducing tees are used where the branch size exceeds one-half of the header size. Branch connections of sizes that are equal to or less than one-half of the header size are of a design that conforms to the USAS rules for reinforcement set forth in the USAS B31.1 Code for Pressure Piping. Bosses for branch connections are attached to the header by means of full penetration welds. All welding is performed by welders and welding procedures qualified in accordance with the ASME Boiler and Pressure Vessel Code, Section IX, Welding Qualifications. The shop fabricator is required to submit all welding procedures and evidence of qualifications for review and approval prior to release for fabrication. All welding materials used by the shop fabricator must have prior approval. All high-pressure piping butt welds containing radioactive fluid at greater than 600°F temperature and 600 psig pressure or equivalent are radiographed. The remaining piping butt welds are randomly radiographed. The technique and acceptance standards are those outlined in UW-51 of the ASME Boiler and Pressure Vessel Code, Section VIII. In addition butt welds are liquid-penetrant examined in accordance with the procedure of ASME Boiler and Pressure Vessel Code, Section VIII, Appendix VIII, and the acceptance standard as defined in the USAS Nuclear Code Case N-10. Finished branch welds are liquid-penetrant examined on the outside, and where size permits, on the inside root surfaces. Chapter 6, Page 29 of 174 Revision 22, 2010

IP2 FSAR UPDATE A postbending solution anneal heat treatment is performed on hot-formed stainless steel pipe bends. Completed bends are then completely cleaned of oxidation from all affected surfaces. The shop fabricator is required to submit the bending, heat treatment, and cleanup procedures for review and approval-prior to release for fabrication. General cleaning of completed piping subassemblies (inside and outside surfaces) is governed by basic ground rules set forth in the specifications. For example, these specifications prohibit the use of hydrochloric acid and limit the chloride content of service water and demineralized water. The packaging of the piping subassemblies for shipment is done so as to preclude damage during transit and storage. Openings are closed and sealed with tight fitting covers to prevent the entry of moisture and foreign material. Flange facings and weld end preparations are protected from damage by means of wooden cover plates and securely fastened in position. The packing arrangement proposed by the shop fabricator is subject to approval. 6.2.2.3.14 Pump and Valve Motors Outside Containment Motor electrical insulation systems are supplied in accordance with IEEE, and NEMA standards and are tested as required by standards. Temperature rise design selection is such that normal long life is achieved even under accident loading conditions. Criteria for motors of the safety injection system require that under any anticipated mode of operation the motor nameplate rating is not exceeded. The pump motors have a 1.15 service factor for normal operation. Design and test criteria ensure that motor loading does not exceed the application criteria. 6.2.2.3.15 Pump and Valve Motors Inside Containment Motors for the recirculation pumps were originally specified to operate in an ambient condition of saturated steam of 270°F and 47 psig pressure for 1 day, followed by indefinite operation at 155°F and 5 psig in a steam atmosphere. These ambient conditions and operating times have been updated and are maintained by the ongoing Environmental Qualification Program discussed in Section 7.1.4. As part of this program, the recirculation pump motors are qualified to withstand containment environmental conditions following the loss of coolant accident so that the pumps can perform their required function during the recovery period (one year). These motors are of a similar design as the containment fan cooler motors. Refer to Section 6.4.2.2.5 for a description and evaluation of the motor design. Per the 12/06/04 NRC generic SER on NEI 04-07 (Reference 5), evaluations of PWR post accident emergency recirculation performance to address the potential impact of debris blockage per GSI-191 (Reference 4) and Generic Letter 2004-02 (Reference 3) will use a mission time of 30 days. The motors for the valves inside containment are designed to withstand containment environment conditions following the loss-of-coolant accident so that the valves can perform the required function during the recovery period. Periodic operation of the motors and tests of the insulation ensure that the motors remain in a reliable operating condition. Chapter 6, Page 30 of 174 Revision 22, 2010

IP2 FSAR UPDATE Although the motors that are provided only to drive engineered safety features equipment are normally run only for tests, the design loading and temperature rise limits are based on accident conditions. Normal design margins are specified for these motors to make sure the expected lifetime includes allowance for the occurrence of accident conditions. 6.2.2.3.16 Valve Motor Operators Environmental Qualification As part of the original plant design, a program of environmental qualifications was performed on valve motor operators important to plant safety. Tests to demonstrate the adequacy of valve motor operators to be functional after exposure to temperature, pressure, and radiation were conducted in two groups. The first group test was the exposure of valve motor operators to both temperature and pressure. Two suppliers, Philadelphia Gear Corporation Limitorque Division and Crane Company Teledyne Division, conducted simulated containment pressure and temperature tests as follows with pressure and temperature similar to that predicted for the incident:

1. Operator located inside a pressure vessel with the operator exposed to approximately 330°F at 90 psig.

2. Operator cycled approximately 3 times under simulated valve operating loads.

3. Pressures and temperatures reduced in step change to 285°F at 60 psig, 219°F at 20 psig, and 152°F at atmosphere or less.

