RS-14-128, LaSalle, Units 1 & 2, Updated Final Safety Analysis Report, Revision 20, Chapter 6.0, Engineered Safety Features

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LaSalle, Units 1 & 2, Updated Final Safety Analysis Report, Revision 20, Chapter 6.0, Engineered Safety Features
ML14113A090
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Issue date: 04/11/2014
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RS-14-128
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LSCS-UFSAR 6.0-1 REV. 13 CHAPTER 6.0 - ENGINEERED SAFETY FEATURES The engineered safety features of LaSa lle County Station are those systems whose actions are essential to a safety action required to mitigate the consequences of postulated accidents. The features can be divided into five general groups as follows: containment system s, emergency core cooling systems (ECCS), habitability systems, fission product removal and control systems and other systems. The LSCS engineered safety features, listed by th eir appropriate general grouping, are given below: GROUP SYSTEM Containment Systems

Primary Containment Secondary Containment Containment Heat Removal System Combustible Gas Control System Containment Isolation System Emergency Core Cooling System High-Pressure Core Spray System (HPCS)

Low-Pressure Core Spray System (LPCS)

Low-Pressure Coolant Injection System (LPCI)

Automatic Depressurization System (ADS)

Habitability Systems Control Room HVAC Fission Product Removal and Control Systems

Standby Gas Treatment System Emergency Make-Up Air Filter System LSCS-UFSAR 6.0-2 REV. 13 GROUP SYSTEM Other Systems Main Steamline Isolation Valve Isolated Condenser Leakage Treatment Method

LSCS-UFSAR 6.1-1 REV. 13 6.1 ENGINEERED SAFETY FEATURE MATERIALS The materials utilized in the LSCS engineered safety feature systems have been selected on the basis of an engineering review and evaluation for compatibility with:

a. the normal and accident service conditions of the (engineered safety feature) ESF system, b. the normal and accident environmental conditions associated with the ESF system, c. the maximum expected normal and accident radiation levels to which the ESF will be subjected, and
d. other materials to preclude material interactions that could potentially impair the operation of the ESF systems.

The materials selected for the ESF systems ar e expected to function satisfactorily in their intended service without adverse effects on the service, performance or operation of any ESF.

6.1.1 Metallic Materials In general, all metallic materials used in ESF systems comply with the material specifications of Section II of the ASME Boiler and Pressure Vessel Code.

Pressure-retaining materials of the ESF systems comply with the stringent quality requirements of their applicable quality group classification and ASME B&PV

Code,Section III classification. Adherence to these requirements assures materials of the highest quality for the ESF systems. In those cases where it is not possible to adhere to the ASME material specifications, metallic materials have been selected in compliance with other nationally recognized standards, e.g., ASTM, where practicable, or chosen in compliance with current industry practice.

6.1.1.1 Materials Selection and Fabrication Metallic materials in ESF systems have, in general, been designed for a service life of 40 years, with due consideration of the effects of the service conditions upon the properties of the material, as required by Section III of the ASME B&PV Code, Article NC-2160.

Pressure retaining components of the ECCS have been designed with the following corrosion allowances, in compliance with the general requirement of Section III of the ASME B&PV Code, Article NC-3120:

a. Ferritic Materials LSCS-UFSAR 6.1-2 REV. 18, APRIL 2010 1. water service 0.08 inches
2. steam service 0.120 inches
b. Austenitic Materials 0.0024 inches

For ESF systems other than ECCS, appropriate corrosion allowances, considering the service conditions to which the material will be subjected, have been applied.

The metallic materials of the ESF syst ems have been evaluated for their compatibility with core and containment spray solutions. No radiolytic or pyrolytic decomposition of ESF material will occur during accident conditions, and the integrity of the containment or function of any other ESF will not be effected by the action of core or containment spray solutions.

Material specification for the principal pressure-retaining ferritic, austenitic, and nonferrous metals in each ESF component ar e listed in Table 6.1-1. Materials that would be exposed to the core cooling water and containment sprays in the event of a loss-of-coolant accident are identified in th is table. Sensitization of austenitic stainless steel is prevented by the following actions:

a. Design specifications for austenitic stainless steel components require that the material be cleaned using halide free cleaning solutions and that special care be exercised in the fabrication, shipment, storage, and construc tion to avoid contaminants.
b. Design specifications call for ASME material, which is to be supplied in the solution annealed condition.
c. Design specifications prohibit the use of materials that have been exposed to sensitizing temperatures in the range of 800° F to 1500° F.

Cold-worked austenitic stainless steels wi th yield strengths greater than 90,000 psi are not utilized in ESF systems. Therefore, there are no compatibility problems with core cooling water or the containment sprays.

Metallic reflective thermal insulation is used exclusively inside the primary containment. Premoulded non-hydrophobic Microtherm MPS Insulation enclosed in a 24 gauge stainless steel jacket is installed on the Unit 2 RVWLIS piping, 2NB86A-3/4" and 2NB88A-3/4", and the ma in steam high-flow instrument piping, 2MSC6AD-3/4" inside primary containm ent. Premoulded non-hydrophobic Microtherm MPS insulation enclosed in LSCS-UFSAR 6.1-3 REV. 14, APRIL 2002 24 gauge stainless steel jacket is insta lled on Unit 1 RVWLIS piping 1NB09A-2", 1NB09B-1", 1NB88A-1", 1NB24A-2", and 1NB24B-1", and the main steam high-flow instrument piping, 1MSC6AK-3/4", inside primary containment. The aforementioned Microtherm Insulation is also installed on the Unit 1 main steam high-flow instrument piping, 1MSC6AK-3/4", inside primary containment.

ARMAFLEX insulation is installed on the chilled water system inside primary containment.

Outside containment, calcium silicate or an engineering approved alternative thermal insulation is utilized. Design specifications on the nonmetallic insulation require that it be in accordance with Regulatory Guide 1.36, in order to avoid the possibility of chloride induced stress corrosion cracking in austenitic stainless steel in contact with the insulation.

To avoid hot cracking (fissuring) during weld fabrication and assembly of austenitic stainless steel components of the ESF, the design specifications require the following:

a. Maximum delta ferrite content for wrought and duplex cast components is 5% - 15%.
b. Chemical analyses are performed on undiluted weld deposits, or alternately, on the wire, consumable insert, etc., to verify the delta ferrite content.
c. Delta ferrite content in weld metal is determined using magnetic measurement devices.
d. Maximum interpass temperatur e shall not exceed 350°F during welding. e. Test results as discussed above are included in the qualification test report.
f. Weld materials meet the re quirements of Section III.
g. Production welds are examined to verify that the specified delta-ferrite levels are met.
h. Welds not meeting these leve ls are unacceptable and must be removed.

LSCS-UFSAR 6.1-4 REV. 14, APRIL 2002 6.1.1.2 Composition, Compatibility and Stability of Containment and Core Spray Coolants The core sprays have two possible sources of coolant. The HPCS system is supplied from either the cycled condensate storage tank or the suppression pool. The normal source of water for HPCS is the suppression pool. The capability remains for the HPCS system to draw a suction on the cycl ed condensate tank because the piping to the tank is installed, but isolated by a b lind flange. Establishment of this flowpath is under administrative control. The LPCS and LPCI are supplied from the suppression pool only. Water quality in both of these sources is maintained at a high level of purity with the possible exception of potentially high soluble-iron metallic impurities. Additional discussion of the water qualities are given in Subsections 6.1.3, 9.2.7, and 9.2.11. Limited corrosion inhibitors or other additives (such as zinc and noble metals) are present in either source.

The containment spray utilizes the suppression pool as its source of supply. No radiolytic or pyrolytic decomposition of ESF materials are induced by the containment sprays. The containment sprays should not be a source of stress-corrosion cracking in austenitic stainless steel during a LOCA.

6.1.2 Organic Materials Table 6.1-2 lists all the organic compounds that exist within the containment in significant amounts. All these materials in ESF components have been evaluated with regard to the expected service conditions, and have been found to have no adverse effects on service, performance, or operation.

The dry well liner and coated exposed metal surfaces inside containment are prime coated with an inorganic zinc compound that has been fully qualified in accordance with ANSI standards N101.2, N101.4, an d N512 , with the exception of a small quantity (44 gallons) used on pipe hangers and snubber attachments and recirculating pump motors. Uncoated metal surfaces shall be evaluated for acceptability. No radiolytic or pyrolytic decomposition or interaction with other ESF materials will occur.

6.1.3 Postaccident Chemistry The post-accident chemical environment inside the primary containment will consist of water from the suppression pool and the cycled condensate storage tank, i.e. water sources for the high pressure core spray, low pressure core spray, low pressure core injection, reactor core isolation cooling and containment spray. The

suppression pool may contain trace amounts of corrosion inhibiting chemicals such as hydrogen, zinc and noble metals. Additionally, portions of the Reactor Building Closed Cooling Water (RBCCW) system and the Primary Containment Chilled Water (PCCW) system are inside the containment. Both systems contain limited LSCS-UFSAR 6.1-5 REV. 14, APRIL 2002 amounts of corrosion inhibitors, and have portions of their piping inside containment classified as Seismic Category 2. During a Design Basis Accident (DBA) either or both of these systems can fail and release the corrosion inhibitors to the suppression pool before isolation. Due to the limited quantity (trace amounts) of these chemicals in the secondary systems and the dilution factor as a result of a DBA, the water will be approximately neutral (pH = 7), and there will be no adverse affect to equipment, coatings or other materials during ECCS or RCIC operation.

LSCS-UFSAR TABLE 6.1-1 (SHEET 1 OF 5) TABLE 6.1-1 REV. 13 PRINCIPAL PRESSURE-RETAINING MATERIAL FOR ESF COMPONENTS

I. Containment Systems A. Primary Containment

1. Containment Walls 4500 psi Concrete
  • 2. Drywell Liner SA-516, Grade 60
  • 3. Suppression Chamber Liner SA-240, Type 304
  • 4. Drywell Head SA-516, Grade 70
  • 5 Penetrations a. Drywell SA-333, Grade 1 or 6 (Seamless) b. Suppression Chamber SA-312, Grade TP 304 (Seamless) *6. Equipment Hatch SA-516, Grade 70
  • 7. Personnel Access Hatch
a. Drywell SA-516, Grade 70 b. Suppression Chamber SA-240, Type 304 *8. Suppression Vent Downcomers SA-240, Type 304 *9. Vacuum Relief Piping a. Drywell to Suppression Chamber Penetration SA-106, Grade B b. Suppression Chamber Penetration SA-312, Grade TP 304 (Seamless) 10. Vacuum Relief Valves SA-105
  • Indicates that material may be subjected to containment spray or core cooling water in the event of a loss-of-coolant accident.

LSCS-UFSAR TABLE 6.1-1 (SHEET 2 OF 5) TABLE 6.1-1 REV. 13

  • 11. Pressure Retaining Bolts a. Drywell SA-320, Grade L43 SA-193, Grade B7 SA-194, Grade 7 b. Suppression Chamber SA-193, Class 2, Grade B8C, Type 347 SA-194, Class 2, Grade 83, Type 347 B. Secondary Containment
1. Ducts A-526
2. Dampers A-285, Grade B A-181, Grade 1 C. Containment Heat Removal System 1. RHR Pumps A-516, Grade 70 2. RHR Heat Exchanger
a. Shell Side SA-516, Grade 70
b. Tube Side SA-249, Grade TP 304L
  • 3. Piping SA-106, Grade B
  • 4. Valves SA-216, Grade WCB or SA-105 *5. Pressure-Retaining Bolting SA-193, Grade B7 *6. Welding Material SFA-5.18E70S-3(F-6, A-1)

D. Containment Isolation System

  • 1. Piping SA-106, Grade B or SA-312, Grade TP 304 *2. Valves SA-216, Grade WCB or SA-105 or SA-182, Grade 316L or Grade F316 or SA-351, Grade C8FM or SA-351 Grade CF3
  • Indicates that material may be subjected to containment spray or core cooling water in the event of a loss-of-coolant accident.

LSCS-UFSAR TABLE 6.1-1 (SHEET 3 OF 5) TABLE 6.1-1 REV. 13

  • 3. Pressure-Retaining Bolting SA-193, Grade B7 *4. Welding Material SFA-5.18E70S-3 (F-6, A-1)

E. Combustible Gas Control System 1. Piping SA-106, Grade B

2. Valves SA-216, Grade WCB
3. Recombiner SA-358, Grade 304
4. Blower 5. Pressure-Retaining Bolting SA-193, Grade B7 6. Welding Material SFA-5.18E70S-3 (F-6, A-1)

II. Emergency Core Cooling System A. High-Pressure Core Spray 1. Pump A-516, Grade 70

2. Piping
  • a. Inside Reactor Building SA-106, Grade B
b. Outside Reactor Building SA-409, Grade TP 304
  • 3. Valves SA-216, Grade WCB or SA-105
  • 4. Pressure-Retaining Bolting SA-193, Grade B7
  • 5. Welding Materials SFA-5.18E70S-3 (F-6, A-1)

B. Low-Pressure Core Spray 1. Pump A-516, Grade 70

  • 2. Piping SA-106, Grade B *3. Valves SA-216, Grade WCB or SA-105
  • Indicates that material may be subjected to containment spray or core cooling water in the event of a loss-of-coolant accident LSCS-UFSAR TABLE 6.1-1 (SHEET 4 OF 5) TABLE 6.1-1 REV. 13
  • 4. Pressure-Retaining Bolting SA-193, Grade B7 *5. Welding Materials SFA-5.18E70S-3 (F-6, A-1)

A. Low-Pressure Coolant Injection 1. RHR Pump A-516, Grade 70 *2. Piping SA-106, Grade B

  • 3. Valves SA-216, Grade WCB or SA-105
  • 4. Pressure-Retaining Bolting SA-193, Grade B7
a. Inlet SA-155, Grade KCF70
b. Outlet SA-106, Grade B
  • 2. Valves

III. Habitability System A. Blowers A-283, A-242 B. Dampers A-285, Grade B A-181, Grade 1 C. Ducts A-526 D. Housing A-36

IV. Fission Product Removal and Control System A. Standby Gas Treatment System 1. a. Piping (Downstream of Filter Unit)

SA-106, Grade B b. Piping (Upstream of Filter Unit)

A-106, Grade B 2. Housing A-36

  • Indicates that material may be subjected to containment spray or core cooling water in the event of a loss-of-coolant accident.

LSCS-UFSAR TABLE 6.1-1 (SHEET 5 OF 5) TABLE 6.1-1 REV. 13

3. Valves SA-216, Grade WCB or SA-105, or SA-516, Grade 7 4. Dampers A-285, Grade B A-181, Grade 1
5. Blowers A-283, A-242
6. Pressure-Retaining Bolting
a. Pressure-Retaining Bolting (Downstream of Filter Unit)

SA-193, Grade B7 b. Pressure-Retaining Bolting (Upstream of Filter Unit)

A-193, Grade B7 7. Welding Materials SFA-5.18E70S-3 (F-6,A-1) B. Emergency Air Filter System

1. Ducts A-526
2. Dampers A-285, Grade B A-181, Grade 1
3. Housing A-36
4. Blower A-283, A-242 V. Other Systems A. Main Steamline Isolation Valve Leakage Control System (Deleted)
  • Indicates that material may be subjected to containment spray or core cooling water in the event of a loss-of-coolant accident LSCS-UFSAR TABLE 6.1-2 (SHEET 1 OF 2) TABLE 6.1-2 REV. 18, APRIL 2010 ORGANIC MATERIALS WITHIN THE PRIMARY CONTAINMENT MATERIAL USE QUANTITY Acrylomitrile Butadiene/PVC Foam Rubber ARMAFLEX Insulation on the Chilled Water Piping Throughout Drywell Chlorosulfinated Polyethylene (Hypalon)

Low Voltage Electrical Power Cable Jacketing and Insulation Material Throughout Drywell Etylene Propylene Rubber (EPR) Low Voltage Electrical Power Cable Jacketing and Insulation Material Throughout Drywell High Temperature Ethylene Propylene Medium Voltage Electrical Power Cable Jacketing and

Insulation Material Throughout Drywell Hypalon/Hypalon Instrumentation Cable Insulation/Jacketing Material Throughout Drywell EPR/Hypalon Instrumentation Cable Insulation/Jacketing Material Throughout Drywell Cross-Linked Polyolefin/Alkaneimide Polymer Instrumentation Coaxial and Triaxial Insulation/

Jacketing Material Throughout Drywell Modified Phenolic Coating for Exposed Carbon Steel Surfaces 16 ft 3 Modified Phenolic Surfacer Coating for Exposed Concrete Surfaces 17 ft 3 Modified Phenolic Finish Coating for Exposed Concrete Surfaces 5 ft 3 LSCS-UFSAR TABLE 6.1-2 (SHEET 2 OF 2) TABLE 6.1-2 REV. 18, APRIL 2010 MATERIAL USE QUANTITY Alkyd Primer and Finish Pipe hangers and Snubber Attachments

and GE Recirculating Pump 44 gal. Lube Oil Reactor Recirculation Pump Motor (2 motors/unit) 145 gal per unit Silicone Fluid (SF 1147, GE) MSIV Hydraulic Fluid (4 valves within containment) 1 1/2 gal. per valve Non-separating high temperature grease Drywell cooling area coolers < 1 gal.

Fyrquel 220/or Fyrquel EHC (stauffer)

Recirculation Control Valve Hydraulic Fluid (2 valves) 118 gal. per valve Silicone Fluid Lisega Hydraulic Snubbers

< 1 1/2 gal. per snubber Fiberglass Reinforced Silicone Fabric 1 (2) RF01 and 1 (2) RE02 Sump Cover Mat 400 ft 2 per unit Silicone Sealant 1 (2) RF01 and 1 (2) RE02 Sump Cover Mat < 1 gal. per unit

LSCS-UFSAR 6.2-1 REV. 13 6.2 CONTAINMENT SYSTEMS 6.2.1 Containment Functional Design This section establishes the design bases for the primary containment structure, describes the major design features of the structure, and presents an evaluation of the capacity of the containment to perform its required safety function during all normal and postulated accident conditions described in this UFSAR.

6.2.1.1 Containment Structure 6.2.1.1.1 Design Bases

The primary containment structure has been designed to meet the following safety design bases:

a. Containment Vessel Design
1. The containment structure has the capability to withstand the peak transient pressures and temperatures that could occur due to the postulated design-basis accident (DBA).
2. The containment has the capability to maintain its functional integrity indefinitely after the postulated DBA.
3. The containment structure also withstands the peak environmental transient pressures and temperatures associated

with the postulated small line break inside the drywell.

4. The containment structure has also been designed to withstand the coincident fluid jet forces associated with the flow from the postulated rupture of any pipe within the containment.
5. The containment has also been designed to withstand the hydrodynamic forces associated with a DBA and safety-relief valve discharge, as described in the LaSalle Design Assessment Report. Design loading combinations are also described in the design assessment report: Design pressure and temperature, and the major containment design parameters are listed in Table 6.2-1.
b. Containment Subcompartment Design The internal structures of the containment have been designed to accommodate the peak transient pressures and temperatures LSCS-UFSAR 6.2-2 REV. 13 associated with the postulated design-basis accident (DBA). The effects of subcompartment pressurization for the postulated pipe ruptures have been evaluated. Subcompartment pressurization is more fully discussed in Subsection 6.2.1.2.
c. Containment Internals Design The drywell floor has been designed to withstand a downward acting differential pressure of 25 psig in combination with the normal operating loads and safe shutdown earthquake (SSE). The drywell floor has also been designed to accommodate an upward acting deck differential pressure of 5 psig, in order to account for the wetwell pressure increase that could occur after a loss-of-coolant accident (LOCA). d. Containment Design for Mass and Energy Release
1. The maximum postulated release of mass and energy to the containment is based upon the instantaneous circumferential rupture of a 24- inch reactor recirculation line or a 26-inch main

steamline.

2. The effects of metal-water reactions and other chemical reactions following the DBA can be accommodated in the containment design.
e. Energy Removal Features The RHR system, through the containment cooling mode, is utilized to remove energy from the containment following a LOCA by circulating the suppression pool water through a residual heat removal (RHR) heat exchanger for cooling, and returning the water to the pool through the low-pressure core injection (LPCI) in the reactor pressure vessel (RPV) or the suppression chamber spray header. The containment spray mode of the RHR system can also be utilized to condense steam and reduce the temperature in the drywell following a LOCA. A more detailed description is available in Subsection 6.2.2. The RHR containment cooling mode energy removal capability is not affected by a single failure in the system, sinc e a completely redundant loop is available to perform this functi on. Two redundant loops of the containment spray system are also provided.

LSCS-UFSAR 6.2-3 REV. 13 f. Pressure Reduction Features The containment vent system dire cts the flow from postulated pipe ruptures to the pressure suppression pool, and distributes such flow uniformly throughout the pool, to condense the steam portion of the flow rapidly, and to limit the pressure differentials between the drywell and wetwell during various postaccident cooling modes.

g. Hydrostatic Loading Design The containment design permits filling the containment system drywell with water to a level 1 foot below the refueling floor to permit removal of fuel assemblies during postaccident recovery.
h. Impact Loading Design The containment system is protected against missiles from internal or external sources and excessive motion of pipes that could directly or indirectly jeopardize containment integrity.
i. Containment Leakage Design The containment limits leakage during and following the postulated DBA to values less than leakage rates that would result in offsite doses greater than 10 CFR 100.
j. Containment Leakage Testability It is possible to conduct periodical leakage tests as may be appropriate to confirm the integrity of the containment at calculated peak pressure resulting from the postulated DBA.

For the purposes of the containment structure design, the design-basis accident (DBA) is defined as a mechanical failure of the reactor primary system equivalent to the circumferential rupture of one of the recirculation lines. During the DBA, the long-term peak suppression pool temperature shall not exceed the design temperature.

6.2.1.1.2 Design Features The primary containment is a concrete stru cture with the exception of the drywell head and access penetrations, which are fabricated from steel. The major components are shown in Figure 3.8-1. The concrete is designed to resist all loads associated with the design-basis accident.

LSCS-UFSAR 6.2-4 REV. 15, APRIL 2004 The primary containment walls have a steel liner, which acts as a low leakage barrier for release of fission products.

The walls of the primary containment are posttensioned concrete; the base mat is conventional reinforced concrete. The dividing floor between the drywell and suppression chamber is conventional reinforced concrete and is supported on a cylindrical base at its center, on a seri es of concrete co lumns and from the containment wall at the periphery of the slab.

The drywell floor is rigidly connected to the primary containment wall. A full moment and shear connection is provided by dowels and shear lugs welded to the reinforced liner plate as shown in Figure 3.

8-4. The thermal expansion is accounted for in the containment design; the resulting forces and moments are accommodated

within the allowable stress limits.

The primary containment walls support the reactor building floor loads and, in addition, also serve as the biological shield. A detailed discussion of the structural design bases is given in Chapter 3.0. The codes, standards, and guides applied in the design of the containment structure an d internal structures are identified in Chapter 3.0.

The walls of the primary containment st ructure are posttens ioned, using the BBRV system of posttensioning utilizing parallel lay, unbonded type tendons. The tendons are fabricated from 90 one-quarter inch diameter, cold drawn, stress relieved, prestressing grade wire. Each tendon is encased in a conduit. The walls are prestressed both vertically and horizontally for floor elevations below 820 feet. The horizontal tendons are placed in a 240 system using three buttresses as anchorages with the tendons staggered so that two-thirds of the tendons at each buttress terminate at that buttress. For floor elevations above 820 feet, the horizontal tendons are placed in a 360 system using two buttresses as anchorages. Access to the tendon anchorages is maintained to allow for periodic inspection. For a typical layout of hoop tendons, see Fi gure 3.8-11. A typical layout of the vertical tendons is illustrated in Figure 3.8-11.

All liner joints have full penetration welds. The field welds have leaktightness testing capability by having a small steel channel section welded over each liner weld. Fittings are provided in the channel for leak testing of the liner welds under pressure. The actual containment leakag e boundary during normal operation and accident conditions consists of the liner and liner joint butt welds when the leak test channel is vented to the containment atmosphere and the combined containment liner, liner joint butt welds, containment liner leak test channels, channel fillet welds and the leak test connections when the leak test channel test connection plugs are installed. The liner anchor age system considers the effects of temperature, negative pressure, prestressing, and stress transfer around penetrations.

LSCS-UFSAR 6.2-5 REV. 13 Drywell The drywell is a steel-lined posttensioned concrete vessel in the shape of a truncated cone having a base diameter of approximately 83 feet and a top diameter of 32 feet.

The floor of the drywell serves both as a pressure barrier between the drywell and suppression chamber and as the support structure for the reactor pedestal and downcomers. The drywell head is bolted at a steel ring girder attached to the top of the concrete containment wall and is sealed with a double seal. The double seal on the head flange provides a plenum for determining the leaktightness of the bolted connection. The base of the ring serves as the top anchorage for the vertical prestressing tendons and the top of the ring serves as anchorage for the drywell head. The drywell houses the reactor and its associated auxiliary systems. The primary function of the drywell is to contain the effects of a design-basis recirculation line break and direct the steam released from a pipe break into the suppression chamber pool. The drywell is designed to resist the forces of an internal design pressure of 45 psig in combination with thermal, seismic, and other forces as outlined in Chapter 3.0.

The drywell is provided with a 12-foot diameter equipment hatch for removal of equipment for maintenance and an air lock for entry of personnel into the drywell.

Under normal plant operations, the equipm ent hatch is kept sealed and is opened only when the plant is shut down for refueling and/or maintenance.

The equipment hatch is covered with a steel dished head bolted to the hatch opening frame which is welded to the steel liner. A double seal is utilized to ensure leaktightness when the hatch is subjected to either an internal or external pressure. The space between the double seal serves as a plenum for leak testing the hatch seal. The personnel air lock is a cylindrical intake welded to the steel liner. The double doors are interlocked to maintain containment integrity during operation.

All welds that make up the vapor barrier have test channels to permit leak testing of the welds: When the leak test channel test connections are plugged, the leak test channel is part of the vapor barrier.

The primary containment ventilation system, as described in Subsection 9.4.9, is provided to maintain drywell temperatures at approximately 135 F during normal plant operation.

LSCS-UFSAR 6.2-6 REV. 14, APRIL 2002 The primary containment vent and purge syst em, as described in Subsection 9.4.10, is designed to purge potentially radioactive gases from the drywell and suppression chamber prior to and during perso nnel access to the containment.

Containment penetration cooling is provided on high temperature penetrations through the primary containment wall by th e reactor building closed cooling water system. The penetrations served by this system and the design basis for the cooling loads are described in Subsection 9.2.3.

Pressure Suppression Chamber and Vent System The primary function of the suppression ch amber is to provide a reservoir of water capable of condensing the steam flow from the drywell and collecting the noncondensable gases in the suppression chamber air space. The suppression chamber is a stainless steel-lined posttensioned concrete vessel in the shape of a cylinder, having an inside diameter of 86 feet 8 inches. The foundation mat serves as the base of the suppression chamber. The suppression chamber is designed for the same internal pressure as the drywell in combination with the thermal, seismic, and other forces. The liner design and te sting are the same as covered previously within this subsection (6.2.1.1.1.2).

The entire suppression chamber is lined with stainless steel.

The drywell floor support columns are also provided with a stainless steel liner on the outside surface. Two 36-inch diameter openings are provided for access into the suppression chamber for inspection. Under normal plan t operation, these access openings are kept sealed. They are opened only when the plant is shut down for refueling and/or maintenance. The access openings are located in the cylindrical walls of the chamber 14 feet 2 inches above the suppr ession pool water level. The access openings are closed using a bolted steel ha tch cover. The hatch cover is designed with a double seal and test plenum to ensure leaktightness.

The suppression chamber vent system consists of 98 downcomer pipes open to the drywell and submerged 12 feet 4 inches below the low water level of the suppression pool, providing a flow path for uncondensed steam into the water. Each downcomer has a 23.5-inch internal diameter. The downcomers project 6 inches above the drywell floor to prevent flooding from a br oken line. Each vent pipe opening is shielded by a 1-inch thick steel deflecto r plate to prevent overloading any single vent pipe by direct flow from a pipe break to that particular vent. The principal parameters for design of the primary containment, suppression pool, reactor

building and the vent downcomers are listed in Table 6.2-1.

LSCS-UFSAR 6.2-7 REV. 14, APRIL 2002 Vacuum Relief System Vacuum relief valves are provided between the drywell and suppression chamber to prevent exceeding the drywell floor negative design pressure and backflooding of the suppression pool water into the drywell.

In the absence of vacuum relief valves , drywell flooding could occur following isolation of a blowdown in the drywell. Condensation of blowdown steam on the drywell walls and structures could result in a negative pressure differential between the drywell and suppression chamber.

The vacuum relief valves are designed to equalize the pressure between the drywell and wetwell air space regions so that the reverse pressure differential across the

diaphragm floor will not exceed the design value of five pounds per square inch.

The vacuum relief valves (four assemblie s) are outside the primary containment and form an extension of the primary co ntainment boundary. The vacuum relief valves are mounted in special piping which connects the drywell and suppression chamber, and are evenly distributed around the suppression chamber air volume to prevent any possibility of localized pres sure gradients from occurring due to geometry. In each vacuum breaker assembly, two local manual butterfly valves, one on each side of the vacuum breaker, are provided as system isolation valves should failure of the vacuum breaker occur.

The vacuum relief valves are instrumented with redundant position indication and are indicated in the main control room. The valves are provided with the capability for local manual testing. The position indication requirements for the vacuum relief valves are located in the Administrative Technical Requirements. (References 21, 22, and 23)

This design provides adequate assurance of limiting the differential pressure between the drywell and suppression cham ber and assures proper valve operation and testing during normal plant operation.

No vacuum relief valves are provided between the drywell and the reactor building atmosphere. The concrete containment structure has the ability to accommodate subatmospheric pressures of ap proximately 5 psi absolute.

6.2.1.1.3 Design Evaluation The key design parameters for the pressure suppression containment being

provided for the LaSalle County Statio n (LSCS) are listed in Table 6.2-1.

These design parameters are not determined from a single accident event but from an envelope of accident conditions. As a result, there is no single design-basis accident (DBA) for this containment system.

LSCS-UFSAR 6.2-8 REV. 15, APRIL 2004 The containment system was analyzed orig inally at 3434 MWt reactor power. Since then, the containment system evaluation was performed for a reactor power of 3559 MWt by analyzing the limiting events at this power level. The results for 3559 MWt power are included in this section, while keeping most of the original analysis results for 3434 MWt power as a referenc e analysis for hist orical purposes.

A maximum drywell and suppression chamber pressure of 39.6 psig and 30.6 psig, respectively is predicted near the end of the blowdown phase of a loss-of-coolant accident (LOCA) transient. Approximately the same peak pressure occurs for either the break of a recirculation line or a main steamline. Both accidents are evaluated at 3434 MWt.

For 3559 MWt reactor power, the maximum cont ainment pressure is predicted to be 39.9 psig in the drywell and 27.9 psig in the suppression chamber for the recirculation line break. The main steamline break was not reevaluated for the uprated power level.

The most severe drywell temperature condition is predicted for a small primary system rupture above the reactor water level that results in the blowdown of reactor

steam to the drywell. Based upon the thermodynamic conditions this would produce high temperature steam in the drywell.

In order to demonstrate that breaks smaller than the rupture of the largest primary system pipe will not exceed the containment design parameters, the blowdown phase of an intermediate size break is evaluated. Containment design conditions are not exceeded for this or the other break sizes.

All of the analyses assume that the primary system and containment are at the maximum normal operating conditions. Re ferences are provided that describe relevant experimental verification of the analytical models used to evaluate the containment response.

Table 6.2-1 provides a listing of the key design parameters of the LSCS primary containment system including the design characteristics of the drywell, suppression chamber and the pressure suppression vent system.

Table 6.2-2 provides the performance parameters of the related engineered safety feature systems which supplement the de sign conditions of Table 6.2-1 for containment cooling purposes during po staccident operation. Performance parameters given include those applicable to full capacity operation and to those reduced capacities employed for containment analyses.

LSCS-UFSAR 6.2-8a REV. 14, APRIL 2002 6.2.1.1.3.1 Accident Response Analysis The containment functional evaluation performed at 3434 MWt is based upon the consideration of several postulated accident conditions resulting in release of reactor coolant to the containment. These accidents include:

a. an instantaneous guillotine rupture of a recirculation line, b. an instantaneous guillotine rupture of a main steam-line, c. an intermediate size liquid line rupture, and
d. a small size steamline rupture.

Energy release from these accidents is reported in Subsection 6.2.1.3.

LSCS-UFSAR 6.2-9 REV. 19, APRIL 2012 The accident response analysis is based on the GE calculations. This is determined based on the containment response being dependent on the amount of energy in the system, the containment design, and the failure modes that allow the pressurization to occur rather than the fuel type. The am ount of energy in the system is based on initial conditions and the assumed blowdown. As the blowdown assumed for the

containment response analysis as show n in Tables 6.2-18 and 6.2-19 bound the blowdown predicted by the SPC LOCA method ology and results, less energy would be released to the containment using the SPC blowdown.

For 3559 MWt reactor power, the limiting even t, an instantaneous guillotine rupture of a recirculation line, was analyzed to perform the containment functional evaluation. The analysis at 3559 MWt was performed in accordance with the Generic Guidelines for General Electric Boiling Water Reactor Power Uprate, NEDC-31897P-A (Reference 24). This analysis employed essentially the same methodology, while taking a more detailed modeling approach for the reactor vessel blowdown evaluation. The analysis re sults for 3559 MWt reactor powe r are included in Section 6.2.1.1.3.1.1 under the heading "Evaluation at 3559 MWt Reactor Power," after a description of the original 3434 MWt analysis which is kept as a reference analysis for historical purposes.

The current licensing basis analysis (Refer ence 31), performed at 3559 MWt (102% of 3489 MWt) bounds the MUR operating condit ions at 3546 MWt (Reference 35). The AREVA MUR containment evaluation (Reference 34) concludes that decay heat of AREVA fuels (ANTRIUM-10 and ATRIUM 10XM) is bounded by the GE fuels used for the LaSalle containment analysis. Therefore, this analysis is applicable for operation with ATRIUM-10 fuel and the ATRIUM 10XM LTAs at MUR conditions.

6.2.1.1.3.1.1 Recirculation Line Rupture The instantaneous guillotine rupture of a main recirculation line results in the maximum flow rate of primary system fluid and energy into the drywell as illustrated in Figure 6.2-1 by the diagram showing th e location of a recirculation line break.

Immediately following the rupture, the flow out of both sides of the break will be limited to the maximum allowed by critical fl ow considerations. Figure 6.2-1 shows a schematic view of the flow paths to the break. Flow in the suction side of the recirculation pump will correspond to critical flow in the 2.565 square foot pipe cross section. Flow in the discharge side of the recirculation pump will correspond to critical flow at the ten jet pump nozzles asso ciated with the broken loop, providing an effective break area of 0.468 ft

2. In addition, there is a 4- inch cleanup line crosstie that will add 0.080 ft 2 to the critical flow area, yielding a total of 3.113 ft
2.

LSCS-UFSAR 6.2-9a REV. 19, APRIL 2012 Assumptions for Reactor Blowdown The response of the reactor coolant syst em during the blowdown period of the accident is analyzed using the following assumptions:

a. At the time the recirculation pipe breaks, the reactor is operating at the most severe condition that maximizes the parameter of interest; that is, primary containment pressure.
b. The recirculation line is considered to be severed instantly. This results in the most rapid coolant loss and depressurization, with coolant being discharged from both ends of the break.
c. The reactor is shut down at the time of accident initiation because of void formation in the core region. Scram also occurs in less than 1 LSCS-UFSAR 6.2-10 REV. 13 second from receipt of the high drywell pressure signal. The difference between shutdown at time zero and 1 second is negligible.
d. The vessel depressurization flow rates are calculated using Moody's critical flow model (Reference 1) assuming "liquid only" outflow, since this assumption maximizes the energy release to the containment:

"Liquid only" outflow requires that all vapor formed in the RPV by bulk flashing rises to the surface rather than being entrained in the existing flow. Some of the vapor would be entrained and would significantly reduce the RPV discharge flow rates. Moody's critical flow model, which assumes annular, isentropic flow, thermodynamic flow, thermodynamic phase equilibrium, and maximized slip ratio, accurately predicts vessel outflows through small diameter orific es. However, actual rates through larger flow areas are less than the model indicates because of the effects of a near homogeneous two- phase flow pattern and phase nonequilibrium. This effect is in addition to the reduction caused by vapor entrainment, discussed previously.

e. The core decay heat and the sensible heat released in cooling the fuel to 545 F are included in the reactor pressure vessel depressurization calculation: The rate of ener gy release is calculated using a conservatively high heat transfer coefficient throughout the depressurization period. By maximizing the assumed energy release rate, the RPV is maintained at nearly rated pressure for approximately 20 seconds. The high RPV pressure increases the calculated blowdown flowrates; this is conservative for containment analysis purposes. With the RPV fluid temperature remaining near 545 F, however, the calculated release of sensible energy stored below 545 F is negligible during the first 20 seconds. The sens ible energy is released later, but does not affect the peak drywell pressure. The small effect of sensible energy release on the long-term suppression pool temperature is included.
f. The main steam isolation valves are assumed to start closing at 0.5 seconds after the accident. They are assumed to be fully closed in the shortest possible time of 3 seco nds following closure initiation.

Actually, the closure signal for th e main steam isolation valves is expected to occur from low water leve l, so these valves may not receive a signal to close for more than 4 seconds, and the closing time could be as long as 5 seconds. By assuming rapid closure of these valves, the RPV is maintained at a high pressure, whic h maximizes the discharge of high energy steam and water into the primary containment: In addition, the rapid closure of the main steam isolation valves cuts off motive power to the steam-driven feedwater pumps.

LSCS-UFSAR 6.2-11 REV. 13 g. Reactor feedwater flow is assumed to stop instantaneously at time zero.

Since cooler feedwater flow tends to depressurize the RPV, thereby reducing the discharge of steam and water into the primary containment, this assumption is considered conservative and consistent with that of assumption f.

With respect to suppression pool temperature, this assumption has been supplemented with an additional evaluation. The purpose being to evaluate the suppression pool long term temperature response. For this evaluation, the feedwater is assumed to have been injected into the suppression pool, by the end of the recirculation piping break blowdown phase (at 600 seconds), in order to assess long term peak pool temperature. See paragraph entitled "Evaluation of Post-LOCA

Feedwater Injection" in this section.

h. A complete loss of offsite power occurs simultaneously with the pipe break. This condition results in th e loss of power conversion system equipment and also requires that all vital systems for long-term cooling be supported by onsite power supplies.

Assumptions for Containment Pressurization The pressure response of the containment during the blowdown period of the accident is analyzed using the following assumptions:

a. Thermodynamic equilibrium exists in the drywell and suppression chamber. Since nearly complete mi xing is achieved, the analysis assumes complete mixing, which is in the conservative direction.
b. The fluid flowing through the drywell-to-suppression chamber vents is formed from a homogeneous mixture of the fluid in the drywell. The use of this assumption results in complete liquid carry-over into the drywell vents. c. The fluid flow in the drywell-to-suppression chamber vents is compressible except for the liquid phase.
d. No heat loss from the gases inside the primary containment is assumed.

