RS-13-009, Clinton, Unit 1, Updated Safety Analysis Report, Revision 15, Chapter 3 - Design of Structures, Components, Equipment and Systems

Clinton, Unit 1, Updated Safety Analysis Report, Revision 15, Chapter 3 - Design of Structures, Components, Equipment and Systems

ML13016A315
Person / Time
Site:	Clinton
Issue date:	01/10/2013
From:	Exelon Generation Co
To:	Office of Nuclear Reactor Regulation
References
RS-13-009
Download: ML13016A315 (833)

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CPS/USAR CHAPTER 03 3.3-1 REV. 11, JANUARY 2005 3.3 WIND AND TORNADO LOADINGS 3.3.1 Wind Loadings 3.3.1.1 Design Wind Velocity A design wind velocity of 85 mph is used in the design of Clinton Power Station (CPS) Seismic Category I structures. This wind velocity is based on a mean recurrence interval of 100 years using the wind speed distribution map given in ANSI A58.1-1972 (See Figure 3.3-1). The effective velocity pressures of different heights are obtained using Table 5 for exposure C of ANSI A58.1-1972. The effective velocity pressures include a variation of wind velocity given by the following formula:

()7/130z30/zVV= (3.3-1) where V 30 = design wind velocity in mph at height of 30 feet above grade.

V z = wind velocity in mph at a height of z feet above grade. The effective velocity pressures also include the effect of gusts through the use of appropriate gust factors as specified in ANSI A58.1-1972. 3.3.1.2 Determination of Applied Forces The design wind velocity V is converted to velocity pressure using the formula:

30 2 30V00256.0q= (3.3-2) where q 30 = basic wind pressure in psf. The effective velocity pressures of winds for buildings and structures, q , and for parts and portions, q , at various heights above the ground are computed using the following formulas: 30fzfqGKq= (3.3-2a) 30pzpqGKq= (3.3-2b) where K z = the velocity pressure coefficient which depends upon the type of exposure and height z above ground, and G f , G p = gust factors which depend upon the type of exposure and dynamic response characteristics of the structure, or parts and portions thereof.

CPS/USAR CHAPTER 03 3.3-2 REV. 13, JANUARY 2009 Tables 5 and 6 of ANSI A58.1-1972 (Reference 1) provide the values of the effective velocity pressures for exposure C. The design wind pressure P is calculated by pimpCqqCP= (3.3-3) where q equals q f or q p whichever is appropriate, C p is the external pressure coefficient, q m is the effective velocity pressure for calculating internal pressures and C pi is the internal pressure coefficient.

The external and internal pressure coefficients, C p and C pi, recommended in ANSI A58.1-1972, are summarized in Table 3.3-1. For the containment wall and the dome, ASCE paper no. 3269, "Wind Forces On Structures" (Ref. 4), has been used to obtain the wind pressure coefficients as shown in Figure 3.3-2. 3.3.2 Tornado Loadings 3.3.2.1 Applicable Design Parameters Clinton Power Station is located in Region I, as defined by Regulatory Guide 1.76.

The following are the parameters of the design-basis tornado (DBT): a. a maximum tangential velocity of 230 mph at a radius of 150 feet, b. a maximum translational velocity of 46 mph, c. an external pressure drop of 1.2 psi at a rate of 0.5 psi/s acting upon fully enclosed areas, and d. a spectrum of tornado-generated missiles and their pertinent characteristics, as delineated in Subsection 3.5.1.4. 3.3.2.2 Determination of Forces on Structures All tornado wind pressure and differential pressure effects are considered static in application, since the natural period of the building structures and their exposed elements is short compared to the time rate of the applied design pressure. Venting has not been adopted as a design measure for Category I buildings. (Q&R 220.02)

The effects of the maximum credible tornado are translated into forces on the structures with the aid of a tornado model (Reference 3) that incorporates the design parameters of Subsection 3.3.2.1. The distribution of wind velocity with cyclonic radius for this model is represented by the following expressions: 0.1 R r For C (3.3-4)

CPS/USAR CHAPTER 03 3.3-3 REV. 11, JANUARY 2005

()t C C V R rVrV+= 0.1 R r For C (3.3-5) ()t C C V r RVrV+= where: r = Radius from the center of vortex (ft), V(r) = Tangential wind velocity at a radius r (mph),

R C = Radius to the maximum wind velocity (ft),

V C = Maximum tangential wind velocity (mph), and V t = Translational wind velocity (mph). The distribution of the pressure drop as a function of the cyclonic radius for the model is represented by the following expressions: 0.1 R r For C (3.3-6) ()=2 C C R r5.01Prp 0.1 R r For C (3.3-7) ()2CC r R 2 Prp= where: p(r) = Pressure drop (psi) at radius r, and P c = Pressure drop at center of vortex. All other terms previously defined.

CPS/USAR CHAPTER 03 3.3-4 REV. 11, JANUARY 2005 The distribution of the differential pressure and tangential velocity plus translational velocity as a function of the distance from the center of the tornado, as represented by the above expressions, is shown graphically in Figure 3.3-3. The tornado velocity is converted into an equivalent static pressure, using the equation 3.3-2. It is assumed there is no variation in velocity with height and that the gust factor is unity.

The pressure coefficients for windward pressure and for leeward, sidewall and roof suction given in Table 3.3-1 are used to determine tornado wind loading. The effective velocity pressure due to wind velocity alone is shown in Figure 3.3-4. Figure 3.3-5 shows the resultant static surface pressure when the pressure drop components and dynamic wind components are combined for rectangular flat-topped structures. For cylindrical and hemispherical structures the effective velocity pressure obtained from Figure 3.3-4 is used to obtain the pressure distribution shown in Figure 3.3-2 which is then combined with the pressure drop shown graphically in Figure 3.3-3 to obtain the resultant static surface pressure. The tornado-generated missile loadings are considered as impactive dynamic loads. The method adopted for designing structures for this impactive load is described in Subsection 3.5.3. The total tornado load is found by combining all three of the above-mentioned individual tornado loadings as follows:

i) W t = W w ii) W t = W p iii) W t = W m iv) W t = W w + W p v) W t = W w + W m vi) W t = W w + W p + W m where: W w = tornado wind pressure, W p = tornado differential pressure, W m = tornado missile load, and W t = total tornado load. Depending on the particular structure under consideration, the most adverse effects of these combinations on the structure are used to derive the total design tornado load. The total design tornado load is then combined with other loadings as per Tables 3.8-1 and 3.8-

2.

CPS/USAR CHAPTER 03 3.3-5 REV. 13, JANUARY 2009 When designing for the postulated tornado, the structure under consideration is placed in various locations of the pressure field to determine the maximum critical effects of shear, overturning moment and torsional moment on the structure. The intent of SRP No. 3.3.2 is to simplify the analysis and design of the structure for a tornado load based on the maximum wind velocity with the corresponding pressure drop without going into the detail of actual load distribution based on its location. On the Clinton Project, the overall analysis of structures for the tornado load is based on the distribution of curves as shown in Figure 3.3-3. At any particular location the total tornado load is equal to the tornado load plus pressure drop. Therefore, this can be formulated as shown in the fourth load combination equation. The overall effect of the total tornado load on the structure is maximized by shifting its tornado center along the structure. Although the load combination equations for the total tornado loads used on the Clinton Project differ from those given in SRP Section 3.3.2.II.3(d), they are acceptable for determining the tornado loads for overall design of the structure. The sixth load combination equation will be revised as follows based on the above explanation:

()m pwtWWWWvi++= For determining the tornado loads for the local design of structures on the Clinton Project, a maximum wind pressure plus its corresponding pressure drop (1.2 psi) as shown in Figure 3.3-3 were used. Although the load combination equations used for total tornado loads differ from those equations given in SRP Section 3.3.2.II.3(d), their numerical values are identical:

SRP equation m pwtWW5.0WW++= ()()m tW14435.0232W

++= m tW448W+= Clinton equation m pwtWW0.1WW++= ()()m tW1445.10.1232W

++= m tW448W+= W m is the same in both equations. (Q&R 220.01)

CPS/USAR CHAPTER 03 3.3-6 REV. 11, JANUARY 2005 3.3.2.3 Effects of Failure of Structures and Components Not Designed for Tornado Loads The turbine building, radwaste building superstructure, circulating water screen house steel superstructure and service building are non-Seismic Category I structures. The turbine building is constructed on reinforced concrete up to elevation 800 feet 0 inch, and structural steel with metal roof decking and metal siding above elevation 800 feet 0 inch. The radwaste building superstructure is constructed of reinforced concrete. The service building is constructed of metal roof decking and precast concrete wall panels. The turbine building siding and roof decking is designed to blow off in an approaching tornado, and the structure is designed to withstand tornado loads on the exposed structural frame and the remaining siding. The remaining siding is assumed to wrap around the center girts (see Figure 3.3-7). Thus, the integrity of the turbine building under the design-basis tornado is assured. The missiles generated in the blowoff of siding and roof decking of the turbine building are evaluated to be less damaging than the postulated tornado-generated missiles discussed in

Subsection 3.5.1.4. The steel superstructure of the circulating water screen house is not designed to withstand DBT. The Seismic Category I portions of this circulating water screen house are designed to withstand the effects of collapsing steel superstructure under DBT. The vent stack on top of the diesel generator and HVAC building is designed to withstand DBT. The vent stack has been designed as a stiffened box girder to resist an equivalent static load due to the 360-mph wind speed of the design tornado with allowable stresses as delineated in Subsection 3.8.4.5. External missiles are considered by allowing them to penetrate the stack shell without impairing the overall integrity of the stack. (Q&R 220.04) Turbine building siding is designed to blow in or out within a pressure range of 50 psf to 90 psf. The turbine building siding is composed of double span blow-in/blow-out panels consisting of face and liner sheets interconnected by subgirts. There are no fasteners used at the lap joints between the adjoining face sheets to permit blow-in or blow-out (see Figure 3.3-6 sketch A). Liner sheet is fastened to the girt at center support with screws penetrating the subgirt. At panel end supports, 3-inch wide, 12-gauge plate clamps down both liner ends. No screws penetrate the liner sheet (see Figure 3.3-6 sketch B). Under wind pressure, because of large displacement, the liner ends slip out of the clamping plate and start to wrap around center support or start to bend away from center support for inward or outward pressure respectively (see Figure 3.3-7 sketches C and D). After the siding panel collapses, it remains anchored to, and wrapped around the center support or remains anchored and bent outward for inward or outward pressure respectively (See Figure 3.3-7 sketches C and D). Turbine building roof deck is designed to blow out within a pressure range of 55 psf to 70 psf for corners, and 40 psf to 60 psf for all other areas. This is assured through the use of pressure release fasteners. (Q&R 220.03) The radwaste building superstructure has been designed for the SSE and tornado wind loads. Its failure is not deemed to be a credible design condition. The mass of the service building is CPS/USAR CHAPTER 03 3.3-7 REV. 11, JANUARY 2005 small relative to that of the adjacent Category I structures. Therefore, its failure in the event of the SSE will not add sufficient lateral force to affect the integrity of the Category I Structures. The CGCB has been designed to withstand the SSE. (Q&R 220.05)

3.3.3 References

1. American National Standards Building Code for Minimum Design Loads in Buildings and Other Structures, ANSI A58.1-1972, Section 6. 2. Design Basis Tornado for Nuclear Power Plants, USAEC Regulatory Guide 1.76.

3. J. D. Stevenson, "Engineering and Marketing Guide to Tornado, Missile Jet Thrust and Pipe Whip Effects on Equipment and Structures," Appendix B,:1-3, Report prepared for: Nuclear Structural Associates, Pittsburgh, Pennsylvania. 4. ASCE Paper No. 3269, " Wind Forces on Structures", Transaction of the American Society of Civil Engineers, Vol. 126, Part II, 1961.

CPS/USAR CHAPTER 03 3.3-8 REV. 11, JANUARY 2005 TABLE 3.3-1 ANSI WIND PRESSURE COEFFICIENTS ITEM EXTERNAL COEFFICIENT(C p) INTERNAL COEFFICIENT(C pi) Windward wall 0.8 +0.3 Leeward wall

-0.5 -0.3 Side wall

-0.7 -0.3 Roof -0.7

-0.3 Corners -2.0 -0.3 Cylinders and spheres 0.5 -0.3 CPS/USAR CHAPTER 03 3.4-1 REV. 12, JANUARY 2007 3.4 WATER LEVEL (FLOOD) DESIGN The Probable Maximum Flood (PMF) level elevation at the Clinton Power Station (CPS) site is 708.9 feet MSL. Refer to Subsection 2.4.3.5 for a complete description of this calculation. For local intense Probable Maximum Precipitation (PMP) refer to Subsection 2.4.2.3. Compliance to Regulatory Guide 1.59 is discussed in Subsection 2.4.3.1. For design purposes, the groundwater table at the CPS site is conservatively taken as elevation 730 feet. For actual groundwater elevations, refer to Subsection 2.4.13.

3.4.1 Flood Protection 3.4.1.1 Flood Protection Measures for Seismic Category I Structures The effects of probable maximum flooding do not influence the station design because of the large difference in elevation between station grade (736 feet) and probable maximum flood elevation (708.9). However, the circulating water screen house is designed to withstand the effects of flooding. The following protection measures are adopted for Seismic Category I systems and components located in the circulating water screen house and located below the probable maximum flood level. The flood protection arrangement of the Circulating Water Screen House is shown in Figure D3.6-134. a. Water stops are provided in all construction joints up to the maximum flood level. b. Water seal rings are provided for all penetrations in exterior walls below the maximum flood level. c. Watertight doors designed to withstand the hydrostatic head of the maximum flood level are provided for all doorways located on both the entrance walls and the internal walls of the SSW pump rooms which are below the maximum flood level. d. A hatch is provided on the roof of the essential service water pump structure (elevation 730 feet) for access during PMF. The measures listed above are not required (except as noted below) at the CPS site, because grade at the station site is 27.1 feet above the PMF level. However, the measures are adopted for the portions of the structures at the station site located below the maximum groundwater level. ECCS pump cubicles (RHR, LPCS, HPCS) located on the 707.5-foot elevation of the auxiliary and fuel (HPCS) building are protected from internal floods. All the walls, penetrations and doors in these cubicles are watertight to elevation 731 feet 5 inches (high water level in suppression pool). Watertight doors are shown on drawing M01-1105, sht 5. Also accounted for in the design is that flooding in one cubicle as a result of the rupture of a pump suction line from the suppression pool will not result in the flooding of other cubicles.

CPS/USAR CHAPTER 03 3.4-2 REV. 11, JANUARY 2005 3.4.1.2 Permanent Dewatering System The plant structures are designed to withstand the effects of groundwater conditions at the site. Therefore, a permanent dewatering system is not required and has not been installed. 3.4.2 Analysis Procedures All substructures below elevation 730 feet 0 inch at the CPS site are designed to withstand full hydrostatic head of groundwater. The walls in the circulating water screen house that are exposed to floodwater are designed for hydrodynamic forces also. The wind wave forces are determined in accordance with the "Shore Protection Manual," Volume II (U.S. Army Coastal Engineering Research Center, Department of the Army Corps of Engineering, 1973 and later Editions). The hydrodynamic loads resulting from the seismic forces are determined in accordance with the procedures delineated in "Dynamic Pressures on Fluid Containers," Nuclear Reactor and Earthquakes, TID 7024, USAEC (August 1963). The structural stability of Seismic Category I structures with flotation, overturning and sliding is investigated under the combined effects of the PMF and the wind wave forces. These criteria are the same as those given in Subsection 3.8.5.5.2. (Q&R 220.06)

CPS/USAR CHAPTER 03 3.5-1 REV. 11, JANUARY 2005 3.5 MISSILE PROTECTION 3.5.1 Missile Selection and Description 3.5.1.1 Internally Generated Missiles (Outside Containment) All safety-related structures, systems and components in the auxiliary, radwaste, control, fuel, diesel generator and HVAC and circulating water screen house buildings are protected, to the extent practical, from the effects of postulated internal missiles. This is achieved by proper equipment layout where possible, otherwise suitable physical barriers are provided to isolate the missile or to shield the critical system or component. Missile selection is done for pressurized as well as rotating type equipment. For pressurized equipment the following potential missiles are investigated: a. valve bonnets (large and small), b. valve stems, c. thermowells, and d. pressurized vessel head bolts. The pressurized missiles are discussed in detail in Section 3.5.1.7. For rotating equipment which has a potential for being subjected to an overspeed condition in excess of design limitations the potential missiles investigated are as follows: a. pump blade, b. pump impeller, c. small flanges, and d. coupling bolts. 3.5.1.2 Internally Generated Missiles (Inside Containment) All safety-related structures, systems and components within the containment are protected to the extent practical from the effects of postulated internal missiles. This is achieved by proper equipment layout where possible, otherwise suitable physical barriers are provided to isolate the missile or to shield the critical system or component. The following potential missiles from pressurized equipment were investigated: a. valve bonnets (large and small), b. valve stems, CPS/USAR CHAPTER 03 3.5-2 REV. 11, JANUARY 2005 c. thermowells, and d. vessel head bolts. These pressurized missiles are discussed in detail in Section 3.5.1.7. In addition, the potential missiles investigated for rotating equipment are from the following: a. pump blades, b. pump impellers, c. small flanges, and d. coupling bolts. In all cases investigated, it was shown that the kinetic energy of the potential missile was contained by the strain energy capacity of the equipment casing. Thus, it is demonstrated that the postulated missile is contained within the equipment casing. The most substantial pieces of NSSS rotating equipment are the recirculation pump and motor. This potential missile source is covered in detail in Reference 9. It is concluded in Reference 9 that destructive pump overspeed can result in certain types of potential missiles, but no damage is possible to any safety-related equipment because these missiles would not escape from the interior of either the pump or the motor. With regard to evaluation of the probabilistic consequences of pump impeller missiles ejected from pipe breaks, it is concluded in Attachment 3 of Reference 9 that no damage is possible to primary containment, any major piping system, or an inboard main steam isolation valve. Absence of damage is because trajectories of postulated missiles do not interact with these systems. The above discussion demonstrates that the probability of significant damage from recirculation pump or other motor missiles is so low that no protection other than pipe restraints is recommended. The following potential missiles due to gravitational forces have been investigated: a. Systems, components, and structures classified as seismic Category I are designed to withstand the applicable dynamic loads without failing. Therefore, they are not considered potential gravitational missile sources. b. Non-seismic Category I items inside the containment are designed as follows: 1. Structural Non-Category I classified items are designed not to fall; therefore, they cannot become a gravitational missile during reactor operation and following a LOCA. 2. Lighting fixtures which are Non-Category I classified will be restrained in areas of safety-related items.

CPS/USAR CHAPTER 03 3.5-3 REV. 11, JANUARY 2005 3. Non-Category I classified piping and piping supports have been designed to assure that the Class D piping and other non-safety piping will not fall or interact to impair the capability of the essential systems to perform their intended functions. 4. HVAC ductwork and hangers located in the Containment Building have been designed seismically with minor exceptions where failure of the duct system could not affect safety-related components. 5. Mechanical Non-Category I classified items such as unit heaters, area coolers, and HVAC instrumentation are designed seismically and, therefore, not considered gravitational missiles. Subsections 3.5.1.1 and 3.5.1.2 have been supplemented by responses given to Grand Gulf Questions 211.13, 211.14, 211.15, and 211.16. We have evaluated the effect of postulated missiles, blades, impellers, and shafts generated by typical rotating equipment such as pumps, fans, compressors, and turbines. It was concluded that in the unlikely event that internal missiles were generated, they would be contained by the equipment housing. One exception to the above protection analysis is certain fans supplied by Buffalo Forge Company, because calculations show that missiles could escape their housings. To protect essential equipment from internally generated missiles, fan housing reinforcement to prevent missile penetration, was provided for fans lVT06CA, lVT06CB, lVR04CA, lVR04CB, OVA05CA, OVA05CB, OVL02CA, OVL02CB, 1VF03CA, lVF03CB, OVA04CA, OVA04CB, OVW03CA, OVW03CB, lVR03CA, lVR03CB, lVT03CA, and lVT03CB. Similarly we postulated missiles such as nuts, bolts, and valve stems from failure of pressurized components and conclude that the energy content of such missiles would be insufficient to cause damage or failure to safety-related equipment, components, or structures. (Q&R 410.1) 3.5.1.3 Turbine Missiles The Nuclear Regulatory Commission (NRC) requires consideration of the effects of turbine missiles on the operation of nuclear power plants. A probability-based analysis is used to demonstrate protection against the effects of turbine missles at CPS. A proprietary GE report to the NRC dated January 1984 and entitled, "Probability of Missile Generation in General Electric Nuclear Turbines," (NUREG 1048, Appendix U) described GE's methodology for evaluating the probability of wheel missle generation for nuclear turbines manufactured by GE. The methodology includes consideration of the probability of unit overspeed, wheel materials, in-service inspection capabilities and the potential for wheel containment by stationary turbine structures. The analysis methodology considers two fundamental failure modes that can lead to missile generation, brittle fracture failures and ductile tensile failures. The turbine originally supplied to CPS utilized conventional built-up rotors with shrunk-on wheels and axial keyways. The original built-up rotors have been replaced with rotors manufactured from monoblock forgings. The following is a discussion of the two fundamental failure modes and their applicability to the monoblock rotors for CPS:

CPS/USAR CHAPTER 03 3.5-4 REV. 11, JANUARY 2005 Brittle Fracture Failure- For the original built-up rotors with shrunk-on wheels operated in the speed ranges considered by GE, the probability of bursting, and thus of missile generation, is dominated by the brittle fracture mechanism. However, the replacement rotors for CPS are of monoblock construction and do not have shrunk on wheels. Therefore, the formerly dominant brittle fracture failure mechanism is not applicable to the new rotors. Ductile Failure- The probability of ductile failure for a rotor of any type is considered to be a function of speed, temperature and material tensile strength. With stress below ultimate strength, the probability of a ductile failure is negligible. The brittle and ductile failure modes are statistically independent. The GE probabilistic analysis of turbine overspeed was also documented in the 1984 NRC report, and is applicable to units with LP monoblock rotors. The overspeed analysis considers the characteristics of the turbine control system, the unit configuration, and test requirements for the steam valves and other overspeed protection devices. This overspeed analysis showed that the probability of attaining a given overspeed decreases rapidly as the overspeed increases. As long as the control system is maintained in accordance with GE's recommendations (discussed below), the annual probability at CPS of attaining an overspeed of 120% or greater is 1.7 x 10-6. To keep the probability of a significant overspeed event very low, periodic maintenance and inspection of valves and other overspeed protection components are required. The intervals are established to maintain system reliability. The 1.7 x 10-6 per year probability assumes the longest permissible interval between valve inspections and would be lower with more frequent inspections. The NRC has developed guidelines that limit the maximum annual probability for various hypothetical events. In the case of CPS, the limit for the annual probability for generation of a turbine missile is 1 x 10-4 (NUREG 1048, Appendix U, Table U-l). Since the CPS probability estimate of 1.7 x 10-6 is below the NRC threshold for probability of missile generation, protection against missile generation for the replacement CPS rotors can be shown by avoiding the potential for ductile failure at any operating speed below 120%. GE has evaluated the tensile stresses in designing the rotating components of the monoblock turbine for CPS. All of the rotating components have sufficient margin to tensile strength at design component temperatures to support operating speeds well in excess of 120% of normal. For example, the overspeed capability of the un-bucketed HP and LP rotors is over 200%. The limiting components, per design, for the bucketed rotors are the LP L-0 buckets which have overspeed capability of 170%. For the CPS rotors, the probability of attaining an overspeed of 120% is at or below 1.7 x 10-6 per year and there is a negligible probability of ductile failure at 120%. Therefore, the probability of turbine missile generation caused by ductile failure is well below the maximum NRC probability of 1 X 10-4 per year and may be ignored.

CPS/USAR CHAPTER 03 3.5-5 REV. 13, JANUARY 2009 3.5.1.4 Missiles Generated by Natural Phenomena Tornadoes are the only natural phenomenon occurring in the vicinity of CPS that can generate missiles. Tornado impact velocities are obtained from the Tennessee Valley Authority Report TVA-TR74-1. The mathematical model in this report is developed on Hocker's observation of the Dallas tornado of April 2, 1957. The parameter values of this model are consistent with those of the design-basis tornado (DBT) specified in Regulatory Guide 1.76. The tornado missile trajectory is calculated by solving the equations of motion with the assumptions that (1) a potential missile does not tumble during the short period of time between its at-rest position and the point of its injection into the tornado wind field, and (2) the missile moves in a tumbling mode beyond its point of injection. The object that eventually becomes a tornado missile is assumed to be injected into the tornado wind field in the aerodynamic mode. The sample of missiles covers the wide spectrum of objects which might be windborne by a tornado and is listed in Table 3.5-3. The impact velocities of these missiles resulting from the design basis tornado (see Subsection 3.3.2) are shown in Table 3.5-3. Missiles A, and B, are considered at all elevations, and missile C is postulated at elevations up to 30 feet above grade level. These missiles are assumed to be capable of striking in all directions. Table 3.5-5 lists (in general terms) the protection systems and/or component and the barrier for its protection from tornado missiles. Table 3.5-6 lists the structures/barriers, the concrete thickness strength, and the curing time on which the strength is based. 3.5.1.5 Missiles Generated by Events Near the Site Based on a review of the nearby industrial, transportation, and military facilities (as described in Section 2.2), it is concluded that there are no potential missiles resulting from accidental explosions in the vicinity of the site. 3.5.1.6 Aircraft Hazard The airports and airways in the vicinity of CPS are described in Subsection 2.2.2.5. The aircraft hazard does not constitute a design-basis event because of the following reasons: a. There are no airports within 10 miles with projected operations greater than 500d2, nor are there any airports outside of 10 miles having the number of operations per year greater than 1000d2 (where d = distance in miles from the plant). Only two airports are known to accommodate commercial aircraft, the Decatur and Bloomington-Normal Airports. The Bloomington-Normal Airport is 22.5 miles north northwest of the plant and is served by American Eagle, American Flagship, Northwest Airlink, and TW Express. Aircraft common to this airport are the Jetstream 31, Saab 340, DeHaviland 8, ATR 41 and 72, private jets, helicopters, and general aviation. Occasionally, charter flights will bring in Boeing 727 and 737, and McDonnell-Douglas DC-9. The number of operations (takeoffs and landings) during the 1994 calendar year was 89,457. The Decatur Airport is 24 miles south of the plant. It is normally served by American Eagle and TW Express, private and military helicopters, private jets CPS/USAR CHAPTER 03 3.5-6 REV. 11, JANUARY 2005 and general aviation. The number of operations during the 1994 calendar year was 61,212. The Rantoul National Aviation Center, formerly the Chanute Air Force Base (37 miles east northeast of the plant) is a municipal airport with an I5 (instrument classification) rating with daily flights of Citation 5s through general aviation. There are no regular commuter flights at this time. b. Two of the private airstrips within five miles of the station (see Subsection 2.2.2.5) are: Martin Airstrip located about 4.5 miles south of the station, and Thorp Airstrip located about 4.75 miles northwest of the station. Information obtained from the owners indicates that both airstrips are used for personal use only, and if needed for emergency. Martin has one turf runway 2,000 feet long oriented north-south and averages about 4 to 6 operations per week. Thorp has two grassy runways 1,500 feet long each, oriented north-south and east-west, and averages about 4 to 5 operations per week on each runway. There are no commercial flights at these airports (Aircraft Owners and Pilots Association, 1975). The Springfield office of the Division of Aeronautics states that the four airports within a 10-mile radius of CPS are small private airstrips that have had no recorded accidents in the last two decades. The effective area of Category I Buildings is 0.00842 square miles and the probability of a fatal crash is taken as 1.2 x 10-8 per square mile per aircraft movement (SRP 3.5.1.6). Using the procedures in SRP 3.5.1.6, the probability of an aircraft crashing into the Category I Buildings is 0.315 x 10-7 per year for the Martin Airstrip and 0.525 x 10-7 per year for the Thorp Airstrips. The total impact probability by aircraft from the two airstrips is 0.42 x 10-7 per year, per unit, which meets the acceptance criteria of SRP 2.2.3. c. The four low altitude federal airways described in Subsection 2.2.2.5 are: Airway V313 with an average daily traffic of 20 flights; Airway V233 with an average daily traffic of 20 flights; Airway V434 with an average daily traffic of 15 flights; and Airway V72 with an average daily traffic of 10 flights. The actual width of each of these five airways is 8 nautical miles, or 9.21 statute miles. Calculation of the probability of impact on the station by aircraft flying along these airways required evaluation of inflight crash rates (C), projected number of flights per year (N), effective area of the station (A), and width of airway (plus twice the distance from the airway edge to the station when the station is outside the airway) (W). As recommended by SRP 3.5.1.6 for light (less than 100 flights per day) commercial traffic, a value of C = 3 x 10-9 per aircraft mile has been used for each airway. The values used for N account for an increase in air traffic of about CPS/USAR CHAPTER 03 3.5-7 REV. 11, JANUARY 2005 89% during the 40 years life expectancy of the station. The 89% increase over the 40 year period is equivalent to the 21% increase in air carrier aircraft operations in FAA Aviation Forecasts for the 1980-1992 period. The effective area (A) used in the calculations include the actual plant area, a shadow area upon horizontal plane and a skid area of and around the Category I Buildings. The shadow area and the skid area and consequently the effective area depend on the orientation of the airways with respect to the buildings of the station. The distance from center of the station, the present yearly traffic (Np), the values of N, A, W, and the probability of impact (PFA), on the Category I Buildings obtained for each airway are in Table 3.5-7. The sum of the four individual probabilities is 0.54 x 10-7 per year per unit, which meets the acceptance criteria of SRP 2.2.3. d. There are three Guard aviation units within about fifty miles of the plant. There is an Army National Guard unit at the Decatur Airport (24 miles south) with an assortment of ten helicopters. There is an Air National Guard unit at the Springfield Airport (52 miles southwest) with F-16 fighters. And there is another Air National Guard unit at Peoria Airport (58 miles northwest) with C-130 aircraft. 3.5.1.7 Internally Generated Missiles From Pressurized Components The station has been designed to assure that in the event of an internally generated pressurized missile: 1. There is no loss of containment function. 2. The reactor coolant pressure boundary is not breached as a result of postulated missile generation. 3. There is no loss of function to systems required to shut down the reactor and maintain it in a safe shutdown condition or to mitigate the consequences of such a postulated event: a. No equipment in one safety-related division is allowed to be damaged by a missile generated by equipment in another division. b. Missiles generated from non-safety related equipment shall not damage any safe shutdown equipment. c. Loss of offsite power is assumed concurrent with missile generation. Those systems required to perform safety functions for this event are listed in Table 3.6-1. A system is considered capable of generating a missile if, during normal plant conditions as defined in Subsection 3.6.1.1.1.d, it meets either of the following criteria for more than 2% of the time that the system is in operation: 1. The maximum operating temperature exceeds 200° F. 2. The maximum operating pressure exceeds 275 psig.

CPS/USAR CHAPTER 03 3.5-8 REV. 11, JANUARY 2005 Those systems capable of missile generation are shown on Table 3.6-2. The sections of those systems capable of missile generation are further defined as the "dotted" lines in Figure 3.6-1. The methods used to protect safety-related equipment from potential missiles generated internally due to pressurized components are as follows: 1. Provide design feature on pressurized equipment to prevent missile generation. 2. Locate high energy systems in separate missile-proof rooms. 3. Locate redundant system or components outside of the missile range and trajectory. 4. Orient the potential missile source to prevent unacceptable consequences of missile generation. The following is the method used to select potential missiles generated by high energy lines as defined above: 1. Thermometers or other detectors installed on piping or in wells are evaluated as potential missiles if the failure of a single circumferential weld would cause their ejection. 2. Unrestrained sections of piping such as vents, drains, and test connections are evaluated as potential missiles if required due to postulated pipe breaks shown in Figures B3.6-1 through B3.6-34. 3. Valves of ANSI 900-pound standard rating and above and some valves of ANSI 600-pound standard rating, constructed in accordance with the ASME Code, Section III, are pressure seal bonnet-type valves. For these valves, the bonnets are prevented from becoming missiles by the retaining ring, which would have to fail, and by the yoke, which would capture the bonnet or reduce bonnet energy. Because of the highly conservative design of the retaining ring of these valves, bonnet ejection is highly improbable. Hence, bonnets are not considered credible missiles. 4. Most valves of ANSI rating 600-pounds standard and below are valves with bolted bonnets. Valve bonnets are prevented from becoming missiles by limiting the stresses in the bonnet-to-body bolting material by the rules set forth in the ASME Code, Section III, and by designing the flanges in accordance with the applicable code requirements. Even if bolt failure were to occur, the likelihood of all bolts experiencing a simultaneous complete severance failure is very remote. The widespread use of valves with bolted bonnets and the low historical incidence of complete severance failure of the entire valve bonnet confirm that bolted valve bonnets are not credible missiles. 5. Single nuts, bolts, nut and bolt and nut and stud combinations are not credible missiles since all the valve body to bonnet connections are bolt through connections with a nut on one side. Also, all studs and bolts on primary system pressure boundary equipment are designed to ASME Section III standards and CPS/USAR CHAPTER 03 3.5-9 REV. 11, JANUARY 2005 are torqued to the stress levels allowed by ASME Section III. At these stress levels, the stored strain energy in the studs or bolts is too low to permit the stud or bolt to become a missile. 6. Valve stems are not considered as potential missiles if at least one feature in addition to the stem threads is included in their design to prevent ejection. Valves with backseats are prevented from becoming missiles by this feature. In addition, air and motor operated valve stems are effectively restrained by the valve operators. No credible valve stem missiles were identified on Clinton. 7. Dead end flanges are evaluated as potential missiles if required due to postulated pipe breaks shown in Figures B3.6-1 through B3.6-34. Our analysis of thermowells shows that the energy associated with either a piston type or jet propelled type missile is low and will not cause damage to any essential components. Piston-Type Missile The velocity of a piston-type missile (e.g., valve stem, etc.) is calculated by assuming that the missile, with no losses of energy due to friction, air resistance, etc. Work is the integral of force times displacement, while the kinetic energy of the missile is one-half the produce of missile mass times the square of the missile velocity. Assuming the force constant (and equivalent to PAO) and equating the kinetic energy to the work done results in a missile velocity given by the expression (Reference 10). 2/1Og/wAP2V= (3.5-6) where P = Pressure acting on area AO (lb/ft2) AO = Area of missile under pressure (ft2) = Displacement of length of "piston" stroke (ft) W = Weight of the missile (lb) V = Velocity of the missile (ft/sec) g = Acceleration of gravity (ft/sec2) Jet-Propelled Missiles Jet-propelled missiles (valve bonnets and thermowells) are missiles propelled by fluid escaping from a pressurized system in which there is essentially no lateral constraint on the fluid. Thus, the escaping jet of fluid will not only impinge on the missile during the period of missile acceleration, but will also flow around and past the missiles. The velocity of such a missile is estimated by employ in the jet property solution as given by Moody (Reference 11) for saturated steam blowdowns. The work of Reference 11 was directed toward the prediction of blowdown thrust and jet forces on stationary targets; however, by making a few simplifying assumptions and apply in the CPS/USAR CHAPTER 03 3.5-10 REV. 11, JANUARY 2005 principle of momentum, this work can be applied to the determination of velocity-displacement relationships for jet-propelled missiles. The specific assumptions are: (1) the asymptotic properties of the jet exist over the entire region of travel of the missile; (2) the missile is completely surrounded by the fluid jet during its time of flight. Applying these assumptions and the principles of momentum to the relative velocity of the jet and the missile, the following expression results relating the missile displacement and velocity: ()=uVuVnAWyV/11/ (3.5-7) where W/A y = Distance traveled by the missile from the break (ft) W = Missile weight (lb) A = Frontal area of missile (ft2) u = Asymptotic velocity of jet (ft/sec) V = Asymptotic specific volume of jet (ft3/lbm) V = Velocity of missile (ft/sec) Neglecting friction we used equation 3.5-7 by expanding it in the power series neglecting higher order terms to solve for velocity of the missile. Missile velocities are calculated at three regions: (a) at the break, (b) at the asymptote or transition area, (c) at the fully developed flow region. These are calculated based on specific stagnation properties (temperatures and pressures). Those areas of the plant where high energy lines were located within the same room as safety-related components required for safe shutdown are identified in Table 3.5-8. These are the only areas of the plant where internally generated missiles could potentially impact essential components. In the control building, the fuel building and auxiliary building identified in Table 3.5-8, there are no potential missile sources from high energy lines which could impact essential components. All high energy lines in the remaining areas of the plant, the steam tunnel, the containment and the drywell were reviewed for potential missile sources. These potential missile sources and all the targets that they could hit are listed in Table 3.5-9. As shown in Table 3.5-9, no potential missile could hit and disable components which are essential for plant safe shutdown. Therefore, the design of the Clinton Power Station incorporates sufficient separation between divisional equipment to assure that the containment can be isolated, the reactor coolant CPS/USAR CHAPTER 03 3.5-11 REV. 11, JANUARY 2005 boundary is not breached, and the reactor can be safely shut down even with a loss of offsite power. 3.5.2 Structures, Systems and Components To Be Protected from Externally Generated Missiles 3.5.2.1 General Missile selection and description for those external missiles which, if generated, could damage plant structures, systems or components important to safety, are identified in Subsections 3.5.1.4 and 3.5.1.5. 3.5.2.2 Structures Providing Protection Against Externally Generated Missiles Seismic Category I structures are designed to withstand postulated external missiles, thereby protecting the systems and components located within. Openings in the Unit 1-Unit 2 (Unit 2 has been cancelled) interface walls of Category I structures are closed with tornado missile resistant concrete barriers. These concrete barriers are considered part of the Category I structure, and the missile proof walls are shown in Figure 3.5-3. Protective characteristics of Seismic Category I Structures are summarized in Table 3.5-6. 3.5.2.3 Barriers (Other Than Structures) Providing Protection Against Externally Generated Missiles Those structures, systems and components to be protected from externally generated missiles, and the missile barrier associated with each, are identified in Table 3.5-5. The missile barriers indicated are designed for tornado-generated missiles using the procedures given in Subsection 3.5.3. Structures which protect plant systems and components from missiles generated outside the plant are identified in Subsection 3.5.2.2. Protection for those safety related systems and components not located within Seismic Category I structures (i.e. outdoors) is also identified in Table 3.5-5. The location of missile barriers is shown in Figures 3.5-3 through 3.5-5. 3.5.2.4 Systems/Components Not Requiring Unique Tornado Missile Protection A limited amount of safety related systems and components located near penetrations in Seismic Category I structures or located outside of such structures are evaluated as not requiring unique tornado missile protection barriers. Two approaches were used in the evaluation (reference 15): 1. Certain safety related systems and components are screened out using the criteria of Regulatory Guide 1.117, Tornado Design Classification, including its Appendix, which together, detail the items that should be protected from the effects of tornadoes. The criteria in the Regulatory Guide are summarized as important systems and components required to ensure the integrity of the reactor coolant pressure boundary, ensure the capability to shut down the reactor and maintain it in a safe shutdown condition, and those whose failure could lead to radioactive releases resulting in calculated offsite exposures greater than 25% of the guideline exposures of 10 CFR Part 100 using appropriately conservative CPS/USAR CHAPTER 03 3.5-12 REV. 11, JANUARY 2005 analytical methods and assumptions. The safety related systems and components not required to support these Regulatory Guide 1.117 guidelines are evaluated as not requiring unique tornado missile protection. 2. "Important" systems and components (as discussed in Regulatory Guide 1.117) are generally protected. The limited amount of unprotected portions of important systems and components are analyzed using a probabilistic missile strike analysis as permitted in Standard Review Plan 3.5.1.4, "Missiles Generated By Natural Phenomena". This analysis is conducted to determine the total (cumulative) probability per year of missiles striking important structures, systems, and components due to postulated tornadoes. This information is then utilized to determine the specific design provisions that must be provided to maintain the estimate of strike probability below an allowable level. The allowable level established for the protection of such systems and components at CPS is consistent with the acceptance criteria in Standard Review Plan 2.2.3, "Evaluation of Potential Accidents", i.e., that a probability of occurrence of initiating events (those that could lead to potential consequences in excess of the 10 CFR Part 100 guidelines) of approximately 1x10-6 per year is acceptable if, when combined with reasonable qualitative arguments, the realistic probability can be shown to be lower. The CPS-specific acceptance criteria is that the total probability of tornado missiles simply striking an important system or component must be shown by analysis to be less than 1x10-6 per year. This CPS-specific criteria contains the following inherent conservatisms:

• It is assumed that an important system or component simply being struck by a tornado missile would result in damage sufficient to preclude it from performing its intended safety function, although this is not realistic for all cases.

• The analysis calculates the probability of tornado missiles striking penetration openings. The openings themselves are not targets. The true targets are the safety-related components located inside the buildings. Some of the missile types listed in Table 3.5-3 cannot enter the openings and damage the components.

• The missile population is conservatively estimated.

• All postulated missiles are conservatively estimated to have minimal restraints. The analysis uses an NRC approved methodology (Reference 13) developed by the Electric Power Research Institute (EPRI)(Reference 14). The methodology is implemented using the computer program TORMIS, which is described in Section 3.5.2.5. Should CPS evaluations using the TORMIS methodology provide results indicating that the probability of tornado damage exceeds the acceptance criteria of 1x10-6 per year, then unique barriers are utilized to reduce the total probability to below the acceptance criteria. Temporary removal of a protective feature is permitted under administrative controls, if removal is determined to be necessary.

CPS/USAR CHAPTER 03 3.5-13 REV. 11, JANUARY 2005 3.5.2.5 TORMIS Description TORMIS implements a methodology developed by the Electric Power Research Institute. TORMIS determines the probability of striking walls and roofs of buildings on which penetrations or exposed portions of systems/components are located. The probability is calculated by simulating a large number of tornado strike events at the site for each tornado wind speed intensity scale. After the probability of striking the walls or roof is calculated, the exposed surface area of the particular components are factored in to compute the probability of striking a particular item. The TORMIS analysis for the CPS site (Reference 15) is in accordance with the TORMIS program, as described in Reference 14, using site-specific parameters described below: 1. The probability of a tornado strike at CPS is based upon the broad region values associated with the Fujita F-scale. 2. The Fujita scale (F-scale) wind speeds are used in lieu of the TORMIS wind speeds (F-scale) for the F0 through F5 intensities. In addition, a wind speed range from 320 to 360 mph is used for the F6 intensity to correspond to the tornado wind speed described in Section 3.3.2.1. 3. A more conservative near-ground profile was used than the base case in TORMIS, resulting in a higher tornado ground wind speed. The profile has a ground wind speed equal to 82% of the wind speed at 33 feet (i.e., V0/V33=0.82). 4. The number of missiles used in the CPS TORMIS analysis is a conservative value for CPS-specific sources, such as laydown, parking, and warehouse areas. These are postulated by general walkdown information at CPS. 3.5.3 Barrier Design Procedure Two types of structural response to missile impact have been investigated, as follows: a. Local effect in the impacted area which includes estimation of the depth of penetration and, in the case of concrete barriers, the potential for secondary missiles by scabbing. b. Overall response of the barrier, which includes the calculation of deflection due to missile impact. Generally, all missiles (internal or external) are considered as impacting instantaneously with a very short rise time relative to the natural period of the impacting structure. Two types of barriers are designed to resist missile impact, as follows: a. Reinforced concrete barriers: The depth of penetration into a concrete barrier is calculated using either the modified Petry equation (Reference 5) or the modified NDRC formula (Reference 12). The material property constant used in the penetration formula is 4.76x10-3 ft3/lb. Concrete barriers are designed such that the missile penetrates no more than two-thirds of the thickness of the barrier, preventing scabbing (Reference 6). The overall deformation of the panel is investigated using methods presented in Reference 6. Reference 6 presents an CPS/USAR CHAPTER 03 3.5-14 REV. 11, JANUARY 2005 equation of motion which makes it possible to calculate an impact force time-history consistent with the calculated penetration depth. To establish the capacity of the barrier to absorb energy, the deflection due to static loads is first calculated. The deflection due to missile impact is then determined by integrating the equation of motion or by using a simplified expression adopted from the equation of motion. This is compared with the maximum allowable deflection (of allowable ductility ratio) per ACI 349. Elements encased in concrete with 5 inches of cover and buried in 4 feet of soil are investigated using the soil penetration equations given in the reference. This investigation showed that most of the missiles listed in Table 3.5-4 did not penetrate the four feet of soil cover. For those missiles with penetration depths of more than four feet, a striking velocity at the concrete surface is estimated based on energy balance, and a penetration depth into the reinforced concrete is calculated using Equation 3.5-5. The penetration depths thus obtained are found to be much less than the minimum avaiable depth of reinforced concrete cover. For elements with less than the required amount of soil or concrete cover, the probability analysis was performed. The analysis showed that the probability of a tornado generated missile striking or damaging these elements is less than 1x10-7. Therefore, the tornado missile is not considered a design basis for these elements.

REFERENCE:

Young, C. W., "Depth Prediction for Earth-Penetrating Projectiles," Journal of the Soil Mechanics and Foundations Division, ASCE, Vol. 95, No. SM3, Proc. Paper 6588, May 1969, pp. 803-817. (Q&R 220.11) b. Steel plate barriers - The thickness of steel plate required to resist the impacting missile is calculated using the Stanford formula (Reference 7). The overall structure response, including structural stability and deformation, is investigated using concepts and methods presented in Reference 8. The only steel barriers provided are those employed for tornado missile protection of pipe sleeves in exterior wall and roof. The design was based on the modified Stanford Research Institute (SRI) formula, as noted in subsection 3.5.3, with its inherent ductility requirements. The pipe sleeve caps are sufficiently similar to the test specimens used in developing the SRI formula; therefore, overall behavior is included in the results. (Q&R 220.10) It is to be noted that the location of the reactor shield wall inside the containment structure is such that no potential missile will strike the reactor pressure vessel. In Reference 6, the maximum displacement, Ym, of a structural element under the impact of a rigid missile of mass m is given by:

CPS/USAR CHAPTER 03 3.5-15 REV. 11, JANUARY 2005 2/12o2mV'DMmY+= (1) where: M = equivalent mass of structural element D' = penetration of missile into structure Vo = missile velocity prior to impact ciruclar frequency of structural element The method specified in SRP Section 3.5.3 for treating this problem is given in the reference cited at the end of this response. This reference specifies the equivalent static design load, F, as: KFFi= (2) ='D2VmFoi (3) o1V'D2t= (4) 11tT7.015.01tT12K+µ+µ= (5) where: Fi = peak value of impact force developing between missile and target t1 = duration of impact T = 2 µ = ductility ratio = yYm CPS/USAR CHAPTER 03 3.5-16 REV. 11, JANUARY 2005 y = static yield deflection, when element is idealized by an SDF dynamic system To compare the two methods, a square reinforced conrete panel 26 feet long and 1.5 feet thick, with reinforcement details and material properties as summarized in the attached Figure 3.5-6 was considered. The results obtained by the two methods for the response of this panel to the impact of a rigid mass at the center are summarized below. The missile corresponds to the utility pole described in Tables 3.5-3 and 3.5-4 of the USAR, with Vo = 241 ft/sec ft2.32seclb500,1m2= Head-on concrete area = 143.10 in2 For this missile, the modified Petry formula with Kp = 0.0032 ft3/lb yields a missile penetration value of D' = 5.976 in. The panel is approximately as an SDF system for response evaluation, with a mass of =gweightpanel32.0M where g = acceleration gravity. The period of the system is determined on the basis of a cracked section moment of inertia to be 0.0961 second. The yield force Qy and the corresponding panel deflection y are 358.5 kips and 0.680 inches, respectively. The maximum displacement for Method 1 (the method given in the USAR) is obtained directly from Equation 1: []inches37.1Y1methodmax= To obtain the maximum displacement for Method 2 (the method given in SRP Section 3.5.3), a trial and error procedure is used for Equations 2 through 5 to obtain the ductility parameter. This value is []68.22method=µ Therefore, []methodmaxY = 2.68 x 0.0680 = 1.82 inches The two results are within 25% of each other. To provide an indication of the difference between the results from the two methods when other ductility ratios are used, the velocity of the missile is arbitrarily increased from 241 ft/sec to 364 ft/sec. This yields a missile penetration CPS/USAR CHAPTER 03 3.5-17 REV. 11, JANUARY 2005 of 12 inches, which is 2/3 of the panel thickness (the maximum penetration permitted per the USAR). Under this condition, []inches08.2Y1methodmax= []48.52method=µ []2methodmaxY = 5.48 x 0.680 = 3.73 inches The maximum displacements agree to within 44%. It should be noted that in Equation 3 above, the work of the missile on the target during the penetration process has not been evaluated consistently with the linear variation of contact force and velocity with time, which is the starting point of derivation in the reference. When the equation is done consistently with these assumptions, Equation 3 changes to Equation 3*: 'D2Vm43Foi= (3*) The values obtained for Ymax for the two cases described above then become []sec/ft241Vfor12.1*Yomax2method== []sec/ft364Vfor24.2*Yomax2method== The ratios of 2method1methodmaxmaxY/Y are given below: Missile Velocity ft/sec Penetration Depth inches Max. Deflection Ratio of the Two Methods 241 5.9 1.22 364 12 .927 Since the underestimation of method 1 is less than 8%, the use of this method is considered acceptable. REFERENCE (From SRP 3.5.3): R. A. Williamson and R. R. Alvy, "Impact Effect of Fragments Striking Structrual Elements," Homes and Narver, Inc., Revised November 1973. (Q&R 220.07) 3.5.4 References 1. NUREG 1048, Supplement 6, Appendix U, "Probability of Missile Generation in General Electric Nuclear Turbines," (Hope Creek SSER 6), 7/86. 2. Not used.

CPS/USAR CHAPTER 03 3.5-18 REV. 11, JANUARY 2005 3. S. H. Bush, "Probability of Damage to Nuclear Components Due to Turbine Failure," Nuclear Safety, 14(3), pp. 187-201, May-June 1973. 4. Not used. 5. A. Amirikian, "Design of Yards Protective Structures," Bureau of Yards and Docks, Publication No. NAVDOCKS p-51, Department of the Navy, Washington, D.C., August 1950. 6. J. M. Doyle, M. J. Klein, and H. Shah, "Design of Missile Restraint Concrete Panels," 2nd International Conference on Structural Mechanics in Reactor Technology, Berlin, Germany, September 1973. 7. W. B. Cottell and A. W. Savolainen, "U.S. Reactor Containment Technology," ORNL-NSIC-5, Vol. 1, Chapter 6, Oak Ridge National Laboratory, August 1965. 8. R. A. Williamson and R. R. Alvy, "Impact Effect of Fragments Striking Structural Elements," Holmes and Naver, Inc., Revised November 1973. 9. GE Letter Report, "Analysis of the Recirculation Pump Under Accident Conditions," Revision 2, March 30, 1979. 10. R. C. Gualtney, "Missile Generation and Protection in Light Water Cooled Power Reactor Plants," USAEC Report ORNL-NSIC-22, September 1968. 11. F. J. Moody, "Prediction of Blowdown Thrust and Jet Forces," ASME Publication G9-HT-31, August 1969. 12. George E. Sliter, "Assessment of Empirical Concrete Impact Formulas," ASCE Journal, Volume 106, No. ST5, May 1980. 13. Letter, Rubenstein (NRC) to Miraglia (NRC) entitled "Safety Evaluation Report - Electric Power Research Institute (EPRI) Topical Reports Concerning Tornado Missile Probabilistic Risk Assessment (PRA) Methodology", dated October 26, 1983. 14. Twisdale, L.A. and Dunn, W.L., EPRI NP-2005, Tornado Missile Simulation and Design Methodology, Volumes I and II, Final Report dated August 1981. 15. Sargent & Lundy Report SAD-465, Revision 2, September 16, 1998, Tornado Missile Hazard Assessment for Clinton Power Station.

CPS/USAR CHAPTER 03 3.5-19 REV. 11, JANUARY 2005 Table 3.5-1 Has Been Deleted CPS/USAR CHAPTER 03 3.5-20 REV. 11, JANUARY 2005 Table 3.5-2 Has Been Deleted CPS/USAR CHAPTER 03 3.5-21 REV. 13, JANUARY 2009 TABLE 3.5-3 TORNADO-GENERATED MISSILES AND THEIR PROPERTIES MISSILE WEIGHT (lb) DIMENSIONS CdA/m (ft2/lb) HORIZONTAL IMPACT VELOCITY (Note 1) (ft/sec) A. Solid Steel Sphere 0.147 1 in. diameter 0.0166 26 B. 6-inch Schedule 40 steel pipe 287 6.625-in. OD x 15 feet long 0.0212 135 C. Automobile (Note 2) 4,000 16.4 ft x 6.6 ft x 4.3 ft 0.0343 135 Note 1. Vertical impact velocities are taken equal to 67% of the horizontal impact velocities. Note 2. The automobile missile is considered to impact at all altitudes less than 30 feet above all grade levels within 0.5 mile of the plant structures.

CPS/USAR CHAPTER 03 3.5-22 REV. 11, JANUARY 2005 Table 3.5-4 Has Been Deleted CPS/USAR CHAPTER 03 3.5-23 REV. 11, JANUARY 2005 TABLE 3.5-5 PROTECTED COMPONENTS AND ASSOCIATED MISSILE BARRIERS FOR EXTERNALLY GENERATED MISSILES COMPONENT BARRIER A. Protected Components Within the Plant 1. Reactor coolant pressure boundary and other safety-related equipment inside containment Containment structure, dry- well, internal structures, and beams 2. Emergency core cooling, containment spray, cooling water, ventilation, electrical, instrumentation, control, and other safety-related equipment in auxiliary building Containment building, auxiliary building, and internal structures 3. Control room and protected electrical, instrumentation, control, and ventilation equipment in control building Control building 4. Spent fuel pool Fuel pool walls, fuel building 5. Emergency diesel generators and diesel fuel oil system (1) Diesel-generator building 6. Shutdown service water pumps and associated piping SSW portions of the Circulating Water Screen House protects the SSW Pumps and associated piping. Some SSW piping is located beneath steel plates in construction openings which haven't been demonstrated to be complete missile barriers. The openings are analyzed in the TORMIS analysis, described in 3.5.2.4 and 3.5.2.5. 7. Portion of the reactor coolant pressure boundary in the auxiliary building Auxiliary building and auxiliary building steam tunnel B. Protected Components Outdoors 1. Electrical manholes (Category I) Protected by a reinforced 1-foot- thick concrete cover with steel plate manhole covers (1-inch-thick galvanized plate) 2. Electrical duct banks (Category I) Protected by a minimum of 5 inches of reinforced concrete, buried a minimum of 4 feet below finish grade with exceptions described in Section 3.5.3, item a. Reinforced concrete barriers.