4. Operator cycled approximately 3 times at each of the levels of change. Full recordings of pertinent data were taken throughout the tests.

5. Unit was examined after completion of test and operating data compared to data prior to exposure.

The second group test was the radiation test on a motor from the valve operator.

1. Two production line motors were used for this test; one exposed to 1.5 x 108 rads of gamma radiation for an approximate period of 1 month, the other motor used for the final comparative analysis.

2. Both units were tested for coil resistance, insulation meggering both before and after motor vibration, and reversing operations.

More recently, a program of environmental requalifications of items important to plant safety has been initiated using the "Division of Operating Reactors" or NUREG-0588 guidelines. See Section 7.1.4 for a discussion of this ongoing program. In response to IE Information Notice 86-03, all limitorque motor operators on the EQ Master List (see Section 7.1.4) were inspected and serviced to assure that wiring, limit switches and torque switches have been environmentally qualified. Chapter 6, Page 31 of 174 Revision 22, 2010

IP2 FSAR UPDATE In response to the IE Bulletin 85-03, the operability of key safety Motor Operated Valves was verified with associated full differential pressure. 6.2.2.4 Electrical Supply Details of the normal and emergency power sources for the safety injection system are presented in Chapter 8. 6.2.2.5 Protection Against Dynamic Effects The injection lines penetrate the containment adjacent to the primary auxiliary building. For most of the routing, these lines are outside the crane wall, and hence are protected from missiles originating within these areas. Each line penetrates the crane wall near the injection point to the reactor coolant pipe. In this manner, maximum separation and hence protection it provided in the coolant loop area. Coolant loop supports are designed to restrict the motion in one loop due to rupture in another loop to about one-tenth of an inch, whereas the attached safety injection piping can sustain a 3-in. displacement without exceeding the working stress range. The analysis assumes that the injection flow to the ruptured loop is spilled on the containment floor. In 1989, the NRC approved changes to the design basis with respect to dynamic effects of postulated primary loop ruptures, as discussed in Section 4.1.2.4 All hangers and anchors are designed in accordance with USAS B31.1, Code for Pressure Piping. This code provides minimum requirements on materials, design, and fabrication with ample safety margins for both dead and dynamic loads over the life of the equipment. Concrete missile barriers, bumpers, walls, and other concrete structures are designed in accordance with ACI 318-63, Building Code Requirements for Reinforced Concrete. Specifically, these standards require the following:

1. All materials used are in accordance with ASTM specifications that establish quality levels for the manufacturing process, minimum strength properties, and for test requirements that ensure compliance with the specifications.

2. Welding processes and welders must be qualified for each class of material welded and for types and positions of welds.

3. Maximum allowable stress values are established, which provide an ample safety margin on both yield strength and ultimate strength.

6.2.3 Design Evaluation 6.2.3.1 Range of Core Protection The measure of effectiveness of the safety injection system is the ability of the pumps and accumulators to keep the core flooded or to reflood the core rapidly when the core has been uncovered for postulated large area ruptures. The result of this performance is to limit sufficiently any increase in clad temperature below a value where emergency core cooling objectives are met (Section 6.2.1). The sequence of events involving safety injection actuation for small and large breaks of a reactor coolant pipe are presented in Section 14.3.2. Chapter 6, Page 32 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.2.3.2 System Response To provide protection for large area ruptures in the reactor coolant system, the safety injection system must respond rapidly to reflood the core following the depressurization and core voiding that is characteristic of large area ruptures. The accumulators act to perform the rapid reflooding function with no dependence on the normal or emergency power sources and also with no dependence on the receipt of an actuation signal. The operation of this system with three of the four available accumulators delivering their contents to the reactor vessel (one accumulator spilling through the break) prevents fuel clad melting and limits metal-water reaction to an insignificant amount (<1-percent). The function of the safety injection and residual heat removal pumps is to complete the refill of the vessel and ultimately return the core to a subcooled state. Moreover, there is sufficient excess water delivered by the accumulators to tolerate a delay in starting the pumps. Initial response of the injection systems is automatic, with appropriate allowances for delays in the actuation of circuitry and active components. The active portions of the injection systems are automatically actuated by the safety injection signal (Chapter 7). In addition, manual actuation of the entire injection system and individual components can be accomplished from the control room. In the analysis of system performance, delays in reaching the programmed trip points and in the actuation of components are conservatively established on the basis that only emergency onsite power is available. The starting sequence of the safety injection and residual heat removal pumps and the related emergency power equipment is discussed in sections 7.2 and 8.2.3.4 and their analyzed performance is discussed in the various Chapter 14 safety analyses. 6.2.3.3 Single-Failure Analysis A single active failure analysis is presented in Table 6.2-10. All credible active system failures are considered. This analysis is based on the worst single failure (generally a pump failure) in both the safety injection and residual heat removal pumping systems. The analysis shows that the failure of any single active component will not prevent fulfilling the design function. The analysis of the loss-of-coolant accident presented in Section 14.3 is consistent with this single-failure analysis. In addition to active failures, an alternative flow path is available to maintain core cooling if any part of the recirculation flow path becomes unavailable due to a single passive failure. This is evaluated in Table 6.2-11. The procedure followed to establish the alternative flow path also isolates the spilling line. A valve is provided in the containment recirculation line to the residual heat removal pumps to isolate this line should it be required. Therefore, the ECCS design incorporates redundancy of components such that neither a single active component failure during the injection phase nor an active or passive failure during the recirculation phase will degrade the ECCS function. Only active failures are assumed to occur within the first 24 hours following the initiating event. Failure analyses of the component cooling and service water system under loss-of-coolant accident conditions are described in Sections 9.3 and 9.6, respectively. Chapter 6, Page 33 of 174 Revision 22, 2010