This adds extra conservatism to the analysis; that is, the analysis will tend to predict higher containment pressures than would actually result.

Assumptions for Long-Term Cooling Following the blowdown period, the emergency core cooling systems (ECCS) discussed in Section 6.3 provide water for core flooding and long-term decay heat LSCS-UFSAR 6.2-12 REV. 13 removal. The containment pressure and temperature response during this period are analyzed using the following assumptions:

a. The LPCI pumps are used to flood the core prior to 600 seconds after the accident. The high-pressure core sp ray (HPCS) is assumed available for the entire accident.
b. After 600 seconds, the LPCI pump flow may be diverted from the RPV to the containment spray. This is a manual operation. Actually, the containment spray need not be activated at all to keep the containment pressure below the containment design pressure. Prior to activation of the containment cooling mode (arbit rarily assumed at 600 seconds after the accident), all of the LPCI pump flow will be used only to flood the core. c. The effect of decay energy, stor ed energy, and energy from the metal-water reaction on the suppression pool temperature are considered.
d. During the long-term containment response (after depressurization of the reactor vessel is complete) the suppression pool is assumed to be the

only heat sink in the containment system.

e. After approximately 600 seconds, the RHR heat exchangers are activated to remove energy from the containment via recirculation cooling from the suppression pool wi th the RHR service water systems.
f. The performance of the ECCS equipment during the long-term cooling period is evaluated for each of the following three cases of interest:

Case A - Offsite Power Available All ECCS equipment and containment spray operating.

Case B - Loss of Offsite Power Minimum diesel power available for ECCS and containment spray.

Case C - Same as Case B (except no containment spray) Initial Conditions for Accident Analyses

Table 6.2-3 provides the initial reactor coolant system and containment conditions used in all the accident response evaluations. The tabulation includes parameters for the reactor, the drywell, the suppressi on chamber and the vent system. A supplementary safety evaluation has also been performed, as discussed in LSCS-UFSAR 6.2-13 REV. 13 Section 6.2.1.8, to evaluate an increase in the initial suppression pool temperature value to 105 F. Table 6.2-4 provides the initial conditio ns and numerical values assumed for the recirculation line break accident as well as the sources of energy considered prior to the postulated pipe rupture. The assumed conditions for the reactor blowdown are also provided.

The mass and energy release sources and rates for the containment response analyses are given in Subsection 6.2.1.3. Short-Term Accident Response The calculated containment pressure and temperature responses for the recirculation line break are shown in Figures 6.2-2 and 6.2-3 respectively. The calculated peak drywell pressure is 39.6 psig, which is 12% below the containment design pressure of 45 psig. The suppression chamber is pressurized by the carryover of noncondensables from the drywell and by heatup of the suppression pool. As the vapor formed in the drywell is condensed in the suppression pool, the temperature of the suppression chamber water approaches 150 F and the suppression chamber pressure stabilizes at approximately 30 psig. The drywell pressure stabilizes at a slightly higher pressure, the difference being equal to the downcomer submergence. During the RPV depressurization phase, most of the noncondensable gases in the drywell initially are forced into the suppression chamber. However, following the depressurization the noncondensables will redistribute between the drywell and suppression chamber via the vacuum breaker system. This redistribution takes place as pressure is decreased by the steam condensation process occurring in the drywell.

The LPCI and LPCS systems supply sufficient core cooling water to control core heatup and limit metal-water reaction to less than 0.2%. After the RPV is flooded to the height of the jet pump nozzles, th e excess flow discharges through the recirculation line break into the drywell. This flow of water (steam flow is negligible) transports the core decay heat out of the RPV, through the broken recirculation line, in the form of hot water which flows into the suppression chamber via the drywell to suppression chamber vent system. This flow, in addition to heat losses to the drywell walls, provides a heat sink for the drywell atmosphere, LSCS-UFSAR 6.2-14 REV. 14, APRIL 2002 causes a depressurization of the containment, and redistributes the noncondensables as the steam in the drywell is condensed.

Table 6.2-8 provides the peak pressure, temperature, and time parameters for the recirculation line break as predicted for the conditions of Table 6.2-1 and in correspondence with Figures 6.2-2 and 6.2-3.

The transient peak calculated drywell floor (deck) differential pressure is 24.2 psid, which is 3.2% below the design sustained differential pressure of 25 psid.

During the blowdown period of the LOCA, the pressure suppression vent system conducts the flow of the steam-water gas mixture in the drywell to the suppression pool for condensation of the steam. The pressure differential between the drywell and suppression pool controls this flow vers us time. Figure 6.2-4 provides the mass flow versus time relationship through the vent system for this accident. A supplementary evaluation has been performed for the addition of feedwater to the suppression pool to assess the impact on long term pool temperature. This evaluation estimates that the peak short term pool temperature will increase by an additional 15.4 F. This results in a short term pool temperature (at 600 seconds) of approximately 166 F . For further discussion, s ee Section 6.2.1.1.3.1.1 in the paragraph titled, "Evaluation of Post-LOCA Feedwater Injection."

Long-Term Accident Responses In order to assess the adequacy of the containment following the initial blowdown transient, an analysis was made of the long-term temperature and pressure response following the accident. The analysis assumptions are those discussed previously for the three cases of interest. The initial pressure response of the containment (the first 600 seconds after th e break) is the same for each case.

Case A - All ECCS Equipment Oper ating (with containment spray) This case assumes that offsite a-c power is available to operate all cooling systems.

During the first 600 seconds following the pi pe break, the high-pressure core spray (HPCS), low-pressure core spray (LPCS), and all three LPCI pumps are assumed operating. All flow is injected directly into the reactor vessel.

After 600 seconds, both RHR heat exchangers are activated to remove energy from the containment. During this mode of operation the flow from two of the LPCI pumps is routed through the RHR heat exchanger, where it is cooled before being discharged into the containment spray header.

The containment pressure response to this set of conditions is shown as curve A in Figure 6.2-5. The corresponding drywell and suppression pool temperature responses are shown as curve A in Figure s 6.2-6 and 6.2-7. After the initial blowdown and subsequent depressurization due to core spray and LPCI core LSCS-UFSAR 6.2-15 REV. 13 flooding, energy addition due to core decay heat results in a gradual pressure and temperature rise in the containment. When the energy removal rate of the RHR exceeds the energy addition rate from the decay heat, the containment pressure and temperature reach a second peak valu e and decrease gradually. Table 6.2-5 summarizes the cooling equipment operation, the peak containment pressure following the initial blowdown peak, and the peak suppression pool temperature.

Case B - Loss of Offsite Power (with containment spray) This case assumes no offsite power is available following the accident with only minimum diesel power. The containment spray is operating and injecting into the drywell after 600 seconds. During this mode of operation the LPCI flow through one RHR heat exchanger is discharged into the containment spray nozzles.

The containment response to this set of co nditions is shown as curve B in Figure 6.2-5. The corresponding dyrwell and suppression pool temperature responses are shown as curve B in Figures 6.2-6 and 6.2-7. A summary of this case is given in Table 6.2-5.

Case C - Loss of Offsite Power (no containment spray)

This case assumes that no offsite power is available following the accident, with only minimum diesel power. For the fi rst 600 seconds following the accident, one HPCS and two LPCI pumps are used to cool the core. After 600 seconds the spray may be manually activated to further reduce containment pressure if desired. This analysis assumes that the spray is not activated.

After 600 seconds, one RHR heat exchanger is activated to remove energy from the containment. During this mode of operation, one of the two LPCI pumps is shut down and the service water pumps to the RHR heat exchanger are activated. The LPCI flow is cooled by the RHR heat exchanger before being discharged into the reactor vessel.

The containment pressure response to this set of conditions is shown as curve C in Figure 6.2-5. The corresponding drywell and suppression pool temperature responses are shown as curve C in Figures 6.

2-6 and 6.2-7. A summary of this case is given in Table 6.2-5.

When comparing the "spray" Case B with th e "no spray" Case C, the same duty on the RHR heat exchanger is obtained since the suppression pool temperature response is approximately the same as shown in Figure 6.2-7. Thus, the same amount of energy is removed from the pool whether the exit flow from the RHR heat exchanger is injected into the reactor vessel or into the drywell as spray. However, the peak containment pressure is higher for the "no spray" case, but the pressure is LSCS-UFSAR 6.2-16 REV. 13 still much less than the containment design pressure of 45 psig. (Subsection 6.2.2.3 describes the containment cooling mode of the RHR system.)

A supplemental evaluation has been performed for the purpose of evaluating the suppression pool long term temperature response. For this evaluation, the feedwater is assumed to have been injected into the suppression pool, by the end of the recirculation piping break blowdown phase (at time t = 600 seconds), in order to assess long term peak pool temperature. See paragraph entitled "Evaluation of Post-LOCA Feedwater Injection" in this se ction. Additionally, a slightly reduced RHR pump flow rate of 7200 gpm (versu s 7450 gpm) has been evaluated, as discussed in Section 6.2.2.3.4. Both of these evaluations are evaluated for the DBA-LOCA in Reference 18. The results indicate an increase in the long term peak suppression pool temperature of approxim ately 8 F due to the feedwater injection and an approximately 1.5 F increase due to the lower RHR flow rate. The 200 F peak pool temperature given in Table 6.2-5 is not exceeded. Plant specific safety evaluations have been performed and have concluded that the existing DBA-LOCA analyses referenced above bounds thes e effects on the containment response.

Energy Balance During Accident In order to establish an energy distribution as a function of time (short term, long term) for this accident, the following energy sources and sinks are required:

a. blowdown energy release rates, b. decay heat rate and fuel relaxation energy, c. sensible heat rate, d. pump heat rate, and
e. heat removal rate from suppression pool.

Items a, b, and c are provided in Subsection 6.2.1.3. The pump heat rate value that has been used in the evaluation of the containment response to a LOCA for Case A is 4881 Btu/sec. A complete energy balance for the recirculation line break accident is given in Table 6.2-6 for the reactor system, the containment, and the containment cooling systems at time zero, at the time of peak drywell pressure, at the end of reactor blowdown, and at the time of the long-term second peak pressure reached in the containment.

The energy and mass balance have been annotated to include the effects of feedwater coastdown/injection on the long te rm peak suppression pool temperature. See paragraph entitled "Evaluation of Post-LOCA Feedwater Injection" in this section and footnote in Table 6.2-6.

LSCS-UFSAR 6.2-17 REV. 13 Chronology of Accident Events The complete description of the containment response to the design-basis recirculation line break has been given above. Results for this accident are shown in Figures 6.2-2 through 6.2-7. A chronological sequence of events for this accident from time zero is provided in Table 6.2-7.

The original and 1988 General Electric co ntainment analysis (references 8 & 17), assumed feedwater flow stopped at the initiation of the LOCA. This assumption is conservative for an assessment on the peak cladding temperature (PCT) or containment pressure and temperature response. However, in order to make a more conservative analysis on the suppression pool predicted temperatures, the feedwater energy due to feedwater pump coastdown, or depressurization and resulting feedwater liquid carryover to the pool, should be taken into account in the suppression pool energy balance. A supplementary evaluation was performed to assess the impact on peak suppression pool temperature due to the addition of energy from the feedwater system. (Reference 18)

For this evaluation, the feedwater mass downstream of the 2nd Low Pressure Feedwater Heater is injected into the vessel. The feedwater upstream of this feedwater heater is at a temperature less than 212 F and would not be expected to be injected into the vessel during a DBA-LOCA. The mechanism for FW injection into the vessel during a LOCA with loss of onsite power is fl ashing of feedwater liquid when the vessel drops below the sa turation pressure corresponding to the feedwater liquid temperature. Thus, only feedwater initially at a temperature above 212 F is assumed to flash and be injected into the vessel. This is conservative since vessel pressures are expected to remain higher than atmospheric pressure during the period when the peak pool temperature occurs. The latest revision of plant piping drawings were used as input to determine the feedwater volume. Additionally, the sensible energy in the feedwater system metal is also added to the feedwater liquid injected into the vessel. It is conservatively assumed that the feedwater flowing into the vessel and coming into contact with hotter feedwater piping metal downstream, will instantaneously achieve thermal equilibrium with the hotter feedwater system metal. This maximizes the metal sensible energy transfer to the feedwater.

For the analysis, all feedwater mass and energy is injected to the vessel and subsequently transferred to the suppre ssion pool by 600 seconds into the LOCA event. This is modeled by adding all the feedwater mass and energy input at time t

= 600 seconds. Based on this previous discussion, this analysis provides a conservative estimate of the amount of energy addition to the pool due to feedwater injection.

LSCS-UFSAR 6.2-18 REV. 18, APRIL 2010 The results indicate an increase in the long term peak suppression pool temperature of approximately 8 F (Reference 18). The 200 F peak pool temperature given in Table 6.2-5 is not exceeded.

Evaluation at 3559 MWt Reactor Power The analysis of an instantaneous guilloti ne rupture of a recirculation line at 3559 MWt reactor power, Reference 25, employed essentially the same methodology as the 3434 MWt analysis, except for the RPV blow down calculation in the short-term containment response analysis. The blowdown calculation was performed using the LAMB break flow model (Reference 26), which models physical phenomena in the pipe and vessel in a more detailed manner. The LAMB break flow rate and enthalpy calculated at initial reactor power of 3559 MWt and initial pressure of 1025 psig were used as input to the containment analysis model in the short-term analysis. For the analysis of the long-term containment response, Case C, which was the limiting case among the three cases (Cases A, B, and C) analyzed at 3434 MWt reactor power, was analyzed at 3559 MWt. The analysis of Case C at 3559 MWt had the same assumptions as the original analysis at 3434 MWt with respect to the availability of the ECCS pumps and RHR heat exchanger.

The key input assumptions updated for the analysis at 3559 MWt are: a) the core decay heat is based on the ANSI/ANS 5.1-1979 decay heat model with a two-sigma uncertainty adder (the decay heat calculations also include contributions from miscellaneous actinides and activation products consistent with the recommendation of GE SIL 636.); and b) the water in the feedwater system continues to flow into the RPV until all feedwater above 212ºF is depleted to maximize pool heat-up.

Table 6.2-3a shows initial conditions assumed for the analysis of the design basis recirculation line rupture at 3559 MWt.

The analysis results are tabulated and plotted, as follows. Tables 6.2-5a and 6.

2-8a show a summary of the analysis results for the long-term and short-term responses, respectively. The short-term containment pressure and temperature responses are plotted in Figures 6.2-2a and 6.2-3a, respectively. Figure 6.2-5a provides the long-term containment pressure response. The long-term drywell airspace and pool temperature responses are given in Figure 6.2-6a and 6.2-7a respectively.

Evaluation at 3559 MWt Reactor Power and Considering GE SC06-01 General Electric (GE) Safety Communication SC06-01 identified an alternate single failure that can be more limiting with respect to peak suppression pool temperature during the Design Basis LOCA than reported in the existing license basis analysis.

The current licensing basis analysis (Reference 31) assumes the single failure of an emergency diesel generator. In this si tuation one residual heat removal (RHR) division is lost and only minimum emer gency core cooling systems (ECCS) and RHR containment cooling pumps are available (reference UFSAR 6.2.

1.1.3.1.1 Case C).

LSCS-UFSAR 6.2-18a REV. 20, APRIL 2014 This results in minimum suppression pool cooling, however this assumed failure also minimizes the pump heat to the suppression pool. As described in GE SC06-01, an alternate worst-case accident scenario with respect to suppression pool temperature may exist where the postulated single failure results in loss of one RHR division, but with all ECCS pumps remaining available. In this configuration, the pump heat to the suppression pool is maximized and can result in a higher peak suppression pool temperature. An analysis was performed in Reference 33 that determines the impact of the concerns of GE SC06-01. The analysis uses the same inputs and assumptions of Reference 31 except for the following significant differences: 1. All ECCS pumps are assumed to be available and operate in accordance with their design requirements for reactor vessel coolant make-up. 2. A single active failure is assumed, which results in only one RHR heat exchanger being operable for containment cooling for the duration of the event. 3. A RHR service water temperature of 107F and a RHR heat exchanger K-factor of 438 Btu/sec-F have been evaluated in Reference 36. The resultant maximum long-term post DBA-LOCA suppression pool temperature is 197F. 6.2.1.1.3.1.2 Main Steamline Break The main steamline break, which is not the limiting event with respect to the containment response, was not analyzed at a reactor power of 3559 MWt. The original analysis at 3434 MWt is presented in this subsection. The sequence of events immediately following the rupture of a main steamline between the reactor vessel and the flow limiter has been determined. The flow on both sides of the break will accelerate to the maximum allowed by critical flow considerations. In the side adjacent to the reactor vessel, the flow will correspond to critical flow in the 2.98-ft2 steamline cross section. Blowdown through the other side of the break can occur because the steamlines are all interconnected at a point upstream of the turbine by the bypass header. This interconnection allows primary system fluid to flow from the three unbroken steamlines, through the header and back into the drywell via the broken line. Flow will be limited by critical flow in the 0.94-ft2 steamline flow restrictor. The total effective flow area is thus 3.92 ft2, which is the sum of the steamline cross-sectional area and the flow restrictor area. Subsection 6.2.1.3 provides information on the mass and energy release rates.

LSCS-UFSAR 6.2-18b REV. 18, APRIL 2010 Immediately following the break, the total steam flow rate leaving the vessel would be approximately 12,000 lb/sec, which exceeds the steam generation rate in the core of 4,500 lb/sec. This steam flow to steam generation mismatch causes an initial depressurization of the reactor vessel at a rate of 50 psi/sec. The void formation in the reactor vessel water causes a rapid rise in the water level, and it is conservatively assumed that the water level reaches the vessel steam nozzles 1 second after the break occurs. The water level rise time of 1 second is the minimum that could occur under any reactor operating condition. From that time on, a two-phase mixture would be discharged from the break. During the first second of the blowdown, the blowdown flow will consist of saturated reactor steam. This steam will enter the containment in a super-heated condition of approximately 330 F. Figures 6.2-8 and 6.2-9 show the pressure and temperature response of the drywell and containment during the primary system blowdown phase of the accident.

Figure 6.2-9 shows that the drywell atmosphere temperature approaches 330 F after 1 second of primary system steam blowdown. At that time, the water level in the vessel will reach the steamline nozzle elevation and the blowdown flow will change to a two-phase mixture. This increased flow causes a more rapid drywell pressure rise. However, the peak differential pressure is 24.2 psid, which occurs shortly after the vent clearing transient. As the blowdown proceeds, the primary system pressure and fluid inventory will decrease and this will result in reduced break flow rates.

LSCS-UFSAR 6.2-19 REV. 14, APRIL 2002 As a consequence, the flow rate in the vent system also starts to decrease, and this results in a decreasing differential pressure between the drywell and containment.

Table 6.2-8 presents the peak pressures, peak temperatures, and times of this accident as compared to the recirculation line break.

Approximately 50 seconds after the start of the accident, the primary system pressure will have dropped to the drywell pressure and the blowdown will be over. At this time the drywell will contain pure steam, and the drywell and suppression chamber pressures will stabilize at approximately 30 and 25 psig, respectively; the difference corresponds to the hydrostatic pressure at the lower end of the submerged vents.

The drywell and containment will remain in this equilibrium condition until the reactor pressure vessel refloods. During this period, the emergency core cooling pumps will be injecting cooling water from the suppression pool into the reactor. This injection of water will eventu ally flood the reactor vessel to the level of the steamline nozzles, and at this time, the ECCS flow will spill into the drywell. The water spillage will condense the steam in the drywell and thus reduce the drywell pressure.

As soon as the drywell pressure drops below the suppression chamber pressure, the drywell vacuum breakers will open and noncondensable gases from the suppression chamber will flow back into the drywell. This process will continue until the pressures in the two regions equalize and stabilize at approximately 7.5 psig.

6.2.1.1.3.1.3 Intermediate Breaks The intermediate-size break, which is not the limiting event with respect to the containment response, was not analyzed at a reactor power of 3559 MWt. The original analysis at 3434 MWt is presented in this subsection.

The failure of a recirculation line results in the most severe pressure loading on the drywell structure. However, as part of the containment performance evaluation, the consequences of intermediate breaks are also analyzed. This classification covers those breaks for which the blowdown will result in reactor depressurization and operation of the ECCS. This subsection describes the consequences to the containments of a 0.1-ft 2 break below the RPV water level. This break area was chosen as being representative of the intermediate size break area range. These breaks can involve either reactor steam or liquid blowdown.

Following the 0.1-ft 2 break, the drywell pressure increas es at approximately 1 psi/sec. This drywell pressure transient is sufficiently slow so that the dynamic effect of the water in the vents is negligible and the vents will clear when the drywell-to-wetwell differential pressure is equal to the ve nt submergence pressure. For the LSCS containment design, the maximum distance between the pool surface and the bottom of the vents is 12 feet 10 inches. Thus, the water level in the vents will reach this point when the drywell-to-containment pressure differential reaches 5.2 psid.

LSCS-UFSAR 6.2-20 REV. 14, APRIL 2002 Figures 6.2-10 and 6.2-11 show the drywell and wetwell pressure and temperature response, respectively. The ECCS respon se is discussed in Section 6.3.

Approximately 5 seconds after the 0.1-ft 2 break occurs, air, steam, and water will start to flow from the drywell to the suppression pool; the steam will be condensed and the air will enter the wetwell free space. After 5 seconds there will be a constant pressure differential of 5.2 psid between the drywell and wetwell. The continual purging of drywell air to the suppression chamber will result in a gradual pressurization of both the wetwell and drywell to about 22 and 27 psig, respectively. Some continuing containment pressurization will occur because of the gradual pool heatup. The ECCS will be initiated by the 0.1-ft 2 break and will provide emergency cooling of the core. The operation of these systems is such that the reactor will be depressurized in approximately 600 seconds. This will terminate the blowdown phase of the transient. The drywell w ill be at approximately 27 psig and the suppression chamber at approximately 22 psig.

In addition, the suppression pool tempera ture will be the same as following the DBA because essentially the same amount of primary system energy would be released during the blowdown. After reactor depressurization, the flow through the break will change to suppression pool water that is being injected into the RPV by the ECCS. This flow will condense the drywell steam and will eventually cause the drywell and containment pressures to equalize in the same manner as following a recirculation line rupture.

The subsequent long-term suppression pool and containment heatup transient that follows is essentially the same as for the recirculation break.

From this description, it can be concluded that the consequences of an intermediate size break are less severe than those from a recirculation line rupture.

6.2.1.1.3.1.4 Small Size Breaks The small-size break, which is not the limiting event with respect to the containment response, was not analyzed at a reactor power of 3559 MWt. The original analysis at 3434 MWt is presented in this subsection.

Reactor System Blowdown Considerations This subsection discusses the containment transient associated with small primary system blowdowns. The sizes of primary sy stem ruptures in this category are those blowdowns that will not result in reactor depressurization due either to loss of LSCS-UFSAR 6.2-20a REV. 14, APRIL 2002 reactor coolant or automatic operation of the ECCS equipment. Following the occurrence of a break of this size, it is assumed that the reactor operators will initiate an orderly plant shutdown and depressurization of the reactor system. The thermodynamic process associated with th e blowdown of primary system fluid is one of constant enthalpy. If the primary system break is below the water level, the blowdown flow will consist of reactor water. Blowdown from reactor pressure to the drywell pressure will flash approximately one-third of this water to steam and two-LSCS-UFSAR 6.2-21 REV. 13 thirds will remain as liquid. Both ph ases will be at saturation conditions corresponding to the drywell pressure. Thus, if the drywell is at atmospheric pressure, the steam and liquid associated with a liquid blowdown would be at 212 F. Similarly, if the containment is assumed to be at its design pressure, the reactor coolant will blow down to approximately 293 F steam and water.

If the primary system rupture is located so that the blowdown flow consists of reactor steam only, the resultant steam temperature in the containment is significantly higher than the temperature associated with liquid blowdown. This is because the enthalpy of high-energy saturated steam is nearly twice that of saturated liquid. The higher enthalpy will result in a superheat condition. For example, decompression of 1000-psia steam to atmospheric pressu re will result in 298 F superheated steam (86 F of superheat).

Based upon this thermodynamic process, it is concluded that a small reactor steam leak will impose the most severe temperatu re conditions on the drywell structures and the safety equipment in the drywell. For larger steamline breaks, the superheat temperature is nearly the same as for small breaks, but the duration of the high-temperature condition is less. This is because the larger breaks will depressurize the reactor more rapidly than the orderly reactor shutdown that is assumed to terminate the small break.

Containment Response For drywell design consideration, the following sequence of events is assumed to occur. With the reactor and containment operating at the maximum normal conditions, a small break occurs that allows blowdown of reactor steam to the drywell. The resulting pressure increase in the drywell will lead to a high drywell pressure signal that will scram the reacto r and activate the containment isolation system. The drywell pressure will continue to increase at a rate dependent upon

the size of the steam leak. This pressure increase will lower the water level in the vents until the level reaches the bottom of th e vents. At this time, air and steam will start to enter the suppression pool. The steam will be condensed and the air will be carried over to the suppression chamber free space. The air carry-over will result in a gradual pressurization of the containment at a rate dependent upon the size of the steam leak. Once all the drywell air is carr ied over to the suppression chamber, pressurization of the containment will cease and the system will reach an equilibrium condition with the drywell pressure at 27 psig and the suppression chamber at approximately 22 psig. The drywell will contain only superheated steam, and continued blowdown of reactor steam will condense in the suppression pool.

LSCS-UFSAR 6.2-22 REV. 13 Recovery Operations The reactor operators will be alerted to the incident by the high drywell pressure signal and the reactor scram. For the purposes of evaluating the duration of the superheat condition in the drywell, it is assumed that their response is to cool down the reactor in an orderly manner using any method, but limiting the reactor cooldown rate to 100 F per hour. The normal method to achieve recovery is by use of the high pressure core spray in conjunction with the automatic depressurization system. This feed and bleed process can be utilized until the reactor is depressurized. Depending upon their availability and the situation, other methods such as the use of turbine bypass valves in conjunction with the main condenser can be utilized to achieve depressurization. This will result in the reactor primary system being depressurized within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />. Drywell Design Temperature Considerations For drywell design purposes, it is assumed that there is a blowdown of reactor steam for the 6-hour cooldown period. The corresponding design temperature is determined by finding the combination of primary system pressure and containment pressure that produces the maximum superheat temperature. Thus for design purposes, this results in a temperature condition of 340 F. 6.2.1.1.3.2 Accident Analysis Models The short-term pressurization analytical models, assumptions, and methods used by GE to evaluate the containment response during the reactor blowdown phase of a LOCA are described in References 2 and 3. Once the RPV blowdown phase of the LOCA is over, a fairly simple model of the drywell and suppression chamber may be used. During the long-term, post-blowdown containment cooling mode, the ECCS flow path is a closed loop and the suppression pool mass will be constant. Schematically, the cooling model loop is shown in Figure 6.2-12. Since there is no storage other than in the suppression pool (the RPV is reflooded during the blowdown phase of the accident), the mass flowrates shown in the figure are equal, thus: eccsSDmmmOO LSCS-UFSAR 6.2-23 REV. 13 Analytical Assumptions The key assumptions employed in the model are as follows: a. The drywell and suppression chamber atmosphere are both saturated (100% relative humidity). b. The drywell atmosphere temperature is equal to the temperature of the coolant spilling from the RPV, or to the spray temperature if the sprays are activated. c. The suppression chamber atmosphere temperature is equal to the suppression pool temperature or to the spray temperature if the sprays are activated. d. No credit is taken for heat losses from the primary containment or to the containment internal structures. Energy Balance Considerations The rate of change of energy in the suppression pool, Ep, is given by: shswMdtdpEdtd .s.ssshdtdMMdtdhww Since _d_d_t (Mws) = 0 (because there is no storage), and for water at the conditions that will exist in the containment: where: Cp = 1.0 for the specific heat of pool water, Btu/ lb-F Ts = pool temperature, F. The pool energy balance yields: h m h m T dtd C Mso sDo D sps w This equation can be rearranged to yield: sTdtdpCshdtd LSCS-UFSAR 6.2-24 REV. 13 An energy balance on the RHR heat exchanger yields (6.2-3) where: hc = enthalphy of ECCS flow entering the reactor, Btu/lb. Similarly, an energy balance on the RPV will yield: Combining Equations 6.2-1, 6.2-2, 6.2-3, and 6.2-4 gives This differential equation is integrated by finite difference techniques to yield the suppression pool temperature transient. Containment Thermodynamic Conditions Once the energy equations are solved, the drywell and suppression chamber atmospheric temperatures can be calculated. swMsh osm Dh oDm sT dtd oxsHscmqhh eccsmeqDqchDh swMXHqeqDqsTdtd LSCS-UFSAR 6.2-25 REV. 13 For the case in which no containment spray is operating, the suppression chamber temperature, Tw, at any time will be equal to the current temperature of the pool, Ts, and the drywell temperature, Td, will be equal to the temperature of the fluid leaving the RPV. Thus: and Tw = Ts. For the case in which the containment spray is assumed to be operating, both the drywell and suppression chamber atmosphere will be at the spray temperature, Tsp where: eccsmxHqsTspT and, TD = Tw = Tsp. Using the suppression chamber and drywell atmosphere temperatures, and assumption (a) (drywell and suppression chamber saturated), it is possible to solve for the containment total pressures, since: (6.2-6) (6.2-7) where: PD = drywell total pressure, psia, PaD = partial pressure of air in drywell, psia, PvD = partial pressure of water vapor in drywell, psia, Ps = suppression chamber total pressure, psia, Pa s = partial pressure of air in the suppression chamber, psia, eccsmxHqeqDqsTDT DvPDaPDP svPsaPsP LSCS-UFSAR 6.2-26 REV. 13 Pv s = partial pressure of water vapor in the suppression chamber, psia, and, from the Ideal Gas Law: (6.2-8) (6.2-9) where: MaD = mass of air in drywell, lb, Mas = mass of air in the suppression chamber, lb, R = gas constant ft-lbf/lb VD = drywell free volume, ft3. Vs = suppression chamber free volume, ft3. With known values of TD and Tw , Equations 6.2-6, 6.2-7, 6.2-8 and 6.2-9 can be solved by transient analysis and iteration. This iteration procedure is also used to calculate the unknown quantities MaD and Mas. Solution of Equations The transient analysis is based on successive time step integration of the suppression pool temperature. When this integration has been performed and the value of Ts at the end of a time step has been calculated, a pressure balance is made. Using values of MaD and Mas from the end of the previous time step and the updated values of TD and Ts, a check is made to see if Ps is greater than or equal to PD using Equations 6.2-6, 6.2-7, 6.2-8, and 6.2-9. If Ps is greater than or equal to PD, then the two values are made equal. The vacuum breakers between the drywell and suppression chamber are provided to ensure that Ps cannot be greater than PD. 144DVDRTDaMDaP 144sVwRTsaMsaP LSCS-UFSAR 6.2-27 REV. 13 Hence, with PD = Ps and knowing that: MaD + Mas = constant; (6.2-10) where the constant is the known total initial mass of air in the suppression chamber and drywell prior to the accident, Equations 6.2-6, 6.2-7, 6.2-8, and 6.2-9 can be solved for Mas, MaD , and Ps/PD. It is conservatively assumed that the total mass of air remains constant, which ignores any containment leakage that might occur during the transient. If, as a result of the end-of-time-step pressure check, where: H = submergence of vents, ft, and Vw = specific volume of fluid in vent, ft3/lb then the pressure in the drywell is higher than the pressure in the suppression chamber but not sufficiently so to depress the water to the bottom of the vents and thus permit air to flow from the drywell to the suppression chamber. Under these circumstances, no air transfer is assumed to have occurred during the time step, and Equations 6.2-6, 6.2-7, 6.2-8, and 6.2-9 are solved using the updated temperatures with the same Mas and MaD values from the previous time step. If the end-of-time step pressure check shows: then the drywell pressure is set to the value: (6.2-11) 'wVHsPDPsP wVHsPDP VHsPDP LSCS-UFSAR 6.2-28 REV. 15, APRIL 2004 This requires that the drywell pressure never exceed the suppression chamber pressure by more than the hydrostatic head associated with the submergence of the vents. To maintain this condition, some transfer of drywell air to the suppression chamber will be required. The amount of air transfer is calculated by using Equation 6.2-10 and combining Equations 6.2-6, 6.2-7, 6.2-8, 6.2-9 and 6.2-11 to give: wswsasvDDDaDvvHV144RTMPV144RTMP which can be solved for the unknown air masses. The total pressures can then be determined. 6.2.1.1.4 Negative Pressure Design Evaluation Containment negative pressure has been addressed in Chapter 3.0 and in the Design Assessment Report. 6.2.1.1.5 Suppression Pool Bypass Effects Protection Against Bypass Paths The pressure boundary between drywell and suppression chamber including the vent pipes, vent header, and downcomers are fabricated, erected, and inspected by nondestructive examination methods in accordance with and to the acceptance standards of the ASME Code Section III, Subsection B, 1971 (Summer 1972 Addenda). This special construction, inspection and quality control ensures the integrity of this boundary. The design pressure and temperature for this boundary was established at 25 psid and 340 F, which is substantially greater than conditions during a DBA. Actual peak accident differential pressure and temperature across this boundary will be less than their design values during a LOCA. In addition a stainless steel liner has been provided between the drywell and the wetwell as described in Chapter 3.0. All penetrations of this boundary except the vacuum breaker seats and suppression pool temperature monitoring probe penetrations and testing penetrations are welded. All penetrations are available for periodic visual inspection. The following paragraphs describe the evaluation of the steam bypass event at 3434MWt. The limiting event was analyzed for a reactor power level of 3559 MWt, and it was concluded that this reactor power has no significant impact on the suppression pool steam bypass.

LSCS-UFSAR LU2000-027I 6.2-28a REV. 14, APRIL 2002 Reactor Blowdown Conditions and Operator Response In the highly unlikely event of a reactor depressurization to the drywell accompanied by a simultaneous open bypass path between the drywell and suppression chamber, several postulated conditions may occur. For a given primary system break area, the maximum allowable leakage capacity can be determined LSCS-UFSAR 6.2-29 REV. 17, APRIL 2008 when the containment pressure reaches the design pressure at the end of reactor blowdown. The most limiting conditions would occur for those primary system break sizes which do not cause rapid reactor depressurization. This corresponds to breaks of less than approximately 0.4 ft 2 which require some operator action to terminate the reactor blowdown.

Immediately after the postulated conditions given above for a small primary system break, there would be a fairly rapid rise in containment pressure as the noncondensable gases in the drywell are carried over to the suppression chamber. During this portion of the transient, it is assumed that the plant operators are unaware that a leakage path exists. Under normal circumstances, the maximum pressure that can occur in the suppression chamber is approximately 25 psig. This is the pressure that would result if all of the noncondensable gases initially in the

containment are carried over to the suppression chamber free space. For the maximum allowable leakage calculations, it was assumed that the plant operators realize a leakage path exists only when the suppression chamber pressure reaches 30 psig. For conservatism, an additional 10-minute delay is assumed before any corrective action is taken to terminate the transient. The corrective action is also assumed to take 5 minutes to be effective. At that time, the containment pressure would be equal to the design pressure if the allowable leakage had occurred. The specific type of corrective action taken after 10 minutes is not accounted for in the analysis. The operators have several options available to them. If the source of the leakage is undefined, they could depressurize the primary system via either the main condenser or relief valves, or they could activate the containment sprays.

Analytical Assumptions

When calculating the allowable leakage capacities for a spectrum of break sizes, the following assumptions are made:

a. Flow through the postulated leakage path is pure steam. For a given leakage path, if the leakage flow consists of a mixture of liquid and vapor, the total leakage mass flow rate is higher, but the steam flowrate is less than for the case of pure steam leakage. Since the steam entering the suppression chamber free space results in the additional containment pressurization, this is a conservative assumption.
b. There is no condensation of the le akage flow on either the suppression pool surface or the containment an d vent system structures. Since condensation acts to reduce the suppression chamber pressure, this is a conservative assumption. For an actual containment there will be condensation, especially for the larger primary system breaks where vigorous agitation at the pool surface will occur during blowdown.

Analytical Results

LSCS-UFSAR 6.2-30 REV. 17, APRIL 2008 The LSCS containment has been analyzed to determine the allowable leakage between the drywell and suppression chamber. Figure 6.2-13 shows the allowable leakage capacity )K/A( as a function of primary system break area. A is the area of the leakage flow path and K is the total geometric loss coefficient associated with the leakage flow path. The maximum allowable leakage capacity is at )K/A( = .030 ft2. Since a typical geometric loss factor would be 3 or greater, the maximum allowable leakage area would be .052 ft2. This corresponds to a 3-inch line size. Figure 6.2-13 is a composite of two curves. If the break area is greater than approximately 0.4 ft2, reactor depressurization will terminate the transient and allow higher leakage. However break areas less than 0.4 ft2 result in continued reactor blowdown which limits the allowable leakage. Figure 6.2-14 shows the containment response associated with breaks larger than 0.4 ft2. The containment pressure would reach design pressure at the end of reactor blowdown. Figure 6.2-15 shows the same response for a typical small break less than 0.4 ft2. The containment pressure would reach design conditions, in this case, approximately 5 minutes after operator action. 6.2.1.1.6 Suppression Pool Dynamic Loads The manner in which suppression pool dynamic loads resulting from postulated loss-of-coolant accidents, transients, and seismic events have been integrated into the LSCS design is completely described in the LaSalle Design Assessment Report, which was submitted with the FSAR as a reference document. The load histories, load combinations, and analyses are all presented in detail in this referenced report. A safety relief valve in-plant test was conducted on unit 1 as committed by Commonwealth Edison per NUREG-0519. A report entitled "Commonwealth Edison Proprietary LaSalle County I In-Plant S/RV Test Initial Evaluation Report" was submitted March 4, 1983 (C. W. Schroeder to A. Schwencer) and resubmitted October 14, 1983 (C.W. Schroeder to H.R. Denton). The document contains information and data demonstrating the adequacy of existing design basis hydrodynamic loads resulting from safety/relief valve actuation. Supplementary evaluations have been performed, as discussed in Section 6.2.1.8, to verify that an increase in the initial suppression pool temperature (from 100 F to 105 F) would not significantly impact the dynamic loading scenarios associated with containment response to postulated LOCAs and SRV operation. Containment Dynamic Loads were evaluated for power uprate to 3489MWt in Reference 25. The evaluation shows the LOCA and SRV loads remain within the defined limits.

LSCS-UFSAR 6.2-31 REV. 13 6.2.1.1.7 Asymmetric Loading Conditions The manner in which potential asymmetric loads were considered for LSCS is fully described in the Design Assessment Report.

A description of the analytical models utilized for these analyses, as well as a description of the containment testing program, is also presented in this report.