CPS/USAR TABLE 3.5-5 (Cont'd) CHAPTER 03 3.5-24 REV. 14, JANUARY 2011 COMPONENT BARRIER 3. Control building a. Ventilation air intakes Control Building ventilation air intakes are protected by a minimum 2-foot-thick reinforced concrete missile barrier (see Figure 3.5-3) except for the VA and VR system air intakes which are protected by two layers of thick heavy duty grating. Other openings in the Control Building walls, which are not protected and are analyzed in the TORMIS analysis described in 3.5.2.4 and 3.5.2.5, are for unused Unit 2 air intakes. b. Ventilation air exhausts Protected by a minimum 2-foot-thick reinforced concrete missile barrier (see Figure 3.5-3) c. External access doors Doors are designed to withstand tornado missiles or they are protected by a minimum 2-foot-thick reinforced concrete missile barrier, except for steel roll-up door which is protected by two layers of thick heavy duty grating. 4. Auxiliary building a. Access doors Access doors to the Auxiliary Building are internal to the Turbine Building, Control Building, and Fuel Building and are protected by these buildings. 5. Diesel-generator building a. Ventilation air intakes Protected by a minimum 2-foot-thick reinforced concrete missile barrier (see Figure3.5-3) b. Ventilation air exhausts (for the diesel generator combustion air exhaust exception see Section 9.5.8.1.1.b.) c. Standby gas treatment system exhaust Protected by a minimum 2-foot-thick reinforced concrete missile barrier (see Figure 3.5-3) Barrier not required for those portions external to diesel-generator building CPS/USAR CHAPTER 03 3.5-25 REV. 11, JANUARY 2005 COMPONENT BARRIER d. Station common exhaust vent Barrier not required because missile trajectory through this opening cannot endanger safety-related structures,systems or components e. Access doors Doors are designed to withstand tornado missiles or they are protected by a minimum 2-foot-thick reinforced concrete missile barrier f. Diesel fuel oil storage tank fill lines Protected by diesel-generator building except as noted in Subsection 9.5.4.3 6. Fuel building a. Access doors Doors are designed to withstand tornado missiles or they are protected by a minimum 2-foot-thick reinforced concrete missile barrier CPS/USAR CHAPTER 03 3.5-26 REV. 11, JANUARY 2005 TABLE 3.5-6 CONCRETE BARRIER PARAMETERS Structures Minimum Concrete Thickness Design Strength at 91 Days (ksi) Auxiliary Building Walls 2'-0" 3.5 Auxiliary Building Roof 1'-6" 3.5 Fuel Building Walls 2'-0" 3.5 Fuel Building Roof 2'-0" 3.5 Control Building Walls 2'-0" 3.5 Control Building Roof 2'-0" 3.5 Diesel Generator Building Walls 2'-0" 3.5 Diesel Generator Building Roof 2'-0" 3.5 Containment Wall 3'-0" 4.0 Containment Dome 2'-6" 4.0 Circulating Water Screen House Walls 2'-0" 3.5 Circulating Water Screen House Roof 1'-6" 3.5 Note: Penetrations in exterior walls and roofs of Safety Related buildings are analyzed using the TORMIS analysis described in USAR Sections 3.5.2.4 and 3.5.2.5.

CPS/USAR CHAPTER 03 3.5-27 REV. 11, JANUARY 2005 TABLE 3.5-7 PROBABILITY OF AIRCRAFT IMPACT FROM FEDERAL AIRWAYS AIRWAY DISTANCE FROM THE STATION IN MI., d PRESENT TRAFFIC PER YEAR, Np PROJECTED TRAFFIC PER YEAR1240)21.1(xnNp= EFFECTIVE AREA IN SQ, MI., A (2 Units) WIDTH IN MI., w PROBABILITY OF IMPACT PER YEAR (2 Units) PFA=CxNxA/w V313 1.5 7,300 13,780 0.00842 9.21 0.378 x 10-7 V233 2.0 7,300 13,780 0.00842 9.21 0.378 x 10-7 V434 6.0 5,475 10,335 0.00690 12.0 0.178 x 10-7 V72 4.75 3,650 6,890 0.00690 9.5 0.150 x 10-7 TOTAL 1.084 x 10-7 Probability of Impact Per Year Per Unit = 0.542 x 10-7 NOTE: UNIT 2 HAS BEEN CANCELLED.

CPS/USAR CHAPTER 03 3.5-28 REV. 11, JANUARY 2005 TABLE 3.5-8 AREAS WHERE ESSENTIAL COMPONENTS AND HIGH ENERGY LINES OCCUR TOGETHER Drywell Containment Steam tunnel Auxiliary building Fuel building Control building CPS/USAR CHAPTER 03 3.5-29 REV. 11, JANUARY 2005 TABLE 3.5-9 POTENTIAL MISSILE SOURCES WHICH COULD IMPACT SAFETY RELATED COMPONENTS POTENTIAL THERMOWELL MISSILE LOCATION REMARKS 1B33-N023A 1B33-N023B 1B33-N028A 1B33-N028B Drywell The reactor recirculation pump which is hit is not required for shutdown nor would it be pierced by the missile 1B33-N021 1B33-N022 Drywell Hanger 1RT28007S - Not essential for shutdown Conduit C74523 - Will withstand missile Hanger 1RT01020S - Will withstand missile Weir wall - Will withstand missile 1G33-N004 Containment FW Guard Pipe - Will withstand missile 1G33-N015 Containment Valve 1G33-F031 and associated Flex Conduit - Not essential for shutdown. Valve itself will withstand missile 1G33-N007 Containment Pipe 1RT11A4 - Will withstand missile 1G33-N019 Containment Pipe 1RT06A6 - Will withstand missile Pipe 1RT05D4 - Will withstand missile Steam Tunnel Wall - Will withstand missile 1G33-N020 Containment Pipe 1RT05D4 - Will withstand missile Steam Tunnel Wall - Will withstand missile 1G33-N006A Containment Hanger 1RE16033G - Not essential for shutdown Steam Tunnel Wall - Will withstand missile CPS/USAR TABLE 3.5-9 (Cont'd) POTENTIAL MISSILE SOURCES WHICH COULD IMPACT SAFETY RELATED COMPONENTS CHAPTER 03 3.5-30 REV. 11, JANUARY 2005 POTENTIAL THERMOWELL MISSILE LOCATION REMARKS 1G33-N006B Containment Valve 1G33-F022E - Will withstand missile Hanger 1RE20049X - Will withstand missile Hanger 1RE20060S - Not essential for shutdown Steam Tunnel Wall - Will withstand missile 1B21-N040 Auxiliary Steam Tunnel Hanger 1RI08020R - Will withstand missile 1B21-N059 Auxiliary Steam Tunnel Pipe 1FW25AB 3/4 - Not essential for shutdown Valve 1B21-F466B - Not essential for shutdown Pipe 1RI24A2 - Will withstand missile 1B21-N060 Auxiliary Steam Tunnel Pipe 1MS35A3 - Not essential for shutdown CPS/USAR CHAPTER 03 3.6-1 REV. 12, JANUARY 2007 3.6 PROTECTION AGAINST THE DYNAMIC EFFECTS ASSOCIATED WITH THE POSTULATED RUPTURE OF PIPING Piping failures in normally operating high- and moderate-energy fluid systems are postulated. The direct results of a piping failure are pipe whip, fluid impingement, environment pressurization, temperature and humidity effects, and flooding. Coincident with the piping failure, the functional failure of any single active component, a seismic event the level of the safe shutdown earthquake and a loss of offsite power are assumed to occur. Given the above, the safety function of essential systems and components will not be impaired beyond that required to bring the plant to a safe shutdown. This section describes the design bases and protective measures which ensure that the containment, essential systems, components and equipment, and other essential structures are adequately protected from the effects associated with the postulated rupture of high-energy piping and cracks of high- and moderate-energy piping both inside and outside the containment. 3.6.1 Postulated Piping Failures in Fluid Systems The following is a summary of applicable definitions, criteria employed, potential sources and locations of piping failures, identification of systems and components essential to safe plant shutdown, limits of acceptable loss of function or damage and effect on safe shutdown, habitability of critical areas following postulated piping ruptures, and the impact of the plant design on inservice surveillance and inspection. The steam tunnel subcompartment pressure analysis inside containment is given in Subsection 6.2.1.2. The results of the steam tunnel subcompartment pressure analysis outside containment are presented in Subsection 3.6.1.2.2 and Table 3.6-3. The details of the steam tunnel pressure analysis outside containment are presented below. The steam tunnel outside containment is subject to pressure differentials when a high energy line with the tunnel is postulated to rupture. A simultaneous break of one main steamline and one feedwater line was considered in this analysis to determine the accident conditions within

the steam tunnel. The steam tunnel is a passageway starting at the primary containment boundary, which is dead-ended, and the ending in the turbine building. It is made of two sections; one section is horizontal run which starts at the containment boundary and ends in the turbine building, and the other section is a vertical section located in the turbine building (Drawings M01-1107, M01-1111 Sheet 1, and Figure 6.2-132 Sheets 3, 4, and 6.) The only exit from the steam tunnel is into the turbine building at the grade floor (elevation 737 feet 0 inches). All of the lines present in the steam tunnel are considered to be high energy lines. There are four main steamlines and two feedwater lines. The simultaneous break of one main steamline and one feedwater line was assumed to be an instantaneous guillotine rupture within the steam tunnel. The pipe breaks inside the tunnel were postulated to occur approximately 10 feet downstream of the outer main steam isolation valves since this results in the most severe pressure transient in the tunnel. General Electric data on mass and energy release for one main steamline downstream of the outer main steam isolation valves was used to determine the CPS/USAR CHAPTER 03 3.6-2 REV. 11, JANUARY 2005 break flow for the main steamline. The conditions assumed in the feedwater line were 1,077 psig, 430

°F and an enthalpy of 408.5 Btu/lbm. The feedwater line break mass and energy releases were also evaluated at a reduced feedwater temperature (RFWT) of 380

°F and an enthalpy of 354.8 Btu/lbm. It was also conservatively assumed that the feedwater flowed out of the break with sonic velocity for the entire period of the transient. The feedwater break flow was calculated using the Moody critical flow model. The feedwater flow was assumed to be limited from the reactor side by an instantaneous closure of the check valve in the line, but unimpeded from the heater side. For conservatism, it was assumed that only one check valve located in the drywell operates under accident conditions so that approximately 87 cubic feet of feedwater volume (in 62 feet of pipe of 1.407 square feet flow area) was available from the reactor side for flow out of the feedwater break. The mass and energy release rates are shown in Table 3.6-7. The nodalization of the steam tunnel analysis required the nodalization of the steam tunnel, the basement, grade floor, mezzanine floor, and the turbine floor of the turbine building. In addition, another node was required to model the atmosphere outside the turbine building. The nodalization is shown in Figure 3.6-15. It was calculated that the turbine room siding on the turbine floor of the turbine building would yield if the pressure on the turbine floor approached 15.4 psia. This would open a flow path equal to 2,000 square feet between the turbine floor and the atmosphere. The volume and vent path characteristics were determined in a similar manner as described in Subsection 6.2.1.2.3.2.1 and all of the nodal and vent path descriptions are shown in Tables 3.6-8 and 3.6-9, respectively. The initial conditions within the steam tunnel were assumed to be at 14.7 psia and a relative humidity of 100%. The initial temperatures within the steam tunnel nodes were assumed to be 150

°F, and the initial temperatures within the turbine building and atmospheric nodes were assumed to be 104°F. The simultaneous rupture of one main steamline and one feedwater line within the steam tunnel resulted in a maximum pressure of 13.8 psig, 0.25 seconds after the initiation of the accident.

The pressure history is shown in Figure 3.6-16. The first Feedwater Isolation Valve outside the containment within the steam tunnel is designed as a Class 1 valve for severe duty application with environmental effects more severe that the accident conditions derived from the simultaneous break of one main steamline and one

feedwater line.

Information on the main steamline isolation valves is given in Subsection 6.2.4.2, System Design. (Q&R 410.2) 3.6.1.1 Design Bases 3.6.1.1.1 Definitions

a. Essential Systems and Components Systems and components required to shut down the reactor and/or mitigate the consequences of a postulated piping failure, without offsite power.

CPS/USAR CHAPTER 03 3.6-3 REV. 11, JANUARY 2005 b. High-Energy Fluid Systems Fluid systems that, during normal plant conditions, are either in operation or maintained pressurized under conditions where either or both the following are

met: 1. Maximum operating temperature exceeds 200

° F. 2. Maximum operating pressure exceeds 275 psig. c. Moderate-Energy Fluid Systems Fluid systems that, during normal plant conditions, are either in operation or maintained pressurized (above atmospheric pressure) under conditions where both of the following conditions are met: 1. Maximum operating temperature is 200

° F or less, and 2. Maximum operating pressure is 275 psig or less.

NOTE: Piping which operates above the high-energy limits less than 2% of the time may be classified as moderate energy lines. d. Normal Plant Conditions Plant operating conditions normally experienced during reactor startup, operation at power, hot standby, or reactor cooldown to cold shutdown condition. e. Upset Conditions Plant operating conditions during system transients that may occur with moderate frequency during plant service life and are anticipated operational occurrences, but not during system testing. f. Postulated Piping Failures Longitudinal and circumferential breaks in high energy fluid system piping and through-wall leakage cracks in high and moderate-energy fluid system piping postulated according to the provisions of Branch Technical Position (BTP) MEB 3-1, attached to Standard Review Plan (SRP) 3.6.2. g. S h and S a Allowable stresses at maximum (hot) temperature and allowable stress range for thermal expansion, respectively, as defined in Article NC-3600 of the ASME

Code, Section III. h. S m Design stress intensity as defined in Article NB-3600 of the ASME Code, Section III.

CPS/USAR CHAPTER 03 3.6-4 REV. 11, JANUARY 2005 i. Single Active Component Failure Malfunction or loss of function of a component of electrical or fluid systems. The failure of an active component of a fluid system is considered to be a loss of component function as a result of mechanical, hydraulic, pneumatic, or electrical malfunction, but not the loss of component structural integrity. The direct consequences of a single active component failure are considered to be a part of the single failure. j. Terminal Ends Extremities of piping runs that connect to structures, large components (e.g., vessels, pumps) or pipe anchors that act as rigid constraints to piping movement, including rotational movement, from static or dynamic loading. A branch connection to a main piping run is a terminal end of the branch run, except for cases where the piping model includes both the run and branch piping. Intersections of runs of comparable size and stability are not considered terminal ends when the piping stress analysis model includes both the run and the branch piping, and the intersection is not rigidly constrained to the building structure. k. Leakage Crack A postulated opening in the piping system, the consequences of which are evaluated on the basis of pressure and temperature differential conditions and flooding effects. l. Fluid Systems High- and moderate-energy fluid systems that are subject to the postulation of piping failures against which protection of essential systems and components is needed. 3.6.1.1.2 Criteria The pipe failure protection conforms to Appendix A of 10 CFR 50, General Design Criterion 4, Environmental and Missile Design Bases. The overall design for this protection is in compliance with NRC BTP APCSB 3-1 (attached to SRP 3.6.1), MEB 3-1 (attached to SRP 3.6.2) and NRC NUREG-1061, Volume 3 (Reference 12), the implementation of which is discussed herein. 3.6.1.1.3 Objectives Protection against pipe failure effects was provided to fulfill the following objectives: a. Assure that the reactor can be shut down safely and maintained in a safe shutdown condition or mitigate the consequences of a loss-of-coolant accident (LOCA). b. Assure that containment integrity is maintained.

CPS/USAR CHAPTER 03 3.6-5 REV. 12, JANUARY 2007 c. Assure that a pipe break does not directly or indirectly cause a loss of reactor coolant beyond makeup capability. d. Assure that the radiological doses of a postulated piping failure will remain below the limits of 10 CFR 50.67. 3.6.1.1.4 Assumptions The following assumptions were used to determine the protection requirements: a. Pipe breaks or cracks were postulated to occur during normal plant operation (i.e., reactor startup, operation at power, hot standby or reactor cooldown to a cold shutdown). b. Only high-energy piping as defined in Subsection 3.6.1.1.1.b and shown on Figure 3.6-1 was considered in the determination of the potential pipe break location. Moderate-energy piping as defined in Sub-section 3.6.1.1.1.c and shown on Figure 3.6-1 was capable of producing only cracks. c. Pipe breaks were evaluated for the effects of pipe whip, jet impingement, flooding, room pressurization, and other environmental effects such as

temperature. d. Pipe cracks were evaluated for flooding and effects from spray. e. Each longitudinal or circumferential break in high-energy fluid system piping, or leakage crack in moderate-energy fluid system piping, was considered separately as a single postulated initial event occurring during normal plant conditions. (Applicable to seismic and nonseismic piping). f. Pipe failures (breaks or cracks) inside the containment were not evaluated concurrently with failures outside the containment. g. Offsite power was assumed to be unavailable if a trip of the turbine-generator system or reactor protection system was a direct consequence of the postulated piping failure, unless it was more conservative to assume that offsite power was available (e.g., a feedwater line break with offsite power available leads to a larger inventory of water for flooding considerations). h. A single active component failure was assumed in systems used to mitigate consequences of the postulated piping failure and to safely shut down the reactor, except as noted in paragraph i below. The single active component failure was assumed to occur in addition to the postulated piping failure, such as unit trip and loss of offsite power. i. Where the postulated piping failure was assumed to occur in one of two or more redundant trains of a dual-purpose, moderate-energy essential system (i.e., one required to operate during normal plant conditions as well as to shut down the reactor and mitigate the consequences of the piping failure), single failures of components in the other train or trains of that system were not assumed provided the system was designed to the following criteria: (1) Seismic Category I CPS/USAR CHAPTER 03 3.6-6 REV. 11, JANUARY 2005 standards: (2) has both offsite and onsite power sources; and (3) was constructed, operated, and inspected to quality assurance, testing, and inservice inspection standards appropriate for nuclear safety systems. Examples of systems that qualify as moderate-energy dual-purpose essential systems are the shutdown service water system and the residual heat removal system. j. All available systems, including those actuated by operator actions, were employed to mitigate the consequences of a postulated piping failure to the

extent clarified in the following paragraphs: 1. In judging the availability of systems, account was taken of the postulated failure and its direct consequences such as unit trip and loss of offsite power, and of the assumed single active component failure and its direct consequences. The feasibility of carrying out operator actions was judged on the basis of ample time and adequate access to equipment being available for the proposed actions. 2. Only Seismic Category I equipment could be used to mitigate the consequences of the failure and bring the plant to a safe shutdown. k. A whipping pipe was not considered capable of rupturing impacted pipes of equal or greater nominal pipe diameter and equal or greater thickness. l. Pipe movement was assumed to occur in the direction of the jet reaction.

m. Absorption of the fluid internal energy associated with the pipe break reaction could take into account any line restrictions (e.g., flow limiter) between the pressure source and break location and absence of energy reservoirs, as applicable. n. Initial pipe break events are not assumed to occur in pump and valve bodies because of their greater wall thicknesses and their locations in the low stress portions of the piping systems. o. Piping which is physically separated (or isolated) from structures, systems, or components important to safety by protective barriers, such as concrete incasements, will not initiate another pipe break event beyond the separated area. 3.6.1.1.5 Identification of Systems Important to Plant Safety For a given postulated piping failure, the systems which may be required to shut down the plant and maintain it in a safe condition are identified in Table 3.6-1. Figures FP-8 thru FP-20 of Clinton Power Station Fire Protection Evaluation Report dated April 12, 1978, show the location on general arrangement drawings of all safety-related equipment including electrical cable trays. Typical piping runs with postulated failure points indicated and the design approach to protect essential components are illustrated in Subsection 3.6.2. High-energy fluid systems and divisional separation are shown in Figure 3.6-1 for each system important to plant safety.

CPS/USAR CHAPTER 03 3.6-7 REV. 11, JANUARY 2005 Those portions of non-safety-related systems (i.e., service air, service water) which penetrate the primary containment were evaluated individually for divisional separation and postulated piping failure effects as described in Subsections 3.6.1.1.3 and 3.6.1.1.4.

3.6.1.2 Description 3.6.1.2.1 Potential Sources for Piping Failure A list of systems considered high-energy per Subsection 3.6.1.1.1.b is given in Table 3.6-2. The boundaries for these high-energy, as well as for moderate-energy lines, are shown in Figure 3.6-1. The piping and instrumentation diagrams (P&ID's) in Figure 3.6-1 should not be used for detailed information, i.e. vents, drains. Detail information should be taken from P&ID's for each applicable system located throughout this USAR. Locations, orientations, and size of piping failures within high-moderate-energy piping systems are postulated per the criteria given in Subsection 3.6.2. The effects of pipe whip, jet impingement, spraying and flooding on essential systems, components, and equipment are discussed in Attachment D3.6, Failure Mode Analysis. There are no credible secondary missiles formed from the postulated rupture of piping.

3.6.1.2.2 Structures and Compartments Used to Protect Against Piping Failure Pressure response analyses were performed fo r the subcompartments containing high-energy piping. For a detailed discussion of the line breaks selected, vent paths, room volumes, analytical methods, pressure results, etc., refer to Subsection 6.2.1.2 for containment subcompartments and Table 3.6-3 for subcompartments located outside the containment. Several subcompartments analyzed are also used for divisional separation and therefore contain only safety-related systems and components of one safety division. Subcompartments used as such were analyzed to determine the environmental and pressurization effects of pipe failure; but, in these cases, only system failure, not component failure, was analyzed in the Failure Mode Analysis (Attachment D3.6). Those subcompartments used for divisional separation are listed in Table 3.6-4 for the auxiliary building and the fuel building. (The circulating water screen house has no high-energy piping systems and therefore no subcompartment evaluations are included.) 3.6.1.2.3 Pipe Failure Effects on Control Room There is no high-energy piping capable of producing impact or jet impingement effects in or near the control room. There are no effects upon the habitability of the control room by pipe break either from pipe whip, jet impingement, or transport of steam. Further discussion on control room habitability systems is provided in Section 6.4. 3.6.1.2.4 Impact of Plant Design for Postulated Piping Failures on Inservice Inspection Access has been provided for inservice inspection as dictated by the ASME Boiler and Pressure Vessel Code, Section XI, "Inservice Inspection of Nuclear Power Plant Components."

CPS/USAR CHAPTER 03 3.6-8 REV. 11, JANUARY 2005 3.6.1.3 Safety Evaluation 3.6.1.3.1 General In the plant design, consideration was given to the effects of postulated piping breaks with respect to the limits of acceptable damage/loss of function, to assure that, even with coincident single loss of active component, an earthquake equal to the safe shutdown earthquake, and loss of offsite power, the remaining structures, systems, and components would be adequate to safely shut down the plant. This subsection summarizes the Structural, Mechanical, Instrumentation, Electrical, and HVAC items that are safety-related and therefore designed to remain functional for the following: (1) a high-energy line rupture with resulting whip, impingement, compartment pressurization and temperature rise, wetting of compartment surfaces, and flooding, or (2) a moderate-energy break with resulting impingement, wetting of compartment surfaces, and flooding. By means of the design features such as separation, barriers, and pipe whip restraints, all of which are discussed below, it has been assured that the design function or necessary component operability of essential items will not be impaired by the effects of postulated breaks and cracks. Specific design features used for protecting the essential systems, components, and equipment are summarized in Attachment D3.6. The ability of specific safety-related systems to withstand a single active failure concurrent with a postulated event is discussed as applicable. If a review of the pipe layout and plant arrangement drawings showed that the effects of the postulated breaks/cracks, on a reasonable basis, are isolated, physically remote, or restrained by plant design features from essential systems or components, no further evaluation was performed. 3.6.1.3.2 Protection Methods

a. General The effects associated with a particular break/crack must be mechanistically consistent with the failure. Thus, actual pipe dimensions, piping layouts, material properties, and equipment arrangements were considered in defining the specific measures for protection against actual pipe movement and other associated consequences of postulated failures. Protection against the dynamic effects of pipe failures was provided in the form of pipe whip restraints, equipment shields, and physical separation of piping, equipment, and instrumentation. The precise method chosen depended largely upon limitations placed on the designer such as accessibility, maintenance, and proximity to other pipes and equipment. b. Separation The plant arrangement provides distance separation to the extent practicable between redundant safety systems (including their auxiliaries) in order to prevent CPS/USAR CHAPTER 03 3.6-9 REV. 11, JANUARY 2005 loss of safety function as a result of the dynamic effects of pipe break or crack. Separation is the basic protective measure incorporated in the design to protect

against dynamic effects. c. Barriers, Shields, and Enclosures Protection requirements were met with walls, floors, columns, and foundations in many cases. d. Piping Restraints Measures for protection against pipe whipping as a result of high-energy pipe breaks were not provided where any one of the following applied: 1. The piping was physically separated (or isolated) from all essential safety-related structure, systems, or components required to place the plant in a safe shutdown condition following the postulated rupture, or was restrained from whipping by plant design features such as concrete

encasement. 2. Following a single break, the unrestrained movement of either end of the ruptured pipe could not damage, to an unacceptable level, any structure, system, or component required to place the plant in a safe shutdown condition following the postulated rupture. 3. The energy associated with the whipping pipe was demonstrated to be insufficient to impair, to an unacceptable level, the safety function of any structure, system, or component required to place the plant in a safe shutdown condition following the postulated rupture. The design criteria for restraints are given in Subsection 3.6.2. 3.6.1.3.3 Specific Protection Measures

a. The general layout of the facility followed a multi-step process to ensure adequate separation. 1. Safety-related systems were located away from most high-energy piping. 2. Redundant (e.g., "A" and "B" trains) safety subsystems were located in separate compartments. 3. As necessary, specific components were enclosed to retain the redundancy required for those systems that must function as a consequence of specific piping failure. 4. Drainage systems were reviewed to assure their adequacy for flooding prevention.

CPS/USAR CHAPTER 03 3.6-10 REV. 11, JANUARY 2005 b. For high-energy piping systems which penetrate the containment and drywell, isolation valves are located as close to the containment as possible to facilitate design of moment guides and maintain isolation valve operability. c. The pressure, water level, and flow sensor instrumentation for those essential systems which are required to function following a pipe rupture are protected. d. High-energy fluid system piping restraints and protective measures were designed such that a postulated break in one pipe could not, in turn, lead to a rupture of other nearby pipes or components if the secondary rupture could result in consequences considered unacceptable for the initial postulated break. e. Deleted f. For any postulated pipe rupture, the structural integrity of the containment was assured. In addition, for those postulated ruptures classified as a loss of reactor coolant, the design leak tightness of the containment fission product barrier was

assured. g. Safety/relief valves and the RCIC steamline were located and restrained so that a pipe failure would not prevent depressurization. h. Separation was provided to preserve the independence of the low-pressure core spray (LPCS) and low-pressure coolant injection (LPCI) portions of the RHR systems. i. High-energy piping which penetrated both the drywell and the containment was provided with guard pipes in accordance with Subsection 3.6.2.4. The encapsulated piping was designed in accordance with the criteria of Subsection 3.6.2.5. 3.6.2 Determination of Break Locations and Dynamic Effects Associated with the Postulated Rupture of Piping Described herein are the design bases for postulating piping breaks and cracks inside and outside of containment, the procedures used to define the jet thrust reaction at the break location, the jet impingement loading criteria, and the piping dynamic response models. 3.6.2.1 Criteria Used to Define Break and Crack Location and Configuration 3.6.2.1.1 Definition of High-Energy Fluid System The definition of a high-energy fluid system is found in Subsection 3.6.1.1.1b.

3.6.2.1.2 Definition of Moderate-Energy Fluid System The definition of a moderate-energy fluid system is found in Subsection 3.6.1.1.1c.

3.6.2.1.3 Postulated Pipe Breaks and Cracks A postulated pipe break is defined as a sudden, gross failure of the pressure boundary either in the form of a complete circumferential severance (guillotine break) or as development of a CPS/USAR CHAPTER 03 3.6-11 REV. 11, JANUARY 2005 sudden longitudinal split and is postulated for high-energy fluid systems only. For moderate-energy fluid systems, pipe failures are confined to postulation of controlled cracks in piping and branch runs. These cracks affect the surrounding environmental conditions only and do not result in whipping of the cracked pipe. The following high-energy piping systems (or portions of systems) have been considered in the determination of a postulated pipe break during normal plant conditions and are evaluated for potential damage resulting from dynamic effects. a. All piping which is part of the reactor coolant pressure boundary and subject to reactor pressure continuously during station operation. b. All piping which is beyond the second isolation valve but which is subject to reactor pressure continuously during station operation. c. All other piping systems or portions of piping systems considered high-energy systems. A high-energy piping system break is not postulated as simultaneous with a moderate-energy piping system crack nor is any pipe break or crack outside containment postulated concurrently with a postulated pipe break inside containment. A list of high-energy fluid systems is provided in Table 3.6-2. 3.6.2.1.4 Exemptions from Pipe Break Evaluation and Protection Requirements The following are exemptions for postulating pipe breaks: a. Piping is classified as moderate-energy piping.

b. The nominal pipe size is 1 inch or less.

c. The operation period is short. (An operation period is considered "short" if the fraction of time that the system operates within the pressure-temperature conditions specified for high-energy fluid systems is about 2 percent or less of the time that the system operates as a moderate-energy fluid system: e.g., systems such as the reactor decay heat removal system). Piping in this category is classified as moderate energy. d. Portions of high-energy piping systems that are isolated from the source of the high-energy fluid during normal plant conditions are exempted from consideration of postulated pipe breaks. This would include portions of piping systems beyond a normally closed valve. This type of piping is classified as moderate-energy

fluid system. e. Pump and valve bodies are exempted from consideration of pipe break because of their greater wall thickness. f. CRD insert lines are exempted per Reference 7.

CPS/USAR CHAPTER 03 3.6-12 REV. 11, JANUARY 2005 Protection from pipe whip dynamic effects associated with pipe break is not provided if, following a single postulated pipe break, piping for which the unrestrained movement of either end of the ruptured pipe in any feasible direction about a plastic hinge, formed within the piping, cannot cause a loss of function of any structur e, system, or component important to safety. 3.6.2.1.5 Types of Breaks and Leakage Cracks in Fluid System Piping Except as noted in Subsection 3.6.2.1.4, the following types of breaks are postulated in high-energy fluid system piping: a. Circumferential breaks are postulated only in piping exceeding a 1-inch nominal pipe diameter. Where break locations are selected in piping without the benefit of stress calculations, breaks are postulated concurrently at the piping welds to each fitting, valve, or welded attachment (not applicable to recirculation piping). Circumferential breaks are assumed to result in pipe severance separation amounting to at least a one-diameter lateral displacement of the ruptured piping sections. Cases for which the resulting pipe separation is less than one pipe diameter because of physical limitation by piping restraints, structural members, or piping stiffness will be identified. b. Longitudinal splits are postulated only in piping having a nominal diameter equal to or greater than 4 inches. Longitudinal breaks are assumed to result in an axial split without pipe severance. Splits are oriented (but not concurrently) at two diametrically opposed points on the piping circumference such that the jet reaction causes out-of-plane bending of the piping configuration. Alternatively, a single split is assumed at the section of highest tensile stress as determined by detailed stress

analysis. c. Circumferential breaks are assumed at all terminal ends and at intermediate locations identified by the criteria in Subsections 3.6.2.1.6.1 and 3.6.2.1.6.2.1.1.

At each of the intermediate postulated break locations identified to exceed the stress and usage factor limits of the criteria in Subsections 3.6.2.1.6.1 and 3.6.2.1.6.2.1.1, either a circumferential or a longitudinal break, or both, are postulated per the following: 1. Circumferential breaks are postulated at fitting joints.

2. Longitudinal breaks are postulated in the center of the fitting at two diametrically opposed points (but not concurrently) located so that the reaction force is perpendicular to the plane of the piping and produces out-of-plane bending. 3. Consideration is given to the occurrence of either a longitudinal or circumferential break. The state of stress in the vicinity of the postulated break location may be used to identify the most probable type of break. If the maximum stress range in the longitudinal direction is greater than CPS/USAR CHAPTER 03 3.6-13 REV. 11, JANUARY 2005 1.5 times the maximum stress range in the circumferential direction, only the circumferential break is postulated, and, consequently, if maximum stress range in the circumferential direction is greater than 1.5 the stress range in the longitudinal direction, only the longitudinal break is postulated. If there are no significant differences between the circumferential and longitudinal stresses, then both types of breaks are considered. 4. At intermediate locations chosen to satisfy the minimum break location criteria, only circumferential breaks are postulated (not applicable to recirculation piping). d. For design purposes, a longitudinal break area is assumed to be the equivalent of one circumferential pipe area for all longitudinal breaks postulated in piping. e. For both longitudinal and circumferential breaks, after assessing the contribution of upstream piping flexibilities, pipe wh ipping is assumed to occur in the plane defined by the piping geometry and configuration for circumferential breaks and out-of-plane for longitudinal breaks and to cause pipe movement in the direction of the jet reaction. f. For a circumferential break, the dynamic force of the jet discharge at the break location is based upon the effective cross-sectional flow area of the pipe and on a calculated fluid pressure as modified by an analytically or experimentally determined thrust coefficient. Thrust coefficients have been determined in accordance with ANS-58.2 (ANSI N176). Justifiable line restrictions, flow limiters, and the absence of energy reservoirs are used, as applicable, in the reduction of the jet discharge. The following through-wall leakage cracks are postulated in high and moderate-energy fluid system piping at the locations specified in this position: a. Cracks are postulated in moderate-ener gy fluid system piping and branch runs exceeding a nominal pipe size of 1 inch. b. Fluid flow from a crack is based on a circular opening of area equal to that of a rectangle one-half pipe-diameter in length and one-half pipe wall thickness in

width. c. The flow from the crack is assumed to result in an environment that wets all unprotected components within the compartment, with consequent flooding in the compartment and communicating compartments. Flooding effects are determined on the basis of a conservatively estimated time period required to effect corrective actions. Evaluation of jet impingement effects is not considered for postulated through-wall leakage cracks. d. Through-wall leakage cracks instead of breaks are postulated in the piping of those fluid systems that qualify as high-energy fluid systems for only short operational periods as defined in Subsection 3.6.2.1.4.

CPS/USAR CHAPTER 03 3.6-14 REV. 11, JANUARY 2005 3.6.2.1.6 Location for Postulated Pipe Breaks and Leakage Cracks 3.6.2.1.6.1 Criteria for Reactor Recirculation Piping System Inside Containment - Within the Scope of the NSSS Supplier Postulate pipe break locations are selected in accordance with the U.S. Nuclear Regulatory Commission (NRC) Branch Technical Position APCSB 3-1, Appendix B NRC Branch Technical Position MEB 3-1, and NUREG-1061, Volume 3. For the ASME Section III, Class 1 recirculation piping system which is classified as high-energy, the postulated break locations are

as follows: a. At the terminal ends of the pressurized portions of the run.

b. Where the maximum stress range between any two load sets (including zero load set) according to Subarticle NB-3600, ASME Code Section III for service level B (upset plant conditions), and an independent OBE event transient, exceeds the following: 1. The maximum stress range between any two loads sets (including the zero load set) should not exceed 2.4 S m and should be calculated by Eq. (10) in NB-3653, ASME Code Section III. If the calculated maximum stress range of Eq. (10) exceeds 2.4 S m , the stress ranges calculated by both Eq. (12) and Eq. (13) in Paragraph NB-3653 should meet the limit of 2.4 S

m. 2. The cumulative usage factor should be less than 0.1.

3. The maximum stress, as calculated by Eq. (9) in NB-3652 under the loadings resulting from a postulated piping failure beyond these portions of piping should not exceed the lesser of 2.25 S m and 1.8 S y except that following a failure outside containment, the pipe between the outboard isolation valve and the first restraint may be permitted higher stresses provided a plastic hinge is not formed and operability of the valves with such stresses is assured in accordance with the requirements specified in SRP Section 3.9.3. Primary loads include those which are deflection limited by whip restraints. Intermediate breaks are no longer postulated where the calculated cumulative usage factors and stress intensity ranges are lower than the limits specified in subparagraph b above (see

Reference 11). There are no Class 1 moderate energy piping systems.

There is no piping component in this subsection other than ASME Class 1. 3.6.2.1.6.2 Piping Other Than Reactor Recirculation Piping in Subsection 3.6.2.1.6.1 This subsection applies to all high- and moderate-energy piping inside and outside containment with the exception of the reactor recirculation piping in Subsection 3.6.2.1.6.1.

CPS/USAR CHAPTER 03 3.6-15 REV. 11, JANUARY 2005 3.6.2.1.6.2.1 High-Energy Fluid System Piping 3.6.2.1.6.2.1.1 Fluid System Piping Not in the Containment Penetration Area

a. Breaks in ASME, Section III, Class 1 piping are postulated at the following locations in each piping run or branch run: 1. At terminal ends of the run; 2. At locations where the primary plus secondary stress intensity range between any two load sets (including the zero load set) as calculated by Equation (10) and either Equation (12) or (13) in paragraph NB-3653 of ASME, Section III exceeds 2.4 S m for loadings resulting from normal and upset plant conditions, including SRV discharge and suppression pool vibratory loads and an OBE event; and 3. At any intermediate locations between terminal ends where the cumulative usage factor derived from the piping fatigue analysis under the loadings resulting from plant normal, upset, and testing conditions, SRV discharge and suppression pool vibratory loads, and an OBE event exceeds 0.1. 4. In the event that two intermediate locations cannot be determined by the stress or usage factor limits just described, the two locations of highest stress, as calculated by equation (10) in paragraph NB-3653 of ASME Section III, which are separated by a change in direction of the pipe run, are selected. If the piping run has only one change or no change of direction, only one intermediate break is postulated. A given elbow or other fitting (tee, reducer, etc.) is considered as a single-break location regardless of the number of types of breaks postulated at the fitting. 5. Breaks postulated using the criteria in item 4 are generally referred to as "arbitrary intermediate breaks" (AIBs). Initially AIBs were postulated for the following Class I piping (see Attachment B3.6 Figures): MS0l HP0l RH05 NB01 MS02 LP01 RH34 MS05 MS03 RH01 RI01 RR32 MS04 RH03 SC07 RR33 However, recently, AIBs have been eliminated for this piping, since it is not susceptible to stress corrosion cracking, steam or water hammer effects, or thermal fatigue in fluid mixing situations (See Reference 11). b. Breaks in ASME Section III Class 2 and 3 piping and seismically qualified ANSI B31-1 piping are postulated at the following locations-in each piping run or branch run:

CPS/USAR CHAPTER 03 3.6-16 REV. 11, JANUARY 2005 1. At terminal ends of the run, and 2. At each intermediate location where the stresses under the loadings resulting from normal and upset plant conditions, SRV discharge and suppression pool vibratory loads (if applicable), and an OBE event, as calculated by the sum of equations (9) and (10) in paragraph NC-3652 of ASME Section III, exceed 0.8 (1.2 S h + S a). 3. In the event that two intermediate locations cannot be determined by the stress limits described above, the two locations of highest stress as calculated by the sum of equations (9) and (10) in paragraph NC-3652 of ASME Section III which are separated by a change in direction of the pipe run are selected. If the piping run has only one change or no change of direction, only one intermediate break is postulated. A given elbow or other fitting (tee, reducer, etc.) is considered as a single-break location regardless of the number of types of breaks postulated at the fitting. 4. Breaks postulated using the criteria in item 3 are referred to as "arbitrary intermediate breaks" (AIBs). Initially AIBs were postulated for the following Class 2 and 3 piping (see Attachment B3.6 Figures):

RT02 RT07 RH14

RT05 RT08 RH07 IS03 RI02 RH08 RT06 However, recently AIBs have been eliminated for this piping since it is not susceptible to stress corrosion cracking, steam or water hammer effects, or thermal fatigue in fluid mixing situations (see Reference 11). 5. As an alternative to the foregoing, intermediate locations are assumed at each location of potential high stress or fatigue such as pipe fittings, valves, flanges, and welded attachments. c. Longitudinal breaks are not postulated at terminal ends. d. Leakage cracks in high-energy ASME Section III Class 1 piping are postulated at locations where the primary and secondary stress intensity range, as calculated by equation (10) or equations (12) and (13) in paragraph NB-3653 of ASME Section III, exceed 1.2 Sm for loadings resulting from normal and upset plant conditions. e. Leakage cracks in high-energy ASME Section III Class 2 and 3 piping and seismically qualified ANSI B31.1 piping are postulated at locations where the stresses under the loadings resulting from normal and upset plant conditions and an OBE event, as calculated by the sum of equations (9) and (10) in paragraph NC-3652 of ASME Section III, exceed 0.4 (1.2S h + S a).

CPS/USAR CHAPTER 03 3.6-17 REV. 11, JANUARY 2005 f. Breaks and cracks in nonseismically qualified piping are postulated at locations to produce the controlling design basis events. 3.6.2.1.6.2.1.2 Fluid System Piping in Containment Penetration Areas This subsection applies to the fluid system piping between the containment isolation valves, including the valves and the valve-to-process piping welds and any connection to the containment penetration. Breaks are not postulated in the containment penetration area as defined above where the following design requirements are met. a. The following design stress and fatigue limits are not exceeded for ASME Code Section III Class 1 piping: 1. The primary plus secondary stress intensity range between any two load sets (including the zero load set) for normal and upset conditions, SRV discharge and suppression pool vibratory loads, and an OBE event, as calculated by equation (10) or equations (12) and (13), does not exceed

2.4 S m. 2. The cumulative usage factor U derived from the piping fatigue analysis under the loadings associated with normal, upset, and testing conditions, SRV discharge and suppression pool vibratory loads, and an OBE event is less than 0.1. 3. The maximum stress, as calculated by equation (9) in paragraph NB-3652 under the loadings resulting from internal pressure, dead weight, and a postulated piping failure of fluid systems beyond these portions, does not exceed 2.25 S

m. b. The following design stress and fatigue limits are not exceeded for ASME Code Section III Class 2 piping: 1. The maximum stress range, as calculated by the sum of equations (9) and (10) in paragraph NC-3652, ASME Code Section III, under the loadings resulting from the normal and upset plant conditions (i.e., sustained loads, occasional loads including SRV discharge and suppression pool vibratory loads, and thermal expansion) and OBE event, does not exceed 0.8 (1.2 S h +S a). 2. The maximum stress, as calculated by equation (9) in paragraph NC-3652 under the loadings resulting from internal pressure, dead weight, and a postulated piping failure of fluid system piping beyond these portions of piping does not exceed 1.8 S

h. c. Following a piping failure outside the first pipe whip restraint, the formation of a plastic hinge is not permitted in the piping between the containment penetration and the first pipe whip restraint. Bending and torsion limiting restraints are installed, as necessary, at locations selected to optimize overall piping design, to prevent formation of a plastic hinge as noted, to protect against the impairment of CPS/USAR CHAPTER 03 3.6-18 REV. 12, JANUARY 2007 the leaktight integrity of the containment, to assure isolation valve operability, and to meet the stress and fatigue limits in the containment penetration area. d. Leakage cracks in the containment penetration area are postulated in accordance with Subsection 3.6.2.1.6.2.1.1. e. The number of circumferential and longitudinal piping welds and branch connections is minimized as much as practical. f. The length of these portions of piping is reduced to the minimum length practical.

g. An augmented ISI will be performed as discussed in Subsections 5.2.4.12 and 6.6.8. The break exclusion areas are typically shown on the B3.6 Figures.

3.6.2.1.6.2.1.3 Details of the Containment Penetration Details of the containment penetrations are discussed in Subsections 3.8.1 and 3.8.2.

3.6.2.1.6.2.2 Moderate-Energy Fluid System Piping Inside and Outside Containment Leakage cracks in moderate-energy piping are postulated individually at locations that would result in the maximum effects from fluid spraying and flooding, with the consequent hazards or environmental conditions developed from the spray or flood. The consequences of these postulated breaks were analyzed for each room in the safety-related buildings. The results of

these analyses are presented in USAR Section D.3.6.3. 3.6.2.1.7 Definitions Throughout this section, applicable definitions are located in Subsection 3.6.1.1.1.

3.6.2.2 Analytical Methods to Define Forcing Functions and Response Models 3.6.2.2.1 Reactor Recirculation Loop Piping - Inside Containment 3.6.2.2.1.1 Analytical Methods to Define Blowdown Forcing Functions The criteria that are used for calculation of fluid blowdown forcing functions include the following: a. Circumferential breaks are assumed to result in pipe severance and separation amounting to at least a one diameter lateral displacement of the ruptured piping sections. The details of the inelastic pipe whip analysis are provided in CPS-USAR Subsection 3.6.2.2.1.2, which describe methodol-gy of the GE computer program PDA used for this analysis. b. The dynamic force of the jet discharge at the break location is based on the effective cross-sectional flow area of the pipe and on a calculated fluid pressure as modified by an analytically or experim entally determined thrust coefficient. Limited pipe displacement at the break location, line restrictions, flow limiters, CPS/USAR CHAPTER 03 3.6-19 REV. 11, JANUARY 2005 positive pump-controlled flow, and the absence of energy reservoirs are taken into account, as applicable, in the reduction of jet discharge. ANS-58.2 (Reference 13) is the basic document that is used for determining the thrust coefficients in evaluating the dynamic force due to jet discharge. General Electric is using a value of 1.26 for a main steam line break and 2.0 for a recirculation line break. These values are conservative upper bound based on the following theoretical and experimental methods. (1) Saturated and Superheated Steam (Main Steam Line) For calculating the thrust force, saturated or superheated steam is treated as an ideal gas with a ratio of specific heat equal to 1.3. Considering the flow to be isentropic, the thrust coefficient, per Reference 14, for frictionless flow is given by:

o a T P/P - 1.26 C= where: P a = ambient pressure around pipe P o = pressure in the pipe C T = thrust coefficient For the main steam high energy pipe, since P a<<P o , C T 1.26 (2) Sub-Cooled Water (Recirculation Line) As the degree of subcooling increases, the thrust coefficient for frictionless subcooled water increases from 1.26 in steam condition to a maximum of 2.0 for non-flashing water. Normalization of the steady state thrust coefficient for frictionless flow of subcooled water based on the Henry-Fauske model (Reference 15) results in the following expression for C T: 0.75 *h 0 ;0.861h - 2.0 C*2 T= 1.0 *h 0.75 ;0.97h 3.0h - 3.22 C*2*T<+= where: h* = (h o - 180) / (hsaturated - 180) h o = stagnation enthalpy (Btu/lbm)

Hsaturated = saturated water enthalpy at the stagnation pressure (Btu/lbm).

CPS/USAR CHAPTER 03 3.6-20 REV. 11, JANUARY 2005 The experimental comparisons made by Hanson (Reference 16) showed good agreement with the theoretical prediction of thrust coefficient. In evaluating the dynamic force due to jet discharge, a conservative thrust coefficient of 2.0 is used for the recirculation piping. (MEB (DSER) 14) c. Breaks are postulated to occur instantaneously (MEB (DSER) 14). Blowdown forcing functions are determined by either of the two following methods: a. The predicted blowdown forces on pipes fed by a pressure vessel are described by transient and steady-state forcing functions. The forcing functions used are based on methods described in Reference 3. These are simply described as

follows: 1. The transient forcing functions at points along the pipe, result from the propagation of waves (wave thrust) along the pipe, and from the reaction force due to the momentum of the fluid leaving the end of the pipe (blowdown thrust). 2. The waves cause various sections of the pipe to be loaded with time-dependent forces. It is assumed that the pipe is one-dimensional, in that there is no attenuation or reflection of the pressure waves at bends, elbows, and the like. Following the rupture, a decompression wave is assumed to travel from the break at a speed equal to the local speed of sound within the fluid. Wave reflections will occur at the break end, and the pressure vessel until a steady flow condition is established. Vessel and free space conditions are used as boundary conditions. The blowdown thrust causes a reaction force perpendicular to the pipe break. 3. The initial blowdown force on the pipe is taken as the sum of the wave and blowdown thrusts and is equal to the vessel pressure (P o) times the break area (A). After the initial decompression period (i.e., the time it takes for a wave to reach the first change in direction), the force is assumed to drop off to the value of the blowdown thrust (i.e., 0.7 P o A). 4. Time histories of transient pressure, flow rate, and other thermodynamic properties of the fluid are used to calculate the blowdown force on the pipe using the following equation:

()A g P P F c 2 a+= where: F = Blowdown Force

P = Pressure at exit plane

P a = Ambient pressure CPS/USAR CHAPTER 03 3.6-21 REV. 11, JANUARY 2005 u = Velocity at exit plane = Density at exit plane A = Area of break g c = Gravitational constant 5. Following the transient period a steady-state period is assumed to exist. Steady-state blowdown forces are calculated including frictional effects.

For saturated steam, these effects reduce the blowdown forces from the theoretical maximum of 1.26 PoA. The method of accounting for these effects is presented in Reference 3. For subcooled water, a reduction from the theoretical maximum of 2.0 PoA is found through the use of Bernoulli's and standard equations such as Darcy's equation, which account for friction. b. The following is an alternate method for calculating blowdown forcing functions: The computer code RELAP3 (Reference 4) is used to obtain exit plane thermodynamic states for postulated ruptures. Specifically, RELAP3 supplies exit pressure, specific volume and mass rate. From these data the blowdown reaction load is calculated using the following relation:

AE x AE T R g v G P P AE T c E 2 E E=+= where: T/AE = thrust per unit break area - lbf/ft 2 , P E = exit pressure - lbf/ft 2 , P = receiver pressure - lbf/ft 2 , G E = exit mass flux - lb/sec ft 2 , E v = exit specific volume - ft 3/lbm, g c = gravitational constant - 32.174 (ft-lb) / and (lbf - sec 2), R = reaction force on the pipe - lbf.

CPS/USAR CHAPTER 03 3.6-22 REV. 11, JANUARY 2005 3.6.2.2.1.2 Pipe Whip Dynamic Response Analyses The criteria used for performing the pipe whip dynamic response analyses include the following: a. A pipe whip analysis is performed for each postulated pipe break. However, a given analysis is used for more than one postulated break location if the blowdown forcing function, piping and restraint system geometry and piping and restraint system properties are conservative for other break locations. b. The analysis includes the dynamic response of the pipe in question and the pipe whip restraints which transmit loading to the structures. c. The analytical model adequately represents the mass/inertia and stiffness properties of the system. d. Pipe whipping is assumed to occur in the plane defined by the piping geometry and configuration, and to cause pipe movement in the direction of the jet

reaction. e. Piping within the broken loop is no longer considered part of the RCPB. Plastic deformation in the pipe is considered as a potential energy adsorber. The maximum strain in the pipe is limited to 25% of the ultimate uniform strain of the pipe material. This limit is the same as that imposed on the energy absorbing, plastically deforming pipe whip restraints. Piping systems are designed so that plastic instability does not occur in the pipe at the design dynamic and static loads unless damage studies are performed which show that direct damage to

any essential system or component does not result. f. Components such as vessel safe ends and valves which are attached to the broken piping system and do not serve a safety function or whose failure would not further escalate the consequences of the accident, are not designed to meet ASME Code imposed limits for essential components under faulted loading.

However, if these components are required for safe shutdown, or serve a safety function to protect the structural integrity of an essential component, limits to meet the Code requirements for faulted conditions and limits to ensure operability if required are met. The pipe whip analysis was performed using the PDA computer program (Reference 5). PDA is a computer program used to determine the response of a pipe subjected to the thrust force occurring after a pipe break. The program treats the situation in terms of generic pipe break configuration, which involves a straight, uniform pipe fixed at one end and subjected to a time dependent thrust-force at the other end. A typical restraint used to reduce the resulting deformation is also included at a location between the two ends. Nonlinear and time-independent stress-strain relations are used for the pipe and the restraint. Similar to the popular plastic-hinge concept, bending of the pipe is assumed to occur only at the fixed end and at the location supported by the restraint. Shear deformation is also neglected. The pipe bending moment deflection (or rotation) relation used for these locations is obtained from a static nonlinear cantilever beam analysis. Using the moment-rotation relation, nonlinear equations of motion of the pipe are formulated using an CPS/USAR CHAPTER 03 3.6-23 REV. 11, JANUARY 2005 energy consideration, and the equations are numerically integrated in small time steps to yield time-history information of the deformed pipe. A comprehensive verification program has been performed to demonstrate the conservatisms inherent in the PDA pipe whip computer program and the analytical methods utilized. Part of this verification program included an independent analysis by Nuclear Services Corporation, under contract to the General Electric Company, of the recirculation piping system for the 1969 Standard Plant Design. The recirculation piping system was chosen for study because of its complex piping arrangement and assorted pipe sizes. The NSC analysis included elastic-plastic pipe properties, elastic-plastic restraint properties and gaps between the restraint and pipe and is documented in Reference 6. The piping/restraint system geometry and properties and fluid blowdown forces were the same in both analyses. However, a linear approximation was made by NSC for the restraint load-deflection curve supplied by GE. This approximation is demonstrated in Figure 3.6-2. The effect of this approximation is to give lower energy absorption of a given restraint deflection. Typically, this yields higher restraint deflections and lower restraint to structure loads than the GE analysis. The deflection limit used by NSC is the design deflection at one-half of the ultimate uniform strain for the GE restraint design. The restraint properties used for both analyses are provided in Table 3.6-5. A comparison of the NSC analysis with the PDA analysis, as presented in Table 3.6-6, shows that PDA predicts higher loads in 15 of the 18 restraints analyzed. This is due to the NSC model including energy absorbing effects in secondary pipe elements and structural members.

However, PDA predicts higher restraint deflections in 50% of the restraints. The higher deflections predicted by NSC for the lower loads are caused by the linear approximation used for the force - deflection curve rather than by differences in computer techniques. This comparison demonstrates that the simplified modeling system used in PDA is adequate for pipe rupture loading, restraint performance and pipe movement predictions within the meaningful design requirements for these low-probability postulated accidents. 3.6.2.2.2 Piping Other Than Reactor Recirculation Loop Piping - Inside Containment This subsection applies to all high-energy piping, both inside and outside containment, excluding the piping considered to be part of the reactor recirculation loop. 3.6.2.2.2.1 Determination of Pipe Thrust and Jet Loads 3.6.2.2.2.1.1 Circumferential Breaks The dynamic force of the jet discharge at the break location is based on the effective cross-sectional flow area of the pipe and on a calculated fluid pressure as modified by an analytically determined thrust coefficient. Line restriction flow limiters, positive pump controlled flow, and the absence of energy reservoirs are taken into account, as applicable, in the reduction of the jet discharge. Pipe whipping is assumed to occur in the plane defined by the piping geometry and configuration and to cause pipe movement in the direction of the jet reaction. 3.6.2.2.2.1.2 Longitudinal Breaks The dynamic force of the fluid jet discharge is based on a circular break area equal to the cross-sectional flow area of the pipe at the break location and on a calculated fluid pressure modified by an analytically determined thrust coefficient as determined for a circumferential break at the same location. Line restrictions, flow limiters, positive pump-controlled flow, and the absence of CPS/USAR CHAPTER 03 3.6-24 REV. 11, JANUARY 2005 energy reservoirs are taken into account, as applicable, in the reduction of jet discharge. Piping movement is assumed to occur in the direction of the jet reaction unless limited by structural members, piping restraints, or piping stiffness. 3.6.2.2.2.1.3 Pipe Blowdown Force and Wave Force The calculation of the magnitude and duration of the wave force acting on bounded pipe segments is based on a design guide for estimating discharge forces by Moody (Reference 1). The calculation of the blowdown force is based on either an exact computer model (Reference 9) or on the following simplified conservative methodology. The calculation of the blowdown force is consistent with Reference 1, and with Section 6.0 of ANSI N176 dated January, 1978 (Reference 2). If there is a fluid reservoir having sufficient capacity to develop a steady jet for a significant interval, the magnitude of the steady-state blowdown force used for saturated steam, saturated water, or a saturated steam and water mixture is equal to 1.26 P o A e for frictionless fluid flow (where P o equals the stagnation pressure of the initial vessel fluid and A e equals the break area). The magnitude of the steady-state blowdown force used for subcooled water varies from 1.26 P o A e to 2.0 P o A e for frictionless fluid flow depending on the degree of subcooling. However, the steady-state blowdown force is reduced by taking frictional effects into consideration as per Reference 2. For break locations where the frictional effects are significant, the blowdown force on the broken pipe segment is further reduced by considering the effect of wave propagation and reflection. Figure 3.6-3

shows the blowdown force on the pipe versus time for circumferential breaks. The pipe thrust used for longitudinal breaks is equal to the largest circumferential blowdown force at the same break location in accordance with Subsection 3.6.2.2.2.1.2. Nomenclature used in Figure 3.6-3 is defined below. a. Three different blowdown magnitudes are calculated: 1. F impulse = Fimp = P o A e 2. Fintermediate = F int = (P o A e - F w) F imp = F int implies F w = 0 where F w = wave force (transient),

A e = pipe flow area, and P o = line pressure. 3. Fsteady state = F ss b. Fw initial is determined from Figure 9-23 of Reference 1. F w initial for flashing water for pressures not shown in Figure 9-23 is equal to (P - 1.26 P sat) A (where P sat equals the saturation pressure of the initial pipe fluid). c. Fsteady state = Fss is determined in accordance with Reference 2.