IP2 FSAR UPDATE 6.2.3.4 Reliance on Interconnected Systems During the injection phase, the high-head safety injection pumps do not depend on any portion of other systems, with the exception of the suction line from the refueling water storage tank and the component cooling loop as a heat sink for bearing and lube oil cooling. During the recirculation phase of the accident for small breaks, suction to the high-head safety injection pumps is provided by the recirculation pumps or, should backup capability be required, the residual heat removal pumps. The residual heat removal (low-head) pumps are normally used during reactor shutdown operations. Whenever the reactor is at power, the pumps are aligned for emergency duty. 6.2.3.5 Shared Function Evaluation Table 6.2-12 is an evaluation of the main components, which have been previously discussed, and a brief description of how each component functions during normal operation and during the accident. 6.2.3.6 Passive Systems The accumulators are a passive safety feature in that they perform their design function in the total absence of an actuation signal or power source. The only moving parts in the accumulator injection train are in the two check valves. The working parts of the check valves are exposed to fluid of relatively low boric acid concentration. Even if some unforeseen deposition accumulated, a reversed differential pressure of about 25 psi can shear any particles in the bearing that may tend to prevent valve functioning. This is demonstrated by calculation. The isolation valve at each accumulator is only closed when the reactor is intentionally depressurized or momentarily for testing when pressurized. The isolation valve is normally open and an alarm in the control room sounds if the valve is inadvertently closed. It is not expected that the isolation valve will have to be closed due to excessive leakage through the check valves. The check valves operate in the closed position with a nominal differential pressure across the disk of approximately 1650 psi. They remain in this position except for testing or when called upon to function. Since the valves operate normally in the closed position and are therefore not subject to the abuse of flowing operation or impact loads caused by sudden flow reversal and seating, they do not experience any wear of the moving parts, and therefore function as required. When the reactor coolant system is being pressurized during the normal plant heatup operation, the check valves can be tested for leakage as soon as there is about 150 psi differential across the valve. This test confirms the seating of the disk and whether or not there has been an increase in the leakage since the last test. When this test is completed, the discharge line test valves are opened and the reactor coolant system pressure increase is continued. There should be no increase in leakage from this point on since increasing reactor coolant pressure increases the seating force and decreases the probability of leakage. Chapter 6, Page 34 of 174 Revision 22, 2010

IP2 FSAR UPDATE The accumulators can accept leakage from the reactor coolant system without effect on their availability. Table 6.2-13 indicates what inleakage rates, over a given time period, require readjusting the level at the end of the time period. In addition, these rates are compared to the maximum allowed leak rates for manufacturing acceptance tests (20 cm3/hr i.e., 2 cm3/hr/in). Inleakage at a rate of 5 cm3/hr-in., 2.5 times test, would require that the accumulator water volume be adjusted approximately once every 30 months. This would indicate that level adjustments can be scheduled for normal refueling shutdowns and that this work can be done at the operator's convenience. At a leak rate of 30 cm3/hr-in. (15 times the acceptance leak rate), the water level will have to be readjusted approximately once every 5 to 6 months. This readjustment will take about 2 hr maximum. The accumulators are located inside the reactor containment and protected from the reactor coolant system piping and components by a missile barrier. Accidental release of the gas charge in the accumulator would cause an increase in the containment pressure. This release of gas has been included in the containment pressure analysis for the large break loss-of-coolant accidents, (Section 14.3.3.3 and 14.3.5.1.1). During normal operation, the flow rate through the reactor coolant piping is approximately 5 times the maximum flow rate from the accumulator during injection. Therefore, fluid impingement on reactor vessel components during operation of the accumulator is not restricting. 6.2.3.7 Emergency Flow to the Core Special attention is given to factors that could adversely affect the accumulator and safety injection flow to the core. These factors are as follows:

1. Steam binding in the core, including flow blockage due to loop sealing.

2. Loss of accumulator water during blowdown.

3. Short circuiting of the accumulator from the core to another part of the reactor coolant system.

4. Loss of accumulator water through the breaks.

All of the above are considered in the analysis of the Loss of Coolant Accident which is discussed in Section 14.3. 6.2.3.8 External Recirculation Loop Leakage Table 6.2-9 summarizes the maximum potential leakage from the leak sources of the external recirculation loop, which goes through the residual heat removal pumps, a residual neat exchanger, and the high-head safety injection pumps. In the analysis, a maximum leakage is assumed from each leak source. For conservatism, 3 times the maximum expected leak rate from the pump seals was assumed, even though the seals are acceptance tested to essentially zero leakage, and a leakage of 10 drops/min was assumed from each flange although each flange would be adjusted to essentially zero leakage. The total maximum potential leakage resulting from all sources is 999 cm3/hr to the auxiliary building atmosphere and 21 cm3/hr to the drain tank. During external recirculation, significant margin exists between the design and operating conditions of the residual heat removal system components, as shown in Table 6.2-14. In Chapter 6, Page 35 of 174 Revision 22, 2010

IP2 FSAR UPDATE addition, during normal plant cooldown, operation of the residual heat removal system is initiated when the primary system pressure and temperature have been reduced to below 365 psig (the upper limit to prevent RHR system overpressurization) and 350°F, respectively. Even assuming a conservative maximum RHR System pressure of 232 psig and a conservative maximum RHR System temperature of 277°F during recirculation as shown in Table 6.2-14, significant margin also exists between normal operating and accident conditions. In view of the above margins, it is considered that the leakage rates tabulated in Table 6.2-9 are conservative. The radiological consequences of external recirculation loop leakage following a design basis accident are presented in Section 14.3.6.6. 6.2.3.9 Pump Net Positive Suction Head Requirements 6.2.3.9.1 Residual Heat Removal Pumps The net positive suction head (NPSH) of the residual heat removal pumps is evaluated for normal plant shutdown operation and the operation of both the injection and recirculation phases of the design-basis accident. The residual heat removal pumps are used as backup to the internal recirculation pumps in the event of failures to the normal recirculation path; this duty provides the pumps with the minimum NPSH condition. For the design case of [Deleted] one pump recirculating through one heat exchanger path, the available NPSH exceeds the NPSH required, assuming saturated fluid and no operator action to throttle back the flow. There is no postulated failure that requires both RHR pumps to operate in the recirculation phase. 6.2.3.9.2 Safety Injection Pumps The NPSH for the safety injection pumps is evaluated for both the injection and recirculation phase operation of the design-basis accident. The end of injection phase operation gives the limiting NPSH requirement; the NPSH available is determined from the elevation head and vapor pressure of the water in the refueling water storage tank and the pressure drop in the suction piping from the tank to the pumps. At the end of the injection phase, greater than 20-percent NPSH margin is available assuming all three pumps running together with two residual heat removal pumps at maximum flow conditions permitted by the system alignment. 6.2.3.9.3 Recirculation Pumps The NPSH for the recirculation pumps is evaluated for recirculation operation. The NPSH available is determined considering the elevation head of the water above the pump NPSH reference line (eye of the 1st stage impeller) in the sump, level drawdown due to the flow path in the containment, fluid temperature adjustments, and strainer head losses. The NPSH determination met the requirements of GL 2004-02. The containment water level is confirmed to be above the minimum level required for NPSH, prior to starting the recirculation pumps during the changeover from the injection phase to the recirculation phase. The RWST level is confirmed to be less than 2 feet prior to stopping the remaining operating containment spray pump and establishing simultaneous recirculation flow to the core and the spray headers. This maximizes the available NPSH to the recirculation pumps in this mode of operation. The internal recirculation pumps are conventional vertical condensate pumps and are of double suction design, requiring less NPSH. At the initiation of the recirculation phase, the NPSH requirement is met with one or both pumps operating at the pump design flow. When Chapter 6, Page 36 of 174 Revision 22, 2010