6.2.1.1.8 Containment Ventilation System The primary containment ventilation system is discussed in Section 9.4.

6.2.1.1.9 Postaccident Monitoring A description of the postaccident monitoring system is provided in Section 7.5.

6.2.1.1.10 Drywell-to-Wetwell Vacuum Breaker Valves Evaluation for LOCA Loads During the pool swell phase of a loss-of-coolant accident, air flows from the drywell through the vent pipes and the suppression pool into the suppression chamber air space resulting in a rise of the suppression pool surface and compression of the air space region above it. This transient wetwell air space pressurization may cause the vacuum breaker valves to experience hi gh opening and closing impact velocities. To estimate the valve disc actuation velocities, the Mark II Owner's Group developed a vacuum breaker valve dynamic model described in NEDE-22178-P(1), "Mark II Containment Drywell-to-Wetwell Vacuum Breaker Models," August 1982, which describes the generic methodology us ed to calculate the response of the drywell-to-wetwell vacuum breaker to certain transients in the Mark II containment. The LaSalle plant, however, is unique in that it is the only domestic Mark II plant which has its vacuum breakers located outside containment. Because of this feature, the Mark II Owners Group model was modified to take credit for the pressure losses associated with the exte rnal piping and isolation valves which connect the vacuum breaker between the wetwell and drywell at LaSalle. In a letter dated December 28, 1982, CECo submi tted a report to the NRC, CDI-82-33, "Reanalysis of the LaSalle Wetwell-to-Drywell Vacuum breakers under Pool Swell Loading Condition," December 1982, out lining the valve modeling improvement which have been made to take credit for the pressure losses associated with vacuum breaker piping. This report documents the reduction of the valve impact velocities during pool swell which are attributed to the use of a more realistic hydrodynamic torque on the valve disc. This analysis has been accepted by the NRC. However, because the hydrodynamic loads associated with a loss-of-coolant accident were not considered in the original design of th e vacuum breaker, CECo decided to modify the vacuum breakers to improve performance and reliability, and to further increase the margin of safety. The modifi cations included material upgrade and/or dimensional changes to strengthen eccentric shaft, hinge arms, hinge plates, fasteners and a load distribution device to reduce the severity of the vacuum LSCS-UFSAR 6.2-32 REV. 14, APRIL 2002 breaker pallet opening impact loading. Th e modified design was tested under an applied mechanical force which produced an opening pallet impact velocity of 20.2 radians/second and a closing impact veloci ty of 25.8 radians/second. The predicted pallet impact velocities for LaSalle are an opening impact velocity of 16.6 radians/second and a closing impact velocity of 24.2 radians/second. After testing, the vacuum breaker leak rate was verified to be within the acceptable limit. The test results verified the operability and fu nctional capability of the vacuum breaker well in excess of the predicted opening and closing impact velocities, and, thus, demonstrated that the modified LaSalle vacuum breakers will function properly under pool swell induced impact loadings with a considerable margin of safety.

6.2.1.1.11 Impact of Increased Initial Suppression Pool Temperature

Supplementary safety evaluations have been performed, as discussed in Section 6.2.1.8, to verify that an increase in the initial suppression pool temperature (from 100 F to 105 F) would not significantly impact the consequences of the various containment line break analyses.

6.2.1.2 Containment Subcompartments For the most part, the drywell is a large continuous volume interrupted at various locations by piping, grating, ventilation ducting, etc. The only two volumes within the drywell which can be classified as subcompartments are the annular volume between the biological shield and the reactor pressure vessel, and the volume bounded by the drywell head and the reactor vessel head. These regions are referred to as the biological shield a nnulus and head cavity, respectively, and require special design consideration resulting from the postulation of line breaks in these volumes.

6.2.1.2.1 Design Bases The methodology used to determine the containment subcompartment pressurization loads and the results pertaining to the pressurization loads documented herein are applicable to reactor operation at or below the bounding thermal power level of 3559 MWt (Reference 30).

Biological Shield Annulus Pressure transients within the biologic al shield annulus are important for two considerations: (1) determination of the design conditions for the shield wall, and (2) determination of the tipping forces on the reactor pressure vessel. It is not a priori clear that one line break will yield the most severe conditions for both considerations. Therefore, consequences of two line breaks were studied: (a) a LSCS-UFSAR 6.2-32a REV. 14, APRIL 2002 complete circumferential failure of one of the two recirculation outlet lines at the safe end to pipe weld, and (b) a complete circumferential failure of one of the six feedwater lines at the safe end to pipe weld. While it was assumed that the recirculation line break with its high mass and energy blowdown rates yields most severe shield wall loads, the break of the feedwater line was added to determine the most severe conditions on the vessel. The pressure transien ts following either LSCS-UFSAR 6.2-33 REV. 13 postulated break were used in determination of shield wall and pressure vessel design adequacy.

The performed pressurization analyses for the postulated recirculation line break and feedwater line break were based on the nodalization schemes depicted on Figures 6.2-16 and 6.2-17, respectively. Both nodalization schemes were given careful consideration to assure correct local and overall pressure responses.

Recirculation Line Break The sudden injection of the subcooled liquid into the shield penetration (Node 35) and adjoining annulus initially causes a sign ificant fraction of the liquid to flash to steam, pressurizing the penetrations and annulus. The responses of the penetration volume and adjoining subcompa rtments are shown on Figure 6.2-18. Within 10 milliseconds after the postulated break both flows out of the penetration have choked. Some 10 milliseconds later, both the penetration pressure and the pressure in the surrounding annulus node peak, reflecting subcooling and inventory effects addressed in the blowdown flow rates. Flow into the annulus initially proceeds in all directions, but soon swings preferentially upward in response to increasing pressures within the dead-ended skirt region. By 0.1 second into the transient, the pressures in and about th e penetration have stabilized and shortly after (by 0.5 seconds), the differential pressures across the shield wall have begun to decrease (Figure 6.2-21). The differential pressure across the sh ield wall peaks at 115 psid in the region immediately around the penetration. Peak differential pressure across the shield door in the penetration, however, reaches 325 psid.

Feedwater Line Break Pressurization effects of the postulated feedwater line break are much less pronounced than for the recirculation break. Much of the injected fluid finds its way up and out of the annulus and over the top of the shield wall and into the drywell. Nevertheless, the differential pressure across the shield wall surrounding the penetration peaks at 50 psid, while the differential pressure across the shield door in the penetration reaches 205 psid (F igure 6.2-22). By 0.5 second into the transient all the differential pressures across the shield wall have peaked and are decreasing (Figure 6.2-23).

The break area for the recirculation line break was assumed to be time dependent and limited by effects of pipe restraints (see Attachment 6A). The feedwater line break was assumed to provide instantaneous full size break area. Both break models included the effects of subcooled liquid inventory in the determination of

mass and energy flux data.

No margins were applied to the calculated differential pressures for this final pressurization analysis.

LSCS-UFSAR 6.2-34 REV. 13 Head Cavity The head cavity area was analyzed for specific line breaks. They were: 1) a break of the recirculation outlet line within the drywell; and 2) a break of the main steamline within the drywell; and, 3) a simultaneous break of the head spray line and the RPV head vent line within the head cavity. These analyses were carried out to establish the pressure differentials that would exist across the refueling bulkhead plate as a result of these a ccident conditions.

The break of the recirculation outlet line, the drywell DBA, was found to produce the highest pressure differential across the refuelin g bulkhead plate, a value of 9.0 psid upward. The simultaneous break of the he ad spray line and RPV head vent line caused a pressure differential of 7.0 psid downward. The main steamline data are not presented due to the fact that the recirculation line break produced the higher differential pressure value.

The break size, mass flow rate, and energy content for the recirculation line were defined in Subsection 6.2.1.

1.3.1 and Table 6.2-18. The supporting assumptions for these data are also supplied in the same su bsection. The break size, mass flow rate, and energy content for the head spray line were determined using Moody's flow

through the 3.72-inch diameter head spra y nozzle at reactor conditions with a multiplier of 1.0. Flow from the other side of the head spray line break was neglected. In addition, the simultaneous break of the RPV head vent line was considered because of the lack of whip restraints on the head spray line. The break size, mass flow rate, and energy conten t for the RPV head vent line were determined using Moody's flow at reactor co nditions with a multiplier of 1.0. The RPV head vent line was postulated to rupture at the four-to-two inch reducer in the line located in the head cavity. The flow occurred at both ends of the break, one having a diameter of 4.0 inch es and the other 2.0 inches.

No margin was applied to the results, since the analysis was done for the final design, and a margin is not required for that situation. However, a margin does exist, and this is indicated in Tables 6.2-11 and 6.2-12.

6.2.1.2.2 Design Features Biological Shield Annulus The biological shield annulus is an annul ar space 48.7 feet high and about 2 feet thick formed by the reactor pressure vessel and its skirt and the biological shield wall. The shield wall is provided with 32 penetrations to allow for routing for the lines connected to the vessel.

The shield wall is also pierced to provide 2 HVAC openings and 2 reactor skirt access doors. The 3-1/2 inch thermal insulation divides the shield annulus, except for the lower skir t portion, into 2 almost equal annului. The inner steel shell of the annulus is spanned with vertical and horizontal LSCS-UFSAR 6.2-35 REV. 13 stiffeners which extend 5 inches into the annulus. Egress to the drywell at the top of the shield is partially blocked by the gusset plates supporting the reactor vessel stabilizers (Figures 3.8-23). The penetratio ns in the shield wall are designed with shield doors with a gap of approximately 3 inches between the doors and the thermal insulation on the penetrating lines.

Figure 3.8-39 provides an exterior wall stretchout of the shield wall.

In the annulus pressurization analysis , it was assumed that following the postulated line break the vessel insulation within the annulus was instantaneously displaced to the shield wall. The vessel insulation support structure remains in its original configuration. Venting of the annulus into the drywell was possible through the annulus between the pipe and shield doors in the 32 nozzle penetrations in the shield wall and by mean s of an opening at the top of the shield wall above which the insulation was assume d to blow out instantaneously when the pressure across the insulation above the shield wall reaches 3 psid. Other possible vent paths such as HVAC openings, reac tor skirt access doors, and insulation blowout panels were assumed to remain closed.

Head Cavity

Note: The current flow paths have been changed to include the two manholes between the head cavity and the drywell and the four ducted HVAC vents have been modified by the addition of discharge nozzles. The impact of this change has been evaluated and it has been determined that the analysis presented here is bounding.

The physical system, shown in Figure 3.8-1, was modeled as three node with two flow paths for this analysis. The head cavity, drywell, and wetwell are all described by single volumes. The model for the simultaneous break of the head spray and RPV head vent lines in the head cavity is shown in Figure 6.2-19, and that for the recirculation line break in the drywell in Figure 6.2-20. The pertinent data regarding the volumes and flow paths are given in Tables 6.2-11 through 6.2-14.

There are eight HVAC vents in the refuelin g bulkhead plate: four sixteen-inch diameter supply vents, and four eighteen-inch diameter return vents. The return vents have ductwork attached to them. All of the HVAC (supply and return) were modeled for the postulated break in the head cavity since the pressure in the return vents with the ductwork would always be greater than the drywell pressure.

However, only the supply vents were considered to allow flow for the breaks in the drywell. It was assumed that the HVAC re turn ductwork would be crushed by the fast rising drywell pressure. The downcomer vents between the drywell and wetwell were modeled as one flow path with a valve in the path set to open at 0.824 second for the recirculation line brea

k. The 0.824 second was taken as a conservative estimate of the time normally required to clear the downcomer vents. At this time, the entire vent area beco mes available for pressure relief of the drywell and head cavity region. The si multaneous head spray line and RPV head LSCS-UFSAR 6.2-36 REV. 13 vent line break is a much smaller break and results in a relatively slow pressurization of the drywell. A valve was again used in the flow path, but in this instance, the valve opening was dependent upon the drywell pressure exceeding the hydrostatic head at the downcomer exit. The opening differential pressure used was 5.2 psid which is equivalent to a 12-foot downcomer submergence. The flow was carried over directly into the wetwell air volume. No credit was taken for condensation. The flow through both flow paths was taken to be a completely homogeneous mixture.

6.2.1.2.3 Design Evaluation Biological Shield Annulus

The RELAP 4 Mod 3 computer code was used to perform the analyses. The assumptions made in modeling the problem were in accordance with the applicable USNRC guidelines.

The mass and energy blowdown rates were determined according to the methods described in Attachment 6.A.

Initial conditions in the annulus and drywell are indicated in Tables 6.2-9 and 6.2-10. In subsonic flow conditions, two flow models were used, as defined in RELAP 4 Mode 3: (a) compressible flow, single stream model was used for the path of major flow direction, and (b) incompressible flow without momentum flux model was used for flow paths other than the paths of the major flow direction. For sonic flow conditions the Moody or sonic choking model were specified with the multiplier 0.6 for the Moody choking model. Homogeneous flow was assumed for the vent mixture. The biological shield annulus between th e reactor pressure vessel and the shield wall was modeled differently for each of th e two postulated line breaks. In either case, advantage was taken of the near sy mmetry of the annular space across the vertical plane passing through the centerline of the failed line.

Nodalization of the biological shield annulus was determined on the basis of natural geometric boundaries and the constraint that the pressure drop within a node be reasonably low as compared to pressure drop across the boundaries of the node. Nodal boundaries were suggested by the pr esence of the reinforcing steel, thermal insulation support structure and nozzles.

Significant pressure drops near the break suggested smaller nodes (by and large limited with two successive obstructions) around the penetration than elsewhere (Fig ures 6.2-37 and 6.2-38). Therefore the assumption was made that sinc e RELAP 4 allows input of lo ss coefficients only at the junctions between nodes, the junctions should be placed at points where major LSCS-UFSAR 6.2-37 REV. 13 pressure losses occur. Furthermore, it may be concluded that increasing the number of junctions (by making smaller nodes) beyond this point will yield no improvement in the accuracy of the results. To test this hypothesis, a sensitivity study was performed on the sacrificial shield nodalization. Using the original nodalization (Figure 6.2-39) as a basis, an "equivalent" model was run which maintained the nodalization near the break but drastically reduced the number of nodes further from the break (Figure 6.2-40). This model demonstrated identical pressure response close to the break and only minor differences away from the break (Figures 6.2-41 and 6.2-42). This indicated that the nodalization far from the break was sufficiently refined in the original model and that the "equivalent" model could be used to simulate a response close to the break. Two additional models were run. The first combined the nodes closest to the break into one large node (Figure 6.2-43). The pressure response was not consistent with the original runs (Figures 6.2-44 and 6.2-45). This indicated that a model which does not locate node boundaries at all flow restrictions close to the break is not acceptable. The last model substituted six nodes for the three original nodes, causing junctions to occur at locations which coincide with no actual flow restriction (Figure 6.2-46). This model showed a net increase of 5% in the force caused by the pressures in the area being investigated. An examination of the axial and circumferential pressure distributions showed only minor differences (Figures 6.2-47 and 6.2-48). The sensitivity study indicates that the original nodalization provides an adequate description of the pressurization of the sacrificial shield annulus. An increase in the complexity of the RELAP 4 model would not result in a significant change in the results. As previously indicated, half of the annulus was nodalized in case of either postulated line break; for the recirculation line break half-annulus consisted of 35 nodes and the half-drywell of 3 nodes (Table 6.2-9), while for the feedwater line break the half-annulus consisted of 29 nodes and the half-drywell of 3 nodes (Table 6.2-10). Volume of each node was calculated as a net volume, that is, the respective volume of the annulus including the volume of penetrations (if any) was corrected for the volume of the insulation and nozzles. The junctions, 85 and 69 for the recirculation line break and feedwater line break respectively, were assigned the smallest flow area anywhere between the centers of two volumes. All partial loss coefficients, kj's, were derived from Reference 6. The total loss coefficient kt was then determined by adding the weighted partial loss coefficients in series: 2iAtAiKitk LSCS-UFSAR 6.2-38 REV. 13 where At is the junction area and Ai is the area within the junction and pertaining to the partial loss coefficient k. When parallel paths, j, were combined, the following relations were utilized: Only similar junctions were combined in this manner (like 2 or more penetrations connecting drywell with the same volume of the annulus), other junctions were modeled separately. Inertia coefficients were similarly calculated using simplified conservative approximations to the integrated junction characteristics. Thus, for the junctions with only minor variations, in cross-sectional flow area along the junction, the inertia, I, was approximated by: where Li is the distance along the junction where junction's cross-sectional area is Ai. In cases where there appear major variations in the cross-sectional flow area (constriction in the conduit) the inertia was estimated by: where d is a "characteristic" diameter of the constriction of length Lo and with area Ao (for an orifice the characteristic diameter is taken to be the diameter of the orifice). L1, A1 and L2, A2 are the length and flow area of the conduit partitioned by the constriction. In special cases, where the constriction is not an ordinary orifice, a variation of the above relation was used to evaluate I. jAjtA 2 ik1tAiAitK iLitA1I 2Ad2LoAd2oL1Ad1LI LSCS-UFSAR 6.2-39 REV. 13 Parallel paths were characterized by: To further illustrate methods of determination of the junction characteristics, treatment of selected representative junctions will be shown in detail. The junctions are those for the recirculation line break nodalization scheme: 9, 47, 72. Junction 9 connects the break volume (node 35), which consists of the half-annulus in the recirculation line penetration extended from the shield door to the reactor vessel, with the surrounding annular node (34). The minimum junction area was in this case within the break volume, half of the annular area formed by the recirculation line and the penetration wall was calculated to be 7.04 ft2. In determining the loss coefficient for this junction, Diagram 11-9, Reference 6, was utilized. An upper limit value was set at 0.85 and considered the only loss for this junction. The inertia coefficient, I, for the junction was calculated as a sum of two contributions: (a) inertia through the half-annulus of the penetration (0.23), and (b) an upper limit estimate of the inertia within the annulus, node 34 (0.07), totaling 0.30 ft-1. Junction 47 is a vertical junction connecting nodes 16 and 21. The junction area is the related annulus cross-section area reduced by two constrictions, stiffener and the thermal insulation support structure. Although the constrictions appear at different elevations (11 inches apart), they were assumed at the same elevation. This assumption leads to the junction area of 7.72 ft2 (upstream volume flow area is 11.87 ft2 and the flow area of the downstream volume is 12.36 ft2). The loss coefficient was estimated using Diagram 4-9 of Reference 6, at 0.66 for flow area 7.72 ft2. The total junction loss coefficient is therefore 0.67. The junction area is characterized by the radial width of 1.45 feet. This width was taken as the characteristic length, d, for the purposes of the inertia coefficient determination. Then, using a variation of the above described relation for I, it was found that I = 0.45 ft -1. 1jI1jI 2AdLoAdI LSCS-UFSAR 6.2-40 REV. 13 Junction 72 is an example of the vent path through the line penetration and connects annular node 28 with the containment node 37. The actual penetration is located on the boundary between nodes 28 and 29. For this reason, only half of the penetration was treated as the junction 72.

The minimum area of the junction is the cr oss-sectional area of the half of annulus between the shield door and penetration line. It was determined to be 9.71 ft

2. Half-penetration flow area was calculated at 5.33 ft
2. The inertia coefficient for this junction was determined on the basis of the above areas and the characteristic diameter as being the hydraulic diameter at the penetration exit (3.3 ft

-1). The loss coefficient for the junction was, however, determined for the whole penetration and it consisted of a friction loss (0.02 for A = 10.65 ft 2), turning losses at the nozzle and contraction-expansion losses at the shield doors. The turning losses were approximated with losses in the branch of a tee section as shown in Diagram 7-21, Reference 6, and estimated at 1.05 bas ed on the penetration area 10.65 ft

2. The loss at the shield door was approximated with a loss due to a discha rge from a straight conduit through a thick-walled orifice or grid, Diagram 11-28, Reference 6, and calculated at 1.69 based on the penetration exit area 1.424 ft
2. Then the total loss coefficient based on the area 1.424 ft 2 is 1.71, which is the loss coefficient of the junction.

A complete review of all volume and junction parameters as used in the analyses is given in Tables 6.2-9, 6.2-10, 6.2-24, and 6.2-25. Tables of junction characteristics include an indication whether the junction was choked during the analysis. The junctions closer to the break volume choked very early in the transient; an indication that the pressurization was hardly a function of either assigned loss coefficients or inertia coefficients.

Mass and energy blowdown rates used in th e analysis are given in Tables 6.2-26 and 6.2-27.

Figure 6.2-18 depicts the calculated differ ential pressures across the biological shield wall (doors) for the postulated recirculation line break. Figures 6.2-49 and 6.2-50 show final pressure distribution in axial and circumferential direction, respectively also for the recirculation line break. Figures 6.2-22, 6.2-51, and 6.2-52 give the same information for the postulated feedwater line break.

Head Cavity Note: The current flow paths have been changed to include the two manholes between the head cavity and the drywell and the four ducted HVAC vents have been modified by the addition of discharge nozzles. The impact of this change has been evaluated and it has been determined that the analysis presented here is bounding.

LSCS-UFSAR 6.2-41 REV. 13 The computer code utilized for this invest igation was RELAP4/Mod 5 (Reference 7) as received from the Argonne Code Center. A listing of the input for each case (Tables 6.2-15 and 6.2-16) is provided to demonstrate the options of the code that were utilized to obtain a solution. The mass and energy inputs were taken from Table 6.2-18 for the recirculation line brea k, and calculated based on Moody's flow model with a multiplier of 1.0 for the simultaneous head spray line and RPV head vent line break. The details regarding the data contained in Table 6.2-18 are given in Subsection 6.2.1.1.3.1. The basic assumptions utilized in the analysis are given below. a. Thermodynamic equilibrium exists in each containment subcompartment. The containmen t option of the RELAP4/MOD5 computer code was utilized which allows for the flow of air, water vapor, and liquid between the nodes.

b. The constituents of the fluid flowing through the subcompartment vents are based on a homogeneous mixture of the fluid in the subcompartment. The consequences of this assumption result in complete liquid carry-over through subcompartment vents.
c. No heat loss from the gases inside the primary containment is assumed. This adds extra conservatism to the analysis, i.e., the analysis will tend to predict higher containment pressures than would actually exist.
d. Incompressible single-stream flow without momentum flux was used for all junctions.
e. The Moody model for critical flow was used when choking occurred in a junction.
f. The stagnation properties which include dynamic velocity effects were used to determine the flow rate in conjunction with the Moody model.
g. A contraction coefficient of 0.6 was implemented with the junction flow areas which reduces the flow and retains higher pressures closer to the break. In addition, a contraction coefficient of 1.0 was utilized for the fill junction which was used to simulate the break.
h. The reactor pressure vessel head insulation remains in place and retains its structural integrity during any postulated accident. This is conservative since the RPV head cavity volume is minimized which will result in higher pressures in the head cavity.

LSCS-UFSAR 6.2-42 REV. 13 i. The manholes between the head cavity and the drywell are assumed to be closed. This reduces the flow area between the volumes increasing the differential pressure across the bulkhead.

j. All of the HVAC vents (supply and return) are modeled for the postulated break in the head cavity since the pressure in the return vents with the ductwork would alwa ys be greater than the drywell pressure. However, only the supply vents are considered to allow flow for the breaks in the drywell. It is assumed that the HVAC return ductwork would be crushed by the rising drywell pressure.
k. To simplify the input to RELAP4/MOD5, the flow area properties of the HVAC vents are combined into one equivalent vent.
l. The downcomers are represented by an equivalent single flow path with a flow area equal to the sum of the actual flow areas.
m. The modeling of downcomer clearing the initiation of flow into the wetwell was modeled in two ways. In the case of the recirculation line break, the downcomer clearing is extremely rapid. To accurately simulate this, the model would have to be rather complex due to the large inertial and frictional effects present in the downcomer. This complexity was avoided by making use of an accident chronology shown in Table 6.2-7 which found th e vent clearing time to be 0.824 second. A valve was placed in the flow path and opened 0.824 second after the line break. The simultan eous head spray line and RPV head vent line break is a much smaller break and results in a relatively slow pressurization of the drywell. A va lve was again used in the flow path, but in this instance, the valve opening was dependent upon the drywell pressure exceeding the hydrostatic head at the downcomer exit. The opening differential pressure used was 5.2 psid which is equivalent to a 12-foot downcomer submergence.
n. No significant depressurization of the reactor pressure vessel occurs during the postulated break.
o. The simultaneous pipe break of the head spray line and the RPV head vent line was considered because of the lack of whip restraints on the head spray line. The resultant whip of the head spray line is assumed to rupture the RPV head vent line. Neither the RCIC nor the RHR system is operating during the time of the head spray line break, i.e., the RHR-RCIC stop valve is assumed to be closed during the time of the accident. The RPV head vent line is connected at the RPV head and at the main steam header. Therefore, a break in this line results in a two direction blowdown, one side feeds directly from the RPV, and LSCS-UFSAR 6.2-43 REV. 14, APRIL 2002 other feeds from the main steamline. The head spray line has a limiting flow area at the head sp ray nozzle which has a diameter of 3.72 inches. The RPV head vent line is postulated to rupture at the 4-inch to 2-inch reducer in the line located in the head cavity. The steam flow occurs at both ends of the break, one having a diameter of 4.0 inches and the other 2.0 inches. The total flow area was determined to be 0.163 square feet. All of the fl ows are assumed to have the same RPV conditions which are a pressure of 1050.0 psia and an enthalpy of 1190.0 Btu/lbm. Utilizing Moody' s choked flow tables from RELAP4/MOD5, a maximum flow of 2200.0 lbm/sec-ft 2 or 357.9 lbm/sec was calculated. This is used as a constant flow rate for the break in the head cavity.
p. The mass and energy release rates used for the recirculation line break are those given in Table 6.2-18. The break sizes are specified in Subsection 6.2.1.1.3.1.1 and the details regarding line size, break size, orifice size, etc., are given in Table 6.2-4.
q. RELAP4/MOD5 lacks the ability to model steam condensation in the suppression pool. This limitation has no effect on the results obtained prior to vent clearing but will re sult in an overestimation of the pressure rise in the wetwell after vent clearing. Since the maximum differential pressure across the refu eling bulkhead occurs very shortly after downcomer vent clearing in the case of the recirculation line break, the effect is negligible. However, it is noted that the long-term pressure values are not realistic because of this modeling method. In the case of the break in the head cavity, flow through the downcomers does not begin until long after the peak differential pressure across the refueling bulkhead plate occurs.
r. The initial conditions are taken to be the normal operating conditions as given in Table 6.2-3 except with a relative humidity of 0.1%. In the head cavity and drywell the initial pressure is 15.45 psia, the initial temperature is 135 F and the relative humidty is 0.1%. In the wetwell the initial pressure is 15.45 psia, the initial temperature is 100 F and the relative humidity is 0.1%.

The node and flow path data specifics are given in Tables 6.2-11 and 6.2-12 for the simultaneous break of the head spray and RPV head vent lines and Tables 6.2-13 and 6.2-14 for the recirculation line break. The nodes and flow paths are graphically depicted in Figure 6.2-19 for the simultaneous break of the head spray line and RPV head vent line, and Figure 6.2-20 for the recirculation line break.

A description of the loss coefficient determination for the flow paths is provided. This problem has only two flow paths to co nsider. The first path connects the head LSCS-UFSAR 6.2-44 REV. 14, APRIL 2002 cavity to the drywell and consists of eight ports through the bulkhead plate. Four of these ports are the HVAC supply ports for the head cavity and do not have any ductwork attached to them. The remain ing four ports are the HVAC return ducts from the head cavity and have ductwork attached to them. All of the HVAC vents (supply and return) were modeled for the po stulated break in the head cavity since the pressure in the return vents with th e ductwork would always be greater than the drywell pressure. The losses considered were the turning losses of the fluid around the RPV head from the break to the HVAC ports in the bulkhead. These losses are very small since the turning radius around the RPV head is so large.

Therefore, this loss was neglected. The ports without the ductwork were considered as thick-edged orifices. This loss coefficient was determined using Diagram 4-14 of Reference 6 and was calculated to be 1.52. The ports with the ductwork consist of a 24-inch to 18-inch diameter reducer followed by ductwork which includes a series of elbows and one tee. The fl ow finally exits into the dr ywell through one of the tee branches. Diagrams 3-9, 6-1, and 7-25 of Reference 6 were used to calculate the loss coefficient and it was determined to be 4.62. Since the flow through the ports with and without ductwork is parallel, the losses were combined for parallel flow and the total loss coefficient was calcul ated, as described in Subsecti on 6.2.1.2.3, to be 2.62. The flow area for this case is the total of the minimum flow areas through each of the eight HVAC vents. The total flow area was determined to be 11.12 square feet. For the recirculation line break within the drywell, only the supply vents which are without ductwork were considered to allo w for flow. It is assumed that the HVAC return ductwork would crush because the drywell pressure would be greater than the pressure in the ductwork. The loss coefficient for this case is calculated for the ports without the ductwork. The loss coefficient was determined as mentioned earlier and was calculated to be 1.52. The flow area for this case was determined to be 4.92 square feet.

The loss coefficient for the second flow path, through the downcomers, was taken from Table 6.2-1 and is 5.2. No attempt was made to model the inertial effects of the clearing transient. The path was treated as a valve that opened at a prespecified time of 0.824 second for the recirculation line break. For the simultaneous head spray line and RPV head vent line break, the path was treated as a valve that opened when the drywell pr essure exceeded the hydrostatic head of 5.2 psid which is equivalent to a 12-foot downcomer submergence. The path model considers no inertial effects; this is a cons ervative approach, since it has the effect of making the pressure differentials across the bulkhead plate higher.

Figure 6.2-24 depicts the pressure histories of the head cavity and drywell for the break in the head cavity and the recirculati on line break in the head cavity and the recirculation line break in the drywell. The pressure differential histories across the bulkhead plate for the break in the head cavity and the recirculation line break in the drywell are shown in Figure 6.2-25. The peak pressure differential for each break was found to be 9.0 psid upward fo r the recirculation line break and 7.0 psid downward for the simultaneous head spray line and RPV head vent line break. The LSCS-UFSAR 6.2-45 REV. 14, APRIL 2002 differential pressure history as shown for the simultaneous break of the head spray line and RPV head vent line shows two differential pressure peaks. The first differential pressure peak is due to the su dden pressurization of the head cavity and the second peak is due to the sudden op ening of the downcomers at a pressure differential between the drywell and wetwell of 5.2 psid. This second peak is erroneous because no inertial effects were modelled in the downcomer flow path and therefore was not considered as the design downward differential pressure. The design pressure differential is 10.6 psid in both directions. This provides for a margin factor of approximately 1.2 at the final design stage.

6.2.1.2.4 Impact of Increased Initial Suppression Pool Temperature Supplementary safety evaluations have been performed, as discussed in Section 6.2.1.8, to verify that an increase in the initial suppression pool temperature would not significantly impact the consequences of this accident scenario.

6.2.1.3 Mass and Energy Rele ase Analyses for Postulated Loss-of-Coolant Accidents This section contains a description of th e transient energy release rates from the reactor primary system to the containment system following a LOCA with minimum ESF performance. In general, a very conservative analytical approach is taken in that all possible sources of energy are accounted for, whereas the suppression pool is assumed to be the only available heat sink. No credit is taken for either the heat that will be stored in the suppression chamber and drywell structures, or the heat that will be transmitted through the containment and dissipated to the environment.

The analysis at 3559 MWt used essentially th e same methodology as the analysis at 3434 MWt, except for the RPV blowdown in the short-term containment response analysis, as noted in Subsecti on 6.2.1.1.3. The break flow rate and enthalpy used for the short-term containment response analysis at 3559 MWt are given in Table 6.2-18a. For the analysis of the long-t erm containment response, one of the key input assumptions updated for the analysis at 3559 MWt is that the core decay heat is based on the ANSI/ANS 5.1-1979 decay he at model with a two sigma uncertainty adder. The core decay heat values used in the 3559 MWt analysis are provided in Table 6.2-20a. The following subsections explain how the transient mass and release rates from the vessel to the containment were determined for the original analysis at 3434 MWt.

6.2.1.3.1 Mass and Energy Release Data Table 6.2-18 provides the mass and enthalpy release data for the containment DBA, recirculation line break. Blowdown steam an d liquid flow rates and their respective enthalpies are reported for a 24-hour period following the accident. Figures 6.2-26 LSCS-UFSAR 6.2-45a REV. 14, APRIL 2002 and 6.2-27 show the blowdown flow rates for the recirculation lines break graphically. This data was employed in the DBA containment pressure-temperature transient analyses repo rted in Subsection 6.2.1.1.3.1.

Table 6.2-19 provides the mass and enthalpy release data for the main steamline break. Blowdown data is presented for a 24-hour period following the accident. Figure 6.2-28 shows the vessel blowdown flow rates for the main steamline break as a function of time after the postulated rupture. This information has been employed in the containment response analys es presented in Subsection 6.2.1.1.3.1.

LSCS-UFSAR 6.2-46 REV. 13 6.2.1.3.2 Energy Sources The reactor coolant system conditions prior to the design basis recirculation line break are presented in Tables 6.2-3 and 6.

2-4. Reactor blowdown calculations for containment response analyses are based upon these conditions during a loss-of-coolant accident.

Following each postulated accident event, the stored energy in the reactor system and the energy generated by fission product decay will be released. The rate of release of core decay heat for the evaluation of the containment response to a LOCA is provided in Table 6.2-20 as a function of time after accident initiation. This data is based upon a normalization factor of 3440 MWt and includes the energy of fuel relaxation.

Following a LOCA, the sensible energy stored in the reactor primary system metal will be transferred to the recirculating ECCS water and will thus contribute to the suppression pool and containment heatup. Figure 6.2-29 shows the temperature transients of the various primary system structures which contribute to this sensible energy transfer. Figure 6.2-30 shows the variation of the sensible heat content of the reactor vessel and internal structures during a recirculation line break accident based upon the temperature transient responses.

6.2.1.3.3 Effects of Metal-Water Reaction The containment systems shall accommodate the effects of metal-water reactions and other chemical reactions following a postulated DBA. The amount of metal-water reaction is limited to values consiste nt with the performance objectives of the emergency core cooling systems (ECCS).

6.2.1.3.4 Impact of Increased Initial Suppression Pool Temperature Supplementary safety evaluations have been performed, as discussed in Section 6.2.1.8, to verify that an increase in the initial suppression pool temperature would not significantly impact the consequences of this accident

scenario.

6.2.1.4 Mass and Energy Release Analysis for Postulated Secondary System Pipe Ruptures Inside Containment (PWR)

Not applicable.

6.2.1.5 Minimum Containment Pressure Analysis for Performance Capability Studies on Emergency Core Cooling System (PWR)

Not applicable.

LSCS-UFSAR 6.2-47 REV. 17, APRIL 2008 6.2.1.6 Testing and Inspection Containment testing and inspection programs are fully described in Subsection 6.2.6 and in Chapter 14.0 of the FSAR. The requirements and bases for acceptability are outlined completely in the Technical Specifications.

6.2.1.7 Instrumentation Requirements A complete description of the instrumentation employed for monitoring the containment conditions and actuating those systems and components having a safety function is presented in Chapter 7.0.

6.2.1.8 Evaluation of 105 F Suppression Pool Initial Temperature Temperature limits on the suppression pool for Boiling Water Reactors (BWR) with Mark II containment were implemented to minimize the potential for high amplitude loads on the pool during accide nt events. However, some of the limits were implemented with excessive conservatism because the loading phenomena were not completely understood. This suppression pool temperature limit has therefore been historically chosen based on the maximum expected service water temperature. For LaSalle County Station Units 1 and 2, the licensing safety evaluations were based upon a 100 F suppression pool water temperature, which was equivalent to the Ultimate Heat Sink design temperature limit.

Hot weather in Illinois can cause the temperature of the ultimate heat sink to rise to the point where the suppression pool temperature limit of 100 F may be exceeded. However, the ultimate heat sink design limit will not be exceeded. To prevent an unnecessary plant sh utdown during a period of high electrical demand, plant specific safety evaluations have been performed (References 10-20) to demonstrate that plant operation with higher suppression pool temperature is acceptable, i.e., the plant safety limits will still be met with the higher temperatures.

The suppression pool was designed to function as both a heat sink and an emergency water source during transient and accident events as discussed throughout section 6.2. Therefore, performance of the following evaluations were required to support a 5 F increase in the initial suppression pool temperature as LaSalle County Station Units 1 and 2:

a) Containment loads associated with SRV operation including air clearing loads and steam condensation loads.

b) Containment response associated with LOCA events including the peak pressure and temperature design limits, condensation capability, condensation oscillation load s (CO), and chugging loads.

LSCS-UFSAR 6.2-48 REV. 20, APRIL 2014 c) Equipment performance for design basis events including the impact on the core cooling capability of the ECCS and the parameters which could impact the operability of the ECCS pumps (such as NPSH availability, etc.).

d) Equipment and ECCS performance for other non-LOCA events, e.g., ATWS. For each of these cases the evaluation showed that the increase of the initial suppression pool temperature would have an insignificant impact on the existing design margin for the suppression pool and ECC systems. Peak local pool temperature will increase by 3 F at a 105 F initial pool bulk temperature for SRV related events.*

The results of this evaluation were subm itted to the NRC (Reference 11), and an approved license amendment to change the maximum suppression pool temperature limit to 105 F was received (Reference 12). The Ultimate Heat Sink design temperature limit is changed to 107 F in Reference 36.

6.2.2 Containment Heat Removal System The containment heat removal system func tion is accomplished by the containment cooling mode of the RHR system. The system is also equipped with spray headers in the drywell and suppression chamber areas. However, no credit was taken for these spray headers for either heat removal or fission product control following a LOCA. 6.2.2.1 Design Bases The containment heat removal system, consisting of the suppre ssion pool cooling system, is an integral part of the RHR syst em. It meets the following safety design bases: a. The source of water for restoring RPV coolant inventory is located within the containment to establish a closed cooling-water path.

b. A closed loop flow path between the suppression pool and the RHR heat exchangers is established so that the heat removal capability of these heat exchangers can be utilized.
c. This system, in conjunction with the ECC systems, has such diversity and redundancy that no single failure can result in its inability to cool the core adequately (Subsection 6.3.1).
  • Peak bulk suppression pool temperature, in the case of LOCA events, is still approximately 10 F below the allowable values.

LSCS-UFSAR 6.2-49 REV. 13 d. To ensure that the RHR contai nment cooling subsystem operates satisfactorily following a LOCA, each active component shall be testable during operation of the nuclear system.

6.2.2.2 System Design

The containment cooling subsystem is an integral part of the RHR system, as described in Subsection 5.4.7. The piping and instrumentation diagram is given in Drawing Nos. M-96 (sheets 1-4) and M-142 (sheets 1-4). Re dundancy is achieved by having two complete containment cooling systems.

Consideration of the fouling of heat exch angers and the selection of temperatures for heat exchanger design are di scussed in Subsection 5.4.7.

6.2.2.3 Design Evaluation In the event of the postulated LOCA, the short-term energy release from the reactor primary system will be dumped to the suppression pool. This will cause a pool temperature rise of approximately 46 F. Subsequent to the accident, fission product decay heat will result in a continui ng energy dump to the pool. Unless this energy is removed from the primary containment system, it will eventually result in unacceptable suppression pool temperatures and containment pressures. The containment cooling mode of the RHR system is used to remove heat from the suppression pool.

A supplementary evaluation has been performed for the addition of feedwater to the suppression pool to assess the impact on long term pool temperature. This evaluation estimates that the peak short term pool temperature will increase by an additional 15.4 F. This results in a short term pool temperature (at 600 seconds) of approximately 166 F. Further details are given in Section 6.2.1.1.3.1.1 in the paragraph titled, "Evaluation of Post-LOCA Feedwater Injection".