CPS/USAR CHAPTER 03 3.6-25 REV. 11, JANUARY 2005 d. T imp = Time to Fintermediate for circumferential breaks and is determined by dividing the distance to breaks and is determined the first elbow from the break by the sonic speed of the significant fluid wave. The sonic wave speed (C) is

determined from Figure 9-29 of Reference 1. e. F final = The larger of F int or F ss. 3.6.2.2.2.2 Methods for the Dynamic Analysis of Pipe Whip Pipe whip restraints provide clearance for thermal expansion during normal operation. If a break occurs, the restraints or anchors nearest the break are designed to prevent unlimited movement at the point of break (pipe whip). Simplified models of the local region near the break were analyzed to calculate displacement of the pipe and restraint. These calculated displacements were then used to calculate strains in the pipe, and were compared to allowable restraint deflection. A finite difference model was used (Reference 10) for the pipe moment-curvature and the restraint resistance-displacement functions. The simplified models shown in Figure 3.6-5 were used to represent the local region near the break and to calculate the displacement in the restraint as well as the displacements and strains in the pipe. 3.6.2.2.2.2.1 Finite Difference Analysis A finite difference formulation specialized to the case of a straight beam and neglecting axial inertia and large deflection effects is used for the analysis of pipe whip of stainless and carbon steel pipes. The dynamic analysis is performed by direct numerical time integration of the equations of motion. The equations of motion are of the form:

1-M - 2M M - )y" m - (P h k k 1 k k k k+=+ (3.6-15) where: h k is the node spacing, P k is the externally applied lateral loads at node k, m k is the lumped mass at node k, y k is the lateral deflection at node k, and M k is the internal resisting moment in the beam at node k.

Power law moment-curvature relationship is assumed and the central difference approximation for the curvature

()1 k k 1 k 2 y y 2 y h 1++ is used.

CPS/USAR CHAPTER 03 3.6-26 REV. 11, JANUARY 2005 A timewise central-difference scheme is used to solve the dynamic equations ()()()()t t t t 2 t t y y" y t y++= (3.6-16) and for the first time step

()()0 2 y t t y= (3.6-17) A time step not more than 1/10 the shortest period of vibration is used in the integration. 3.6.2.2.2.2.1.1 Elastic-Plastic Moment Curvature Law The pipe is assumed to obey an elastic-strain hardening plastic moment-curvature law with isotropic strain hardening. The symbols used are defined as follows: M = moment, M= current yield moment, E = elastic modulus of material at temperature, I = moment of inertia,

Z = EI, = curvature, = M/z = elastic curvature, = increment of plastic curvature, = = effective plastic curvature, and o = = permanent set curvature. At the end of each integration step, new values of o are calculated at each node.

The known values of , and M at the start of the step are used to calculate M, , M and by the following procedure:

if , Z/M o< M = z (o), and = 0; if , Z/M o>

CPS/USAR CHAPTER 03 3.6-27 REV. 11, JANUARY 2005

()(), sin F M M o o+== and = o - M / Z, where F() = K()n 3.6.2.2.2.2.1.2 Power Law Moment Curvature Relationship The following stress strain law is assumed in the plastic range:

n o R K= (3.6-18) The corresponding moment curvature law is

()n K M= (3.6-19) where: ()()()2/3 2/n r K 1 2/n r R R n 3 2 K n 3 i n 3 o+++=++ (3.6-20) or, to a good approximation,

()()n 3 i n 3 o 2 R R n 076.n 291.0 1 n 3 K 4 K+++= (3.6-21) in which:

R o = pipe outside radius, and R i = pipe inside radius. In the elastic range, the moment-curvature law is:

M = EI (3.6-22) The transition from elastic to plastic behavior on initial loading occurs at:

()K EI 1 n 1= (3.6-23) 3.6.2.2.2.2.1.3 Strain Rate Effects The effect of strain rate in carbon steel is accounted for by using a rate dependent stress strain law of the form CPS/USAR CHAPTER 03 3.6-28 REV. 11, JANUARY 2005

()()()+=G 4.40 1 , 5/1 (3.6-24) where G( is the static stress strain relationship. For stainless steels, the effect of strain rate is less pronounced so that a 10% increase in yield and ultimate strengths is used. The selection of material properties is discussed in Attachment A3.6. 3.6.2.2.2.2.1.4 Restraint Behavior The analysis is capable of handling the bilinear or power law restraint behavior as shown in Figure 3.6-7. The behavior of the restraint is unidirectional. The restraint unloads elastically only to zero state, being left with a permanent set, and reloads along the same curve as shown in Figure 3.6-7. 3.6.2.2.2.3 Method of Dynamic Analysis of Unrestrained Pipes The impact velocity and kinetic energy of unrestrained pipes is calculated on the basis of the assumption that the segments on each side of the break act as rigid-plastic cantilever beams subject to piecewise constant blowdown forces. The hinge location is fixed either at the nearest restraint or at a point determined by the requirement that the shear at an interior plastic hinge is zero. The kinetic energy of an accelerating cantilever segment is equal to the difference between the work done by the blowdown force and that done on the plastic hinge. The impact

velocity V is found from the ex pression for the kinetic energy:

2 I eq V M (1/2) KE= (3.6-25) where M eq is the mass of the single degree of freedom dynamic model of the cantilever. The impacting mass is assumed equal to M eq. 3.6.2.3 Dynamic Analysis Methods to Verify Integrity and Operability 3.6.2.3.1 Jet Impingement Analyses and Effects on Safety Related Components 3.6.2.3.1.1 Jet Impingement Criteria and Characteristics The criteria used for evaluating the effects of fluid jets on safety related structures, systems and components are as follows: a. Safety-related structures, systems and components are not impaired so as to preclude essential functions. For any given postulated pipe break and consequent jet, those structures, systems and components needed to safely shut down the plant are identified. b. Safety related structures, systems and components which are not necessary to safely shut down the plant for a given break are not protected from the consequences of the fluid jet. c. Safe shutdown of the plant due to postulated pipe ruptures of the reactor coolant pressure boundary (RCPB) is not jeopardized by sequential failures of CPS/USAR CHAPTER 03 3.6-29 REV. 11, JANUARY 2005 safety-related piping. The required emergency core cooling system performance is maintained. d. Offsite dose limits specified in 10 CFR 100 are complied with.

e. Postulated design-basis breaks resulting in jet impingement loads are assumed to occur in high-energy lines at full (100%) power operation of the plant. f. Postulated through-wall leakage cracks are postulated in moderate-energy lines and are assumed to result in wetting and spraying of safety-related structures, systems and components. g. Reflected jets are considered only when there is an obvious reflecting surface (such as a flat plate) which directs the jet onto a safety-related target. Only the first reflection is considered in evaluating potential targets. h. Potential targets in the jet path are considered or the full extent of pipe displacement up to the calculated final position of the broken end of the ruptured pipe. This selection of potential targets is considered adequate due to the large number of breaks analyzed and the protection provided from the effects of these postulated breaks. Jet impingement load calculations where the stagnation condition at the postulated break location is steam or saturated water between 60 to 170 BARS pressure (1 BAR = 14.7 psi), that were prepared after April 1, 1984 are based on NUREG/CR-2913 (Reference 8).

NUREG/CR-2913 is a multidimensional computer study, which also accounts for the shock effects at the jet/target interface. When using the NUREG procedure, the impingement force includes the shape factor, K, as defined in Reference 2.

Jet impingement load calculations outside the range of pressure where the NUREG procedure is applicable (less than 60 BARS), or calculations where the stagnation condition at the postulated break location is subcooled water, or calculations that were prepared before April 1, 1984, are based on the following simplified, one-dimensional procedure. The analytical methods used to determine which targets are impingement upon by a fluid jet and the corresponding jet impingement load include: a. The impinging jet proceeds along a straight path.

b. The total impingement force acting on any cross-sectional area of the jet is time and distance invariant, with a total magnitude equivalent to the fluid blowdown

force as defined below. c. The jet impingement force is uniformly distributed across the cross-sectional area of the jet, and only the portion intercepted by the target is considered. d. The circumferential and longitudinal br eak opening is assumed to be a circular orifice of cross-sectional flow area equal to the effective flow area of the break.

CPS/USAR CHAPTER 03 3.6-30 REV. 11, JANUARY 2005 e. The jet impingement force is equal to the steady state value of the fluid blowdown force as calculated by the methods described in Subsection 3.6.2.2.1.1. f. The distance of jet travel is divided into two or three regions. Region 1 (see Figure 3.6-8) extends from the break to the asymptotic area. Within this region the discharging fluid flashes and undergoes expansion from the break area pressure to the atmospheric pressure. In Region 2 the jet remains at a constant diameter. In Region 3 interaction with the surrounding environment is assumed to start and the jet expands at a half angle of 10

°. g. Moody (Reference 1) has developed a simple analytical model for estimating the asymptotic area for steam, saturated water, and steam-water blowdown conditions. For fluids discharging from a break which are below the saturation temperature at the corresponding room pressure or have a pressure at the break area equal to the room pressure, free expansion does not occur. In these cases, the jet can be assumed to have a constant cross-sectional area equal to the break area. h. For fluids which are above the saturation temperature at room pressure, the jet model expands at a half angle of 45

° from the break to the asymptotic area (Region 1) for fully separated circumferential and longitudinal breaks. Assuming a linear expansion from the break area to the asymptotic area, the jet shape can be defined for Region 1 as well as Regions 2 and 3. Reference 2 is used to

determine the asymptotic area. i. Both longitudinal and fully separated circumferential breaks are treated similarly. The value of fl/D used in the blowdown calculation is also used for jet

impingement. j. Circumferential breaks with partial (i.e., l<D/2) separation between the two ends of the broken pipe, not significantly offset (i.e., no more than one pipe wall thickness lateral displacement) are more difficult to quantify; therefore, use of this procedure will be identified if applied. For these cases the following assumptions

are made: 1. The jet is uniformly distributed around the periphery. 2. The jet cross section at any cut through the pipe axis has the configuration depicted in Figure 3.6-8(B) and the jet regions are as therein delineated. 3. The jet force Fj = total blowdown F.

4. The pressure at any point intersected by the jet is:

R j j A F P=

CPS/USAR CHAPTER 03 3.6-31 REV. 11, JANUARY 2005 where: A R = the total 360

° area of the jet at a radius equal to the distance from the pipe centerline to the target. 5. The pressure of the jet is then multiplied by the area of the target submerged within the jet in the manner explained in Paragraphs k and l. 6. The area (A R) of the jet at target intersection distance r t from pipe centerline is calculated by using Reference 2 to determine r A (the distance from the pipe centerline to the plane of asymptoticity) and the relationship

] (B) 8-3.6 Figure [See l r 2 A R t R= where: A A = asymptotic jet area A B = break area D = pipe inside diameter l = distance of pipe separation

l A = width of jet at r A and infinitely outward l R = width of jet at r T k. Target loads are determined using the following procedures and assumptions: 1. For both the fully separated circumferential breaks and the longitudinal breaks, the jet is assumed to reach its asymptoticity expanding at a half angle of 45

° from the break. (Region 1, Figure 3.6-8(a)). For design purposes, the jet is assumed to have linear expansion within this region. The distance L from the break to the asymptotic area is calculated by =1 A A 2 D L 2/1 B A B where: A A = asymptotic jet area A B = jet cross sectional area at break D B = diameter of jet at break area The ratio of A A/A B is determined from Reference 2.

CPS/USAR CHAPTER 03 3.6-32 REV. 11, JANUARY 2005 2. The area within Region 2 can be assumed to be constant out to the beginning of Region 3 which starts at the intersection of a line drawn at a 10° half angle dotted line Figure 3.6-8(a) and (c) and the boundary of the jet. In Region 3 the area expands at a constant 10

° half angle. 3. After determination of the total area of the jet at the target, the jet pressure is calculated by x j i A F P= where: P = incident pressure, and A = area of the expanded jet at the target intersection. 4. The total force on any target which intercepts a portion of the jet is A P K F t jet o target= where: A t = the area of the target intercepted by the jet K o = the shape factor. The shape factor is related to the drag coefficient, C d , by K = 1/2 C d. Values of C d are given in Reference 2. l. For the partially separated circumferential breaks described in Paragraph j above, the target loads are calculated similarly, with the exception that the jet geometry is different according to Paragraph j and Figure 3.6-8(B). Evaluation of the potential targets to withstand the jet impingement loads is performed.

For analysis of piping systems as targets, evaluation of design adequacy is based on the following load combination for the faulted condition: Pressure + Weight + (SSE 2 + Jet 2)1/2 Functional capability is evaluated when required.

3.6.2.3.1.2 Protective Measures 3.6.2.3.1.2.1 Protection and Analyses Guidelines Protection against the dynamic effects of a pipe break is provided in the form of pipe whip restraints, equipment shields as required, and physical separation of piping, equipment, and instrumentation. The precise method used in choosing the kind of protection depends on other CPS/USAR CHAPTER 03 3.6-33 REV. 11, JANUARY 2005 limitations placed on the designer, such as accessibility, maintenance, and proximity to other pipes. The following are examples of present designs intended to better protect safety-related equipment from the consequences of the pipe breaks: a. The lines as described in Attachment B3.6 of the following systems inside the containment and dry well were analyzed for restraint against pipe whip and assessed for jet impingement: 1. main steam 2. feed water

3. RHR

4. RCIC 5. LPCS 6. HPCS

7. RWCU

8. reactor recirculation

9. nuclear boiler

10. standby liquid control b. The lines as described in Attachment B3.6 of the following systems outside of the containment were assessed for jet impingement and analyzed against pipe whip: 1. main steam 2. feed water

3. RCIC

4. RWCU 5. MSIV-LCS Dynamic effects associated with the LOCA do not compromise the integrity of the containment and drywell. The consequences of jet impingement do not result in any of the following: a. inability to insert control rods,

b. inability to isolate the reactor coolant pressure boundary, and c. inability to meet the core cooling system requirements. Valves which are normally closed and are not signalled to be open were assumed to be closed.

Impacted active equipment (e.g., valves and instruments) are considered able to perform their intended functions if loads are shown to be within allowable limits, otherwise, shields must be provided. Impacted passive equipment (pipes, restraints, and structures) are considered capable of continuing to perform their intended functions. Protection of the reactor pressure vessel from the surface impact effects of a pipe whip need not be considered because the impact energy is insufficient to cause loss of the functional integrity of the vessel.

CPS/USAR CHAPTER 03 3.6-34 REV. 11, JANUARY 2005 3.6.2.3.1.2.2 Equipment Shields for Isolation Equipment shields are selectively provided as required in order to isolate the equipment necessary to ensure segregation of the redundant systems of an accident and prevent it from causing a further chain accident. These shields are designed to with stand the rupture forces from piping and jets. 3.6.2.3.1.2.3 Jet Impingement Shields Jet impingement shields are also selectively provided as required to limit the consequence of rupture of the piping and are designed to withstand the resultant jet forces. 3.6.2.3.1.2.4 Separation Independence of redundant safety systems and components is maintained in most cases by separating the redundant components so that no single postulated event can prevent the safety-related function from occurring. This is achieved by the following: a. physical separation of source and target,

b. routing of cables so that different penetrations and paths are utilized to ensure that one event will not preclude both the primary and backup components from fulfilling their design function, c. deflection utilized to redirect a jet spray from an essential component, d. utilization of intermediate components and structure to intercept and defray forces, and e. location of duplicate instrument lines to ensure that one cause will not preclude each of the redundant systems from fulfilling its design function. 3.6.2.3.1.2.5 Acceptability of Analysis The postulation of high energy line break locations and the conservative analysis of resulting jet thrust and impingement have been used to identify areas where restraints or other protection devices are required to protect safety-related systems and components. 3.6.2.3.2 Pipe Whip Effects on Safety-Related Components This section of the USAR provides the criteria and methods used to evaluate the effects of pipe displacements on safety-related structures, systems and components following a postulated pipe rupture. The criteria which are used for determining the effects of pipe displacements on components are as follows: a. Components such as vessel safe ends and valves which are attached to the broken piping system and do not serve a safety function or whose failure would not further escalate the consequences of the accident, are not designed to meet ASME Code Section III imposed limits for essential components under faulted loading.

CPS/USAR CHAPTER 03 3.6-35 REV. 11, JANUARY 2005 b. If these components are required for safe shutdown, or serve a safety function to protect the structural integrity of an essential component, limits to meet the Code requirements for faulted conditions and limits to ensure operability, if required, are met. 3.6.2.3.3 Pipe Whip Restraints 3.6.2.3.3.1 Functional Requirements Pipe whip restraints differentiated from piping supports are designed to control the movement of a postulated ruptured pipe for an extremely low probability gross failure in a piping system carrying high-energy fluid. The piping integrity usually does not depend on the pipe whip restraints during normal, upset, emergency, or faulted conditions as defined in Section III of the ASME Boiler and Pressure Vessel Code. When piping integrity is lost because of a postulated break, the pipe whip restraint acts to limit the movement of the broken pipe to an acceptable

distance. The probability of pipe break accidents warrants that breaks be postulated in high-energy lines and that measures be taken to prevent consequential damage. The jet reaction force at the break is so large that snubbers and hangers not designed for pipe rupture loadings will usually be unable to prevent large displacement of the pipe. This large displacement of the pipe may cause damage to other mechanical, electrical, and structural systems necessary for safe shutdown of the plant. Also, if unrestrained, the blowdown thrust could produce strains equal to or greater than the ultimate strains in the pipe, resulting in local collapse of the pipe. In order to mitigate the effects of pipe break, restraints are provided near the points of the postulated breaks. The restraints are designed to absorb the impact energy and to resist the steady-state blowdown force after absorbing the pipe whip energy. Pipe whip restraints allow free thermal movements at all times, during operation and shutdown of the plant. 3.6.2.3.3.2 Types of Pipe Whip Restraints Three different types of pipe whip restraints are used as discussed below: a. Tension Restraints A typical tension restraint is shown in Figure 3.6-11. The tension restraint is composed of a variable assembly of U-shaped steel rods joined together to form a restraint of a specified resistance. b. Crushable Material Restraints Figure 3.6-12 shows a typical crushable material restraint. The crushable material absorbs energy only under compressive loading. c. Two-legged Restraints Figure 3.6-13 shows a typical two-legged restraint. Rods serve as yielding members for tensile loads. The energy associated with compressive loads is absorbed by the crushable energy-absorbing material.

CPS/USAR CHAPTER 03 3.6-36 REV. 11, JANUARY 2005 3.6.2.3.3.3 Loading and Load Combinations The pipe whip restraints are designed for the governing load combinations under abnormal/severe or abnormal/extreme loading conditions as per Table 3.8-2. 3.6.2.3.3.4 Design Requirements For reactor recirculation piping, the dynamic analysis for pipe whip restraints is performed using the Pipe Dynamic Analysis (PDA) program as described in Subsection 3.6.2.2.1.2. For other piping, the dynamic analysis for pipe whip restraints is performed using the Pipe Whip Restraint Reaction Analysis (PWRRA) programs. This program provides resultant force-time histories which can then be input into the Response Spectrum Generation (RSG) program to generate

dynamic load factors. The yielding portion of the restraint is designed for the peak dynamic load. The non-yielding portion of the restraint is designed for the equivalent static load. The functions of PWRRA are explained in detail in Subsection 3.6.2.2.2.2 and Appendix C, Section 25. The description of RSG is presented in Appendix C, Section 16. 3.6.2.3.3.5 Design Limits Allowable steel stresses for non-yielding members are taken as 1.6 times AISC allowable but not more than 0.95 F y where F y = specified minimum yield stress. Yielding in tension rods is limited to 50% of the ultimate strain.

Crushable material design is based on energy absorption principles. Deflection is controlled by the design energy. The honeycomb material thickness is designed such that the strain under this deflection is limited to 80% of the strain at which the honeycomb starts to strain-harden, and lies within the horizontal portion of the stress strain curve of the material. This ensures that the honeycomb material will not experience a deflection in excess of that defined by the horizontal portion of the load deflection curve. 3.6.2.4 Guard Pipe Assembly Design Criteria The design of Code Class MC guard pipes and mechanical penetration assemblies conforms to ASME BPV Code, Section III, Division 1, 1974, including applicable addenda and Code Cases through Summer 1974. Details and design considerations of the Class MC guard pipe assemblies are discussed in Subsections 3.8.1.1.3.1, 3.8.1.5.3, and 3.8.1.5.5. The inservice inspection of the guard pipe assemblies is in accordance with ASME BPV Code, Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components (applicable edition/addenda as required by 10CFR50.55a). Penetration assembly components are arranged in a way that accessibility for periodic weld examinations and other inservice inspections, as applicable, are provided.

CPS/USAR CHAPTER 03 3.6-37 REV. 11, JANUARY 2005 3.6.2.5 Material to be Submitted for the Operating License Review 3.6.2.5.1 Implementation of Criteria for Defining Pipe Break Location and Orientation 3.6.2.5.1.1 Postulated Pipe Breaks in Recirculation Piping System - Inside Containment The criteria for selection of postulated pipe breaks in the recirculation piping system, inside containment, are provided in Subsection 3.6.2.1.6.1. The postulated pipe break locations and types selected in accordance with these criteria are shown in Figure B3.6-31. Conformance with these criteria is shown by Table B3.6-31. 3.6.2.5.1.2 Pipe Whip Restraints for Recirculation Piping System Inside Containment The pipe whip restraints provided for this recirculation piping system are also shown in Figure B3.6-31. This system of restraints prevents unrestrained pipe whip resulting from a postulated rupture at any of the identified break locations. 3.6.2.5.1.3 Jet Effects for Postulated Ruptures of Recirculation Piping System - Inside Containment The effects of jet impingement from breaks in the reactor recirculation piping are detailed in Subsection D3.6.2.4. 3.6.2.5.2 Piping Other Than Reactor Recirculation Piping The following material pertains to the dynamic analyses applicable to piping systems inside and outside containment with the exception of the reactor recirculation loop piping. 3.6.2.5.2.1 Implementation of Criteria for Defining Pipe Break Locations and Configurations The locations and number of design-basis breaks associated with whip restraints, including postulated rupture orientations for the high-energy piping systems, are based on the criteria delineated in Subsection 3.6.2.1 and are shown in Attachment B3.6. 3.6.2.5.2.2 Implementation of Criteria Dealing With Special Features Special protective devices in the form of pipe whip restraints and impingement shields are designed in accordance with Subsection 3.6.2.3. Pipe whip restraint locations, configurations, and orientations in relation to break locations are included in Attachment B3.6. Where special protective devices are located in the vicinity of welds requiring augmented inservice inspection, one or both of the following criteria are met: a. Special protective devices are located at such a distance from all welds so as to allow inservice inspection. b. Special protective devices are removable so that inservice inspection can be performed.

CPS/USAR CHAPTER 03 3.6-38 REV. 11, JANUARY 2005 3.6.2.5.2.3 Acceptability of Analyses Results The postulation of break locations for high energy piping systems and analyses of the resulting jet thrust, impingement and pipe whip effects have been considered. Postulated pipe break results are included in Attachment B3.6. 3.6.2.5.2.4 Design Adequacy of Systems, Components, and Component Supports For each of the postulated breaks, the equipment and systems necessary to mitigate the consequences of the break and to safely shut down the plant (i.e., all essential systems and components) are identified in Subsection 3.6.1. The equipment and systems are protected against the consequences of each of the postulated breaks and cracks to ensure that their design-intended functions will not be impaired to unacceptable levels. Where it is necessary to restrict the motion of a pipe that would result from a postulated break, pipe whip restraints are included in the respective piping systems, or structural barriers or walls are designed to prevent the whipping of the pipe. Design adequacy of the restraints is included in Attachment B3.6. Results of a typical restraint analyses are given in Table B3.6-35 The structure and structural barriers are designed to withstand the effects of jet impingement loads. The loading combinations and allowable design limits are given in Tables 3.8-1.1, 3.8-1.2, and 3.8-2. The evaluation of essential components under dynamic effects associated with jet impingement is presented in Attachment D3.6. 3.6.2.5.2.5 Implementation of Criteria Related to Protective Assembly Design Guard pipes are discussed in Subsection 3.6.2.4.

3.6.3 References

1. P. T, Lahey, Jr. and F. J. Moody, "Pipe Thrust and Jet Loads," The Thermal-Hydraulics of a Boiling Water Nuclear Reactor, Section 9.2.3. pp. 375-409, Published by American Nuclear Society, Prepared for the Division of Technical Information, United States Energy Research and Development Administration, 1977. 2. ANSI N176 Design Basis for Protection of Nuclear Power Plants Against Effects of Postulated Pipe Rupture , Draft, January 1978 3. GE Spec. No. 22A2625 - "System Criteria and Applications for Protection Against the Dynamic Effects of Pipe Breaks." 4. RELAP3 - A computer Program for Reactor Blowdown Analysis IN-1321, issued June 1979, Reactor Technology TID-4500. 5. GE Report NEDE-10313 - "PDA - Pipe Dy namic Analysis Program for Pipe Rupture Movement" (Proprietary Filing)

CPS/USAR CHAPTER 03 3.6-39 REV. 11, JANUARY 2005 6. Nuclear Services Corporation Report No. GEN-02-02, "Final Report Pipe-Rupture Analysis of Recirculation System for 1969 Standard Plant Design." 7. GE Safety Evaluation Report for the Design of GESSAR-238, NSSS (Docket No. STN50-550), page 3-4. 8. NUREG/CR-2913, SAND 82-1935, R4, "Two-Phase Jet Loads." 9. RELAP4/MOD5, Computer Program User's Manual, 09.8.026-5.5.

10. Pipe Whip Restraint Reaction Analysis User's Manual, 09.5.125-2.1.

11. NUREG-0853, "Safety Evaluation Report Related to the Operation of Clinton Power Station, Unit No. 1, "Supplement No. 5, January 1986. 12. NUREG-1061, Volume 3, November 1984. 13. ANS-58.2 (ANSI N176), "Proposed American National Standard Design Basis for Protection of Light Water Nuclear Power Plants Against Effects of Postulated Pipe

Rupture." 14. Shapiro, A.H., "The Dynamics and Thermodynamics of Compressible Fluid Flow", Vol, Ronald Press, NY, 1965. 15. Webb, S.W., "Evaluation of Subcooled Water Thrust Forces", Nuclear Technology, Vol. 31, October 1976. 16. Hanson, G.H., "Subcooled - Blowdown Forces on Reactor System Components: Calculation Method and Experimental Configuration", Idaho Nuclear Corporation Report IN-1354, June 1970.

CPS/USAR CHAPTER 03 3.6-40 REV. 11, JANUARY 2005 TABLE 3.6-1 SYSTEMS IMPORTANT TO PLANT SAFETY Piping Systems Main Steam (including the Automatic Depressurization System) (MS)

Feedwater (FW) Reactor Recirculation (RR)

Diesel Oil (DO) Fuel Pool Cooling & Cleanup (FC)

Shutdown Service Water (SX)

Combustible Gas Control (HG)

Suppression Pool Makeup (SM) MSIV Leakage Control (IS)

Low Pressure Core Spray (LP)

High Pressure Core Spray (HP)

Residual Heat Removal (including the Low Pressure Coolant Injection) (RH)

Control Rod Drive (RD)

Standby Liquid Control (SC) Reactor Core Isolation Cooling (RI)

Control Room HVAC (VC)

Diesel Generator Room Ventilation (VD)

Standby Gas Treatment (VG)

Shutdown Service Water Ventilation (VH)

Essential Switchgear Heat Removal (VX) ECCS Pump Room Cooling (VY) Refrigeration Piping Switchgear Heat Removal (RG)

Other Systems Neutron Monitoring

Containment Atmosphere Monitoring Leak Detection Process Radiation Monitoring

Safety Related Standby Power (i.e., Diesel Generator, Auxiliary DC Power System) Containment Isolation Systems

CPS/USAR CHAPTER 03 3.6-41 REV. 11, JANUARY 2005 TABLE 3.6-2 HIGH-ENERGY FLUID SYSTEMS SYSTEMS NOTES Main Steam (MS) Extraction Steam (ES) (1)

Feedwater (FW)

Condensate (CD) (1)

Condensate Booster (CB) (1)

Control Room HVAC (VC)

Heater Drains (HD) (1)

Misc. Vents & Drains (DV) (1) Turbine Drains (TD) (1) Turbine Gland Steam Seal Steam (GS) (1)

MSIV Leakage Control (IS)

Reactor Recirculation (RR) (2)

Low Pressure Core Spray (LP) (2) High Pressure Core Spray (HP) (2) Nuclear Boiler (NB) (2)

Residual Heat Removal (RH) (2)

Reactor Water Cleanup (RT)

Standby Liquid Control (SC) (2) Reactor Core Isolation Cooling (RI) Control Rod Drive (RD) (2)

Off Gas (OG)

(1) Radwaste Chemical Waste Process (WF) (1)

Chemical Radwaste Reprocessing & Disposal (WZ) (1)

Radwaste Sludge Process (WX) (1) Auxiliary Steam (AS) (1) Post Accident Sampling (PS) (2)

Containment Monitoring (CM) (2)

Laboratory HVAC (VL)

(1) Not considered an initiating system for piping failure because of complete physical separation from safety related systems, components and structures. (These systems are located in Non-Category I structures where there are no safety related components or systems.) (2) The only high-energy portions of these systems are those portions which make up the reactor coolant boundary. (See Figure 3.6-1 for exact boundaries.)

CPS/USAR CHAPTER 03 3.6-42 REV. 11, JANUARY 2005 TABLE 3.6-3

SUMMARY

OF SUBCOMPARTMENT PRESSURIZATION ANALYSES INITIAL CONDITIONS DBA BREAK CONDITIONS CUBICLE DESIGNATION SUBCOMPARTMENT DESCRIPTION VOLUME (ft 3) HEIGHT (ft) FLOW CROSS SECTIONAL AREA (ft 2) BOTTOM ELEVATION (ft) TEMP (°F) PRESS. (psia) HUMID.(%) BREAK LOC. VOL. NO. BREAK LINE BREAK AREA (ft 2) BREAKTYPE* CALC.PEAK PRESS.DIFF. (psig) DESIGN PEAK PRESS. DIFF. (psig) 1 Aux. Bldg. Floor Drain Sys. 14,650.0 22.50 250.0 712.00 122.0 14.7 0.1 1 1RH04A14 0.0091 L 1.25 1.25 2 RHR Pump Rm. 'C' 33,650.0 28.50 700.0 707.50 122.0 14.7 0.1 2 1RH08A14 0.0091 L 0.60 0.60 3 RHR Pump Rm. 'B' 24,650.0 26.75 610.0 707.50 104.0 14.7 0.1 4 1RH40AB10 0.7882 C 5.07 5.07 4 RHR Heat Exchanger Rm. 'B' 50,500.0 91.00 555.0 707.50 104.0 14.7 0.1 4 1RH40AB10 0.7882 C 5.54 5.54 5 RCIC Pump and Turbine Rm. 26,000.0 26.50 750.0 707.50 108.0 14.7 0.1 5 1RI05A4 0.1660 C 6.16 6.16 6 RHR Heat Exchanger Rm. 'A' 49,100.0 91.00 535.0 707.50 104.0 14.7 0.1 6 1RH40AA10 0.7882 C 5.54 5.54 7 RHR Pump Rm. 'A' 23,700.0 26.75 450.0 707.50 104.0 14.7 0.1 6 1RH40AA10 0.7882 C 5.07 5.07 8 LPCS Pump Rm. 48,892.0 28.25 660.0 707.50 104.0 14.7 0.1 8 1RH03AA14 0.0091 L 0.90 0.90 9 Corridor 82,900.0 28.25 465.0 707.50 104.0 14.7 0.1 14 1RT02AB3 0.0751 C 0.34 0.34 10 Ground Floor West 104,000.0 23.50 1100.0 737.00 104.0 14.7 0.1 10 1RH04A14 0.0091 L 0.48 0.48 11 MSIV Rms. 9,486.0 13.00 200.0 737.00 150.0 14.7 0.1 5 1RI05A4 0.1660 C 2.34 2.34 12 Ground Floor East 59,700.0 23.50 1100.0 737.00 104.0 14.7 0.1 14 1RT02AB3 0.0751 C 0.40 0.40 13 Personnel Access Area 10,500.0 10.50 165.0 737.00 104.0 14.7 0.1 13 1RH03AA14 0.0091 L 0.60 0.60 14 RWCU Pump Rms. 1,275.0 10.00 35.0 737.00 104.0 14.7 0.1 14 1RT02AB3 0.0751 C 4.7 4.7 15 Pipe Chase 485.0 9.00 14.0 750.00 122.0 14.7 0.1 15 1RT02F6 0.3300 C 10.95 10.95 16 Pipe Tunnel 14,000.0 8.50 250.0 750.50 122.0 14.7 0.1 16 1RT02F6 0.3300 C 3.84 3.84 17 Mezzanine Floor West 78,700.0 17.50 785.0 762.00 104.0 14.7 0.1 10 1RH04A14 0.0091 L 0.18 0.18 18 Mezzanine Floor East 78,400.0 17.50 785.0 762.00 104.0 14.7 0.1 18 1RH03AA14 0.0091 L 0.48 0.48 19 Main Floor West 80,950.0 17.75 810.0 781.00 104.0 14.7 0.1 18 1RH03AA14 0.0091 L 0.36 0.36 20 Main Floor East 86,650.0 19.00 865.0 781.00 104.0 14.7 0.1 18 1RH03AA14 0.0091 L 0.36 0.36 A Auxiliary/Main Steam Tunnel 141,356 Varies Varies Varies 150 14.7 0.1 A 1MS01EA24 1FW02EA20 3.1300 9.318 C 13.8 13.8

• L = 'through-wall leakage crack' C = 'circumferemtial pipe rupture'

CPS/USAR CHAPTER 03 3.6-43 REV. 11, JANUARY 2005 TABLE 3.6-4 SUBCOMPARTMENTS USED FOR DIVISIONAL SEPARATION SUBCOMPARTMENT LOCATION RHR Pump "A" Cubicle (and RHR Heat Exchanger "A" Cubicle)

El. 707 ft-6 in., Auxiliary Building RHR Pump "B" Cubicle (and RHR Heat Exchanger "B" Cubicle)

El. 707 ft-6 in., Auxiliary Building RHR Pump "C" Cubicle El. 707 ft-6 in., Auxiliary Building LPCS Pump Cubicle El. 707 ft-6 in., Auxiliary Building RCIC Pump Cubicle El. 707 ft-6 in., Auxiliary Building HPCS Pump Cubicle El. 707 ft-6 in., Fuel Building Safety-Related Division 1 Switchgear Area El. 781 ft-0 in., Auxiliary Building (Area bounded by column row 117 wall on the west, containment and column row AD walls on the south, column row 124 wall on the east, and column row S wall on the north.) Safety-Related Division 2 Switchgear Area El. 781 ft-0 in., Auxiliary Building (Area bounded by

column row 102 wall on

the west, column row AD and containment walls on the south, column row 107 wall on the east, and column row S wall on the north.)

CPS/USAR CHAPTER 03 3.6-44 REV. 11, JANUARY 2005 TABLE 3.6-5 RESTRAINT DATA General Restraint Data for 1 Bar of a Restraint F = C 2 ( restraint) n Where restraint = pipe - total clearance

Pipe Size (In) Rest Load Direction C 2 n Limit Restraint Initial Clearance Effective Clearance Total Clearance 12 0° 27,733 -24 6.129 4 1.941 5.941 12 90° 14,795 -401 9.063 4 12.247 16.247 16 0° 109,265 -24 6.278 4 1.934 5.934 16 90° 62,599 -377 8.978 4 12.187 16.187 24 0° 102,228 -24 8.222 4 1.984 5.984 24 90° 55,531 -375 11.972 4 13.685 17.685 24 38°* 109,888 -24 5.588 4 5.698 9.698 24 52°* 109,835 -24 5.473 4 8.462 12.462

• Applies to Restraint RCR 3 only.

CPS/USAR CHAPTER 03 3.6-45 REV. 11, JANUARY 2005 TABLE 3.6-6 COMPARISON OF PDA AND NSC CODE No. Bars Load (kips) Restraint Deflection (in.) Percent of Design Restraint Deflection Pipe Deflection (in.) Break ID No.* Restraint ID No.* Force Vector (degrees)(a) PDA NSC PDA NSC PDA NSC PDA NSC PDA NSC RC1J RCR1 0 5 5 803.3 788.3 6.6 7.9 79.9 96.4 17.7 15.6 RC2LL RCR1 90 5 5 766.4 458.4 15.0 7.5 125.2 62.6 35.8 24.5 RC3LL RCR2 0 6 6 747.0 639.7 2.3 3.7 27.7 45.4 17.2 20.1 RC3LL RCR2 90 6 6 796.6 780.3 10.2 10.5 85.4 88.1 41.5 43.0 RC4LL RCR3 0 5 5 846.0 838.4 8.2 8.1 99.2 98.0 18.9 16.4 RC4LL RCR3 52 8 8 1,319.0 1,073.9 5.4 4.2 99.2 76.9 23.4 17.3 RC4CV RCR3 38 8 8 1,260.7 1,275.0 4.5 5.6 80.4 99.9 22.6 18.7 RC6AV RCR3 38 8 8 928.5 722.5 1.3 1.8 22.5 31.7 23.7 95.4 RC7J RCR7 0 6 6 953.3 801.6 6.3 5.8 76.4 70.1 16.5 21.6 RC8LL RCR6 90 4 4 599.0 NA** 8.3 NA 69.2 NA 26.8 NA RC8LL RCR7 90 6 6 895.0 NA 8.2 NA 68.2 NA 29.3 NA RC9CV RCR6 0 4 4 575.8 520.2 4.2 5.5 50.6 67.3 13.2 14.6 RC9LL RCR8 90 6 6 830.2 546.8 11.4 6.8 95.3 56.9 36.7 26.2 RC11A RCR8 90 6 6 818.3 493.6 11.0 6.0 91.7 50.1 31.4 23.7 RC12 RCR9 0 6 6 NA 832.9 NA 6.3 NA 76.9 NA 15.7 RC13 RCR10 0 4 4 668.4 478.0 5.9 3.7 93.5 58.4 13.4 10.4 RC14CV RCR20 0 8 8 285.0 309.6 2.8 5.9 46.3 95.9 15.5 14.0 RC14LL RCR20 90 8 8 116.3 129.9 1.0 3.4 10.5 37.1 22.0 23.6 RC16 RCR11 0 4 4 687.4 518.4 6.6 4.4 105.1 69.9 15.4 10.2

• See Figure 3.6-2 (a) Force Vector Represented as - ** NA - Data Not Available

CPS/USAR CHAPTER 03 3.6-46 REV. 11, JANUARY 2005 TABLE 3.6-7 MASS AND ENERGY RELEASE RATE DATA LINE BREAK IN STEAM TUNNEL OUTSIDE CONTAINMENT TIME (sec) LIQUID MASS FLOW RATE (lb m/sec) STEAM MASS FLOW RATE (lb m/sec) LIQUID ENTHALPY (BTU/lb m) STEAM ENTHALPY (BTU/lb m) TOTAL MASS RELEASE RATE (lb m/sec) TOTAL ENERGY RELEASE RATE (BTU/sec) FWL* MSL* FWL MSL FWL MSL FWL MSL 0 24041 0 0 10600 408.5 550 0 1189.9 34641 2.24E + 07 0.028 24041 0 0 10600 408.5 550 0 1189.9 34641 2.24E + 07 0.029 18603.5 0 0 10600 408.5 550 0 1189.9 29203.5 2.02E + 07 0.34 18603.5 0 0 10600 408.5 550 0 1189.9 29203.5 2.02E + 07 0.341 5307.5 0 0 10600 408.5 550 0 1189.9 15907.5 1.48E + 07 1.000 5307.5 0 0 10600 408.5 550 0 1189.9 15907.5 1.48E + 07 1.001 5307.5 17395 0 4605 408.5 550 0 1189.9 27307.5 1.72E + 07 1.750 5307.5 17395 0 4605 408.5 550 0 1189.9 27307.5 1.72E + 07 1.751 5307.5 20400 0 1600 408.5 550 0 1189.9 27307.5 1.53E + 07 4.000 5307.5 20400 0 1600 408.5 550 0 1189.9 27307.5 1.53E + 07 5.500 5307.5 0 0 0 408.5 550 0 1189.9 5307.5 2.17E + 06 6.000 5307.5 0 0 0 408.5 550 0 1189.9 5307.5 2.17E + 06 The feedwater line break mass and energy releases were also evaluated at a reduced feedwater temperature (RFWT) of 380

°F and an enthalpy of 354.8 Btu/lbm. The subcompartment analyses were also completed for the case of reduced feedwater temperature (RFWT) at LPU conditions. The resulting feedwater line break mass release increased by 5% and the energy releases increased by 3%.

• FWL - Feedwater Line MSL - Main Steam Line CPS/USAR CHAPTER 03 3.6-47 REV. 11, JANUARY 2005 TABLE 3.6-8 SUBCOMPARTMENT NODAL DESCRIPTION LINE BREAK IN STEAM TUNNEL OUTSIDE CONTAINMENT INITIAL CONDITIONS DBA BREAK CONDITIONS VOLUME NO. DESCRIPTION CROSS- SECTIONAL AREA (ft 2) VOLUME (ft 3) TEMP.(°F) PRESS.(psia) HUMID. (%) BREAK LOC. VOL. NO. BREAK LINE BREAK AREA (ft 2) BREAK TYPE CALC.**PEAK PRESS DIFF. (psig) 1 Main Steam Tunnel 562 67549 150 14.7 .1 CASE 1* 13.8 2 Main Steam Tunnel 590 44948 150 14.7 .1 CASE 2* 8.2 3 Main Steam Tunnel 1100 28859 150 14.7 .1 CASE 3* 6.0 4 Turbine Building: Basement, Grade Flr. Mezzanine Flr. 2000 780300 104 14.7 .1 - 5 Turbine Building:

Main Flr. 3000 275000 104 14.7 .1 - 6 Atmosphere 10000 10 10 104 14.7 .1 -

________________________

• Indicates the break location for each case. The break was a simultaneous double-ended guillotine break of one main steam line and one feedwater line. ** The calculated peak pressure difference = the peak node pressure =14.7 psia.

CPS/USAR CHAPTER 03 3.6-48 REV. 11, JANUARY 2005 TABLE 3.6-9 SUBCOMPARTMENT VENT PATH DESCRIPTION LINE BREAK IN STEAM TUNNEL OUTSIDE CONTAINMENT VENT PATH NO. FROM VOL. NODE NO. TO VOL. NODE NO. DESCRIPTION OF VENT PATH FLOW CHOKED UNCHOKED AREA (ft 2) LENGTH (ft) INERTIA (ft-1) HYDRAULIC DIAMETER (ft) FRICTION K, ft/d TURNING LOSS, K EXPANSION, K CONTRACTION, K TOTAL 1* 1 2 UNCHOKED 406.5 31.7 .078 22.8 2.5 2 2 3 UNCHOKED 590.0 40.1 .068 27.4 1.6 3 3 4 UNCHOKED 635.0 8.3 .013 28.4 2.1 4** 4 5 UNCHOKED 976.5 10.7 .011 35.3 4.1 5*** 5 6 UNCHOKED 2000.0 16.0 .008 50.5 2.9 6**** 0 1 1.0 00.0 .000 00.0 0.0

___________________________

• Opened on a differential pressure in either direction of 0.37 psid. ** Opened on a differential pressure of 0.4 psid from 4 to 5. *** Opened on a differential pressure of 0.7 psid from 5 to 6. **** The break flow: CASE 1 into Node 1, CASE 2 into Node 2, CASE 3 into Node 3.

CPS/USAR CHAPTER 03 REV. 11, JANUARY 2005 ATTACHMENT A3.6 SELECTION OF PIPE MATERIAL PROPERTIES FOR USE IN PIPE WHIP ANALYSIS

CPS/USAR CHAPTER 03 A3.6-1 REV. 11, JANUARY 2005 ATTACHMENT A3.6 SELECTION OF PIPE MATERIAL PROPERTIES FOR USE IN PIPE WHIP ANALYSIS The selection of yield and ultimate strength values for piping for use in pipe whip analysis is discussed in Section 6.3.2.3 of Reference 1. This part of the standard permits the use of representative or actual test data values of material properties. Minimum ASME code values can also be used with more conservative results. A substantial amount of elevated temperature test data for A106 Grade B carbon steel is given in Reference 2. Material property values based on this data are obtained and used. Since little test data is available for TP304, TP304L, and TP316 stainless steels, ASME code specified values are used with the realization that they are very conservative. The power law stress strain relationship is used for all steels.

n K s= (1) The effect of strain rate in carbon steels is accounted for (as suggested in Reference 3) by modifying Equation (1) as follows:

()()n p/1 K D 1 ,+= (2) D = 40.4 sec p = 5 This modification has been widely used (see for examples References 3 and 4). For stainless steels, the effect of strain rate is less pronounced (Reference 5) so that the use of a 10% increase in yield and ultimate strengths as suggested in Reference 1 is used. A106 Grade B Material Properties at 600

° F The results of tests on 73 specimens are given in Reference 2. Twenty-one were tested at 600° F, twenty-two at room temperature, and the rest at temperatures between 200

° F and 585° F. Yield stress (0.2% offset values were measured) was shown to decrease, ultimate stress to increase, with increasing temperature. The minimum yield stress of any of the 73 specimens tested was 31.6 ksi, and the average for the 21 tested at 600

° F was 36.01 ksi with a sigma value of 3.6 ksi. The minimum ultimate stress value for all specimens was 64.4 ksi, and the average for the 22 tested at room temperature 71.79 ksi with a sigma value of 4.72 ksi. The strength coefficient K and the hardening exponent n can be evaluated from the following equations:

(.002) K y= (3) n Kn u= (4)

CPS/USAR CHAPTER 03 A3.6-2 REV. 11, JANUARY 2005 Values of K and n obtained in this way are given in the following tabulation:

Yield Stress y (ksi) Ultimate Stress u (ksi) K (ksi) n Minimum 31.60 64.40 86.486 0.16201 Mean-Sigma 32.41 67.07 90.277 0.16484 Mean 36.01 71.79 96.080 0.15792 (Material properties in the tabulation, for A106 Grade B at 600

° F, are based on data from Reference 2.) The mean-sigma values of K = 90.277 ksi and n = 0.16484 are used for all temperatures 600

° F and below.

CPS/USAR CHAPTER 03 A3.6-3 REV. 11, JANUARY 2005 References

1. ANSI N176 Design Bases for Protection of Nuclear Power Plants Against Effects of Postulated Pipe Rupture , Draft, January 1978. 2. R. J. Eiber, et. al. Investigation of the Initiation and Extent of Ductile Pipe Rupture , Battelle Memorial Institute, Report BMI-1866, July, 1969. 3. S. R. Bodner and P. S. Symonds, "Experimental and Theoretical Investigation of the Plastic Deformation of Cantilever Beams Subjected to Impulsive Loading," JAM, December, 1962. 4. J. C. Anderson and A. K. Singh, "Inelastic Response of Nuclear Piping Subjected to Rupture Forces," ASME Paper No. 75-PVP-21. 5. C. Albertini and M. Montagnani, "Wave Propagation Effects in Dynamic Loadings," Nuclear Engineering and Design, 37-115-124, 1976.

CPS/USAR ATTACHMENT B3.6 POSTULATED PIPE BREAK RESULTS

CPS/USAR CHAPTER 03 B3.6-1 REV. 11, JANUARY 2005 ATTACHMENT B3.6 POSTULATED PIPE BREAK RESULTS Attachment B3.6 presents specific details discussed in Subsection 3.6.2.5.

The data presented is described below: 1. Location of all break locations that are postulated using the stress criteria or that require whip restraints are shown in Figures B3.6-1 through B3.6-34. These figures also show all required whip restraints. 2. The type of pipe break postulated, the pipe stresses, and the allowable stresses at the postulated break locations i are shown in Tables B3.6-1 through B3.6-34. Typically, the allowable stresses are lower bound values based on a temperature that envelopes all piping locations in the subsystem. These allowables are conservative for pipe rupture analysis purposes, and do not necessarily correspond to the allowable stresses used in the code stress analysis of these subsystems. 3. Typical results of pipe whip restraint analyses inside containment for high pressure core spray system are identified in Table B3.6-35. 4. Typical results to demonstrate design adequacy of those portions of high-energy piping penetrating containment for which additional stress criteria apply (i.e., break exclusion piping), and for which valve operability requirements must be met, are shown in

Table B3.6-36.

CPS/USAR CHAPTER 03 B3.6-2 REV. 11, JANUARY 2005 TABLE B3.6-1 BREAK DATA, FEEDWATER SUBSYSTEM FW-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR FW-C21 C 42480 58165 8160 43491 0.086 FW-C31 C 42480 Terminal End Break FW-C32 C,L 42480 47075 28735 23042 0.211 FW-C33 C,L 42480 60181 47159 14654 0.312 FW-C34 C,L 42480 69546 56638 13573 0.435 FW-C40 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-3 REV. 11, JANUARY 2005 TABLE B3.6-2 BREAK DATA, FEEDWATER SUBSYSTEM FW-02 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR FW-C1 C 42480 58165 8160 43491 0.086 FW-C11 C 42480 Terminal End Break FW-C12 C,L 42480 47075 28735 23042 0.211 FW-C13 C,L 42480 60181 47159 14654 0.312 FW-C14 C,L 42480 69546 56638 13573 0.435 FW-C20 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-4 REV. 11, JANUARY 2005 TABLE B3.6-3 BREAK DATA, FEEDWATER SUBSYSTEM FW-03 OUTSIDE CONTAINMENT BREAK** NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8(1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) 3 C Not Applicable Not Applicable 4 L Not Applicable Not Applicable 5 C Not Applicable Not Applicable 6 C Not Applicable Not Applicable 7 L Not Applicable Not Applicable 8 C Not Applicable Not Applicable 9 C Not Applicable Not Applicable A3 C Not Applicable Not Applicable A4 L Not Applicable Not Applicable A5 C Not Applicable Not Applicable A6 C Not Applicable Not Applicable A7 L Not Applicable Not Applicable A8 C Not Applicable Not Applicable A9 C Not Applicable Not Applicable FW03-1 C Not Applicable Terminal End FW03-2 C Not Applicable Terminal End

______________________

• Break type: C = circumferential, L = longitudinal

• • Breaks were based on fitting criteria.

CPS/USAR CHAPTER 03 B3.6-5 REV. 11, JANUARY 2005 TABLE B3.6-4 BREAK DATA, HPCS SUBSYSTEM HP-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR HP-C9 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-6 REV. 11, JANUARY 2005 TABLE B3.6-5 BREAK DATA, LPCS SUBSYSTEM LP-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR LP-C9A C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-7 REV. 11, JANUARY 2005 TABLE B3.6-6 BREAK DATA, MAIN STEAM SUBSYSTEM MS-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR MS-C68 L,C 42480 67675 49264 18661 0.0423 MS-C69 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-8 REV. 11, JANUARY 2005 TABLE B3.6-7 BREAK DATA, MAIN STEAM SUBSYSTEM MS-02 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR MS-C33 L,C 42480 64551 47389 23406 0.0306 MS-C34 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-9 REV. 11, JANUARY 2005 TABLE B3.6-8 BREAK DATA, MAIN STEAM SUBSYSTEM MS-03 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR MS-C56 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-10 REV. 11, JANUARY 2005 TABLE B3.6-9 BREAK DATA, MAIN STEAM SUBSYSTEM MS-04 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR MS-C20 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-11 REV. 11, JANUARY 2005 TABLE B3.6-10 BREAK DATA, MAIN STEAM DRAIN SUBSYSTEM MS-05 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR MS-C69 C 42480 Terminal End Break MS-C78 C 42480 Terminal End Break MS-C87 C 42480 Terminal End Break MS-C93 C 42480 39658 -- -- .130 MS-C94 C 42480 Terminal End Break

• Break type: C = circumferential..

CPS/USAR CHAPTER 03 B3.6-12 REV. 11, JANUARY 2005 TABLE B3.6-11 BREAK DATA, MAIN STEAM SUBSYSTEM MS-06 OUTSIDE CONTAINMENT BREAK** NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8(1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) MS-C201 C Not Applicable Not Applicable MS-C202 C Not Applicable Not Applicable

_______________________

• Break type: C = circumferential.

• • Breaks were based on fitting criteria.

CPS/USAR CHAPTER 03 B3.6-13 REV. 11, JANUARY 2005 TABLE B3.6-12 BREAK DATA, MAIN STEAM SUBSYSTEM MS-07 OUTSIDE CONTAINMENT BREAK** NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8(1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) 3L L Not Applicable Not Applicable 4C C Not Applicable Not Applicable 5L L Not Applicable Not Applicable 6C C Not Applicable Not Applicable 9L L Not Applicable Not Applicable 10C C Not Applicable Not Applicable 11L L Not Applicable Not Applicable 12C C Not Applicable Not Applicable A3L L Not Applicable Not Applicable A4C C Not Applicable Not Applicable A5L L Not Applicable Not Applicable A6C C Not Applicable Not Applicable A9L L Not Applicable Not Applicable A10C C Not Applicable Not Applicable A11L L Not Applicable Not Applicable A12C C Not Applicable Not Applicable

______________________

• Break type: C = circumferential, L = longitudinal **Breaks were based on fitting criteria.

CPS/USAR CHAPTER 03 B3.6-14 REV. 11, JANUARY 2005 TABLE B3.6-13 BREAK DATA, MAIN STEAM DRAIN SUBSYSTEM MS-38A OUTSIDE CONTAINMENT All High Energy Piping in Subsystem MS-38A is Break Exclusion. There are no postulated breaks on this subsystem.

CPS/USAR CHAPTER 03 B3.6-15 REV. 11, JANUARY 2005 TABLE B3.6-14 BREAK DATA, RHR SUBSYSTEM RH-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RH-C10 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-16 REV. 11, JANUARY 2005 TABLE B3.6-15 BREAK DATA, RHR SUBSYSTEM RH-03 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RH-C26 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-17 REV. 11, JANUARY 2005 TABLE B3.6-16 BREAK DATA, RHR SUBSYSTEM RH-05 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RH-C18 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-18 REV. 11, JANUARY 2005 TABLE B3.6-17 BREAK DATA, RHR SUBSYSTEM RH-34 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RH-C35 C 47580 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-19 REV. 11, JANUARY 2005 TABLE B3.6-18 BREAK DATA, RHR SUBSYSTEM RH-07 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + RELIEF VIV + EQ.10B)

• Break type: C = circumferential, L = longitudinal. Footnote: Terminal End Break at Penetration 1AB-0204 is not postulated as explained in Section D3.6.4.

CPS/USAR CHAPTER 03 B3.6-20 REV. 11, JANUARY 2005 TABLE B3.6-19 BREAK DATA, RHR SUBSYSTEM RH-08 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + RELIEF VIV + EQ.10B)

• Break type: C = circumferential, L = longitudinal. Footnote: Terminal End Break at Penetration 1AB-0202 is not postulated as explained in Section D3.6.4.

CPS/USAR CHAPTER 03 B3.6-21 REV. 12, JANUARY 2007 TABLE B3.6-20 BREAK DATA, RCIC SUBSYSTEM RI-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RI-C11 C 42240 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-22 REV. 11, JANUARY 2005 TABLE B3.6-21 BREAK DATA, RCIC SUBSYSTEM RI-02/RH-14 OUTSIDE CONTAINMENT BREAK** NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) RH-CA5 C 32400 Terminal End Break RH-CA55 C 32400 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-23 REV. 12, JANUARY 2007 TABLE B3.6-22 HAS BEEN DELETED.

CPS/USAR CHAPTER 03 B3.6-24 REV. 14 JANUARY 2011 TABLE B3.6-23 BREAK DATA, RWCU SUBSYSTEM, RT-01 INSIDE CONTAINMENT Note: See Calculation 066204(EMD) "Fatigue Analysis of Piping Subsystem 1RT-01" for Stress and Usage Factors for RWCU Subsystem, RT-01. BREAK CALCULATED STRESS (psi) NUMBER TYPE* 2.4S m (psi) EQ.10 EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RT-C1 C TERMINAL END BREAK RT-C28 C,L RT-C28B C TERMINAL END BREAK RT-C28C C RT-C35A C TERMINAL END BREAK RT-C39 C RT-C40 C RT-C40A C RT-C58A C,L RT-C27B C,L RT-C39A C RT-C79 C TERMINAL END BREAK

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-25 REV. 11, JANUARY 2005 TABLE B3.6-24 BREAK DATA, RWCU SUBSYSTEM RT-02 INSIDE CONTAINMENT BREAK** NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) RT-C151A C 32400 Terminal End Break RT-C161A C 32400 Terminal End Break RT-C251A C 32400 Terminal End Break RT-C255A C 32400 Terminal End Break RT-C266A C 32400 Terminal End Break RT-C315A C 32400 Terminal End Break RT-C365A C 32400 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-26 REV. 11, JANUARY 2005 TABLE B3.6-25 BREAK DATA, RWCU SUBSYSTEM RT-05 INSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ.9B + EQ.10) RT-C173B C 32400 Terminal End Break RT-C174C C 32400 Terminal End Break RT-C181 C 32400 Terminal End Break RT-C191 C 32400 Terminal End Break RT-C203A C 32400 Terminal End Break RT-C211A C 32400 Terminal End Break RT-C221A C 32400 Terminal End Break RT-C203B C 32400 46700 RT-C203C C 32400 34700

• Break type: C = circumferential, L = Longitudinal.