IP2 FSAR UPDATE simultaneous recirculation flow to the core and spray headers is established, the available NPSH requirement will be met at expected pump flows. [Deleted] 6.2.4 Minimum Operating Conditions The Technical Specifications establish limiting conditions regarding the operability of the system when the reactor is critical. 6.2.5 Inspections and Tests 6.2.5.1 Inspection All components of the safety injection system are inspected periodically to demonstrate system readiness. The pressure-containing components are inspected for leaks from pump seals, valve packing, flanged joints, and safety valves during system testing. Current requirements for safety injection system surveillance are discussed in Sections 3.5 and 5.5.6 of the facility Technical Specifications and in UFSAR Section 1.12, "Inservice Inspection and Testing Programs". 6.2.5.2 Preoperational Testing 6.2.5.2.1 Component Testing Preoperational performance tests of the components were performed in the manufacturer's shop. The pressure-containing parts of the pump were hydrostatically tested in accordance with Paragraph UG-99 of Section VIII of the ASME Code. Each pump was given a complete shop performance test in accordance with Hydraulic Institute Standards. The pumps were run at design flow and head, shutoff head, and at additional points to verify performance characteristics. Net positive suction head was established at design by means of adjusting suction pressure for a representative pump. This test was witnessed by qualified Westinghouse personnel. The remote-operated valves in the safety injection system are motor-operated. Shop tests for each valve included a hydrostatic pressure test, leakage tests, a check of opening and closing time, and verification of torque switch and limit switch settings. The ability of the motor operator to move the valve with the design differential pressure across the gate was demonstrated by opening the valve with an appropriate hydrostatic pressure on one side of the valve. The recirculation piping and accumulators were initially hydrostatically tested at 150-percent of design pressure. The service water and component cooling water pumps were tested prior to initial operation. 6.2.5.2.2 System Testing An initial functional test of the core cooling portion of the safety injection system was conducted during the hot-functional testing of the reactor coolant system before initial plant startup. The purpose of the initial systems test was to demonstrate the proper functioning of instrumentation Chapter 6, Page 37 of 174 Revision 22, 2010

IP2 FSAR UPDATE and actuation circuits and to evaluate the dynamics of placing the system in operation. This test was performed following the flushing and hydrostatic testing of the system. The functional test was performed with the water level below the safety injection setpoint in the pressurizer and with the reactor coolant system initially cold and at low pressure. The safety injection system valving was set initially to simulate the system alignment for plant power operation. To initiate the test, the safety injection block switch was moved to the unblock position to provide control power allowing the automatic actuation of the safety injection relays from low-pressure signals from the pressurizer instrumentation. Simultaneously, the breakers supplying outside power to the 480-V buses were tripped manually and operation of the emergency diesel system automatically commenced. The high-head safety injection pumps and the residual heat removal pumps were started automatically following the prescribed diesel loading sequence. The valves were operated automatically to align the flow path for injection into the reactor coolant system. The rising water level in the pressurizer provided indication of system delivery. Flow into the reactor coolant system terminated with the filling of the pressurizer, and the operation of the safety injection systems was terminated manually in the control room. This functional test provided information to confirm valve operating times, pump motor starting times, the proper automatic sequencing of load addition to the emergency diesels, and delivery rates of injection water to the reactor coolant system. The functional test was repeated for the various modes of operation needed to demonstrate performance at partial effectiveness, that is, to demonstrate the proper loading sequence with two of the three emergency diesels and to demonstrate the correct automatic starting of a second pump should the first pump fail to respond. These latter cases were performed without delivery of water to the reactor coolant system, but included starting of all pumping equipment involved in each test. The systems were accepted only after the demonstration of proper actuation and after the demonstration of flow delivery and shutoff head within design requirements. Flow was introduced into the reactor coolant loops through the accumulator discharge line to demonstrate the operability of the check valves and remotely actuated stop valve, and to confirm length to diameter (L/D) ratios of accumulator discharge lines used in the calculation. 6.2.5.3 Post-operational Testing 6.2.5.3.1 Component Testing Routine periodic testing of the safety injection system components and all necessary support systems at power is done. No inflow to the reactor coolant system will occur whenever the reactor coolant pressure is above the safety injection pump shutoff head. If such testing indicates a need for corrective maintenance, the redundancy of equipment in these systems permits such maintenance to be performed without shutting down or reducing load under conditions defined in the Technical Specifications. These conditions include such matters as the period within which the component is to be restored to service and the capability of the remaining equipment to meet safety limits within such a period. Chapter 6, Page 38 of 174 Revision 22, 2010