6.2.2.3.1 RHR Containment Cooling Mode When the RHR system is in the containment cooling mode, the pumps draw water from the suppression pool, pass it throug h the RHR heat exchangers, and inject it back either to the suppression pool or to the RPV.

In order to evaluate the adequacy of the RHR system, the following limiting case is postulated:

a. Reactor initially at maximum power.
b. Isolation scram occurs.

LSCS-UFSAR 6.2-50 REV. 17, APRIL 2008

c. Manual depressurization discharges heat to suppression pool.
d. Suppression pool cooling is established approximately 10 minutes after the technical specification limit for pool water temperature is reached.

A complete discussion of the suppression pool temperature transients is contained in Chapter 6 of the LSCS-DAR.

The suppression pool temperature transients have been analyzed based on an increased initial suppression pool temperature of 105 F as discussed in Section 6.2.1.8. The scenarios analyzed are based on those spec ified in NUREG-0783, Reference 15 provides the results of this analysis. For all analyzed cases the long term suppression pool temperature is less than 200 F. 6.2.2.3.2 Summary of Co ntainment Cooling Analysis When calculating the long term, post LOCA pool temperature transient, it is assumed that one RHR heat exchanger loop is not available, the suppression pool level initially is at the technical specification minimum, the suppression pool temperature initially is at the technica l specification maximum, and the design RHR heat exchanger fouling factors are used. No credit is taken for heat loss to

environs or to the pool structures.

The resultant suppression pool transient maximum te mperature for 3434 MWt is 200 F (see References 8, 15, 16, 17, and 18). It is concluded that even with the very conservative assumptions described above, the RHR system in the containment cooling mode can meet its design objectiv e of safely terminating the limiting case temperature transient. See subsection 6.2.

2.3.5 for impact of power uprate to 3489 MWt.

6.2.2.3.3 Impact of Increased Initial Suppression Pool Temperature Supplementary evaluations have been performe d, as discussed in Section 6.2.1.8, to verify that an increase in the initial suppression pool temperature would not impact the ability of the RHR containment cooling system to meet its design objective.

6.2.2.3.4 Impact of Re duced RHR Suppression Pool Cooling Flow Rate The original and 1988 General Electric co ntainment analyses (references 8 & 17), has been supplemented with an evaluation which considers an RHR pump flow rate during the suppression pool cooling of 7200 gpm. As noted in Table 6.2-2, the previous analysis used a flow rate of 7450 gpm. Although the RHR pump is capable of such performance, the minimum requ ired Technical Specification flow per specification SR 3.6.2.3.2 is only 7200 gpm.

Since suppression pool cooling is only initiated after 600 seconds into the DBA-LOCA, the affect of this lower flow rate LSCS-UFSAR 6.2-51 REV. 18, APRIL 2010 will be seen as slightly lower efficiency for the RHR heat exchanger and a higher long term suppression pool temperature.

The results of the Reference 18 General Electric analysis indicate an increase in the long term pool temperature of 1.5 F for the DBA-LOCA case.

For cases which involve SRV blowdown to the suppression p ool the lower RHR pump flow rate was asse ssed in S&L Calculat ion 3C7-0181-003, Rev. 3 (Reference 15) and the effect on the peak suppression pool temperature was an increase of less than or equal to 1 F in the peak suppression pool temperature. For all cases examined, the highest peak pool temperature calculated is 195 F which is still less than 200 F peak temperature for all cases analyzed. Thus, complete steam condensation is assured with these elevated pool temperatures.

6.2.2.3.5 Impact of Power Uprate The resultant post-LOCA maximum suppre ssion pool temperature at 102% of uprated reactor thermal power, 3559 MWt, is 196.1º F, as shown in Table 6.2-5a.

The resultant maximum long-term post DBA-LOCA suppression pool temperature with the concerns of SC06-01 addressed in 197 F as shown in Table 6.2-5A. The maximum suppression pool temperature at 3559 MWt for NUREG-0783 events is 190.7º F as evaluated in Reference 31.

The suppression pool limit for events with SRV discharge is evaluated in References 25 and 27. In the NRC's Safety Evaluation of Reference 28 for the elimination of local suppression pool temperature limits for plants with T-Quenchers, an additional concern was raised on the pote ntial transfer of non-condensed SRV steam plumes to ECCS suction strainers. An an alysis was performed in Reference 29 that modeled the steam plume formation, determined the extent of steam plume projection, and verified that the plume can not enter ECCS suction strainers.

However, the analysis determined the ex istence of a potential steam ingestion concern for the "K" SRV and the Reactor Core Isolation Cooling (RCIC) suction strainer, if the temperature of the suppressi on pool is above 200º F. Administrative controls have been implemented to caution the operators on the use of "K" SRV and RCIC simultaneously when the suppression pool temperature is above 200º F.

6.2.2.3.6 Sensitivity of Initiation Time of RHR Containment Cooling Mode A one-time sensitivity analysis was performed to determine the impact on the peak suppression pool temperature, if the star t of the RHR Containment Cooling Mode is delayed for longer than 10 minutes, fo llowing a DBA-LOCA. Manual operator action from the main control room is needed, in order for Suppression pool cooling to be initiated. These actions could require up to a few minutes to accomplish (accounting for valve stroke times, etc.). The impact on peak suppression pool temperature was studied if the start of suppression pool cooling is delayed from 10 minutes to 30 minutes.

LSCS-UFSAR 6.2-51a REV. 18, APRIL 2010 The study utilized power uprate decay heat loads. The results of this study indicate there is a very small impact on peak suppression pool temperature. The 30 minute case results in an increase of 2.0 deg-F, added to the current analysis peak of 197 deg-F, results in a postulated peak temper ature of 199 deg-F. This peak temperature does not challenge the suppression pool design limits. The operator actions to re-align RHR are anticipated to require much less time than the additional 20 minutes of this analysis. The increase in peak suppression pool temperature is concluded to be negligible (i.e. less than 1 deg-F) for these anticipated starting times which are only a few minutes longer than 10 minutes.

6.2.2.4 Test and Inspections The operational testing and the periodic inspection of components of the containment heat removal system are descri bed in Subsection 5.4.7.4.

6.2.2.5 Instrumentation Requirements Suppression pool cooling by the RHR system is manually initiated from the control room where sufficient instrumentatio n is provided for that purpose.

6.2.3 Secondary Containment Functional Design The Secondary Containment consists of the Reactor Building, the equipment access

structure, and a portion of the main steam tunnel and has a minimum free volume of 2,875,000 cubic feet.

The reactor building completely encloses th e reactor and its primary containment.

The structure provides secondary containment when the primary containment is closed and in service, and primary containment when the primary containment is open, as it is during the refueling period. The reactor building houses the refueling and reactor servicing equipment, the new and spent fuel storage facilities, and other reactor auxiliary or service equipment, including the reactor core isolation cooling system, reactor water cleanup demineralizer system, standby liquid control system, control rod drive system equipment, the emergency core cooling system, and electrical equipment components.

6.2.3.1 Design Bases The functional capability of the ventilation system to maintain negative pressure in the secondary containment with respect to outdoors is discussed in Subsection 9.4.2.

6.2.3.2 System Design The reactor building is designed and constructed in accordance with the design criteria outlined in Chapter 3.0. The reactor buildin g exterior walls and superstructure up to the refueling floor are constructed of reinforced concrete.

LSCS-UFSAR 6.2-52 REV. 15, APRIL 2004 Above the level of the refueling floor, the building structure is fabricated of structural steel members, insulated siding and a metal roof. Joints in the superstructure paneling are detailed to assure leaktightness.

Penetrations of the reactor building are designed with leakage characteristics cons istent with leakage requirements of the entire building. The reactor building is de signed to limit the inleakage to 100% of the reactor building free volume per day at a negative interior pressure of 0.25 inch H 2 0 gauge, while operating the standby gas tr eatment system. The building structure above the refueling floor is also designed to contain a negative interior pressure of 0.25 inch H 2 0 gauge. Personnel entrance to the reactor building is through an interlocking double door airlock. Rail car access openings in the reactor building at elevation 710 feet 6 inches provided with double doors to assure that building access will not interfere with maintaining integrity of the secondary containment.

Ventilation for the reactor building is provided by means of a once-through ventilation system. Outdoor air is filtered then evaporatively or chilled glycol cooled to *reduce the supply air dry bulb temperature to increase the sensible cooling capacity of this air. This air is then preheated as required to satisfy the plant operating conditions.

The equipment is arranged as follows: outs ide air inlet, filter, chilled glycol/heating coil evaporative *cooler (abandoned-in-place), resistive heating coils, and supply fans. Three 50% vane axial fans are provided, two of which normally operate and one which serves as a standby.

Supply air is distributed to the reactor building by means of a duct system to provide

equipment cooling in various areas as requir ed. Air is routed from clean areas to areas with progressively greater contamination potential. Pressure differential control dampers are used as required to maintain negative pressures in potentially contaminated cubicles. All exhaust air is ro uted through a return duct system to the exhaust fans.

All supply air delivered to the refueling floor level is exhausted from the periphery of the spent fuel and equipment storage pools and the reactor well. This air is routed directly to the main system exhaust duct.

Three vane axial exhaust fans are provided, two of which normally operate and one of which serves as a standby. The discharge from the exhaust fans is routed to the plant vent where the air is discharged to the atmosphere. All exhaust air is monitored for radiation.

Normal ventilation systems are not required to operate during accident conditions and are automatically shut down whenever the standby gas treatment system starts. The equipment for this system is not powered from essential buses. To

  • Note: The evaporative coolers are abandoned-in-place.

LSCS-UFSAR 6.2-53 REV. 13 maintain the integrity of the secondary containment, two isolation dampers are provided in the supply air duct between the supply fan discharge and the penetration through the secondary containment wall.

The secondary containment structure protects the equipment in the building from externally generated missiles. Piping syst ems within the secondary containment have been analyzed for high energy pipe breaks outside primary containment and pipe whip restraints are provided as required. The effects of jet impingment have also been analyzed and included in the design of the structure and pipe whip restraints. For more information on high energy pipe breaks outside primary containment see Appendix C.

The isolation features and isolation signals for secondary containment are discussed in Section 6.5, Chapter 7.0 and Subsection 9.4.2.

6.2.3.3 Design Evaluation The design evaluation of secondary containment ventilation system and atmospheric cleanup system is given in Sect ion 6.5 and Subsection 9.4.2.

6.2.3.4 Test and Inspections The program for initial performance testing is outlined in the Technical Specifications. Periodic functional testing of the second ary containment and secondary containment isolation system is described in the Technical Specifications.

6.2.3.5 Instrumentation Requirements The instrumentation to be employed for the monitoring and actuation of the standby gas treatment system is fully described in Chapter 7.0.

The instrumentation used for the monitoring and actuation of the ventilation and cleanup system is discussed in Subsections 7.3.8 and 7.6.1.2.

6.2.4 Containment Isolation System The primary objective of the containment is olation system is to provide protection against the release of radioactive materials to the environment through the fluid system lines penetrating the containment.

This objective is accomplished by ensuring that isolation barriers are provided in all fluid lines that penetrate primary containment, and that automatic closure of the appropriate isolation valves occurs.

LSCS-UFSAR 6.2-54 REV. 13 6.2.4.1 Design Bases The design requirements for containment isolation barriers are:

a. The capability of closure or isolation of pipes or ducts that penetrate the containment is provided to ensure a containment barrier sufficient to

maintain leakage within permissible limits.

b. The arrangements of isolation valving and the criteria used to establish the isolation provisions conform to the requirements of General Design Criteria 54 through 57, as discussed in Section 3.1.
c. The design of all containment isolation valves and associated piping and penetrations is Seismic Category I.
d. Containment isolation valves and associated piping and penetrations meet the requirements of the ASME Boiler and Pressure Vessel Code,Section III, for Class 1 or 2 components, as applicable.
e. Isolation valves, actuators, and co ntrols are protected against loss of safety function from missiles and accident environments.
f. Containment isolation valves prov ide the necessary isolation of the containment in the event of accidents or other conditions to limit the untreated release of radioactive materi als from the containment in excess of the design limits.
g. Appropriate isolation valves are auto matically closed by the signals listed in Table 6.2-21. The criteria for assigning isolation signals to their associated isolation valves is described in Subsection 7.3.2. Once the isolation function is initiated, it goes to completion.
h. Redundancy and physical separation are required in the electrical and mechanical design to ensure that no single failure in the system prevents the system from performi ng its safety function.

The governing conditions under which cont ainment isolation becomes mandatory are high drywell pressure or low water level in the reactor vessel. One or both of these signals initiate closure of isolation valves not required for emergency shutdown of the plant. These same signals also initiate the ECCS. The valves associated with an ECCS may be closed remote manually from th e control room or cl ose automatically, as appropriate.

Excess flow check valves are used as a means of automatic isolation on all static instrument sensing lines that penetrate the drywell containment and connect to LSCS-UFSAR 6.2-55 REV. 16, APRIL 2006 either the reactor pressure boundary or the drywell atmosphere. The valve is located downstream of the root valve and as close as practical to the outside surface of the containment. This valve is automatically closed to restrict flow in case of a sensing line break outside containment.

Backfill Injection lines have been added to the reference legs originating from Condensing Chambers 1(2) B21-D004A/B/C/D to comply with NRC Bulletin 93-03.

These lines use two simple check valves in series to accomplish the outboard containment isolation function. It is acceptable to use the two simple check valves instead of one excess flow check valve for the backfill injection lines because these lines would not need the built-in bleed flow path in an excess flow check valve to reopen when appropriate. The 4 lbs./hr. CRD flow would reopen the check valves when it is available. If it is not availabl e, it is not appropriate to reopen the check valves. This meets the Regulatory Guide 1.11 "... the valve should reopen automatically or be capable of being reopened readily under the conditions that prevail when reopening is appropriate. It should not be necessary to break a line to reopen a closed valve." In addition, there is no instrument reading that will be significantly effected by the closure of these check valves.

Dead-end instrument sensing lines that are in communication with the reactor pressure boundary and penetrate the primary containment are equipped with 1/4 inch orifice as close to the proces s as possible inside the drywell.

6.2.4.2 System Design Table 6.2-21 presents the design informat ion regarding the containment isolation provisions for fluid system lines and instrument lines penetrating the containment. Containment isolation signals are identified in Table 6.2-21 and valve

arrangements are represented in Figure 6.2-31.

The plant protection system signals that initiate closure of the containment isolation valves are listed in Table 7.3-2.

The isolation provisions follow the requirements of General Design Criteria 54, 55, 56, and 57. General Design Criteria 54 applies to all of the containment isolation valves. Compliance with General Design Crit eria 55, 56, and 57 is described below. The justification for this design is also presented.

6.2.4.2.1 Evaluation Agains t General Design Criterion 55

Feedwater Line Each feedwater line forming a part of the reactor coolant pressure boundary is provided with a swing type check valve on Unit 1 and a swing type check valve on Unit 2 inside the containment, and a nonslam type, air operated testable check valve outside the containment, as close as LSCS-UFSAR 6.2-56 REV. 14, APRIL 2002 practicable to the containment wall. In addition, a motor-operated gate valve is installed upstream of the outside isolat ion valve to provide long-term isolation capability.

During a postulated LOCA, it is desirable to maintain reactor coolant makeup from all available sources. Therefore, it would not improve safety to install a feedwater isolation valve that closed automatically on signals indicating a LOCA, and, thereby, eliminate a source of reactor makeup. The provision of the check valves, however, ensure the prevention of a significant lo ss of reactor coolant inventory and offer immediate isolation should a break occur in the feedwater line. For this reason, the outermost valve does not automatically is olate upon signal from the protection system. The valve is remote manually closed from the main control room to provide long-term leakage protection upon operat or determination that continued makeup from the feedwater system is unavailable or unnecessary.

In addition, the outboard check valve is provided with a special actuator that performs the following functions:

a. The actuator is capable of partially moving the valve disc into the flow stream during normal plant operation in order to ensure that the valve is not bound in the open position. The actuator is not capable of fully closing the valve against flow, however, and there is no significant disruption of feedwater flow.
b. The actuator is capable of applying a seating force to the valve at low differential pressures and abnormal conditions. This improves the leaktightness capability of the valves. The actuator will be utilized during leak testing.

ECCS Lines to the RPV The subject penetration(s) meet the alternate primary containment isolation criteria of NUREG 0800 "Standard Review Plan for the review of Safety Analysis Reports for Nuclear Power Plants" (SRP) instead of the explicit requirements of GDC 55.

The HPCS, LPCS, and LPCI lines penetrate the drywell and inject coolant directly into the reactor pressure vessel. Isolation is provided on each of these lines by a normally closed check valve inside the containment and a normally closed motor-operated gate valve located outside the cont ainment, as close as practicable to the exterior wall of the containment. If a loss-of-coolant accident occurred, each of these valves would be required to open to suppl y coolant to the RPV. The motor-operated gate valves are automatically opened by their appropriate signals, and the check valves are opened by the coolant flow in th e line. The opening capability of the check valve can be tested by monitoring flow through the valve into the reactor vessel.

LSCS-UFSAR 6.2-57 REV. 16, APRIL 2006 Control Rod Drive Lines The control rod drive system, has two type s of lines to the RPV; the insert and withdraw lines that penetrate the drywell and connect to the control rod drive.

The control rod drive insert and withdraw lines can be isolated by the solenoid valves outside the primary containment.

These lines that extend outside the primary containment are small, and termin ate in a system that is designed to prevent out-leakage. Solenoid valves normally are closed, but open on rod movement and during reactor scram. In addition, a ball check valve located in the control rod drive flange housing automatically seals the insert line in the event of a break. RHR and RCIC Head Spray Lines The subject penetration(s) meet the al ternative primary containment isolation criteria of NUREG 0800 "Standard Review Plan for the review of Safety Analysis Reports for Nuclear Power Plants" (SRP) in stead of the explicit requirements of GDC 55. The RHR and RCIC head spray lines meet outside the containment to form a common line which penetrates the drywell and discharges directly into the reactor pressure vessel. The testable check valve inside the drywell is normally closed. The testable check valve is located as close as practicable to the reactor pressure vessel.

Three types of valves, a testable check valve, a normally closed motor-operated remote manual gate valve, and a normally closed motor-operated automatic globe valve, are located outside the containmen

t. The check valve assures immediate isolation of the containment in the event of a line break. The globe valve on the RHR line receives an automatic isolation signal while the gate valve on the RCIC line is remote manually actuated to provide long-term leakage control.

Standby Liquid Control System Lines The standby liquid control system line penetrates the drywell and connects to the reactor pressure vessel. In addition to a simple check valve inside the drywell, a check valve together with an explosive actuated valve are located outside the drywell. Since the standby liquid control li ne is a normally closed, nonflowing line, rupture of this line is extremely remote. The explosive actuated valve, though, functions as a third isolation valve. This valve provides an absolute seal for long-term leakage control as well as preventing leakage of sodium pentaborate into the reactor pressure vessel during normal reactor operation.

LSCS-UFSAR 6.2-57a REV. 14, APRIL 2002 Reactor Water Cleanup System The reactor water cleanup (RWCU) pumps, heat exchangers, and filter demineralizers are located outside the primary containment. The return line from the filter demineralizers connects to the feedwater line outside the containment between the outside containment feedwate r check valve and the outboard motor-operated gate valve. Isolation of this lin e is provided by the feedwater system check LSCS-UFSAR 6.2-58 REV. 14, APRIL 2002 valve inside the containment, the feedwater check valve outside the containment, and a motor-operated gate valve which provides a long term isolation capability.

During the postulated loss-of-coolant accide nt, it is desirable to maintain reactor coolant makeup. For this reason, valves wh ich automatically isolate upon signal are not included in the design of the system. Consequently, a third valve is required to provide long-term leakage control. Should a break occur in the reactor water cleanup return line, the check valves would prevent significant loss of inventory and offer immediate isolation, while the outermost isolation valve would provide long-term leakage control.

Recirculation Pump Seal Water Supply Line

The recirculation pump seal water line extends from the recirculation pump through the drywell and connects to the CRD supply line outside the primary containment. The seal water line forms a part of the reactor coolant pressure boundary, therefore the consequences of failing this line have been evaluated. This eval uation shows that the consequences of breaking this line is less severe than that of failing an instrument line. The recirculation pump seal water line is 3/4-inch Class B from the recirculation pump through the second check valve (located outs ide the containment). From this valve to the CRD connection the line is Class D. Sh ould this line be postulated to fail and either one of the check valves is assumed not to close (single active failure), the flow rate through the broken line has been calcul ated to be substantially less than that permitted for a broken instrument line. Th erefore, the two check valves in series provide sufficient isolation capability for postulated failure of this line.

RHR Shutdown Cooling Return Line The subject penetration(s) meet the alternative primary containment isolation criteria of NUREG 0800 "Standard Review Plan for the review of Safety Analysis Reports for Nuclear Power Plants" (SRP) instead of the explicit requirements of GDC 55.

The shutdown cooling return lines are connected to the reactor recirculation pump discharge lines. The isolation valve arrangement on these lines is identical to that on the ECCS lines connected to the RPV. Ho wever, the motor-operated valve outside containment closes automatically upon receipt of an isolation signal.

RHR Shutdown Cooling Suction Line The penetration (M-7) has been protected by a relief valve mounted between the inboard automatic isolation and the containm ent penetration. This relief valve was added in response to NRC Generic Letter GL 96-06 concerns for isolated line overpressurization during a LOCA. Because the RHR Shutdown Cooling piping up to and including the outer containment penetration automatic isolation valve is part of the RCPB, the penetration configuration must meet GDC 55.

LSCS-UFSAR 6.2-59 REV. 13 Reactor Recirculation System Sample Line The Reactor Recirculation sample line is a 3/4" line that is an extension of the RCPB to the outboard isolation valve. The containment penetration (M-36) has an automatic isolation inside containment and an automatic isolation outside containment. A 3/4" bypass line with a check valve has been added around the inboard isolation valve in response to Generic Letter 96-06. The check valve will open to relieve penetration overpressurization following a LOCA. Manual valves between the check valve and the RR 24" process line will be maintained locked open, when required for overpressure protection, to assure a vent path for overpressure protection.

The two automatic valves and the inboard check valve meet the requirements of GDC 55. 6.2.4.2.2 Evaluation Agains t General Design Criterion 56 Primary Containment Chilled Water System

The Primary Containment Chilled Water System (PCCW) consists of two independent trains of cooling for the primary containment atmosphere. Each train penetrates the containment with a supply and return line. Each line has an inboard and an outboard automatic isolatio n valve. Each penetration (M-25, M-27, M-28, M-26) has been protected by a relief valve mounted between the inboard automatic isolation and the containment pe netration. These relief valves were added in response to NRC Generic Letter GL 96-06 concerns for isolated line overpressurization during a LOCA.

The penetration configuration must meet GDC 56.

RCIC Turbine Exhaust Vacuum Breaker Line Minimum Flow Bypass

The RCIC turbine exhaust line is provided with a vacuum breaker system to prevent condensation of the exhaust steam from inducing a vacuum in the line. The vacuum relief line connects the turbine exhaust line to the suppression chamber atmosphere. Two check valves in-series in the line prevent stea m from exhausting to the vapor space above the pool, and two motor-operated globe valves, one on either side of the aforementioned check valves, provide remote manual isolation capability for the RCIC turbine exhaust vacuum breaker line.

Combustible Gas Control and Post-LOCA Atmosphere Sampling Lines The post-LOCA sampling system lines which penetrate the containment and connect to the drywell and suppression chamber air volume are each equipped with LSCS-UFSAR 6.2-60 REV. 13 a single divisional fail-open, solenoid operated isolation valve located outside and as close to the containment as possible. The combustible gas control system lines which penetrate the containment are equipped with two normally closed motor-operated valves in series, located outside containment, remote manually actuated from the control room. These valves prov ide assurance of isolating these lines in the event of a break and also provide long-term leakage control. In addition, the piping is considered an extension of containment boundary since it must be available for long-term usage following a de sign basis loss-of-coolant accident, and, as such, is designed to the same quality standards as the primary containment.

Thus, the need for isolation is conditional.

Containment Vent and Purge and Containment Drain Lines The drywell and suppression chamber vent and purge and containment drain lines have test isolation capabilities commensurate with the importance to safety of isolating these lines. Each line has two normally closed, instrument air powered, air cylinder actuated valves located outside the primary containment. The air cylinders are operated by solenoid valves connected to the control logic. Containment isolation requirements are met on the basis that the purge and drain lines are normally closed, low-pressure lines constructed to the same quality standards as the containment and meet th e Branch Technical Position CSB 6-4.

These isolation valves are interlocked to preclude opening of the valves while a containment isolation signal exists. Furt hermore, the consequences of a break in these lines result in no significant safety consideration.

Drywell and Suppression Chamber Air Sampling Lines The air sampling lines are used for continuously drawing containment air during normal operation as part of the leak detection system. These lines are equipped with two normally open, solenoid operated, spring to close valves in series, located outside and as close as possible to the containment. This manner of routing the system piping reduces the number of containment penetrations and minimizes the potential pathways for radioactive material release. In addition, the piping

upstream of the air sampling isolation valv es is considered an extension of the containment since it must be available fo r long-term usage following a design basis loss-of-coolant accident. The piping is pa rt of the post-LOCA atmosphere sampling system, and as such, is designed and fabr icated to the same quality standards as the containment. Containment isolation requirements are met on the basis that these lines are low-pressure lines constructed to the same quality standards as the containment furthermore, the consequences of a break in these lines result in no significant safety consideration.

LSCS-UFSAR 6.2-61 REV. 13 Service Air and Clean Condensate Supply Lines The Service Air and Clean Condensate supply lines, which penetrate the containment, provide air and water servic e connectors inside the drywell during reactor shutdown and outages. These lines are equipped with two manually operated valves which are locked closed during reactor operations. In addition, each line is equipped with a spool piece which is removed and respective blank flanges installed during reactor operations. The va lves and spool pieces are located outside of and as close as possible to the contai nment. This manner of routing the system piping reduces the number of containment penetrations. Since these lines are isolated during reactor operations, the potential pathways for radioactive material release is minimized. Furthermore, the consequences of a break in these lines result in no significant safety consideration.

Reactor Building Closed Cooling Water System The Reactor Building Closed Cooling Water System (RBCCW) inside containment consists of a closed loop providing cooling for the reactor recirculation pump heat loads and penetration heat loads. The system penetrates the containment with a supply and return line. Each line has an inboard and an outboard automatic isolation valve. Each penetration (M-16, M-17) has been protected by a relief valve mounted between the inboard automatic isolation and the containment penetration. These relief valves were added in response to NRC Generic Letter GL 96-06 concerns for isolated line overpressurization during a LOCA.

The penetration configuration must meet GDC 56.

Primary Containment Chilled Water System The Primary Containment Chilled Water System (PCCW) consists of two independent trains of cooling for the primary containment atmosphere. Each train penetrates the containment with a supply and return line. Each line has an inboard and an outboard automatic isolatio n valve. Each penetration (M-25, M-27, M-28, M-26) has been protected by a relief valve mounted between the inboard automatic isolation and the containment pe netration. These relief valves were added in response to NRC Generic Letter GL 96-06 concerns for isolated line overpressurization during a LOCA.

The penetration configuration must meet GDC 56.

6.2.4.2.3 Evaluation Agains t General Design Criterion 57 Lines penetrating the primary containment for which neither Criterion 55 nor Criterion 56 govern comprise the closed system isolation valve group.

LSCS-UFSAR 6.2-62 REV. 14, APRIL 2002 Influent and effluent lines of this group are isolated by automatic or remote manual isolation valves located as closely as possible to the containment boundary.

ECCS Pump Test Lines and Minimum Flow Bypass Lines The LPCS, HPCS, and RHR pump test and minimum flow bypass lines have

isolation capabilities. All the pump test lines are equipped with normally closed motor-operated globe valve outside the containment that is opened only during pump testing. The RHR pump test lines discharge below the surface of the suppression pool. Thus, the lines are not directly open to the containment atmosphere, since the pool acts to seal the discharge from the containment. The LPCS and HPCS lines discharge into the air space above the suppression pool surface. All the test lines are low-pressure lines, constructed to the same quality standards as the containment. All valves can be remote manually operated from the main control room, and close automatically on a system start signal.

The minimum flow bypass line on the HPCS has a normally closed motor-operated gate valve located outside the containment while the LPCS and RHR are minimum flow bypass lines equipped with a normally open motor-operated gate valve. A high speed valve is utilized to assure that pump minimum flow requirements are met.

The LPCS and RHR valves are closed when adequate flow in the pump discharge lines is established. The minimum flow bypass lines connect into the associated pump test lines outside the containment. This reduces the number of penetrations through the primary containment, thus minimizing the potential pathways for radioactive material release.

RCIC Turbine Exhaust, Vacuum Pump Discharge and RCIC Pump Minimum Flow Bypass The RCIC turbine exhaust and vacuum pu mp discharge lines which penetrate the containment and connect to the suppression chamber are equipped with a normally open, motor-operated, remote manually actuated valve located as close to the containment as possible. The RCIC turbin e exhaust line motor-operated isolation valve is a gate valve and the RCIC vacuum pump discharge line moter-operated isolation valve is a globe valve. In addi tion, there is a simple check valve upstream of the motor-operated valve which provides positive actuation for immediate isolation in the event of a break upstream of this valve. The gate valve in the RCIC turbine exhaust is designed to be locked open in the control room and interlocked to preclude opening of the inlet steam valve to the turbine while the turbine exhaust valve is not in a full open position. The RCIC vacuum pump discharge line is also normally open but has no requirement for interlocking with the steam inlet valve to the turbine. The RCIC pump minimum flow by pass line is isolated by a normally closed motor-operated globe valve with a check valve installed upstream. This valve is controlled by sensors in the RCIC pump discharge line flow and pressure. The valve is also remote manually controlled from the main control room.

LSCS-UFSAR 6.2-63 REV. 14, APRIL 2002 The RCIC turbine exhaust line is also provided with a vacuum breaker system to prevent condensation of the exhaust steam from inducing a vacuum in the line. The vacuum relief line connects the turbine exhaust line to the suppression chamber atmosphere.

Two check valves in-series in the line prev ent steam from exhausting to the vapor space above the pool, and two motor-operat ed globe valves provide remote manual isolation capability for the vacuum breaker line.

ECCS and RCIC Safety/Relief Valves The safety/relief valves which serve the RHR shutdown cooling line located outside primary containment, RHR Pumps A and C suction lines, RHR Pumps A, B, and C discharge lines, RHR Heat Exchanger drain lines to the RCIC System, LPCS and HPCS suction drain lines, RHR Pumps A and B suction drain lines and discharge drain lines, RHR Pump C discharge drain line, LPCS Pump suction and pump discharge lines, and the HPCS Pump suction line and water leg pump discharge line, discharge water into the air space above the suppression pool surface. The safety/relief valve on RHR Pump B suction line discharges water below the suppression pool surface. The safety/re lief valves on the RHR Heat Exchangers Shell Side and the RCIC steam supply lines to the RHR Heat Exchangers discharge steam below the suppression pool surface. The safety/relief valves are normally closed and provide a containment barrier in the lines. The thermal expansion safety/relief valve on the Unit 1 HPCS pump discharge line discharges water to the reactor building equipment drains and is normally closed. The thermal expansion safety/relief valve on the Unit 2 HPCS pump discharge line discharges water to the Unit 2 HPCS Pump Room and is normally closed. The safety/relief valves on the RCIC Lube Oil Cooler Supply Line, the RCIC System Pump suction line, and the RCIC Barometric Condenser discharge water to the reactor building equipment drains and are normally closed. Block valves cannot be added to the safety/relief valve discharge lines because they would preclude proper operation of the safety/relief valves, and are prohibited by the piping codes.

ECCS and RCIC Pump Suction Lines The RHR, RCIC, LPCS, and HPCS suction lines contain motor-operated, remote manually actuated, gate valves which provide assurance of isolating these lines in the event of a break. These valves also provide long-term leakage control. In addition, the suction piping from the suppression chamber is considered an extension of containment since it must be available for long-term usage following a design basis loss-of-coolant accident, and as such is designed to the same quality LSCS-UFSAR 6.2-63a REV. 14, APRIL 2002 standards as the containment.

Thus, the need for isolation is conditional since the ECCS pumps take suction from the suppre ssion pool in orde r to mitigate the consequences of LOCA. Therefore, their proper position for performing their safety fuction is open, not closed.

It should also be noted that the suction line of the ECCS pumps serves as the source of supply to the water leg pumps, which keep the ECCS discharge lines filled to avoid hydrodynamic effects on ECCS pump initiation. Isolating these water leg pumps from their supply source would de grade rather than improve the safe operation of the plant. However, the suction lines are provided with a motor-operated gate valve that can be remote manually closed from the control room, if required by a system line break or other highly unlikely event.

LSCS-UFSAR 6.2-64 REV. 17 APRIL 2008 6.2.4.2.4 Miscellaneous Compliance with regulatory guides is addressed in Appendix B.

The isolation valves have been designed against loss of function from missiles, jet forces, pipe whip, and earthquake. The containment isolation valves and valve operators have been designed to oper ate under normal plant and postulated accident conditions. The effects of radiation, humidity, pressure and temperature both inside and outside the containment, as defined in Chapter 3.0, have been accounted for in the valve design.

Containment isolation valves are provided with adequate mechanical redundancy to preclude common mode failures. The power supplies to the inboard isolation valves are provided from a separate electrical di vision than those that supply the outboard isolation valves. Therefore, a common mode failure in one electrical division would not prevent containment isolation. The vent and purge valves consist of Air Operated Valves and Motor Operated Valv es. See Table 6.2-21 for specific valve characteristics.

A complete list of Primary Containment Isolation Valves is contained in Table 6.2-28.

A leak detection system has been provided to detect leakage for determining when to isolate the affected systems that requ ire remote manual isolation. This leak detection system is described in Subsection 5.2.5.

The design provisions for testing the leakage rates of the containment isolation valves are shown in the valve arrangement drawings, Figure 6.2-31 as referenced in Table 6.2-21. The test connections indicated consist of a double-valved test line with provision for a pressure gauge attachment.

The design provision for testing the leakage rates of the containment isolation valves 2FC086 and 2FC115 is shown on va lve arrangement drawing, Figure 6.2-31, Sheet 10C, Detail "AD". The test connection indicated consists of a single valve test line with a provision for a pressure gauge attachment.

6.2.4.3 Design Evaluation The main objective of the containment isolat ion system is to provide protection by preventing releases to the environment of radioactive materials. Redundancy is provided in design aspects to satisfy the requirement that an active failure of a single valve or component does not prev ent containment isolation: Mechanical components are redundant, as shown by the isolation valve arrangements.

LSCS-UFSAR 6.2-65 REV. 17 APRIL 2008 Electrical redundancy is provided in is olation valve arrangements to eliminate dependence on a single power source to attain isolation. Electrical cables for isolation valves in the same process line have been routed separately. Cables have been selected based upon the specific environment to which they will be subjected.

Provisions ensure that the po sition of all nonpowered isol ation valves is maintained. For all powered valves, the position is indicated in the main control room. A discussion of the instrumentation and contro ls associated with the isolation valves is given in Chapter 7.0.

In single failure analysis of electrical systems, no distinction is made between mechanically active or passive components; all fluid system components such as valves are considered "electrically active" whether or not "mechanical" action is required.

Electrical systems as well as mechanical systems are designed to meet the single failure criterion for both mechanically ac tive and passive fluid system components regardless of whether that component is required to perform a safety action. Even though a component such as an electrically operated valve is not designed to receive a signal to change state (open or closed) in a safety scheme, it is assumed as a single failure that the system component changes state or fails. Electrically operated valves include valves that are electrically piloted but air operated as well as valves that are directly operated by an electrical device. In addition, all electrically operated valves that are automatically actu ated also can be manually actuated from the main control room. Therefore, a single failure in any electrical system is analyzed regardless of whether the loss of a safety function is caused by component failing to perform a requisite mechanical motion or a component performing an

unnecessary mechanical motion.

6.2.4.4 Tests and Inspections A discussion of the testing and inspection pertaining to isolation valves is provided in Subsection 6.2.6, the Technica l Specifications, and Table 6.2-21.

6.2.5 Combustible Gas Control in Containment In order to assure that the containment integrity is not endangered due to the generation of combustible gases following a postulated LOCA, systems for controlling the relative concentrations of su ch gases are provided within the plant. The system includes subsystems for mixing the containment atmosphere, monitoring hydrogen concentration, reduci ng combustible gas concentrations, and, as a backup, purging. The hydrogen recombining function of the hydrogen recombiners is abandoned in place.

LSCS-UFSAR 6.2-66 REV. 17, APRIL 2008 6.2.5.1 Design Bases The hydrogen recombining function of th e hydrogen recombiners is abandoned in place. The valves that provide RHR coo ling water to the hydrogen recombiners are also abandoned in place in the closed posi tion. The blower an d associated piping are not abandoned and remain operational to maintain the drywell mixing function. The design basis information for the hydrogen recombination function remains for historical reference.

The following design bases were used for the combustible gas control system design:

a. A double-ended rupture of a main recirculation line results in the most rapid coolant loss and reactor depressurization, with the coolant being

discharged from both ends of th e break. The noncondensable gas initially in the drywell is forced into the suppression chamber during the RPV depressurization phase. This transfer process takes place through downcomers that connect the drywell and suppression chambers. The postulated metal-water reaction begins in the core region and is assumed to produce hydrogen immediately after the recirculation pipe breaks. The reac tion would last 2 minutes during which 0.945% of the active Zircaloy fuel cladding has reacted. The radiolysis of the coolant in the core region, water sump on the drywell floor and suppression pool also is assumed to begin immediately. The hydrogen and oxygen thus generated will evolve to drywell and suppression chamber atmospheres.

b. The combustible gas control system has the capability for monitoring the hydrogen concentration in drywell and suppression chamber and alarming as the hydrogen concentrat ion reaches 4%. It also has the capability of mixing the atmospheres of both drywell and suppression chamber. It also will control the combustible gas concentrations in the primary containment without reliance on purging and without the release of radioactive material to the environment.
c. The primary systems for combustibl e gas control, including measuring, meet the design, quality assurance, redundancy, energy source, and instrumentation requirements for an engineered safety feature system according to Appendix A of 10 CFR 50.
d. The combustible gas control system will be activated after a LOCA in time to assure that the hydrogen concentration does not exceed 4

volume percent of hydrogen in either the drywell or wetwell atmospheres. In addition, the LSCS containment is nitrogen inerted to

LSCS-UFSAR 6.2-67 REV. 17, APRIL 2008 an oxygen concentration of 4% by volume. This is below the combustible limit of oxygen in hydrogen but still provides enough oxygen to react with all the hydrogen that would be produced by the metal water reaction.

e. One recombiner system is provid ed for each nuclear unit. Each recombiner is capable of being cross-connected to the other unit to provide 100% redundancy. The recomb iners are located outside of the primary containment in an accessible area and, therefore, routine maintenance, testing and/or inspection can be performed during normal plant operation or shutdown conditions.
f. The components of the combustible gas control system are protected from missiles and pipe whip to assu re proper operation under accident conditions as required for safety-related systems. The system has been designed to perform in the event of failure of any one of its active components.
g. The combustible gas control systems are designed as Seismic Category I devices. As previously mentioned, the units are capable of being cross-connected to provide redundancy and are further capable of withstanding the temperature and pressure transients resulting from a LOCA. All components that can be subjected to containment atmosphere are capable of withstanding the humidity, temperature, pressure, and radiation conditions in the containment following a LOCA. h. The combustible gas control system is designed to remain operable in the postaccident environment in the reactor building. Components subjected to the reactor containment postaccident environment are likewise designed for those conditions.
i. The combustible gas control system recombiner units are located outside of the primary containment in an accessible area. They can be inspected or tested during normal plant operation or during shutdown conditions.
j. The hydrogen recombiner units are fixed units that are permanently installed; therefore, it is not necessary to have the ability to transport them. k. The recombiner units are remotely started from the control room and the local control panel in the auxiliary electric equipment room. They are designed such that there are no local operating adjustments required on a unit operating in a post-LOCA environment. This fact eliminates the necessity of biological shielding.