CPS/USAR CHAPTER 03 B3.6-27 REV. 11, JANUARY 2005 TABLE B3.6-26 BREAK DATA, RWCU SUBSYSTEM RT-06 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ. 9B + EQ. 10) RT-603 C 32400 Terminal End Break RT-604 C 32400 Terminal End Break RT-605 C 32400 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-28 REV. 11, JANUARY 2005 TABLE B3.6-27 BREAK DATA, RWCU SUBSYSTEM RT-07 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ. 9B + EQ. 10) RT-C703 C 32400 Terminal End Break RT-C704 C 32400 Terminal End Break RT-C705 C 32400 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-29 REV. 11, JANUARY 2005 TABLE B3.6-28 BREAK DATA, RWCU SUBSYSTEM RT-08 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ. 9B + EQ. 10) RT-C801 C 32400 Terminal End Break RT-C803 C 32400 Terminal End Break RT-C805 C 32400 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-30 REV. 11, JANUARY 2005 TABLE B3.6-29 BREAK DATA, RWCU DRAIN SUBSYSTEM RR-32 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RT-C85 C 32940 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-31 REV. 11, JANUARY 2005 TABLE B3.6-30 BREAK DATA, RWCU DRAIN SUBSYSTEM RR-33 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR RT-C95 C 32940 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-32 REV. 11, JANUARY 2005 TABLE B3.6-31 BREAK DATA, REACTOR RECIRCULATION (RR)

INSIDE CONTAINMENT*

STRESS RATIO PER ASME EQNS.

BREAK IDENT. EQ(10) S 2.4 S m EQ(12) S 2.4 S m EQ(13) S 2.4 S m USAGE FACTOR BREAK TYPE*** BREAK BASES SECTION NO. RS1 0.654 0.230 0.464 0.0 C 3.6.2.1.6.1.a RS3** LL 1.676 0.911 0.619 0.30 L 3.6.2.1.6.1.b RD1 0.620 0.351 0.409 0.0 C 3.6.2.1.6.1.a RD2 0.796 0.185 0.374 0.0 C 3.6.2.1.6.1.a RD3 0.701 0.094 0.394 0.0 C 3.6.2.1.6.1.a RD4 0.650 0.279 0.411 0.0 C 3.6.2.1.6.1.a RD5 0.747 0.506 0.376 0.0 C 3.6.2.1.6.1.a

_____________________________

• Loop A same as Loop B except as noted. ** Loop B only. Subscript "LL" indicates longitudinal break.

• • • Break type: C - circumferential, L - longitudinal CPS/USAR CHAPTER 03 B3.6-33 REV. 11, JANUARY 2005 TABLE B3.6-32 BREAK DATA, MSIV-LEAKAGE CONTROL SUBSYSTEM IS-03 OUTSIDE CONTAINMENT BREAK NUMBER TYPE* ALLOWABLE STRESS (psi) 0.8 (1.2S h + S A) CALCULATED STRESS (psi) (EQ. 9B + EQ. 10) 1S-C1 C 32400 Terminal End Break 1S-C2 C 32400 Terminal End Break 1S-C3 C 32400 Terminal End Break 1S-C4 C 32400 Terminal End Break

• Break type: C = circumferential, L = Longitudinal.

CPS/USAR CHAPTER 03 B3.6-34 REV. 11, JANUARY 2005 TABLE B3.6-33 BREAK DATA, SLCS SUBSYSTEM SC-07 OUTSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR SC-C2 C 40080 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-35 REV. 11, JANUARY 2005 TABLE B3.6-34 BREAK DATA, NUCLEAR BOILER SUBSYSTEM NB-01 INSIDE CONTAINMENT BREAK NUMBER TYPE*

2.4S m (psi) EQ.10 CALCULATED STRESS (psi)

EQ. 12 EQ. 13 CUMULATIVE USAGE FACTOR NB-C1 C 42480 Terminal End Break NB-C20 C 42480 Terminal End Break

• Break type: C = circumferential, L = longitudinal.

CPS/USAR CHAPTER 03 B3.6-36 REV. 11, JANUARY 2005 TABLE B3.6-35 RESULTS OF WHIP RESTRAINT ANALYSIS FOR HIGH PRESSURE CORE SPRAY INSIDE CONTAINMENT PIPING SYSTEM RESTRAINT INFORMATION*

POSTULATED BREAK ID RESTRAINT ID F imp (kips) T imp (10-3 sec.)FFINAL (kips) GAP (inches) TIP DIS-PLACEMENT inches) ACTUAL DEFLEC-TION (INCHES) PEAK DYNAMIC LOAD (kips) ALLOWABLE DEFLECTION (inches) HP-C9:C HP-R5 - - 100.95 7.3 19.25 4.42 278.5 6.29

• Restraint information is based on current analysis.

CPS/USAR CHAPTER 03 B3.6-37 REV. 11, JANUARY 2005 TABLE B3.6-36 RESULTS OF CONTAINMENT PENETRATION PIPING ANALYSIS FOR FEEDWATER INSIDE CONTAINMENTNote:

STRESS (psi)

BREAK NUMBERS RESTRAINT NUMBER (GUIDE) PEAK RESTRAINTREACTION (kips) MAXIMUM PIPE STRESS IN CONTAIN- MENT PENE- TRATION AREA ALLOWABLEFW-C32(C) FW-R16 270 & 23845 44381 FW-C32(L) FW-R16A 462 Note: Mirror image of feedwater line has the same stress and load results CPS/USAR CHAPTER 03 B3.6-38 REV. 11, JANUARY 2005 TABLE B3.6-37 BREAK DATA, FWLCS, SUBSYSTEMS RH-85 AND RH-86 OUTSIDE CONTAINMENT RH-85 BREAK NUMBER BREAK TYPE* ALLOWABLE STRESS (psi) 0.8(1.2S h + S a) CALCULATED STRESS (psi) (EQ. 9B + EQ. 10) RH-85-1 C N/A Terminal End RH-85-2 C N/A Terminal End RH-86 NO BREAKS POSTULATED AS THIS PORTION OF HIGH ENERGY PIPING IS IN THE BREAK EXCLUSION AREA.

• BREAK TYPE: C = circumferential, L = longitudinal

CPS/USAR CHAPTER 03 C3.6-1 REV. 11, JANUARY 2005 ATTACHMENT C3.6 EVALUATION OF ESSENTIAL COMPONENTS UNDER DYNAMIC EFFECTS OF JET IMPINGEMENT This attachment has been deleted. The evaluation of essential components under the effects of jet impingement is covered in Section 3.6, which discusses jet forces and geometries. Attachment B3.6 lists the break locations in high energy piping. Attachment D3.6 discusses piping layout.

CPS/USAR ATTACHMENT D3.6

SUMMARY

OF FAILURE MODE ANALYSIS FOR PIPE BREAKS AND CRACKS

CPS/USAR CHAPTER 03 D3.6-1 REV. 11, JANUARY 2005 ATTACHMENT D3.6

SUMMARY

OF FAILURE MODE ANALYSIS FOR PIPE BREAKS AND CRACKS D3.6.1 GENERAL This Attachment describes the specific pipe failure protection provided to satisfy the requirements of Subsection 3.6.1 and demonstrates that essential systems, components and equipment are not adversely affected by pipe breaks or cracks. The information is divided into three categories: (1) a discussion of high-energy pipe breaks and the dynamic effects of pipe whip and jet impingement (Subsection D3.6.2); (2) a discussion of moderate-energy pipe cracks and the effects of spraying or wetting (Subsection D3.6.3); and (3) a discussion of flooding as a result of breaks or cracks (Subsection D3.6.4). The primary method used throughout the plant to protect essential systems, components and equipment was physical separation. In order to effect physical separation of safety systems, certain generalized procedures were followed during the design stages of the project. All piping, mechanical equipment, electrical components and instrumentation for each of the safety systems in each of the divisions (3 Divisions of Engineered Safety Features and 4 Divisions of Electrical Components) were numbered to indicate the division in which they belong. In the design of the systems and in the layout of the general arrangements, all equipment and components in each of the safety divisions were marked with different colors so that it was easily determined from looking at drawings that they have been given separation from equipment and components in the other divisions. This technique facilitated the location of systems not in the safety division which need be checked for their impact on safety systems. Improper interconnections between safety divisions also were located by this mechanism. For those safety-related systems located outside the containment, physical separation was the primary mode of protection in all but a few isolated cases. These cases are described in the following subsections. In general, the following design techniques applicable to the equipment in each safety division were applied throughout the plant: a. The equipment within a single safety division was maintained together physically throughout the plant. b. The distance to be maintained between equipment within different safety divisions precluded disabling or degrading of more than one nuclear safety-

related division from a single event. c. Non-nuclear safety-related equipment which contains high-energy pipelines or presents the potential for plant flooding (see Section 3.4) was located to preclude the disabling or degrading of more than one nuclear safety-related division by a single event. d. In areas where adequate distance could not be maintained between two or more safety-related divisions or between high-energy non-divisional safety-related equipment and a nuclear safety related division, a failure mode analysis approach was taken to determine that the safety-related equipment involved was CPS/USAR CHAPTER 03 D3.6-2 REV. 11, JANUARY 2005 not adversely affected by pipe breaks or cracks in the immediate area. If the analysis indicated that a nuclear safety related system required protection, barriers, i.e., cubicles, walls, and tunnels were employed (see e. below). e. In areas where cubicles or wall-type barriers are impractical, restraints were employed between high energy lines and the nuclear safety-related equipment. The implementation of separation barriers was as follows: (1) distance separation; and (2) general or area barriers, such as rooms and walls (see Figures 3.5-3 through 3.5-5, Missile-Proof Walls). Color-coded piping and instrument diagrams and color-coded composite diagrams were used to ensure that the routing of high-energy lines throughout the plant did not adversely affect essential systems, components and equipment. It should be noted here that the pipe whip analysis was performed for high-energy lines as discussed in Section D3.6.2. Figures D3.6-1 through D3.6-132 show only the high-energy piping (cross-hatched) and divisional piping, equipment, ductwork and instrument lines. Two inch and smaller piping and instrument tubing or electrical conduit which is presently being routed will be protected in accordance with the requirements of Subsection 3.6.1. For a further description of pipe/crack locations and types, break exclusion areas (no-break zones, that is, areas where a pipe break is not postulated due to meeting the criteria of the Standard Review Plan (SRP) and Branch Technical Position (BTP) MEB 3-1, guard pipes, and pipe whip restraints, refer to Subsection 3.6.2. Attachment B3.6 identifies all pipe whip restraints and associated break locations on isometric drawings. High-energy lines 8-inch nominal diameter and larger in the areas of the containment drywell are restrained; therefore, the dynamic effects of pipe whip are minimal on essential

components. The most limiting problem in these areas is jet impingement. To protect from jet impingement, essential components were separated from other divisions and from high energy lines. D3.6.2 HIGH-ENERGY PIPING High-energy fluid systems are considered to be pipe rupture initiating systems. These systems are listed in Table 3.6-2. Several of the high-energy systems are located in areas or buildings which house no safety-related systems, equipment or components. Therefore, these high-energy systems cannot impact on safety-related equipment and were not analyzed. Those systems are as follows: a. extraction steam, b. condensate, c. condensate booster,

d. heater drains, CPS/USAR CHAPTER 03 D3.6-3 REV. 11, JANUARY 2005 e. miscellaneous vents and drains, f. turbine drains,

g. turbine gland steam seal steam,

h. radwaste chemical waste process, i. chemical radwaste reprocessing and disposal, j. radwaste sludge process, and

k. auxiliary steam. The remaining high-energy fluid systems are described in the following subsections. Appropriate isometric drawings with break locations and restraints are shown in Attachment B3.6 and Figures D3.6-1 through D3.6-133. These figures depict areas both inside and outside the containment. For this analysis, the movement of the restrained piping (tip displacement) was calculated using the PWRRA program as described in Subsection 3.6.2.3.3. D3.6.2.1 Main Steam Piping The location of the postulated pipe breaks and the pipe whip restraints for the main steam system is shown on Figures B3.6-6 through B3.6-13. The stress analysis used for the main steam system is summarized in Tables B3.6-6 through B3.6-13. D3.6.2.1.1 General Each of the four 24-inch main steamlines is welded to the appropriate reactor nozzle at elevation 797 feet-1/2 inch. This is approximately 7 feet above the top of the shield wall. After an elbow, the pipe is routed downward to elevation 771 feet, then horizontally around the reactor to the area between azimuthal angles 341

° and 19°, where all four main steamlines then pass through their respective inboard MSIV's and the drywell wall penetrations. The main steamlines then pass through the containment steam tunnel, inside guardpipes, exiting the north wall of the containment and entering the auxiliary building main steam tunnel. Within the auxiliary building main steam tunnel, the main steamline passes through the out board MSIV, a third safety-related isolation valve, and runs horizontally through the steam tunnel into the turbine building. No breaks were postulated from the inboard main steam isolation valve and extending outside containment beyond the third isolation valve through the north wall of the auxiliary building steam tunnel up to the first elbow fitting from the isolation valve (Figures B3.6-6 through B3.6-13). The piping from the reactor vessel through the second isolation valve satisfies all the requirements of ASME Code, Section III, Class 1, Quality Group A. Between the second and third isolation valves, the piping is Class 2, Quality Group B. After the third isolation valve, the piping complies with ANSI Standard B31.1, Quality Group D. A total of 16 safety/relief valves are mounted on the horizontal runs between the reactor and the first isolation valve inside the drywell. The discharge piping and vent lines from these CPS/USAR CHAPTER 03 D3.6-4 REV. 11, JANUARY 2005 safety/relief valves are normally unpressurized; therefore, there is no potential for dynamic pipe whip or similar hazards. In addition, an 8-inch line branching from main steamline A supplies steam to the RCIC turbine. This line, which passes through the containment and auxiliary building steam tunnels, is discussed in the analysis of the RCIC System in Subsection D3.6.2.10. D3.6.2.1.2 Inside Drywell The dynamic load of a nonrestrained whipping main steam pipe could impact several systems: the 8-inch RCIC, the 12-inch LPCI, the 10-inch LPCS, the 10-inch HPCS, the CRD hydraulic

system, and the ADS relief valve system. A ruptured main steamline would rapidly depressurize the reactor as discussed in Chapters 6 and 15, therefore, the RCIC, the HPCS and the ADS systems would be unnecessary in mitigating the consequences of a main steamline break. To preclude any likelihood of loss of a system required for safe plant shutdown, pipe restraints and guides have been installed inside the drywell as shown on the isometric and the composite drawings (see Attachment B3.6 Figures). The environmental conditions are the same as the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). The effects of jet impingement from breaks in the main steam piping have been evaluated. In the event of a postulated main steam line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand jet force. D3.6.2.1.3 Inside the Containment Steam Tunnel The main steam piping from the deflection-limiting restraint inboard of the inboard isolation valve, through the steam tunnel in the containment, has been qualified as a no-break zone.

This piping is inside guard pipes, consequently, no analysis of failure modes was performed. D3.6.2.1.4 Inside the Auxiliary Building Steam Tunnel The main steam piping in the auxiliary building steam tunnel has no breaks from the containment wall and extending outside containment beyond the third isolation valve and through the north wall of the auxiliary building steam tunnel into the turbine building. Consequently, no analysis of failure modes was performed. Due to the location of the postulated main steamline break and the location and design of the main steam and feedwater guide structures, jets from a break in the main steamline do not impact on any equipment required to mitigate the consequences of the main steamline break. D3.6.2.1.5 Inside the Turbine Building After the no-break zone, the piping is not seismically qualified, consequently breaks have been postulated using the fitting criteria according to the requirements of Subsection 3.6.2. Because there is no essential equipment in the turbine building and because of the installation of the third CPS/USAR CHAPTER 03 D3.6-5 REV. 11, JANUARY 2005 main steam isolation valve, there are no areas of concern for dynamic effects of pipe break in the turbine building. D3.6.2.2 Feedwater System Piping The location of the postulated pipe breaks and the pipe whip restraints for the feedwater system is shown on Figures B3.6-1 through B3.6-3. The stress analysis used for the feedwater system is summarized in Tables B3.6-1 through B3.6-3. The essential equipment which is required for mitigating the consequences of a feedwater line break, and which may be hit by a jet from a break in the feedwater line has been evaluated for the jet impingement effects from postulated feedwater breaks. In the event of a postulated feedwater line break, any equipment which is required for the break has sufficient redundant equipment not hit by the jet or is sufficiently separated from the break so that plant safety is not affected. D3.6.2.2.1 General From turbine building, each of the two 20-inch feedwater lines passes through the north wall of the auxiliary building into the auxiliary building steam tunnel. Inside the auxiliary building steam tunnel, feedwater lines pass horizontally through the tunnel, through the motor-operated isolation valves, and through the air-assisted check valves into the containment building. Once inside the containment building steam tunnel, the feedwater lines, enclosed in guard pipes, pass through the tunnel into the drywell, through a check valve and a manual maintenance valve.

The line then splits into two 12-inch risers which terminate at elevation 784 feet-3-1/2 inches. At this elevation termination, each feedwater line passes through the shield wall and connects to a reactor nozzle. The only other high-energy lines connecting to the feedwater lines are the 10-inch residual heat removal lines which connect inside the auxiliary building steam tunnel to the feedwater lines between the motor-operated isolation valve and the air-assisted check valves and the RHR branch lines for the FWLC mode of RHR. Refer to D3.6.2.7.6 for further description of FWLC piping. No breaks were postulated in the feedwater system piping extending from the check valve in the drywell through the containment steam tunnel, through the auxiliary building steam tunnel up to the turbine building. Breaks were postulated only inside the turbine building steam tunnel after the first elbow fitting and inboard of the check valve in the drywell. As previously mentioned, the only other high-energy lines analyzed as part of the feedwater system are the residual heat removal 10-inch lines which connect to the feedwater lines in the auxiliary building steam tunnel between the motor-operated isolation valves and the air-assisted check valves. From the results of the pipe rupture analysis, these two 10-inch RHR lines are postulated to have no break from the feedwater line up to the auxiliary building steam tunnel wall, except that terminal-end breaks are postulated at the auxiliary building steam tunnel wall. D3.6.2.2.2 Inside the Drywell Pipe breaks were postulated at points of high stress or usage factor and at each feedwater connection to the reactor nozzles. To ensure that no unacceptable damage could result, restraints were installed to prevent pipe movement from damaging nearby safety-related systems, specifically the ADS system and the 10 inch RCIC steam piping which branches off the main steamline. Spray from a break inside the shield wall at the reactor pressure vessel connections is localized in the vicinity of the nozzles and poses no safety hazard. The CPS/USAR CHAPTER 03 D3.6-6 REV. 12, JANUARY 2007 pressurization of the annulus between the RPV and its shield wall is discussed in Subsection 6.2.1.2.3.2. For the location of these breaks, see B3.6 Figures. Components which are in close proximity to the feedwater system risers are the ADS valves and their discharge lines, the 12-inch RHR (LPCI) injection lines Division 2, and the LPCS and HPCS injection lines. The breaks, which were postulated in the horizontal runs of the feedwater piping (outboard of the check valve and before the risers), could endanger the following safety-related lines: the 8-inch RCIC steamline, the 12-inch RHR (LPCI) injection line Division 1, and the combustible gas control system discharge lines. To preclude any likelihood of loss of a system required for safe plant shutdown, restraints and guides have been installed inside the drywell as shown on the isometric drawings. The break of a feedwater line inside the drywell will create conditions no worse than those following a LOCA. All Class 1E electrical equipment inside the drywell whose operation, during or after a LOCA, is required for safe shutdown is qualified for the post-LOCA drywell environment as discussed in Section 3.11. D3.6.2.2.3 Containment Steam Tunnel All feedwater piping inside the containment steam tunnel is within the boundaries of the no-break zone and enclosed in the previously mentioned guardpipes. Consequently, a piping failure is not postulated to occur in any of these lines. Guides and restraints are located to minimize the effects of pipe movement of the feedwater pipe within the guardpipes if a break occurs outside the no-break area. The guides and restraints limit any movement to an acceptable level such that damage does not occur to either the guardpipe or any other piping in the containment steam tunnel. D3.6.2.2.4 Outside Containment For the feedwater piping, no breaks are postulated from the containment wall through the auxiliary building steam tunnel and into the turbine building. As previously discussed, the piping in the turbine building is not seismically analyzed and breaks are thus postulated at each fitting. There is, however, no essential equipment in the turbine building. The entire run of feedwater piping in the auxiliary building steam tunnel has no postulated breaks, therefore, no analysis of failure modes was performed. Terminal end breaks were postulated for the 10-inch residual heat removal lines analyzed as part of the feedwater system. These postulated breaks are inside the auxiliary building steam tunnel, at the wall, adjacent to each of the residual heat removal heat exchanger rooms. The effects of jet impingement and pipe whip from these breaks has been evaluated. Due to the remote location of these breaks, jets and pipe whips would not impact sufficient redundant equipment to prevent a safe shutdown of the reactor. Pressurization of the auxiliary building steam tunnel is described in Section 3.6.1, and is based on a bounding simultaneous rupture of a main steam line and a feedwater line. The reactor water cleanup lines are discussed in Section D3.6.2.8 of this

attachment. D3.6.2.3 Main Steam Isolation Valve Leakage Control System (MSIV-LCS)

Note: As a result of the re-analysis of the Loss of Coolant Accident (LOCA) using Alternative Source Term (AST) Methodology, it is no longer necessary to credit the Main Steam Isolation Valve Leakage Control System (MSIVLCS) for post-LOCA activity leakage mitigation. The system has been left in place as a passive system and is not required to perform any safety function.The composite drawings which show the piping for the main steam isolation valve leakage control system are Figures D3.6-8 and D3.6-14. The piping shown on Figure D3.6-8 is not high energy, as the isolation valves terminate the high-energy portion within the auxiliary building steam tunnel.

CPS/USAR CHAPTER 03 D3.6-7 REV. 11, JANUARY 2005 D3.6.2.3.1 General The high-energy portion of the MSIV-LCS, shown on Figure D3.6-14, is that between the main steam isolation valve drain line and the isolation valves for the MSIV-LCS located just above the auxiliary building steam tunnel floor. For the inboard MSIV-LCS, these high-energy lines consist of four 1 1/2-inch lines. For the outboard system, these high-energy lines consist of four 2-inch lines terminating in a 2 1/2-inch header which is isolated inside the main steam tunnel by normally closed motor-operated isolation valves. The size of the high-energy lines in the MSIV-LCS precludes the likelihood of

their damaging any other safety-related systems in the near vicinity. Jets from postulated breaks in MSIV-LCS lines do not load any equipment required to mitigate the consequences of a break in the MSIV-LCS or affect the ability to achieve safe shutdown. The environmental conditions associated with the breaks are the same as the local environment in the auxiliary building steam tunnel. All Class 1E electrical equipment in the steam tunnel has been qualified (refer to Section 3.11 for environmental qualification). D3.6.2.4 Reactor Recirculation System Reactor recirculation system Figures B3.6-29 through B3.6-31, show the locations of the postulated pipe breaks and pipe whip restraints. The stress analysis used for the reactor recirculation system is summarized in Tables B3.6-29 through B3.6-31. The piping in this system was analyzed for pipe break and pipe restraint locations by General Electric Company. The effects of jet impingement from breaks in the reactor recirculation piping have been evaluated. In the event of a postulated reactor recirculation line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet forces. D3.6.2.4.1 General Each of the two reactor recirculation loops leaves the reactor pressure vessel at elevation 757 feet 10-1/2 inches at azimuthal angles 0 and 180 for Loops B and A, respectively. Each 20-inch suction line then drops vertically to elevation 726 the inlet isolation valve, the reactor recirculation pump, the flow control valve and the outlet isolation valve. At this point, the discharge line runs vertically up to elevation 744 feet 7-1/2 inches, where it joins the C-shaped 16-inch horizontal header. From this header, five 10-inch vertical lines run up to elevation 758 feet 3 1/2 inches, where they turn and horizontally enter the reactor pressure vessel. In addition, from Loop B only, an 18-inch RHR suction line branches off from the vertical run between the reactor outlet and the inlet isolation valve at elevation 733 feet 6 inches. This line is considered high energy up to the isolation valve; however, it is not a part of the ECCS systems. The line passes horizontally from the connection approximately 4 feet, then turns vertically upward and rises through an isolation valve to elevation 757 feet 6 inches, where it turns and runs horizontally out of the north wall of the drywell through the containment building steam tunnel enclosed in a guardpipe and into the auxiliary building steam tunnel. This line is further discussed in Subsection D3.6.2.7.

CPS/USAR CHAPTER 03 D3.6-8 REV. 11, JANUARY 2005 D3.6.2.4.2 Recirculation Loop "A" The GE analysis resulted in the postulated circumferential and longitudinal breaks shown in the Figure B3.6-31. Restraints were installed on each riser as shown on Figure B3.6-31 to limit the travel of potential ruptured piping in the direction of the RPV radius. Restraints were also installed on the 16 inch header to limit travel. In addition, one restraint was installed near the RPV nozzle of the suction line. D3.6.2.4.3 Recirculation Loop "B" The postulated break locations and resulting restraints are the same for both LOOPS A and B, with the exception that Loop B has the RHR connection (see Subsection D3.6.2.7.3). This connection results in the postulation of another longitudinal failure at the tee joint. To protect against this postulated failure, two additional restraints are installed, one just above and one just below the tee. D3.6.2.4.4 Inside the Drywell All the piping associated with the reactor recirculation system is contained within the drywell except for the RHR suction line tapping off Loop B.

The risers associated with the reactor recirculation piping could conceivably damage the RHR injection lines (LPCI) and the RCIC steamline (from the main steamline to the RCIC turbine). The RHR (LPCI) lines are protected from pipe whip by restraints and guides and are redundant in function. The RCIC steamline could be impacted by a rupture of the reactor recirculation piping. However, it is protected against the dynamic effects of pipe break by restraints and guides. In addition, it is redundant in function to the high-pressure core spray system. As noted above, restraints have been installed to preclude any likelihood of the loss of a system required for safe plant shutdown. The environmental conditions are the same as the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). D3.6.2.5 Low-Pressure Core Spray (LPCS)

The low-pressure core spray system, Figure B3.6-5, show the locations of the postulated pipe breaks and of the pipe whip restraints. The stress analysis used for the low-pressure core spray system is summarized in Table B3.6-5. The effects of jet impingement from breaks in the low pressure core spray piping have been evaluated. In the event of a postulated low pressure core spray line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet force. D3.6.2.5.1 General The portion of the LPCS system which is considered to be high energy is that piping between the reactor nozzle and the inboard isolation check valve. This system does not operate during normal plant operation; consequently, only that part of the piping which is normally exposed to reactor pressure is classified as high-energy.

CPS/USAR CHAPTER 03 D3.6-9 REV. 11, JANUARY 2005 The high-energy portion of the LPCS piping begins at the reactor nozzle at elevation 782 feet 9 inches at azimuthal angle 90

°. The line passes through the shield wall penetration and drops to elevation 769 feet 5 1/4 inches, where it runs horizontally, turning at a 90

° angle and passing through the locked-open maintenance valve. The line then passes through the inboard isolation valve, which is a check valve, and passes out of the drywell. No breaks are postulated after the isolation valve. D3.6.2.5.2 Inside the Drywell Systems which are in the vicinity of the LPCS and could be impacted by the dynamic effects of a pipe break or crack are portions of the ADS system including electrical conduits, operators and accumulators, the ADS system discharge line and the RCIC steamline. To ensure that no movement of the vertical leg can impact these systems, the vertical leg has been restrained in

the event of a pipe break. The environmental conditions are the same as the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). D3.6.2.6 High-Pressure Core Spray (HPCS)

HPCS Figure B3.6-4 shows the locations of the postulated pipe breaks and pipe whip restraints. The stress analysis used for the HPCS system is summarized in Table B3.6-4. D3.6.2.6.1 General The portion of the HPCS system which is considered to be high energy is that piping between the reactor nozzle and the inboard isolation check valve. This system does not operate during normal plant operation; consequently, only that part of the piping which is normally exposed to reactor pressure is classified as high-energy. The high-energy portion of the HPCS system begins at the reactor nozzle at elevation 767 feet 5 1/4 inches and azimuthal angle 270

°. The line passes through the shield wall penetration, turns vertically downward and runs to elevation 771 feet 11 3/8 inches, where it passes through one pipe bend. From here it runs horizontally through a 90

° elbow at elevation 769 feet 5 1/4 inches, turns 90

° again passing through the locked open manual maintenance valve and through the inboard isolation check valve. It then makes two 75

° turns and passes horizontally out of the drywell at elevation 769 feet 5 1/4 inches. D3.6.2.6.2 Inside the Drywell There is no essential equipment in the immediate vicinity of the HPCS piping reactor nozzle or vertical riser until the line reaches elevation 769 feet 5 1/4 inches. A break in the HPCS system piping at this elevation could impact the following systems: the ADS system discharge line, the combustible gas control system compressor discharge line and the inboard isolation valves for the drywell cooling system. To preclude the likelihood of loss of any of these systems, the piping was restrained as shown on the isometric and composite drawings. The environmental conditions are the same as for the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification).

CPS/USAR CHAPTER 03 D3.6-10 REV. 11, JANUARY 2005 The effects of jet impingement from breaks in the high pressure core spray piping have been evaluated. In the event of a postulated high pressure core spray line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet force. D3.6.2.7 Residual Heat Removal System (RHR)

The RHR system Figures B3.6-14 through B3.6-19, B3.6-21, and B3.6-35 show the locations of the postulated pipe breaks and of the pipe whip restraints. The stress analysis used for the RHR system is summarized in Tables B3.6-14 through B3.6-19, B3.6-21, and B3.6-37. D3.6.2.7.1 General The piping for the RHR system is divided into three (3) parts as follows: a. the LPCI injection lines (RHR A Loop, RHR B Loop and RHR C Loop), b. the shutdown cooling suction line from reactor recirculation Loop B.

c. the feedwater leakage control mode (FWLC) lines to main feedwater. Each of the above is discussed in the following subsections.

D3.6.2.7.2 Low-Pressure Coolant Injection (LPCI)

The LPCI subsystems are not in use during normal plant operation; consequently, only the piping between the reactor pressure vessel and the first normally closed valve is pressurized and thus classified as high-energy for pipe rupture analysis. D3.6.2.7.2.1 LPCI "A" The high-energy portion of the LPCI A Loop piping begins at the reactor nozzle at elevation 778 feet 3 1/2 inches at azimuthal angle 45

°. The line passes through the shield wall penetration and drops to elevation 761 feet 3 1/2 inches. At this elevation the piping turns and passes horizontally through the locked open manual maintenance valve, the inboard isolation check valve and exits the drywell. The only safety-related systems which are in the vicinity of this loop and could be impacted by the dynamic effects of a pipe break are portions of the ADS system including the operators and accumulators and the ADS system discharge lines. To ensure that no movement of the vertical or horizontal legs can occur, restraints have been installed as shown in Figure B3.6-14. The effects of jet impingement from breaks in the LPCI-"A" piping have been evaluated. In the event of a postulated LPCI-"A" line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet force. D3.6.2.7.2.2 LPCI "B" The high-energy portion of the LPCI B Loop, Figure B3.6-15, begins at the reactor nozzle at elevation 778 feet 3-1/4 inches at azimuthal angle 225. The line passes through the shield wall penetration and down through the locked-open manual maintenance valve and the inboard CPS/USAR CHAPTER 03 D3.6-11 REV. 11, JANUARY 2005 isolation check valve. Following the check valve, the line exits the drywell. No breaks are postulated after the isolation valve. There ar e no safety related systems in the immediate vicinity of the LPCI B Loop that could be impacted by the dynamic effects of pipe break in the high energy portion of the piping. The ADS discharge lines and the isolation valves and associated electrical conduits for the drywell cooling system will not be impacted by a break in the LPCI B Loop. Although these systems are located in this quadrant of the containment and are in the near vicinity of the LPCI B Loop piping, they are not in the vicinity of the high-energy portion. The restraints for breaks located on the high-energy portion of the piping will prevent any damage to these systems in the event of a break. The effects on jet impingement from breaks in the LPCI-"B" piping have been evaluated. In the event of a postulated LPCI-"B" line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet force. D3.6.2.7.2.3 LPCI "C" The high-energy portion of the LPCI C Loop Figure B3.6-16 begins at the reactor nozzle at elevation 778 feet 3-1/4 inches at azimuthal angle 135

°. The line passes through the shield wall and drops through the manual maintenance valve to elevation 769 feet 5 inches and through the inboard isolation check valve. Following the isolation valve, it drops down to elevation 764 feet 1/2 inch, where it passes out of the drywell. No breaks are postulated after the isolation valve.

Systems which are in the vicinity of the LPCI C Loop and could be impacted by the dynamic effects of a pipe break are portions of the ADS system discharge piping, the LPCS injection line and the drywell purge system isolation valve. To ensure that no movement of the high-pressure portion of the LPCI C Loop piping can impact these systems, restraints have been installed to

prevent pipe whip. The effects of jet impingement from breaks in the LPCI-"C" piping have been evaluated. In the event of a postulated LPCI-"C" line break, any equipment hit by the break and required for safe shutdown has sufficient redundant equipment not hit by the jet, or is sufficiently separated from the break so that the equipment can withstand the jet force. D3.6.2.7.3 RHR Suction from the Reactor Recirculation Loop B For the shutdown cooling mode of the RHR system , Figure B3.6-17, suction is taken from reactor recirculation Loop B. This suction line begins at the tee at elevation 733 feet 6 inches and azimuthal angle 0 of the reactor recirculation loop. The line runs vertically upward through the manual maintenance valve and through a motor-operated isolation valve. After the isolation valve, the line continues upward until at Elevation 757 feet 6 inches it turns 90

° and passes out of the drywell, through a guardpipe and out of the containment. The high energy portion of this suction line is a short L-shaped run extending about 4-1/2 feet horizontally and about 14-1/2 feet vertically. Breaks were postulated as shown in Figure B3.6-17. The only system which is in the near vicinity of this suction line is the RCIC steamline. To ensure that no movement of the vertical leg of the suction line could impact this system, restraints were located to prevent pipe movement. D3.6.2.7.4 RCIC Steamlines to RHR Heat Exchangers Out of Service CPS/USAR CHAPTER 03 D3.6-12 REV. 11, JANUARY 2005 D3.6.2.7.5 RHR Heat Exchanger Rooms RHR Heat Exchanger Rooms A and B have been designed to withstand jet impingement and pipe whip effects of postulated breaks in the high energy RHR piping in these rooms. However, postulation of such line breaks in the RHR Heat Exchanger Rooms A and B is not required (refer to NRC letter dated January 7, 1986, Docket No. 50-461). D3.6.2.7.6 RHR Feedwater Leakage Control Mode (FWLC) Piping to Main Feedwater The FWLC piping of the RHR system is not in use during normal operation. The portions of piping from the RHR pump rooms to the check valves in the auxiliary building steam tunnel are classified as moderate - energy piping per Section 3.6.2.1.4.d. Consequently, only the piping between the main feedwater headers and the first c heck valve in each line is classified as high enery piping. The two portions of piping (3/4 inch) connected to the feedwater header down stream of valves 1B21-F032A/B are considered to be within the break exclusion boundary for piping between containment isolation valves. Therefore no pipe breaks are postulated in these sections of piping. The two portions of piping (2 inch) connected to the feedwater header up stream of valves 1B21-F032A/B are postulated to have terminal end breaks at the connection to the main feedwater header. The pipe whip and jet impingement resulting from these breaks have been evaluated and do not affect any adjac ent components or equipment in the steam tunnel. D3.6.2.8 Reactor Water Cleanup System (RWCU)

The reactor water cleanup system isometric drawings, Figures B3.6-23 through B3.6-30, show the location of the whip restraints and associated postulated pipe breaks. The stress analysis used for the reactor water cleanup system is summarized in Tables B3.6-23 through B3.6-30. D3.6.2.8.1 General For many purposes, the reactor water cleanup system can be considered as two nearly separated subsystems: (1) the loop from the reactor, through the heat exchangers, through the filter demineralizers, back through the regenerative heat exchangers, and return to the reactor through the feedwater line; and (2) the auxiliary subsystem which removes the used demineralizer resin and replaces it with new resin. The first subsystem is almost entirely classified as a high-energy system; the second subs ystem, even during the small time it is in operation, is classified as moderate energy. Neither is required for a safe plant shutdown.

From the standpoint of piping failure, the only concern is the possible detrimental effect of a RWCU pipe rupture on other equipment required for safe shutdown. One 4-inch line taps off of the bottom of each reactor recirculation loop at azimuthal angles 155

° and 335°. Each line runs axially out from the reactor at elevation 724 feet 8 inches. Each line passes through a motor-operated isolation valve, turns and runs vertically upward to elevation 732 feet. At this point, both lines turn and circle towards one another around the inside of the weir wall, meeting at a tee at an azimuthal angle of approximately 58

°. Two feet before the tee, both lines increase in diameter to 6 inches and, from the tee, the combined suction 6-inch line turns and runs back to azimuthal angle 19.6. At this point, the pump suction line drops horizontally to elevation 725 feet 4 inches, turns, and runs through a motor-operated isolation valve. Immediately after the motor-operated isolation valve, the line tees with the pump suction line from the bottom of the reactor vessel. The pump CPS/USAR CHAPTER 03 D3.6-13 REV. 11, JANUARY 2005 suction line from the bottom of the reactor vessel is routed from elevation 742 feet 6 inches at zimuthal angle 210 northward to where it drops to 725 feet 4 inches and joins with the pump suction line coming from the reactor recirculation lines. For the actual routing of this reactor drain line, see Figure B3.6-23. One line branches off of the pump suction line from the bottom of the reactor vessel. It contains a normally closed motor-operated bypass line used in the hot standby mode. Another branch from the bypass is a 2-inch reactor drain line which runs to a sump in the drywell. The combined pump suction line (consisting of the combined line from the reactor recirculation loops and the line from the reactor pressure vessel) runs vertically upward from elevation 725 feet 4 inches to elevation 756 feet 5 inches. There the line turns and runs horizontally through the inboard motor-operated containment isolation valve. It then exits the drywell, passing through the containment inside the containment steam tunnel enclosed in a guardpipe. Upon exiting the containment, the line passes through the outboard containment motor-operated isolation valve, turnsand is routed around the outside of the containment wall to the reactor water clean up pump room cubicles. These are located on elevation 737 feet in the area bounded by column rows 117-123 and column rows Z-AB. The pump discharge line then returns to the auxiliary building steam tunnel (along the same general routing), where it passes through the outboard motor-operated containment isolation valve into the containment building steam tunnel. Then the line passes through the inboard containment isolation valve and runs up to the floor above the steam tunnel where the reactor water cleanup heat exchangers are

located. Discharge lines from heat exchangers join together in one line which goes through the steam tunnel and then splits into two lines each terminating in separate RWCU filter demineralizers. They are located in cubicles at azimuth 270 and elevation 803 feet 3 inches. Discharge lines from filter demineralizers go back to the steam tunnel in the containment building and both are connected to one pipe. This pipe passes through the inboard containment isolation valve and out of the containment by way of the outboard containment isolation valve.

At this point, the line is no longer considered high-energy; however, it does continue northward out of the auxiliary building steam tunnel, into the turbine building and goes to the main condenser. The elevation at which this penetration exits the containment is 762 feet 3 inches. The second line returns reactor water from the reactor water cleanup system to the feedwater system. Headers from the heat exchanger drop vertically downward into the containment building steam tunnel, along the same path as the reactor water cleanup line to the condenser.

At elevation 763 feet 8 7/8 inches, the line passes through the inboard containment isolation valve and through the outboard containment isolation valve, where it tees into two branch lines, each routed to one of the feedwater lines. One leg of the tee runs westward through a motor-

operated isolation valve and becomes the RHR line (at elevation 763 feet 8-7/8 inches) which terminates in the feedwater line. The other branch of the tee runs across the steam tunnel, where it turns southward through a motor-operated valve and becomes the RHR line which terminates in the second feedwater line. D3.6.2.8.2 Inside the Drywell In the pump suction lines, circumferential breaks have been postulated as shown in Attachment B3.6. The dynamic load of a ruptured RWCU line could impact several systems:

CPS/USAR CHAPTER 03 D3.6-14 REV. 11, JANUARY 2005 the RHR system (suction line for shutdown c ooling mode), the main steamlines, and the feedwater lines. The RHR line itself, an 18-inch line, will not be affected by the rupture of a 6-inch RWCU line. Also in this same area are the main steam and feedwater lines. Both are of such size that they would not be affected by the rupture of the 6-inch RWCU line. The environmental conditions in the drywell are the same as the local environment. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). The effects of jet impingement and pipe whip from postulated breaks in the RWCU lines have been evaluated. Due to the small size and remote location of the RWCU lines, jets and pipe whip from any RWCU line break do not hit sufficient redundant equipment to prevent safe shutdown of the reactor. D3.6.2.8.3 Inside and Outside Containment The RWCU piping is located in the following areas:

1. Auxiliary Building Main Steam Tunnel;

2. Auxiliary Building RWCU Pump Cubicles; 3. Auxiliary Building Radwaste Pipe Tunnel; 4. Containment RWCU Cubicles and Tunnels;

5. Containment Main Steam Tunnel.

The effects of jet impingement and pipe whip from postulated breaks in the RWCU lines have been evaluated. Due to the small size and remote location of the RWCU lines, jets and pipe whip from any RWCU line break do not hit sufficient redundant equipment to prevent safe shutdown of the reactor. D3.6.2.9 Standby Liquid Control System The postulated pipe breaks for the standby liquid control system are shown in Figure B3.6-33.

D3.6.2.9.1 General The high-energy portion of the standby liquid control system is that which is shown on Figure B3.6-33. The high-energy portion is between the connection to the reactor pressure vessel at elevation 742 feet 3 inches and the first check valve. The line runs from the bottom of the reactor pressure vessel at azimuthal angle 225

° outward axially approximately 1 foot, where it drops to elevation 743 feet 3 inches, turns and passes out through the shield wall at elevation 738 feet. Once outside the shield wall, the piping turns and runs vertically upward to elevation 741 feet, where it passes through a motor-operated isolation valve and a check valve. This check valve will terminate the high-energy portion of the line analyzed for pipe rupture.

CPS/USAR CHAPTER 03 D3.6-15 REV. 12, JANUARY 2007 D3.6.2.9.2 Inside the Drywell The only systems within the immediate vicinity of the standby liquid control system high-energy piping are the HPCS system, the reactor recirculation system and the control rod drive system. The 20-inch reactor recirculation line and the 10-inch HP line will be unaffected by the rupture of the 3-inch standby liquid control system line. The standby liquid control line is routed such that a break within the shield wall cannot impact the control rod drive system insert or withdrawal lines. The environmental conditions are the same as the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). The effects of jet impingement and pipe whip from potential breaks in the SLCS piping have been evaluated. Due to its routing and small size, jets and pipe whip from postulated breaks in the SLCS do not hit sufficient redundant equipment to prevent safe shutdown of the reactor. D3.6.2.10 Reactor Core Isolation Cooling System (RCIC)

The RCIC isometric drawing shown in Figures B3.6-20 through B3.6-22 show the postulated pipe breaks in this system. The stress analysis used for the reactor core isolation cooling system is summarized in Tables B3.6-20 through B3.6-21 and for subsystem RI-11, the stress analysis is summarized in calculation EMB-049168.

D3.6.2.10.1 General Two portions of the RCIC system piping are considered to be high-energy: (1) the steamline to the RCIC turbine, and (2) the reactor pressure vessel head spray line. An 8-inch branch line from main steamline A delivers steam to drive the RCIC turbine. The 8-inch line branches off main steamline A at elevation 788 feet 3 3/4 inches. The line circles around the outside of the shield wall to azimuthal angle 0

°, where it drops vertically downward to elevation 761 feet 5 inches. After two turns, the line passes through the normally open inboard

containment isolation valve and drops to elevation 758 feet 4 3/8 inches, where it passes horizontally out of the drywell and through the containment building steam tunnel. The line exits the containment building steam tunnel and passes through the outboard motor-operated

containment isolation valve in the auxiliary building steam tunnel. After the outboard containment isolation valve, the line runs northward approximately 19 feet. The original line reduces in size from 8 inches to 4 inches. The 4-inch line turns vertically downward and drops from elevation 757 feet 6 1/2 inches through the auxiliary building steam tunnel floor to elevation 730 feet, where it goes into the RCIC turbine cubicle. Jets and pipe whip from postulated breaks in the RCIC turbine steam supply line do not impact any equipment required to mitigate the consequences of the break. The compartment at elevation 707 feet 6 inches is the RCIC turbine cubicle. All equipment in this area is for support of the RCIC turbine. The portion of the RCIC system connected to the reactor pressure vessel head spray which is considered high-energy is that portion between the reactor pressure vessel nozzle and the first check valve. The vessel nozzle is located at elevation 813 feet 10 3/8 inches. Welded to the nozzle is a 6-inch tee. One branch of the tee (at azimuthal angle 150

°) is connected to a 4-inch pipe that is reduced to 2-inch size and connected to main steamline "A" at elevation 790 feet 6-1/2 inches. The function of this line is to vent noncondensable gas from the reactor pressure vessel head. The other branch connection located at azimuthal angle 330

°, is the inlet to the CPS/USAR CHAPTER 03 D3.6-16 REV. 12, JANUARY 2007 RPV head spray line from the RCIC pump. After the tee, this line reduces from 6 inches to 4 inches and passes through a check valve. This terminates the portion of the line considered high-energy for pipe rupture. D3.6.2.10.2 Inside the Drywell In the immediate vicinity of the steamline to the RCIC turbine are the 3-inch main steamline drains, the 12-inch LPCI A injection line, and the 12-inch feedwater line. The dynamic load of the steamline to the RCIC turbine could not affect the larger LPCI injection line or the feedwater line. However, it could impact the 3-inch main steamline drains. The main steamline drains are not required to mitigate the consequences of an accident. To limit the movement of the vertical run and thereby the moments on the inboard containment isolation valve, restraints were located on both the vertical run and the horizontal run inboard of the inboard containment isolation valve. Outboard of the containment isolation valve between the valve and the guardpipe, restraints were located to prevent unacceptable movement of that piping and conceivably impairment of operation of the isolation valve. Also, in the drywell is the RCIC discharge to the reactor pressure vessel head spray. This line is fitting-to-fitting from the reactor nozzle to the inlet of the check valve. The actual routing is nozzle to tee to reducer to back-to-back 45 degree bends to check valve. The RCIC system Figure B3.6-22 shows the location of the postulated pipe breaks and the pipe whip restraints.

The actual length of piping precludes damaging any safety-related component in this area. Of primary concern in this area is the drywell head. The drywell head is designed to accommodate jet impingement and pipe whip loads from the RCIC head spray line. The environmental conditions from either a reactor pressure vessel head spray line break or a break of a steamline to the RCIC turbine line are the same as the local environment in the drywell. All Class 1E electrical equipment in the drywell has been qualified (refer to Section 3.11 for environmental qualification). The effects of jet impingement and pipe whip from potential breaks in the RCIC piping have been evaluated. Both the RCIC steam supply and head spray lines are routed such that jets from potential breaks in these lines do not hit sufficient redundant equipment to prevent safe shutdown of the reactor. D3.6.2.11 Control Rod Drive System Breaks are not postulated in the CRD lines. As noted in Subsection 3.6.2.1.4f, CRD insert lines are exempted. Breaks are not postulated in the CRD withdrawal lines due to the small size and energy content of the lines. D3.6.3 MODERATE ENERGY PIPING Through-wall leakage cracks were postulated to occur in accordance with Subsection 3.6.2. These cracks were assumed to result in wetting of any equipment in the area whether above or below the crack and at substantial distances from the crack. In addition, the possibility of compartment flooding was considered and is discussed in Subsection D3.6.4. Each room, compartment and/or area in each seismic Category I building has been evaluated. Class 1E electrical components are not evaluated for the spray effects associated with a postulated moderate energy line crack in areas where the environmental conditions of a CPS/USAR CHAPTER 03 D3.6-17 REV. 11, JANUARY 2005 postulated high energy line break exist. The environmental conditions of a high energy line break are more severe than for a moderate energy line crack. Class 1E electrical components are qualified for the environmental conditions (steam, 100% relative humidity and condensation) resulting from a postulated high energy line break. This environmental condition (with associated pressure) envelopes the spray effects of a moderate energy line crack. In areas where the environmental conditions caused by a high energy line break do not exist, the Class 1E electrical components were conservatively assumed to become nonfunctional due to a moderate energy line crack (spray) with the exception of valve actuators, junction boxes, pullboxes, conduits, cable trays, limit switches, solenoid valves and instruments which are environmentally qualified for steam, 100% relative humidity and condensation. Each area of the auxiliary building, control building, diesel generator building, fuel building, and circulating water screenhouse were evaluated for moderate energy line cracks. These areas can be identified by referring to the general arrangement drawings cited in Chapter 1 and the composite drawings in this attachment. The drywell and certain areas of the containment are subject to high energy line breaks and were not evaluated for moderate energy line cracks.

Concurrent with the moderate energy line crack, a single failure and SSE were postulated. For some systems (e.g., RHR and shutdown service water), a single failure in the redundant system has been excluded in accordance with Subsection B.3.b(3) of BTP ASB 3-1. In all such cases the redundant system train meets the qualification of that NRC position. Loss of offsite power was assumed where a turbine generator or reactor protection system trip is a result of the postulated line crack. Using the above postulates and assumptions, an analysis was done which identified all equipment in the same room as a MELB and which may be required for plant shutdown. A single failure analysis was performed to determine if the identified equipment was required for safe shutdown or to mitigate the consequence of the cracks in question. The results of this analysis are summarized below. D3.6.3.1 Containment All equipment within the containment and drywell which must operate during or after a LOCA is qualified for the appropriate accident environmental conditions as described in Section 3.11. The wetting associated with a postulated failure of any moderate energy piping is within the bounds of that qualification. D3.6.3.2 Auxiliary Building Each compartment or significant area is discussed in detail starting at the lowest elevation, 707 feet 6 inches. Most of the essential equipment is located on the lower two levels in compartments north of the containment. D3.6.3.2.1 RHR Pump Room A RHR pump room A, including the compartment housing RHR heat exchanger A, is located on elevation 707 feet 6 inches between column rows 114, 121 and U,AA. It has been assumed that any piping failure in this room would disable all the equipment in the room. This would result in the loss of the RHR pump A and its associated support equipment, such as the instrument panel and area cooling coil cabinet. Loss of this equipment would not impair in any way the RHR B system, the RHR C system, or the reactor core isolation cooling system. All CPS/USAR CHAPTER 03 D3.6-18 REV. 11, JANUARY 2005 other ECCS would remain operable, subject to the additionally postulated single active failure. The consequences would remain within the range of events analyzed in Chapters 6 and 15. The electrical equipment is not adversely affected due to wetting caused by a MELB spray when the equipment in the room is qualified for 100% humidity, which is no worse than the wetting of the equipment due to a MELB. Also of concern in this area are several containment isolation valve assemblies. These valve assemblies are qualified to withstand a water spray and still remain functional. However, even

should any or all of these valves fail in the worst position, no significant release of radioactivity would occur. The standby gas treatment system and the auxiliary building gas control boundary would still remain functional. D3.6.3.2.2 RHR Pump Room B The RHR pump room B is located on elevation 707 feet 6 inches between column rows 105, 110, and U,AA. Because of the similarity between the arrangements of RHR A and RHR B pump rooms, the basic analysis and conclusions of RHR A (listed in the preceding subsection) apply. D3.6.3.2.3 RHR Pump Room C RHR pump room C is located on elevation 707 feet 6 inches between column rows 102, 105 and U.8,AB. All of the equipment in this room is associated with the RHR C pump, except for some leak detection and containment monitoring instrumentation. Therefore, because of the similarity of RHR pump room C and RHR pump rooms A and B, the basic analysis and conclusions drawn previously apply. D3.6.3.2.4 RCIC Pump Room The RCIC pump room is located on elevation 707 feet 6 inches between column rows 110, 114 and U,Z. It was assumed that all equipment in this room would be inoperable from the effects of pipe rupture spray. The only equipment in this room is that associated with the RCIC system, which is not required for safe shutdown or the mitigation of the effects of any piping failure in this room. Failure of the RCIC system would not prevent safe shutdown of the reactor. D3.6.3.2.5 LPCS Pump Room The LPCS pump room is located on elevation 707 feet 6 inches between column rows 121, 124 and U.8, AA. All of the equipment in this room is associated with the LPCS, except for some leak detection, containment monitoring and suppression pool monitoring instrumentation. This instrumentation, as well as all electrical circuits and all piping in this area, are primarily Division 1 except for a few Division 2 components that are required. The electrical equipment is not adversely affected due to wetting caused by a MELB spray when the equipment in the room is qualified for 100% relative humidity, which is no worse than a wetting of the equipment due to a MELB. A failure of LPCS or a failure of Division 1 is compensated for by the use of Division 2 RHR LPCI Loops B and C.

CPS/USAR CHAPTER 03 D3.6-19 REV. 11, JANUARY 2005 D3.6.3.2.6 MSIV Leakage Control System Rooms The MSIV leakage control system rooms are located on elevation 737 feet 0 inch between column rows 110, 112, and W,Z. The inboard system (Division 1) and the outboard system (Division 2) are physically separated from one another by a barrier wall halfway between column rows 110 and 112. This wall physically ensures that any steam or water line break in one room will affect that division only, leaving the other division to function. It was conservatively assumed that a water spray from a broken line in either room would disable all the equipment in

that room. D3.6.3.2.7 Pipe Tunnel The auxiliary building tunnel is a vertical tunnel that passes through elevation 737 feet 0 inch from the auxiliary building steam tunnel to the RCIC pump room. The tunnel is located between column rows 110, 114 and U,U.8. This pipe tunnel contains high energy piping and the valves associated with operation of the RCIC turbine. There are no other safety-related components in this tunnel. D3.6.3.2.8 Auxiliary Building Steam Tunnel The auxiliary building steam tunnel was not evaluated for a MELB, since HELBs are postulated in this area. The condition of a HELB is more severe than that of a MELB. D3.6.3.2.9 Electrical Switchgear/Motor Control Center Rooms Two rooms in the Auxiliary Building Elevation 781'-0" contain Divisional Class 1E switchgear and motor control centers. The room to the West of the Auxiliary Building Steam tunnel is located between Column Rows 102, 107, S and AD; and contains Division 2 equipment. The room to the East of the Auxiliary Building Steam tunnel is located between Column Rows 117, 124, S and AB; and contains Division 1 equipment. The only piping in these areas that is of concern, is the shutdown service water piping supplying the room area coolers. This portion of the piping in the area of the switchgear has been classified as a no-break area due to the low piping stresses. D3.6.3.3 Fuel Building The fuel building contains relatively little electrical equipment to be protected from a water spray. The only essential equipment in the fuel building is in the high-pressure core spray cubicle containing its associated support equipment. The other item in this area which is supplied with electric power is the low-frequency motor generator sets, which supply electrical power to the reactor recirculation pumps. This item is not required for the safe shutdown of the plant and was therefore not evaluated for damage from water spray. D3.6.3.3.1 High-Pressure Core Spray Pump Room The high-pressure core spray pump room is located on elevation 707 feet 6 inches between column rows 102, 106 and AD,AH. All of the equipment in this room is associated with the high pressure core spray system which is Division 3. The electrical equipment of HPCS is not adversely affected by wetting caused by a MELB spray when the equipment in the room is qualified for 100% relative humidity, which is no worse than the wetting of the equipment due to CPS/USAR CHAPTER 03 D3.6-20 REV. 11, JANUARY 2005 a MELB. Even assuming a failure, the plant can be safely shut down without the use of the high-pressure core spray. D3.6.3.4 Control Building Piping in the control building was located with the intention of minimizing the probability of equipment damage from pipe leaks. This approach, combined with the special separation of redundant equipment in the system, results in very low probability of any hazardous effects of leakage. The plant service water and chilled water systems are the primary sources of wetting in the control building. The chilled water lines are located in the basement of the control building. The hydrogen recombiner skids are the only essential equipment located in the basement area. The recombiners are designed to remain operable within the environment of a water spray. Piping on other floors is limited to plant service water and cooling water for area coolers and chillers. The lines are routed to the extent feasible in areas where there is no other equipment.