IP2 FSAR UPDATE Test Circuits are provided to examine periodically the leakage back through the check valves and to ascertain that these valves seat whenever the reactor system pressure is increased. The recirculation pumps are normally in a dry sump. These pumps can only be started and allowed to reach full speed with the plant at cold shutdown. Flow testing of these pumps is performed during refueling operations by filling the recirculation sump and directing the flow back to the sump through the valve on the discharge of the pump. The service water and component cooling pumps not running during normal operation may be tested by alternating with the operating pumps. The contents of the accumulators and the refueling water storage tank are sampled periodically to determine the boron concentration. 6.2.5.3.2 System Testing System testing is conducted during plant shutdown to demonstrate proper automatic operation of the safety injection system. A test signal is applied to initiate automatic action and verification made that the safety injection and residual heat removal pumps receive start signals. The test demonstrates the operation of the valves, pump circuit breakers and automatic circuitry. Isolation valves in the injection line will be blocked closed so that flow is not introduced into the reactor coolant system. The test is considered satisfactory if control board indication and visual observations indicate all components have operated and sequenced properly. The safety injection piping up to the final isolation valve is maintained sufficiently full of borated water at refueling water concentration while the plant is in operation to ensure the system remain operable and perform properly. The safety injection pumps recirculate refueling water through the injection lines via a small test line provided for this purpose. Flow in each of the safety injection headers and in the main flow line for the residual heat removal pumps is monitored by a local flow indicator. Pressure instrumentation is also provided for the main flow paths of the safety injection and residual heat removal pumps. Accumulator isolation valves are blocked closed for this test. The high pressure safety injection pumps are run and the variable orifices and injection line valves are adjusted to balance flowrates within the specified range. The eight-switch sequence for recirculation operation is tested to demonstrate proper sequencing of valves and pumps. The recirculation pumps are blocked from starting during this test. The external recirculation flow paths are hydrotested during periodic retests at the operating pressures. This is accomplished by running each pump, which could be used during external recirculation (safety injection and residual heat removal pumps) and checking the discharge and recirculation test lines. The suction lines are tested by running the residual heat removal pumps and opening the flow path to the safety injection pumps in the same manner as described above. During the above test, all system joints, valve packings, pump seals, leakoff connections, or other potential points of leakage can be visually examined. Valve gland packing, pump seals, and flanges are adjusted or replaced as required to reduce the leakage to acceptable proportions. For power-operated valves, final packing adjustments are made, and the valves are Chapter 6, Page 39 of 174 Revision 22, 2010

IP2 FSAR UPDATE are put through an operating cycle before a final leakage examination is made. The entire recirculation loop, except the recirculation line to the residual heat removal pumps, is pressurized during periodic testing of the engineered safety features components. The recirculation line to the residual heat removal pump is capable of being hydrotested during plant shutdown, and it is also leak-tested at the time of the periodic retests of the containment. REFERENCES FOR SECTION 6.2

1. M. J. Bell, et al., Investigations of Chemical Additives for Reactor Containment Spray Systems, WCAP-7153, Westinghouse Electric Corporation, March 1968.

2. Deleted

3. NRC Generic Letter 2004-02, Potential Impact of Debris Blockage on Emergency Recirculation During Design Basis Accidents at Pressurized-Water Reactors, dated September 13, 2004.

4. NRC Generic Safety Issue (GSI) 191, Assessment of Debris Accumulation on PWR Sump Performance.

5. NRC Generic SER dated 12/06/04 on the NEI 04-07 Guidance Report entitled, Pressurized Water Reactor Sump Performance Evaluation Methodology.

6. A. E. Lane, et al., Evaluation of Post Accident Chemical Effects in Containment Sump Fluids to Support GSI-191, WCAP-16530-NP-A, Westinghouse Electric Corporation LLC, March 2008.

7. T. S. Andreychek, et al., Evaluation of Downstream Sump Debris Effects in Support of GSI-191, WCAP-16406-P-A, Revision 1; Westinghouse Electric Corporation LLC, March 2008.

8. T. S. Andreychek, et al., Evaluation of Long-Term Cooling Considering Particulate, Fibrous and Chemical Debris in the Recirculating Fliud, WCAP-16793-NP, Revision 0; Westinghouse Electric Corporation LLC, May 2007.

TABLE 6.2-1 Safety Injection System - Code Requirements Component Code Refueling Water Storage Tank AWWA D100-65 Residual Heat Exchanger Tube Side ASME Section III Class C Chapter 6, Page 40 of 174 Revision 22, 2010

IP2 FSAR UPDATE Shell Side ASME Section VIII Accumulators ASME Section III Class C Valves USAS B16.5 (1955) Piping USAS B31.1 (1955) Chapter 6, Page 41 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-2 (Sheet 1 of 2) Instrumentation Readouts On The Control Board For Operator Monitoring During Recirculation Valves System Valve No. SIS MOV 1802 A, B SIS MOV 885 A, B SIS MOV 889 A, B SIS MOV 888 A,B SIS MOV 866 A, B, C, D SIS MOV 851 A, B SIS MOV 856 A, B, C, D,E,F SIS MOV 882 SIS MOV 842 SIS MOV 843 ACS MOV 744 ACS MOV 745 A,B ACS MOV 746 ACS MOV 747 ACS MOV 1810 ACS HCV 638 ACS HCV 640 Instruments System Channel No SIS FI 945 A, B SIS FI 946 A, B, C,D SIS FI 924 SIS FI 925 SIS FI 926 SIS FI 927 SIS LI 938 SIS LI 939 [Deleted] SIS LI 941 SIS LT 3300 SIS LT 3301 SIS LT 3302 [Deleted] SIS LT 3304 SIS PI 922 SIS PI 923 SIS PI 947 ACS FI 640 ACS LI 628 Chapter 6, Page 42 of 174 Revision 22, 2010