LSCS-UFSAR 6.2-67a REV. 17, APRIL 2008 6.2.5.2 System Design The combustible gas control system consists of four subsystems: a mixing system, a hydrogen monitoring system, two hydrogen recombiners, and a purge system. The design features of these four systems are described in the following sections.

The hydrogen recombining function of the hydrogen recombiners is abandoned in place. The valves that provide RHR coo ling water to the hydrogen recombiners are also abandoned in place in the closed posi tion. The blower an d associated piping are not abandoned and remain operational to maintain the drywell mixing function. The design basis information for the hydrogen recombination function remains for historical reference.

LSCS-UFSAR 6.2-68 REV. 14, APRIL 2002 Hydrogen Mixing System The function of the mixing subsystem is to ensure that local concentrations with greater than 4% hydrogen cannot occur within the primary containment following a LOCA. The atmospheres of both drywell proper and suppression chamber area, each of

which is a single compartment, are well mixe

d. The mixing is achieved by natural convection processes. Natural convection occurs as a result of the temperature difference between the bulk gas space in the vessel and the containment wall. The natural convective action is enhanced by the momentum of steam emitted from the point of rupture. There are two interior subcompartments where gases may not achieve thorough mixing with the bulk containment atmosphere. The drywell head area, which is for reactor vesse l refueling purposes, is one such subcompartment. The other is the control rod drive area immediately below the reactor pressure vessel. The physical arrangements and/or location of the monitoring system and the hydrogen recombiner system are such that concentrations above the 4% limit of combustible gases will not occur.

The atmosphere between the drywell and suppression pools will be mixed during the depressurization phase of the LOCA.

The hydrogen recombiner units will also serve to affect mixing between these two compartments. The hydrogen recombiner will take suction on the drywell and discharge to the suppression pool. This will in turn cause the atmosphere from the suppression pool to circulate into the drywell via the vacuum breaker lines.

The monitoring system will alert the operator of the concentration within these subcompartments and the positions of the effluent and suction points of the recombiner will preclude the building of concentrations above the limit in these areas as well as the drywell and wetwell proper.

Hydrogen Monitoring System The hydrogen monitoring system forms a part of the primary containment monitoring system which is di scussed in Subsection 7.5.2.

Hydrogen Recombiner System The concentration of combustible gases in the primary containment (drywell and suppression pool areas) following a LOCA is controlled by the hydrogen recombiner system. The combustible gas control system contains one hydrogen recombiner per reactor unit. The hydrogen recombiner is located outside of the primary containment. The amount of Hydr ogen in the effluent gas being returned to the wetwell shall not exceed 0.1% by volume. The system will process the primary containment atmosphere at a rate of at least 125 scfm using a blower to supply containment gases to the recombiner. The recombination process LSCS-UFSAR 6.2-69 REV. 14, APRIL 2002 takes place within the recombiner as a resu lt of an exothermic reaction. The steam is then cooled and the resulting water and remaining gases are returned to the primary containment. Suction is taken from the drywell area, and the discharge is returned to the suppression po ol area above water level.

The hydrogen recombiner unit is skid mounted and is an integral package. All pressure containing equipment including pi ping between components is considered as an extension of the containment and, ther efore, is designed as ASME III Class 2. The skid and the equipment mounted on it are designed to meet Seismic Category I requirements. The hydrogen recombiner system is designed to accommodate conditions present in the containment (temperature and pressure) following a LOCA event. Piping and instrumentation for the system are shown in Drawing No.

M-130. The hydrogen recombiner unit, whic h requires a 1-2 hour warmup period, is initiated manually from the control room and the local control panel in the aux. electric equipment room. It is initiate d prior to primary containment hydrogen concentration reaching 3 volume percent which occurs approximately 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after the accident. Based on the original core loading, the time at which containment hydrogen generation reaches 4 volume percent varies with fuel types located in the core. However, this is acceptable based on Design Basis described in Section 6.2.5.1.d. Once placed in operation, th e system continues to operate until it is manually shut down when an adequate ma rgin below the hydrogen concentration design limit is reached. The operation of the system can be tested from the control room or the auxiliary equipment room. The test consists of energizing the blower and heaters and observing system operation to see if components are performing properly. Flow and pressure measuremen t devices are periodically calibrated.

The hydrogen recombiner system is servic ed by electrical power and cooling water systems, which are placed in operation concurrent with a loss-of-coolant accident.

Cooling water required for the operation of the system is taken from the residual heat removal system. The cooling water is utilized to cool the water vapor and the residual gases leaving the recombiner prior to returning them to the containment. All hydrogen recombiner unit cooling water is returned to the suppression pool.

Each recombiner unit has the capability of serving either containment; therefore, there is 100% redundancy of all components and controls.

All functions and controls necessary to st art the combustible gas control system are also located in the control room and in the auxiliary electric equipment room which is readily accessible from the control room.

LSCS-UFSAR 6.2-70 REV. 17, APRIL 2008 6.2.5.3 Design Evaluation The hydrogen recombining function of the hydrogen recombiners is abandoned in place. The valves that provide RHR coo ling water to the hydrogen recombiners are also abandoned in place in the closed posi tion. The blower an d associated piping are not abandoned and remain operational to maintain the drywell mixing function. The design basis information for the hydrogen recombination function remains for historical reference.

6.2.5.3.1 General In evaluating the combustible gas control system design, it was found necessary to

consider:

a. hydrogen generated in the post-LOCA environment, b. resultant drywell and containment concentrations, and
c. the functional requirements of the combustible gas control system.

The following analytical results are provided:

a. The beta, gamma, and beta plus ga mma energy release rates plotted as functions of time (Figure 6.2-32).
b. The integrated beta, gamma and beta plus gamma energy release plotted as functions of time (Figure 6.2-33).
c. The integrated production of combustible gas within the containment (drywell and suppression chamber) plotted as a function of time for each source (i.e., metal-water reacti on and radiolysis) (Figure 6.2-34).
d. The concentration of combustible gas in the drywell and suppression chamber plotted as a function of time , if uncontrolled (Figure 6.2-35). This curve establishes the basis for activation of the combustible gas control system.
e. The combustible gas concentration in the containment (drywell and suppression chamber) plotted as a function of time with (125 scfm) 100% recombiner capacity initiated at 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after LOCA (Figure 6.2-36).

LSCS-UFSAR 6.2-71 REV. 14, APRIL 2002 6.2.5.3.2 Sources of Hydrogen Short-Term Hydrogen Generation In the period immediately after the LOCA, hy drogen is generated by both radiolysis and metal-water reaction. However, in ev aluating short-term hydrogen generation, the contribution from radiolysis is insignificant when compared to the hydrogen generated by the metal-water reaction. The only metal-water reaction considered to be significant is reaction of water with the zirconium fuel cladding which produces hydrogen by the following reaction:

Zr + 2H 2 O ZrO 2 + 2H 2 Based on loss-of-coolant accident calculat ional procedures and the analyses of emergency core cooling system (ECCS) pe rformance in conformance with 10 CFR 50.46 and Appendix K, the extent of the above chemical reaction is estimated to be 0.1% of the fuel cladding material. However, the metal-water reaction-generated hydrogen based on a core-wide penetratio n of 0.00023 inches for 764 bundles with each bundle containing 101 pounds of zircon ium in the active fuel cladding, results in a 0.945% metal-water reaction. Theref ore, 0.945% of fuel cladding, which is greater than five times the maximum amount calculated in accordance with 10 CFR 50.46, is assumed to react with water to produce hydrogen. The duration of this reaction is assumed to be 120 seconds with a constant re action rate. The resulting hydrogen is assumed to be uniformly distributed in the drywell containment. This assumption is supported by the test data reported in BNWL 1592 of July 1971.

Figure 6.2-34 presents the accumulated hydr ogen generation as a result of this chemical reaction.

Long-Term Hydrogen Generation Hydrogen is also produced by decomposition of water due to absorption of the fission product decay energy immediately after LOCA.

2H 2 O 2H 2 + O 2 Generation of hydrogen and oxygen due to radiolysis of coolant water is an important factor in determining the long-term gas mixture composition within the containment compartments. Conservative assumptions were used to determine the fission product distribution model that applies after the accident and, therefore, the hydrogen generation rates. The incore radiolysis contributes hydrogen to the drywell, and radiolysis of the suppression pool water contributes hydrogen directly to the suppression chamber. Hydrogen is also discharged from the radiolysis of sump water on drywell floor into the drywell atmosphere. The total decay energy utilized in the analyses was based on American Nuclear Society Standard ANS 5.1-1979 multiplied by a factor of 1.2, conservatively assuming a 1000-day reactor LSCS-UFSAR 6.2-72 REV. 14, APRIL 2002 operating time at constant full power level to determine the fission product buildup.

Halogen and noble gas inventories were determined from TID-14844.

Hydrogen can also be formed by corrosion of metals in the containment. The significant portion of this source is from the corrosion of zinc and aluminum. Since the spray system uses only demineralized water for the purpose of reducing temperature and pressure inside the drywell, the corrosion of aluminum and zinc will contribute a negligible amount of hydrogen to the containment atmosphere. Hydrogen is, during normal operation of the plant, dissolved in the primary system water. Figure 6.2-35 presents the accumulated hydrogen and oxygen generation from both chemical reaction and radiolysis decomposition of water.

6.2.5.3.3 Accident Description A complete description of the post-LOCA cond itions is found in Subsection 6.2.1 and Section 6.3.

Following the postulated LOCA, the postulated metal-water reaction begins in the core region and is assumed to produce hydrogen immediately after the recirculation pipe breaks. The reaction lasts 2 minut es during which 0.945% of the active zircaloy fuel cladding reacts. The radiolysis of the coolant in the core region, water sump on the drywell floor and suppression pool is assumed to begin immediately.

The hydrogen and oxygen thus generated will evolve to drywell and suppression chamber atmospheres. The hydrogen conc entration in the drywell would, after about 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br />, approach the flammability limit if uncontrolled. The hydrogen recombiner system is manually activated be fore the hydrogen concentration reaches 3 volume percent. The recombiner system takes gases from the drywell atmosphere, recombines the hydrogen with oxygen to form water vapor, and returns the resulting cooled water and remaining gases to the suppression chamber. The pressure buildup in the suppression cham ber due to the operation of recombiner system taking suction on the drywell and discharging to the suppression pool will cause the opening of the vacuum brea ker valves between the drywell and suppression chamber. As a result, the flow of the gas mixture from the wetwell to the drywell will balance the negative pressure differential between two volumes and will also result in lower concentrations due to the influx of the wetwell gases.

6.2.5.3.4 Analysis Based on the above hydrogen sources and the accident description, the hydrogen concentration in the drywell and suppression chamber is calculated as a function of time. In formulating the model of the Mark II containment for these calculations, a conservative assumption is made, namely the interchange of mass between the drywell and the suppression chamber through downcomers which takes place during blowdown process is neglected, that is, no hydrogen is removed from the drywell except through the recombiner system. This assumption is conservative, as LSCS-UFSAR 6.2-73 REV. 15, APRIL 2004 it results in a shorter time for the dryw ell hydrogen concentration to reach the flammability limit. Furthermore, the hydr ogen and oxygen gases can flow back to the drywell from suppression chamber through vacuum breakers due to pressure increase in the suppression chamber by th e operation of the recombiner system.

Table 6.2-22 gives all of the necessary parameters used to determine the amount of

hydrogen generation in the LSCS analysis. The results of the analyses are presented in Figures 6.2-35 and 6.2-36. It was determined that the uncontrolled hydrogen concentration in the drywell even tually reaches 4% by volume (dry basis) approximately 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br /> after the LOCA. The suppression chamber hydrogen concentration was determined to be 3.0% by volume due to radiolytic hydrogen generation. Prior to the drywell concentrat ion reaching 3% by volume, a recombiner system is activated. A single system is designed to keep the hydrogen concentration below 4% by volume at all times until ra diolytic generation has ceased. The performance of the recombiner system, which is initiated 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after LOCA, is shown in Figure 6.2-36. The hydrogen conc entration is 3.0% by volume at the time of initiation. Thus, the use of a sing le 125 scfm recombiner system provides effective control of hydrogen concentrat ion and, therefore, would prevent the formation of combustible gas mixture in both drywell and suppression chamber.

6.2.5.4 Testing and Inspections Each active component of the combustible gas control system is testable during normal reactor power operation.

The combustible gas control systems and the containment purge system will be tested periodically to assure that they will operate correctly.

Preoperational tests of the combustible gas control system are conducted during the final stages of plant construction prior to initial startup (Chapter 14.0). These tests assure correct functioning of all controls, instrumentation, recombiners, piping, and valves. System reference characteristics, such as pressure differentials and flow rates, are documented during the preoperational te sts and are used as base points for measurements in subsequent operational tests.

6.2.5.5 Instrumentation Requirements The instrumentation provisions for actuat ing the combustible gas control system and monitoring the system are described in Subsection 7.3.5.

6.2.6 Containment Leakage Testing

This section presents the testing program for the reactor containment, containment penetrations and containment isolation barriers that comply with the requirements of the General Design Criteria and Append ix J to 10 CFR 50. Each of the tests LSCS-UFSAR 6.2-74 REV. 19, APRIL 2012 described in this Subsection was performed as a preoperational and will be performed as a periodic test.

6.2.6.1 Containment Integrated Leakage Rate Test Following the completion of the construction, repair, inspection, and testing of welded joints, penetrations, and mechanical closures including the satisfactory completion of the structural integrity test s as described in Subsection 3.8.1.7, a preoperational containment leakage rate test was performed to verify that the actual containment leak rate does not exceed the design limits. In order to ensure a successful integrated leak rate test, loca l leakage tests (Type B and C tests) were performed on penetrations and isolation valves, and repairs are made, if necessary, to ensure that leakage through the contai nment isolation barriers does not exceed the design limits.

An integrated leakage rate test is then performed on the entire containment in order to determine that the total leakage (exclusive of MSIV leakage) through containment isolation barriers does not ex ceed the maximum allowable leakage rate of 1.0% per day at the calculated peak cont ainment internal pressure at 39.9 psig. The pertinent test data, including test pressures and acceptance criteria, is presented in Table 6.2-23.

Pretest requirements have been describe d in the preoperational test abstract included in Chapter 14.0 of the FSAR. As stated therein, power operated isolation valves will be closed by their actuators prior to the start of the integrated leakage rate test.

During the integrated leak rate test the containment systems are configured as follows; a. Reactor building closed cooling water - lined up for normal operation; isolation valves closed and system filled.

b. Primary containment chilled water - lined up for normal operation; isolation valves closed and system filled.
c. Residual heat removal - One loop lined up in shutdown cooling mode. Other loops lined up in low-pressure coolant injection standby mode and isolated, containment and suppr ession pool spray flow paths isolated, full flow test lines isolated, reactor head cooling flow path isolated, minimum flow isolated, shutdown cooling discharge line isolated on standby system and condensate discharge from RHR heat exchangers shell side flow path isolated; system filled. May be lined up in normal standby injection mode.
d. Low-pressure core spray - system filled and isolated. May be lined up in normal standby injection mode.

LSCS-UFSAR 6.2-75 REV. 18 APRIL 2010 e. High-pressure core spray - system filled and isolated. May be lined up in normal standby injection mode.

f. Reactor core isolation coolin g - isolation valves closed; RCIC condensate filled and isolated.

RCIC full flow test return line to suppression pool filled and isolated.

g. Reactor water cleanup - suction line filled and isolated; return line filled and isolated.
h. Standby liquid control - lines filled and isolated.
i. Control rod drive - system filled. Vented outboard of HCU directional control valves.
j. Reactor recirculation system - pumps off, system filled.
k. RPV and primary containment instrumentation - lines filled and vented to containment instrumentation to the RPV or drywell will be opened. l. Neutron monitoring system (TIP) - TIPs will be fully withdrawn and the ball valves closed.
m. Floor and equipment drains - sumps pumped down to low water level, isolation valves closed.
n. Clean condensate - drained and ve nted, isolation valves closed, spool piece removed and blind flange installed or filled and isolated and system leakage added to type A result.
o. Service air - vented, isolation va lves closed, spool piece removed and blind flange installed.
p. Feedwater - filled and isolated.
q. Main steam - filled, isolation valves closed.
r. Containment monitoring - post-LOCA monitoring system open to containment, pumps off, valves op en; drywell monitoring and sampling system isolated, pumps off.
s. Post-LOCA hydrogen control - lined up for unit operation, isolation valves open or isolated and system leakage added to type A result.

LSCS-UFSAR 6.2-76 REV. 18, APRIL 2010 t. Primary containment instrument air - all accumulators vented, isolation valves closed. u. Fuel Pool Cooling - Cycled Condensate to Refueling Bellows filled and isolated, Reactor Well Drain filled and isolated. v. All accessible liner leak test channel plugs are verified installed. The Type C leak rates for the following penetrations are added to the Type A test results on a Minimum-Path Basis: a. reactor building closed cooling water, b. primary containment chilled water, c. RHR shutdown cooling suction, d. reactor core isolation cooling steam supply, e. reactor water cleanup suction, f. reactor water sample, g. floor and equipment drains, h. inboard MSIV drain, i. Feedwater Lines, j. RCIC Full Flow Test Return Line to Suppression Pool. k. Cycled Condensate to Refueling Bellows l. Reactor Well Drain Measures will be taken to ensure stabilization of the containment conditions prior to containment leakage rate testing. The test method utilized is the absolute method, as described in ANSI/ANS 56.8-1994. The test procedure, test equipment and facilities, period of testing, and verification of leak test accuracy also follow the recommendations of ANSI/ANS 56.8-1994. The acceptance criteria for the preoperational containment integrated leakage rate test are in compliance with the criteria given in Appendix J of 10 CFR 50. except as LSCS-UFSAR 6.2-77 REV. 13 noted below. Structural verification test acceptance criteria are described in Subsection 3.8.1.7.

The acceptance criteria for the periodic containment integrated leakage rate test are in compliance with the criteria given in 10CFR50 Appendix J Option B, NRC Reg Guide 1.163, NEI-94-01, Rev. 0, and ANSI/ANS 56.8-1994. The As-Found Type A test leakage must be less than the acceptance criterian of 1.0 La (Primary Containment overall leakage rate acceptance criterion). During the first unit startup following testing (prior to entering a mode where containment integrity is required) the As-Left Type A leakag e rate shall not exceed 0.75 La.

6.2.6.2 Containment Penetration Leakage Rate Test

Containment penetrations whose design incorporates resilient seals, gaskets, or sealant compounds; air lock door seals, equipment and access doors with resilient seals or gaskets; and other such penetrations received a preoperational and will be periodically leak tested in accordance with Appendix J of 10 CFR 50 except as noted in the following paragraph.

The following penetrations were preoperationally and will be periodically tested to Type B criteria:

a. equipment access hatch, b. personnel air lock, by (when co ntainment integrity is required, the personnel airlock should be tested within 7 days after each containment access except when the airlock is being used for multiple entries, then at least once per 30 da ys, by verifying seal leakage to be less than or equal to 5 scfh when the gap between the door seals is pressurized to greater than or equal to 10 psig - exception to 10 CFR 50 Appendix J) overall air lock leakage rate is less than or equal to 0.05 La when tested at greate r than or equal to Pa.
c. drywell head,
d. suppression chamber access hatches, e. CRD removal hatch, f. electrical penetrations, g. TIP penetration flanges, SA flange and MC flange, h. Drywell to suppression pool vacuum breaker and associated manual isolation valves flanges and actuator seals, LSCS-UFSAR 6.2-78 REV. 13 i. Vent and purge isolation valve flanges, and packing
j. HPCS minimum flow line branch line 1(2)HP20C-2" Blind flanges
k. RCIC spectacle flange 1(2)E51-D316 blind flange half when required.

See Table 6.2-21 note 49.

l. ECCS Relief Valves Discharg ing to Suppression Pool Inbound (Containment Side) Flanges.

It should be noted that no pipe penetrations are provided with expansion bellows.

The containment penetration is an anchor point in the system, and the thermal movements have been accounted for on th is basis. Therefore, no leakage rate testing of expansion bellows penetration assemblies will be required.

Test methods utilized to determine containment penetration leak rates are described as follows:

a. Equipment Access, CRD Removal, and Suppression Chamber Acess The equipment access hatch has been furnished with a double-gasketed flange and bolted dished door, as shown in Figure 3.8-34. The CRD removal and suppression chamber access hatches have been furnished with a double-gasketed flan ge and bolted door. Provision is made to test pressurize the space between the double gaskets of the door flanges and the doors.
b. Personnel Air Lock The personnel lock is constructed as a double-door, latched, welded steel vessel, as shown in Figure 3.8-33. The space between the air doors can be pressurized to peak containment pressure through the test connections provided. Each of the doors are provided with a test connection for pressurizing between the seals.

In addition, all four shaft seal assemblies are provided with a test connection to allow for individual shaft seal leak test.

c. Drywell Head A double-gasketed seal and test tap, as shown in Figure 3.8-5, is provided for leak rate testing of the drywell head.
d. Electrical Penetrations

LSCS-UFSAR 6.2-79 REV. 13 Each electrical penetration, as represented in Figure 3.8-21 and listed in Table 3.8-1 (with an "E" penetration number), is provided with a pressure gauge to monitor leakage. The double-gasketed and O-ring seals are provided with a test connection for leak rate testing.

e. Tip Penetration Flanges, Clean Co ndensate (MC) and Service Air (SA)

Penetrations Each TIP MC or SA penetration fl ange is provided with a double-gasketed seal and a test connection for type B leak testing.

f. Drywell to Suppression Pool Vacuum Breakers Each drywell to suppression pool vacuum breaker has two double-gasketed flanges and a manual actuat or O-ring and shaft seal. These seals are provided with test connecti ons for leak testing. The Vacuum Breaker line manual isolation valves have a double-gasketed flange on the inboard or containment side provid ed with test connections for leak testing. The outboard flanges on the manual isolation valves are leak tested by pressurizing the entire vacuum breaker line and performing

soap bubble test on the outboard flan ge. The stem seal or packing of these valves will be tested either locally or by primary containment pressurization and subsequent soap bubble inspection.

g. Vent and Purge Isolation Valves Each inboard vent and purge valve has a double-gasketed flanged seal

on its containment side. These seals are provided with test connections for leak testing. The st em packing of these valves is also provided with a test connection for packing leak test. See also Table 6.2-21 Note 41.

h. HPCS Minimum Flow Line Blind Flanges One double-gasketed blind flange is installed on each of the HPCS minimum flow line branch connections 1(2)HP20C-2". These flanges are provided with a test connection for type B leak testing.
i. RCIC Spectacle Flange 1(2)E51-D316 The installed blind flange half of spectacle flange 1(2)E51-D316 is tested by pressurizing with air the upstream RCIC full flow test return line to Condensate Storage Tank and then check for leaks at the flange upstream gasket joint. Done when required per Table 6.2-21 note 49.

LSCS-UFSAR 6.2-80 REV. 13 j. ECCS Relief Valves' Containment Side Flanges are Type B tested by one of the following methods: Test Port/Testable Gasket; Primary Containment Pressurization and su bsequent soap bubble inspection; Special Test Equipment mounted over the flange thus pressurizing against the gasket.

Test pressures are given in Table 6.2-23.

The acceptance criteria for the preoperational containment penetration leakage rate test is in compliance with the criteria given in Appendix J of 10 CFR 50. The periodic test acceptance criteria is established in accordance with the LaSalle County Station Local Leak Rate Test Program, and also is in agreement with Appendix J Option B of 10 CFR 50, NRC Regulatory Guide 1.163, Nuclear Energy Institute NEI-94-01 Rev.

0, and ANSI/ANS-56.8-1994.

6.2.6.3 Containment Isolat ion Valve Leakage Rate Test Those containment isolation valves that are to receive a Type C test are so indicated in Table 6.2-21.

Test taps for leakage rate testing have be en provided on the lin es associated with the containment isolation valves. These taps are indicated on the valve arrangement drawings associated with Ta ble 6.2-21. The test method is as described in Appendix J of 10 CFR 50. Te st pressures are shown in Table 6.2-23.

The acceptance criteria for the leakage rate testing is given in Table 6.2-23 and the Primary Containment Leak Rate Testing Program.

6.2.6.4 Scheduling and Reporting of Periodic Tests The periodic leakage test schedule is give n in the LaSalle County Station Leak Rate Test Program.

6.2.6.5 Special Testing Requirements The secondary containment will be tested as required by the Technical Specifications.

6.2.7 References

1. F. J. Moody, "Maximum Two-Phase Vessel Blowdown from Pipes," Topical Report APED-4827, General Electric Company, 1965.

LSCS-UFSAR 6.2-81 REV. 14, APRIL 2002

2. A. J. James, "The General Electric Pressure Suppression Containment Analytical Model, (NEDO-10320), April 1971.
3. A. J. James, "The General Electric Pressure Suppression Containment Analytical Model," April 1971, Su pplement 1, (NEDO-10320), May 1971.
4. K. V. Moore and W. H. Ratting, "RELAP 4-A Computer Program for Transient Thermal-Hydraulic Analysis, "ANCR-1127, Aerojet Nuclear Company, December 1973.
5. F. J. Moody, "Maximum Rate of a Single Component, Two Phase Mixture," Journal of Heat Transfer, Transactions, American Society of Mechanical Engineers, Vol. 87, No. 1, February 1965.
6. I. E. Idelchik, Handbook of Hydraulic Resistance, AEC-TR-6630, 1966.
7. "RELAP 4/MOD5 A Computer Progra m for Transient Thermal- Hydraulic Analysis of Nuclear Reactors and Related Systems," ANCR-NUREG-1335, Aerojet Nuclear Company, September 1976.
8. NEI 94-01, Rev. 0, July 26, 1995, Nuclear Energy Institute Industry Guideline for Implementing Performance-Based Option of 10CFR Part 50 Appendix J.
9. ANSI/ANS 56.8-1994, American Nation al Standard for Containment System Leakage Testing Requirements.
10. GE Document EAS-49-0888, "Justification of Continued Operation With Increased Suppression Pool Temperature at LaSalle County Station," Revision 1, August 1988. (Proprietary)
11. Technical Specification Submittal Lette r Sections 3.6.2.1 and 4.6.2.1, dated 10-07-88.
12. Amendment 67 for Unit 1 (Facilit y Operating License NFP-11), and Amendment 49 for Unit 2 (Facility Operating License NFP-18), dated July 7, 1989. 13. Calc. L001799, Rev. 0, "Assessment of Containment Line Base Mat Reactor Pedestal, Downcomer Bracing, Drywell Floor & Suppression Pool Columns for Suppression Pool Temperature Increase." 14. Calc. L001800, Rev. 0, "Assessment of Containment Wall for Suppression Pool Temperature Increase" LSCS-UFSAR 6.2-81a REV. 14, APRIL 2002 15. Calc. L001810, Rev. 0, "Impact of Increase in the Suppression Pool Temperature at LaSalle on Design Basis Suppression Pool Dynamic Loads." 16. Letter from ComEd NFS dated 5-07-98, Nuclear Fuel Services Letter, NFS:BSA:98-055, dated 5-08-98, from R.W. Tsai to G. Campbell, "Impact of Initial Suppression Pool Temperature on Hydrogen Generation" 17. Calc. 3C7-0181-003, Rev. 3, "Suppression Pool Temperature Transient Studies" 18. General Electric Letter Report GE-NE-B13-01920-013, January 1998, "Current Suppression Pool Water Temperatures Following a Design Basis Accident for LaSalle County Station Units 1 and 2" 19. General Electric Report EAS-083-1188, "Elimination of the High Suppression Pool Temperature Limit for LaSalle County Station Units 1 & 2", dated November 1988. 20. General Electric Letter Report GE-NE-T23-00762-00-01, July 1998, "Evaluation of Peak Suppression Pool Temperature with Assumption of Feedwater Coastdown and Reduced RHR Flow Rate During Long-Term Containment Cooling" 21. Letter from J. A. Benjamin (ComEd) to U. S. NRC, "Request for a Change to the Technical Specifications, 'Vacuum Relief System'" dated August 6, 1999. 22. Letter from J. A. Benjamin (ComEd) to U. S. NRC, "Supplemental Information to Request for a Change to the Technical Specifications to Vacuum Relief System" dated November 15, 1999. 23. Letter dated December 21, 1999 from D. M. Skay to O. D. Kingsley, "Issuance of Amendments, approved amendment 138 for LaSalle Unit 1 and amendment 122 for LaSalle Unit 2." 24. Licensing Topical Report, "Generic Guidelines for General Electric Boiling Water Reactor Power Uprate," NEDC-31897P-A, May 1992. 25. LaSalle County Station Power Uprate Project, Task 400, "Containment System Response," GE-NE-A1300384-02-01R1, Revision 1, October 1999 (and Task Report Changes based on Steam Plume Analysis, GE-LPUP-332, dated 5/4/2000).

LSCS-UFSAR 6.2-82 REV. 20, APRIL 2014 26. General Electric Company, "General Electric Company Analytical Model for Loss-of Coolant Analysis in Accordance with 10CFR50 Appendix K," NEDO-20566A, September 1986. 27. ComEd letter to NRC, "Response to Request for Additional Information License Amendment Request for Power Uprate Operation," dated 3/31/2000. 28. General Electric Company, NEDO-30832, "Elimination of Limit on Local Suppression Pool Temperature for SRV Discharge with Quenchers," Class I, December 1984, (NRC approved version with NRC Safety Evaluation Report issued as NEDO-30832-A, Class I, May 1995). 29. General Electric Analysis of LaSalle Steam Plume Ingestion Potential, NSA 00-116, dated 3/29/2000. 30. LaSalle County Station Power Uprate Project, Task 401, "Annulus Pressurization," GE-NE-A1300384-06-01, Revision 0, June 1999. 31. Design Analysis No. L-002874, Rev. 0, "LaSalle County Station Power Uprate Project Task 400: Containment System Response (GE-NG-A1300384-02-01 R3) Revision 3". 32. EC #334017, Rev. 0, "Increased Cooling Water Temperature Evaluation to a new Maximum Allowable of 104F." 33. Design Analysis L-003352, Rev. 0, "Evaluation for GE Safety Communication SC06-01 Containment System Response (GEH 0000-0069-6598-R0)." 34. Design Analysis L-003509, Revision 0, "Evaluation of Appendix R, Station Blackout, Containment and Source Terms for LaSalle MUR Power Uprate," July 2010. 35. Design Analysis L-003566, Revision 0, "T1000 Series - S/U Test and Generic Applicability," July 2010. 36. EC #388666, Rev. 0, "Revise Design Analyses for UHS Temperature of 107F"

  • *
  • LSCS-UFSAR TABLE 6.2-15 SIMULTANEOUS BREAK OF THE HEAD SPRAY LINE AND RPV HEAD VENT LINE IN THE HEAD CAVITY INPUT DATA* , SALl E -IIEIIO CAVITY PRESSUR I ZA TION -3C1'0416-003 REV 0 4266-00
  • PRORU:M DIMENSIONS 010 0 01 '2 9 5 3 3 0 0 3 0 0 0 0 0 0 0 3
  • PRUBLEM CONSTIINTS 010002 0.0 t .0
  • TIME STEP DATA 030010 I 50 0 0 0.01 0.00005 2.0 030020 50 0 0 0.002 0.00005 3.5 030030 50 0 0 0.0005 0.00005 3.9 030040 50 0 0 0.01 0.00005 8.0 030050 1 50 0 0 0.1 0.00005 30.0
  • TRTP CONTROLS 0*10010 1 1 0 0 20.0 0.0 040020 2 4 :2 3 5.2 0.0 040010 3 1 0 0 0.0 0.0
  • VOlUME DATA 0!;01) 1 I 00 15.45 135. . 001 4072 . 15.51 O. 0 261.5 18.3 819.73 0 050021 00 15.45 135. . 001 184664 79.74 O . 0 2315.0 54.3 740.00 0 050031 o 0 15.45 100. . 001 116085. 33.87 O . 0 5198.0 81.4 706.14 0
  • JUNe TJ ON Oil Til 080011 1 2 0 0 O. 11.12 819.73 .55 2.62 2.62 0 1 o 3 O. .6 -1 o O. 080021 2 3 0 1 0.0 2Q5.000 740.00 .24 1.9 1. 9 0 1 0 3 O. .6 -1 00. 080011 0 1 1 o 351.9 0.16261 821.52 .00 0.0 0.0 0 1 0 3 O. 1. -1 o O.
  • VAlVE DATA CARDS 110010 -2 0 0 O. O. O. O.
  • FIl.L TIlBlE CARDS 130100 3 1 2 3 'LBS/SEC' 550. 1. O. 130101 (' 2200. 30. 2200.
  • RELAP4/MOD5 computer code utilized for analysis . TABLE 6.2-15 REV. 0 -APRIL 1984
  • *
  • LSCS-UFSAR TABLE 6.2-16 RECIRCULATION LINE BREAK INPUT DATA* lA SAllE -HEAD CAVITY PRESSURIZATION

-3C7-0476-003 REV 0 4266-00

  • RECIRCULATION lINE BREAK
  • 4 HVAC INLET VENTS AVAilABLE FOR FLOW INTO HEAD CAVITY
  • PROBLEM DIMENSIONS 010001 -2 9 2 3 3 0 0 3 0 0 . 1 0 0 0 0 0 3
  • PROBLEM CONSTANTS 010002 0.0 1.0
  • EDIT VARIABLES 020000 AP I AP 2 AP 3 JW 2 AH 2 JW 1 TO 1 FO 1 TO 2
  • rIME STEP DATA 030010 1 50 0 0 0.005 0.00005 2.0 030020 1 50 0 0 0.01 0.00005 30.0
  • TRIP CONTROLS 1 1 0 0 10.0 0.0 2 l' 0 0 0.0 040030 3 1 0 0 0.0 0.0
  • VOLUME OAT A 050011 00 15.45 135 .. 001 4072. 15.57 0.0 261.5 18.3 819.73 0 050021 0 0 15.45 135 .. 001 177049. 79.74 O. 0 2315.0 54.3 0 050031 0 0 15.45 100 .. 001 176085. 33.87 O. 0 5198.0 81.4 706.14 0
  • JUNCTION DATA 080011 1 2 000. 4.92 819.73 .83 1.52 1.52 0 1 030 .. 6 -10 O. 080021 2 3 0 1 0.0 295.000 740.00 .24 1.9 1.9 0 1 030 .. 6 -, 0 O. 080031 0 2 1 0 25690. 1. 770 .. 00 0.0 0.0 0 1 030. 1. -1 0 O.
  • VALVE DATA CARDS 110010 -2 0 0 o. O. O. O.
  • FIll TABLE CARDS 130100 3 4 9 'LBS/SEC' TIME FLOW ENTHAlPV 130101 0.0000 22710.0 532.0 130102 0.0016 22710.0 532.0 130103 0.0017 34060.0 532.0 130104 1.5500 34060.0 532.0 130105 1.5600 27550.0 532.0 130106 1.7500 27550.0 532.0 .30107 1.7600 547.0 130108 1.9800 24840.0 547.0 130109 10.1100 24320.0 538.0
  • RELAP4/MOD5 computer code utilized for analysis . TABLE 6.2-16 REV. 0 -APRIL 1984
  • *
  • I.... < u C I.... .. .... ct o o \.? Z .... III 1-.. w .... '" z ... 1: cr o ,.., I ..0 N c., ,.., a C') 1 ..0 I' ,... o 1 " '-' c ::: .... tr. I.... -' -' "" ffl ... x: o 1 !Xl C cr :l.
  • a c. a o c -',., n:: ';'" "'" 0 I .... 0 -l: V' ., "" .... I,/' 4,' C"-If L.. I.: Z:;-l .:::. 0-N r '" " 0-... 0-.. ,.... c * :l.I 0.. 4,/"1 O!:)NVO .... O" 1 W T I.... -' c" r:: ,",,0 en 0'" C> C '=, r: r r.:'
  • c c.
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<t C ('j ..... ,1:. n .. 0.. C' o 00""0 x: N .:x: t' 0 0 0 X Cl.uCO .... CCO-l6.:N t""':f""':

LSCS-UFSAR 0"" o c:"

  • 00 o ** a: ..c "'"00 c,,"1'" I' ex: . . ,...... 0-rt'. ..,n ..... f'"") 1.1'. ::T::xJo...

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  • c CU' * ,.... i..-' c r'""'I :r ...c. t.t t'. a... Lt IJ"' :.r t I.!: * * *
  • 0 C:
  • x * ..0 n C* ::J C' Ow ,... r--. ,.... ,......,....
  • t""'; .,....",.....,....

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  • 1 c 1 ,. ,. x x . ,., c. Ci * * -c c.:
  • __ eC)O_ . ,. *** x
  • C C C )or c:. l."l COC"l ,j"'luLJif
r::r:rQO L ...... ' * !:r OI.l'J"'1.I'

< u 0 Q '" U <f ,-, * <0'" ("\J ('0.. C".' ('." V C 0-<000 C '" NNOOOOO .... ",,,-1-' cr .. .... c <1 ;:, II' 0 ,., 0 ..0 l1' t-'i 11' t".: ,... "3" CC ,.... ;T i.!': ..c"'" C* ,... N * * '" .;r ""'C""'O"!.P::r,.....

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,.,.. -I""': o ,.., o 00;; '" cc c:=,::. ::l _N"'" ....i -N QOUOOn CQW:>CC' :rT':r 0-V "" :::>C-JO-C-O-c-c-0000 ,...,..."'

00 Cl C* MMPjt""')

I"'f1 Z -'" t""'):;r 1..."1 -J::JOOQOO C. -; c.; C r .. C'.., '-' L{" ",cr",::::

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  • C
  • C
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  • C C

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TABLE 6.2-17 REV. 0 APRIL 1984

LSCS-UFSAR 6.3-34 REV. 20, APRIL 2014 6.3.5.3 LPCS Actuation Instrumentation

The LPCS is automatically actuated by the following sensed variables: reactor vessel low water level, or drywell high pressure.

In addition the LPCS can be manually actuated from the control room.

6.3.5.4 LPCI Actuation Instrumentation

The LPCI is automatically actuated by th e following sensed variables: reactor vessel low water level, or drywell high pressure. Reactor vessel low water level or drywell high pressure also stops other modes of RHR system operation so that LPCI is not inhibited.

In addition, the LPCI can be manually actuat ed from the control room. Subsection 7.3.1.3.2.3 discusses conformance to IEEE-279 and other applicable regulatory requirements for the ECCS instrumentation and controls.