Therefore, a postulated failure could not wet essential switchgears, motor control centers or equipment. No piping is located in or near the main control room or in the cable spreading room

below the control room. D3.6.3.5 Diesel-Generator Building Each of the two standby diesel generators and the one HPCS diesel generator has its own room within the building. The existence of doors between the divisional diesel generator rooms has no safety significance. If a postulated line break occurs in the shutdown service water piping (which is the largest fluid source in the room), the floor drain system provided in the diesel generator cubicles is sized to accommodate the expected flow from the pipe break. The doors have an 8" curb. No water accumulation would occur such that the remaining other divisional equipment could be impaired. Consequently, a moderate energy line break in the piping for any one diesel and a single failure will leave sufficient redundant sources of power to safely shut

down the reactor. In addition, most of the piping is located relatively low in the building, and by the arrangement of the electrical equipment, sensitive items would not be significantly wetted. D3.6.3.6 Circulating Water Screen House The circulating water screen house contains the shutdown service water pumps of Divisions 1 and 2 and the cooling water pump for the high-pressure core spray diesel. These pumps and their associated equipment are the only essential items located in the circulating water screen

house. The shutdown service water pumps are located on elevation 699 feet. The pumps for Unit 1 are located in the northeast corner of the building between column rows 1 and 2, close to column C. Each pump is located in its own cubicle and physically separated from all other pumps. All associated support equipment for each pump is located in its respective cubicle. Consequently, no postulated pipe failure in either pump room would disable the redundant pump for that unit.

In addition, failure of the Shutdown Service Water System (Division 3) to supply cooling water to the high-pressure core spray diesel cooling system would disable only the high-pressure core spray diesel. This would not prevent safely shutting down the plant.

CPS/USAR CHAPTER 03 D3.6-21 REV. 11, JANUARY 2005 D3.6.4 INTERNAL FLOODING An analysis has been performed to ensure that flooding as a result of postulated high and moderate energy line breaks will not compromise the safe shutdown capability of the Clinton Power Station. Flood levels for various areas of the plant were calculated. These flood levels were used as input to structural load and safe shutdown assessments. See Subsections 3.4.1 and 3.8.4 for structural load assessment due to flooding. Safe shutdown for postulated flooding due to internal piping failures is analyzed for areas containing safe shutdown equipment and/or areas where flooding could potentially impact associated circuits of electrically operated safe shutdown equipment. The buildings which were analyzed for safe shutdown following postulated internal flooding are: a. auxiliary building, b. containment (including drywell),

c. control building,

d. diesel-generator building,

e. fuel building,

f. radwaste building, g. screenhouse, and h. turbine building. The auxiliary, containment, control, diesel-generator, fuel, and screenhouse buildings contain safety-related equipment. The flood protection arrangement of the Circulating Water Screen House is shown in Figure D3.6-134. The radwaste and turbine buildings do not house safety-related equipment, but flooding in these buildings has potential impact on buildings which do house safety-related equipment. Analyses were performed to determine the flood level response of the various areas of the station to postulated failures of moderate and hi gh energy fluid systems. To accomplish this, the station was divided into flood zones. Many of these zones, termed "general areas," are areas that exhibit large open spaces within the plant and often contain stairwells and hatches that are open to lower levels. Other areas, termed "subcompartments," are smaller areas (generally) enclosed by Seismic Category 1 walls that open to the general areas only through doorways or hatches. Maximum flood levels for any particular flood zone were calculated assuming a single piping failure as the initiating event (Branch Technical Position ASB3-1). For each zone, high and moderate energy piping was evaluated to determine which single postulated line failure would produce limiting flood levels. The fluid release rates were calculated based on crack or break sizes determined from Standard Review Plan 3.6.2 and the duration of the release was generally taken to be 60 minutes. Duration of the release was CPS/USAR CHAPTER 03 D3.6-22 REV. 11, JANUARY 2005 generally taken to be 46 minutes for the circulating water expansion joint failure and 30 seconds for isolable RWCU failures. Certain breaks (e.g., ECCS suction lines) were assumed to be non-isolable. In most cases (i.e., MELB) the blowdown rate was considered to be constant. In other cases (i.e., HELB) an initial high blowdown rate (based on line inventory) was followed by a smaller rate based on an upstream limiting area. Potential/pipe and jet/impingement piping failures following a HELB could contribute to the flood source and these were considered in the

analysis. Fluid removal from general areas was by means of centrally located stairwells or open hatches and to a lesser extent, floor drains. Where no such removal paths were present, the fluid was assumed to accumulate within the area. Fluid removal from subcompartments was by means of floor drains and flow under doors leading to general areas. Flood levels for the zones were calculated for breaks within and outside of the areas of interest and the limiting flood levels were tabulated. The calculated flood levels were found to be less than 2 inches for most areas of the plant. The zones that experience flood levels exceeding 10 inches are concentrated in the lower elevations of the plant while flood zones in the upper floors of the plant outside containment would not experience flooding above 10 inches. This result reflects the general design of systems within the plant with large lines and high energy systems located primarily on the lower floors. One potential source of major flooding is a postulated failure of the circulating water expansion joint in the turbine building. In the event that the non-safety-related flood mitigation systems fail to perform their function, a postulated failure of this expansion joint could result in flooding up to elevation 719 feet in all areas of the power block except for the auxiliary building, the HPCS pump room, and the diesel oil tank rooms. Another major source of flooding is the postulated failure of any of the non-isolable portions of the ECCS pump suction lines to the suppression pool. Postulated failure of one of these lines could result in flooding of a single ECCS cubicle to the high water level in the suppression pool (elevation, 731 feet 5 inches). Flooding events are analyzed by one of two methods for the safe shutdown assessment: a. zone-by-zone basis b. flooding source-by-source basis The zone-by-zone method examines the maximum flood level in each zone (for any postulated piping failure) and determines which, if any, electrically operated safe shutdown equipment is subject to submergence. This method is applied to all areas of the plant except for portions of the lowest elevation (elevation, 707 feet 6 inches/712 feet) in the auxiliary building. The source-by-source method examines safe shutdown on a flood source-by-source basis. For each postulated piping failure, the effect of the calculated flood level in that flood zone as well as concomitant levels in other zones is evaluated for safe shutdown. In addition to identifying which, if any, safe shutdown equipment is subject to submergence, the potential for reactor/

turbine trip as a direct consequence of piping failures was also evaluated. This source-by-source method is applied to portions of the lowest floor (elevation, 707 feet 6 inches/712 feet) of the auxiliary building.

CPS/USAR CHAPTER 03 D3.6-23 REV. 15, JANUARY 2013 For each flooding event, the most restrictive SAF was assumed. When the flooding event was due to a postulated failure in a dual purpose moderate energy piping system, a SAF in the redundant system was not considered per Paragraph B3.b(3) of BTP-ASB 3-1 of Standard Review Plan 3.6.1. Credit for use of the redundant dual-purpose system for achieving safe shutdown is contingent upon the availability of a diverse power supply (i.e., offsite power and emergency onsite power). This dual purpose ex clusion criteria is applied to the RHR heat exchanger trains (RHR-A, RHR-B) and the directly associated supporting systems which include the shutdown service water train, the cubicle coolers, and the Class 1E divisional power supply. Loss-of-offsite power is assumed when reactor/turbine trip is a direct consequence of the initial flooding event. In certain instances, LOOP is not a conservative assumption, and thus was not assumed (e.g., the postulated circulating water expansion joint rupture in the turbine building requires offsite power to drive the flooding source - the circulating water pumps). Postulation of line breaks in high energy RHR piping in the RHR Heat Exchanger Rooms A and B is not required (refer to NRC letter dated January 7, 1986, Docket No. 50-461). Therefore, flooding from these sources is not included in the design basis. Flooding from postulated high energy line breaks within the Auxiliary Building Main Steam Tunnel have been included in the design basis, and safe shutdown has been demonstrated for these postulated breaks. Safe shutdown has been assured by analysis for design basis flooding events.

The height of the stairway landing/door bottom (elevation 715 feet) between the turbine building (elevation 712 feet) and the auxiliary building (elevation 707 feet 6 inches) is above the water level in the turbine building that would be caused by the unlikely rupture of a line from the two condensate storage tanks and the consequent release of 800,000 gallons of water into the turbine building. In the event of a simultaneous rupture of this line and a line from the demineralized water storage tank (400,000 gallons), water will not be contained in the turbine building. However, water will not enter the core spray cooling pump (RHR, LPCS, HPCS, and RCIC) cubicles in the auxiliary building, since they are watertight to elevation 731 feet 5 inches. In the event that an entire circulating water expansion joint fails, leaving a 6.25 inch gap between the piping and the waterbox, 267,400 gallons of water per minute will be released to the turbine building. Each condenser cavity, designed to contain flooding to elevation 715 feet, is equipped with a redundant system of level switches which will alarm in the control room if the water level in the condenser cavity reaches an elevation of more than 1 foot (elevation 710 feet) above the condenser cavity floor at elevation 709 feet. Additionally, thes e level switches will close a motor-operated valve in the floor drain piping between the condenser cavity and the turbine building floor drain sump to slow flooding of the turbine building. A second system of redundant level switches will automatically stop the circulating water pumps if the flood water reaches an elevation of 714 feet within the condenser cavity. An additional foot, from elevation 714 feet to elevation 715 feet remains to contain the water flow due to the coastdown of the circulating water pumps after they are initially signaled to stop. The level switches that stop the circulating water pumps are powered by CW Pump A control power. Therefore, if the CW Pump A control fuses are pulled, these level switches would not stop CW Pumps B & C, if required. In the event of failure of an expansion joint and both redundant sets of level switches, the turbine building could be flooded above 715 feet. Then, because of flow areas between the turbine building and radwaste and control buildings, they could be flooded also. The limiting level of 719 feet could be reached only after 46.6 minutes with no operator action to stop the circulating water pumps. The turbine building water level could reach 726.7 feet in this time, but CPS/USAR CHAPTER 03 D3.6-24 REV. 15, JANUARY 2013 no essential equipment would be affected. In addition, even assuming the failure of level switches, postulated above, the control room operator will still have adequate warning of

flooding in the turbine building. By elevation 719, the CRD pumps and turbine building MCC's 1A, 1B, 1C and 1D, among other things, will be flooded, and before 726.7 feet the condensate and condensate booster pumps will be lost. The control building is protected up to 719 feet. The radwaste building contains no equipment required for safe shutdown.

CPS/USAR CHAPTER 03 3.7-1 REV. 11, JANUARY 2005 3.7 SEISMIC DESIGN Safety-related structures, systems, and components that are designed to remain functional in the event of a safe shutdown earthquake (SSE) are designated as Seismic Category I. All Seismic Category I items are analyzed and designed through the use of appropriate methods of dynamic analysis as described in the following subsections. 3.7.1 Seismic Input 3.7.1.1 Design Response Spectra The horizontal design response spectra defined at the ground surface (surface spectra) are shown on Figure 3.7-12 through 3.7-15 and Figures 3.7-20 through 3.7-23 for OBE and SSE, respectively. The vertical design response spectra defined at the ground surface (surface spectra) are shown on Figures 3.7-16 through 3.7-19 and Figures 3.7-24 through 3.7-27 for OBE and SSE, respectively. The maximum horizontal and vertical ground accelerations at the foundation level corresponding to the above site response spectra are 25% of gravity for the

safe shutdown earthquake (SSE) and 10% of gravity for the operating basis earthquake (OBE). The design response spectra comply with Regulatory Guide 1.60. In accordance with agreements reached on September 24 and October 15, 1981 meetings between IP and the NRC staff, IP performed a seismic hazard analysis for the CPS site to demonstrate that the plant is located in an area which does not pose any higher seismic hazard than anywhere else in the Central Stable Region. Consistent with the results of this hazard analysis, site specific response spectra will be developed which will be used as the basis for comparing response spectra (derived from a soil spring analysis) to the current design response

spectra which were derived from a finite element analysis. This comparison is expected to show that the finite element analysis results either bound or are very close to the soil spring

analysis results. In order to resolve the concern identified in this question, as well as in questions 220.15, 220.21, and 220.26, the following work has been done: a. Site specific response spectra were developed for the Clinton site for a 5.8 magnitude earthquake, as described in the Weston Geophysical report, "Site Specific Response Spectra for the Clinton Power Station - Unit 1", submitted to the NRC staff with References 16 and 17. b. The soil-structure interaction analysis was performed using the soil spring method; variation in soil properties was also considered in this analysis; and no deconvolution

was used. c. The critical plant structures, piping and equipment were reevaluated to the new seismic loads, and it was concluded that the Clinton plant design is conservative and can withstand the new seismic load. d. The details of the analysis described in paragraph (b) above, the responses obtained from these analyses, and the reevaluation results are contained in References 18, 19

and 20. (Q&R 220.14)

CPS/USAR CHAPTER 03 3.7-2 REV. 11, JANUARY 2005 3.7.1.2 Design Time History The following two-step procedure is used for generating the foundation and rock level (base on the soil-structure interaction system) time histories. In the soil-structure interaction analysis, the rock time history is applied at the base of the soil-structure model. STEP 1: Generation of Design Time History The north-south and vertical components of the 1940 El Centro earthquake records are modified using the program RSG (see Appendix C for description of RSG) so that the response spectra generated using these synthetic records match closely the Regulatory Guide 1.60 response spectra for the horizontal and vertical directions. The frequencies used in generating the response spectra from the modified synthetic time histories are spaced as follows: Frequency Range (Hz) Increment (Hz) 0.5-3.0 0.10 3.1-3.6 0.15 3.6-5.0 0.20

5.0-8.0 0.25 8.0-15.0 0.50 15.0-18.0 1.0 18.0-22.0 2.0 22.0-34.0 3.0 The comparison of response spectra obtained from horizontal and vertical synthetic time histories and the corresponding Regulatory Guide 1.60 design spectra for 0.2g ground acceleration is presented in Figures 3.7-1 through 3.7-10 for 2, 3, 4, 5, and 7 percent damping ratios. The synthetic time histories are then scaled to the required ground surface design response spectra acceleration levels. STEP 2: Generation of Foundation and Rock Motion The soil profile above the rock is modeled as a one-dimensional continuous shear layer system. The layering scheme of the 246 feet of soil below the plant complex is shown in Figure 3.7-11. The soil below the circulating water screen house is shown in Figure 3.7-76. The design time history obtained in Step 1 is applied at the ground surface, and the foundation and rock time histories are obtained using the program SHAKE (see Appendix C for description of SHAKE program). The strain-dependent soil properties for various soil layers used in the SHAKE analysis are given in Table 3.7-9. A comparison of the free field foundation and design spectra for the main plant for 1, 2, 3, and 4 percent damping ratios for OBE and 1, 3, 4, and 7 percent damping ratios for SSE are given in Figures 3.7-12 through 3.7-27. Variation in soil properties at the site has been taken into account in the soil-structures interaction analyses using soil spring method. A detailed review of the dynamic soil properties was undertaken. The goal of the review was to define the upper and lower bound curves of soil CPS/USAR CHAPTER 03 3.7-3 REV. 11, JANUARY 2005 shear modulus valves. At the same time it was decided to develop site specific response spectra for the Clinton site to resolve the their seismic soil-structure interaction issue identified by NRC staff in questions 220.14, 220.21, and 220.26. For this purpose, an estimate of shear wave velocities for soils present below the foundation mat was required. A review of the shear wave velocities give in Figures 2.5-369 through 2.5-371 suggested that in light of the knowledge gained from comparable soil deposits, the shear wave velocities given in these figures were high. The shear wave velocities given in the FSAR were computed from the measured compressional wave velocities and estimated Poisson's ratio. In view of the current knowledge, the estimated values of Poisson's ration are considered low. Based on the above, a thorough review of the shear wave velocities and the low-strain soil moduli was performed by Dames & Moore. Based on the results of this review, Figures 2.5-369 through 2.5-371, and Table 2.5-46 and 2.5-48 were revised in Amendment 12, dated January 1982. The rationale and references that were used to estimate the shear velocity for the glacial soils are summarized in Table A3.7-1, Evaluation of Geophysical Data. This table summarizes the geophysical measurements made at five nuclear plant sites. In all cases, both the compressional and shear wave velocities were measured in the field. As shown in the table, Poisson's ratios for glacial soils range from 0.45 to 0.48 based on the values calculated from the measured velocities presented in the table. As a result, it was estimated that Poisson's ratios for the Illinoian glacial till and Wisconsinan glacial till are 0.46 and 0.48, respectively. These values, along with the measured compressional velocity, were utilized to calculate the estimated

shear wave velocity at Clinton. For structural fill, the normalized shear modulus factor (k2) versus shear strain relationship was established from the laboratory test data. In addition, k 2max was calculated using the Hardin and Drnevich equation, Reference 21. Since the k 2max value obtained from the lab data is less than the k2max value calculated from the Hardin and Drnevich equation, Professor Drnevich was consulted. Based on the discussions with Professor Drnevich, it was concluded the k 2 values obtained from the laboratory tests should be multiplied by a factor of about two, resulting in a k 2max of 100. The recommended values for use are given in Table 2.5-48. Attachment A3.7 is a letter from Professor Drnevich that is attached for reference and indicates that a k 2max of 100 is realistic. In the soil-structure interaction analysis of the plant structures using soil spring method, a range of soil properties has been used. Q&R Figures 220.15-2 through 220.15-5, extracted from Reference 22, show the upper bound and lower bound soil properties curves used. A comparison of these curves with the values given in Table 2.5-48 are bounded by the curves shown in these figures. Since the structural evaluation has used the curves shown in the above figures, it is concluded that there is no effect of the revisions in Table 2.5-48 on plant structures.

The two earth structures, i.e., the natural slopes surrounding the ultimate heat sink and the submerged dike, have Table 2.5-48, and have been found to have adequate factors of safety.

(Q&R 220.15)

CPS/USAR CHAPTER 03 3.7-4 REV. 11, JANUARY 2005 3.7.1.3 Damping Values 3.7.1.3.1 Critical Damping Values The damping ratios (expressed as a percentage of critical) used in the analysis of various Seismic Category I structures, systems, and components are listed in Table 3.7-1. These damping ratios conform to Regulatory Guide 1.61. The damping values used in the analysis of Category I structures are in compliance with the requirements of Regulatory Guide 1.61, Section C.3. Since the maximum combined stress in Category I structures due to static, seismic, and other dynamic loading are not significantly lower than the yield stress and one-half yield stress for SSE and OBE, respectively; Regulatory Guide 1.61, Table 1 damping values were used. (Q&R 220.16) Alternative critical damping values for piping may be used as described in Section 3.7.1.3.2. 3.7.1.3.2 Alternative Critical Damping Values for NSSS Piping Alternative critical damping values, as provided in ASME Boiler and Pressure Vessel Code, Section III, Division 1 Code Case N-411-1 may be used. When used, the following provisions are applied. 1. The code case damping is applied only to uniform (or envelope) response spectra loading analysis for seismic and seismic-like building filtered hydrodynamic loads and the annulus pressurization loading. 2. The code case damping is applied to a spectral analysis load case in its entirety and is not mixed with other damping values within that one load case. 3. Modal and direction combination of the three earthquake directions are combined in accordance with Regulatory Guide 1.92. 4. Consideration of a sufficient number of modes such that the inclusion of additional modes would not result in more than a 10% increase in response. 5. Assurance that the predicted piping displacements are such that adequate clearance exists with respect to adjacent components and equipment. 6. Line mounted equipment is designed to withstand the increased pipe motion. 7. The code case damping is not applied to piping analytical models that incorporate equipment with natural frequencies below 20 hertz. 3.7.1.4 Supporting Media for Seismic Category I Structures The description of the supporting media for Seismic Category I structures is presented and discussed in Section 2.5. The soil properties used in the design basis seismic analysis (finite element soil model) discussed in answers to Questions 220.14, 220.15, 220.21, and 220.26 are based on the properties given in Table 2.5-48. The following is a list of Seismic Category I structures with the embendment depth, the depth of soil between bedrock and foundation, the foundation width, and the structural height.

CPS/USAR CHAPTER 03 3.7-5 REV. 11, JANUARY 2005 Foundation Width (ft)

Direction Structure Embedment Depth (ft)

Depth of Soil (ft)* E-W N-S Structural Height Above Grade (ft)

Containment Building 34 199.6 130 (diameter) 191 Auxiliary Building 38.2 195.4 178 122 64 Fuel Handling

Building 33.7 200.7 182 151 64 Circulating Water

Screen House 44.5 152.9 176 238 32.5 Control Building 43 189.2 219 100 112.2 Diesel Generating and HVAC Building 33 199.2 221.1 106 52 Radwaste Building (substructure only) 43 186.1 232 321 45

• Values obtained using representative parameters for elevation of bedrock. Refer to Figure 2.5-282 and Figure 2.5-373. 3.7.2 Seismic System Analysis The seismic analysis of the containment and other main building structures is carried out for two building models, one for the single unit building complex presently being constructed, and the second for the two-unit building model. The second unit was planned for construction later on the extension of the single unit basemat. Two separate soil-structure interaction analyses are carried out with the two building models and the corresponding foundation interaction time histories are used for generating various floor spectra in each building model. The envelopes of the floor response spectra from the two analyses are used for the analysis of the equipment and piping supported on various floors. 3.7.2.1 Seismic Analysis Methods The seismic analysis is performed using the modal superposition method. The member forces and accelerations of mass points are determined by the response spectrum method while the response spectra at various floor elevations for subsystem analysis are generated by the time history method. All modes with frequencies less than 33 Hz were included in the analysis except when the number of modes required to reach 33 Hz exceeded 30. For these cases, 30 modes were used in the analysis. The following provides the number of modes and the highest frequency considered for the various models:

CPS/USAR CHAPTER 03 3.7-6 REV. 11, JANUARY 2005 BUILDING DIRECTION NUMBER OF MODES HIGHEST FREQUENCY Main Plant (1 unit) horizontal

vertical 25 26 52.9533 32.5488 Main Plant

(2 unit) horizontal

vertical 25 27 42.8550 32.2245 Containment horizontal vertical 30 30 27.9994 26.0777 Note: Unit 2 has been canceled. (Q&R 220.17) Dynamic modeling of the building structures is described in Subsection 3.7.2.3. The computer program DYNAS (Dynamic Analysis of Structures) is used to analyze the Seismic Category I building structures. The description of this program is given in Appendix C. Figures 3.7-28 through 3.7-30 are typical sketches of the horizontal seismic model for one and two unit stations. Figure 3.7-79 is the sketch of the horizontal seismic model for the circulating water screen house. Rigid slabs at various floor elevations are connected by shear wall springs. The containment is modeled as a lumped mass spring model. The horizontal models are analyzed for X (E-W direction) and Y (N-S direction) excitations and the results are combined as described in Subsection 3.7.2.6. As in the horizontal analysis, both response spectrum and time history methods of analysis are performed on the vertical model also using the DYNAS program. 3.7.2.2 Natural Frequencies and Response Loads 3.7.2.2.1 Horizontal Excitation The periods, mode shapes, and dynamic responses of the structural lumped mass models shown in Figures 3.7-28 through 3.7-30, and 3.7-79 are computed using the program DYNAS.

Summaries of the modal frequencies and modal participation factors for the containment model, single unit main plant model, two unit main plant model, and circulating water screen house model are presented in Tables 3.7-2 through 3.7-4, and 3.7-12. Seismic response loads for the safe shutdown earthquake for the containment wall and major Seismic Category I shear walls are shown in Figures 3.7-31 through 3.7-37. Response spectra for the design of subsystems consist of envelopes of the responses obtained for the single unit and two unit plant models. Design horizontal acceleration response spectra for the SSE in the East-West and North-South directions at the base slab (elevation 712 feet 0 inch), top of the reactor pedestal (elevation 742 feet 8 inches), drywell floor (elevation 803 feet 3 inches), grade floor (elevation 737 feet 0 inch), and mezzanine floor (elevation 762 feet 0 inch) are shown in Figures 3.7-38 through 3.7-47. The design horizontal acceleration response spectra for the circulating water screen house in the east-west and north-south directions at the base slab (elevation 653 feet 6 inches), intermediate floor (elevation 682 feet 6 inches), main floor (elevation 699 feet 0 inch) and roof (elevation 730 feet 0 inch) are shown in Figures 3.7-82 through 3.7-89 for a SSE excitation.

CPS/USAR CHAPTER 03 3.7-7 REV. 11, JANUARY 2005 3.7.2.2.2 Vertical Excitation The modal frequencies and participation factors of the containment, main structure, and circulating water screen house lumped mass models shown in Figures 3.7-48, 3.7-49, 3.7-80, and 3.7-81 are presented in Tables 3.7-5 through 3.7-7, and 3.7-13. Forces in the structures resulting from a vertical excitation are obtained by a response spectrum method of analysis. Seismic response loads in the containment wall for the single unit and two unit lumped mass systems are shown in Figures 3.7-50 and 3.7-51. Enveloped vertical acceleration response spectra for the main plant at the base slab (elevation 712 feet 0 inch), top of the reactor pedestal (elevation 742 feet 8 inches), drywell floor (elevation 803 feet 3 inches), the grade floor (elevation 737 feet 0 inch), and the mezzanine floor (elevation 762 feet 0 inch) are shown in Figures 3.7-52 through 3.7-56. The design vertical wall acceleration response spectra for the circulating water screen house at the base slab (elevation 653 feet 6 inches), intermediate floor (elevation 682 feet 6 inches),

main floor (elevation 699 feet 0 inch), crane level (elevation 719 feet 0 inch), and roof (elevation 730 feet 0 inch) are shown in Figures 3.7-90 through 3.7-94 for SSE excitation. 3.7.2.3 Procedure Used for Modeling 3.7.2.3.1 Designation of System Versus Subsystem Analysis of a nuclear power plant complex subjected to seismic excitations is divided into two parts. The first is analysis of "seismic systems" comprising major buildings and structures which house and/or support Seismic Category I systems and components. The second part is an analysis of "seismic subsystems," which include Seismic Category I systems and components. Major structures which are analyzed as seismic systems are: a. main plant complex - control, diesel and HVAC, radwaste and Unit 1 and 2 fuel, auxiliary and turbine buildings; b. containment wall; c. containment inner structures - drywell wall, shield wall, weir wall and RPV pedestal; and d. circulating water screen house. 3.7.2.3.2 Decoupling Criteria for Subsystems All subsystems such as equipment and piping are decoupled from the floors on which they were supported, since the mass of the structures is large relative to the subsystem masses. However, the masses of these subsystems are included with the structural mass of the supporting floor slabs. No specific ratios between subsystem mass and system mass, R m , or between fundamental frequencies of subsystem and system, R f, used in subsystem decoupling, since no quantitative CPS/USAR CHAPTER 03 3.7-8 REV. 11, JANUARY 2005 criteria were available at the time the seismic model was generated. All subsystems, except the RPV mass, were lumped at the appropriate location in the seismic model. The RPV was

modeled as part of the seismic model. The subsystems, with the exception of the RPV, generally have a small mass ratio (R m is less than 0.01) or frequencies away from resonance with the system (0.8 R f 1.25). (Q&R 220.18) 3.7.2.3.3 Lumped Mass Consideration Two independent models are used to obtain responses to horizontal and vertical excitations. Since the response in the horizontal direction due to a vertical excitation (and vice versa) is negligible, the models for horizontal and vertical analyses are decoupled. Horizontal responses are obtained for excitations along two principal horizontal axes. The results of these analysis are then combined using a square-root-of-the-sum-of-the squares (SRSS) method. For the main building horizontal model, all shear walls and slabs have been modeled at their physical location. This assures the adequacy of the model. The containment horizontal model consists of 270 degrees of freedom. Thirty modes were extracted from this model for the modal seismic analysis. This meets the SRP requirements that the degrees of freedom be at least two times the number of modes with frequencies less than 33 CPS. For the main building and containment vertical models, each slab location and each of the slab panel dynamic characteristics is modeled. This assures the adequacy of the model. (Q&R 220.19) 3.7.2.3.3.1 Model for Horizontal Excitation In the lumped mass idealization, the entire mass of the structure is concentrated at a number of discrete points. In general, each mass point has six degrees of freedom. However, certain degrees of freedom may be neglected depending on the configuration of the structure and the type of excitation. The concrete elements connecting the lumped masses are modeled as linear elastic members. The main plant complex and circulating water screen house are modeled as a shear structure consisting of rigid concrete slabs interconnected with shear walls. The predominant mode of deformation is shear deformation. Consequently, the only significant rotations are those about the vertical axis, and for each slab only three degrees of freedom are considered, two horizontal translations and rotation about the vertical axis. Three mass parameters, corresponding to the three degrees of freedom, are associated with each slab. The mass parameters associated with the two horizontal translations are the same and are equal to the mass of the slab. The mass parameter associated with rotation about the vertical axis is equal to the mass polar moment of inertia of the slab about a vertical axis through its centroid. These mass parameters are computed from the mass distribution of the slabs, equipment, piping and tributary walls. Since the shear walls are distributed horizontally, the model slabs are actually treated as rigid bodies with only horizontal dimensions. Shear wall forces are a function of wall stiffness, location of the wall with respect to the centroids of the slabs and the relative motion of the slabs to which the shear walls are attached. Concrete shear CPS/USAR CHAPTER 03 3.7-9 REV. 11, JANUARY 2005 walls are treated as deep beams for computation of stiffness, while conventional frame or truss analysis methods are used to compute the stiffness of steel framing. The containment structures is modeled as a frame structure with all six degrees of freedom considered at the lumped mass locations. The discrete mass at each node includes equipment, piping, wall, and slab masses. The mass of the water in the suppression pool and containment pool is lumped with the containment wall and drywell masses at appropriate elevations. The stiffnesses of the structural elements are computed based on the geometry of the structure and the assumption of linear elastic behavior. The containment model includes a model of the reactor pressure vessel (RPV) and internals (see Figure 3.7-30). A detailed description of the RPV analysis is included in Subsection 3.7.3.14. The shear structures and frame models for the main building complex and containment building are coupled for use in the soil-structure interaction analysis described in Subsection 3.7.2.4.

After the interaction time history and response spectrum at the foundation have been computed, the models are decoupled and further analysis is performed on each separately. The sloshing effects of the water in the suppression pool, containment pool, and fuel pools in the fuel handling building were evaluated in accordance with Housner's method explained in Reference 23. It was found that because of the pool geometry, the convective wall pressure was less than the impulsive pressure. In the design of the pools, the impulsive pressure was considered acting throughout the depth of the pool which is conservative. (Q&R 220.20) 3.7.2.3.3.2 Model for Vertical Excitation A frame system is used to model the main plant complex, the containment building, and the circulating water screen house for analysis of vertical excitations. Since axial deformations are the dominant mode for vertical excitations, only the vertical degrees of freedom are considered in the analysis. Several single-degree-of-freedom systems are connected to the wall system (See Figures 3.7-48, 3.7-49, and 3.7-81) to simulate the multiperiod characteristics of the slab

and beam systems. Masses are concentrated at wall-slab intersections and at the center of the slabs. Wall masses are distributed equally to the adjacent slabs. One-third of the total slab mass, including piping and equipment, is assumed to be effective for slab vibrations. The remaining mass is lumped with the wall at that elevation. After the vertical interaction time history and response spectrum at the foundation are obtained, further analyses of the containment and main plant are performed separately. The RPV and internals are included in the containment model (see Figure 3.7-48). Detailed analysis for the RPV is discussed in Subsection 3.7.3.14. 3.7.2.4 Soil-Structure Interaction The supporting media for various Seismic Category I structures are given in Subsection 3.7.1.4. The horizontal soil-structure interaction analysis is done using a finite element soil model.

CPS/USAR CHAPTER 03 3.7-10 REV. 11, JANUARY 2005 Strain-dependent soil parameters used in the interaction analysis are presented in Table 2.5-48. The design time history is applied at the free field ground surface level. The SHAKE program (described in Appendix C) is used to analyze the model shown in Figure 3.7-11 for the main plant and containment building and Figure 3.7-76 for the circulating water screen house in order to obtain the strain-compatible shear modulus and damping values for each layer for both OBE and SSE earthquakes. Corresponding compatible rock motions for OBE and SSE are also obtained from this analysis. Strain compatible shear modulus (G) and damping values for each layer for horizontal OBE and SSE excitations are presented in Table 3.7-14 for the main plant and containment. Vertical ground motion is considered to travel as a compression wave. In order for "SHAKE" to perform a compression wave analysis, the shear moduli obtained from horizontal excitations are multiplied by a factor

()(),21,12µµ where µ is Poisson's ratio. (This factor is equal to the ratio of compression wave and shear wave velocities squared.) Damping values used for the vertical analysis are the same as for the horizontal analysis. The strain levels in TAble 3.7-14 are based upon the design time history and are consistent with the design earthquake (0.26g, Regulatory Guide 1.60 spectrum at the ground surface for SSE) and are not considered high. For soil properties variation, see the response to Question 220.15. (Q&R 220.21) For the main plant and containment building, a set of soil properties and the corresponding rock motions for both OBE and SSE earthquakes are obtained, then the axisymmetric finite element soil model shown in Figure 3.7-57 is used for extracting normalized modes of soil using the computer program DYNAX (see Appendix C for description of DYNAX program). In the DYNAX soil model shown in Figure 3.7-57, the structural basemat is modeled through the use of massless rigid shell elements at the foundation elevation. The model has both translational and rotational degrees of freedom at the interface node. The rigid elements at the foundation element simulate the actual distributed interface between the structural basemat and the soil, even though the structure is connected at only one point in the mathematical model. The effects of neglecting the interaction between the walls and the soil is insignificant because: i) The embedment depth (d) is small compared to the structural base dimension (B). For the one unit model, the average d/B ratio is 0.07. For such small embedment ratios, the embedment effects on soil impedance are negligible. ii) The soil-structure interaction responses using perfect contact between the foundation and the side soil may not be realistic, because a partial loss of contact between the soil and the foundation leads to a pronounced reduction in rocking stiffness. iii) Neglecting the effect of side soil leads to more rocking of the structure, and thus is conservative. It should be noted that even though the interaction between the side soil and the walls is neglected for the soil-structure interaction analysis, the walls are designed to resist dynamic earth pressure loads expected during the seismic event. (Q&R 220.22)

CPS/USAR CHAPTER 03 3.7-11 REV. 11, JANUARY 2005 The three dimensional axisymmetric soil media model is shown in Figure 3.7-57. In this model the boundary condition are as follows: (i) All surface nodes are free to model zero traction condition at the surface; (ii) The soil basemat interfact nodes are connected by rigid elements to simulate the rigid foundation interface; (iii) Nodes at the bottom boundary are fixed. This boundary coincides with the physical rock-soil interface at the CPS site. (iv) The lateral free field boundary nodes are free to move in the horizontal direction but are contained in the vertical direction. This boundary condition simulates the shear beam behavior for the soil media, consistent with the concept that horizontal seismic motions can be modeled by vertically propagating shear waves. To assure that the imposed lateral boundary condition has no practical effect on the soil structure interaction responses, the lateral boundary was chosen to be sufficiently far from the structure. To verify that the lateral boundary location is far enough from the structures in the coupled soil structure interaction model, response at the foundation elevation at the lateral boundary was computed in the coupled soil structure interaction model. This response was compared to the foundation elevation free field response obtained from the SHAKE analysis. This comparison is presented in Figure 3.7-95. It can be observed that the soil model for SSI is large and does simulate the free field conditions at distances sufficiently far from the structure (Q&R 220.24) Using the modal synthesis technique, the three dimensional building models (one and two unit) shown in Figures 3.7-58 and 3.7-59 are analyzed in the two orthogonal horizontal directions X and Y using the program DYNAS for OBE and SSE earthquakes. One discrete torsional soil spring and corresponding mass are included to account for possible torsional interaction due to the non-symmetric nature of the building complex. The torsional spring constant for each soil layer is calculated as (Reference 13): 3Gr 3 16 k o= where r o is the radius of the effective area. For a rectangular base with dimensions B and L,

()422 o 6LBBL r+= G = shear modulus of soil layer obtained from SHAKE. Finally, the total spring constant is calculated by adding the torsional stiffnesses of springs in series for all layers. The effective mass inertia (I) of the soil participating in torsional vibration is taken as (Reference 13):

o 5 gr3.0 I=

CPS/USAR CHAPTER 03 3.7-12 REV. 11, JANUARY 2005 where = unit weight of soil. The modal damping values required in DYNAS are obtained from the DYNAX program by using the strain-compatible damping values obtained from SHAKE for each layer. The DYNAX program computes the mixed modal damping for the modal synthesis soil-structure interaction analysis option as follows: a. Various layers in the DYNAX soil model (Figure 3.7-57) are assigned damping values consistent with those obtained in the SHAKE analysis. b. The mixed modal damping is generated using Equation 3.7-3. (Q&R 220.23) The resulting interaction spectra in the two orthogonal X and Y directions at the foundation of the structures for one and two unit building models are obtained for both OBE and SSE earthquakes. These may be found in Calculation SDQ51-14AS02. The horizontal design response spectra at relevant locations of the structure are generated in the two models using a fixed base model subjected to the corresponding interaction time history. This decoupled fixed base analysis is justified as the foundation torsion, and rocking is found to be insignificant.

Forces in the shear walls are generated by subjecting the two fixed-base building models to the corresponding 15% widened interaction response spectrum. Shear walls are finally designed for the maximum of the two unit building models. Two-dimensional finite element models shown in Figures 3.7-77 and 3.7-78 were used for the circulating water screen house for the coupled soil-structure interaction analysis. The computer program LUSH (described in Appendix C) was used for this analysis. For excitation in the vertical direction, a lumped mass multiple spring model as shown in Figure 3.7-68 is used instead of the finite element soil model. The portion of soil from foundation to rock level is modeled as a prismatic column of soil equal in area to the area of the basemat. The layering scheme of the soil is the same as that used for horizontal soil-structure interaction analysis. The Young's modulus E for the soil layers in the vertical direction is calculated from

the strain compatible shear modulus values obtained from the horizontal SHAKE analysis using the relation

()µ+=1G2E where µ = Poisson's ratio. The SHAKE program is used to obtain a compatible rock motion for both OBE and SSE earthquakes (using the program to solve the problem of compression wave propagation instead of shear wave) and specifying the design vertical time history at the free field surface level. Each layer of soil is then represented by an axial spring with its mass lumped at its two ends. The stiffness of each of these axial springs is computed as (AE/L) where A is the surface area of the layer (equal to the building basement area) and L is the thickness of that layer. The above soil model, together with the vertical building models (one and two unit) described in Figures 3.7-69 and 3.7-70 are analyzed for OBE and SSE earthquakes using the program DYNAS. The interaction spectra at the foundation of the building are generated by subjecting the coupled soil structure model to the compatible rock motion. These are shown in Figures 3.7-71 through 3.7-74. The vertical design spectra are at relevant locations of the structure are CPS/USAR CHAPTER 03 3.7-13 REV. 11, JANUARY 2005 generated for the two models, using a fixed-base model subjected to the corresponding vertical interaction time history. All the power block structures in CPS are supported by a single common basemat and interconnected by shear walls in both the north-south and east-west directions. This structural configuration results in a small height-to-base dimension and also leads to a large torsional moment of inertia. This leads to relatively small rocking and torsional motions. To assure that the torsional and rocking motions are insignificant, thus justifying the use of a decoupled fixed base model, the response obtained from the coupled soil-structure interaction model were compared to those obtained from the uncoupled fixed base model. Typical comparison of the coupled vs. uncoupled model response are provided in Figures 3.7-96 (el. 781 feet 0 inch), 3.7-97 (el. 825 feet 0 inch), and 3.7-98 (el. 874 feet 0 inch). The locations of these slabs are shown in Figure 3.7-29. It can be observed that the two models lead to approximately the same response, thus justifying the use of the decoupled fixed base model. The CPS structural layout is such that the horizontal excitation leads primarily to shear and bending in shear walls, columns, and the containment. The vertical excitation leads to axial loads in shear walls and columns and bending of floor slab panels. In the absence of any significant rocking, as shown above, the vertical and horizontal responses are uncoupled, and thus can be computed using separate horizontal and vertical models. For the CPS project, separate vertical and horizontal models are thus used. It should be noted that even though the horizontal and vertical excitation responses are obtained using separate models, all structures, piping, and equipment are designed for the combined effects of the three components of earthquake. (Q&R 220.25) 3.7.2.5 Development of Floor Response Spectra 3.7.2.5.1 Introduction If a structure is subjected to an earthquake, the base of a subsystem (or equipment) mounted on a floor slab or wall experiences the motion of the slab or wall. This motion may be significantly different from the input motion at the base of the structure. Therefore, the response spectra used in the analysis of the structure are not directly applicable to the analysis of subsystems mounted in the structure unless the subsystem element is modeled in the dynamic model of the structure. Also, unless the subsystem element is a rigid mass, rigidly connected to the slab or wall, the motion of the subsystem is different from the motion of the slab or wall, because the subsystem element is a flexible elastic system which responds dynamically to the motion of the slab. For these reasons, the motion experienced by a subsystem is the structure's base excitation modified as a function of the structure's characteristics, and the mode of attachment to the structure. To establish explicit slab or wall motions applicable to development of subsystem design criteria, time history forcing functions are used to excite the building models used in the system analysis. Resulting time history slab or wall motions are used to generate response spectra for the analysis of subsystems supported in the building. Final design response spectra are obtained by enveloping the response spectra from the analyses of the single unit and two unit systems.

CPS/USAR CHAPTER 03 3.7-14 REV. 11, JANUARY 2005 3.7.2.5.2 Horizontal Response Spectra Time history analyses of each building system are performed on the horizontal seismic models as discussed in Subsection 3.7.2.1. The following general procedure was used to develop the horizontal seismic subsystem input: a. The responses at each slab of interest were obtained by exciting the structure separately along the two principal axes (X and Y) of the structure. The responses obtained from the two components are combined on a square-root-of-the sum-of-the-squares basis (SRSS). The justification of decoupling vertical and horizontal models is provided in response to Question 220.25. It should be noted that even though the horizontal and vertical excitation responses are obtained using separate models, all structures, piping, and equipment are designed for the combined effects of the three earthquake components. (Q&R 220.27) b. Response spectra are generated at 0.5, 1, 2, 3 and 4% of critical damping for the OBE and 1, 3, 4, 7, and 10% of critical damping in the main plant and 1, 2, 3, 4, 7 and 10% of critical damping in the containment building for the SSE. c. Response spectra are generated at each slab which supports Seismic Category I subsystems or components. Fifty periods from 0.02 to 2.0 seconds are used to define each spectrum curve. The periods used to generate response spectra at each slab are shown in the attached Table 3.7-15. As can be seen from the table, the periods selected are at very fine intervals and included the natural period of most of the supporting structures. In cases where the period selected for spectrum generation differed from the natural period of the supporting structure, the deviation is negligible (the maximum deviation between the spectral period and the natural period of the supporting structure is only 0.01 second). (Q&R 220.28) d. For the design of subsystems, the peaks of the response spectra are widened by 15% to either side of the peak. 3.7.2.5.3 Vertical Response Spectra The procedure for determining subsystem response spectra in the vertical direction is the same as that for determining responses in the horizontal direction. Response spectra are generated for uncoupled time history motion in the vertical direction at slabs and discrete mass points at the wall/slab junction. For the design of subsystems, the response spectra are widened by 15%

to either side of the peak. 3.7.2.6 Three Components of Earthquake Motion Seismic responses resulting from analysis of systems due to three components of earthquake motion are combined in the following manner as per Regulatory Guide 1.92: 2R2R2RR z y x++= (3.7-1)

CPS/USAR CHAPTER 03 3.7-15 REV. 11, JANUARY 2005 where: R = design seismic response.

R x , R y, and R z are probable maximum, codirectional seismic responses of interest (strain, displacement, stress moment, shear, etc.) due to earthquake excitations in x, y, and z directions, respectively. 3.7.2.7 Combination of Modal Responses 3.7.2.7.1 Systems Other Than NSSS When a response spectrum method of analysis is used to analyze a system, the maximum response (accelerations, shears, and moments) in each mode is calculated independent of time, whereas actual modal responses are independently time dependent and maximum responses in different modes do not occur simultaneously. Based on Regulatory Guide 1.92 and References 5 and 8, the final response R is computed as: 2/1 N1K N111k1RRk R=== (3.7-2) where: 1 2 11w1kk1k1k++= in which:

()2/1]1[kk 2 k= kd t 2kk+= where k and k are the modal frequency and damping in the kth mode, respectively and t d is the duration of the earthquake. For the time history method of seismic analysis, the displacements, acceleration, shears, and moments due to each mode are added algebraically at each instant of time to obtain the final response. 3.7.2.7.2 NSSS In a response spectrum modal dynamic analysis, if the modes are not closely spaced (i.e., if the frequencies differ from each other by more than 10 percent of the lower frequency), the modal responses are combined by the square root of the sum of the squares (SRSS) method as described in Subsection 3.7.2.7.2.1. If some or all of the modes are closely spaced, a double sum method, as described in Subsection 3.7.2.7.2.2, is used to evaluate the combined response. In a time-history method of dynamic analysis, the algebraic sum at every step is CPS/USAR CHAPTER 03 3.7-16 REV. 11, JANUARY 2005 used to calculate the combined response. The use of the time-history analysis method precludes the need to consider closely spaced modes. 3.7.2.7.2.1 Square Root of the Sum of the Squares Method Mathametically, this SRSS method is expressed as follows:

()2/1 2 n1i iRR== where: R = Combined Response R i = Response in the i th mode n = Number of Modes considered in the analysis. 3.7.2.7.2.2 Double Sum Method This method is defined mathematically as: 2/1 N1k N1sskksRR R=== where: R = Representative maximum value of a particular response of a given element to a given component of excitation R k = Peak value of the response of the element due to the k th mode N = Number of significant modes considered in the modal response combination Rs = Peak value of the response of the element attributed to s th mode where: ()1 2sskksk1ks++= in which:

[]2/1 2 k1kk=

CPS/USAR CHAPTER 03 3.7-17 REV. 11, JANUARY 2005 kdwt 2kk+= where k and k are the modal frequency and the damping ratio in the kth mode, respectively, and t d is the duration of the earthquake. Subsection 3.7.2.7.1 describes the double sum expressions used by Sargent & Lundy to combine modal responses, whereas Subsection 3.7.2.7.2.2 describes the double sum expressions used by General Electric. The differences in these Subsections are minor and of no engineering significance. The following paragraphs summarize the differences and evaluate their significance: a. For the second summation, index variable i is used in Subsection 3.7.2.7.1, and s is used in Subsection 3.7.2.7.2.2. This difference is of no significance because the response R is not dependent on the summation index. b. The coupling term (or ks) appears within the absolute sign in Subsection 3.7.2.7.1 and outside the absolute sign in Subsection 3.7.2.7.2.2. The coupling term (or ks) is always a positive quantity; thus it does not matter whether e k is outside or inside the absolute sign. Both equations will lead to the identical response R. c. In defining the term k Subsection 3.7.2.7.1 uses k, whereas Subsection 3.7.2.7.2.2 uses 'k. It is debatable whether 'k or k should be used in computing 'k; in S&L's opinion the use of 'k in the equation is correct. However, the concern over the use of k or 'k is academic and does not affect the seismic design, because the damping values used in the CPS design are small (1% to 7%). For these small damping values, the variance in design responses computed using k or 'k is of the order of 0.1% and of no engineering significance. (Q&R 220.29) 3.7.2.8 Interaction of Non-Seismic Category I Structures with Seismic Category I Structures When Seismic Category I and non-Seismic Category I structures are integrally connected, the non-Seismic Category I structure is included in the model when determining the forces on Seismic Category I structures. The non-Seismic Category I structure is designed under the criteria that ensure that a failure of any part of the non-Seismic Category I structure does not affect the seismic behavior or structural integrity of Seismic Category I structures or systems. 3.7.2.9 Effects of Parameter Variations on Floor Response Spectra To account for the expected variation in structural properties, damping and soil properties, the peaks of various floor response spectra curves are widened by 15% on the period scale to either side of the peak for horizontal as well as vertical components, as per Regulatory Guide

1.122. 3.7.2.10 Use of Constant Vertical Static Factors The Seismic Category I structures, systems, and components are analyzed in the vertical direction using the methods described in Subsection 3.7.2.1. However, beams in a floor slab are designed using a constant vertical acceleration equal to 1.5 times the acceleration value CPS/USAR CHAPTER 03 3.7-18 REV. 11, JANUARY 2005 corresponding to the fundamental frequency of the beam from the applicable wall response spectrum (see Reference 14). Subsection 3.7.2.10 is in conformance with the provisions of the SRP. There is no deviation between the analysis and design method stipulated in the FSAR and SRP Section 3.7.2-II.1.b, which permits the use of any rational and justifiable equivalent static load method. Justification is given below. 1. SRP Section 3.7.2-II.1.b(3) applies only to the design of floor attached structures, equipment, and components, and is based on a static load method which involves no analysis, i.e., no frequency calculation or modeling of the component. 2. The equivalent static load design method stated in Subsection 3.7.2.10 for design of floor framing is a more comprehensive and realistic method. It involves modeling each main floor framing member and determination of the fundamental frequency of the member, consideration of the source of seismic excitation, and includes the effect of higher mode participation. The adequacy and conservatism of this method has been evaluated by comparing the results with a dynamic analysis for a typical floor framing. The results of this analysis were published in the Proceedings of the ASCE Spring Convention in Dallas, Texas in April 1979 (FSAR Reference 14, Subsection 3.7.5). 3. The justification for using the wall response spectrum rather than the floor response spectrum is as follows: The floor framing members are supported by steel columns. In the vertical seismic model, the columns are included with the walls. The seismic response of the floor framing members is given by the response spectra of the supporting columns. The column response spectrum is given by applicable wall spectrum. Therefore, it is appropriate to use the wall response spectra for the design of the floor framing

members. 4. Since the floor framing member is modeled as a single degree of freedom system, an amplification factor of 1.5 is used to account for higher mode participation. This factor is a conservative value. The behavior of a typical steel floor framing member is close to that of a single degree of freedom system, in which higher mode participation is insignificant. The use of 1.5 as the amplification factor for flexible beams with frequencies lower than 33 Hz will ensure the design adequacy of the equivalent static load method used herein. (Q&R 220.30) 3.7.2.11 Method Used to Account for Torsional Effects The complex building structures with heavy equipment and concrete slabs at the various floor elevations have asymmetrical mass-stiffness distribution. Consequently, the slab rotation about the vertical axis occurs when this type of structure is subjected to lateral loads. This torsional effect in slabs is accounted for by including a torsional degree of freedom in each slab of the horizontal building structure model. In the CPS structural design, the additional 5% accidental torsion was not included because this requirement did not exist when the plant was designed and constructed. However, the CPS structures have been designed to resist large torsional loads due to the asymmetry of the equipment and structural layout. Thus, the additional accidental torsion equal to 5% of the CPS/USAR CHAPTER 03 3.7-19 REV. 11, JANUARY 2005 dimension does not result in significant additional forces in the lateral load resisting structural elements (shear walls). This conclusion is based on a CPS-unique analysis where the SSE design shear wall forces were compared to those resulting from the consideration of accidental torsion as required by the NRC staff. This analysis shows that the shear wall load increased an average of 7% with a range of 0 to 15%. Typical values of these increases are presented in Table 3.7-16. The small magnitude of the additional loads leads us to conclude that the effect of the 5% accidental torsion on the CPS design is insignificant. (Q&R 220.31) 3.7.2.12 Comparison of Responses The time history method of analysis is used to generate acceleration response spectra at the lumped mass locations for the design of the structural subsystems and piping components. The response spectrum method of analysis is used to generate forces and moments for the design of structural components. A comparison of the forces and moments generated by the response spectrum and the time history methods of analysis for the single unit containment wall

due to an east-west SSE excitation is given in Table 3.7-8. For the CPS design, a response spectrum analysis was performed to compute structural forces and moments. A time history analysis was performed to obtain floor response spectra and inter-story drift. No structural forces and moments were computed from the time history method; thus, a comparison of time history and response spectra analysis procedures used on the CPS project are consistent with Regulatory Guide 1.92 requirements. Given the acceptability within the engineering profession of using the time history and response spectra methods for seismic analysis, we do not believe a comparison of forces obtained for the CPS project using the two methods should be required. The lower acceleration values for slabs close to the support obtained with the response spectrum method are to be expected because the acceleration values computed in the response spectrum method are pseudo-absolute values obtained by multiplying the modal displacements (relative to the supports) by the square of the modal frequency, 2. In the time history method, the acceleration values are the absolute values and are obtained by adding the base acceleration to the relative acceleration values computed in the modal method. This condition arises because responses in the modal seismic analysis are expressed by fixed base mode shapes. Because structural forces and moments are a function of the relative displacement of the node points, the response spectra and time history methods lead to very close forces. Reference 2 of Regulatory Guide 1.92 (Reference 8) presents a comparison of the forces obtained from the time history and the response spectra methods using various modal combination rules and a wide range of structural frequencies. Note that with the double sum method (the method used in design), a close comparison between the time history and the response spectrum method results is achieved for a wide range of structural configurations.

(Q&R 220.32) 3.7.2.13 Method for Seismic Analysis of Dams The method of analysis used for evaluating the seismic stability of Seismic Category I dams is described in Subsection 2.5.5.2.4.

CPS/USAR CHAPTER 03 3.7-20 REV. 11, JANUARY 2005 3.7.2.14 Determination of Seismic Category I Structure Overturning Moments The Seismic Category I structure overturning mo ments are determined from the relation of the shear force of the structure and the height of the structure for each mode separately. The overturning moments for each mode are then combined by the double sum method to determine the probable maximum overturning moment in the structure. 3.7.2.15 Analysis Procedure for Damping In case of structures with components of different damping characteristics, there are two approximate techniques of computing composite modal damping values to lead to a normal mode solution. These are based on weighting the damping factors according to the mass or the stiffness of each element. The two formulations are:

()()()()()()j T j n1i j i i T j M M j== (3.7-3) ()()()()()()j T j n1i j i i T j K K j== (3.7-4) where: n = total number of components, j = composite modal damping for mode j, i = critical modal damping associated with component i,

[j]i = mode shape vector corresponding to element i, region and mode j,

[M]i , [K]i = subregion of mass or stiffness matrix associated with component i, [M] and [K] are the mass and stiffness matrices of the system. In cases where the stiffness and mass matrices are both diagonal, both Equations 3.7-3 and 3.7-4 would give identical results. In complex structural system where the previous condition is not met, the two methods would give different results and it is not possible to project the superiority of one technique over the other. Since both methods provide rational approximate results, Equation 3.7-4 is used in the analysis of fixed-base dynamic models.