IP2 FSAR UPDATE ACS TR 636 RCS LI 459 TABLE 6.2-2 (Sheet 2 of 2) Instrumentation Readouts On The Control Board For Operator Monitoring During Recirculation Instruments (continued) System Channel No. RCS LI 460 RCS LI 461 RCS LI 462 Pumps System Pumps SIS Safety Injection SWS Service Water ACS Component Cooling CS Containment Spray RS Recirculation ACS Residual Heat Removal Key: ACS - Auxiliary Coolant System CS - Containment Spray System RCS - Reactor Coolant System RS - Recirculation SIS - Safety Injection System SWS - Service Water System Chapter 6, Page 43 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-3 (Sheet 1 of 3) Quality Standards Of Safety Injection System Components Residual Heat Exchanger A. Tests and inspections

1. Hydrostatic test

2. Radiograph of longitudinal and girth welds (tube side only)

3. Ultrasonic testing of tubing or eddy current tests

4. Dye penetrant test of welds

5. Dye penetrant test of tube to tube sheet welds

6. Gas leak test of tube to tube sheet welds before hydro and expanding of tubes B. Special manufacturing process control

1. Tube to tube sheet weld qualifications procedure

2. Welding and NDT and procedure review

3. Surveillance of supplier quality control and product Component Cooling Heat Exchanger A. Tests and inspections

1. Hydrostatic Test

2. Dye penetrant test of welds B. Special Manufacturing Process Control

1. Welding and NDT and procedure review

2. Surveillance of supplier quality control and product Safety Injection, Recirculation, and Residual Heat Removal Pumps A. Test and inspections 1 Performance test 2 Dye penetrant of pressure retaining parts1

3. Hydrostatic test B. Special manufacturing process control

1. Weld, NDT, and inspection procedures for review

2. Surveillance of suppliers quality control system and product Accumulators A. Tests and inspection

1. Hydrostatic test

2. Radiography of longitudinal and girth welds

3. Dye penetrant/magnetic particle of weld B. Special manufacturing process control

1. Weld, fabrication, NDT, and inspection procedure review

2. Surveillance of suppliers quality control and product Chapter 6, Page 44 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-3 (Sheet 2 of 3) Quality Standards Of Safety Injection System Components Valves A. Tests and inspections

1. 200 psi and 200°F or below (cast or bar stock)

a. Dye penetrant test

b. Hydrostatic test

c. Seat leakage test

2. Above 200 psi and 200°F

a. Forged valves (2-1/2-in. and larger)

(1) Ultrasonic tests of billet prior to forging (2) Dye penetrant 100-percent of accessible areas after forging (3) Hydrostatic test (4) Seat leakage test

b. Case valves (1) Radiograph 100-percent 2 (2) Dye penetrant all accessible areas 2 (3) Hydrostatic test (4) Seat leakage

3. Functional tests required for:

a. Motor operated valves

b. Auxiliary relief valves B. Special manufacturing process control

1. Weld, NDT, performance testing, assembly and inspection procedure review

2. Surveillance of suppliers quality control and product

3. Special weld process procedure qualification (e.g., hard facing)

Piping A. Tests and inspections Class 1501 and below Seamless or welded. If welded 100-percent radiography is required, shop-fabricated and field-fabricated pipe weld joints are inspected as follows: 2501R - 610R: 100-percent radiographic inspection and penetrant examination 301R - 302R: 20-percent random radiographic inspection 151R - 152R: 100-percent liquid penetrant examination B. Special manufacturing process control Surveillance of suppliers quality control and product Refueling Water Storage Tank A. Tests and inspections

1. Vacuum box test of tank bottom seams Chapter 6, Page 45 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-3 (Sheet 3 of 3) Quality Standards Of Safety Injection System Components

2. Hydrostatic test of tank

3. Hydrostatic test of tank heater coil

4. Spot radiography of longitudinal and girth welds B. Special manufacturing process control

1. Weld, fabrication, NDT, and inspection procedure review

2. Surveillance of suppliers quality control and product

3. Material chemical and physical properties certification Notes:

1. Except Internal Recirculation Pump.

2. For valves with radioactive service only.

TABLE 6.2-4 Accumulator Design Parameters Number 4 Type Stainless steel lined/ carbon steel Design pressure, psig 700 Design temperature, °F 300 Operating temperature, °F 100-150 Normal operating pressure, psig Note 2 Total volume, ft3 (each) 1100 Water volume at operating conditions, ft3 (each) Note 2 Minimum boron concentration (as boric acid), ppm 2000 Relief valve setpoint, psig1 700 Notes:

1. The relief valves have soft seats and are designed and tested to ensure exceptional tightness.

2. Minimum and maximum operating pressure and volume are controlled by Technical Specifications.

Chapter 6, Page 46 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-5 DELETED TABLE 6.2-6 Refueling Water Storage Tank Design Parameters Number: 1 Material: Stainless Steel Nominal Capacity, gal. 350,000 Volume Required by Technical Specifications 345,000 (solution), gal. Normal pressure, psig Atmospheric Operating temperature, oF 40-110 Design pressure, psig Atmospheric Design temperature, oF 120 Boron concentration (as boric acid), ppm 2400 (minimum) Type of heating Steam Chapter 6, Page 47 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-7 Pump Design Parameters Safety injection pump Number 3 Design pressure, discharge, psig 1750 Design pressure, suction, psig 200 Design temperature, °F 285 Design flow rate, gpm 400 Maximum flow rate, gpm 650 Design head, ft 2500 Shutoff head, ft 3550 Material Martensitic stainless steel Motor, hp 400 Type Horizontal centrifugal Recirculation pump Number of pumps 2 Type Vertical centrifugal Design pressure, discharge, psig 250 Design temperature, oF 300 Design flow, gpm 3000 Design head, ft 360 Material Austenitic stainless steel Maximum flow rate, gpm 4428 Shutoff head, ft 476 Motor, hp 350 Residual heat removal pump Number of pumps 2 Type Vertical centrifugal Design pressure, discharge, psig 600 Design temperature, oF 400 Design flow, gpm 3000 Design head, ft 350 Material Austenitic stainless steel Maximum flow rate, gpm 5500 Shutoff head, ft 390 Motor, hp 400 Chapter 6, Page 48 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-8 Residual Heat Exchangers Design Parameters Heat Exchangers Number 2 Design heat duty, Btu/hr (Normal) 30.8 x 106 Design UA1, Btu/hr-°F 1.2 x 106 Design cycles (85°F - 350°F) 200 Type Vertical shell and U-tube Normal condition Tube side Shellside Design pressure, psi 600 150 Design flow, lb/hr 1.44 x 106 2.46 x 106 Inlet temperature, °F 135 88.3 Outlet temperature, °F 113.5 100.8 Notes:

1. Total heat transfer coefficient x Area.

Chapter 6, Page 49 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-9 Estimated External Recirculation Loop Leakage3 Items Number Type of Leakage Leakage to Leakage to of Units Control and Unit Atmosphere Tank (cm3/hr) Leakage Rate (cm3/hr) Residual heat 2 Mechanical seal 01 6 removal pumps with leakoff-(low-head safety drop/min injection) High-head safety 3 Same as residual 01 9 injection pumps heat removal Flanges: Gasket-adjusted to zero leakage following any test

                                        - 10 drops/min per flange

a. Pump 15 4501 0

b. Valves - 16 4802 0 Bonnet to Body (larger than 2-in.)

Valves - stem 6 Backseated, 02 6 leakoffs double packing with leakoff 1 cm3/hr-in.stem diameter Misc. small valves 23 Flanged body 692 0 packed stems - 1 drop/min Totals 999 21 Notes:

1. Total estimated leakage from RHR and SI pump mechanical seals and flanges is 450 cc/hr.

2. The total leakage estimated from all sources including valve stem leakage, packing leakoffs, flanged body packed stems and other potential sources (pumps and flanges) of External Recirculation Loop is 999 cc/hr.

3. Actual measured leakage is limited by Technical Specifications. The radiological consequences of external recirculation loop leakage following a design basis accident are presented in Section 14.3.6.6.

Chapter 6, Page 50 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-10 (Sheet 1 of 3) Single Active Failure Analysis - Safety Injection System Component1 Malfunction Comments A. Accumulator (injection phase) Deliver to broken Totally passive system with loop one accumulator per loop. Evaluation based on three accumulators delivering to the core and one spilling from ruptured loop. B. Pump: (injection phase)

1. Safety injection Fails to start Three provided. Evaluation based on operation of two.

2. Residual heat removal Fails to start Two provided. Evaluation based on operation of one.

3. Essential service water Fails to start Three provided. Evaluation based on operation of two.

4. Component cooling2 Fails to start A total of 1 of 3 required during recirculation.

5. Nonessential service water2 Fails to start A total of 1 of 3 required during recirculation.

6. Recirculation2 Fails to start Two provided. One required to operate during recirculation.

7. Auxiliary component cooling Fails to start Two provided. One required pump to operate during recirculation.

Chapter 6, Page 51 of 174 Revision 22, 2010

IP2 FSAR UPDATE TABLE 6.2-10 (Sheet 2 of 3) Single}}