LSCS-UFSAR TABLE 6.5-1 (SHEET 1 OF 4) TABLE 6.5-1 REV. 13 STANDBY GAS TREATMENT SYSTEM COMPONENTS

NAME OF EQUIPMENT TYPE, QUANTITY AND NOMINAL CAPACITY (per component)

A. Filter Unit

1. Equipment Numbers 1VG01S, 2VG01S
2. Type Package
3. Quantity 2
4. Components of Each Unit
a. Fan Type Centrifugal

Quantity 1 Drive Direct Capacity (ft 3/min) 4000 (nominal)

Static Pressure (in. H 2O) 14.8 b. Demister

Type Impingement Quantity 1 Bank with 4 elements Static resistance

clean (in. H 2O) 0.95 dirty (in. H 2O) 1.7 c. Heater Type Electric, sheathed, single stage

LSCS-UFSAR TABLE 6.5-1 (SHEET 2 OF 4) TABLE 6.5-1 REV. 17, APRIL 2008 NAME OF EQUIPMENT TYPE, QUANTITY AND NOMINAL CAPACITY (per component)

Quantity 1 Capacity (kW) 23

Accessories Overload cutout

d. Prefilter

Type High Efficiency Quantity 1 Bank With 4 Elements Efficiency (per ASHRAE) Dust Spot Test) 90% Static resistance clean (in. H 2O) 0.35 dirty (in. H 2O) 1 e. HEPA Filters

Type Absolute High Efficiency Quantity 4 Elements per Bank. Two Banks per Train Media Glass Fiber, Waterproof, Fire Resistant Bank Efficiency (% with 0.3 micron particles) 99.97 (Purchased) 99.95 (Operational Requirement)

Static Resistance clean (in. H 2O) 0.7 dirty (in. H 2O) 2 LSCS-UFSAR TABLE 6.5-1 (SHEET 3 OF 4) TABLE 6.5-1 REV. 15, APRIL 2004 NAME OF EQUIPMENT TYPE, QUANTITY AND NOMINAL CAPACITY (per component)

f. Charcoal Adsorber Bed Type Vertical gasketless

Quantity 8 - 8 in. thick Media Impregnated Charcoal

Iodine Removal Efficiency (%) 99 (Operational Requirement) 99 (Operational Requirement)

Quantity of Media (lb) 5800

Depth of Bed (in.) 8 Residence Time for 8 in. bed (sec) 2.0

Static Resistance (in. H 2O) 4.6

g. Standby Cooling Air Fan Type Centrifugal Quantity 1

Drive Direct Capacity (ft 3/min) 200

Static Pressure (in. H 2O) 5 LSCS-UFSAR TABLE 6.5-1 (SHEET 4 OF 4) TABLE 6.5-1 REV. 13 NAME OF EQUIPMENT TYPE, QUANTITY AND NOMINAL CAPACITY (per component)

B. Secondary Containment Isolation Dampers

1. Equipment Numbers 1VQ037, 1VQ038 2VQ037, 2VQ038 1VR04YA&B, 1VR05YA&B 2VR04YA&B, 2VR05YA&B
2. Type Special
3. Quantity 8
4. Operator Air Cylinder
5. Diameter (in.) 72 LSCS-UFSAR TABLE 6.5-2 TABLE 6.5-2 REV. 0 - APRIL 1984 STANDBY GAS TREATMENT SYSTEM EQU1PMENT FAILURE ANALYSIS COMPONENT FAILURE FAILURE DETECTED BY ACTION Primary Fan Motor Burnout, Drive Shaft Break, etc. Flow Monitor - Low-Flow Switch Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down. Electric Heating

Coil Element Overheat High Temperature Protection Circuit on Coil Main Control Board Indication. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down.

Standby Cooling Fan No Startup Results In High Charcoal Adsorber Temperature Temperature Switch If temperature switch trips, then alarm sounds in main control room (Station operator manually actuates deluge valves). Redundant train started after its isolation valves are positioned properly. Operating train is then shut down.

Flow Control Valve Fails Open Flow Monitor - High-Flow Switch Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down.

Flow Control Valve Fails Shut Flow Monitor - Low-Flow Switch Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down. Isolation Valve Fails Open None - Redundant valves or backflow dampers provided as required. Fails Shut Flow Monitor - Low-Pressure Switch Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down. HEPA Filter High Particulate Loading High P Alarms Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down.

Duct Destruction by Equipment Missile or Flailing Pipe Flow Monitor - High-Flow Switch Main Control Board Alarm. Redundant train started after its isolation valves are positioned properly. Operating train is then shut down. Deluge Valve Fails Closed No Detection None required, two valves provided to flood bed.

LSCS-UFSAR 6.6-1 REV. 17 APRIL 2008 6.6 INSERVICE INSPECTION OF ASME CODE CLASS 2 AND 3 COMPONENTS 6.6.1 Components Subject to Examination All ASME Class 2 components (pressure vessels, piping, pumps, and valves) are inservice inspected according to ASME, B&PVC,Section XI, Subsection IWC, with appropriate addendum(s). The main steamlines (four) are inspected from the first outside containment isolation valve to the turbine stop valves. Inspection requirements are the same as for ASME Class 2 components.

All ASME Class 3 components (pressure vessels, piping, and valves) are inservice inspected according to ASME, B&PVC,Section XI, Subsection IWD, with appropriate addendum(s).

6.6.2 Accessibility

The design and arrangement of the ASME Class 2 and ASME Class 3 piping, pump, and valve components have been made acce ssible for inspection and examination as follows: Pipe and Equipment Welds Location and clearance envelopes have been established for inspection and examination. Co ntours and surface finish are acceptable for inspection and examination.

Insulation Removal

Piping or components to be inspected according to the Section XI code which are insulated, have been designed with removable numbered insulation panels.

Shielding Piping or components to be inspected according to the Section XI code and are radiologically shielded have been designed with removable or accessible shields.

6.6.3 Examination Techniques and Procedures Inservice inspection will be in acco rdance with ASME, B&PV Section XI.

6.6.4 Inspection Intervals The initial 10-year inspection program for LaSalle units 1 and 2 was submitted to the NRC on July 13, 1982 and December 21, 1982, respectively. The inservice LSCS-UFSAR 6.6-2 REV. 17 APRIL 2008 inspection program for both units 1 and 2 are based on the requirements of the ASME,Section XI 1980 edition including addenda through winter 1980. The inservice examinations conducted during the second 120 month Inspection Interval will comply with the 1989 Edition of ASME Section XI, except in cases where relief has been granted by the NRC. The inservice examinations conducted during the third 120 month Inspection Interval will comply with the 2001 Edition through the 2003 addenda, including the December of 2003 Erratum of ASME Section XI, except in cases where relief has been granted by the NRC.

6.6.5 Examination Categories and Requirements The inservice inspection categories and requirements for Class 2, and Class 3 components are in agreement with ASME Section XI.

Specific written requests for relief from ASME code requirements determined to be impractical were contained in the initial in service inspection program. Relief from those requirements was granted by the NRC, detailed evaluation is included in Appendix C of NUREG-0519, Supplement No. 5, Safety Evaluation Report related to the operation of LaSalle County Station, Units 1 and 2.

6.6.6 Evaluation of Examination Results The evaluation of Class 2 components ex amination results will comply with the requirements of Section XI.

The repair procedures for Class 2 and 3 components comply with the requirements of Section XI.

6.6.7 System Pressure Tests All Class 2 system pressure testing complies with the criteria of Code Section XI, Article IWC-5000. All Class 3 system pres sure tests comply with the criteria of Article IWD-5000.

6.6.8 Augmented Inservice Inspection to Protect Against Postulated Piping Failures This inspection has been adequately cove red by the requirements of Section XI already adhered to previously.

LSCS-UFSAR 6.7-1 REV. 13 6.7 MAIN STEAM ISOLATION VALVE LEAKAGE CONTROL SYSTEM (MSIV-LCS)

Unit 2 deleted, Unit 1 abandoned in place The Main Steam Isolation Valve Leakage Control System provided originally has been deleted. The valve leakag es are processed by the Isolated Condenser Leakage Treatment Method as discussed in Section 6.8.

LSCS-UFSAR 6.8-1 REV. 13 6.8 Main Steam Isolation Valve - Isolated Condenser Leakage Treatment Method The Main Steam Isolation Valve - Isolated Condenser Leakage Treatment Method (MSIV - ICLTM) (Also called the MSIV Al ternate Treatments Leakage Paths) controls and minimizes the release of fiss ion products which could leak through the closed main steam isolation valves (MSIV's) after a LOCA. The system provides this control by processing valve leakage through the main steamlines, main steamline drains, and the main condenser.

6.8.1 Design Bases 6.8.1.1 Safety Criteria The following general and specific design criteria represent system design, safety, and performance requirements imposed upon the MSIV-ICLTM:

a. The safety function of the main steamlines and main steamline drains are described in LSCS-UFSAR Section 10.3.
b. The safety function of the main condenser is described in LSCS-UFSAR Section 10.4.1.

6.8.1.2 Regulatory Acceptance Criteria The classification of the components and piping of the main steam supply system is listed in Table 3.2-1. All components and piping for the main steam supply system are designed in accordance with the code s and standards listed in Table 3.2-2 for the applicable classification.

The classification of the main condenser is described in LSCS-UFSAR Section 10.4.1.3.

6.8.1.3 Leakage Rate Requirements The MSIV-ICLTM has been incorporated as an integral part of the BWR plant design. The design features employed with this systems are established to reduce the leakage rate of radioactive materials to the environment during a postulated LOCA. Leakage control requirements are imposed upon the MSIV-ICLTM in order to:

a. eliminate the possibility of secondary containment bypass leakage of accident induced radioactive releases, b. allow for higher MSIV leakage limits, and LSCS-UFSAR 6.8-2 REV. 18, APRIL 2010
c. assure reasonable leakage verification test frequencies (once per fuel cycle).

The design and operational requirements imposed upon the MSIV-ICLTM relative to the foregoing criteria are established to:

a. allow MSIV leakage rates up to a total of 400 scfh for all MSIV valves, b. allow a MSIV leakage rate verification testing frequency compatible with the requirements of plant operating technical specifications, and
c. assure and restrict total plant dose impacts below 10 CFR 100 guidelines.

6.8.2 System Description 6.8.2.1 General Description The system provides this control by pr ocessing valve leakage through the main steamlines, main steamline drains, and the main condenser.

6.8.2.2 System Operation (U2 MSIV LCS delete, U1 Abandon-in-place)

With the deletion of the MSIV-LCS, MSIV leakage will pass from the outboard MSIV, through the main steamlines, main steamline drains and into the condenser. The large wetted volume in the main cond enser plates out inorganic iodine and holds up other fission products that esca pe through the MSIVs, limiting release to the environment. This alternate pathwa y is more reliable than the MSIV-LCS since less equipment is employed. The alternate pathway also has a much higher capacity for processing leakage than does the MSIV-LCS, with a capacity of only 100 scfh. In addition, the MSIV-LCS will on ly operate at less than 35 psig reactor vessel steam dome pressure, whereas the alternate pathway is independent of reactor pressure.

To properly align the pathway, in addition to closing the MSIVs and the containment isolation valves, operators will close valves to isolate the leakage pathway from the auxiliary steam supplies. The operating drains will remain open and either one of two startup drains will be opened. All of the remote manually operated valves that need to be moved are powered from Class 1E power supplies. Although these valves and their power supplies (with the exception of the MSIVs) are not maintained as safety-related, design control for all of these valves is maintained with respect to their importance to safety. Appropriate changes to station LSCS-UFSAR 6.8-3 REV. 13 procedures have been made to reflect deletion of the MSIV-LCS and use of the alternate leakage treatment method.

6.8.2.3 Equipment Required The following equipment components are pr ovided to facilitate system operation:

a. piping - process piping is carbon steel throughout;
b. valves - motor-operated, standard closing speeds;
c. main condenser

6.8.3 System Evaluation An evaluation of the capability of the MSIV-ICLTM to prevent or control the release of radioactivity from the main steamlin es during and following a LOCA has been conducted. The results of this evaluation are presented in LaSalle County Nuclear Power Stations Units 1 and 2 Applicatio n for Amendment of Facility Operating Licenses NPF-11 and NPF-18, Appendix A, Technical Specifications, and Exemption to Appendix J of 10CFR50 Regarding Elimin ation of MSIV Leakage Control System and Increased MSIV Leakage Limits , NRC Docket Nos. 50-373 and 50-374. Additionally, Sargent & Lundy performed an evaluation on the piping, condenser and turbine building, to assure they would remain functional during a seismic event to mitigate the radiologically consequenc es of MSIV leakage (Reference Sargent & Lundy Calculation 068078 (EMD), Rev. 2, dated 8/9/95 for Unit 1 and 067927 (EMD), Rev. 2 dated 8/10/95 for Unit 2).

See Section 15.6.5.5 for more informat ion in the dose analysis and dose consequences.

6.8.4 Instrumentation Requirements

The instrumentation necessary for contro l and status indication of the MSIV-ICLTM is designed to function under Seis mic Category I and environmental loading conditions appropriate to its installation with the control circuits designed to satisfy separation criteria. MSIV closed indication is powered from Class 1E power and is maintained as safety-related.

6.8.5 Inspection and Testing Preoperational tests for the main steamlines, main steamline drains, and the main condenser are discussed in LSCS-UFSAR Sections 10.3.4 and 10.4.1.4. No additional testing is required to support this operating mode.

LSCS-UFSAR TABLE 6.8-1 REV. 13 TABLE 6.8-1 DOSE CONSEQUENCES OF MSIV LEAKAGE LEAKAGE 30 DAYS FO LLOWING LOCA-UNIT 1 (100 SCFH per line)

WHOLE BODY DOSE (rem) THYROID DOSE (rem)

Exclusion Area (509 meters) 1.451E-3 3.14E-2 Low Population Zone (6400 meters) 3.3E-2 10.47 LSCS-UFSAR REV. 13

ATTACHMENT 6.A ANNULUS PRESSURIZATION

LSCS-UFSAR 6.A-i REV. 18, APRIL 2010 ATTACHMENT 6.A TABLE OF CONTENTS PAGE 6.A ANNULUS PRESSURIZATION 6.A-1 6.A.1 INTRODUCTION 6.A-1 6.A.2 SHORT-TERM MASS ENERGY RELEASE 6.A-1

6.A.2.1 Instantaneous Guillotine Break 6.A-3 6.A.2.2 Break Opening Flow Rate 6.A-4 6.A.2.3 Combined Break Flow 6.A-5 6.A.2.4 Determination of the Mass Flux, G 6.A-5 6.A.2.5 Biological Shield Wall 6.A-5 6.A.2.6 Comparison of the GE Model with the Henry/Fauske Correlation 6.A-6

6.A.3 LOAD DETERMINATION 6.A-10

6.A.3.1 Acoustic Loads 6.A-10 6.A.3.2 Pressure Loads 6.A-10 6.A.3.3 Jet Loads 6.A-11 6.A.3.4 Dynamic and Seismic Analysis (DYSEA) Computer Program 6.A-12 6.A.4 PRESSURE TO FORCE CONVERSION 6.A-14 6.A.5 SACRIFICIAL SHIELD ANNULUS PRESSURIZATION AND RPV LOADING DATA 6.A-16

6.A.6 JET LOAD FORCES 6.A-18 6.A.7 RECIRCULATION AND FEEDWATER LINE BREAK FORCING FUNCTION 6.A-19 6.A.8 REFERENCES 6.A-20

LSCS-UFSAR 6.A-ii REV. 18, APRIL 2010 ATTACHMENT 6.A LIST OF TABLES NUMBER TITLE 6.A-1 Time History for Postulated Recirculation Suction Pipe Rupture 6.A-2 Acoustic Loading on Reactor Pressure Vessel Shroud 6.A-3 RPV Coordinates of Nodal Points 6.A-4 Maximum Member Forces Due to Annulus Pressurization 6.A-5 Maximum Acceleration Due to Annulus Pressurization 6.A-6 RELAP4 Input Data, Recirculation Line Outlet Break 6.A-7 RELAP4 Input Data, Feedwater Line Break 6.A-8 Force Constants and Load Centers For Recirculation Line Outlet Break 6.A-9 Force Constants and Load Centers For Feedwater Line Break 6.A-10 DYSEA01 Program Input For Jet Load Forces LSCS-UFSAR 6.A-iii REV. 13 ATTACHMENT 6.A LIST OF FIGURES NUMBER TITLE 6.A-1 Safe End Break Location 6.A-2 Break Flow Vs. Time - Feedwater Line Break 6.A-3 Geometry 6.A-4 Wave Speed 6.A-5 Mass Flux, Moody Steady Slip Flow 6.A-6 Break Flow Vs. Time 6.A-7 Nomenclature for Time History Computer Printout 6.A-8 Feedwater Line System Nodalization - Leg EA 6.A-9 Feedwater Line System Nodalization - Leg EB 6.A-10 Recirculation Line System Nodalization 6.A-11 Comparison of the GE and RELAP4/MOD5 Methods -

Feedwater Line Break, Leg EA 6.A-12 Comparison of the GE and RELAP4/MOD5 Methods -

Feedwater Line Break, Leg EB 6.A-13 Comparison of the GE and RELAP4/MOD5 Methods -

Recirculation Line Break, Finite Opening Time 6.A-14 Horizontal Model for Annulus Pressurization 6.A-15 Annulus Pressurization Loading Description 6.A-16 Annular Space Nodalization For Recirculation Line Break 6.A-17 Annular Space Nodalization For Feedwater Line Break

LSCS-UFSAR 6.A-1 REV. 13 6.A ANNULUS PRESSURIZATION 6.A.1 INTRODUCTION Annulus pressurization refers to the load ing on the shield wall and reactor vessel caused by a postulated pipe rupture at the reactor pressure vessel nozzle safe-end to pipe weld. The pipe break assumed is an instantaneous guillotine rupture which allows mass/energy release into the drywell and annular region between the biological shield wall and the reactor pressure vessel (RPV).

The mass and energy released during the postulated pipe rupture cause:

a. A rapid asymmetric decomp ression acoustic loading of the annular region between the vessel and shroud from the pipe break at or beyond the vessel nozzle safe-end weld.
b. A transient asymmetric differential pressure within the annular region between the biological shield wall and the reactor pressure vessel (annulus pressurization).
c. A jet-stream release of the reactor pressure vessel inventory and the impact of the ruptured pipe against the whip restraint attached to the biological shield wall.

The results of the mass and energy release evaluation are then used to produce a dynamic structural analysis (force-time history) of the RPV and shield wall. The force time history output from the dyna mic analysis is subsequently used to compute loads on the reactor components. The following is a more detailed description of the annulus pressurization calculation performed for the LaSalle County Station.

6.A.2 SHORT-TERM MASS ENERGY RELEASE The postulated pipe rupture at the weld between recirculation or feedwater piping and the reactor nozzle safe end leads to a high rate of water and steam mixture into the annulus between the RPV and the shie ld wall. Figure 6.A-1 illustrates the location of this break. Calculation of the mass/energy release is performed using the generic method for short-term mass releases. This method and a sample calculation are described below. Figure 6.A-2 illustrates a typical mass flux vs. time for a feedwater line break.

The purpose of this procedure is to document the method by which short-term mass release rates are calculated. The flow ra tes which could be produced by a primary system line break for the first 5 seconds include the effects of inventory and subcooling. Optionally, credit may be taken for a finite break opening time.

LSCS-UFSAR 6.A-2 REV. 13 ASSUMPTIONS The assumptions are as follows:

a. The initial velocity of the fluid in the pipe is zero. When considering both sides of the break, the effects of initial velocities would tend to cancel out.
b. Constant reservoir pressure.
c. Initial fluid conditions inside the pipe on both sides of the break are similar.
d. Wall thickness of the pipe is small compared to the diameter.
e. Subcompartment pressure

~ 0. f. Mass flux is calculated using the Moody steady slip flow model with subcooling.

g. For steamline breaks, level swell occurs at 1 second after the break with a quality of 7%.

NOMENCLATURE (See Figure 6.A-3)

A BR - Break area.

A L - Minimum cross-sectional area between the vessel and the break. This can be the sum of the areas of parallel flow paths.

C - Sonic velocity (see Figure 6.A-4).

D - Pipe inside diameter at the break location.

F I - Inventory flow multiplier.

F I = 0.75 for saturated steam.

FI = 0.50 for liquid and saturated steam-liquid mixtures.

g c - Proportionality constant (=32.17 2 lbm-ft/lbf-sec 2). G - Mass flux.

LSCS-UFSAR 6.A-3 REV. 13 G C - Maximum mass flux (see Figure 6.A-5).

h O - Reservoir or vessel enthalpy.

h P - Initial enthalpy of the fluid in the pipe.

h 7 - Enthalpy at P O and a quality of 7%.

L I - Inventory length. The distance between the break and the nearest area increase of A L whichever is less.

M - Mass flow rate.

I M - Mass flow rate during the inventory period.

P O - Reservoir or vessel pressure.

PSAT - Saturation pressure for liquid with an enthalpy of h P. t - Time.

t I - Length of the inventory period.

v - Specific volume of the fluid initially in the pipe.

V I - Volume of the pipe between the break and A L . X - Separation distance of the break.

6.A.2.1 Instantaneous Guillotine Break

The following method should be applied to each side of the break and the results summed to determine the total flow.

LSCS-UFSAR 6.A-4 REV. 14, APRIL 2002 c L 2 t F A A If I I' I BR L=> v F G A V t F A A If I BR I I' I BR L=<Inventory Period Prior to a pipe break, the fluid in the pipe is moving at a relatively low velocity. After the break occurs, a finite time is required to accelerate the fluid to steady-state velocities. The length of this time period is conservatively estimated as follows:

a. (6.A-1) b.

(6.A-2) where G is calculated as shown in Subsecti on 6.A.2.4 for a large separation distance and t < t I.

During this time period, the mass flow rate is calculated as Steady-State Period Following the inventory period, the flow is assumed to be choked at the limiting cross-sectional flow area. For t I < t < 5.0 seconds, (6.A-4) 6.A.2.2 Break Opening Flow Rate See Table 6.A-1 for the pipe displacement time history for postulated recirculation suction pipe rupture and Figure 6.A-7 for the nomenclature used.

Inventory Period The inventory period is determined as des cribed in Subsection 6.A.2.1. The flow rate as a function of pipe separation distance is given by where G is obtained by using the methods of Subsection 6.A.2.4 (a or b).

I F BR A G I M= G L A M=X D G M=(6.A-3) (6.A-5)

LSCS-UFSAR 6.A-5 REV. 13 Determining Flow Rate Following the inventory period, equation 6.A-5 is used to deter mine the flow rate where the mass flux, G, is determined from Subsection 6.A.2.4 (a, c, or d).

6.A.2.3 Combined Break Flow To determine the total flow rate released from the break, the results of Subsections 6.A.2.1 and 6.A.2.2 are compared and whichever produces the smallest flow rate at any time is used (see Figure 6.A-6). Both methods produce maximum flow rates based on different limiting areas. The transfer from one curve to the other represents a change in the point where the flow is choked.

6.A.2.4 Determination of the Mass Flux, G Depending on the time period, fluid conditions, and break separation distance, the mass flux is determined as follows:

a. If X < X B , b. If X > X B and t < t I G = G c (P o , h p) from Figure 6.A-5
c. If X > X B and t > t I G = G c (P o , h o) from Figure 6.A-5
d. If the break is a steamline and T > 1.0, level swell occurs.

G = G c (P o , h 7) from Figure 6.A-5 Note that for complete break separation (Subsection 6.A.2.1), X is always greater than X B, and for saturated water, X B is equal to zero.

6.A.2.5 Biological Shield Wall For the purpose of analyzing the biological shield wall pressurization, credit may be taken for flow which escapes through the wall penetration. If the initial break location is in the annulus region between the wall and the vessel, no flow is assumed to escape through the penetration. If, however, it is located within the penetration itself, some of the flow may be assumed to escape. It is recommended

()()2 D o P SAT P 1 B X= v o P c 2g G= (6.A-6)

LSCS-UFSAR 6.A-6 REV. 13 that the fraction of the flow which escapes be calculated based on the ratio of the minimum annular flow area between the pe netration and pipe surface and between the penetration and pipe surface and between the penetration and the safe-end nozzle. 6.A.2.6 Comparison of the GE model with the Henry/Fauske Correlation The GE methodology for calculating the mass energy release from a recirculation line break which results in an annulus pressurization event was provided the NRC's Mr. Denwood F. Ross, Assistant Director for Reactor Safety, via GE letter dated May 2, 1978, from Mr. E. D. Fuller of BWR Licensing. This methodology was used in the adequacy assessment made for LSCS.

The definition of the annulus pressurization is given in the introduction (Subsection 6.A.1). A description of the time aspect s of the calculated mass and energy flow rates followed by a description of the modeling for the feedwater line and separately for the recirculation line is provided below. A comparison is then made between GE's analytical method and the method used in RELAP4/MOD5. Finally, both graphical and numerical results of this comp arison are provided to substantiate the conclusion that the resulting break flows using the GE methods are much more conservative than those predicted by the use of RELAP for the LaSalle plant.

Timing Aspects of Mass and Energy Flow Rates The GE method for calculating the short-term mass/energy release assumes that the overall time for mass release may be divided into two periods, the inventory period and the quasi-steady period. The inventory period is defined as the time required to accelerate the pipe fluid to steady-state velocities, at which time the flow is assumed to choke at minimum flow cross sections. During this time, the mass flux is based on initial thermodynamic conditions exis ting within the pipe. In the quasi-steady period, during which the flow is choked, the mass flux is based on thermodynamic conditions upstream from th e choke points. For both time periods the mass flux is determined from a graph of critical mass flux versus enthalpy, as calculated by the Moody Slip Flow Method. Each side of the break is analyzed separately and the results summed to give the total mass release rate.

Method for Feedwater Line Modeling The feedwater system for LaSalle County St ation consists of the pumps, heaters, valves, and piping necessary for the tran smission of hotwell condensate to the reactor vessel as part of the closed cycl e cooling loop. LSCS has three feedwater pumps, two steam- driven and one electric-driven. During normal operation, the electric pump is in standby. The flow passes through a complex series of pipes and components from the feedwater pumps to the reactor vessel.

LSCS-UFSAR 6.A-7 REV. 13 The break location for the feedwater line break is the safe-end to the pipe weld housed within the vessel to shield wa ll subcompartment. For the feedwater line break, instantaneous break opening is assumed. Flow for the vessel side passes through the feedwater nozzles of the broken line and out the break. Flow from the system side passes through the feedwater piping network and out the break.

The nodalization of the feedwater system is shown in Figures 6.A-8 and 6.A-9. A series of 24 modes was selected after sensitivity studies were completed which demonstrated that a 24-node model was conservative relative to higher-noded systems.

The broken feedwater leg to be analyzed was chosen by multiple RELAP runs to determine the limiting break location. The critical assumptions in the analysis are as follows:

a. The feedwater pumps are simulated as (constant) mass flow sources. b. The reactor pressure vessel (RPV) is an infinite reservoir at constant temperature and pressure.
c. The temperature of the pump-side hydraulic network remains constant.
d. Appropriate sections of the hydraulic network are combined by means of "Ohm's Law" expressions for series and parallel circuits, assuming constant fanning friction actions.
e. The RPV thermodynamics stat e is subcooled at the prevailing temperature in the lower plenum (532

° F). The break is modeled as an instantaneous guillotine pipe break with complete pipe offset. Before the break occurs, a fully open valve connects, Volumes 18 and 19. Closed valves connect those volumes to Volume 1, an infinite sink at constant pressure and temperature (atm ospheric conditions). The break is initiated at time zero by closure of the valve between Volumes 18 and 19 simultaneous opening of the valves to Volume 1.

Method of Recirculation Line Modeling The recirculation system for LaSalle County Station is similar to the recirculation system of other BWR's. Flow is taken from the lower jet pump diffuser region, passed through 21-inch lines to a constant-speed pump, and then through a flow control valve to a header which feeds flow to five risers which provide flow to two jet pump nozzles each.

LSCS-UFSAR 6.A-8 REV. 13 The nodalization for the recirculation line leak is shown in Fi gure 6.A-10. The system has been modeled using 21 nodes. The break is located at the vessel nozzle safe-end to pipe weld on the recirculation pump suction side. The type of break considered here has a finite break opening time. For this case the break opening is complete after 30 milliseconds, at which time the pipe offset longitudinal distance is 5.8 inches. The break area is modeled as the surface area of an imaginary volume having a length of 5.8 inches and a diameter equal to that of the recirculation pipe ID. This volume (#18) is connected by a valve (Type 3) to an infinite reservoir (volume #19), and also by valves (Type 2) to the vessel side volume (#27) and pump side volume (#21). Both valves (Type 1) also connect Volumes 17 and 21. It is normally open before the break, and at the initiation of the break, closes at the same rate as the other valves open. The sum of the areas of the Type 2 valves equals the pipe area.

This network of valves best represents the break with finite opening time. Valves of Type 2 are opened at the same rate as Type 3 to ensure that choking occurs at Junctions 21 and/or 22. Junction 23 (having valve Type 3) is in reality a fluid surface, and choking cannot physically occur there. Choking must at least occur at Junctions 21 and/or 22, where the fluid is constrained by the pipe diameter.

Other assumptions in th e analysis include:

a. Negligible effects of core reactor kinetics on rated heat transfer to the core volume (Volume 2).
b. Constant flow of steam from the steam dome (Vol ume 5) at rated conditions.
c. Constant flow of feedwater at rated conditions.
d. Recirculation pumps trip at the time zero and are modeled via pump characteristic curves for coastdown.
e. Jet pump hydraulics were modeled as one equivalent pump per recirculation loop.

Comparison of General Electr ic Analysis to RELAP4/MOD5 For the annulus pressurization event, th e NRC has questioned General Electric's method for computing mass and energy flow rates following a postulated LOCA from long lines containing subcooled fluid. A program was developed to expedite the licensing of the LaSalle County Station to perform RELAP analyses using appropriate assumptions and to compare the results with those obtained using General Electric's method.

LSCS-UFSAR 6.A-9 REV. 13 RELAP4/MOD5 is a general computer prog ram which can be used to analyze the thermal hydraulic transient behavior of a water- cooled nuclear reactor subjected to postulated accidents such as loss-of-coolant accidents. The program simultaneously solves the fluid flow, heat transfer, the reactor kinetics equations describing the behavior of the reactor.

Numerical input data is utilized to describe the initial conditions and geometry of the system being analyzed. This data includes fluid volume, geometry, pump characteristics, power generation, heat exchanger properties, and nodalization of fluid flow paths. Once the system has been described with initial flow, pressure, temperature, and power level boundary co nditions, transients such as loss-of-coolant accident can be simulated by control action inputs. RELAP then computes fluid conditions such as flow, pressure, mass inventory an d quality as a function of time. For the brief transients considered here (t 0.5 seconds), appreciable simplification of the overall thermal-hydraulic system, including the reactor pressure vessel, is justified owing to the relatively longer time constants which apply for heat transfer.

Brief summaries of the modeling approaches for feedwater and recirculation line breaks are given below.

The assumptions applied to th ese analyses are as follows:

a. Feedwater line:
1. LaSalle RELAP deck as basis.
2. Henry-Fauske-Moody flow model is used.
3. Instant break opening.
4. Mass flux terms between ve ssel and break (short side) are eliminated.
b. Recirculation line:
1. LaSalle RELAP deck as basis.
2. Finite break opening time is allowed for.
3. Henry-Fauske-Moody flow model is used.
4. Momentum flux terms in RELAP between vessel and break (short side) are eliminated.

Results of the Analysis

LSCS-UFSAR 6.A-10 REV. 13 The resulting break flows using the GE methods are much more conservative than those obtained by the use of RELAP. This is indicated graphically in Figures 6.A-11 through 6.A-13.

Conclusions The mass release result for the GE mass energy release method and the RELAP4/Mod 5 calculations are compared in Figures 6.A-11 through 6.A-13 for the postulated feedwater line break and reci rculation line break respectively. The analyses show that the GE method is conservative relative to RELAP 4/Mod 5 for both cases. The ration (r) of the GE method flow rates to those from RELAP/MOD5 is as follows:

Break Location r(t = 0.1 sec) r(t = 0.5 sec) Feedwater (Leg EA) 2.300 1.70 Feedwater (Leg EB) 2.200 1.60 Recirculation Line 1.065 1.21 6.A.3 LOAD DETERMINATION 6.A.3.1 Acoustic Loads Because the boiling water reactor (BWR) is a two-phase system that operates at or close to saturation pressure (1000 psi), th e differential pressure across the reactor shroud is of short duration, and the BWR system is not subjected to a significant shock-type load with respect to structural supports. This short- duration acoustic load is confined to a bending moment and shear force on the reactor pressure vessel and reactor shroud support. Typical results of the integrated force acting on the reactor pressure vessel shroud are given in Table 6.A-2.

6.A.3.2 Pressure Loads The pressure responses of the RPV-shield wall annulus for a recirculation suction line and a feedwater line were investigated using the RELAP4 computer code. An asymmetric model using several nodes and flow paths was developed for the analysis of the recirculation and feedwater line breaks. Further description of these analytical models and detailed discussion of the analyses may be found in Section 6.2.

The pressure histories generated by the RELAP4 code were in turn used to calculate the loads on the sacrificial wall and the reactor pressure vessel. The

annulus was divided into seven zones and an eighth-order Fourier fit to the output LSCS-UFSAR 6.A-11 REV. 14, APRIL 2002 pressure histories made for each zone to produce the Fourier coefficients required for the structural analysis of the shield wall. The specific loading data consisted of the time-pressure (psia) hist ories for each node within the annulus. Time-force histories representing the resultant loads on the RPV for each node through its geometric center were generated by taking the product of the node pressure and its "effective" surface area.

A sample pressure-to-force ca lculation is shown in Subsection 6.A.4. Subsection 6.A.5 shows the nodalization schemes and pressure areas used in this calculation. The time-force histories (forcing functions) calculated at each nodal point for both a recirculation and a feedwater line break are shown in Subsection 6.A.7. The nodal points are illustrated in Figure 6.A-14.

6.A.3.3 Jet Loads To address structural loads on the vessel and internals completely, jet thrust, jet impingement, and pipe whip restraint loads must be considered in conjunction with the above mentioned pressure loads. Jet thrust refers to the vessel reaction force with results as the jet stream of liquid is released from the break. Jet impingement refers to the jet stream force which leaves the broken pipe and impacts the vessel.

The pipe whip restraint load is the force which results when the energy-absorbing pipe whip restraint restricts the pipe separation to less than one full pipe diameter. This restricted separation is accounted for as a finite break opening time in the mass/energy release calculation. These je t loads are calculated as described in ANSI 176 (draft), "Design Ba sis For Protection Of Nuclear Power Plants Against Effects Of Postulated Pipe Ruptures", January 1977.

The jet load forces used in this analysis are shown in Subsection 6.A.6. Although these values have been calculated for a re circulation line break only, they are also conservatively used for the feedwater load evaluation. This is conservative because the calculation of these jet effects depends largely on the area of the break, and the recirculation line is about 2.5 times larger in area. Figure 6.A-15 illustrates the location of the pressure loads and jet loads with respect to the RPV and shield wall.

The pressure loads and jet loads describe d above are then combined to perform a structural dynamic analysis. Both of these loads are appropriately distributed along a horizontal beam model, which is shown in Figure 6.A-14. The vessel coordinates of these nodal points are described in Table 6.A-3.

The force time histories are then applied to a composite lumped- mass model of the pedestal, shield wall, and a detailed repres entation of the reactor pressure vessel and internals. The DYSEA01 computer program is used for this analysis. This computer program is described in Subsec tion 6.A.3.4. The analysis produces acceleration time histories at all nodes for use in evaluating the reactor pressure vessel and internal components. Response spectra at all nodes are also computed.

LSCS-UFSAR 6.A-12 REV. 13 The peak loading on the major components used to establish the adequacy of the component design is shown in Tables 6.A-4 and 6.A-5.

6.A.3.4 Dynamic and Seismic Anal ysis (DYSEA) Computer Program The DYSEA (Dynamic and Seismic Analysis) program is a GE proprietary program developed specifically for seismic and dynamic analysis of RPV and internals/building systems. It calculates the dynamic response of linear structural systems by either temporal modal superposition or response spectrum method. Fluid- structure interaction effect in the RPV is taken into account by way of hydrodynamic mass.

The DYSEA program was based on the program SAP-IV with added capability to

handle the hydrodynamic mass effect. St ructural stiffness and mass matrices are formulated similar to SAP-IV. Solution is obtained in the time domain by calculating the dynamic response mode by mode. Time integration is performed by using Newmark's -method. Response spectrum solution is also available as an option. Program Version and Computer

The DYSEA version now operating on the Honeywell 6000 computer of GE, Nuclear Energy Systems Division, was developed at GE by modifying the SAP-IV program.

Capability was added to handle the hydrodyn amic mass effect due to fluid-structure interaction in the reactor. The progra m can handle three-dimensional dynamic problems with beam, trusses, and springs. Both acceleration time histories and response spectra may be used as input.

History of Use The DYSEA program was developed in the su mmer of 1976. It has been adopted as a standard production program since 1977 an d it has been used extensively in all dynamic and seismic analysis of the RPV and internals/building systems.

Extent of Application The current version of DYSEA has been used in all dynamic and seismic analysis since its development. Results from test problems were found to be in close agreement with those obtained from either verified programs or analytic solutions.

LSCS-UFSAR 6.A-13 REV. 13 Test Problems Problem 1:

The first test problem involves finding the eigenvalues and eigenvectors from the following characteristic equation:

(2 [M]-[K]) {x} = 0 where is the circular frequency, x is th e eigenvector, and [K] and [M] are the stiffness and the mass matrices given by:

(6.A-8) The analytical solution and the solution from DYSEA are:

a) Eigenvalues i: i DYSEA SOLUTION ANALYTIC SOLUTION 1 5.7835 5.7837 2 30.4889 30.4878 3 75.0493 75.0751

[]

=2 25 4-1 Symmetric 2 4 2 q 4 1 2 q 4 2 4 2 4 1 M[]

+++=4 2 25 1 Symmetric 15 4 2 g 1 q 5 3 4 2 1 K (6.A-7)

LSCS-UFSAR 6.A-14 REV. 13 b) Eigenvectors i: 1. DYSEA SOLUTION ANALYTIC SOLUTION

0.0319

0271.2 0666.0 0072.0 2105.1 5536.1 0319.0 000.1 000.1 000.1 027.2 0666.0 0072.0 211.1 554.1 0319.0 000.1 000.1 000.1 Problem 2: The second test problem compares the dynamic responses of the reactor pressure vessel, internals and reactor building subjected to earthquake ground motion.

The mathematical model of the reactor pressure vessel, internals and reactor building is given in Figure B-1. The inputs in the form of ground spectra are applied at the basement level. Response spectr um analysis was used in the analysis.

Natural frequencies of the system and the maximum responses at key locations have been calculated by both DYSEA and SAMIS. Result comparison are given in B-1 and B-2. It can be seen that the results calculated by DYSEA agree closely with those obtained by SAMIS.