CPS/USAR CHAPTER 03 3.7-21 REV. 11, JANUARY 2005 3.7.3 Seismic Subsystem Analysis 3.7.3.1 Seismic Analysis Methods 3.7.3.1.1 Seismic Analysis Methods for Piping Each pipeline is idealized as a mathematical model consisting of lumped masses connected by elastic members. This means that the weight properties of the subsystem have been lumped on discrete joints and the members connecting these joints are assumed to have all geometric and elastic properties but no weight. The piping subsystem is treated as a space frame having six degrees of freedom for each joint (three translations and three rotations). The displacements of a joint in space can be defined by the above-mentioned six degrees of freedom. The stiffness matrix of the piping system is determined using the elastic properties of the pipe. This includes the effects of torsional, bending, shear and axial deformations, as well as changes in stiffness due to curved members. Next, the mode shapes and the undamped natural frequencies are obtained. The dynamic response of the system is calculated by using either the time-history or the response spectrum method of analysis. With the response spectrum method of analysis, when the piping subsystem is anchored and/or supported at points with different excitations, the analysis is performed using the enveloped response spectra of all response spectra which apply. 3.7.3.1.1.1 Modal Method of Analysis The modal method of analysis is divided into three basis steps: a. Generation of stiffness and mass matrices for the complete piping subsystem from its geometrical mechanical property. b. Formulation and solution of the eigenvalue problem to get eigenvalues and eigenvectors (the eigenvalues are the frequencies, and the eigenvectors are the mode shapes of natural vibration of the system). c. Solution of uncoupled equations for response due to the specified excitation. 3.7.3.1.1.2 Stiffness Matrix Generation The stiffness matrix of the piping subsystem depends only on the geometric and elastic properties of the piping subsystems. The joint stiffness matrix is obtained by summing the individual member stiffness matrices connected to that particular joint. 3.7.3.1.1.3 Mass Matrix Generation The mass matrix for a piping subsystem modeled as a lumped mass system with N degrees of freedom is a diagonal matrix of size N x N. The weight components at the joints are computed from the given weight properties of the members connected to the particular joints.

CPS/USAR CHAPTER 03 3.7-22 REV. 11, JANUARY 2005 3.7.3.1.1.4 Differential Seismic Movements of Interconnected Supports Systems that are supported at points which undergo certain displacements due to a seismic event are designed to remain capable of performing their Seismic Category I functions. The displacements obtained from a time-history analysis of the supporting structure cause moments and forces to be induced into the piping system. Since the resulting stresses are self limiting, it is justified to place them in the secondary stress category. Therefore, these stresses exhibit properties much like a thermal expansion stress, and a static analysis is used to obtain them. The analysis of piping subsystems due to relative seismic support motions consists of two phases. In the first phase, the structural time history responses are generated and reduced to a format suitable for piping analysis; in the second phase, the piping responses due to these structural movements are computed. The details of the various steps are as follows: a. The structural building model is analyzed using the DYNAS Program (Appendix C) to obtain three sets of floor displacement time histories, one each for the two horizontal and vertical excitations. These displacement time histories typically have 1,000 discrete points. For each excitation, a selection of floor responses at 100 random time instances for each of the time history sets is done. These random selections are stored in computer files for use in the piping analysis. b. For each excitation, the piping subsystem is analyzed for the 100 relative support displacements obtained in a. above; the responses of that analysis are enveloped and then the enveloped responses are multiplied by 1.3. c. The responses from each excitation are combined by the square root of the sum of the squares method (SRSS). d. The piping secondary stresses are evaluated using the OBE relative displacement. However, in evaluating the maximum support loads, both the OBE and SSE relative displacements are considered as primary loads in support design at the specified service level. The above procedure yields a reasonable estimate of the maximum responses and support reactions due to seismic relative support movements. 3.7.3.1.2 Seismic Analysis Methods for Equipment The qualification procedure for equipment is discussed in Section 3.9.2.2. 3.7.3.2 Determination of Number of Earthquake Cycles 3.7.3.2.1 BOP Piping Five occurrences of Operating Basis Earthquake (OBE) Loadings are assumed for piping fatigue analysis. Each occurrence of an OBE loading event results in 10 equivalent maximum stress cycles (ASME B&PV Code, Section III, Appendix N-1214). A total of 50 maximum stress cycles are used.

CPS/USAR CHAPTER 03 3.7-23 REV. 11, JANUARY 2005 3.7.3.2.2 BOP Equipment For those pieces of equipment where testing is an acceptable method of qualification, the test duration shall simulate the effect of five OBEs followed by one SSE with each test duration at least equivalent to the strong motion component of the earthquake or a minimum of 10 seconds. The approach used is recommended by IEEE 344. For all other BOP equipment, the qualification method is discussed in Section 3.9. 3.7.3.2.3 NSSS Piping and Component 3.7.3.2.3.1 NSSS Piping Sixty-peak OBE cycles are postulated for fatigue evaluation.

3.7.3.2.3.2 Other NSSS Equipment and Components To evaluate the number of cycles which exist within a given earthquake, a typical boiling water reactor building-reactor dynamic model was excited by three different recorded time histories: May 18, 1940, El Centro NS component, 29.4 seconds; 1952, Taft N 69

° W component, 30 seconds; and March 1957, Golden Gate 80 component, 13.2 seconds. The modal response was truncated such that the response of three different frequency bandwidths could be studied, 0+-10 Hz, 10-20 Hz, and 20-50 Hz. This was done to give a good approximation to the cyclic behavior expected from structures with different frequency content. Enveloping the results from the three earthquakes and averaging the results from several different points of the dynamic model, the cyclic behavior as given in Table 3.7-11 was formed.

Independent of earthquake or component frequency, 99.5% of the stress reversals occur below 75% of the maximum stress level, and 95% of the reversals lie below 50% of the maximum stress level. In summary, the cyclic behavior number of fatigue cycles of a component during an earthquake is found in the following manner: a. The fundamental frequency and peak seismic loads are found by a standard seismic analysis. b. The number of cycles which the component experiences are found from Table 3.7-11 according to the frequency range within which the fundamental frequency

lies. c. For fatigue evaluation, 1/2% (0.005) of these cycles are conservatively assumed to be at the peak load, 4.5% (0.045) at three-quarter peak. The remainder of the cycles will have negligible contribution to fatigue usage. The safe shutdown earthquake has the highest level of response. However, the encounter probability of the SSE is so small that it is not necessary to postulate the possibility of more than one SSE during the 40-year life of a plant. Fatigue evaluation due to the SSE is not necessary since it is a faulted condition and thus not required by ASME Section III.

CPS/USAR CHAPTER 03 3.7-24 REV. 11, JANUARY 2005 The OBE is an upset condition and therefore, must be included in fatigue evaluations according to ASME Section III. Investigation of seismic histories for many plants show that during a 40-year life it is probable that five earthquakes with intensities one-tenth of the SSE intensity, and

one earthquake approximately 20%

of the proposed SSE intensity, will occur. To cover the combined effects of these earthquakes and the cumulative effects of even lesser earthquakes, 10 peak stress cycles are postulated for fatigue evaluation. Subsection 3.9.1.1 presents the number of fatigue cycles used in the design of GE-supplied subsystems. 3.7.3.3 Procedure Used for Modeling 3.7.3.3.1 Modeling of the Piping System 3.7.3.3.1.1 Modeling of the Piping System for BOP Systems The continuous piping system is modeled as an assemblage of beams. The mass of each beam is lumped at nodes which are connected by weightless, elastic members representing the physical properties of each segment. Concentrated weights on the piping system, such as

motor operated valves are modeled as lumped masses. The torsional effects of the valve operators or other equipment with offset center of gravity with respect to the centerline of the pipe are included in the analytical model. 3.7.3.3.1.2 Modeling of NSSS Piping Systems 3.7.3.3.1.2.1 Modeling of Piping Systems The continuous piping system is modeled as an assemblage of three dimensional straight or curved pipe elements. The mass of each pipe element is lumped at the nodes connected by weightless elastic member, representing the physical properties of each segment. The pipe lengths between mass points will be no greater than the length which would have a natural frequency of 33 Hz when calculated as a simply supported beam. In addition, mass points are located at all points on the piping system where concentrated weight such as valves, motors, etc. are located and also at points of significant change in the geometry of the system. All concentrated weights on the piping system such as main valves, relief valves, pumps, and motors are modeled as lumped masses. The torsional effects of the valve operators and other equipment with offset center of gravity with respect to center line of the pipe is included in the analytical model. If the torisional effect is expected to cause pipe stresses less than 500 psi, this effect may be neglected. The criteria employed for decoupling the main steam and recirculation piping systems for establishing the analytical models to perform seismic analysis is given below: a. The small branch lines (6-in. diameter and less) are decoupled from the main steam and recirculation piping systems and analyzed separately because the dynamic interaction is insignificant due to the disparity in the masses of the two

lines. b. The stiffness of all the anchors and its supporting steel is large enough to effectively decouple the piping on either side of the anchor for analytic and code jurisdictional boundary purposes. The RPV is very stiff compared to the piping CPS/USAR CHAPTER 03 3.7-25 REV. 11, JANUARY 2005 system and thus during normal operating conditions the RPV is also assumed to act as an anchor. Penetration assemblies (head fittings) are also very stiff compared to the piping system and are assumed to act as an anchor. The

stiffness matrix at the attachment location of the process pipe (i.e., main steam, RCIC, RHR supply or RHR return) head fitt ing is sufficiently high to decouple the penetration assembly from the process pipe. GE analysis indicates that a satisfactory minimum stiffness for this attachment point is equal to the stiffness in bending and torsion of a cantilever equal to a pipe section of the same size as the process pipe and equal in length to three times the process pipe outer

diameter. 3.7.3.3.1.3 Modeling of NSSS Equipment For dynamic analysis, seismic Category I equipment is represented by lumped mass systems which consist of discrete masses connected by weightless springs. The criteria used to lump masses are: a. The number of modes of a dynamic system is controlled by the number of masses used. Therefore, the number of masses is chosen so that all significant modes are included. The modes are considered as significant if the correseponding natural frequencies are less than 33 Hz and the stress calculated from these modes are greater than 10% of the total stresses obtained from lower

modes. b. Mass is lumped at any point where a significant concentrated weight is located. Examples Are: The motor in the analysis of pump motor stand, the impeller in the analysis of pump shaft, etc. c. If the equipment has a free-end overhang span whose flexibility is significant compared to the center span, a mass is lumped at the overhang span. d. When a mass is lumped between two supports, it is located at a point where the maximum displacement is expected to occur. This tends to conservatively lower the natural frequencies of the equipment. Similarly, in the case of live loads (mobile) and a variable support stiffness, the location of the load and the magnitude of support stiffness are chosen so as to yield the lowest frequency content for the system. This is to ensure conservative dynamic loads since equipment frequencies are such that the floor spectra peak is in the lower frequency range. If such is not the case, the model is adjusted to give more

conservative results. 3.7.3.3.2 Field Location of Supports and Restraints In Seismic Category I buildings, only non-Seismic Category I piping has field located supports and restraints. Field locating is done by the contractor in accordance with the design specifications. A walkdown of this piping will ensure that the pipe supports and restraints are installed in accordance with the design specifications and drawings.

CPS/USAR CHAPTER 03 3.7-26 REV. 11, JANUARY 2005 3.7.3.4 Basis for Selection of Frequencies 3.7.3.4.1 Introduction-Frequency Range For seismic response the frequency range of interest is approximately in the range from 1 to 33 hertz. 3.7.3.4.2 Significant Dynamic Response Modes All modes within a frequency range of interest are included in the dynamic analysis. Generally, the number of modes which are to be considered for the analysis of any given subsystem is dependent on the subsystem characteristics and the amplitude/ frequency content of the input forcing functions. The criterion is to choose the number of modes to cover the peak responses of the applicable loads as much as possible to totally represent the actual piping subsystem responses at the peak response frequency ranges. 3.7.3.5 Use of Equivalent Static Load Method of Analysis Static Analysis: If it can be shown that the fundamental natural frequency of the equipment is equal to, or higher than, 33 hertz, a static analysis shall be performed to determine the stresses and deflections due to seismic loads. In this case, the seismic forces shall be determined by multiplying the mass of the subassembly or part of the equipment times the maximum floor seismic acceleration at the base of the equipment (zero period acceleration from the response spectra). These forces shall be applied through the center of gravity of the subassembly or the part of the equipment. The stresses resulting from each force (in each of the three directions) shall be combined by taking the square root of the sum of the squares (SRSS) to yield the seismic stresses. The seismic deflections (deflections due to seismic loads) shall be calculated in the same way. These seismic stresses and deflections shall be added to all stresses and deflections resulting from all applicable loads, to obtain the final resultant stresses and deflections, which shall be compared with the design limits.

Simplified Dynamic Analysis: A simplified dynamic analysis may be performed, for flexible equipment, applying the same method as the static analysis but using different values for the accelerations. The accelerations to be used shall be obtained by multiplying the g values corresponding to the fundamental natural frequency from the appropriate response spectra curves by 1.5. If the fundamental natural frequency is not known, a static analysis using 1.5 times the maximum peak of the applicable floor response spectra, as applied to seismic accelerations, is acceptable. The 1.5 factor will conservatively account for possible participation of higher modes. After this, the analysis will follow the same procedure described for static

analysis. Detailed Dynamic Analysis: When acceptable justification for static analysis cannot be provided, a dynamic analysis shall be required, and unless a conservative factor is used to account for the participation of higher modes, a detailed dynamic analysis shall be performed.

A mathematical model may be constructed to represent the dynamic behavior of the equipment.

The model can be analyzed using the response spectrum modal analysis or time-history (modal or step-by-step) analysis. The maximum inertia forces, at each mass point, from each mode, shall be applied at that point, to calculate the modal stresses and modal deflections. The various modal contributions shall be combined by taking the square root of the sum of the squares of the individual modal stresses or deflections. The case of closely spaced modes shall be defined and considered as stated in NRC Re gulatory Guide 1.92. The stresses and CPS/USAR CHAPTER 03 3.7-27 REV. 11, JANUARY 2005 deflections resulting from each of the three directions shall be combined by taking the square root of the sum of the squares to obtain the seismic stresses and deflections. These seismic stresses and deflections shall be added to all stresses and deflections resulting from all applicable loads and then compared with the design limits. CPS is in compliance with Regulatory Guide 1.92 in regard to closely spaced modes (See Subsections 3.7.2.7 and 3.7.3.7). 3.7.3.6 Three Components of Earthquake Motion Seismic responses resulting from analysis of components due to three components of earthquake motions are combined in the same manner as the seismic response resulting from the analysis of building structures (Subsection 3.7.2.6). 3.7.3.7 Procedure for Combining Modal Responses The method for combining modal responses for systems or subsystems is discussed in Subsection 3.7.2.7. 3.7.3.8 Analytical Procedures for Piping Systems 3.7.3.8.1 Introduction All Seismic Category I piping is seismically analyzed by either a simplified analysis or a multidegree dynamic analysis, depending on its quality group and nominal size.

3.7.3.8.2 Input Criteria Seismic responses resulting from analysis of systems due to three components of earthquake motion are combined in the following manner as per Regulatory Guide 1.92:

2 z 2 y 2 xRRRR++= (3.7-11) where: R = design seismic response.

R x , R y, and R z are probable maximum, codirectional seismic responses of interest (strain, displacement, stress, moment, shear, etc.) due to earthquake excitations in x, y, and z directions, respectively. In cases where more than one response spectrum may be applied to a subsystem, e.g., if the system is supported from locations in the structure having different response spectra, the response spectra used in the analysis of the subsystem will be an envelope of the applicable response spectra. The total seismic loading obtained from the subsystem analysis consists of two parts, the inertial loading and the loading due to differential anchor movement. Determination of the applicable seismic loading (moment range) depends on the stress being checked (per Section III of ASME B&PV code). For example, when analyzing Class 1 piping, one-half of the moment range due to inertial effects only is used when satisfying Equation 9 of NB-3652, and the total moment range due to both inertial and anchor movements is used when calculating for S n and S p of Equations 10 and 11 of NB-3653.

CPS/USAR CHAPTER 03 3.7-28 REV. 11, JANUARY 2005 Similarly for Class 2 and 3 piping stress analysis, the choice of the seismic loading and constituents (inertial or inertial plus differential anchor movements) is determined by the applicable loading condition and the equations being used as per Subsection NC-3652 of

Section III. The loading range due to differential anchor movements is obtained by performing a static analysis of the affected subsystem with the anchor movements acting on each corresponding terminal end. The anchor movements are determined from the seismic analysis of structural systems. See Section 3.7.3.1.1.4. 3.7.3.8.3 Dynamic Analysis Each pipeline is idealized as a mathematical model consisting of lumped masses connected by elastic members. Appendages having sufficient dynamic effects on the piping system, such as motors attached to motor-operated valves, are included in the model. Using the elastic properties of the pipe, the stiffness matrix for the piping system is determined. This includes the effects of torsional, bending, shear, and axial deformations, as well as changes in stiffness due

to curved members. Next, the frequencies and mode shapes for all the significant modes of vibration are calculated. After the frequency is determined for each mode, the corresponding horizontal and vertical spectral accelerations with appropriate damping are read from the appropriate response spectrum curves. For each mode, the horizontal and vertical displacement and acceleration responses are calculated. The resultant displacement and acceleration responses are determined by combining the maximum response for each mode using the square root of the absolute double sum method. The responses are calculated for each of the three orthogonal directions. Finally, the inertial forces for each direction of earthquake motion for each mode are determined. The stresses due to the inertial forces are determined using the square root of the absolute double sum of the moments for each mode.

Horizontal and vertical earthquake excitations are assumed to occur simultaneously.

Calculations outlined in this subsection are performed using the PIPSYS computer program for the analysis of a 3-dimensional piping system. A more detailed explanation of the method used to combine modal responses is provided in Section 3.7.2.7.1. The relative displacement between anchors corresponding to the elevation of seismic supports and the reactor pressure vessel at the elevation of the nozzles is determined from the dynamic analysis of the structures and vessel. The results of the relative anchor-point displacement are used in a static analysis to determine the additional stresses due to relative anchor-point

displacements. 3.7.3.8.4 Allowable Stresses Allowables for stresses in the piping caused by an earthquake are in accordance with Section III of the ASME B&PV code. Allowables for stresses in the earthquake restraint components such as shock suppressors are in accordance with the allowable stress limits that may have been established by ASME Section III for B&PV Code at the time the restraint components are

purchased. 3.7.3.8.5 Amplified Seismic Responses The two horizontal and one vertical response spectrum curves are derived for all floor elevations. These curves are used in the design of the piping and its components.

CPS/USAR CHAPTER 03 3.7-29 REV. 11, JANUARY 2005 3.7.3.8.6 Use of Simplified Dynamic Analysis For Seismic Category I non-Class 1 systems, 2-inch diameter and smaller, and Class 1 systems l-inch diameter, which are located in Seismi c Category I buildings, a simplified dynamic analysis may be used. This spectra includes OBE with 1% critical damping in accordance with Regulatory Guide 1.61. In order to obtain the emergency condition loads, the upset condition loads are always multiplied by minimum factor of 1.5. (Q&R MEB (DSER) 48) This method yields only response due to inertial effects. The effects of dynamic end

displacement must be considered separately. In this method, piping spans between rigid supports and/ or restraints are treated as independent, simply supported beams. No restraint credit is taken for hangers or restraints not offering stiffness in the direction of the seismic excitation. The span period, maximum midspan deflection, allowable midspan deflection, and end restraint forces are determined for a given span length. The maximum midspan deflection and restraint forces are a function of the floor response spectrum of the building structure in the vicinity of the piping. The spectra used are for the OBE with 1% critical damping as per Regulatory Guide 1.61. In the cases where the other dynamic loads affect the piping, their responses are also considered as identified by individual building elevation and associated response spectra. The data described previously are used to: a. assure that seismic stresses, in conjunction with other primary and secondary stresses, are not greater than the allowable, per ASME Section III, Subsection NC/ND-3600; b. assure that seismic deflections are not large enough to cause contact between pipe and surroundings; and c. provide seismic restraint design loads. 3.7.3.8.7 Modal Period Variation The modal period variation has been considered in the derivation of the response spectrum curves by widening the peaks of those curves (Subsection 3.7.2.9). 3.7.3.8.8 Piping Outside the Containment Structure Seismic Category I piping located outside the containment, but not buried, is analyzed for seismic effect and differential seismic movement at support points, containment penetrations and at any entry points into other structures, as specified in Subsections 3.7.3.8.1 through

3.7.3.8.7. 3.7.3.8.9 Seismic Category I Subsystem Equipment and Components The methods of analysis performed for Seismic Category I equipment and components are presented in Subsection 3.9.2.2.

CPS/USAR CHAPTER 03 3.7-30 REV. 11, JANUARY 2005 3.7.3.9 Multiple Supported Components With Distinct Inputs When the component is supported at points with different elevations, the envelope of each applicable elevation response spectrum is developed and conservatively used for the seismic qualification of the component. The criteria used for considering the piping response due to relative seismic support motions is discussed in Subsection 3.7.3.1.1.4. BOP equipment, if supported at multiple and different locations, is analyzed to the upper bound of the envelope of the individual response spectra. In addition, the effect of relative support displacements, if applicable, is considered. The responses due to inertia effect and relative displacements are combined by the square root of the sum of the squares (SRSS) method. (Q&R 220.35) 3.7.3.10 Use of Constant Vertical Static Factors In general, Seismic Category I piping systems are analyzed in the vertical direction using the methods specified in Subsection 3.7.3.1.1, and Seismic Category I equipment is analyzed in the vertical direction using the methods specified in Subsection 3.7.3.5. Vertical static factors used for equipment are discussed in Subsection 3.7.3.5. 3.7.3.11 Torsional Effects of Eccentric Masses All concentrated loads in the piping system such as valves and valve operators are modeled as massless members, with the mass of the component lumped at its center of gravity. A rigid member is modeled connecting the center of gravity to the piping so that the torsional effects of the eccentric masses are considered. For valve/operator assemblies with natural frequencies greater than or equal to 33 Hz, a simplified (though sufficiently adequate) valve/operator assembly model is considered in the piping analysis to account for eccentricities, thus accounting for bending and torsional effects.

Sargent & Lundy has performed a generic study. Representative piping systems were considered using a detailed finite element representation of the valve assembly to account for its flexibility. The results of this study were compared with similar cases where valves were modeled as rigid in the piping analysis. Amplification factors resulting from this comparison will be used to evaluate any flexible valves. Sargent & Lundy will use the results of this study to qualify the flexible valves. (Q&R MEB (DSER 49) 3.7.3.12 Buried Seismic Category I Piping System and Tunnels Many underground elements like piping, tunnels, reinforced concrete electrical cable ducts, etc., of vital importance need to be designed for accidental conditions such as seismic shock waves passing through the soil medium supporting the element. The buried piping was designed using the ASME Section III, 1977 through Winter 1978 Addenda, stress equations (refer to Tables 3.2-1 and 3.2-3 for the applicable safety class and ASME code). The following stresses were calculated for buried elements: a) for straight portions of the element, the axial stress; b) at bends, the axial and bending stresses.

CPS/USAR CHAPTER 03 3.7-31 REV. 11, JANUARY 2005 For the straight section of elements, Newmark (Reference 12) has presented the following relationship for the maximum axial strain in the element: a. When particle displacement, P, is along the direction of propagation of wave, c Vm m= (3.7-12) b. When particle displacement, p, is perpendicular to the direction of propagation of wave, c2 Vm m= (3.7-12) where: m = maximum axial strain in the homogenous element, V m = maximum particle velocity, C = apparent shear wave velocity in soil.

The stress, , is given by E m= (3.7.13) Where E is the modulus of elasticity for the element. The values used for the maximum particle velocity and the apparent shear wave velocity, were V m = 0.4 ft/sec for OBE = 1.0 ft/sec for SSE C = 2,500 ft/sec The maximum particle velocity was chosen based on the equation given in Reference 15.

6 v ad 2= (3.7.15) where a is the acceleration, v the velocity, and d the displacement. Using the Regulatory Guide 1.60 value of 36 in/g for d, this equation gives a velocity of 48 in/sec/g. For 0.25 g, therefore, the maximum particle velocity is 12 in/sec. The value of the apparent shear wave velocity of 2,500 ft/sec was chosen based on the recommended value given in Reference 12; this is a conservative value for the Wisconsinan glacial till and the interglacial zone present at the site. At bends, the stresses due to the moment, M, induced in the element are included in addition to the axial stress given by Equation 3.7.13. This moment is given by Reference 11.

CPS/USAR CHAPTER 03 3.7-32 REV. 11, JANUARY 2005 2 2 k M= (3.7.14) where: smm= bkk o= ()4EI4/k= m = maximum slippage length k o = modulus of subgrade reaction for fill at the bend b = width of element on elastic foundation I = moment of inertia of element The strains in the buried elements were determined using the effects of a shear wave propagating at 45

°F to the buried elements. This maximizes the strains (see Reference 12 from Section 3.7). Piping that enters the building foundation is rigidly connected to the foundation penetration sleeve as shown by Detail of Pipe Attachment (Figure 3.7-99). The pipe is modeled as a beam on an elastic foundation and the pipe stresses are checked for the relative displacements between the supports. Buried electrical duct runs are rigidly connected to electrical manholes as shown by the typical connection detail (Figures 3.7-100 and 3.7-101). Differential movement between manhole and duct is considered in the concrete duct reinforcing steel design by modeling the duct as a beam on an elastic foundation. (Q&R 220.36) 3.7.3.13 Interaction of Other Piping With Seismic Category I Piping The seismic-induced effects of non-Seismic Category I piping on Seismic Category I piping are accounted for by including in the analysis of the Seismic Category I piping a length of the non-Category I piping to the first anchor beyond the point where the change in category occurs. A sufficient number of restraints on the non-Seismic Category I piping are seismically designed.

At least one restraint in each global direction is required. The axial direction restraint can be located on the Seismic Category I piping adjacent to the pipe category change, since this will also restrain the non-Seismic Category I piping in the axial direction. These criteria meet the requirement of Regulatory Guide 1.29. 3.7.3.14 Seismic Analyses for Reactor Internals This mathematical modeling of the RPV and internals consists of lumped masses connected by elastic (linear) members. Using the elastic properties of the structural components, the stiffness properties of the model are determined and the effects of both bending and shear are included.

CPS/USAR CHAPTER 03 3.7-33 REV. 11, JANUARY 2005 Mass points are located at all points of critical interest such as anchors, supports, and points of discontinuity, etc. In addition, mass points are chosen such that the total mass of the structure is generally uniformly distributed over all the mass points and the full range of frequency of response of interest is adequately represented. Further, in order to facilitate hydrodynamic mass calculations, several mass points (fuel, shroud, vessel), are selected at the same elevation. The various lengths of control rod drive housings are grouped into the two representative lengths shown. These lengths represent the longest and shortest housings in order to adequately represent the full range of frequency response of the housings. The high fundamental natural frequencies of the CRD housings result in very small seismic load. Furthermore, the small frequency differences between the various housings due to the length differences result in negligible differences in dynamic response. Hence, the modeling of intermediate length members becomes unnecessary. Not included in the mathematical model are light components such as jet pumps, in-core guide tubes and housing, sparger, and their supply headers. This is done to reduce the com plexity of the dynamic model. If the seismic responses of these components are needed, they can be determined after the system response has been found. The presence of a fluid and other structural components (e.g., fuel within the RPV) introduces a dynamic coupling effect. Dynamic effects of water enclosed by the RPV are accounted for by introduction of a hydrodynamic mass matrix, which will serve to link the acceleration terms of the equations of motion of points at the same elevation in concentric cylinders with a fluid entrapped in the annulus. The details of the hydrodynamic mass derivation are given in Reference 10. The seismic model of the RPV and internals has two horizontal coordinates for each mass point considered in the analysis. The remaining translational coordinate (vertical) is excluded because the vertical frequencies of RPV and internals are well above the significant horizontal frequencies. Furthermore, all support structures and building and containment walls have a common centerline, hence the coupling effects are negligible. A separate vertical analysis is performed. Dynamic loads due to vertical motion are added to or subtracted from the static weight of components, whichever is more conservative. The two rotational coordinates about each node point are excluded because the moment contribution of rotary inertia from surrounding nodes is negligible. Since all deflections are assumed to be within the elastic range, the rigidity of some components may be accounted for by equivalent linear

springs. The seismic analysis of the RPV and internals employs a linear dynamic model consisting of a detailed representation of the RPV and internals combined with an overall model of the RPV support and containment structure. Such a configuration accounts for the dynamic interaction under dynamic loadings such as seismic. The composite dynamic model is referred to as the "primary structure" model. Sufficient details of the RPV internals components are included in the primary structure dynamic model to enable the generation of representative component interface loads. The interface loads are, in turn, applied to more detailed component stress model in which inelastic behavior is allowed per the ASME III Code. The inelastic response of the components do not alter the linear response of the primary structure. On the substructure and component levels, the ASME Code allows stresses above the elastic limits depending on the subsystem. Typical examples are the core support plate and top guide which are traditionally represented by single mass points in the RPV portion of the primary system dynamic models. Such inelastic stresses are quite localized and have insignificant CPS/USAR CHAPTER 03 3.7-34 REV. 11, JANUARY 2005 effect on the linear response of the primary structure. All local inelastic stresses are verified to be within the ASME Code allowables by appropriate detailed substructure or component analyses. (Q&R MEB (DSER) 51) The shroud support plate is loaded in its own plane during a seismic event and is hence extremely stiff. It may therefore be modeled as a rigid link in the translational direction. The shroud support legs and the local flexibilities of the vessel and shroud contribute to the rotational flexibilities and are modeled as an equivalent torsional spring.

The damping values are given in Table 3.7-1.

3.7.3.15 Analysis Procedure for Damping Damping values used for seismic subsystem analysis are in accordance with Subsection 3.7.1.3. Composite damping is not used in the analysis of subsystems. Alternate damping criteria as specified in ASME Code Case N-411 (Reference 25) can be used for piping subsystems within the limitations documented in References 26 and 27. In particular, the following conditions apply for the use of Code Case N-411 damping: 1) Code Case N-411 damping values are applicable only to building-filtered response spectrum loads (seismic and hydrodynamic); the Code Case damping values shall not be used in time-history analyses such as annulus pressurization, hydraulic transient due to mains steam stop valve closure, etc. 2) When using Code Case N-411 damping values in an analysis, they shall be used in their entirety, and shall not be a mixture of Code Case criteria and Regulatory

Guide 1.61 criteria. 3) If, as a result of using Code Case N-411 criteria, pipe supports are moved, modified, or eliminated, any increased piping displacements due to increased system flexibility shall be checked for adequate clearance with adjacent structures, components and equipment. 4) Code Case N-411 damping is limited to frequencies below 33 H

z. 3.7.4 Seismic Instrumentation 3.7.4.1 Comparison with Regulatory Guide 1.12 The following seismic instrumentation program is provided. It is designed in accordance with Regulatory Guide 1.12, "Instrumentation for Earthquakes", for plants with an SSE of less than 0.3 g. The SSE maximum ground acceleration value at the foundation level for Clinton Power Station has been set at 0.25 g. 3.7.4.2 Location and Description of Instrumentation The instrumentation locations have been chosen to allow meaningful correlation between the recorded accelerations and those calculated using the analytical model of the structure. In addition, the quantities and locations of the instruments are in conformity with Regulatory Guide 1.12.

CPS/USAR CHAPTER 03 3.7-35 REV. 11, JANUARY 2005 3.7.4.2.1 Time-History System The triaxial accelerometers are oriented such that the three axes correspond to the major axes of the analytical model used in the seismic design of the station. Each of five triaxial accelerometers provides input into the Central Recording Unit in the main control room. Four of these triaxial accelerometers (b,c,d,e below) are specifically monitored by the Central Recording Unit in the main control room. At an acceleration of 0.02 g along any of the three axes, the Central Recording Unit starts recording. This acceleration is chosen to screen out minor disturbances while at the same time allowing sufficient sensitivity for appreciable tremors. These accelerometers are located as follows: a. A triaxial accelerometer is placed approximately 1640 feet north of the plant east-west baseline and 210 feet west of the plant north-south baseline on a small underground concrete pad.

b. A triaxial accelerometer is located at the containment basemat elevation (712 feet) at the azimuth 90

°, near the containment wall on the north side of the wall separating the auxiliary building and fuel building. The seismic response at this point is the same as the basemat inside the containment. This sensor also provides data input to the response spectrum analyzer described in Subsection

3.7.4.2.3. c. A triaxial accelerometer is located on the containment wall inside of containment at elevation 851 feet, azimuth 90

°. d. A triaxial accelerometer is located on the control building floor at approximate elevation 737 feet near column-row AC-128. e. A triaxial accelerometer is located in containment on the drywell wall, elevation 779' 10", azimuth 90

°. Data from this sensor is automatically recorded by the Seismic Central Recorder and may be viewed by the Seismic Data Analyzer described in Section 3.7.4.2.3. When the Seismic Central Recorder is started for an event, an annunciator in the main control room and an indicating light on the Seismic Warning Panel are actuated. Detailed time-history analysis of the seismic event can be obtained usi ng the Seismic Data Analyzer and/or Seismic Data Printer. The seismic information can be viewed on the Seismic Data Analyzer (PC screen) and/or printed. 3.7.4.2.2 Seismic Switch A triaxial seismic switch is placed at the same location as the triaxial accelerometer located at the containment basemat (elevation 712 feet, azimuth 90

°) just outside the containmentwall (in the auxiliary building). The central control unit actuates an indicating light and an annunciator in the main control room if the zero period acceleration (of the OBE response spectrum at that location) is exceeded in any of the three axes. The axes are oriented identically with those of the time-history accelerograph sensors.

CPS/USAR CHAPTER 03 3.7-36 REV. 11, JANUARY 2005 3.7.4.2.3 Response Spectrum Analysis A passive response spectrum recorder is located at the circulating water screen house (Seismic Category I Structure independent of the power plant structure). This instrument consists of 12 reeds tuned to different frequencies encompassing the significant portion of the seismic design spectrum for the plant. Peak responses are recorded on scratch plates. A Seismic Central Recorder, Seismic Data Analyzer, and Seismic Data Printer capable of computing and plotting desired response spectra is provided in the Main Control Room. When an event of sufficient magnitude occurs, the Seismic Central Recorder is automatically activated. Data from all five accelerometers identified in Section 3.7.4.2.1 are automatically recorded and once down loaded, may be viewed with the Seismic Data Analyzer. The Seismic Data Analyzer compares the sensor data against the analytical response spectra for the sensor location which is permanently stored in the Data Analyzer. If any axis of the sensors exceeds these stored values, an indicating light and an annunciator in the main control room are actuated. One annunciator will inform the operator if any OBE value is exceeded, and another annunciator if any SSE value is exc eeded. The Seismic Data Analyzer provides detailed information as to the frequencies, axis, and sensor which resulted in actuation of the annunciator.

Time-history data from any of the five sensors identified in Section 3.7.4.2.1 can be manually downloaded form the Seismic Central recorder to the Seismic Data Analyzer for viewing or printing. 3.7.4.2.4 Peak Accelerographs Three triaxial peak accelerographs, each of which measures the absolute peak acceleration in three orthogonal directions coinciding with the principal axes of the analytical model, are placed at the following locations: a. Standby Liquid Control Tank (Seismic Category 1 Reactor Equipment), b. Seismic Category 1 piping connected to an RHR heat exchanger, and

c. Diesel generator oil storage tank (Seismic Category I equipment outside of containment). 3.7.4.2.5 Instrument Performance All instruments are designed to perform their functions satisfactorily over the expected range of environmental conditions, including temperature, humidity, pressure and radiation. A panel in the main control room contains the Seismic Central Recorder, indicating lights, Seismic Data Analyzer, and a power supply. The cabinet and this equipment are qualified in accordance with the requirements of IEEE 344. Battery backup power is automatically provided to operate the equipment on loss of the normal power source with the exception of the printer. The printer is not required in the event of power loss because the seismic information is also available from the Seismic Data Analyzer (PC)

CPS/USAR CHAPTER 03 3.7-37 REV. 11, JANUARY 2005 Display. An annunciator in the main control room is actuated on loss of the normal power source. Built-in test equipment is provided to allow complete in-place testing Central Recorder, Seismic Data Analyzer, and the five triaxial accelerometers. 3.7.4.3 Control Room Operator Notification As described previously, annunciators will alert the operator when a. A control unit starts the Central Recorder b. The seismic switch senses that the OBE acceleration has been exceeded. c. After data is downloaded to the Seismic Data Analyzer and has detected that an OBE acceleration for the on-line accelerometers described in Subsection 3.7.4.2.3 has been exceeded. d. After data is downloaded to the Seismic Data Analyzer and has detected that an SSE acceleration for these sensors has been exceeded. e. Loss of the normal power source occurs. The operator can download the data from the Central Recorder to the Seismic Data Analyzer and examine the Seismic Data Analyzer Display and/or the printer output of the time history records in the control room. In addition, the data from the passive response spectrum analyzer in the circulating water screen house and the three peak accelerographs may be examined. If the OBE maximum acceleration has been exceeded, the operator initiates shutdown of the station. Besides the annunciators, the recorded data is available for analysis of the seismic

event. 3.7.4.4 Comparison of Measured and Predicted Responses The measured response spectra will be compared with the corresponding predicted response spectra for the locations noted in Subsection 3.7.4.2. Agreement between the measured response spectra and the predicted response spectra, or measured response spectra being smaller than the predicted response spectra, would authenticate the capability of the plant to continue operation without undue risk to the health and safety of the public. In the event that the measured response spectra greatly exceed the predicted response spectra, additional evaluation will be performed. This evaluation could include detailed analyses using recorded time histories, remodeling and physical inspection.

3.7.5 References

1. N. Mononobe and H. Matsuo, On the Determination of Earth Pressure During Earthquakes, Proceedings of the World Engineering Congress, Tokyo, Japan (1929). 2. H. B. Seed and R. V. Whitman, Design of Earth-Retaining Structures for Dynamic Loads, ASCE 1970, Specialty Conference, Cornell University (June 22-24, 1970).

CPS/USAR CHAPTER 03 3.7-38 REV. 11, JANUARY 2005 3. H. Matsuo and S. O'Hara, Lateral Earth Pressure and Stability of Quay Cells During Earthquakes, Proceedings of the Second World Conference on Earthquake Engineering, Vol. I (1960). 4. H. M. Westergard, Water Pressures on Dams During Earthquakes, Transactions ASCE , Vol. 98 (1933). 5. E. Rosenblueth and J. Elorduy, Response of Linear Systems to Certain Transient Disturbances, Proceedings, Fourth World Conference on Earthquake Engineering, Vol.

1, Santiago, Chile (1969). 6. Seismic Soil-Structure Interaction Analysis of Nuclear Power Plants, SL-3026, Sargent & Lundy, Chicago, Illinois (May 9, 1973). 7. Dynamic Pressure on Fluid Containers, Chapter 6, Nuclear Reactor and Earthquake , TID-7024, U.S. Atomic Energy Commission (August 1963). 8. A. K. Singh, S. L. Chu and S. Singh, Influence of Closely Spaced Modes in Response Spectrum Method of Analysis, ASCE Specialty Conference on Structural Design of Nuclear Plant Facilities, Chicago, Illinois (December 17-18, 1973). 9. ASCE, Manuals of Engineering Practice, No. 42, Design of Structures to Resist Nuclear Weapons Effects, ASCE Committee on Structural Dynamics of the Engineering Mechanics Division, p. 111. 10. L. K. Liu, Seismic Analysis of the Boiling Water Reactor, Symposium on Seismic Analysis of Pressure Vessel and Piping Components, First National Congress on Pressure Vessel and Piping, San Francisco, California (May 1971). 11. H. H. Shah and S. L. Chu, Seismic Analysis of Underground Structural Elements, Journal of Power Division, ASCE, July 1974. 12. W. J. Hall and N. M. Newmark, Seismic Design Criteria for Pipelines and Facilities, Journal of the Technical Councils of ASCE, Proceedings of the American Society of Civil Engineers, Vol. 104, No. TCI, November 1978. 13. R. V. Whitman and F. E. Richart, Jr., "Design Procedure for Dynamically Loaded Foundations," Journal of the Foundation Division, ASCE, Vol. 93, No. SM6, November 1967, pp. 169-193. 14. Y. L. Tien, V. Kumar and S. J. Fang, Design of Composite Floors for Vertical Seismic Load, ASCE Spring Convention and Exhibit, Dallas, Texas, April 25-29, 1979 Reprint No. 2886. 15. W. J. Hall, B. Mohraz, and N. M. Newmark, Statistical Studies of Vertical and Horizontal Earthquake Spectra, Report NUREG-0003, January 1976, for the U.S. Nuclear Regulatory Commission, Washington, DC. 16. IP letter No. U-0416, from G. E. Wuller to J. R. Miller, NRC, dated February 16, 1982. 17. IP letter No. U-0488, from G. E. Wuller to J. R. Miller, NRC, dated May 24, 1982.

CPS/USAR CHAPTER 03 3.7-39 REV. 11, JANUARY 2005 18. IP letter No. U-0374, from J. D. Geier to J. R. Miller, NRC, dated December 3, 1981. 19. IP letter No. U-0446, from G. E. Wuller to J. R. Miller, NRC, dated March 23, 1982.

20. Presentation at the ACRS Subcommittee meetings held at Decatur, Illinois on February 25-26, 1982. 21. Hardin and Drnevich, "Shear Modulus and Damping in Soils: Design Equations and Curves," ASCE J. of the Soil Mech. and Foundations Div., Vol. 98, No. SM7, July 1972. 22. Illinois Power Co. letter No. U-0374, from J. D. Geier to J. R. Miller, NRC, dated December 3, 1981. 23. Dynamic Pressure on Fluid Containers, Nuclear Reactor and Earthquakes TID-7024, USAEC (August 1963). 24. Y. L. Tien, V. Kumar, and S. J. Fang, "Design of Composite Floors for Vertical Seismic Load," presented at the ASCE Spring Convention, Dallas, Texas, April 1977. ASCE paper No. 2886. 25. American Society of Mechanical Engineers, Section III, Division 1, Code Case N-411, "Alternate Damping Valves for Seismic Anlaysis for Classes 1, 2 and 3 Piping Sections. 26. Letter from A. Schwencer (NRC) to F. A. Spangenberg (IPC), dated April 5, 1985,

Subject:

Use of ASME Code Cases N-397 and N-411 for the Clinton Power Station Unit

1, Docket No. 50-461. 27. Letter from A. Schwencer (NRC) to F. A. Spangenberg (IPC), dated July 19, 1985,

Subject:

Use of ASME Code Case N-411 for the Clinton Power Station Unit No. 1, Docket No. 50-461.

CPS/USAR CHAPTER 03 3.7-40 REV. 11, JANUARY 2005 TABLE 3.7-1 DAMPING VALUES DAMPING, PERCENT OF CRITICAL (1) ITEM EQUIPMENT, OR STRUCTURE OBE SSE BALANCE OF PLANT Equipment and large diameter piping systems, pipe diameter greater than 12 inches 2 3 Small diameter piping systems, diameter less than or equal to 12 inches 1 2 Welded steel structures 2 4 Bolted steel structures 4 7 Prestressed concrete structures 2 5 Reinforced concrete structures 4 7 Soil (2) (2) NSSS Welded Structural Assemblies (Equipment and Supports) 2 3 Vital Piping Systems

- Diameter Greater Than 12 in.

2 3 - Diameter Less Than or Equal to 12 in.

1 2 Reactor Pressure Vessel, Support Skirt, Shroud Head, Separator and Guide Tubes 2 4 Fuel 6 6 (1) Alternate critical damping values for piping systems may be used as described in Section 3.7.1.3.2. (2) Since strain-dependent soil properties are used for the soil-structure interaction, no specific damping values are included.

CPS/USAR CHAPTER 03 3.7-41 REV. 11, JANUARY 2005 TABLE 3.7-2 CONTAINMENT MODEL FOR HORIZONTAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS PARTICIPATION FACTORS MODE FREQUENCY (hertz) E-W EXCITATION N-S EXCITATION 1 4.86 26.93 0.00 2 4.87 0.00 -26.91 3 4.98 40.98 -1.55 4 4.98 1.81 35.04 5 5.10 -4.75 -0.01 6 7.68 5.88 -0.00 7 7.80 -0.00 -6.21 8 8.28 -2.77 -0.00 9 8.35 0.00 -2.56 10 11.02 2.67 -0.02 11 11.23 -0.83 15.23 12 13.46 15.24 0.00 13 14.77 0.99 -14.89 14 16.96 21.51 -0.04 15 21.60 0.72 0.61 16 21.79 0.00 0.60 17 24.41 -6.88 0.00 18 24.41 0.00 -6.88 19 26.77 -5.98 0.00 20 26.98 0.00 6.02 21 27.55 0.00 -4.10 22 27.99 5.56 0.00 CPS/USAR CHAPTER 03 3.7-42 REV. 11, JANUARY 2005 TABLE 3.7-3 ONE-UNIT MAIN STRUCTURE MODEL FOR HORIZONTAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS PARTICIPATION FACTORS MODE FREQUENCY (hertz) E-W EXCITATION N-S EXCITATION 1 1.95 -9.00 0.03 2 2.47 -0.06 -9.16 3 4.97 0.85 -0.20 4 6.80 -103.00 8.50 5 7.58 -15.74 -97.69 6 8.75 -24.06 28.23 7 12.82 -9.25 41.82 8 14.72 -37.76 6.50 9 17.57 28.40 12.21 10 18.16 -20.15 0.92 11 20.52 4.10 -46.64 12 22.43 -28.08 0.59 13 23.84 -2.02 3.66 14 27.36 -6.00 -15.65 15 29.42 9.01 8.83 16 30.13 6.32 -19.69 17 33.71 1.02 1.24 CPS/USAR CHAPTER 03 3.7-43 REV. 11, JANUARY 2005 TABLE 3.7-4 TWO-UNIT MAIN STRUCTURE MODEL FOR HORIZONTAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS PARTICIPATION FACTORS MODE FREQUENCY (hertz) E-W EXCITATION N-S EXCITATION 1 1.96 -12.77 .01 2 3.50 -0.21 14.99 3 3.51 0.57 1.52 4 5.49 2.26 -0.04 5 6.60 133.07 -2.48 6 7.62 4.53 133.40 7 9.40 27.83 -9.07 8 12.27 0.30 -40.65 9 13.45 34.63 -0.07 10 15.50 -52.73 0.58 11 17.65 22.52 6.09 12 19.80 3.92 -60.99 13 21.79 28.87 8.28 14 22.71 -24.08 3.59 15 26.89 -1.59 23.14 16 27.97 -2.89 22.40 17 29.99 9.85 8.20 18 31.07 3.04 -2.62 CPS/USAR CHAPTER 03 3.7-44 REV. 11, JANUARY 2005 TABLE 3.7-5 CONTAINMENT MODEL FOR VERTICAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS MODE FREQUENCIES (hertz) PARTICIPATION FACTOR 1 2.00 2.27 2 5.00 -2.58 3 8.00 -3.44 4 10.93 6.08 5 10.94 -6.48 6 12.90 -10.34 7 12.96 2.31 8 12.99 -1.61 9 13.02 2.63 10 13.35 8.81 11 13.54 -21.74 12 13.93 -7.74 13 13.98 2.21 14 14.05 -1.98 15 14.62 -23.10 16 16.20 -26.47 17 16.98 -1.20 18 17.35 14.23 19 18.12 4.48 20 19.80 -2.78 21 19.91 2.59 22 19.97 1.05 23 20.46 -7.16 24 22.09 -2.36 25 24.23 -3.99 26 26.08 -1.75 27 29.25 18.37 28 32.15 -2.51 CPS/USAR CHAPTER 03 3.7-45 REV. 11, JANUARY 2005 TABLE 3.7-6 ONE-UNIT MAIN STRUCTURE MODEL FOR VERTICAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS MODE FREQUENCIES (hertz) PARTICIPATION FACTOR 1 2.00 13.38 2 2.00 7.85 3 3.00 10.19 4 4.00 -13.60 5 4.00 -8.14 6 5.99 14.07 7 6.00 -1.54 8 7.00 -14.30 9 8.97 -16.58 10 8.98 21.48 11 8.99 1.64 12 9.00 5.46 13 11.97 13.77 14 11.98 17.48 15 12.91 -20.80 16 12.93 -22.35 17 13.90 14.03 18 14.67 -34.52 19 14.85 34.37 20 14.98 5.78 21 16.90 33.29 22 17.59 47.56 23 17.7 42.74 24 18.38 -33.69 25 19.00 -61.04 26 19.23 32.77 27 19.96 -6.19 28 20.61 39.36 29 21.28 -45.43 30 22.74 -8.19 31 23.15 10.23 32 24.12 38.00 33 24.92 -7.59 34 26.51 17.71 35 27.20 11.47 36 27.83 -6.42 37 28.85 32.61 38 29.76 -14.67 39 32.55 27.63 CPS/USAR CHAPTER 03 3.7-46 REV. 11, JANUARY 2005 TABLE 3.7-7 TWO-UNIT MAIN STRUCTURE MODEL FOR VERTICAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS MODE FREQUENCY (hertz) PARTICIPATION FACTOR 1 2.00 -9.80 2 2.00 -17.45 3 3.00 13.72 4 4.00 -10.19 5 4.00 17.75 6 5.99 18.63 7 6.00 -22.39 8 6.99 -18.72 9 8.97 -28.17 10 11.96 -18.44 11 12.88 -23.40 12 11.98 29.73 13 13.91 30.63 14 12.92 -15.50 15 14.72 42.69 16 14.82 48.57 17 14.98 7.81 18 16.85 -56.08 19 17.10 71.95 20 17.42 67.11 21 18.55 28.77 22 18.70 69.12 23 19.20 39.32 24 19.96 8.47 25 20.39 32.67 26 21.19 -53.70 27 22.69 -9.60 28 23.08 -6.73 29 24.04 44.60 30 24.89 -12.49 31 26.35 24.21 32 27.15 -11.92 33 27.68 -14.23 34 28.76 38.28 35 29.57 -27.23 36 30.47 19.33 37 32.22 41.22 CPS/USAR CHAPTER 03 3.7-47 REV. 11, JANUARY 2005 TABLE 3.7-8 SSE FORCES AND MOMENTS FOR SINGLE-UNIT CONTAINMENT MODEL RESPONSE SPECTRUM METHOD TIME HISTORY METHOD MEMBER # SHEAR (Kips) MOMENT (Kip-Ft) SHEAR (Kips) MOMENT (Kip-Ft) 88 8.4 x 10 3 1.2 x 10 6 9.8 x 10 3 1.2 x 10 6 89 8.1 x 10 3 1.0 x 10 6 9.1 x 10 3 9.9 x 10 5 90 7.8 x 10 3 8.6 x 10 5 8.4 x 10 3 8.0 x 10 5 91 7.2 x 10 3 7.1 x 10 5 7.5 x 10 3 6.3 x 10 5 92 6.9 x 10 3 5.6 x 10 5 6.4 x 10 3 4.8 x 10 5 93 5.7 x 10 3 4.2 x 10 5 5.3 x 10 3 3.5 x 10 5 94 4.6 x 10 3 3.0 x 10 5 4.2 x 10 3 2.3 x 10 5 95 3.7 x 10 3 1.9 x 10 5 3.3 x 10 3 1.4 x 10 5 96 2.6 x 10 3 1.4 x 10 5 2.3 x 10 3 1.1 x 10 5 102 7.7 x 10 2 5.6 x 10 4 6.6 x 10 2 3.2 x 10 4 (Refer to Figure 3.7-30 for member location)

CPS/USAR CHAPTER 03 3.7-48 REV. 11, JANUARY 2005 TABLE 3.7-9 PARAMETERS FOR ANALYSIS OF ROCK-SOIL-STRUCTURE-INTERACTION (FINITE ELEMENT MODEL)

COHESIONLESS SOIL COHESIVE SOILS COHESIONLESS SOIL COMPACTED STRUCTURAL FILL RECOMPACTED WISCONSINAN GLACIAL TILL OF WEDRON FORMATION TYPE A MATERIAL(AS COMPACTED) RECOMPACTED WISCONSINAN GLACIAL TILL OF WEDRON FORMATION TYPE A MATERIAL (SATURATED) LOESS WISCONSINAN GLACIAL TILL OF WEDRON FORMATION INTER- GLACIAL DEPOSITS SALT CREEK ALLUVIUM INTERGLACIAL SAND DEPOSITS DENSITY (pcf)

Dry density 123 127 128 101 118 115 100 108 Wet density 132 141 144 120 137 131 125 120 POISSON'S RATION Dynamic 0.40 0.40 0.40 0.40 0.40 0.40 0.40 0.40 Static 0.30 0.40 0.40 0.40 0.40 0.40 0.40 0.40 STATIC MODULUS OF ELASTICITY (Es)

In situ modulus (psf) - 8.0 x 10 5 2.0 x 10 5 2.0 x 10 5 13.1 x 10 5 15.1 x 10 5 - - Increase with surcharge

)psf/psf('d dE m s 350 0 0 0 0 0 150 260 DYNAMIC MODULUS OF ELASTICITY (psf) Single amplitude Shear strain = 1.0%

5,600 ()2/1 m' 11 x 10 5 3 x 10 5 3 x 10 5 12 x 10 5 8 x 10 5 2,800 ()2/1 m' 4,200 ()2/1 m' = 0.1% 36,400 ()2/1 m' 39 x 10 5 8 x 10 5 8 x 10 5 34 x 10 5 31 x 10 5 11,000 ()2/1 m' 17,000 ()2/1 m' = 0.01% 95,000 ()2/1 m' 98 x 10 5 34 x 10 5 34 x 10 5 101 x 10 5 84 x 10 5 45,000 ()2/1 m' 62,000 ()2/1 m' = 0.001% 117,600 ()2/1 m' 148 x 10 5 76 x 10 5 76 x 10 5 232 x 10 5 185 x 10 5 53,000 ()2/1 m' 81,000 ()2/1 m' = 0.0001% 126,000 ()2/1 m' 162 x 10 5 95 x 10 5 95 x 10 5 336 x 10 5 280 x 10 5 56,000 ()2/1 m' 84,000 ()2/1 m' STATIC MODULUS OF RIGIDITY (Gs)

In situ modlus (psf) - 3.0 x 10 5 0.7 x 10 5 0.7 x 10 5 4.7 x 10 5 5.4 x 10 5 - - Increase with surcharge

)psf/psf('d dG m s 135 0 0 0 0 0 54 93 DYNAMIC MODULUS OF RIGIDITY (psf) Single amplitude Shear strain = 1.0%

2,000 ()2/1 m' 4 x 10 5 1 x 10 5 1 x 10 5 4 x 10 5 3 x 10 5 1,000 ()2/1 m' 1,500 ()2/1 m' = 0.1% 13,000 ()2/1 m' 14 x 10 5 3 x 10 5 3 x 10 5 12 x 10 5 11 x 10 5 4,000 ()2/1 m' 6,000 ()2/1 m' = 0.01% 34,000 ()2/1 m' 35 x 10 5 12 x 10 5 12 x 10 5 36 x 10 5 30 x 10 5 16,000 ()2/1 m' 22,000 ()2/1 m' = 0.001% 42,000 ()2/1 m' 53 x 10 5 27 x 10 5 27 x 10 5 83 x 10 5 66 x 10 5 19,000 ()2/1 m' 29,000 ()2/1 m' = 0.0001% 45,000 ()2/1 m' 58 x 10 5 34 x 10 5 34 x 10 5 120 x 10 5 100 x 10 5 20,000 ()2/1 m' 30,000 ()2/1 m' ______________________

'm-- mean effective stress (psf).

CPS/USAR TABLE 3.7-9 (CONT'D) CHAPTER 03 3.7-49 REV. 11, JANUARY 2005 COHESIONLESS SOIL COHESIVE SOILS COHESIONLESS SOIL COMPACTED STRUCTURAL FILL RECOMPACTED WISCONSINAN GLACIAL TILL OF WEDRON FORMATION TYPE A MATERIAL (AS COMPACTED) RECOMPACTED WISCONSINAN GLACIAL TILL OF WEDRON FORMATION TYPE A MATERIAL (SATURATED) LOESS WISCONSINAN GLACIAL TILL OF WEDRON FORMATION INTER- GLACIAL DEPOSITS SALT CREEK ALLUVIUM INTERGLACIAL SAND DEPOSITS DAMPING Percent of critical damping single amplitude Shear strain = 1.0%

16 20 20 20 20 20 21 28 = 0.1% 14 9 15 15 9 9 10 13

= 0.01% 6 5 10 10 5 5 3 4

= 0.001% 2 3 6 6 3 3 1 1.5

= 0.0001% 1 2.5 4 4 2.5 2.5 0.5 0.5 NOTES: 1. The static modulus of elasticity values for cohesive soils were calculated based on the constrained modulus derived from the reloading portion of the consolidation curve 2. Pre-Illinoian cohesive deposits include glacial and lacustrine deposits. 3. Pre-Illinoian cohesionless deposits include Mahomet Valley deposits.