6.A.4 PRESSURE TO FORCE CONVERSION The RELAP4 pressure distribution output is converted to equivalent forces which are input into the DYSEA01 computer progra

m. Each pressure is represented by a force acting normal to the RPV or shield wall at the center of the given pressure surface area. These forces are then converte d into resulting forces (x component) acting on the respective DYSEA01 RPV beam nodal points. Mathematically, this is described as:

F R = PA cos where: F R = resultant force (lb), P = RELAP4 node pressure (psia), A = RELAP4 node surface area (in 2 ), and = Component angle.

LSCS-UFSAR 6.A-15 REV. 18, APRIL 2010 The results of these calculations are summarized in Table 6.A-4.

As an example, the pressure to force conv ersion at DYSEA01 node points 31 and 32 is shown below:

Time = 0.0800 seconds NODE ELEV (inches) PRESSURE (lb/in 2) AREA* (in 2) ANGLE (degrees) FORCE (lb) 6 1089.14 43.61 5828.44 15 245516 7 1089.14 35.34 5828.44 45 145660 8 1089.14 39.24 5828.44 75 59188 9 1089.14 41.40 8617.79 112.5

-136539 10 1089.14 39.99 8617.79 157.5

-318367 - 4543

  • See Table 6.A-8

For 360°, the resultant force is 2 times 4543 lb or an inward (positive) force of 9086 lb.

Since DYSEA nodal points 31 and 32 are at Elevations 1065.2 inches and 1125.7 inches respectively, the RELAP4 pressure

/force at Elevation 1089.14 inches is distributed accordingly.

Consequently:

F 31 = 1125.7 - 1089.14 (9086) = 5491 lb, and 1125.7 - 1065.2 F 32 = 1065.2 - 1089.14 (9086) = 3595 lb.

1065.2 - 1125.7 These values can be compared to the co mputer-calculated DYSEA01 results, which are 5832.6 lb and 3252.7 lb respectively (Reference 1).

In the matrix displacement method of stru ctural analysis, externally applied nodal forces and moments are required to produce nodal displacements equivalent to those that would be produced by forces or pressures applied between nodes. GE LSCS-UFSAR 6.A-16 REV. 13 considers the external moment effects for La Salle AP to be negligible because of the close nodal spacing of the LaSalle RPV mathematical model.

6.A.5 SACRIFICIAL SHIELD, ANNULUS PRESSURIZATION, AND RPV LOADING DATA

This subsection provides a brief descri ption of the analyses performed and the nodalization schemes, force constants, and load centers for the recirculation and feedwater line breaks. These data are used as input to the pressure to force conversion calculation.

The pressure responses of the RPV-sacrificial shield wall annulus to postulated pipe breaks at the RPV nozzle safe-end to pipe weld in a recirculation outlet line and a feedwater line were investigated using th e RELAP4 computer code. Throughout the analyses the following assumptions were made:

a. RPV thermal insulation displaces to the shield wall while retaining its original volume and leaving its support structure in place.
b. Insulation above the shield wall yields to elevated pressures and blows out into the drywell allowing venting of annulus at the summit of the shield wall.
c. sacrificial shield penetration doors remain closed, allowing for limited venting of the annulus through all nozzle penetrations.

The nodalization schemes for both studies remain consistent with the guidelines cited above, with the exception of the region directly above the break, where it was anticipated that a finer mesh would be necessary to properly account for the highly localized pressure gradients expected there (see Figures 6.A-16 and 6.A-17). The final nodalization was determined on the ba sis of available sensitivity studies for similar analyses.

The mass and energy release rates were derived with the methods outlined in Subsection 6.A.2. The blowdown rates for the recirculation outlet line break analysis account for actual pipe displa cement, while those for the feedwater line reflect an assumption of instantaneous pipe displacement (see RELAP4 input listings, Tables 6.A-6 and 6.A-7).

The specific loading data compiled for th e NSSS adequacy evaluation for postulated pipe breaks within the annulus consists of the time-pressure history (psia) and two time-force (lbf) histories for each node within the annulus. The latter two histories represent integrated forces acting through the center of each node on the RPV and the sacrificial shield wall respectively. The time-force histories were generated by LSCS-UFSAR 6.A-17 REV. 13 taking the product of the node pressure and a predetermined constant, or ss, which accounts for the curved surface of the RPV and the sacrificial shield respectively (see Tables 6.A-8 and 6.A-9). The two loadin g histories, one for the RPV and one for the shield wall, are defined below.

(6.A-9)

P i v Where: F v i nodal resultant force on RPV (lbf), P i node absolute pressure (psia), i node height (inches), R v RPV radius (inches), azimuthal width of node (degrees), and D p j pipe OD (in.). 4 2 j p D i P j - 0d cos 2 2 v R i i P i v F+=4 2 j p D j () i P - 2 sin v R i 2 i P

LSCS-UFSAR 6.A-18 REV. 14, APRIL 2002 (6.A-10) = P i ss Where: F ss i nodal resultant force on shield wall (lbf), P i node absolute pressure (psia), i node height (inches), R ss shield wall inner radius (inches), azimuthal width of node (degrees), D ss j penetration ID (inches), and proportionality factor 6.A.6 JET LOAD FORCES This subsection provides the jet load forces which result from pipe separation during the postulated accident. The pipe whip schematic is shown in Figure 6.A-7, and the resulting loads are listed in Table 6.A-1.

These loads are applied to the appropriate nodal points for input to the DYSEA01 computer program. The DYSEA01 progra m input is provided in Table 6.A-10.

4 j ss 2 D P - d cos R P i j 2 2 ss i i i s F=+ s () u i P - 2 sin ss R i 2 i P = 4 j ss 2 D j ()2 2 sin=

2 360 LSCS-UFSAR 6.A-19 REV. 18, APRIL 2010 6.A.7 RECIRCULATION AND FEEDWATER LINE BREAK FORCING FUNCTION The time force histories provided in Refere nce 1 are those values converted from the time-pressure histories which were calculated with the RELAP4 computer program.

These time forces histories are used as input to the DYSEA01 computer program.

LSCS-UFSAR 6.A-20 REV. 18, APRIL 2010 6.A.8 REFERENCES

1. Calculation NSLD 3C7-0477-002, Sacrificial Shield Annulus Pressurization and Reactor Pressure Vessel Loading Data for General Electric NSSS Adequacy Evaluation, Rev. 000A.

LSCS-UFSAR REV. 13 ATTACHMENT 6.B RECIRCULATION SYSTEM SINGLE-LOOP OPERATION

LSCS-UFSAR 6.B-i REV. 15, APRIL 2004 ATTACHMENT 6.B TABLE OF CONTENTS PAGE 6.B RECIRCULATION SYSTEM SINGLE-LOOP OPERATION 6.B-1 6.B.1 INTRODUCTION AND

SUMMARY

6.B-1 6.B.1.1 GE Analysis 6.B.1.2 SPC Analysis 6.B.2 MCPR FUEL CLADDING INTEGRITY SAFETY LIMITS 6.B-1 6.B.2.1 Core Flow Uncertainty 6.B-1 6.B.2.2 Core Flow Measurement During Single-Loop Operation 6.B-1 6.B.2.3 Core Flow Uncertainty Analysis 6.B-2 6.B.2.4 TIP Reading Uncertainty 6.B-3

6.B.3 MCPR OPERATING LIMIT 6.B-4 6.B.3.1 Abnormal Operational Transients 6.B-4 6.B.3.2 Feedwater Controller Failure - Maximum Demand 6.B-5 6.B.3.2.1 Identification of Causes and Frequency Classification 6.B-5 6.B.3.2.2 Sequence of Events and Systems Operation 6.B-5 6.B.3.2.3 Effect of Single Failures and Operator Errors 6.B-6 6.B.3.2.4 Core and System Performance 6.B-6 6.B.3.2.5 Barrier Performance 6.B-7 6.B.3.2.6 Radiological Consequences 6.B-7 6.B.3.3 Generator Load Rejection Without Bypass with RPT 6.B-8 6.B.3.3.1 Identification of Causes and Frequency Classification 6.B-8 6.B.3.3.2 Sequence of Events and System Operation 6.B-8 6.B.3.3.3 Results 6.B-9 6.B.3.3.4 Barrier Performance 6.B-10 6.B.3.3.5 Radiological Consequences 6.B-10 6.B.3.4 Recirculation Pump Seizure Accident 6.B-10 6.B.3.4.1 Identification of Causes and Frequency Classification 6.B-10 6.B.3.4.2 Sequence of Events and Systems Operations 6.B-10 6.B.3.4.3 Systems Operation 6.B-11 6.B.3.4.4 Core and System Performance 6.B-11 6.B.3.4.5 Results 6.B-11 6.B.3.4.6 Barrier Performance 6.B-12 6.B.3.4.7 Radiological Consequences 6.B-12 LSCS-UFSAR 6.B-ii REV. 15, APRIL 2004 6.B.3.5 Summary and Conclusions 6.B-12 6.B.4 OPERATING MCPR LIMIT 6.B-12 6.B.5 STABILITY ANALYSIS 6.B-14 6.B.6 LOSS-OF-COOLANT ACCIDENT ANALYSIS 6.B-14 6.B.6.1 Break Spectrum Analysis 6.B-14 6.B.6.2 Single-Loop MAPLHGR Determination 6.B-14 6.B.6.3 Small Break Peak Cladding Temperature 6.B-15 6.B.7 REFERENCES 6.B-16 LSCS-UFSAR 6.B-iii REV. 13 ATTACHMENT 6.B LIST OF TABLES NUMBER TITLE 6.B-1 Input Parameters and Initial Conditions for Transients and Accidents (Analysis of Initial Core) 6.B-2 Sequence of Events for Figure 6.B-3 (Typical) 6.B-3 Sequence of Events for Figure 6.B-4 (Typical) 6.B-4 Sequence of Events for Figure 6.B-5 (Typical, GE) 6.B-5 Summary of Event Results (Typical)

LSCS-UFSAR 6.B-iv REV. 13 ATTACHMENT 6.B LIST OF FIGURES NUMBER TITLE 6.B-1 Illustration of Single Recirculation Loop Operation Flows 6.B-2 Main Turbine Trip With Bypass Manual Flow Control (Typical) 6.B-3 Feedwater CF With One-Pump Operation (Typical) 6.B-4 Load Rejection With One Pump Operation 6.B-5 Seizure of One Recirculation Pump (Typical) 6.B-6 Decay Ratio vs. Power Curve for Two-Loop and Single-Loop Operation (Typical, GE) 6.B-7 Uncovered Time vs. Break Area - LSCS Units 1 and 2 Suction Break LPCS/DG Failure

LSCS-UFSAR TABLE 6.B-1 (SHEET 1 OF 2)

TABLE 6.B-1 REV. 13 INPUT PARAMETERS AND INITIAL CONDITIONS FOR ANALYSIS OF INITIAL CORE TRANSIENTS AND ACCIDENTS FOR SINGLE-LOOP OPERATION (INITIAL CORE VALUES)**

1. Thermal Power Level, Analysis Value, % NBR 78 2. Steam Flow, lb/h 10.71 x 10 6 3. Core Flow, lb/h 68.26 x 10 6 4. Feedwater Flow Rate, lb/sec 2976
5. Feedwater Enthalpy, Btu/lb 367.3
6. Vessel Dome Pressure, psig 1001
7. Vessel Core Pressure, psig 1006
8. Turbine Bypass Capacity, % NBR 25
9. Core Coolant Inlet Enthalpy, Btu/lb 516.8
10. Turbine Inlet Pressure, psig 969.3
11. Fuel Lattice 8 x 8
12. Core Average Gap Conductance, Btu/sec-ft 2-°F 0.1662 13. Core Leakage Flow, % 12 14. Required MCPR Operating Limit 1.41
  • 15. MCPR Safety Limit 1.06 16. Doppler Coefficient, -¢/°F Nominal EOC-1 Analysis Data 0.221 0.221 17. Void Coefficient, -¢/% Voids Nominal EOC-1 Analysis Data for Power Increase Events Analysis Data for Power Increase Events 7.429 12.63 7.01 18. Core Average Rated Void Fraction, % 0.414 19. Scram Reactivity, Analysis Data FSAR Figure 15.0-2 20. Control Rod Drive Speed, position versus time FSAR Figure 15.0-2
  • Dual-pump operation operating limit for 63%

core flow, obtained by applying K f-curve to operating limit CPR at rated condition (1.24).

    • For cycle specific inputs, see the transient analysis input parameters.

LSCS-UFSAR TABLE 6.B-1 (SHEET 2 OF 2)

TABLE 6.B-1 REV. 13 (INITIAL CORE VALUES)

21. Jet Pump M Ratio 3.20 22. Safety/Relief Valve Capacity, % NBR at 1165 psig Manufacturer Quantity Installed 111.5 Crosby 18 23. Relief Function Delay, sec 0.1 24. Relief Function Response, sec 0.1 25. Setpoints for Safety/Relief Valves Safety Function, psig

Relief Function, psig 1150, 1175, 1185, 1195, 1205 1076, 1086, 1096, 1106, 1116

26. Number of Valve Groupings Simulated Safety Function, No. Relief Function, No.

5 5 27. Vessel Level Trips, Inches above Steam Dryer Skirt Bottom (Instrument Zero)

Level 8 - (L8)

Level 3 - (L3) Level 2 - (L2) 55.5 12.5 -50 28. RPT Delay, sec 0.14 29. RPT Inertia Time Constant for Analysis, sec 6.0

LSCS-UFSAR TABLE 6.B-2 TABLE 6.B-2 REV. 13 SEQUENCE OF EVENTS FOR FIGURE 6.B-3 (INITIAL CORE RESULTS)

TIME (sec) EVENT 0 Initiate simulated failure of 160% upper limit on feedwater flow. 5.46 L8 vessel level setpoint trips main turbine and feedwater pumps.

5.47 Reactor scram trip actuated from main turbine stop valve position switches.

5.47 Recirculation pump (RPT) actuated by turbine stop valve position switches.

5.57 Main turbine stop valves closed and main turbine bypass valves start to open. 8.01, 8.29 Relief valves actuated (groups 1, 2). 11.67, 12.23 Relief valves close (groups 2, 1). 29.32 Main turbine bypass valves closed.

48.35 Main turbine bypass valves start to open.

LSCS-UFSAR TABLE 6.B-3 TABLE 6.B-3 REV. 13 SEQUENCE OF EVENTS FOR FIGURE 6.B-4 (INITIAL CORE RESULTS)

TIME (sec) EVENT -0.015 (approx) Turbine-generator dete ction of loss of electrical load 0 Turbine-generator power load unb alance (PLU) devices trip to initiate turbine control valve fast closure 0 Turbine bypass valves fail to operate 0 Fast control valve closure (FCV) initiates scram trip 0 Fast control valve closure (FCV) initiates recirculation pump trip (RPT) 0.039 Turbine control valves closed 0.14 Recirculation pump motor circuit breakers open, causing decrease in core flow to natural circulation 1.98, 2.12, 2.27, 2.45, 2.74 Relief valves actuated (groups 1, 2, 3, 4, 5) 4.58, 4.91, 5.20 (est) Relief valves close (groups 5, 4, 3) 5.30 Vessel level reaches L8 setpoint, feed water pumps tripped (not simulated) 5.50, 5.84 (est) Relief valves close (groups 2, 1) 12.00 Relief valves actuated (group 1) 19.0 (est) Relief valves close (group 1) 33 2 Relief valves actuated (group 1) 38.0 (est) Relief valves close (group 1)

LSCS-UFSAR TABLE 6.B-4 TABLE 6.B-4 REV. 13 SEQUENCE OF EVENTS FOR FIGURE 6.B-5 (INITIAL CORE RESULTS)

TIME (sec) EVENT 0 Single pump seizure was initiated, core flow decreases to natural recirculation 1.23 Reverse flow ceases in the idle loop 4.93 High vessel water level (L8) trip initiates main turbine trip 4.93 High vessel water level (L8) trip initiates feedwater turbine trip 4.93 Main turbine trip initiates bypass operation

4.96 Main turbine valves reach 90% open position and initiate reactor scram trip 5.03 Turbine stop valves closed and turbine bypass valves start to open to regulate pressure 10.0 (est) Turbine bypass valves start to close 25.1 Turbine bypass valves closed 38.6 Turbine bypass valves reopen on pressure increase at turbine inlet

LSCS-UFSAR TABLE 6.B-5 TABLE 6.B-5 REV. 13

SUMMARY

OF EVENT RESULTS SINGLE RECIRCULATION LOOP OPERATION (Typical)

PARAGRAPH F IGURE DESCRIPTION M AXIMUM N EUTRON F LOW (% NBR) M AXIMUM D OME PRESSURE (psig) M AXIMUM VESSEL PRESSURE (psig) M AXIMUM S TEAMLINE PRESSURE (psig) M AXIMUM C ORE AVERAGE SURFACE HEAT F LUX (% of Initial)

MCPR FREQUENCY* C ATEGORY 6.B.3.2 6.B-3 Feedwater flow Controller Failure (Maximum Demand) 119.3 1112 1126 1103 108.8 1.26 a 6.B.3.3 6.B-4 Generator Load Rejection 135.6 1138 1153 1128 103.5 1.29 b 6.B.3.4 6.B-5 Seizure of Active Recirculation Pump 78.0 1021 1031 1018 100.0 1.17 c

__________________________

  • a = incident of moderate frequency; b = infrequent incident; c = limiting faults

" LSCS-UFSAR POINT OF CRITICAL FLOW fA*RECIRCULATION LINE B.CLEANUP LINE C.COMBINED AREA OF ALL JET PUMP NOZZLES ASSOCIATED WITH TEE BROKEN LOOP tn*BOTTOM HEAD DRAIN REACTOR VESS£L RECI RCU LA TION LOOP RECIRCULATION LOOP TO REACTOR WATER CLEANuP SYSTEM SCHEMATIC SHOWIN3 COMPOSITION OF TOTAL RECIRCULATION LINE BREAK AREA LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-1 DIAGRAM OF THE RECIRCULATION LINE BREAK LOCATION FIGURE 6.2-1 REV.11-APRIL 1996 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6,2-2 RECIRCULATION LINE BREAK PRESSURE RESPONSE (At 3434 MWt)REV, 14, APRIL 2002 LSCS*UFSAR

-I W...J I i c;\lJ)1.Il}i'!r tl.C;....!...!\-J....l-r I I lfl I..\I I-\\-\I u 0 0 it'en.......a&iLU i..;i,..;:..t::.<-):...LJ (--J...J uJ...J 0-<c en""'-;!d-.::::..;r....J 1..'""'-'!.o';o u:.'-C-i I Sd]eDSS3;:jd LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-2A SHORT-TERM PRESSURE RESPONSE FOLLOWING A RECIRCUL1\TlON LINE BREAK (At 3559 MWt)REV.14, APRIL 2002 I LSCS*UFSAR 1000 100 10 TIME ISECONDS)1.0 1 DRY Well TEMP.-DEG.F 2 WETWEll TEMP.-OEG.F1 I I/+See Note 1 2_.---....._.__._--.2 2 2--I**III-2 o o 450.0 300.w 9 w a::::I....<C a: If::e150 Notes: L This point represents the projected suppression pool temperature due to the feedwater coastdownlinjection.

This point is a starting temperature for the assessment of peak long term suppression pool temperature.

This evaluation is discussed in detail in Section 6.2.L 1.3.L1 in the paragraph titled,"Evaluation of Post*LOCA Feedwater Injection".

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2*3 TEMPERATURE RESPONSE FOR RECIRCULATION LINE BREAK (At 3434 MWt)REV.14, APRIL 2002 I I LSCS-IJFSAR tJ l!a Uti Sid w;--l':::

..,;.....\

J LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGU E 6.2-3a SHORT-TERIVI TETvIPERATURE RESPONSE FOLLO\"1NG A RECIRCULATION LINE BREAK (At 3559 MWt)REV.14, APRIL 2002 I

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE CONTAINMENT VENT SYSTEM FLOW RATE VS.TIME FOR RECIRCULATION LINE BREAK (At 3434 MWt)REV.14, APRIL 2002 I LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-5 CONTAINMENT PRESSURE RESPONSE (At 3434 MWt)REV.14.APRIL 2002 I LSCS-UFSAR 5.'to 3.LOG TIME-SEC 27000328 121'\.&LA SALLE , ow PRESSURE.CONT RESPONSE TO 2 WW PRESSURE.LOCA CASE.C SIL 836 I c=:?i II I.2---\, I I---e:::::::::

i-!i--i-I1.2.o o'to.60.SA'rLiSS 013002 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-5a LONG TERM CONTAINMENT PRESSURE RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)CASE C (2 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAY)FIGURE 6.2-5a REV 15, APRIL 2004 LSCS-UFSAR I I::tI I I',,." I LI-, 5f\.LLE.I GO.",,,,,, O)cco(',,,"c rr)i_OCA CASE C SCG6-0 1\I r F I I I i h I I I I I-'to.---."'-I".....'.I.*--<.-.U"l..--0-l-I 20.w I-"-cn U"l r LLJ a:: CL o.I I I I 10'10 2 to!10'tO STIME-SEC LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS RESPORT FIGURE 6.2-5b LONG TERM CONTAINMENT PRESSURE RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)Case C with GE SC06-01 CONSIDERATION (5 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAy)REV.18.APRIL 2010 LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-6 TEMPERATURE RESPONSE (At 3434 MWt)REV.14, APRIL 2002 LSCS-UFSAR

Lt\I I , DW AIRSPACE TEMP'CONT RESPONSE TO LOCA CASE C SIL 636 , 1 11 1 I--I--I--I---I--I--e-l-I.L..L.LJ.200 00 300'100.SAYLES 01300(.1l.2700032B 1214.8 2.3.LOG TIME-SEC 4.S.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-6a LONG TERM DRYWELL TEMPERATURE RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)CASE C (2 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAy)FIGURE 6.2-6a REV 15, APRIL 2004 LSCS-UFSAR LA 9U_Ei"""E T" I,

!1<N*...

__-,,--,---_.._,-"._+-*f i'-OCA lASE C 5(1]6'I,),!

-!I ,M f\-c:fl!I--....

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l-------r-------'------+-------.,f.-----1 10$10 I 10 1T1 M't.-SEC I'1-w r\,1=="'5 21,)0.r-lr.------+------..:>...--

'-==___.=----_-----1-----1 f-.<[I w i-I-L 100.F,-'_...l-.....l-..J......J-J-J-l...w..-_..J....--'--l.....LJW-Ww------JL..--L..L...l.I..L.LLW..._...I.-...I.-.L.j...LI..L..l..l-

---I 10 I LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS RESPORT FIGURE 6.2-6b LONG TERM DRYWELL TEMPERATURE RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)Case C with GE SC06-Ql CONSIDERATION (5 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAY)REV.18.APRIL 2010 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.27 POOL TEMPERATURE RESPONSEISOLATION/SCRAM.

1 RHR AVAILABLE (At 3434 MWt)REV.14, APRIL 2002 LSCS-UFSAR 200.LUl.=-1 IME-SEC d ,.,..;,):")(;;'8

)21";LA SALLE , SP TEMP CONT RESPONSE TO LOCA CASE C SIL 636 I 1 I I II-I I f-f-f-l-I I-----f-I , l.2.3.7,-"("---,....-'1.5.o 300..;i\'U:

..:-::':'130{);:" LASALLE COUNTY STATION c--__UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-7a LONG TERM SUPPRESSION POOL RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)CASE C (2 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAY)FIGURE 6.2-7a REV 15, APRIL 2004 3)0.LSCS-UFSAR:"'-Mp I Iii't";A c/\L: E i Ii I............

LWblL.3l:SP.ONSL_IU+

-_,.-.._--.I L eeA CASt C 5C06*0 11 I 200.r--------+------t-----.--=:::::t:=::::::::_..=-._-----,-----j tO l TIME.-SEC LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS RES PORT FIGURE 6.2-Th LONG TERM SUPPRESSION POOL RESPONSE FOLLOWING A RECIRCULATION LINE BREAK (At 3559 MWt)Case C with GE SC06-01 CONSIDERATION (5 PUMPS, 1 HEAT EXCHANGER WITHOUT CONTINUOUS SPRAY)REV.18.APRIL 2010 LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.28 PRESSURE RESPONSE FOR A MAIN STEAMLINE BREAK (At 3434 MWt)REV.14.APRIL 2002 I LSCSUFSAR LASALL.E COUNTY STATION UPDATED FINAL.SAFETY ANAL.YSIS REPORT FIGURE 6.29 TEMPERATURE RESPONSE FOLLOWING A MAIN STEAML.INE BREAK (At 3434 MWt)REV.14, APRIL 2002 I LSCSUFSAR LASALLE COUN1Y STATION UPDATED FINAL SAFE1Y ANALYSIS REPORT FIGURE 6.2-10 PRESSURE RESPONSE FOR O.I FT2 LIQUID LINE BREAK (At 3434 MWt)REV.14, APRIL 2002 LSCS-UFSjlj{

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LA SALLE COUNTY STATION UPDATED FINA[SAFETY ANALYSIS REPORT FIGURE 6.2-24 PRESSURE HISTORIES OF NODES FOR WORST BREAK CASES (SHEET 1 of 4)REV.0-APRIL 1984 10 40 r0-o 30 C/)a......., lJJ 0:::::J C/)C/)20 lJJ 0::: 0..lJJ<.:><t: 0:: lJJ><t: DRYWELL RESPONSE FOR BREAK IN THE HEAD CAVITY o 2 3 4 5 6 7 B TIME (SECONDS)LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-24 PRESSURE HISTORIES OF NODES FOR WORST BREAK CASES (SHEET 2 of 4)REV.0-APRIL 1984 I DRYWELL RESPONSE FOR RECIRCULATION LINE BREAK IN THE DRYWELL 2.0 1.5 1.0 TIME (SECONDS)0.5 60""'CI 50 (/J 0-....., w 40 0:::=>(/J (/J 30 w 0::: 0-W 20 C)<l 0::: W>10<l 0 0.0 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-24 PRESSURE HISTORIES OF NODES FOR WORST BREAK CASES (SHEET 3 of 4)REV.0-APRIL 1984 60.......Cl 50 en a......, w 40 cr::>en Cf)30 w a::: a.w 20 HEAD CAVITY RESPONSE FOR'"/<t RECIRCULATION LI NE BREA K cr w IN THE DRYWELL>10<t 0 0.0 0.5 1.0 1.5 2.0 TIME (SECONDS)LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT I FIGURE 6.2-24 PRESSURE HISTORIES OF NODES FOR WORST BREAK CASES (SHEET 4 of 4)REV.0-APRIL 1984 I 10 9 7.0 PSID PEAK BULKHEAD ,....8 PLATE DIFFERENTIAL P ESSURE 0 (/)a..7""..J<t I-6 z W 0:: W I.L.5 I.L.0 w 4 0::::l (/)(/)w 3 0:: a..2 BREAK IN HEAD CAVITY o 2 3 4 5 6 7 8 TIME (SECONDS)LA SALLE COUNTY STATION UPDATED FINAL ANALYSIS REPORT FIGURE 6.2-25 PRESSURE DIFFERENTIAL ACROSS THE BULKHEAD PLATE FOR THE WORST BREAK CASES (SHEET 1 of 2)REV.0-APRIL 1984 I 2 o 0.5 1.0 1.5 TIME (SECONDS)2.0 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-25 PRESSURE DIFFERENTIAL ACROSS THE BULKHEAD PLATE FOR THE WORST BREAK CASES (SHEET 2 of 2)REV.0-APRIL 1984 LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-26 VESSEL LIQUID BLOWDOWN Ri\TE (At 3434 MWt)REV.14.APRIL 2002 I LSCS-UFSAR LASALLE COUNTY STA nON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-27 VESSEL STEAM BLOWDOWN RATE (At 3434 fV1Wt)REV.I4.APRIL 2002 I LSCSUFSAR LASALLE COUNIY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-28 MAIN STEAMLINE BREAK RESPONSE PARAMETERS BLOWDOWN FLOW (At 3434 MWt)REV.14.APRIL 2002 I

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.229 TEMPERATURE RESPONSE OR REACTOR VESSEL (At 3434 MWt)REV.14.APRIL 2002 I LSCS*UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-30 SENSIBLE ENERGY TRANSIENT IN THE REACTOR VESSEL AND INTERNAL METALS (At 3434 MWt)REV.14.APRIL 2002 I RPV Containment MO I+--"TC DETAIL (a)AO I MO DETAIL (b)TC SO or MO DETAIL (e)SO or MO NOTE: TC DESIGNATES TEST CONNECTION.

LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 1 of 10)REV.9-APRIL 1993 LSCS--UFSAR FIGURE 6.2-31 DETAIL (d)DETAIL I SPECTACLE FLANGE TC DETAIL (f)Note 1: The Air Actuators are removed from Check Valves 2E 12-FO50A/B.

FIGURE 5 2-31 LA SALLE COUNTY STATION UPOATED FINAL SAFETY ANALYSIS REPORT FIGU RE 6.2-3t CONTAINMENT VALVE A R RANGEMENTS (SHEET 2 OF 10)

REVISION 20, APRIL 2014 DETAIL (d) DETAIL (e) DETAIL (f) RPV LSCS -IJFSAR FIGURE 6.2-31 Note 1 AO Note 1: The Air Actuators are removed from Check Valves 2E12-F050AlB.

fIGURE 6 2-31 CONTAINMENT TC TC LA SAl...lE COUNTY STATION UPOATED SAFErt ANALTS!S REPORT' FlGURE It2-31 CONTAINMENT VAlVE ARAANCOIENlS (SHEET 2 OF , 0) REVISION 20, APRIL 2014 RPV LSCS-UFSAR LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT SO MO SO MO DETAIL (9)TC MO MO DETAIL (h)TC 1.40 DETAIL (i)AO AO TC MO TeSALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 3 OF 10)REV.15, APRIL 2004 DETAIL (j)ACCUMULATOR Containment I SO or M Ei I DETAIL (k)DETAIL (1)SO MO...::::::.J I MO SUPP POOL\LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 4 of 10)REV.a-APRIL 1984 I/DETAIL (m)RPV supp POOL Containment MO o J DETAIL (n)1M*MI Te LA SALLE COU NTY ST ATIO N UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 5 of 10)REV.0-APRIL 1984 RPV CONTAINMENT LSCS-UFSAR FIGURE 6.2-31 SEE DETAIL (p)(UNIT 1 ONLY)DETAIL (0)SUPP POOL SEE DETAIL (p)(UNIT 2 ONLY)MO TC DETAIL (p)RPV SUPP POOL CONTAINMENT SEE DETAIL (0)LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 6 OF 10)REV.14, APRIL 2002

..DETAIL (q)RPV SUpp POOL LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT BLIND FLANGED MO MO MO MO DETAIL (r)SUPP POOL OR (FIGURE 5.2-31 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 7 OF 10)REVISION 13 DETAIL (5)DETAIL (t)DETAIL (U)RPV RPV I RPV TC TC Conta i nment r[}--4 i.,...'./';.0/\.0 Ej Containment TC Containment I co I LA SALLE COUNTY STATION UPDATED FINAL*SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 8 of 10)REV.3-APRIL DETAIL (v)RPV LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT INSTRUMENT I DETAIL (w)DETAIL (x)RPV RPV CONTAINMENT RESTRICTING ORIFICE.NOTE 1 CONTAINMENT RESTRICTING ORIFICE*NOTE 1 TC (SEE NOTE 2)EXCESS FLOW CHECK VALVE EXCESS FLOW CHECK VALVE EXCESS FLOW CHECK VALVE INSTRUMENT INSTRUMENT INSTRUMENT NOTE 1: IN THOSE CASES WHERE INSTRUMENT LINES ARE DIRECTLY CONNECTED TO THE CONTAINMENT ATMOSPHERE.

THE INBOARD PORTION IS BETTER REPRESENTED BY THE INBOARD PORTION IN DETAIL (v);HOWEVER, THE OUTBOARD PORTION REMAINS AS SHOWN HERE.ITE 2: WHERE PROVIDED.SEE CURRENT P&ID.FIGURE 6.2-31 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 9 OF 10)REVISION 13 DETAIL (y)DRYWELL WETWELL Containment TO TO Containment MO DETAIL (z)LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10 of 10)REV.0-APRIL 1984 DlleNI.lgpt Tha CLOC rapra.aDt**boUAdaria.(valva**flaDq,**pump.aal**,tc.)which ara Doraally.aalea cloa,a.automatically clo.ea.or are clo.ea vith a remota Manual operator to accOMpli.h cDDtaioneDt iaolation.

Ta.t Moae 1 i.rapra.entaa by.olia liDa..The acrc By.t..i.aliqDaa to taka'UCtiOD from tha CODaaD.ata.tora91 tank (CST)aDd tba full flOW ta.t raturD liDa 1.al190ad to tha CST.Valva.&5l-r362 aDd r363 viII becOMa priMary contaiDMIDt i.olatioD valva*.r.1 DETAIL AA---...-----.,I I I I I , t..__J t"=-=!f--...L_--t........-v--r---l-I 1 ln\.leil-'4f£ l,;,-f,...,]*I-,>"'7 f'rl6'-J.f1#J

-:'..-V1"oI.t:,,;.;;.;;;J,.Lf 1!!l I.'c""oc' CONTAINMENT P&NZTl\ATION M-17 INSIDE I OUTSIDE SUPPRESSION POOLc a> 0>>-r- r-n ri-lc Vi>>;;o-f m-oO;;oZ-l"Tl G)ct-".w Of o:;,Vi3m........<-Ill i a Ta.t Moda 2 1.repre.aatad by da.bad liDe..The aCIC By.t..i.ali90aa to take'UCtiOD from the Buppre**ioD Pool (8P)aDa tbe flow ta.t ratura Ii..i**li9Oaa to the SP.Valve.&51-r362 aaa 851-r363 viII 00 loaqlr be cODtaiDMIat i.olatioD valva**Valva.&51-r022 aDd r059 will becoma coataiDaaat i.olatioa val VI'.IDd.pactacla fl109a 151-D316 (bliDa.ida)will be a cODtaiDMIat i.olatioD boundary.>>""'0r--.0-.0 Wm:<-.0 LSCS-UFSAR FIGURE 6.2-31 DETAIL (AB)DETAIL (AC)RPV RPV CONTAINMENT RESTRICTING ORIFICE CONTAINMENT RESTRICTING ORIFICE TC (NOTE)EXCESS FLOW CHECK VALVE EXCESS FLOW CHECK VALVE INSTRUMENT REF.LEG I_BACKFILL LINE INSTRUMENT REF.LEG I_BACKFILL LINE NOTE: WHERE PROVIDED, SEE CURRENT P&ID.FIGURE 6.2-31 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT.FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 108 OF 10)REVISION 13 RPV LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT DETAIL (AD)REACTOR WELL DRAIN LC TC LC NOTE: THIS FIGURE APPLIES TO UNIT 2 ONLY.LASALLE COUNTY STATION.UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10C OF 10)FIGURE 6.2-31 REV.11-APRIL 1996 RPV LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT DETAIL (AE)TC TC LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 100 OF 10)REVISION 13 FIGURE 6.2-31 I LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT RPV RV DETAIL (AF)RV MO MO I LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10E OF 10)REVISION 13 FIGURE 6.2-31 DETAIL (AG)RPV supp POOL LSCS-UFSAR FIGURE 6.2-31 CONTAINMENT RV FIGURE 6.2-31 VACUUM BREAKER MO MO MO MO MO LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10F OF 10)REVISION 13 DETAIL (AH)RPV LSCS-UFSAR FIGURE 6.2-31 AO OR MO RV CONTAINMENT AO OR MO TC LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10G OF 10)REVISION 13 FIGURE 6.2-31 DETAIL (AI)RPV LSCS-UFSAR FIGURE 6.2-31 fiLII UNIT 1 ONLY ANGLE VALVE (TYPICAL)CONTAINMENT M LC M LC TC FIGURE 6.2-31 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 10H OF 10)REV.13 LSCS-UFSAR FIGURE 6.2-31 RPVCO NTAINMENT DETAIL (AJ)Note 1 AO 11 TC Note 1: The Air Actuators are removed from Check Valves 2E21-F006, 2E22-F005, and 2E12-F041AB/C.

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 101 OF 10)

REV. 20, APRIL 2014 FIGURE 6.2--31 DETAIL (AJ) RPV LSCS-UFSAR FIGURE 6.2-31 TC Note 1 AO CONTAINMENT MO Note 1: The Air Actuators are removed from Check Valves 2E21-F006, 2E22-F005, and 2E12-F041A1B/C.

FIGURE 6.2-31 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-31 CONTAINMENT VALVE ARRANGEMENTS (SHEET 101 OF 10) REV. 20, APRIL 2014 DETAIL (AK)RPV CONTAINMENT RV SUPPRESSION POOL DETAIL (AL)Operator (typical)Accumulator (typical)CONTAINMENT AO TC LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE CONTAINMENT VALVE ARRANGEMENTS (SHEET lOJ OF 10)REV.14, APRIL 2002 I

.....c:<<u 0 0-l a: w r-I..L.<<wi=......--......---..--....---..--.,---.,..-..---,---,r--..----r-..,.--..,.-..,.--..,.-......---,O o 10 o o o 10 o (MW)

LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-32 ENERGY RELEASE RATES AS A FUNCTION OF TIME REV.0-APRIL 1984

'" ,..---,,...----.,.-----r----...,-----r-----,,...----..,.-...,O

'" o.....J::-<t U o...J cr W-21.J..<t W:E r-o o L..---:l::--

__-**__

I()N o N (OMW) 03.L I LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-33 INTEGRATED ENERGY RELEASE AS A FUNCTION OF TIME REV.0-APRIL 1984

'"(,;: 0-I.,l (f)0 Q)(f)-.0::>-3l:..J Q)0:-I 0 15 ,..::E"-<to 0::-(f)(f)>--..J 0 Q):r: 0 Gi<t 3l: 0::..-....s::-u 0...J a:: 2 w l-lL.w::E-I-a o LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-34 INTEGRATED HYDROGEN PRODUCTION AS A FUNCTION OF TIME REV.0-APRIL 1984 1000.0 100.0 10.0 Time After LOCA (hours)1.0'I I4-Wetwell Oxygen/1J3-Wetwell Hydrogen2-Drywell Oxygen4 f---1-Drywell Hydro en t.,"IIV ,41",0"",,,,-.".".,....10-I.-olI..--"",."--o 0.1 8 4 2 10 12:;:: o III§6'0>LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-35 UNCONTROLLED HYDROGEN AND OXYGEN GENERATION REV.14, APRIL 2002 LSCS-UFSAR I I I Rated Power, 6 Hour Start Time 105%Uprale, 5 Hour Start Time--.-.-..--........-.....-"....-/--I Drywall I/0......-......::..:..........----.........

II Walwell I..........

/-5 4 3<II E:='0>2 o o 10 20 30 40 50 60 Time After LOCA (hours)70 80 90 100 Note: The information provide in this figure is historical.

The hydrogen recombining function of the hydrogen recombiners is abandoned in place.LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-36 HYDROGEN CONCENTRATION WITH 125 SCFM REV.17.APRIL 2008

'6:'Z_-'!,.':'\__.J._:-16'a i 11.....,;.......,...,

!"'..LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-37 NODALIZATION OVERLAY FOR RECIRCULATION LINE BREAK REV.0-APRIL 1984 t;;'"I l l: '-I" l'ir---2H LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-38 NODALIZATION OVERLAY FOR FEEDWATER LINE BREAK REV.0-APRIL 1984

,0<")" o 0)en I()0 0\4)rt\Nrt\N!"l-N lX)...<1<t" (fl<<r I")N ("II--, I rt'I<D (r)(()I")N---t-I")N N I I"-l" l'-N N--, 0..9=I'l')N.s:l-N-1i1f\X1 (g I....00 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-39 NODALIZATION FOR ORIGINAL RECIRCULATION LINE BREAK ANALYSIS REV.0-APRIL 1984 o o CO o o 0'1 t-\0 IV)---c-.Jl;r<1l(0)N-...LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-40 IlEQUIVALENT" NODALIZATION (CASE A)REV.0-APRIL 1984

-CD',.,:--Q N"-"..0'r--b-(J)*, 0 I".L 10 T'J r 1"0 I ,.., 0 00 0 0 0 0 IS'09 III'" r"I N N"70 (l..""

v I I Q'<:>--J..0 V)0 N-V'0.I---0 i" I t-'I...I II0 I'?0<0 0<0 g 0 Cl'" ()'ft 0'"'!!"'-V1'" f'7 N r4.'";'b to p-i-4)<:,(,.J..'" o:;6'l.0'b ,.., 0 00 0 0 0 It'>0!'2.It'>to'" N";'.......LA SALLE COU NTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-41 AZIMUTHAL PRESSURE DISTRIBUTION (ATRECIRCULATION OUTLET NOZZLE)ORIGINAL DATA AND CASE A.Kt:V.U-APRIL 5'02,21 o SO 100 t.:-0.01 I;1I I j I I I-, ISO SO 100 t.0.10 seC e, IS', 30 0 3!I f , I I I I!-1 I&'02.(Joo e, OO-IS'SO i:'0 oC;5..c.50 100 t*O.IOsec e.=O**IS" I.13[I t I I'1 I I I I , TH.o LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-42 AXIAL PRESSURE DISTRIBUTION ORIGINAL DATA AND CASE A REV.0-APRIL 1984 0"1 ,....('l)---0--t..9"l'" CO<'\l-NI 1 I I_I LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-43 SIMPLIFIED NODALIZATION (CASE B)REV.0-APRIL 1984

.---c...0-C)I.f-a tJ..g'b i--I"I 00 e.00 0'" (1-;""l" N V':l'!!lI'0'b 0 ll>..-\--l b:---"I'g J<>';$gS0 0 (l.I'>t")N'"-.Jv-b 0 ll><5".+J 3 b..,<:>,,",,0iiiQ0 0-'"'" (.L§...'" LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-44 AZIMUTHAL PRESSURE DISTRIBUTION (ATRECIRCULATION OUTLET NOZZLE)CASE A AND CASE BU-H.J:'.K.LL 50 1)0 t'o,Ol I G): 015'ISO

;....,0-0----1 7 50-t: 0,0',&, 15'*30' t----J'----.,., I 1 1-----'--'---

..----, t----..!-!...--

..-....-----, I I 1---..,,--------1----j ICC t=0,05,8: 15°-30" 1$.29 f---..LL.---.-----,------,.-

755'.29.._---'-.1-----,-

,.---..__o 50 ICO 150 0 J E)-d-ISo P (p)i<t)50/00 t.'0.1 D)e;\5"-30°-1-\: 1----..,I:----'155 o 50 100"trO,j0J n I I I-----I:I I LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-45 AXIAL PRESSURE DISTRIBUTION CASE A AND CASE B REV.0-APRIL 1984

/0)I"-lr)rf)--()--....0'<t q)r:!

N LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-46 COMPLEX NODALIZATION (CASE C)REV.0-APRIL 1984 b J N I<!-v')"0 0 OJ.0£1\-t-J'b II n-_11 C 0 r;6g g 6-0 0 p....a.-N I"I/)-'-'b'b,/)If')"0 0." 0..4J'b I: ").___.ll ()r;;o 0£00<:>(l--III 0'!: In f")NV'l g 0 0...-tJ 0....-0 0 0§0 0.-III'"'"'" II)11..$....."J N LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-47 AZIMUTHAL PRESSURE DISTRIBUTION (AT OUTLET NOZZLE)CASE A AND CASE C REV.0-APRIL 1984 50 1000,01 J*I: 15 0_30.902.2;3

!---'---.........-----..,.----r---1ss:a---l.----,-----.-------.---

o 5D 100\50 0 t:O.OI J

..)Befa--------'I , 150 1><.Sl).-.J B: I-0 p.....I I-----------\I I 9 7$.2 15'0<.)

50 100 t:'O.\O)9-:.0°-15°

.Je;I'O-1<.pSt n j----1--il-------------1$.'29'-----1.-.....------..,.----,..--

7s.1'l\---..L---,-----.-------.---

o 50 100 1'50 0 50 100 t*0 OS 0 S*"P 4)1:00 S" sa 3 0 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-48 AXIAL PRESSURE DISTRIBUTION (CASE A AND CASE C)REV.0-APRIL 1984 TOP I SECTION AT BREAK PLANE I I I I I I 0 50 100 15'0 200 V 300 PRESSUR E (PSIA)TOP SECTION AT 90 0 w/r TO BREAK PLANE I I I o 50 100 150 PRESSURE (PSIA)200 200 100 150 PRESSURE (PSIA)50 I SECTION AT 180 0 w/r TO BREAK PLANE I I I I I I o TOP LA SALl,..E COU NTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-49 AXIAL PRESSURE DISTRIBUTION AT t=0.500 SECONDS REV,0-APRIL 1984

/

LPC I NOZZLE SECTION r-----..------.18oo 90" FEEDWATER NOZZLE SECTION i------:Jll::"""-------.,1800 90 0 MID-SECTION I I o I I SCALE II, 50 PRESSURE (PSIA)I III 100 I I 150 0'------_....---------...,180 0 270 0 00 1800 BREAK PLANE 90°UPPER RECIRCULATION NOZZLE SECTION 90 0 LOWER RECIRCULATION NOZZLE SECTION LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-50 CIRCUMFERENTIAL PRESSURE DISTRIBUTION AT t=0.500 SECONDS (SHEET 1 of 2)REV.0-APRIL 1984

___-------:.:::----------,1800 90 UPPER REACTOR SKIRT SECTION r---:::iIltC"'

---,180°90 0 LOWER REACTOR SKIRT SECTION SCALE

+1 o 50 100 150 PRESSURE (PSIA)LA SALLE COUNTY STATION UPDATED FINAL" SAFETY ANALYSIS REPORT FIGURE 6.2-50 tIRCUMFERENTIAL PRESSURE DISTRIBUTION AT t=0.500 SECONDS (SHEET 2 of 2)REV.0-APRIL 1984 TOP I I I SECTION AT BREAK PLANE J , , o 50 100 150 PRESSURE (PSI A)200 TOP SECTION AT 90" w/r TO BREAK PLANE I J o 50 100 150 PRESSURE (PSIA)200 TOP SECTION AT 180 0 w/r TO BREAK PLANE I o 50 100 150 PRESSURE (PSIA)200 LA SALLE" COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-51 AXIAL PRESSURE DISTRIBUTION AT t=0.500 SECONDS (CASE C)REV.0-APRIL 1984 FEEDWATER NOZZLE SECTION 27, 90°------------.J BREAK PLANE SECTION I 0°LPCI NOZZLE SECTION i---""'Z'"----,90" C1'RECIRCULATION NOZZLE SECTION.--__"":K'__---.90°0" LOWER REACTOR SKIRT SECTION CI'MID-SECTION 270°r--0°UPPER REACTOR SKIRT SECTION SCALE IIIII IIII IIIII I o 50 100 150 PRESSURE (PSIA)LA SALLE COUNTY STAT!ON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.2-52 CIRCUMFERENTIAL PRESSURE DISTRIBUTION AT t=0.500 SECONDS (CASE C)REV.0-APRIL 1984

    • .....*10 I".2-1-...._.r-.I----=----.I----::::--=:=:-::=:-:-::::-:.t.::::::-:-::.-.I----=----.I----=-.---.J---..;z'-

__

__=_..l_.::...__._..l_..L.__..L._L:.:!_

__!..:.:'__-----1-..-l.--II 11.12.APPROXIMATE AND wusr 8£DETERMINED EWiED ON FINAL.S"tSTEJrlI IWUM Al.LOWAEI...£.

G IS THE MINIMUM POSS19l£AND MAY 8£INCREASED BY OTHERS"""""""'00.

@WILl BE DETERMINED IN PRE-OPERATlOI'W.

TEST.1}lE I:1P REQUIREMENTS Of UOOE C.0 AND/OR E.AND TO ullIrr MAXIMUM DIESEl SERVICE WATER PUMP OCCURS IN WODE C AND MUST BE 22 FEU-BWRSJ)SCOPE ME AS NOTED.

WOIlES._1 0 Tlf£ESTIMATED UN[SIZES ARE FOR INF0RM4TION

TVRE ANO PRESStmE AND UHf SIZES AS DETERMINED BY OTHERS DIAGRAM HYDRAUlIC REQUlREMENlS.

11.lH MQ()£0 AND WffJf A VESSEl.PRESSURE Of 14.1 PS!A, THE fl.DW StWJ.NOT EXCEED 7175 GPIlL'I, 1liCOffjN,.VE lEU-Fill (2E22-Flll)

HAVE BEEN REItOVED PER DCP 0511324 (09lJ25).U,/lIIC,Itw.

..._.....-.......10>-1 (ISIU<.I.lO"l'It**)u M 110'1 SALLE COUNTY FI NAL SAFETY llllJl:S;1.

2.

wuc:rMW I: WINlWUM VAlUE Of'PARAMETER FOR THE WOO(SP£CIFIED.

J.£UVATlONS ARt MOT INCLUDED IN IJP VAlUES GIVEN.ElEVATIONS SWU BE INClUDED WHEN DETERMINING fIfW" f'Ofit THE EMPTY PRESSURE BUNKS.4.HEAD WIll.NOT EXCEED 34M FT.l$.TOtf'OfFT.AND A VESSEl PRESSURE Of 215 PSIA lliE FLOW WST BE 52ft GP>>.I, MODE: E BASED ON A MMlWUM CONTAINMENT PRESSURE OF 4S PSIG.BWRSD NMEH'T DESICH IS 9o\SEO ON HIGHER PRESSURE THE IMPACT ON THE HIGH 1, (NPSH)AVAILAfU AT THE CENTERlINE Of THE PUMP 1 FT**I.FIGU OOCUMENTS UNDER THE fOU.OW1NG lDENTlTIES

<<1'0 I(U5EO IN CONJUNCTION WIlH nus14.7 AM.......PRIMARY MODO.ACCIO£Nr;fIV{;.ICroti'

'"",'PP'$UPi'SUCTION J"1i!O'" SUPPII'CSSION PC:ll."S Ii'1 8 g 10" liI!.nJ/"l9lO,.

159 fA 15SO,"" N/fIt, MOOIi:C pc:JlrlON FiOWGPM PR£5S PSI" reMP-,.MODI: 8 POslrlo", FlOWGPM POSITd'N PRESS PSIII TEMP"!'u MODE II Fel,.NOTE ,.TAIl\-'!: 1-TA8l£1'015 19 o VAtVE OPEN C.,ALV.CLOSED VA.LVe fOO'IRlO4if010 roll FOIl jro'S 1/02"1"034 MOD£8C0 CCC 0 C 0 MODE C C 0 C C C 0 C 0 MODE D I C I 0 I C I C c 0 c 0 MODe J" , 0 c CC0C 0 MODIi: G C C CCC000 MODE J'C C IC\C 0 0 C I 0 MOOGC C C C C 0 C 0 PUMP OPERATING ON BYPASS SUCTION fROM SUPPRESSION POOl (4 5..1 IS"""'.,.21 FL.C>>"4 c:..PM ,_,_y-,_lIH.x-x.'><'X'XXX<T Tl':MP"F'ill'%'.'-iO:i'd:-c

..1--1.0_A H K L..o A o R)IS FOIIi NOTE ,. H TAIII..I!:

1-TA8lE VA.LVe fOOl JRlO4ifOiO roll FOIl 1f014 JI02'l"034 MOD£ 8 C 0 C C C 0 C 0 MODE c C 0 C C C 0 C 0 MODE D l'JOICI'I' 0 c 0 MODe F c 0 C C C 0 C 0 MODIi: G C C C I C 1 c 0 0 0 MODE .J' C c lei 'I 0 1 0 J c I 0 MOOG c c c c c ° c 0 K PUMP OPERATING ON BYPASS SUCTION fROM SUPPRESSION POOl C 4 5 .. ? IS "" 11 1& ,. 2' FLC>>"<I c:..PM ,-,-y. , ... ,. .. .x X X X z<.. 2<.. 2<..

<T TI':MP ... 'u' 'a' '% L iflfO'p 1--1.0_ ..

  • PRIMARY MODO. MOOli:C " a .. 10 11. '2. (NPSH) AVAILABLE AT THE CEHTERlINE Of THE PUMP FT. a IS APPROXIMATE AND MUST 8£ DETERMINED MSED ON FINAL SYS'TEM IMUM ALLOWABL£.

G IS THE MINIMUM POSS19l£ AND MAY 8£ INCREASED S'( OTHERS """""""'00.

WIll BE DETERMINED IN PRE-OPERATIONAl TEST. THE I:1P REQUIREMENTS Of UOOE C, 0 AND/OR E. AND TO UMrr MAXIMUM OIESEl SERVICE WATER PUMP OCCURS IN WODE C AND MUST BE 22 FEU 14, -BWRSO SCOPE M£ AS NOTED. 1!l INC VARIOUS OPERATING MODES. 1., TEMPERATURE AND THE ESTIMATED UN[ SIZES ARE FOR INF'0RM4TIOH RE AND PRESStmE AND UHf SIZES AS DETERMINED BY OTHERS HYDRAULIC REQUIR£MENl5.

1/. IN 0 AND Wffff Ai VESSEl. PRESSURE Of 1 ** 1 PSIA, THE fl.DW StWJ. NOT EXCEED 7175 GPIt. ,.. or CI!£!li!

'E22-"'. (2£22-"18)

HAVE IIEEH RE1IOVf.D PER OCP .0HJ24 ( ....... ).

OOCUtdENTS UNDER THE fOU.OWlNG lDENTlTIES

<< 1'0 I( USED IN CONJUNCTION WIlH nus ulllllC..a

    • l... ... _ ..... -.. ..... 1I)j..I (611:& t.lO"I'lt

'1Ii) ." M SYSTEM References 51 and 55.GE LOCA ANALYSESREV. 20, APRIL 2014ASSUMED IN GE LOCA ANALYSES..J < t= Z w a: w LL LL C w a: :::l en en-we a:fI) Q.Q. ..J-..J W :: >-a: 0 0 .... ..J W en en w > LSCS-UFSAR 1200 1000 [;;] '800 600 400 200 o 1000 Source: 2000 3000 4000 5000 6000 HPCS FLOW (GPM) LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3*2 VESSEL PRESSURE VS. HPCS FLOW

  • LSCS-UFSAR g ,..§...,!'1!I i I I I r-niI i co!.I-w t+I I I i I I j I 1 i I I--I I I I I_-+-_.......--:...........__"'"--__.........__**LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-3 HPCS PUMP CHARACTERISTICS REV.13 5 I lotA"OOI"'.'618 I.II ,."" I'0'**o c B=fi ,..N'" o[-,., coocc o00CC"CI0C0 o c.p c o00CC....00((oC0C0_0"001 C MO**I..._I.....,-.COHO'liON...-.!."", ,..-;.....;,;:..;,..,-,.T.-OO=."_=.....,...I.",.-__"""1"'WI-'W.a...ma.a too ua mtlll H'.ua*.,**I.UU""__wa.----u,......**1IlIUAI_1AI,,1!'IICM.t.---
      • .en...lUI....lCAt..:.....1ft",.-Ul--....rtf'Z!$la'M lUI 0IIll......""'".......,.-..._."U....___
  • ...",I&M'..............

fI....

......

..

r:......=_..".Jt c*-..::=t,,':---',....!FaB.=::'t.....I:.c:-att.u::
  • .::...til"'".u,n_,'I-...II.

..h.=.r.=1t'n

    • "r...._ac..__............

..._,......,..

-.-..._---.., , Cftlllt........,.".

=,-v=.-:

....,s.PIlE_DIll:.lIlI: ACllMlII..USDl UlCo'l.WIl_lDh.........CJ****....X x X

r...........','%......'.'........1..........

taOI.....................

,..,...*I*..I*T*, 10" II'1 I II l0'.zoo...Y" iI...'."1 ,.X..,...., t)<'9:

I'0,,"'_0*,*..*t"\D'IIf'_*.........,.....,..........

%

u>c:-".0,,*..*..*r,__I't/A-_......."....'ttl:rmCia:lt*tOll-",'M.....,......., Yo

.*".N Jl'(C f C.,.B o c

,5 t I I..*(,T**10 It" oJ..o'-'*"'All'f QIIIOfC-'WY (tostO'o..n..,.,....-A LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT A...., ,.*>J...." FIGURE 6.3-4 lPCS SYSTEM PROCESS DIAGRAM L-_'--.L._____JL.__-__..*.;:.3....._..2 1 REV.13 LSCS-UFSAR 300 200- -GE.`i 20003000 4000 50006000 LPCS FLOW (GPM) 70000000 0 0 1000 Source: References 51 and 55.

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-5 VESSEL PRESSURE VS. LPCS FLOW ASSUMED IN GE LOCA ANALYSES REV. 20, APRIL 2014 LSCS-UFSAR .J .ffi 250 m is 200 W a: :;) e 150 a.D.; ,.J-ad a: 100 D e uI 50 m > 0 0 .....,; ...... ............

...... ...... .......... , ...... , ...... , ..... .... '", --GE \ \ \ \ \ \ \ \ 1000 2000 . 3000 40c0 5000 eooo 7000 8000 LPCS FLOW (GPM) Source: References 51 and 55. LASALLE COlJN'1Y STAll0N UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-5 VESSEL PRESSURE VS. LPCS FLOW ASSUMED IN GE LOCA ANALYSES REV. 20, APRIL 2014 LSCS-UFSAR

  • U!

I (%)

0 0 0 0 0 0 0 0 0 0 0)IX)I"-ft)It)l")N--0 0 0 I 0)0 0 I 0 IX)0 0 0 I"-o g--1--__.......--1........-+-,+-.1-0.-___.0 g--g N--8 IX)8 ft)88 N o*LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-6 LPCS PUMP CHARACTERISTICS REV.13 LSCS-UFSAR

_ _ GE 10002000300040005000 a w CC 0 0 T000 8000¨N.¨Source: References 51 and 55.

i LASA LE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-7 VESSEL PRESSURE VS. LPCI FLOW ASSUMED IN GE LOCA ANALYSES REV. 20, APRIL 2014 LSCS-UFSAR I " fli It: is iii 200, II: i .50 . 8J is a:iiJ Q.E!::, 100 iii @ 50 " " " , , , , ,. , , , "'-, , , , " , ' .. ", j o 1000 Source: References 51 and 55. 3000 4000 5000 6OCIO 7000 , LPCI, F1.0W (GPM) LASALLE COUN'IY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-7 VESSEL PRESSURE VS. LPC! FLOW ASSUMED IN GE LOCA ANAL YSES REV. 20, APRIL 2014

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--- 90/!H1 J4'/..,1uIc_0-*D 1::......Ie JlIlk.....'UIOJ ,.r..M JI':.Wt'IC,I'" I,)ftC'--,'.IL...,....,,,..til" ltfUt_0.._.,..........c.-.I'"_ff ru:Y\'Ia..lJ..caa&I I_UfU___..U.oa."",C&-.Ll&, ,...._.....,...S."IQl

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..':'M\lr fOIIl ,...,.., i.AULO.U-'.D.._".....L" n.Qltt..,......, If""'......*C" l." Ita.a""lUlL 1111 uftCll....""" I.'oc.......y*RFV9-APRil 1993 FIGURE 6.3-8 (SHEET 1 of 3)RESIDUAL HEAT REMOVAL SYSTEM (RHR)LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT lSlJ:sl.llll.S lS4.S 11""

--XiX'0 Ill)6" il)J.I!.,,1.-...-Z,.7.".Z7.2U 2'1"'., iii".7...7W IZSIY x X:X, IlS XiX MSO-10'....II MODEE 7 T4I*I.10....MfA"5 X'XXXXXxxx 155lio1..-Z*,,!*iI*L"**

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...........

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  • D..OSIT.O....0

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OI5f1ltlr.E I IE lfiiiiiiifji;o trmr (r IHJU II'(TISl liNE Ttl Sl/,,,n5lOll11Ja) a*LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORl FIGURE 6.3-8 RESIDUAL HEAT REMOVAL SYSTEM (RHR)(SHEET 2 of 3)REV.9-APRIL 1993 r p m,vr"ArNo+tl:'Nr fll LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT___,-1(l.':l<<r Yfl." Xl"'lOltl1

"".J3A.§t-d r04u, ,.--AllCACf'OR SCt"O/o/0-4Pt'(DN7AU"M"I:AI r r.wn s£PltI'C£"'Arel$Y<$.FIGURE 6.8 RESIDUAL HEAT REMOVAL SYSTEM (RHR)REV.14, APRIL 2002

  • LSCS-UFSARII!

I (%)klullP!lI3 8 co§-8-8..,

8 0 r-/I)U 8;::;" co 1: S u 1!III.:§.!!0 j 00 c: Q"E..GI 0.'C n.!!,.II)a: 8i.§::z: "§iii a:*l;:1::::!c.,.Cl 0!I 0.{!.*..I w Zo.!++I ii tn*..I 8 N LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT*FIGURE 6.3-9 LPCI PUMP CHARACTERISTICS Sheet 1 of 1 REV.13 LSCS-UFSAR

,_....._.-i'"""'"...200 1"00 1200.1000**1000 Source: Reference 26*LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-10 HPCS MINIMUM REQUIRED PUMP HEAD TO MEET LOCA ANALYSES ASSUMPTIONS REV.13

0** ISO 100 LSCS-UFSAR

..:II*.--......;: ".-..*1000 Source: Reference 26 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-11 LPCS MINIMUM REQUIRED PUMP HEAD TO MEET LOCA ANALYSES ASSUMPTIONS REV.13

  • LSCS-UFSAR I--._._-.,.;r"--._---I I-r-----I I---I'II"'".soo so 450 400 350 ISO 100*1000 2000 3000 4000 FIowCgpm)1000 Source: Reference 26*LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-12 LPCI MINIMUM REQUIRED PUMP HEAD TO MEET LOCA ANALYSES ASSUMPTIONS REV.13 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.347 SCHEMATIC OF THERMAL OVERLOAD BYPASS CIRCUITRY REV.14, APRIL 2002 I (BASED ON ONE SOTS EQUIPMENT TRAIN OPERATINGl SOTS REACHES ,--FULl.CAPAOITY"--,, I I I I I I I LOCA INITIATES!SECONOAIW CONTAINMENT ISOLATION&START OF SOTS

_I I I I I r MINIMUM-+1__I I I I I 0 d:t-0.05 (Il w::r::-0.10 ci z-0.15 a...:5 OJ*0.20 a:: 0 ro-U'-0.25<<w a::-0.30 w a:::::>(Il-0.35 III w a:: II.-0040 100 105 200 I 252.9 300 400 TIME.Se.CONDS(POST-lOCA)

NOTE: This figure was used to support original licensing.

For current licensing requirements for system pressure-time response, see the Technical Specifications.

LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.3-80 POST LOCA TIME-PRESSURE IN SECONDARY CONTAINMENT (BASED ON ONE SGTS EQUIPMENT TRAIN OPERATING)

REV.15, APRIL 2004 LSCSlJFSAR

...o a:: I-z:*om 0111 wi:...c.cO g§00 mill........0 C..I0%&II"".n32....I-g

%-=LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.4-1 CONTROL AND AUXILIARY ELECTRIC ROOM LAYOUT (SHEET 1 OF 2)REV.14, APRIL 2002 I

\ESS SWITHGEAR VENT SYSTEM (VX).-.,,Ill ,.:)I I:

1 TURBINE BLDG.I (l VENT I SYSTEM(VT)

I..(l'Zl FLOOR fL.731'-0" NOTE: SHADED WAllS AUX.ElEe EQUIPMENT ROOM ENVELOPE BOUNDARIES.

LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS HLPORl F[GlJRE 6.4--I CONTROL AND AUXiLIARY ELECTRIC E'1lJ I PMEN T ROOM LA\'OuT (Silt 1.1 2 nr 2)r<I:V.()--!\l'10L84 6LDfa.IB'*6'<1 VENT.STACK UNIT 2.UNIT TURBINE:: CONTROL ROOM AIR INLET ANl>DETECTOR (EACH SIDE)W Q<t<<" 0....w 0 0<<,-0: 0" I"-0'" I-0 ,':TMONITOR AND INTAKE GRILLE EL.847'-0" TOP DIAMETER 18'-6" PLAN VIEW 1/II II II STACK II II-----STANDBY GAS TREATMENT II SYSTEM El(HAU5T PIPE il II II II REACTOR BLDc:..El.894'-0" E:.L..1080'-0"_._---Bih-Q..IOR L U.BQ4'-O' U__NlT_1...l.._U_N_IT_2AUlC.BUILDING 5ECTION"P;'-"A" LA SALLE COUNTY S1 AllON UPOflTED FINAL SAFETY ANALYSIS REPORT FIGURE 6.4-2 LOCATION OF OUTSIOE AIR INTAKES REV, D APRIL 1984 Airborne Reactor Vessel-6" Iron Reactor Shield 2.5" Iron+21.5" Concrete Plate Out*1000 N*0.250 H*0.005 P Leak Rate of O.005/0ay 6'_0" Concrete*0.250 H*0.005 P a (0.13)Continued on Sheet 2 Plate Out 0.5H+0.5P (0.87)Airborne 10 N+0.5H+0.5P A Continued on Sheet 2 I.OON O.IOH REACTOR BUILDING REACTOR BUILDING I.ON+I.OH+IOP STANDBY GAS TREATMENT REACTOR BLDG.FLOORS 36" Concrete 56" Concrete REACTOR BUILDING WEST WALL LEGENDN-Noble GosesH-HalogenP-Particulates

..-Distribution of fission products immediately following 0 LOCA NOTES I.Flows beyond the primory containment ore fractions of the upstream input.2.The.635%per doy leak rote will increase the downstream sources by approximately 25%[(1-8-.00635t)/(I-e-.005t)125]r------, I CONTROL I: ROOM: I'--J LA SALLe: COU NTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.4-3 CONTROL ROOM SHIELDING MODEL (SHEET 1 of 2)REV.0-APRIL 1984 REFUELING FLOOR aeoetor e Continued from She.f I P1of.Out West side of Building LSCS-UFSAR CLOUD Continued from.Shed I.(XIO)CN't AUXILI"RY BUILDING AIR INT"KE CONTROL.ROOM"IR INTAKE"',OOOcfm 8 CeRiIl4l...12.HoI\ow Bloc.wall..CONTROL ROOM CONTROL R()Ot;I."'INTAKE fiLTER UNIT ADJACENT..1flBORNE'S"llfl'N.)

.Out"" W." LASALLE COUN1Y STATION UPDATED FINAL SAFE1Y ANALYSIS REPORT FIGURE 6.4-3 CONTROL ROOM SHIELDING MODEL (SHEET 2 OF 2)REV.14.APRIL 2002 I REACTOR VESSEL SHIELD WALL SAFE END TO VESSEL""-----.....-......NOZZLE SAFE END TO PIPE WELD LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-l SAFE END BREAK LOCATION REV.0-APRIL 1984

o-.::::t".0>-0::: Z W:J:...:.<:-c(w W:.<: 0:::0 l:Q:J: U tv"I (f)-:::>>-0 OQ-WO en ZO Q c(%:z:....0 zw U C(:.<: l..LJ....(f)en (f):::>-ZC(-u..l..LJ-::E:.....J-Oo/l Q.......en c(...I N W-.0::: 0....-i-.o I I I II 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0....-i 00 Lr\N en to tv"I N....-i....-i....-i Zl.:l-J3s/wal LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-2 BREAK FLOW VS.TIME-FEEDWATER LINE BREAK REV.0-APRIL 1984 CD]J eb , L/f P0 2 POI Dr'\.*11AL A BR AL 2 I LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-3 GEOMETRY REV.0-APRIL 1984

-SA UF Al E)N(Dr-FLASHI Nt;WAI ERCJU_f-I i I I'" I RF E(I(IN-I-........I..............

I""'"----------._.SATURA-------.\El2.-.".--I/SATURA"l EO ER.../FLASHIN:;WAT i'....J//.....2, ,/.///'/V a 10 a llJ llJ a..(J)()10 Z o (J)I ()-(J)a..lJ..-...10 100 PRESSURE (PSIA)1000 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-4 WAVE SPEED REV.0-APRIL l':H34 4000 035 00030000 I--.;;;;:a

......

u w Ul I N E-<250 00 r:x.::>.: en H ()CJ x 2000 0

J H r:x.150 a 0 H X10 000 o 200 400 600 800 1000 1200 ENTHALPY h o (BTU/LBM),..------------------.

LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-5 MASS FLUX.MOODY STEADY SLIP FLOW REV.0-APRIL

"'--USING PARA.6.A.3 r------------,II......-.,.?"-_..._..._....j\TOTAL (SEE PARA.6.A.4)c;'----------------

?'7 7:;;-:;.":;;-:;;00" TIME LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-6 BREAK FLOW VS.TIME REV.0-APRIL 1984 RELATIVE DISPLACEME NT OF PIPE END LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT TOTAL DISPL.OF PIPE END=DISPLACEMENT TIMES-/L+L)I2+RELATIVEL I'DISPLACEMENT RECI RCU LATI ON SUCTION LINE DISPLACEMENT OF PIPE AT RESTRAINT, D FIGURE 6.A-7 NOMENCLATURE FOR TIME HISTORY COMPUTER PRINTOUT ORIGINAL t OF PIPE I I ,........,/--OF MOVING PI PE I 1/VESSEL SAFE END\......1 REV.a-APRIL 1984 I e I+,9 20 t 21 14 15+,8 t 23 0-t25 e 0)I@I e C0<0 t26 I@I t 27 I@I t28 f17 18 I@I j 12 t13 0)t29 e.,.leI l 9 t30 10 11 0 I 0 7 8 (0 5 6 0 CD 3 t4 0+1 2 0 CD 31 32 LA SALLE COUNTY.STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-8 FEEDWATER LINE SYSTEM NODALIZATION

-LEG EA REV.a-APRIL 1984 I (0 I+19 21 t20 14 15 t16 I G8).23 (0 I (19)l..-.25 I (2:0 I.26@@@8 I@I G.27 I (22)I+28 I (23)I.29 I (24)I+17 30 12+13 G I e j 9+18 10 11 0 0 I 7 8 0 I 5 6 (0 0 3 4 0)I+1 0 0)31 32 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-9 FEEDWATER LINE SYSTEM NODALIZATION

-LEG EB REV.0-APRIL 1984 0 STEAM.26*I 0 5 0- FEEDWATER.27 l3 5 I 0)12 114 I 1 7 17 e 16-.....8 0 9!-000-I@0 21......-8....._20-....12@0 t-I.1 I......-......t1S I t 13 e 23 22 G@e (0 I--e 3 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-10 RECIRCULATION LINE SYSTEM NODALI ZA TI ON REV.0-APRIL 1984 0'" r-ei 8 r-ei 0'"<D d 8 N<D N d u z.., I 0 W It>It>Cl d iii 0-Of>0-0:: 8<<..'"..J N d w U 0:: Z Cl.., Z I 0 iii w It>.3 a..:s: iii d 0-0:t...J....:;)8..J 0.<<..]I-d 0 I-w 0:t'";: '" d 8'" d 0'"'" d 8 N d g d 8 d 0'" 0 d 0 0@...<<...(J)Cl<<w...!!!o o:t:...w:t w'"'" z iii:;)I..J WW>+:t:;)0.o N Jas/wql (£01 X)31.'0'1;1 MOl::!.....----------------.., LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-ll COMPARISON OF THE GE AND RELAP4/MOD5 METHODS-FEEDWATER LINE BREAK.LEG EA REV.0-APRIL 1984 0 tl)....0 8....0 g...'" ri.0 W<t I-..J<t w l-II:'"<.?)-Z...0 0 Vi N N<<::l u N W I Z U......., 0!aI Z tl)..., d 0 0 w I 0..J N LI.W::I:..J Vi 0....0..u;W<<:E:E I-0..8 0::l>W I-0..cr co<.?0<.?Z u;0;:)tl)I.....J 0 W tJ)tJ)w 8>+"I;]0..0:E::>w 0..:::ii 0 i=co"'l 0 8"'l 0 Iii N ci 8 N ci 0 tl)-d 0 0 dd 0 N LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-12 COMPARISON OF THE GE AND RELAP4/MOD5 METHODS-FEEDWATER LINE BREAK, LEG EB REV.0-APRIL 1984

_--.....------:r-------------------., q I7l-d o o J:W:lE w t:l Ul:::l Ul w IX: ""'o o::r:..,.4-:5 w\.....-d i In-d w::r: i=-d-d-d--d LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-13 COMPARISON OF THE GE AND RELAP4/MOD5 METHODS-RECIRCULATION LINE BREAK, FINITE OPENING TIME REV.0-APRIL 1984 HINGES 51 17 2 MASSLESS.....--.53 52 Nodes 50 Elements 3 Springs tttttl+l#Rigid Link LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-14 HORIZONTAL MODEL FOR ANNULUS PRESSURIZATION SHIELD B.RESULTANT FORCES A.PRESSURE DISTRI BUT ION CALCULATION OF FORCE I F.'.....-""'1 FORCE DESCRIPTION (ALL FUNCTIONS OF TIME)I.PRESSURE LOADS 2.PIPE RESTRAINT LOAD 3.JET REACTION FORCE 4.JET IMPINGEMENT FORCE LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.A-15 ANNULUS PRESSURIZATION LOADING DESCRIPTION REV.0-APRIL 1984 EL.755.29'£L.793.42'£L.783 83'EL.777.42 1 EL.77 2.73'EL.760.36'EL.767;8:;, 1 180 0£L.804.00'

/'-)/135 0 ,135'/-_.--/.90*I'90*..60*45 0 30*I__1_-_________I I-0'0"31 6(32 33 2-26\)27 2829 e 13-212223 9 2425-16 17 18 19 20'9 Z9 3;',r-..r-(&\..11<3 12'is: 1415---6 7 8 9 10*SJ&G-G 1 G9 2 0 3@4'2Y 5-0 0 r 5ao.c._-;.Br.90"" I I'_1/I 0 0 Upper Levels 1?0*/Indicate Pressure Indicate Pipe Locations Load Centers\,"

I O*Lower Levels IflO*/,,--[-I ,...I....., I I 1 I k:-W_90*\\i<o c";J::>or:z::z:;;0 c:[Tl mr o(J)();J::>.,>.....;;0;;0.....r ()Vl:z:r c"., f!:fll r;J::>.....;J::>n G">VlO-l m c I.....;;0 I o:z: m I:z: 0 me I 0 0'\ r:l:>.....r:l:>):>-l I z.....I mN-'z-<:l:>m:l:>OJ-l r(J);;0.....-<-l mo;::::>:l:>z;;<:: Vl-l., 0;;0 ,;:gZ);;0)-l.tIl<: o:J::l'1:1H t"1......co ,f;:>.!

o'Upper Levels Indicate Pipe Locations-EL.777.42'-EL.767.83'-EL.760.36'-EL.793.42'-EL.783.83'180*-EL.804.00'-EL.755.29'ls'Of/'.Pedestal J US*135 J 90 90*60' O*15'30'I I , , ,:.0....!'X Q<Fl-O@(25 26 27 28 (8;P/Q3)18 19 22.00 c<')(13 14 15 16 17)6<r, tX rg".......C5 r;I',- ,-,.'9 10 11 o 12.'9'5Z X 6(5 6 7 8 (>$8 t?3'9 2 4 (', I)O'"-45'/*90**90*Indicate Pressure Load Centers" 180'I/]180'I-f/'/I f\\"-"-O'Lower Levelso c"):> z;;:j>>z c 0(J)*'"TI):>'"TI>>rn;:0......r rn§;r o VI::::: "'"TI'fIl):>):>.....-in I:i1 U'l()rn rn c;;:;'0 Cl;:0;:0<j Z I"T1....0..........0 0)-<z Z):>/'T1*):>):>-1 J.....I§;-<co N....,;:0):>'-J/'T1-i"):>.......(/)-1;;:0<;:0.....>>:l:z U'l-;::l i'"TI;:0-oj 0 rnO;:0;:0-i I>

//W c#TOTALCOREFLOWWA*ACTIVE LOOP FLOW WI&INACTIVE LOOP FLOW LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.8-1 ILLUSTRATION OF SINGLE RECIRCULATION LOOP OPERATION FLOWS REV.0-APRIL 1984 "10 1000"40 0 1120.....<C C a::.....c*"00 II:<C III....!:!i I!1010 100.It;).....!l...=...II: f 5........1010.z:it...2......:>i..z c...II:.....z 1040<C!1020/110 I'lANOE 01'EX"ECTEO----tI""!MA lUMVt\l I l..OOI' O"EAATlON 1:10*10 10 100'OWa1'l LEVEL ,.IfueUAlllOlUl'I"ATIDI NO.....--......1....----'-----'---

....---.....---....---..o LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B*2 FEEDWATER CF WITH ONE PUMP OPERATION TYPICAL (GE)REV.13 LSCS-UFSAR

cl*..:--..1;.---....:.:----.....:_1:-.

_..w......."u.c:i§5\--+----+---..;.Jt:+1'----=l2 1l13ll:ftl IJ J.1GJl:l3cI1 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B-3 FEEDWATER CF WITH ONE PUMP OPERATION TYPICAL (GE)REV.13 LSCS*UFSAR

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""J,.O gC)CQ31tftl:Jj 1 N3J1:13.1 I LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B*4 TYPICAL LOAD REJECTION WITH ONE PUMP OPERATION REV.13

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"-_'"'*n--LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B-5 TYPICAL SEIZURE OF ONE RECIRCULATION PUMP REV.13 r..8cs-lJFSAR 1.2 ,.------.Ul.TIMATE STABILITY l.IMIT 1.0...........___----SINGl.E LOOP.PUMP MINIMUM SPEE'O--BOTH l.00PS.PUMPS MINIMUM SPHD 0.8 o t:<<II:>'l(u...Q 0.6 0..0.2 MIG HEST!'OWEFI ATTAINABl.E FOASINGLE LOOP OPE F1A o
  • For cycle specific decay ratios.SEE the LaSalle Administrative Technical requirements LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B-6 Typical, GE DECAY RATIO VERSUS POWER CURVE FOR TWO*LOOP AND SINGLE-LOOP OPERATION*

REV.13 LSCS*UFSAR

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...J....J.-L..J!!a'I LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 6.B-7 UNCOVERED TIME VS.BREAK AREA-LASALLE 1 AND 2 SUCTION BREAK LPCSlDG F AlLURE REV, 13...'"....g

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