4. The selected parameters reflect both the results of geophysical and laboratory tests performed during this investigation and results published and previously developed for similar soils.

CPS/USAR TABLE 3.7-9 (CONT'D) CHAPTER 03 3.7-50 REV. 11, JANUARY 2005 COHESIVE SOIL COHESIONLESS SOIL ILLINOIAN GLACIAL TILL LACUSTRINE DEPOSITS PRE-ILLINOIAN DEPOSITS PRE-ILLINOIAN DEPOSITS ROCK* DENSITY (pcf)

Dry density 138 123 130 107 156 Wet density 150 134 145 126 159 POISSON'S RATIO Dynamic 0.35 0.35 0.35 0.4 0.29 Static 0.35 0.35 0.35 0.4 0.29 STATIC MODULUS OF ELASTICITY (Es)

In situ modulus (psf) 43.6 x 10 5 24.9 x 10 5 42.4 x 10 5 110 x 10 5 0.7 to 3.8 x 10 8 Increase with surcharge

)psf/psf('d dE m s 0 0 0 1100 0 DYNAMIC MODULUS OF ELASTICITY (psf) Single amplitude 3.6 to 7.8 x 10 8 Shear strain = 1.0%

22 x 10 5 22 x 10 5 22 x 10 5 28,000 ()2/1 m' 0 = 0.1% 81 x 10 5 76 x 10 5 70 x 10 5 95,000 ()2/1 m' = 0.01% 270 x 10 5 192 x 10 5 208 x 10 5 174,000 ()2/1 m' = 0.001% 702 x 10 5 392 x 10 5 540 x 10 5 218,000 ()2/1 m' = 0.0001% 1620 x 10 5 648 x 10 5 923 x 10 5 238,000 ()2/1 m' STATIC MODULUS OF RIGIDITY (Gs)

In situ modlus (psf) 16.1 x 10 5 9.2 x 10 5 15.7 x 10 5 40 x 10 5 0.3 to 1.5 x 10 8 Increase with surcharge

)psf/psf('d dG m s 0 0 0 392 0 DYNAMIC MODULUS OF RIGIDITY (psf) Single amplitude 1.4 to 3.0 x 10 8 Shear strain = 1.0%

8 x 10 5 8 x 10 5 8 x 10 5 10,000 ()2/1 m' 0 = 0.1% 30 x 10 5 28 x 10 5 26 x 10 5 34,000 ()2/1 m' = 0.01% 100 x 10 5 71 x 10 5 77 x 10 5 62,000 ()2/1 m' = 0.001% 260 x 10 5 145 x 10 5 200 x 10 5 78,000 ()2/1 m' = 0.0001% 600 x 10 5 240 x 10 5 342 x 10 5 85,000 ()2/1 m' DAMPING Percent of critical damping single amplitude Shear strain = 1.0%

22 20 20 20 1 to 2 = 0.1% 16 9 12 10

= 0.01% 7.5 4.5 7.5 3

= 0.001% 4 3 4 2 = 0.0001% 3 2.5 3 1 ______________________

• These values are valid for strain levels on the order of 10

-4 to 10-5 percent.

CPS/USAR CHAPTER 03 3.7-51 REV. 11, JANUARY 2005 TABLE 3.7-10 (This Table has been Deleted.)

CPS/USAR CHAPTER 03 3.7-52 REV. 11, JANUARY 2005 TABLE 3.7-11 NUMBER OF DYNAMIC RESPONSE CYCLES EXPECTED DURING A SEISMIC EVENT Frequency Band Hz 0+-10 10-20 20-50 Total Number of Seismic Cycles 168 359 643 No. of Seismic

Cycles 0.5% cycles between 75% and 100%

of Peak Loads 0.8 1.8 3.2 No. of Seismic Cycles

4.5% cycles between 50% and 75% of Peak Loads 7.5 16.2 28.9 CPS/USAR CHAPTER 03 3.7-53 REV. 11, JANUARY 2005 TABLE 3.7-12 CIRCULATING WATER SCREEN HOUSE MODEL FOR HORIZONTAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTORS MODE FREQUENCY (Hertz) PARTICIPATION N-S EXCITATION FACTORS E-W EXCITATION 1 10.94 48.50 0.02 2 17.88 -0.03 50.53 3 20.41 8.85 -0.03 4 24.24 -11.13 -0.11 5 31.88 0.43 -7.90 6 36.91 4.38 0.68 7 38.72 -3.94 0.19 8 40.36 1.14 1.45 9 56.19 -5.24 0.14 CPS/USAR CHAPTER 03 3.7-54 REV. 11, JANUARY 2005 TABLE 3.7-13 CIRCULATING WATER SCREEN HOUSE MODEL FOR VERTICAL EXCITATION - MODAL FREQUENCIES AND PARTICIPATION FACTOR MODE FREQUENCIES (Hertz) PARTICIPATION FACTOR 1 3.012 -14.44 2 3.014 0.00 3 3.641 174.00 4 5.023 0.00 5 5.026 6.15 6 6.934 0.00 7 6.956 7.22 8 6.988 0.00 9 6.991 4.23 10 7.031 0.00 11 7.032 -3.31 12 9.891 -7.89 13 9.892 -0.01 14 9.965 -2.41 15 9.969 0.00 16 10.036 1.99 17 10.039 0.00 18 13.363 51.50 19 15.298 0.00 20 15.929 3.90 21 18.634 -8.46 22 18.684 0.00 23 19.843 -3.51 24 19.883 0.00 25 21.843 22.84 CPS/USAR CHAPTER 03 3.7-55 REV. 11, JANUARY 2005 TABLE 3.7-14 (Q&R 220.21)

STRAIN DEPENDENT SOIL PROPERTIES OBE - HORIZ. SSE - HORIZ.

Layer Strain (%) Damp.

Gx10 3 (ksf) Strain (%) Damp.

Gx10 3 (ksf) 1 .00031 .025 10.25 .00076 .028 8.78 2 .00120 .031 7.89 .00370 .037 5.39 3 .00266 .034 6.08 .00895 .044 3.79 4 .00469 .039 4.93 .01607 .054 3.00 5 .01713 .043 4.17 .02539 .062 2.41 6 .00995 .045 3.61 .03599 .069 2.02 7 .01629 .054 2.51 .05715 .077 1.42 8 .01898 .057 2.36 .06389 .079 1.35 9 .02132 .059 2.24 .06768 .079 1.32 10 .02313 .060 2.16 .07575 .082 1.26 11 .00315 .056 16.71 .01056 .076 9.79 12 .00371 .059 15.75 .01217 .079 9.25 13 .00432 .062 14.86 .01388 .082 8.75 14 .00497 .065 14.04 .01569 .085 8.30 15 .00568 .067 13.24 .01758 .088 7.84 16 .00644 .069 12.50 .01957 .090 7.43 17 .00723 .070 11.82 .02161 .093 7.05 18 .00805 .072 11.22 .02370 .095 6.70 19 .00890 .073 10.65 .02581 .097 6.38 20 .00977 .075 10.13 .02789 .098 6.08 21 .01047 .076 9.82 .02990 .100 5.81 22 .01111 .077 9.60 .03131 .101 5.69 23 .01172 .079 9.39 .03260 .102 5.59 24 .01230 .080 9.21 .03384 .104 5.49 25 .01989 .088 6.01 .05419 .107 3.58 27 .02076 .088 5.91 .05793 .109 3.48 28 .02122 .089 5.85 .06124 .111 3.40 29 .02143 .089 5.83 .06394 .112 3.33 30 .02140 .089 5.83 .06607 .113 3.29 31 Bedrock CPS/USAR CHAPTER 03 3.7-56 REV. 11, JANUARY 2005 TABLE 3.7-15 (Q&R 220.28)

PERIODS OF THE RESPONSE SPECTRA NUMBER PERIOD (seconds) 1 0.020 2 0.030 3 0.040 4 0.045 5 0.050 6 0.055 7 0.060 8 0.065 9 0.070 10 0.075 11 0.080 12 0.085 13 0.090 14 0.095 15 0.100 16 0.110 17 0.120 18 0.130 19 0.140 20 0.150 21 0.160 22 0.170 23 0.180 24 0.190 25 0.200 26 0.220 27 0.240 28 0.260 29 0.280 30 0.300 31 0.320 32 0.340 33 0.360 34 0.380 35 0.400 36 0.450 37 0.500 38 0.550 39 0.600 40 0.700 41 0.800 42 0.900 43 1.000 44 1.100 45 1.200 46 1.300 47 1.400 48 1.500 49 1.700 50 2.000 CPS/USAR CHAPTER 03 3.7-57 REV. 11, JANUARY 2005 TABLE 3.7-16 (Q&R 220.31)

COMPARISON OF TYPICAL SHEAR WALL DESIGN BASIS FORCES TO THOSE INDUCED BY THE 5% ACCIDENTAL TORSION FOR SSE SPRING NO. DESIGN LOAD ACCIDENTAL TORSION LOAD % INCREASE X-Direction Shear Walls 1027 15448.0 1974.0 12.8 102023 14692.0 1589.0 10.8 1019 13426.0 1206.0 9.0 204023 12965.0 1146.0 8.8 204005 11201.0 1.4 0.0 1009 10005.0 38.8 0.4 405019 8577.0 751.0 8.8 708001 7516.0 303.0 4.0 307001 6645.0 929.0 14.0 203031 5023.0 342.0 6.8 204015 4608.0 275.0 6.0 507007 3138.0 9.0 0.3 304005 2295.0 248.0 10.8 507021 993.0 46.0 4.6 Y-Direction Shear Walls 304004 15685.0 82.0 0.5 1018 14298.0 1757.0 12.3 203006 13416.0 1186.0 8.8 102002 12484.0 197.0 1.6 2012 10906.0 314.0 2.9 708002 9940.0 376.0 3.8 405002 8459.0 636.0 7.5 405026 7213.0 619.0 8.6 204016 5358.0 533.0 9.9 102018 4308.0 564.0 13.1 203026 2891.0 121.0 4.2 2042 1591.0 153.0 9.6 1060 800.0 85.0 10.6 CPS/USAR CHAPTER 03 A3.7-1 REV. 11, JANUARY 2005 ATTACHMENT A3.7 (Q&R 220.15) Dr. V. P. Drnevich Letter of February 23, 1982 on GRANULAR STRUCTURAL FILL (Text of letter Dr. Vincent P. Drnevich to Dr. Terje Preber on February 23, 1982, on Granular structural fill, Illinois Power Company, Clinton Station, Job No. 5646-017-07). I have received the background information on the granular structual fill which you sent on February 10, 1982. The information was from the Clinton Power Station-Final Safety Analysis Report, Amendment 3, April 1981. The information included: results of in-place density test measurements, particle size distribution curves, and information on mineral content. From these data, I was able to classify the structrual fill material and to establish the parameters from which to estimate initial tangent shear moduli. I have performed resonant column tests in the past on material similar in nature to this structural fill. In addition, a colleague at the University of Kentucky, Dr. Bobby O. Hardin, completed a fairly detailed study on the Shear M odulus of Gravels for the U.S. Air Force. The final report on Contract F29601-73-0-0064, September 1973, was used to support my calculations to estimate the initial tangent shear modulus for this structural fill. The process of estimating the initial tangent shear modulus is necessary before one can independently establish valves of K2. Shear moduli were calculated by use of two independent empirical equations; one from the State-of-the-Art paper by Hardin in the ASCE Specialty Conference on Earthquake Engineering and Soil Dynamics in Pasadena in 1978 and the other

from the above reference report on gravels. In both methods, the significant parameters were varied to ascertain the sensitivity of shear modulus (and K

2) to the parameters. From these calculations, it is quite evident that a K 2-value of 100 is a very reasonable and realisitic value to use for design purposes. I would expect that if very accurate insitu seismic tests were to be performed on this structural fill, that the values of K 2 back calculated from the measured shear wave propagation velocities would be approximately 100. I am pleased to be of assistance to you on this matter. If clarification on any of the above items is needed, I would be happy to provide it. (The four graphs submitted with the Q&R will not be included in the USAR).

CPS/USAR CHAPTER 03 A3.7-2 REV. 11, JANUARY 2005 1)TABLE A3.7-1 EVALUATION OF GEOPHYSICAL DATA POISSONS RATIO SITE' SOIL CONDITIONS BLOW COUNT DEPTH RANGE (FT)

COMPRESSIONAL WAVE VELOCITY RANGE (fps) SHEAR WAVEVELOCITY RANGE (fps) RANGE AVERAGE Fermi III Till 30-80 0-30 5500-6900 900-1800 0.485-0.463 0.47 (SDIL) Sterling Till 100/6"-refusal 0-80 7400-7900 2150-2700 0.434-.0454 0.44 (WPHOLE)

Attica Clay 4-8 10-40 5200-6400 800-900 .488-0.490 0.49 (SDIL) Till (ML) 50-100 40-50 6600 1100 0.485 0.48 Sands (SP-GM) 20 50-65 7200 2100 0.454 0.45 Till (ML-SM) 50-150 55-90 8000 1300 0.486 0.49 Bailly G1 Lac Clay 5-30 20-100 5600 900 0.48 0.48 (SDIL) G1 Lac Clay 30-200 20-120 5800-6200 1150-1400 0.47-0.48 0.48 G1 Lac Clay 50-200 120-140 5800 1150 0.48 0.48 Glacial Till 50-150 120-160 6200 1600 0.46 0.46 LaSalle Glacial Till 10-100 40-115 5500-6400 950-1450 0.45-0.485 0.475 (SDIL) (Wisconsinan)

Clinton Glacial Till 12-200/6" 50-170 7500 2000-2100 1 0.46 2 (Illinoian)

1) Calculated 2) Estimated

CPS/USAR CHAPTER 03 3.8-1 REV. 12, JANUARY 2007 3.8 DESIGN OF SEISMIC CATEGORY I STRUCTURES 3.8.1 Concrete Containment 3.8.1.1 Description of the Containment 3.8.1.1.1 General The basic objective of the containment system is to provide the capability, in the unlikely event of the postulated design-basis loss-of-coolant accident (LOCA), to limit the release of fission products to the station site environs so that offsite doses are in compliance with the values specified in 10 CFR 50.67. In addition to the containment, a standby gas treatment system (SGTS) is installed to process any leakage from the containment via a filter or purge system, automatically or manually. To meet the basic safety objective, several subsidiary objectives are achieved by the system or one or more of its components, including the following: a. The containment system has the capability of withstanding the conditions which could result from any of the postulated design-basis accidents for which the containment system is assumed to be functional, including the largest amount of energy release and mass flow associated with the accident. b. The containment system is capable of withstanding the effects of the metal-to-water reaction and other chemical reactions subsequent to any postulated design-basis accident for which the c ontainment system is assumed to be functional, consistent with the performance objectives of the nuclear safety systems and engineered safety features. c. The containment system has the capability to maintain its functional integrity during any postulated design event, including protection against missiles from internal or external sources, excessive motion of pipes, jet forces associated with the flow from the postulated rupture of any pipe within the containment, and actuations of safety/relief valves. d. The containment system has the capability to be filled with water to a level above the active core as an accident recovery method for any postulated design-basis accident in which a breach of the nuclear system primary barrier cannot be sealed. e. The containment system, in conjunction with other nuclear safety systems and engineered safety features, has the capability to limit leakage during any of the postulated design-basis accidents for which it is assumed to be functional, such that offsite doses do not exceed the guideline values. f. In the containment, the suppression pool has the means to rapidly condense the steam flow resulting from the design-basis accident which is the rupture of a main steamline inside the drywell. g. The containment system has the means to conduct the flow from postulated pipe ruptures to the suppression pool, to distribute such flow uniformly throughout the CPS/USAR CHAPTER 03 3.8-2 REV. 11, JANUARY 2005 pool, and to limit pressure differentials between the drywell and the containment during the various post-accident cooling modes. h. Rapid closing, redundant isolation valves will maintain containment leakage at or below permissible limits for all postulated conditions by providing an effective barrier in pipes and ducts that penetrate the containment. i. The containment has the capability to be periodically leak tested as may be appropriate to confirm the integrity of the containment at the peak transient pressure resulting from the postulated design-basis accident. j. The containment system has the capability to store sufficient water to supply the requirements of the core standby cooling system. The suppression pool serves as the principal source of coolant for the low-pressure core spray (LPCS) and the low-pressure core injection (LPCI) mode of the residual heat removal (RHR) system. 3.8.1.1.2 Containment Structure The containment, shown in Figures 3.8-1 and 3.8-2, consists of a right circular cylinder with a hemispherical domed roof and a flat base slab. It is constructed of reinforced concrete and completely lined on the inside of the walls and dome with 1/4-inch stainless steel plate below elevation 735 feet 0 inch and with carbon steel plate of at least 1/4 inch thickness above elevation 735 feet 0 inch. The principal dimensions of the containment are: a. height above basemat: 215 feet 0 inch;

b. inside diameter: 124 feet 0 inch; c. wall thickness: 3 feet 0 inch; d. dome thickness: 2 feet 6 inches; and

e. mat thickness: 9 feet 8 inches. The containment structure supports the polar crane, galleries, and the access ramp to the refueling floor. The lower section of the containment acts as the outer boundary of the suppression pool. Two double-door personnel locks, one located at the refueling floor and the other located at the grade floor, permit access to the containment. An equipment hatch is located at the grade floor. The equipment hatch is sealed during normal operation, or at other times when primary containment is required. The containment wall is reinforced in the hoop, diagonal and meridional directions as shown in Figure 3.8-3. Wall reinforcement is deflected around small penetration sleeves to account for localized stress concentrations. The wall around the equipment hatch and personnel locks is thickened to 6 feet 0 inch, and additional reinforcement is provided. Reinforcing details around these penetrations are shown in Figure 3.8-4. Tangential and transverse shear reinforcement

are provided where necessary.

CPS/USAR CHAPTER 03 3.8-3 REV. 11, JANUARY 2005 The dome is reinforced in two directions as shown in Figure 3.8-3. Orthogonal grid type reinforcement is provided within a radius of 45 feet from the apex of the dome. The remaining portion of the dome is reinforced in the hoop and meridional directions. The containment base slab is continuous with the adjacent auxiliary and fuel building base slabs and is reinforced at top and bottom with reinforcing steel as shown in Figures 3.8-5 through 3.8-9. 3.8.1.1.3 Containment Penetrations To maintain the containment pressure boundary, containment penetrations have the following design characteristics: a. capability to withstand peak transient temperatures; b. capability to withstand the forces caused by the impingement of the fluid from the break of the largest local pipe or connection without failure; c. capability to accommodate the thermal and mechanical stresses which may be encountered during all modes of operation without failure; d. capability to withstand the design pressure; and e. capability to act as a pipe support. The approximate number and sizes of the principal containment penetrations are shown in Table 3.8-5. Locations of these penetrations are shown in Figure 3.8-10. 3.8.1.1.3.1 Pipe Penetrations Pipe penetrations for process pipes which pass through the containment and drywell walls may be classified into three types. Type 1 is used for high-energy lines requiring guard pipes when passing through both the containment and drywell walls. Types 2 and 3 are used for the remainder of process pipes which pass through the containment. Figure 3.8-11 shows the basic design of the three penetration types along with the inclined fuel transfer tube detail. Type 1 penetrations consist of a guard pipe anchored at the containment wall and welded to the flued head. The flued head is welded to the process pipe using a gradual buildup weld. The process pipe is allowed free axial thermal movement from the flued head through the drywell. The guard pipe is allowed free axial thermal movement from the containment anchor point through its own sleeve at the drywell wall. Bellows, anchored to the drywell and welded to the guard pipe, will act as a seal for normal drywell environmental conditions. They are designed for thermal guard pipe expansion and relative seismic motion of guard pipe and drywell. Type 2 penetrations consist of a penetration sleeve anchored in the containment and extending to just inside the liner. Full penetration welds are used to weld the flued head to the process

pipe.

CPS/USAR CHAPTER 03 3.8-4 REV. 11, JANUARY 2005 Type 3 penetrations consist of the sleeve anchored in the containment wall and extending just beyond the containment liner. Full penetration welds are used to attach the cover plate to the process pipe. 3.8.1.1.3.2 Electrical Penetrations Dual header plate type electrical penetration assemblies are used to extend electrical conductors through the containment structure pressure boundary. These penetration assemblies are designed, fabricated, tested, and installed in accordance with the requirements of IEEE 317, "Standard for Electrical Penetration Assemblies in Containment Structures for Nuclear Power Generating Stations," dated December 1976. Drawing E27-1310 and Figure 3.8-12 show electrical penetration plans, sections and details. 3.8.1.1.3.3 Personnel and Equipment Access Hatches Two personnel access locks, shown in Figure 3.8-13, are provided for access to the interior of the containment. Each personnel lock consists of an interlocked double door of welded steel assembly. Each door is equipped with a valve for equalizing pressure across the door such that the doors are not operable unless the pressure is equalized. The two doors in each personnel lock are interlocked to prevent both being opened simultaneously and to ensure that one door is completely closed before the opposite door can be opened. An emergency lighting and communication system operating from an external auxiliary energy source is provided within the personnel locks. The equipment hatch, Figure 3.8-13, is fabricated from welded steel and furnished with a double-gasketed flange and bolted dished door. The hatch barrel is welded to the containment liner. Provisions are made to pressure test the space between the double gaskets of the door flanges. The weld seam tests channels at the liner joint and the dished door are provided to monitor any leakage during leak rate testing. Leak testing of the personnel hatches and the equipment hatch is discussed in Subsection 6.2.6.2. 3.8.1.1.3.4 Fuel Transfer Penetration The inclined fuel transfer tube, shown in Figure 3.8-11 along with the three types of process pipe penetrations, penetrates the containment wall through the fuel transfer penetration. This is essentially a 3/4 inch thick carbon steel rolled plate pipe sleeve of 40-inch ID with a 36-inch standard flange on the containment side. The fuel transfer penetration forms a part of the containment boundary. Alternate isolation provisions for this penetration are decribed in Section 9.1.4.2.3.10. 3.8.1.1.4 Containment Liner The containment wall liner is anchored to the wall with structural T sections. Typical wall liner anchorage details are shown in Figure 3.8-14. When a stiffener is cut to avoid interference with an insert assembly, welded studs are provided to restore anchorage of the liner plate.

CPS/USAR CHAPTER 03 3.8-5 REV. 11, JANUARY 2005 Typical spacing of the liner anchors is 15 inches in the containment wall and the dome. Figure 3.8-15 shows details of dome liner anchors and stiffeners. The top of the exposed base slab is lined with 1/2-inch and 1/4-inch stainless steel plate which serves as a leaktight boundary. The drywell wall and the sump floor are anchored through the base liner plate and into the base slab as shown in Figure 3.8-14. The spans of liner panels in the basemat area are: a. pedestal cavity area: 3 feet 0 inch; b. sump floor area: 6 feet 0 inch; and 20 feet 0 inch;

c. Suppression pool area: 3 feet 0 inch to 4 feet 8 1/4 inch (max.). Leak test channels are provided at the liner seams in the suppression pool area and in the containment wall up to elevation 757 feet 0 inch. The containment liner in the wet areas of the suppression pool is of stainless steel to minimize corrosion problems. 3.8.1.1.5 Polar Crane Girder Brackets The polar crane girder is located just below the spring line of the containment and is supported by brackets that are spaced 15

° apart and embedded into the containment wall. Figure 3.8-16 shows the embedment details for the crane girder brackets. 3.8.1.2 Applicable Codes, Standards, and Specifications This section lists codes, specifications, standards of practice, Regulatory Guides, and other accepted industry guidelines which are adopted to the extent applicable, in the design and construction of the containment. The codes, standards and specifications are listed and discussed in Table 3.8-4 and are given with a specification reference number. The reference numbers for the containment are: a. 1 through 5; 7 through 9; b. 11 through 14, 16, 17, 18, 20, 21;

c. 23, 25, 28, 35, 36, 38 through 41;

d. 43, 44, 46, 47, 48, and 50. Table 3.8-9 gives additional details regarding various codes used for design, material, fabrication, and erection of the major structural items within containment. Appendix B gives a detailed discussion on the construction material standards and quality control procedures required during construction. 3.8.1.3 Loads and Loading Combinations The containment structure is designed using the loads, load combinations, and load factors listed in Table 3.8-1.1. Loads and load combinations listed in Table 3.8-1.1 are used for the design of the steel liner and liner anchors, but the load factor for all load cases is 1.0.

CPS/USAR CHAPTER 03 3.8-6 REV. 11, JANUARY 2005 Other steel elements serving pressure vessel functions, such as the hatches, locks, and pipe penetrations are designed for the loads and load combinations in Tables A3.9-6 and A-3.9-7. In addition to the loads defined in these tables, the pipe break loads are included under faulted conditions. Structural steel elements, such as the polar crane brackets, are designed for the loads and load combinations in Table 3.8-2. The seismic loads include the effects of both hydrostatic forces and hydrodynamic forces of the suppression pool water set in motion by seismic accelerations. Also, seismic loads are computed for both the one-and two-unit plant configuration.

Safety/relief valve discharge and LOCA related pool dynamic loads identified in Table 3.8-1.1 are discussed in Attachments A3.8 and A3.9.

The primary parameters used in the containment design are: a. internal design pressure, 15 psig;

b. external design pressure, 3.0 psig;

c. calculated peak pressure, 8.74 psig;

d. test pressure, 17.25 psig; and e. maximum suppression pool water temperature, 185

°F. Time-dependent pressure and temperature loads as discussed in Subsection 6.2.1 are simultaneously applied. The effects of concrete volume changes are minimized by designing the concrete mix for minimum volume change (see Appendix B) and by prescribing construction procedure to minimize differential strains. A one hundred year recurrence interval snow load of 25 psf has been considered in the design. This load is part of the dead load in the Load Combination Table 3.8-1.1, and part of the live load in Tables 3.8-1.2 and 3.8-2. For the extreme precipitation event (PMP) refer to Subsection

2.4.2.3. (Q&R 220.38) 3.8.1.4 Design and Analysis Procedures 3.8.1.4.1 General The containment is analyzed using computer programs which are available in the Sargent & Lundy program library. These programs have all been validated by comparing result with selected problems where a closed-form solution is available or by comparing the solution of a given structure with the solution of the same structure obtained from one or more previously validated programs. These programs have been us ed very effectively on similar containments and have been found to be appropriate for containment analysis. A more detailed description of the various programs named in these subsections can be found in Appendix C.

CPS/USAR CHAPTER 03 3.8-7 REV. 11, JANUARY 2005 Throughout the analysis, special attention is given to the following: a. the intersection between the base slab and the cylinder; b. the intersection between the cylinder and the dome;

c. the area around the large penetrations; d. loading on the base slab from the underlying foundation material; e. stresses due to transient temperature;

f. penetrations and points of concentrated loads; and

g. embedment of polar crane brackets in the containment wall. 3.8.1.4.2 Shell and Base Slab Analysis The method of analysis used is a thin-shell of revolution finite element procedure using the computer program DYNAX. The complete containment with its basemat is modeled with shell elements. The loads applied to the shell model are centerline loads. Consideration is given to the shift of the load from the actual place of application to the centerline of the shell. Overall effects of non-axisymmetric loads such as a pipe break load are analyzed using a series of Fourier harmonics, the summation of which represents the distribution of the load on the structure. Results of the analysis, except for pool dynamic loads, are presented in Figure 3.8-17. Analysis results and design assessment of critical cross sections are presented in Attachment B3.8. The base slab of the containment building is analyzed by a plate finite-element program PLFEM-II. The stiffening contributed by the walls is also included in the finite-element model. Foundation soil is represented by equivalent springs at the nodal points of the basemat elements. A range of soil properties is used to allow for the short-term and long-term characteristics of the soil. The base slab is also analyzed using the computer program CSEF III to confirm the results from the finite-element analysis.

The analytical models used for base slab analysis are shown in Figure 3.8-18. 3.8.1.4.3 Areas Around Large Penetrations To determine the local effects around larger penetrations, such as the equipment hatch, main steam pipes, and personnel locks, the areas around these penetrations are modeled by the finite element program, PLFEM-II. The element nodes lie along the centerline of the containment wall, thus modeling the curvature of the wall. The size of the model is so chosen that the boundary conditions are compatible with those of an axisymmetric shell of revolution. The areas around the large penetrations are designed for loads and loading combinations listed in Table 3.8-1.1.

CPS/USAR CHAPTER 03 3.8-8 REV. 11, JANUARY 2005 3.8.1.4.4 Liner Analysis Loads used for the analysis of steel liner plates are discussed in Subsection 3.8.1.3. Force in a typical liner panel prior to buckling of any panel is determined from the net strain in the liner restrained by the surrounding containment wall. The liner anchorage system is modeled using the computer program LAFD, which calculates the post-buckling force and deflection of the anchors. The post buckling resistance of the panel is evaluated by the method outlined in Reference 1. This method is based on limit analysis and results in an upper bound for the anchor force and displacements. The anchor force-deflection functions are obtained

from tests (Reference 2). The analytical method used to determine the post-buckling force in the liner and its anchor is described in detail in Reference 3. The anchor is sized such that if failure were to occur it would be in the anchor and not the liner. The following cases are considered to produce the worst possible loading conditions on the

anchorage system: a. Case I - an initial inward deflection of 1/16 inch; b. Case II - lower bound yield and 15% decrease in plate thickness of buckled panel; c. Case III - upper bound yield and 15% increase in plate thickness in stable liner panels; and d. Case IV - anchor spacing doubled to simulate failed or missing anchor. This case considers the post-buckling strength of this panel to be zero to maximize the load on the anchor. 3.8.1.4.5 Thermal Analysis The containment is analyzed for thermal effects resulting from both operating and accident design conditions. The containment is designed using the loads and load combinations discussed in Subsection 3.8.1.3. Load combinations are used with and without the temperature loads, and the design is based on the critical case. When considering the thermal effects, the stead y-state gradients (an example of which is shown in Figure 3.8-19) are applied to the design section along with appropriate forces obtained from the containment analysis. The moments resulting from the thermal effects are permitted to change due to cracking of the concrete section. The stresses in the concrete and reinforcing steel are found by using the program TEMCO, which takes into account the extent of cracking of the section. For the transient gradient, an equivalent linear gradient is found by summing moments about the centerline of the section. The section is then analyzed for the equivalent gradient by the same procedure used for the steady-state gradients.

CPS/USAR CHAPTER 03 3.8-9 REV. 11, JANUARY 2005 3.8.1.4.6 Creep and Shrinkage Effects Strains imposed by creep and shrinkage of the containment concrete are included in the design of the steel liner. In addition, minimum reinforcement is ensured throughout the containment structure to carry the effects of creep and shrinkage. 3.8.1.4.7 Suppression Pool Dynamic Load Analysis The methods used for the analysis of safety-relief valve discharge and LOCA-related pool dynamic loads are described in Attachment A3.8. 3.8.1.4.8 Containment Ultimate Capacity This section presents the details of a study to determine the ultimate containment capacity to withstand post-accident internal pressure. 3.8.1.4.8.1 Static Pressure Capacity and Associated Failure Mode The containment is designed for an accident pressure of 15 psig along with appropriate concurrent loads and load factors presented in Table 3.8-1.1. The calculated ultimate pressure capacity of the containment structure considering the liner as a load resisting element is 95 psig. If the liner is not considered for strength, the ultimate capacity of the containment shell and basemat is 75 psig. The upper and lower bound ultimate pressure capacities are based upon a statistical evaluation of mill test reports, as given in Table 3.8-8. For an analysis considering 2 ( = standard deviation) for the containment reinforcing and liner plate meterials, the ultimate pressure capacities are 95

+/-7.5 psig and 75

+/-5.3 psig with and without the liner, respectively. These ultimate pressure capacities correspond to initiation of yielding in the hoop reinforcement around the mid-height of the containment. The ultimate capacity of the major containment penetrations is controlled by the equipment hatch. The static pressure capacity determined on the basis of the buckling of the equipment hatch spherical head is 76 psig. This was determined by using twice the basic allowable buckling stress, as specified in ASME Code Case N-284 Section 1400. To increase the margin of safety a factor of 1.2 can be applied to the analysis reducing the ultimate pressure retaining capability from 76 to 63 psig. Personnel air lock and equipment hatch door seals have been generically tested to 69 psig. 3.8.1.4.8.2 Design Basis The design of the containment structure is based on ASME B&PV Code, Section III, Division 2, 1973 and the details are presented in Subsection 3.8.1.2. The design of the personnel lock and equipment hatch is based on ASME B&PV Code, Section III, Division 1, 1971 with addenda up to and including the Summer 1973 Addenda and the details are presented in Subsection 3.8.1.2. 3.8.1.4.8.3 Probable Failure Modes Ultimate failure of the containment under internal pressure could result from one or a combination of the following:

CPS/USAR CHAPTER 03 3.8-10 REV. 11, JANUARY 2005 a. For Pressure Boundary backed by concrete 1. Fracture of reinforcing steel 2. Punching of concrete around penetrations b. For Pressure Boundary not backed by concrete 1. Fracture of steel components of penetrations such as equipment hatch and personnel locks, etc. 2. Buckling of the equipment hatch spherical head 3.8.1.4.8.4 Failure Criteria for Ultimate Capacity For the purpose of this study, the failure of the containment is defined as attainment of any one of the following limits: a. Stresses in reinforcing steel limited to yield stress; b. strains in liner limited to the requirements of the ASME B&PV Code, Section III, Division 2, Subarticle CC3720, Factored Load Category; c. stresses in portions of pressure boundary not backed by concrete limited to yielding of steel; or d. critical buckling stress in the equipment hatch spherical head not to exceed twice the ASME Code allowable.

These limits are conservative and therefore there is reserve margin beyond the calculated ultimate pressure capacity.

3.8.1.4.8.5 Containment Finite Element Analysis A nonlinear laminated shell finite element model was used to determine the ultimate pressure capacity of the containment. The Sargent & Lundy computer program DYNAX (see Appendix C for a description of the program) is used for the analysis. The program has the capability of analyzing reinforced concrete shells of revolution accounting for cracking of concrete and yielding of steel. A sketch of the finite element model is shown in Figure 3.8-40 (for detailed engineering drawings, see Figures 3.8-3, 3.8-5, 3.8-6, and 3.8-14). The model uses 57 laminated shell elements and 58 nodes to represent the containment basemat, cylinder, and dome. Each shell element is represented as multiple concrete and steel layers. A sketch of the element layering in the sections is also shown in Figure 3.8-40. The number of layers included in any element varies from nine to thirteen. The layering of the elements allows the program to trace the non linear behavior and load distribution due to cracking of concrete and yielding of liner and reinforcement at each element under increasing pressure. Actual mean material properties were used for the concrete, reinforcing steel, and steel liner. The actual material properties were established directly from the mill test reports and concrete CPS/USAR CHAPTER 03 3.8-11 REV. 11, JANUARY 2005 cylinder tests for the material used in the construction of the containment structure. These properties are given in Table 3.8-8. For the nonlinear analysis, the dead load of the structures, hydrostatic load on the suppression pool boundary, and the incremental internal pressure load were applied simultaneously. Even though the liner actually acts as a strength element, two sets of analyses were performed, one considering the liner as strength element and the other deleting the liner as strength element.

The analysis traces the behavior of the entire structure at each pressure increment. 3.8.1.4.8.6 Results of Analysis The results of the finite element analysis include stresses and strains in the steel reinforcing, concrete, and in the liner. These results show the overall axisymmetric response of the containment structure to incremental pressure. Based on a review of the results, the following locations in the basemat, cylinder, and dome are identified as critical sections (see Figure 3.8-39): a. Radial section in the basemat, b. hoop section around the mid-height of containment, and

c. hoop section in the dome. The most critical section was found to be the hoop section around the mid-height of containment where the yielding of reinforcement first occurs at a pressure of 95 psig with liner considered as strength element and at 75 psig when liner is not considered as strength element. At these calculated ultimate pressures, other sections in the containment are at reinforcement stresses which are less than yield and therefore the containment has some reserve margin beyond these calculated pressures. Figures 3.8-41 through 3.8-45 summarize the containment response under increasing internal pressure loading, giving responses with and without the liner as a structural load carrying member. Figure 3.8-41 shows the variation of hoop reinforcing steel stresses in the most critical section around the mid-height of containment. As stated in Subsection 3.8.1.4.8.1, this is the failure mode of the containment which controls the ultimate capacity. As seen from Figure 3.8-41, the hoop reinforcing stresses reach the average yield strength value of 71.1 ksi at 95 psig and 75 psig for the cases with and without the liner, respectively. Also, as seen from Figure 3.8-45, the liner strain at 95 psig is 0.0025 in/in which is smaller than the strain allowed by ASME B&PV Code, Section III, Division 2, Subarticle CC3720, Factored

Load Category. 3.8.1.5 Structural Acceptance Criteria 3.8.1.5.1 Reinforced Concrete Deformations of the structure under factored loads are limited by specifying a maximum allowable concrete strain of 0.002 in./in. and by keeping the strains in the reinforcing steel not CPS/USAR CHAPTER 03 3.8-12 REV. 11, JANUARY 2005 greater than the yield strain. However, the tensile strain in the reinforcing steel is allowed to exceed yield when the effects of thermal gradients through the concrete section are included. For section analysis, the strain in the reinforcing steel and concrete is assumed to be directly proportional to the distance from the neutral axis. The concrete stress-strain relationship is defined by a half parabola whose apex is the point where the strain is 0.002 in./in. and the stress is 0.85 f '

c (f 'c being the specified concrete compressive strength). The tensile strength of the concrete is neglected. Except for the allowable tangential shear stresses listed in Subsection 3.8.1.5.1.1, all reinforced concrete allowables are in accordance with Article CC 3400 of ASME B&PV Code, Sec. III, Div. 2. 3.8.1.5.1.1 Tangential Shear The containment is designed for the peak tangential shear. The tangential shear stress capacity of concrete v c is limited to 40 psi and 60 psi respectively for the service and factored load combinations defined in Table 3.8-1.1. The excess shear is designed to be carried by inclined reinforcement. 3.8.1.5.2 Steel Liner The allowable stresses and strains for the liner plate are limited to values as specified in Article CC 3000 of ASME B&PV Code Sec. III, Div. 2. When subject to SRV discharge loads, the liner plates are designed in accordance with Subsection NE, Section III of the ASME B&PV Code. As described in Subsection 3.8.1.5.2, the containment liner has been designed according to ASME Boiler and Pressure Vessel Code, Section III, Division 2, Article CC-3000. For SRV loadings, the liner in the suppression pool has been designed according to ASME Code, Section III, Division 1, Subsection NE requirements. The design also complies with the applicable provisions of Regulatory Guide 1.57. (Q&R 220.45) 3.8.1.5.3 Steel Pressure-Retaining Components Portions of the containment boundary that are of steel and not backed by concrete, such as the equipment hatch, personnel locks and Code Class MC penetration assemblies including guard pipes, are designed in accordance with Subsection NE, Section III of the ASME B&PV Code. These components are designed for the load combinations shown in Tables A3.9-6 and A3.9-7. The allowable stresses for these load combinations are summarized in the following list of figures from Section III Div. 1 of ASME B&PV Code: a. design conditions, Figure NE-3221-1; b. normal and upset conditions, Figure NB-3222-1;

c. emergency conditions, Figure NB-3224-1; d. faulted conditions, Table F-1322; and e. test conditions, Paragraph NE-3226.

CPS/USAR CHAPTER 03 3.8-13 REV. 11, JANUARY 2005 3.8.1.5.4 Head Fitting Design All head fittings (cover plates of flued heads), which are classified as Seismic Category I components, meet all stress requirements associated with the applicable design, operating, and testing conditions, as stated in the following paragraphs. The allowable (temperature-dependent) stress values, as applicable to items a, b, c, d, and e, are taken from Tables I-1.1 through I-2.2 of the ASME Code, Section III, Div. 1. a. Design Conditions The head fittings are evaluated for design condition loadings and meet all applicable stress requirements set forth in Paragraph NB-3221 of the ASME Code, Section III. These requirements are shown on Figure 3.8-20. The design condition loading is the worst combination of the following loads: 1. design pressure and temperature,

2. weight loads,

3. operating base earthquake (OBE) loads, and

4. hydraulic transients. b. Normal and Upset Conditions The head fittings are evaluated for normal and upset condition loadings and meet all applicable stress requirements described in Paragraphs NB-3222 and NB-3223 of the ASME Code, Section III. These requirements are shown on Figure 3.8-21. The stress evaluation is conducted for the worst combination of the following loads: 1. operating pressure and temperature, 2. weight loads,

3. thermal expansion loads,

4. thermal and pressure transients,

5. OBE loads,

6. pool dynamic effects, 7. hydraulic transients, and 8. loads due to relative dynamic displacements. c. Emergency Conditions The head fittings are evaluated for emergency condition loadings and meet all applicable stress requirements set forth in Paragraph NB-3224 of the ASME CPS/USAR CHAPTER 03 3.8-14 REV. 11, JANUARY 2005 Code, Section III. These stress requirements are summarized in Figure 3.8-22. The stresses are evaluated for the worst combination of the following loads: 1. operating pressure and temperature, 2. weight loads, 3. SSE loads,

4. pool dynamic effects,

5. hydraulic transients. d. Faulted Conditions The head fittings are evaluated for the applicable faulted condition loadings and meet the stress requirements described in F-1324-1, F-1324.6, and Table F-1322 of Appendix F of the ASME Code, Section III, for system inelastic-component analysis. These stress requirements are summarized in Figure 3.8-23. The stresses are evaluated for the following two loading cases:

1. Maximum Operating Pressures and Temperatures, plus loads due to Pipe Rupture and Jet Impingement, when applicable. 2. Process pipe maximum operating pressure applied in the annulus between the process pipe and the penetration sleeve for MC penetration assemblies only. e. Testing Conditions The head fittings are evaluated for test condition loadings in accordance with Paragraphs NB-3226, NB-6222, and NB-6322 of the ASME Code, Section III. 3.8.1.5.5 Penetration Sleeves and Guard Pipes Containment penetration sleeve components including all guard pipes are Code Class MC and are designed and evaluated for design, operating, and testing conditions, in accordance with the following items, a, b, and c. a. Design Conditions Code Class MC penetration sleeves and guard pipes are evaluated for design condition loadings and meet all applicable stress requirements set forth in Paragraph NE-3221 of the ASME Code, Section III. These requirements are summarized in Figure 3.8-20, where the stress values, S m, are the temperature-dependent allowable design stress intensity v alues. These values are taken from Table I-10.0 of Appendix I of the ASME Code, Section III. The stress evaluation is conducted for the loading combination described in item a of Subsection 3.8.1.5.4.

CPS/USAR CHAPTER 03 3.8-15 REV. 11, JANUARY 2005 b. Operating Conditions Code Class MC penetration sleeve components and guard pipes meet all applicable loading and stress requirements set forth in items b, c, and d of Subsection 3.8.1.1.3 for all operating conditions (normal, upset, emergency, and faulted), thus meeting the intent of Paragraph NE-3113 of the ASME Code, Section III. c. Testing Conditions Code Class MC penetration sleeve components are evaluated for testing conditions and satisfy the requirements specified in Paragraphs NE-6222, NE-6322, and NE-3131 of the ASME Code, Section III, for hydrostatic, pneumatic, or leak tests. 3.8.1.5.6 Basemat In addition to the requirements listed in Subsections 3.8.1.5.1 and 3.8.1.5.2, the basemat meets the provisions of Subsection 3.8.5.5 with regard to bearing pressure, overturning, base sliding, and flotation. 3.8.1.6 Materials, Quality Control, and Special Construction Techniques The construction materials and quality control procedures for the containment are specified in Appendix B. Construction dimensional tolerances are specified on fabrication and construction drawings and in fabrication and construction specifications. Criteria for establishing dimensional tolerances include: a. fit-up, and b. safety. Fit-up of components is required to achieve high quality welded and bolted connections and desired interaction of abutting materials and components. Dimensional tolerances which ensure that geometry of completed structures is consistent with theoretical geometry used in design calculations are specified so that structures function safely under design loads. Tolerance for out-of-roundness of the containment liner is specified in fabrication and construction specifications. This dimensional tolerance is set so that the containment vessel will function as a pressure vessel without introduction of stresses not considered in the design.

Dimensional tolerances are set on locations of penetrations and embedment plates attached to the containment liner. These dimensional tolerances are consistent with construction techniques so that fit-up of the components is achieved.

CPS/USAR CHAPTER 03 3.8-16 REV. 11, JANUARY 2005 3.8.1.7 Testing and Inservice Surveillance Requirements 3.8.1.7.1 Structural Acceptance Test The structural acceptance test is performed after the containment is complete with liner, concrete structures, all electrical and piping penetrations, equipment hatch, and personnel locks in place. The structural acceptance test is performed in accordance with Preoperational Test Procedure PTP-SIT-01, Revision 1, which meets the requirements of Article CC 6000 of ASME B&PV Code, Sect. III, Div. 2 (1980 Edition with Summer 1981 Addenda), with the exception that, because of the low design pressure, the pressure is brought up to 115% of the design pressure in only three increments, and tangential deflections at the equipment hatch are not measured because the deflections are negligible. At each pressure level the pressure is held constant for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> before measuring the deflections at locations shown in Drawing S27-1401. The deflection is measured by taut wire extensometers stretched across the containment and kept under a constant tension. At each pressure level, all cracks which exceed 0.01 inch in width and 6 inches in length are mapped at the following four locations: a. near the base-mat/wall intersection, b. midheight of the wall,

c. springline of the dome, and

d. equipment hatch penetration. Table 3.8-6 shows the predicted deflections.

3.8.1.7.2 Leakage Rate Testing Leakage rate testing is discussed in Subsection 6.2.6.

3.8.2 Steel Containment System This subsection applies to the ASME Class MC Components of the concrete containment system described in Subsection 3.8.1. The MC components include the personnel and equipment access hatches, piping and electrical penetrations, and the fuel transfer penetration. 3.8.3 Concrete and Structural Steel Internal Structures of the Containment 3.8.3.1 Description of Internal Structures Internal structures of the containment vessel support and shield the reactor, support recirculation pumps, support piping and auxiliary equipment, form the pressure suppression system, and provide pools and platforms for refueling operations. The internal structures include the following:

CPS/USAR CHAPTER 03 3.8-17 REV. 11, JANUARY 2005 a. reactor shield wall, b. drywell structure,

c. suppression pool weir wall,

d. reactor pedestal, e. miscellaneous platforms and galleries, f. containment pool,

g. refueling floor,

h. equipment rooms,

i. process pipe tunnel, and

j. support system for recirculation pumps. 3.8.3.1.1 Reactor Shield Wall The reactor shield wall (Figure 3.8-25) is an open-ended cylindrical shell 2 feet 0 inch thick placed around the reactor pressure vessel. The primary function of the shield wall is to act as a radiation and heat barrier between the reactor pressure vessel and the drywell wall. It also provides support for pipes, pipe whip restraints, snubbers, and gallery work. The shield wall consists of two concentric steel cylindrical shells, stiffened with radially placed diaphragms and filled with concrete in between the two shells. It is supported on top of the reactor pedestal ring girder (Figure 3.8-26). Openings are provided for pipe penetrations and inservice inspection. The penetration sizes are minimized because inservice inspection is performed inside the shield wall. Since openings for inservice inspection are not in the high radiation area, shielding doors are not provided. Additional stiffeners are provided, wherever nec essary, for various attachments and around openings for local stiffening.

The shield wall is designed as a structural member to support equipment and piping loads as well as to resist pipe rupture, pressure, thermal, and seismic loads. The presence of concrete inside the shield wall is neglected in determining the load-carrying capacity of the wall. 3.8.3.1.2 Drywell Structure 3.8.3.1.2.1 General Description The drywell is a cylindrical reinforced concrete structure which surrounds the reactor pressure vessel and its support structure. The drywell is structurally designed as follows:

CPS/USAR CHAPTER 03 3.8-18 REV. 11, JANUARY 2005 a. to provide structural support to containment pools, main steam tunnel and RWCU compartments; b. to channel steam release from a LOCA through the horizontal vents for condensation in the suppression pool; c. to protect the containment vessel from internal missiles and/or pipe whip; d. to provide anchor points for pipes; and

e. to provide a support structure for the work platforms, monorails, pipe supports, and restraints that are located in the annulus between the drywell and the containment vessel. The inside diameter of the drywell cylinder is 69 feet 0 inch, and the wall thickness is 5 feet 0 inch. The top of the drywell consists of a flat annular slab 6 feet 0 inch thick at elevation 803 feet 3 inches. The drywell wall is rigidly attached to the base slab at elevation 712 feet 0 inch (refer to Figure 3.8-14 for details). A steel head which can be removed to allow access to the reactor is located over the opening in the annular slab. Figures 3.8-1 and 3.8-2 show the drywell structure in plan and elevation.

The drywell is not normally entered during operation, but access is possible during a hot standby with the reactor subcritical.

The lower portion of the drywell wall is submerged in the suppression pool. Three rows of circular vents, 34 vents per row, penetrate the drywell wall below the normal level of the suppression pool. The surfaces of the drywell wall exposed to the suppression pool are lined with stainless steel clad plate 1 inch thick which is designed to act compositely with the drywell wall. Above the level of the suppression pool a carbon steel form plate 1/2 inch thick is provided on the interior surfaces of the cylinder walls and top slab. Structural T's and headed studs are attached to the form plate to provide mechanical anchorage of the plate to the concrete and to stiffen the liner for construction loads. The form plate provides a surface for forming the drywell walls and ceiling and minimizes bypass leakage, if any, through the drywell wall under accident conditions. Details of the reinforcing in the drywell are shown in Figure 3.8-27. Reinforcing is deflected around small penetrations. At large penetrations additional bars are provided to account for concentration of stress. Reinforcing details around these penetrations are shown in Figures 3.8-28 and 3.8-29. 3.8.3.1.2.2 Drywell Penetrations To maintain the drywell pressure boundary, drywell penetrations have the following design characteristics: a. capability to withstand peak transient temperatures; b. capability to withstand the forces caused by the impingement of the fluid from the largest local pipe or connection without failure; CPS/USAR CHAPTER 03 3.8-19 REV. 11, JANUARY 2005 c. capability to accommodate the thermal and mechanical stresses which may be encountered during all modes of operation without failure; and d. capability to withstand the design pressure. The number and sizes of the principal drywell penetrations are shown in Table 3.8-5. Figure 3.8-30 shows locations of these penetrations. 3.8.3.1.2.2.1 Pipe Penetrations Piping penetrations are of the types used in the containment wall and are discussed in Subsection 3.8.1.1.3.1. 3.8.3.1.2.2.2 Suppression Pool Vents There are 102 stainless steel-lined vent openings in three rows of 34 each around the base on the drywell (Figu re 3.8-30). 3.8.3.1.2.2.3 Electrical Penetrations Drawing E27-1310 and Figure 3.8-12 shows a penetration of the general type that is used for the drywell wall. 3.8.3.1.2.2.4 Personnel and Equipment Access Hatches Access to the drywell is provided by the drywell personnel lock, a personnel hatch located in the drywell ceiling, and the drywell equipment hatch shown in Figure 3.8-13. The personnel lock consists of an interlocked, double-door, welded steel assembly. Each door is equipped with a valve for equalizing pressure across the door such that the doors are not operable unless the pressure is equalized. The two doors in the personnel lock are interlocked to prevent both being opened simultaneously, and to ensure that one door is completely closed before the opposite door can be opened. An emergency lighting and communication system operating from an external auxiliary energy source is provided within the personnel lock interior. The equipment hatch is fabricated from welded steel and furnished with a double-gasketed flange and bolted, dished door. Provision is made to pressure test the space between the double gaskets of the door flanges. A shield wall is provided with the same shielding

requirements as the drywell wall. 3.8.3.1.2.2.5 Access for Refueling Operations The drywell head (Figure 3.8-31) is removed during refueling operations. This head is held in place by bolts and sealed with a double seal. It is opened only when the primary coolant temperature is below 212

°F and the core is sub-critical. The double seal provides a method for determining the leak tightness of the seal without pressurizing the drywell.

CPS/USAR CHAPTER 03 3.8-20 REV. 11, JANUARY 2005 3.8.3.1.3 Suppression Pool Weir Wall The suppression pool weir wall, located inside the drywell, acts as the inner boundary of the suppression pool. It is constructed of reinforced concrete and extends from the outer edge of the drywell sump floor. The weir wall is lined with 1/4-inch stainless steel plate on the suppression pool side to protect the concrete from demineralized water. Vertical angles 3 inches x 3 inches x 3/8 inch spaced at 15 inches on center are used to stiffen and anchor the weir wall liner. The weir wall is reinforced (Figure 3.8-27) on both faces with meridional and hoop steel for moments and membrane forces. Additional meridional and hoop steel is provided where required for tangential shear. Radial ties are provided as required for transverse shear. The principal dimensions of the weir wall are: a. Inside diameter: 61 feet; b. Wall thickness: 1 foot 10 inches;

c. Height above basemat: 23 feet 9 inches; and

d. Height above sump floor: 12 feet 7 1/4 inches. 3.8.3.1.4 Reactor Pedestal The reactor pedestal (Figure 3.8-26 and 3.8-32) supports the reactor pressure vessel (RPV) and reactor shield wall. The pedestal shell is a steel structure consisting of two concentric cylindrical shells connected by radially placed steel diaphragms for the entire height of the cylinders. The top of the pedestal consists of a ring girder to which the reactor shield wall is welded. The RPV base is anchored to the ring girder by pretensioned bolts which are designed to carry the loads through friction. Openings are provided through the pedestal for access, control rod drive piping, and nuclear instrumentation. To increase the stability of the structure, the annulus between the steel cylinders is filled with concrete. The concrete is not considered to act compositely with the steel plates. The base of the pedestal is welded to embedded plates anchored in the sump floor with reinforcing bars attached to the plates (Figure 3.8-26). The principal dimensions of the pedestal are: a. outside diameter of top ring girder: 29 feet 10 inches;

b. inside diameter of top ring girder: 16 feet 6 inches;

c. outside diameter of outer steel shell: 29 feet 10 inches;

d. inside diameter of inner steel shell: 18 feet 6 inches; and

e. height above basemat: 31 feet 2 inches.

CPS/USAR CHAPTER 03 3.8-21 REV. 11, JANUARY 2005 3.8.3.1.5 Miscellaneous Platforms and Galleries Miscellaneous platforms and galleries inside the containment serve the dual function of providing access to the electrical and mechanical equipment and providing structural support for this equipment. The platforms and galleries consist of either concrete slabs cantilevered from the drywell wall or structural steel framing supported on containment, drywell, reactor pedestal, and shield walls. Thermal loads in the gallery framing are considered for those beams where thermal expansion is a concern. The layout and configuration of the framing is such that no significant radial thermal loads are imposed on any of the walls. 3.8.3.1.6 Containment Pool The containment pool supported on the drywell walls has the following functions: a. to provide shielding when the reactor is in operation;

b. to provide storage space for the dryer and separator assemblies; and

c. to provide an area for fuel transfer during refueling. This pool forms a rectangular box across the top of the drywell, which is integrated into the design of the top of the drywell. The weight of the pool is transmitted to the foundation mat through the drywell walls. The interior of the pool is lined with stainless steel plate. 3.8.3.1.7 Refueling Floor The refueling floor provides laydown space for reactor components and refueling equipment. The concrete portion of the floor is designed as an integral part of the drywell structure. The grating portions of the floor are supported by structural steel framing which is supported by the containment pool walls and the containment walls. Connections to the containment wall are designed to transfer only vertical and lateral reactions to the containment structure. 3.8.3.1.8 Equipment Rooms Equipment rooms, located near the top of the drywell, are constructed of reinforced concrete. They are designed as an integral part of the drywell structure and are not supported by the containment walls. The rooms are provided with openings that connect directly to the containment volume. The roof of equipment rooms also is part of the refueling floor. 3.8.3.1.9 Process Pipe Tunnel The process pipe tunnel provides shielding for the process piping between the drywell and the containment. It is designed as an integral part of the drywell structure and is constructed of reinforced concrete. The arrangement at the containment wall permits differential movement between the tunnel and the containment. Doorways connect the tunnel to the containment

volume. 3.8.3.1.10 Drywell Sump Floor The drywell sump floor is a thick slab of reinforced concrete which rests on the basemat and supports the suppression pool weir wall and the reactor pedestal. It is anchored through the CPS/USAR CHAPTER 03 3.8-22 REV. 11, JANUARY 2005 containment liner to the basemat with reinforcing bars, as shown in Figure 3.8-14. A stainless steel liner is provided on the suppression pool side to protect the concrete from demineralized

water. The sump floor is reinforced with bars placed in the radial, vertical and hoop directions. The reinforcement details are shown in Figure 3.8-27. The sump floor has the following principal dimensions: a. inside diameter: 18 feet 6 inches:

b. outside diameter: 64 feet 8 inches; and

c. thickness: 11 feet 1-3/4 inches. 3.8.3.1.11 Support System for Recirculation Pumps The recirculation pump and motor assemblies lie above elevation 729 feet 8 inches in the drywell on opposite sides of the reactor pedestal. The pump and motor assemblies are supported by four constant-support spring hangers which attach to built-up box girders at the top. The box girders form a lattice configuration with radial members spanning between the shield and drywell walls, and tangential members bearing on top of the radial - members.

Seven snubbers and two struts are attached to the pump and motor to protect it against dynamic loads. The two struts are attached to the reactor pedestal, three of the snubbers are attached to the reactor shield wall, and the other four snubbers are attached to embedded plates on the drywell sump floor at elevation 723 feet 1 3/4 inch. 3.8.3.2 Applicable Codes, Standards, and Specifications This subsection lists codes, standards of practice, regulatory guides, and other accepted industry guidelines that are adopted, to the extent applicable, in the design and construction of the structures internal to the containment. To eliminate repetitious listing of the codes and standards for each structure, the codes and standards are listed and discussed in Table 3.8-4 and given a reference number. For each structure internal to the containment, the reference numbers are listed in Subsections 3.8.3.2.1 through 3.8.3.2.4. Table 3.8-9 gives additional details regarding various codes used for design, material, fabrication, and erection of the major structural items within containment. Appendix B gives a detailed discussion on the construction material standards and quality control procedures required during construction. 3.8.3.2.1 Reactor Shield Wall and Pedestal The reference numbers are as follows: a. 25b for reactor pedestal; b. 21 and 23 for reactor shield wall; and c. 28, 41, 43, and 47 for both reactor pedestal and reactor shield wall.

CPS/USAR CHAPTER 03 3.8-23 REV. 11, JANUARY 2005 3.8.3.2.2 Drywell Structure The reference numbers are as follows: a. 1 through 5; 7 through 9; b. 11 through 14, 16, 17, 18, 20; c. 21, 23, 25a, 28; d. 35, 38, 41, 43, 44, 46, 47, 48, and 49. The CPS design is in compliance with ACI 349-76 and Regulatory Guide 1.142 with the following clarifications: a. Requirements of Section 10.6.3 of ACI 349 is met as given in ACI 349-80. b. Regarding the bend test requirements for bar nos. 3 through 11, CPS has followed ASTM A615 requirements, which specify slightly bigger bend radii than ACI 349. The following tables gives a comparison of ACI 349 vs. ASTM A615

requirements:

Bar Designation Numbers ACI 349 Radius ASTM A615 Radius 3, 4, 5 3 1/2 d 4 d 6 5 d 5 d 7, 8 5 d 6 d 9, 10, 11 7 d 8 d It should be noted that the primary purpose of the bend test is to assure against cracking during the bending of bars. However, cracking is related directly to the ductility of bars. The ductility of bars at CPS exceeded significantly those required by ACI 349 and ASTM A615. A comparison of the required elongation vs. the actual statistical average elongation, as determined from the Certified Mill Test Reports, is given below:

Bar Designation Number Required Elongation (Percent)

Actual Statistical Average Elongation (Percent) 5 9 12.03 6 9 10.86 7 8 14.44 8 8 12.47 9 7 14.17 10 7 15.10 CPS/USAR CHAPTER 03 3.8-24 REV. 11, JANUARY 2005 Bar Designation Number Required Elongation (Percent)

Actual Statistical Average Elongation (Percent) 11 7 16.17 14 7 15.56 18 7 14.54 Based on the above comparison, the effect of the minor difference in the bend radii requirement, if any, is offset by the much higher elongation of the bars used at CPS. (Q&R 220.49) 3.8.3.2.3 Miscellaneous Platforms and Galleries, Refueling Floor, Equipment Rooms, Suppression Pool Weir Wall, Process Pipe Tunnel, and Structural Support System for Recirculation Pumps The reference numbers are as follows: a. 1 through 5, 7 through 9; b. 11 through 14, 16, 17, 18, 20;

c. 21, 23, 28, 31, 35;

d. 37, 38, 41, 43, 46, 47, and 48. 3.8.3.2.4 Containment Pool The reference numbers are as follows: a. 1 through 5, 7, 8, 9;

b. 11 through 14, 16, 17, 18, 20;

c. 21, 23, 25d, 28, 35;

d. 37, 38, 41, 43, 44, 46, 47, and 48. 3.8.3.3 Loads and Loading Combinations The reinforced concrete internal structures, which include the drywell, suppression pool weir wall, containment pool, equipment rooms, process pipe tunnel, and portions of the refueling floor, are designed using the loads, load combinations, and load factors listed and discussed in Table 3.8-1.1. In the vent area of the drywell, the steel liner is well anchored to the concrete to ensure composite action and is designed using Table 3.8-1.1. The weir wall, the lower portion of the drywell, and the containment pool are designed to resist the effects of both hydrostatic forces and hydrodynamic forces associated with water set in motion by seismic accelerations.

CPS/USAR CHAPTER 03 3.8-25 REV. 11, JANUARY 2005 The drywell is designed for the differential pressure between the drywell and containment as discussed in Subsection 6.2.1. Time-dependent pressure and temperature loads as discussed in Subsection 6.2.1 are simultaneously applied. The upper containment pool is designed for a maximum water temperature of 212

° F. The weir wall and the portion of the drywell located inside the suppression pool are designed for a maximum water temperature of 185

° F. Safety/relief valve discharge loads and LOCA-related pool dynamic loads identified in Table 3.8-1.1 are described in Attachments A3.8 and A3.9. The structural steel elements of the internal structures, which include the reactor pedestal, the reactor shield wall, portions of the refueling floor, the miscellaneous platforms and galleries, and the structural support system for recirculation pumps are designed using the loads and load combinations listed and discussed in Table 3.8-2. The platforms in the annulus between the drywell and the containment are subjected to upward loads from pool swell, P, and seismic effects. These three upward-acting loads form the basis for the design. The differential pressure ( P) and pool swell forces are combined with live, dead, and seismic loads as applicable. The thermal loads associated with the reactor shield wall include temperature gradients under normal operating and accident conditions due to the absorption of gamma and neutron radiation. Internal structures are designed for the reactions of all other structures or equipment that they may support.

The effects of concrete volume changes are minimized by designing the concrete mix for minimal volume changes (see Appendix B) and by prescribing construction procedures to minimize differential strains. Closed compartments are designed to withstand the temperatures and pressures due to the failure of equipment which is inside of the compartments. The reactor shield wall is designed for loads resulting from pipe breaks within the RPV shield wall annulus. These loads are described in Subsection 6.2.1.2.1.2. 3.8.3.4 Design and Analvsis Procedures 3.8.3.4.1 Reactor Shield Wall The reactor shield wall is designed as a cylindrical shell with inner and outer shell plates stiffened by closely spaced vertical stiffener plates. The concrete filling acts mainly as a radiation shield and is not considered as a structural element to carry any load. Two computer programs, DYNAX and SLSAP-4, are used for the analysis of the reactor shield wall. The shield wall is analyzed as a symmetrical shell in the DYNAX run mainly to calculate stresses resulting from thermal and pressure loads. The reactor shield wall is analyzed as an CPS/USAR CHAPTER 03 3.8-26 REV. 11, JANUARY 2005 anisotropic shell for the effects of pipe rupture loads. The finite-element computer program SLSAP-4 is used for this analysis. Appropriate boundary conditions are used at the base connection with the reactor pedestal for both the models. The pipe rupture loads are applied as concentrated loads at the node points of the three-dimensional finite-element model. The stresses due to different loadings are taken from respective computer output and combined manually to arrive at the final design stress. 3.8.3.4.2 Drywell and Attached Structures 3.8.3.4.2.1 General The drywell is treated as a structure with a cylindrical wall and an annular roof, with the containment pool walls, process pipe tunnel and equipment rooms rigidly attached. The steel plates in the vent area are anchored to the drywell wall to ensure composite action and are designed to transfer loads to the foundation. Throughout the analysis, special attention is given to the following: a. the intersection between the base slab and the cylinder;

b. the intersection between the cylinder and the roof; c. the intersection between the pool walls and the roof or cylinder; d. the stresses around penetrations;

e. the stresses caused by transient temperatures in steel liners and concrete; and

f. penetrations and points of concentrated loads. 3.8.3.4.2.2 Shell Analysis The drywell is divided into two portions for analysis, with the match line located near the mid-height of the cylinder at elevation 755 feet 0 inch. A simplified model of the drywell cylinder wall and slab which includes the weight of the compartment walls and slabs and pool walls was analyzed to obtain results in the lower portion. The upper portion is analyzed by applying boundary conditions at the match line such that compatibility of deformations and equilibrium of

forces is maintained. Analysis of the upper portion of the drywell utilizes the computer program SLSAP-4 to determine the distribution of forces. Taking advantage of the symmetry of the structure about the east west axis, half of the upper drywell, including the fuel pools and equipment compartments, is modeled in Figure 3.8-34. Results in the lower portion of the drywell are obtained using the computer programs DYNAX and PLFEM-II. DYNAX is used to analyze the axisymmetric model for both axisymmetric and non axisymmetric loadings. Non-axisymmetric loads are approximated by Fourier series. The effect of discontinuities, e.q. around large penetrations, is considered by analyzing portions of the structure using PLFEM-II.

CPS/USAR CHAPTER 03 3.8-27 REV. 11, JANUARY 2005 The effect of the vents in the suppression pool area is taken into account by modifying the shell stiffness to obtain the force resultants in the area. The stress distribution and the vents are then more closely analyzed for stress concentration using PLFEM-II. Pressure loads applied to the models are centerline loads. Therefore, consideration is given to the shift of the load from the actual place of application to the centerline of the shell. The results of the analysis except for pool dynamic loads are shown in Figure 3.8-33. Analysis results and design assessments of critical cross sections are presented in Attachment B3.8. 3.8.3.4.2.3 Areas Around Large Penetrations To determine the local effects at large penetrations, such as the equipment hatch and main steam pipes, the areas around these penetrations are modeled by a finite-element program, PLFEM-II. The element nodes lie along the centerline of the drywell wall, thus modeling the curvature of the wall. The size of the model is so chosen that the boundary conditions are compatible with those of an axisymmetric shell of revolution. 3.8.3.4.2.4 Thermal Analysis The drywell is analyzed for thermal effects resulting from both operating and accident design conditions. The drywell is designed using the loads and load combinations discussed in Subsection 3.8.3.3. Load combinations are used with and without the temperature loads, and the design is based on the critical case. When considering the thermal effects, the steady-state gradients are applied to each design section along with concurrent axial loads and moments in a design load combination. The stresses in the concrete and reinforcing steel are then analyzed using the computer program TEMCO, which considers the extent of cracking of the section and the resultant change in the thermal moment. For the transient gradient, an equivalent linear gradient is found by summing moments about the centerline of the section. The section then is analyzed for this equivalent gradient by the same procedure used for the steady-state gradients. 3.8.3.4.2.5 Creep and Shrinkage Effects Sufficient reinforcing is provided throughout the drywell structure to carry the effects of creep and shrinkage. 3.8.3.4.2.6 Suppression Pool Dynamic Load Analysis The drywell structure is analyzed for the SRV discharge loads and LOCA related pool dynamic loads as described in Attachment A3.8. 3.8.3.4.3 Reactor Pedestal and Suppression Pool Weir Wall The reactor support pedestal and the suppression pool weir wall are designed as axisymmetric cylindrical shells fixed at their base. Loads from the reactor pressure vessel and the reactor shield wall are applied at the top of the pedestal. Two Sargent & Lundy shell-of-revolution programs, SOR-III and DYNAX, are used for analysis. The seismic and pipe rupture forces transmitted to the pedestal are included in the design as shears and overturning moments.

CPS/USAR CHAPTER 03 3.8-28 REV. 11, JANUARY 2005 The effects of concentrated pipe break load on the weir wall are analyzed using a series of Fourier harmonics, the summation of which represents the distribution of the load on the structure. The capacity of the section under combined loads is checked using the program TEMCO. Thermal analysis is performed as discussed in Subsection 3.8.3.4.2.4. The weir wall thickness and reinforcement are so proportioned that the pressure suppression efficiency is not impaired by the deflection of the weir wall under design loads. 3.8.3.4.4 Refueling Floor, Miscellaneous Platforms and Galleries, and Support System For Recirculation Pumps The platforms, galleries, and structural supports for the recirculation pumps are designed using conventional elastic design methods. 3.8.3.5 Structural Acceptance Criteria 3.8.3.5.1 Reinforced Concrete Deformations of the drywell structure, containment pools, and equipment rooms under factored load conditions are limited by specifying a maximum allowable concrete strain of 0.002 in./in. Yielding of the reinforcing steel in tension is allowed only when the effects of thermal gradients are considered. For section analysis, the strain in the reinforcing steel and concrete is assumed to be directly proportional to the distance from the neutral axis. The concrete stress-strain relationship is defined by a half parabola whose apex is the point where the strain is 0.002 in./in. and the stress is 0.85 f c , where f c is the specified concrete compressive strength. The tensile strength of the concrete is neglected. Except for the allowable tangential shear stresses listed in Subsection 3.8.3.5.1.1, reinforced concrete allowables for the drywell structure, containment pools, equipment rooms, and sump floor and weir walls are in accordance with ASME B&PV Code, Section III, Div. 2. The stresses and strains in all other reinforced concrete internal structures are limited to those specified in ACI 318. Serviceability checks are made in accordance with ACI 318 to assure crack control and to keep deflections below the limits prescribed in ACI 318 or to the manufacturer's recommendations for equipment supported by the reinforced concrete. The factors of safety against material strength are contained in the load factors in Table 3.8-1.2 and in the capacity reduction factors () in ACI 318. 3.8.3.5.1.1 Tangential Shear for Drywell The allowable tangential shear stress for concrete, under the factored load conditions defined in Table 3.8.1-1, is calculated in accordance with Section 11.16 of ACI 318. The shear stress carried by concrete under service load conditions is 50% of the value allowed for the factored load conditions. Reinforcement is provided to carry the shear in excess of the concrete shear capacity in accordance with Section 11.16 of ACI 318 except that for service load conditions, 50% of f y is used.

CPS/USAR CHAPTER 03 3.8-29 REV. 11, JANUARY 2005 The original drywell design for tangential shear is based on Section 11.16 of ACI-318. This question was also addressed in CPS-FSAR Amendment 16, dated July 6, 1974, in response to Question 3-71, Page 3-120 (see below). Tangential shear design for the drywell is also in compliance with Section 11.10 of ACI-349 which the NRC is currently accepting, as stated in Question 220.49. Structural acceptance criteria for drywell stated in Subsection 3.8.3.5.1 is in agreement with the criteria given in SRP, Section 3.8.3.II.5, with the exception of tangential shear as discussed

above. The drywell is different from the containment structure and therefore it is considered that Sections CC-3411-5 and CC-3521-1 of the Containment Code, ACI-ASME 359, do not apply to the drywell structure. Criteria in Section 11.16 of ACI-318 code will be used in the design of the drywell and reactor pedestal. (Q&R 220.51) 3.8.3.5.2 Structural Steel The stresses in the structural steel are limited to those specified in the AISC Specification, Part 1, when designing for the loading combinations in Table 3.8-2, combinations 1, 2, and 4 through

8. For loading combination 3 of Table 3.8-2, the allowable steel stresses are increased to 1.33 time those specified in the AISC Specification. The appropriate factors of safety against yield are as discussed in the Commentary to the AISC Specifications. The allowable steel stresses are increased to 1.6 times those specified above, subject to an upper bound of 0.95 F y (yield stress), when designing for the loading combinations in Table 3.8-2, combinations 9 through 18.

In this situation a minimum factor of safety of 1.0/.95 = 1.05 against yield will be assured. In both cases, deformation of the steel members is limited by keeping the steel stresses within the elastic range for all loading combinations that exclude the effects of LOCA loads. When LOCA loads are considered in the loading combination, local stress may go beyond yield as long as overall function of the structure is not impaired. Connection parts may be allowed to yield locally under abnormal and/or extreme environmental loading conditions as long as the overall capacity is shown to be adequate. 3.8.3.5.3 Suppression Pool Liner Plate The allowable stresses for the suppression pool liner plate in the vent area of the drywell are those specified for reinforcing steel in Article CC 3000 of ASME B&PV Code, Section III, Div. 2. 3.8.3.5.4 Steel Pressure-Retaining Components Portions of the drywell boundary that are not backed by concrete, such as the equipment hatch, personnel lock, drywell head, guard pipes for the eleven high energy process lines, and MC penetration sleeves are designed in accordance with Subsection NE of Section III of the ASME B&PV Code. These components are designed for the loads and load combinations shown in Tables A3.9-6 and A3.9-7. In addition to the loads defined in these tables, the applicable pipe break loads are included under faulted conditions. The allowable stresses for these load combinations are summarized in the following list of figures from Section III of ASME B&PV Code:

CPS/USAR CHAPTER 03 3.8-30 REV. 11, JANUARY 2005 a. design conditions, Figure NE-3221-1; b. normal and upset conditions, Figure NB-3222-1;

c. emergency conditions, Figure NB-3224-1;

d. faulted conditions, Table F-1322; and e. test, Paragraph NE-3226. 3.8.3.5.5 Reactor Pedestal Steel The stresses in the pedestal shell plates and diaphragms are limited to those specified in Figure NF-3221-1 of Subsection NF (components and supports) of ASME B&PV Code Section III, Division 1. 3.8.3.6 Materials, Quality Control, and Special Construction Techniques There is no danger of radiation damage to the structures internal to the containment because, except for the reactor shield wall, they are not in a region of high-energy neutron flux. There is no danger of radiation damage to the steel plates of the reactor shield wall, because damage occurs at a neutron fluence of about 10 22 nvt. It has been determined that in the 40-year life expectancy of the station, the inside face of the wall will experience a neutron fluence of less than 5 x 10 17 nvt. The construction materials and quality control procedure for all concrete and structural steel internal structures conform to the standards set forth in Appendix B. 3.8.3.7 Testing and Inservice Surveillance Requirements After the drywell is complete with liner, concrete and steel structures, electrical and piping penetrations, equipment hatch, and personnel lock, it is tested in accordance with the procedure outlined for a drywell in SRP 3.8.3.7. The drywell is pressurized to the design pressure of 30 psig in four approximately equal increments. At each pressure level, the pressure is held constant for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> before measuring the strains and deflections at locations shown in Drawing S27-1401. The deflection is

measured by taut wire extensometers stretched across the drywell and kept under a constant tension. Electrical strain gauges are used to measure the strains through the thickness of the

drywell wall. The structural integrity of reinforced concrete members is also evaluated by mapping cracks larger than 0.01 inch in critical areas identifiable by design. Table 3.8-7 shows the predicted deflections of the drywell during the pressure test. The acceptance criteria for the drywell are the same as outlined for the primary reactor containment (see Subsection 3.8.1.7.1).

CPS/USAR CHAPTER 03 3.8-31 REV. 11, JANUARY 2005 3.8.4 Other Seismic Category I Structures 3.8.4.1 Description of the Structures The Seismic Category 1 structures, excluding the containment and containment internal structures, are: a. auxiliary building, b. fuel building,

c. control building,

d. diesel generator and HVAC (DG & HVAC) building,

e. radwaste building substructure,

f. containment gas control boundary, g. portions of circulating water screen house (CWSH), and h. UHS discharge structure. Refer to Section 1.2 for general arrangement drawings which show the relative size and location of the above structures. 3.8.4.1.1 Auxiliary Building The reinforced concrete auxiliary building is located adjacent to and fits the contour of the containment building on one side. This structure is supported on a mat foundation which is continuous with the mats under the containment structure, the control building, and the turbine building. Above the foundation mat, the auxiliary building is structurally isolated from the containment structure, but structurally connected to the control and turbine buildings.

The main steam tunnel extends from the containment to the turbine room through the auxiliary building. It houses the process piping and protects it from the effects of external missiles. In the unlikely event of pipe rupture inside the tunnel, it protects the control room and other Seismic Category I equipment and components from the effects of radioactive steam and pipe rupture loads. The tunnel also provides supports and restraints for the process piping. The ECCS pump rooms, in the lowest level of the auxiliary building, are in flood protection compartments with watertight doors. In the event of a pipe rupture, the flooding in one compartment will not result in the flooding of any other compartment, and the failure of a pump suction line will not drain the suppression pool. The second or grade level of the auxiliary building houses pump room access hatches and a cable tunnel. The third and fourth levels are provided for electrical switchgear and electrical penetrations. 3.8.4.1.2 Fuel Building The fuel building fits the contour of the containment structure on one side and is adjacent to the DG and HVAC building. One side of the building accommodates the fuel shipping cask railroad car. The reinforced concrete structure is supported on a concrete mat which is continuous with CPS/USAR CHAPTER 03 3.8-32 REV. 11, JANUARY 2005 the mats under the containment structure and the DG and HVAC building. The fuel building above the mat is structurally isolated from the containment building, but structurally connected to the DG & HVAC and auxiliary buildings. The three pools in the fuel building provide for fuel transfer, spent fuel storage, and cask loading. The pools are lined with seam-welded stainless steel plate welded to reinforced members embedded in the concrete. Channels are located behind the weld seams of the pool liners and are monitored to detect possible leakage from the pools. The reinforced concrete construction above the main floor provides missile and tornado protection. Crane seismic safety features, Figure 3.8-35, are provided for the fuel building cranes. 3.8.4.1.3 Control Building The reinforced concrete control building is located next to the auxiliary building. It is also adjacent to the DG & HVAC and radwaste buildings. The control building is supported by a reinforced concrete basemat which is continuous with the basemats of the adjoining buildings.

The building is structurally connected to the adjoining buildings above the basemats. HVAC equipment is located in the lower two and the top levels of the control building. Laundry facilities and laboratories are located at the grade or third level. The component cooling water heat exchangers are on the fourth level. The fifth level houses switchgear and provides a cable spreading area. The control room is located on the sixth level. The reinforced concrete walls and slabs of the control building provide tornado and missile protection for the control room and other Seismic Category I systems. 3.8.4.1.4 Diesel Generator and HVAC Building The reinforced concrete DG & HVAC building is located next to the fuel building and adjacent to the control building. The DG & HVAC building is supported by a reinforced concrete base mat which is continuous with the basemats of the adjoining buildings. The building is structurally connected to the adjoining buildings above the basemat. The fuel oil storage tanks are located on the lower level, and diesel generators are located at grade level. HVAC equipment is located above the grade floor. The reinforced concrete construction provides tornado and missile protection. 3.8.4.1.5 Radwaste Building Substructure The reinforced concrete radwaste building is located adjacent to the turbine building and the control building. The radwaste building is supported by a concrete basemat which is continuous with the basemats supporting the turbine building and the control building. The building is structurally connected to the adjoining buildings.

The radwaste building houses station systems which are required to process and dispose of radioactive wastes generated during power operation. The reinforced concrete construction prevents the dispersion of waste material by tornadic winds. Only the portion of the radwaste building located below grade is designed as a Seismic Category I structure.

CPS/USAR CHAPTER 03 3.8-33 REV. 11, JANUARY 2005 3.8.4.1.6 Containment Gas Control Boundary The containment gas control boundary is a limited leakage structure which surrounds the containment structure above the auxiliary and fuel buildings. The enclosure conforms to the shape of the containment and is separated from it by a distance of approximately 4 feet. The enclosure is made up of siding supported by structural steel framing attached to the

containment. 3.8.4.1.7 Circulating Water Screen House The circulating water screen house (CWSH) is a reinforced concrete, Seismic Category I structure located northwest of the station site. The CWSH is supported by a reinforced concrete basemat. The substructure is constructed of reinforced concrete. The shutdown service water pump cubicles in the superstructure are of reinforced concrete, but the rest of the superstructure is constructed of structural steel. The CWSH houses the plant service water pumps and strainers, the shutdown service water pumps and strainers, the diesel-driven fire pump and associated equipment, the circulating water pumps, and the traveling screens. All pumps are the vertical wet-pit type. The shutdown service water system equipment is the only equipment in the CWSH that is required to safely shut down the reactor or to maintain it in a safe shutdown condition. The three shutdown service water pumps and strainers are in their own missile-protected cubicles. Each cubicle has its own cooling unit which is electrically segregated from the others. Each cubicle is flood protected by bulkhead doors. No single failure of the equipment associated with one cubicle will have a detrimental effect on the rest of the system. The CWSH basin is constructed with two inlet channels to provide water to the shutdown service water pumps. 3.8.4.1.8 Ultimate Heat Sink Discharge Structure A reinforced concrete structure is located at the ultimate heat sink to accomodate the shutdown service water discharge lines. 3.8.4.2 Applicable Codes, Standards and Specifications The codes, standards and specifications applicable to the design, fabrication, construction, testing and in-service inspection of safety-related structures outside the containment are listed in Table 3.8-4 and include the following specification numbers: a. 1 through 5, 7, 8, 9, 11 through 14, 16, 17, 18, 20; b. 21, 23, 25c, 25d, and 28;

c. 31, 35, 37, 38;

d. 41, 43, 44, and 46 through 48. 3.8.4.3 Loads and Loading Combination The list of loads and their definitions and the loading combinations applicable to the design of Seismic Category I structures outside the containment are given in Table 3.8-1.2 and 3.8-2.

The list of load categories where the types of loads are defined is also given in these tables.

CPS/USAR CHAPTER 03 3.8-34 REV. 11, JANUARY 2005 In addition to their own dead loads including the weight of equipment, piping, cable pans, etc., floors are designed for conservative live loads resulting from the movement of the largest piece of equipment. The roofs are designed for a uniform live load of 25 psf in addition to snow loads and loads from probable maximum precipitation. The roofs are also designed to withstand suction pressure induced by the design wind and tornadic wind as discussed in Section 3.3. Pattern live loads are applied to determine maximum moments and shears in each slab. All slabs are designed for the effects of internal missiles, thermal gradients, and pipe rupture loads, wherever applicable. Floors and roofs are checked for their ability to transfer shear through

diaphragm action. The walls, interacting with the floor slabs, are designed to withstand the effects of seismic induced shears and moments. All walls are designed for external and internal missiles, transient thermal gradients, tornado-induced pressure, lateral soil and hydrostatic pressure and pipe rupture loads, wherever applicable, in addition to their own weight and associated loads from slabs and beams framing into the walls. For the design of subgrade walls a surcharge load of 500 psf, 1000 psf for E-70 or 300 psf for AASHO H-20 wheel loading is considered. The CWSH is also designed for the hydrostatic and hydrodynamic effects of the cooling lake water. The pools in the fuel building are designed for, in addition to applicable loads listed above, hydrostatic loads and hydrodynamic loads associ ated with water set in motion by seismic accelerations. The pools are designed for the effects of a maximum water temperature of 212°F. The containment gas control boundary is designed to be held under a negative pressure equivalent to 1/4-inch of water when infiltration flow rates given in Subsection 6.2.3 are being passed through the standby gas treatment system. The containment gas control boundary is not designed to withstand the effects of missiles. The siding of the enclosure is designed to fail for wind speeds of 200 mph, which is less than the design basis tornado. The gas control boundary is a fission product barrier only, and it is not designed for the high temperatures and pressures which are postulated for the containment. The steel framing for the containment gas control boundary is designed to withstand effects of tornado loading. In all instances, the Seismic Category I structures and structural components are designed for the vertical and horizontal accelerations associated with both SSE and OBE.

In Radwaste, Control and Diesel Generator Buildings, the effects of pool dynamic loads associated with SRV actuation and LOCA are considered negligible and shall not be used for the analysis and design of these buildings. The Category I manholes, buried piping, electric ducts and tunnels were designed for both dead and live load, and seismic loading conditions. In addition, the buried piping was designed for thermal expansion. The manholes, electrical ducts and tunnels are designed using the strength design method. The buried piping is designed using the ASME Section III (1977) stress

equations. The design procedure complies with the criteria contained in SRP, Section 3.8.4. (Q&R 220.54)

CPS/USAR CHAPTER 03 3.8-35 REV. 11, JANUARY 2005 3.8.4.4 Design and Analysis Procedures Conventional elastic techniques are used in the design and analysis of all structural components. All buildings are analyzed basically as shear wall structures, and all significant openings and discontinuities in structural members are included in the structural model. The boundary conditions selected for all structural models are determined by evaluating the stiffnesses (flexural, torsional and axial) of all the members connected to a boundary point and represent, to the extent practical, the actual restraint conditions. The walls, interacting with the floor slabs, are proportioned to resist the combination of seismic-induced overturning moments, vertical loads, and shears in accordance with the special provisions for shear walls of Appendix A.8 of ACI 318. Adequate provisions are made to transfer wall moments, vertical loads, and shears to the foundation. The finite element program SLSAP is used to analyze the basemat and the fuel pool walls. Frame analysis is done using computer program STRUDL-II. Concrete beams and columns are designed using the computer programs CBEAM and PCAUC, respectively. The STAND system is used to analyze and design structural steel beams and columns. For design of plate girders, the computer program PLGIRD is used. Limitation of concrete strain is per ACI 318 for both operating and design-basis loads for all structures, except for structures designed and analyzed for the effects of pipe breaks, including jet impingement, impact, pressurization and flooding outside the containment, where yield line theory (Reference 5) is used. For steel structures, strains are limited to the elastic range under operating and design-basis loadings. Design and analysis procedures for these structures comply with the portions of ACI-349 Code which are based on ACI 318.

3.8.4.5 Structural Acceptance Criteria The stresses and strains in the reinforced concrete walls, floor slabs, beams and equipment supports are limited to those specified in ACI 318, except for local stress due to concentrated loads as given in Note f) of Table 3.8-1.2. Serviceability checks are made in accordance with ACl 318 to assure crack control and to keep deflections below the limits prescribed by ACI 318 or to the manufacturers' recommendations for equipment supported by the reinforced concrete. The factors of safety against material strength are contained in the load factors in Table 3.8-1.2 and in the capacity reduction () factors in ACI 318 for the reinforced concrete. The stresses and strains in the structural steel are limited to those specified in the AISC Specifications, Part 1, when the loading combinations in Table 3.8-2, combinations 1, 2, and 4 through 8, are being designed for. The appropriate factors of safety against yield are those as discussed in the Commentary to the AISC Specifications. The allowable steel stresses are increased to 1.6 times those specified above, subject to an upper bound of 0.95 f y (yield stress), when the loading combinations in Table 3.8-2, conditions 9 through 18, are being designed for. In this situation a minimum factor of safety of 1.05 against yield is assured. In either case, deformations of structural steel members are limited because the stresses are kept within the elastic range, and redistribution of loads due to plastic deformations is not permitted. In addition, the deflections of all critical steel CPS/USAR CHAPTER 03 3.8-36 REV. 11, JANUARY 2005 members are calculated and kept below the limits prescribed by the AISC Specifications or manufacturers' recommendation for equipment supported by the steel. 3.8.4.6 Materials, Quality Control, and Special Construction Techniques Noncombustible and fire resistant materials are used wherever necessary throughout the facilities, particularly in areas containing such critical systems as the control room and components of engineered safety features. The construction materials conform to the standard set forth in Appendix B. Included in this section is a discussion of the quality control procedures employed specifying the frequency and location of sampling, and test requirements for the materials. Cadwelding procedure is also described in detail in this section. See Figures 3.8-36 and 3.8-37 for typical concrete construction details. 3.8.4.7 Testing and Inservice Surveillance Requirements No preliminary structural integrity or performance tests are conducted. However, rigorous inspection techniques and the quality control procedures described in Appendix B are adopted throughout construction. Routine periodic inspections of the concrete structures are conducted to check for possible deterioration, excessive cracking, or spalling of the concrete. Similar inspection is made on structural steel members to check for deterioration of surface coatings and abnormal deformations or warpage. 3.8.5 Foundations and Concrete Supports 3.8.5.1 Descriptions of Foundations and Supports 3.8.5.1.1 Foundations The station is supported by a common reinforced concrete mat. The mat is several feet thick, and is shown on the general arrangement (Section 1.2). The CWSH is supported by a

reinforced concrete mat. The concrete mats bear on the consolidated soil discussed in Section 2.5.

Typical reinforcing patterns at the junctions of the basemat and walls and the basemat and columns are shown in Figure 3.8-37. The mat in the area of the containment, which includes the auxiliary and fuel buildings, is considered part of the containment and is discussed in Subsection 3.8.1. 3.8.5.1.2 Concrete Supports Seismic Category I equipment is adequately anchored to and/or supported by concrete supports. The concrete supports consist of monolithically poured reinforced concrete pads.

The pads are integrally connected to the basemat or floor slabs by dowels. Typical anchor bolt details for Seismic Category I equipment are shown in Figure 3.8-38.

CPS/USAR CHAPTER 03 3.8-37 REV. 11, JANUARY 2005 The reactor pedestal which supports the RPV and the reactor shield wall is discussed in Subsection 3.8.3. 3.8.5.2 Applicable Codes, Standards and Specifications This section lists the codes, specifications, standards of practice, regulatory guides, and other accepted industry guidelines which are adopted to the extent applicable in the design and construction of the foundations and anchorages for Seismic Category I structures and equipment. To eliminate repetition, these codes, standards and specifications are described and discussed in Table 3.8-4 and given a specification reference number. Listed below are the reference numbers for the foundations. a. 1 through 5, 7, 8, 9; b. 11 through 14, 16, 17, 18; c. 23, 24 and 28;

d. 35 through 38;

e. 41, 43, and 46. 3.8.5.3 Load and Loading Combinations The loads and loading combinations listed and discussed in Subsections 3.8.1.3, 3.8.3.3, and 3.8.4.3 are also applicable to the design of foundations. Refer to Tables 3.8-1.1 and 3.8-1.2 for the load definitions and list of loading combinations that are considered in the design. Stability calculation loads and loading combinations are listed in Table 3.8-1.3. 3.8.5.4 Design and Analysis Procedures Conventional elastic techniques are used for the design and analysis of all Seismic Category I foundations. Design is based on the ACI 318 Code. All interior and exterior loads on the buildings are transferred to the basemat through elastic deformation of shear walls and columns. The foundation mats are properly sized to accommodate total overturning moments due to wind, tornado or seismic loads without exceeding the allowable soil bearing stress at any point. Horizontal translation due to wind, tornado or seismic loadings is resisted by frictional force between concrete mat and underlying soil. Passive resistance of soil acting against the subgrade walls is neglected. The uplift force due to hydrostatic pressure is deducted from the building dead load to compute the resultant downward load for calculating frictional resistance against sliding. The design and analysis of the foundation complies with the portions of the AC1-349 Code which are based on ACI-318. In determining the overturning moments due to seismic loads as discussed in Subsection 3.8.5.4 all three components of earthquake are considered acting simultaneous. (Q&R 220.59) The foundation mats are analyzed as a "mat on elastic foundation" using the finite-element computer program SLSAP-4. The boundary conditions selected for all structural models are determined by evaluating the stiffness (flexural, torsional and axial) of all members connected at a boundary point and represent, to the extent practicable, the actual restraint conditions.

Settlements are taken into account by the soil springs modeled for each node point.

CPS/USAR CHAPTER 03 3.8-38 REV. 11, JANUARY 2005 The modulus of subgrade reaction is varied within certain limits to determine the effects on critical sections. In general, lower values are used for long term loads and higher values for short term loads. Structural building supports for rotating or reciprocating (vibratory) Seismic Category I equipment satisfy vibratory design requirements. The equipment foundation design for vibratory equipment satisfies the machine vibration tolerances given in Figure 4 of "Vibration Tolerances" by T. C. Rathbone, Power Plant Engineering, Vol. 43, 1939. This has been accomplished by providing a foundation equipment mass ratio of 2.5 for vibratory equipment which weighs less than 5 kip and by performing appropriate dynamic analyses for vibratory equipment which weighs 5 kip or more. 3.8.5.5 Structural Acceptance Criteria 3.8.5.5.1 Structural Member Design The acceptance criteria for the reactor containment base slab are as specified in Subsection 3.8.1.5. The foundations for the main building complex and other Seismic Category I structures are proportioned according to the criteria set forth in Subsection 3.8.4.5. 3.8.5.5.2 Stability As described in Subsection 3.8.5.4, the basemats are supported on elastic soil springs and overturning is resisted by unequal bearing pressure. Table 3.8-1.3 lists the loads, load combinations, and factors of safety considered in the foundation stability investigation. 3.8.5.6 Materials, Quality Control, and Special Construction Techniques The construction materials for the mat foundations, concrete supports and machinery and equipment anchors conform to the standards set forth in Appendix B. Contained in that appendix is a discussion of the quality control procedures adopted which include the frequency and location of sampling and test requirements for the materials. Cadwelding is described in

detail. 3.8.5.7 Testing and Inservice Surveillance Techniques Routine observations are made of the mat foundations and concrete supports to determine the extent of cracking and settlement. Representative equipment anchor bolts are periodically tested for tightness.

Rigorous inspection during construction in conjunction with the quality control procedures for the structural materials outlined in Appendix B is carried out. Structural integrity and/or performance tests, in addition to those specified herein, are not conducted.

CPS/USAR CHAPTER 03 3.8-39 REV. 11, JANUARY 2005

3.8.6 References

1. H. G. Young and L. A. Tate, Design of Liners for Reactor Vessels, Conference on Prestressed Concrete Pressure Vessels, Institute of Civil Engineers, Paper 57, Group J (1967). 2. Nelson Stud Welding Applications in Power Generating Plants, Nelson Stud Welding Company, Lorain, Ohio. 3. J. M. Doyle and S. L. Chu, Liner Plate Buckling and Behavior of Stud and Rib Type Anchors, Proceedings of the First International Conference on Structural Mechanics in

Reactor Technology, Vol. 4, Part H, Berlin, Germany (September 1972). 4. C. W. Dunham, "Foundations of Structures," Second Edition, p. 199, McGraw Hill (1962). 5. K. W. Johansen, "Yield-Line Formulae for Slabs," Cement and Concrete Association (1972). 6. ASTM A615-76a, Specifications for Deformed and Plan Billet-Steel Bars for Concrete Reinforcement.

CPS/USAR CHAPTER 03 3.8-40 REV. 11, JANUARY 2005 TABLE 3.8-1.1 LOAD COMBINATIONS AND LOAD FACTORS FOR CONTAINMENT STRUCTURES-REINFORCED CONCRETE LOAD COMBINATION LOAD CONDITION SRV LOCA CATEGORY NO. D L H P' Pa P i P s P o T o T a E E' W W' R o R a Y r Y m Y j H' M F 1V2P ADS ALL MV PS CO CH 1 0.75 0.75 0.75 0.75 0.75 I. Construction 2 1.0 1.0 1.0 1.0 II. Test 3 1.0 1.0 1.0 1.0 1.0 4 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 III. Normal/ and and Severe IV. Environmental 5 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 6 1.0 1.0 1.0 1.5 1.0 1.0 1.25 1.0 1.0 1.0 1.0 1-> SERVICE LOADS <-1 -> 7 1.0 1.0 1.0 1.5 1.0 1.0 1.25 1.25 1.0 1.0 8 1.0 1.0 1.0 1.5 1.0 1.0 1.25 1.25 1.0 V. Abnormal 9 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 10 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 11 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 VI. Extreme/

Environmental 12 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 13 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 14 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 15 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 16 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 17 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 VII. Abnormal/

Severe Environmental 18 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 19 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 20 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 FACTORED LOADS VIII. Abnormal/

Extreme Environmental 21 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 22 1.0 1.0 1.0 1.0 1.0 IX. Severe Environ-mental/ Flooded Condition 23 1.0 1.0 1.0 1.0* 1.0 NOTES: *OBE Flooded Condition a) For the Construction Category, the wind load for a 10-year recurrence will be used.

b) Loads not applicable to a particular item may be deleted.

c) If for any load combination, the effect of any load other than D reduces the load, it will be deleted from the combination.

d) Each case of SRV Actuation is to be considered one at a time.

e) Each case of LOCA is to be considered one at a time.

f) The 33 1/3% increase in stresses allowed by the ASME B&PV Code, Section III, Division 2, CC-3420 (ACI-ASME 359, 1973) for members subject to wind or earthquake shall not be considered.

CPS/USAR TABLE 3.8-1.1 (Cont'd) CHAPTER 03 3.8-41 REV. 11, JANUARY 2005 Load Categories I. Construction Category This category includes all loads during construction. II. Test Category This category includes all loads during the structural acceptance test. III. Normal Category This category includes all loads on the structure during normal operation and shutdown. IV. Severe Environmental This category includes very infrequent loading during the station life, such as operating basis earthquake and design wind. V. Abnormal Category This category includes pressure loads and temperature effects from a postulated high-energy pipe break accident within the containment and/or compartment thereof. It includes pipe rupture loads in penetration and impingement loading. It also includes missile effects other than tornado and postulated accident generated missiles. VI. Extreme Environmental Category This category includes events which are credible but highly improbable such as a safe shutdown earthquake, and wind forces due to tornado and forces due to tornado generated missiles. VII. Abnormal/Severe Environmental Category This extremely unlikely loading is a combination of Categories IV and V. VIII. Abnormal/Extreme Environmental Category This extremely unlikely loading is a combination of Categories V and VI. IX. Severe Environmental/Flooded Category This category includes very infrequent loading during station life, such as operating basis earthquake, design winds and containment flooding associated loads.

CPS/USAR TABLE 3.8-1.1 (Cont'd) CHAPTER 03 3.8-42 REV. 11, JANUARY 2005 Explanation of Loading Conditions and Load Categories D = Dead load of the structure plus any other permanent load; including vertical and lateral pressures of liquids, piping, cable pan and weight of permanent equipment and its normal contents under operating and test conditions. L = Conventional floor and roof live loads, movable equip-ment loads and other loads which vary in intensity such as lateral soil pressure. Live load intensities may vary from zero to their maximum values to determine the most critical effect upon the structure for the load combination under consideration. To account for the effects of impact, equipment operating support reactions will be increased by the following

percentages: a. elevator supports, 100%; b. girders and their connections supporting power-operated cranes, 25%; c. girders and their connections supporting hand-operated cranes, 10%; d. supports for light machinery, shaft or motor-driven, 20%; and e. supports for reciprocating machinery or power-driven units, 50%. H = Hydrostatic pressure load in suppression pool.

Note: Reduced intensities of live loads such as conventional floor loads may be associated with accident and/or severe/extreme environmental conditions.

P o = Containment normal operating pressure.

P a* = Containment design accident pressure load due to large size break (DBA).

P i* = Containment design accident pressure load due to intermediate size break (IBA).

P s* = Containment design accident pressure load due to small size break (SBA). P' = Containment test pressure.

R o = Normal operating or shutdown reactions of piping at supports or anchor points, based on the most critical transient or steady-state condition.

R a = Pipe reactions under thermal conditions generated by the postulated break.

• Since these are time-dependent loads, their effect will be superimposted accordingly.

CPS/USAR TABLE 3.8-1.1 (Cont'd) CHAPTER 03 3.8-43 REV. 11, JANUARY 2005 E = Operating basis earthquake (OBE), including dynamic lateral soil pressure and hydrodynamic groundwater pressure. E' = Safe shutdown earthquake (SSE), including dynamic lateral soil pressure and hydrodynamic groundwater pressure. F = Loads associated with containment flooding, both hydrostatic and hydrodynamic loads.

T o = Thermal effects associated with normal operating, shutdown, construction and test conditions; based on the most critical transient or steady-state condition: a. Climatic temperature ranges maximum outside temperature - 100

° F minimum outside temperature - 0

° F b. Operating temperature ranges normal operating temperature inside containment - 104

° F in general areas, 122

° F in some closed compartments, and 95

° F in the suppression pool T a* = Thermal loads under thermal conditions generated by the postulated break. H' = Forces associated with the maximum probable flood or seiche. See Section 3.4 W = Design Wind Load See Subsection 3.3.1 W' = Tornado load and loads from tornado generated missiles See Subsections 3.3.2 and 3.5.23 M = Loads associated with missiles other than tornado and postulated accident generated missiles - See Section 3.5 Y r = Equivalent static load on the structure generated by the reaction on the broken high-energy pipe during the postulated break, and including an appropriate dynamic factor to account for the dynamic nature of the load Y j = Jet impingement equivalent static load on a structure generated by the postulated break, and including an appropriate dynamic factor to account for the dynamic nature of the load Y m = Missile impact equivalent static load on a structure Ym generated by or during the postulated break, like pipe whipping, and including an appropriate dynamic factor for the dynamic nature of the load

• Since these are time-dependent loads, their effect will be superimposted accordingly.

CPS/USAR TABLE 3.8-1.1 (Cont'd) CHAPTER 03 3.8-44 REV. 11, JANUARY 2005 Safety/Relief Valve (SRV) Discharge SRV lV2P = SRV loading due to subsequent actuation of one safety/relief valve. SRV/ADS = SRV loading due to seven (ADS) safety/relief valves discharge. SRV/ALL = SRV loading due to 16 (all) safety/relief valves discharge.

NOTES:

a. The SRV loads are treated as live loads in all load combinations with the exception of the combination that contains 1.5P a where a load factor of 1.25 is applied to the appropriate SRV loads. b. A single active failure causing one SRV discharge is considered in combination with the design-basis accident (DBA). c. Appropriate multiple SRV discharge is considered in combination with the small break accident (SBA) and intermediate break accident (IBA). d. Thermal loads due to SRV discharge are treated as T o* for normal operation and T a* for accident conditions. e. The suppression pool liner is designed in accordance with the ASME Boiler and Pressure Vessel Code, Division 1, Subsection NE, to resist the SRV negative pressure, considering strength, buckling, and low cycle fatigue.

Loss-of-Coolant Accident (LOCA) Loads: MV = LOCA loading due to main vent clearing PS = LOCA loading due to pool swell CO = LOCA loading due to condensation oscillation CH = LOCA loading due to chugging

• As defined in ACI 359-`1974.

CPS/USAR CHAPTER 03 3.8-45 REV. 11, JANUARY 2005 TABLE 3.8-1.2 LOAD COMBINATION AND LOAD FACTORS FOR REINFORCED CONCRETE (STRUCTURES OTHER THAN CONTAINMENT) LOADING COMBINATION LOAD FACTORS L *SRV (h) *LOCA DESCRIPTION NO. D EHL SLL C R o P o T o E E' W Wt R a P a T a F a H' M Y r Y j Y m ADS 1V2P ALL CH CO PS MV Design Strength (a) CONSTRUCTION 1 1.3 1.3 1.3 1.3 1.3 ACI 318 2 1.4 1.4 1.7 1.7 ACI 318 3 1.1 1.3 1.3 1.3 1.3 1.3 ACI 318 TEST 4 1.4 1.7 1.7 1.7 1.7 1.7 ACI 318 NORMAL 5 1.4 1.7 1.7 1.7 1.7 1.7 1.7 1.7 ACI 318 6 1.4 1.7 1.7 1.7 1.7 1.7 1.7 ACI 318 SEVERE 7 1.4 1.7 1.7 1.7 1.7 1.7 1.7 1.7 1.7 ACI 318 ENVIRONMENTAL 8 1.2 1.7 1.7 1.7 1.7 1.7 1.7 ACI 318 9 1.4 1.7 1.7 1.7 1.7 1.7 1.9 1.7 1.7 ACI 318 10 1.2 1.7 1.7 1.7 1.9 1.7 1.7 ACI 318 11 1.0 1.0 1.0 1.0 1.5 1.0 1.25 1.0 1.0 1.0 1.0 ACI 318 ABNORMAL 12 1.0 1.0 1.0 1.0 1.5 1.0 1.25 1.0 1.0 ACI 318

13 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 14 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 EXTREME 15 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 ENVIRONMENTAL 16 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 17 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 ABNORMAL/ 17A 1.0 1.0 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 SEVERE 18 1.0 1.0 1.0 1.25 1.0 1.25 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 18A 1.0 1.0 1.0 1.25 1.0 1.0 1.0 1.0 1.0 ACI 318 19 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 ABNORMAL/ 19A 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 EXTREME 20 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 20A 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 ACI 318 NOTES: a) For construction combination, wind load for a 10-year recurrence interval shall be used. b) T a is based on a temperature corresponding to the pressure, P

a. c) Loads not applicable to a particular system may be deleted. d) If for any load combination, the effect of any load other than D reduces the load, it will be deleted from the combination. e) For E, E', W t, M & R a, the resultant effects for both horizontal and vertical force components shall be determined by combining the individual effects by the square root of the sum of the squares. f) For combinations 13, 17, 18, 19 and 20, local stresses due to concentrated loads Y r , Y j , Y m & M may exceed allowable stresses provided that there will be no loss of function of any safety-related system. g) For loading combinations 1 through 10, the load factors shown shall be applied using zero values for R o and T o. These loads combinations shall also be checked using the values for R o and T o, but multiplying the combination by 0.75. h) SRV and LOCA loads are considered negligible in the radwaste, control, and diesel generator buildings and shall not be used in the analysis and design of these buildings.

• Only one load under each of these loadings shall be considered at one time.

CPS/USAR TABLE 3.8-1.2 (Cont'd) CHAPTER 03 3.8-46 REV. 11, JANUARY 2005 Load Categories I. Construction Category This category includes all loads during construction. II. Test Category This category includes all loads on the structure during a test of station equipment or systems. III. Normal Category This category includes all loads on the structure during normal operation. IV. Severe Environmental This category includes very infrequent loadings during the station life such as operating basis earthquake and design wind. V. Abnormal Category This category includes pressure loads and temperature effects from a postulated high energy break accident within a building and/or component thereof. It includes pipe rupture loads in penetration and impingement loading. It also includes missile effects other than tornado and postulated accident generated missiles. VI. Extreme Environmental Category This category includes events which are credible but highly improbable such as a safe shutdown earthquake, and wind forces due to tornado and forces due to tornado generated missiles. VII. Abnormal/Severe Environmental Category This extremely unlikely loading is a combination of Categories IV and V. VIII. Abnormal/Extreme Environmental Category This extremely unlikely loading is a combination of Categories V and VI. Explanation of Loading Conditions and Load Categories D = Dead Load of the structure plus any other permanent load; including vertical and lateral pressure of liquids, piping, cable pan and weight of permanent equipment and its normal contents under operating and test conditions. L = Conventional floor and roof live loads, movable equipment loads (EHL), and other loads which vary in intensity, such as lateral soil pressure. Live load intensities may CPS/USAR TABLE 3.8-1.2 (Cont'd) CHAPTER 03 3.8-47 REV. 11, JANUARY 2005 vary from zero to their maximum values to determine the most critical effect upon the structure for the load combination under consideration. To account for the effects of impact, equipment operating support reactions will be increased by the following

percentages: a. elevator supports, 100%;

b. girders and their connections supporting power operated cranes, 25%;

c. girders and their connections supporting Hand operated cranes, 10%; d. supports for light machinery, shaft or motor driven, 20%; and e. supports for reciprocating machinery or power-driven units, 50%. The portion of Tables 3.8-1.1, 3.8-1.2 and 3.8-2 addressing "reduced intensities of live load..." does not deviate from the SRP Procedure. Provision of live load "having its full value or being completely absent..." is made in Tables 3.8-1.1 and 3.8-1.2 under Explanation of Loading Conditions and Load Categories. "Live load intensities may vary from zero to their maximum values to determine the most critical effect upon the structure for the load combination under consideration." The reduced intensities of live loads referred to in the note found in Table 3.8-1.1 are in fact the actual live loads postulated during plant operation. Higher intensities of live load are postulated during plant shutdown to account for dismantled equipment handling and major maintenance

operations. The selection of certain percentages to provide for additional impact live load of specific equipment is in accordance with the general guideline found in Section 1.3.3 of the AISC Specification for the Design, Fabrication and Erection of Structural Steel for Buildings. (Q&R

220.39) EHL = Movable equipment loads C = Crane-lifted load, including impact SLL = Reduced intensities of floor live loads used with seismic loading combinations.

P a* = Pressure equivalent static load within or across a compartment and/or building, generated by the postulated break, and including an appropriate dynamic factor to account for the dynamic nature of the load R o = Normal operating or shutdown reactions of piping at supports or anchor points based on the most critical transient or steady-state condition R a = Pipe reactions under thermal conditions generated by the postulated break Y r = Equivalent static load on the structure generated by the reaction on the broken high-energy pipe during the postulated break, and including an appropriate dynamic factor to account for the dynamic nature of the load CPS/USAR TABLE 3.8-1.2 (Cont'd) CHAPTER 03 3.8-48 REV. 11, JANUARY 2005 Y j = Jet impingement equivalent static load on a structure generated by the postulated break, and including an appropriate dynamic factor to account for the dynamic nature of the load Y m = Missile impact equivalent static load on a structure generated by or during the postulated break, like pipe whipping and including an appropriate dynamic factor to account for the dynamic nature of the load E = Operating basis earthquake (OBE), including dynamic lateral soil pressure and hydrodynamic groundwater pressure E' = Safe shutdown earthquake (SSE), including dynamic lateral soil pressure and hydrodynamic groundwater pressure T o = Thermal effects associated with normal operating, shutdown, construction and test conditions, based on the most critical transient or steady-state conditions a. Climatic temperature ranges maximum outside temperature - 100

° F minimum outside temperature - 0

° F b. Operating temperature ranges ambient temperature inside the fuel building, auxiliary building, control building, DG & HVAC radwaste building, and CWSH -

70° F T a* = Thermal loads thermal conditions generated by the postulate break F a = Flood load generated by a high energy or moderate Energy pipe break outside the containment H' = Forces associated with the maximum probable flood or seiche (see Section 3.4) W = Design Wind Load (see Subsection 3.3.1)

W t = Tornado load and loads from tornado generated missiles Subsections 3.3.2 and 3.5.2.3) M = Loads associated with missiles other than tornado and postulated accident generated missiles (see Section 3.5)

• Since these are time-dependent loads, their effect will be superimposted accordingly.

CPS/USAR TABLE 3.8-1.2 (Cont'd) CHAPTER 03 3.8-49 REV. 11, JANUARY 2005 SRV IV2P = SRV loading due to one safety/relief valve subsequent actuation SRV ADS = SRV loading due to seven (ADS) safety/relief valves discharge SRV ALL = SRV loading due to 16 (all) safety/relief valves discharge LOCA MV = LOCA loading due to main vent clearing LOCA PS = LOCA loading due to pool swell LOCA CO = LOCA loading due to condensation oscillation LOCA CH = LOCA loading due to chugging

CPS/USAR CHAPTER 03 3.8-50 REV. 11, JANUARY 2005 TABLE 3.8-1.3 LOAD COMBINATIONS FOR STRUCTURAL STABILITY OF FOUNDATIONS LOADING COMBINATION LOAD FACTORS SAFETY FACTORS SRV (Notes 3,5) LOCA (Notes 3,5) DESCRIPTION NO. D s D e D 1 L 1 E E' W W t H' IV 2P ADS ALL MVC PS CO CH OVERTURNING SLIDING FLOTATION 1 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.5 1.5 1.5 SEVERE ENVIRONMENTAL 2 1.0 1.0 1.0 1.0 1.0 1.0 1.0 3 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.1 1.1 1.1 4 1.0 1.0 1.0 1.0 1.0 1.0 EXTREME ENVIRONMENTAL 4a 1.0 1.0 1.0 1.0 1.0 1.0 1.0 5 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.1 1.1 1.1 ABNORMAL SEVERE 5a 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 6 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.1 1.1 1.1 ABNORMAL EXTREME 7 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 NOTES: 1. D s = Self Weight of Structure D l = Vertical and Lateral Pressure of Liquid, Groundwater, and Vertical Soil Pressure D e = Actual equipment loads from manufacturers' drawings L l = Lateral soil pressure 2. If for any load combination, the effect of any load other than the dead load reduces the load, it will be deleted from the c ombination. 3. Only one load under each of these loadings shall be considered at one time. 4. For definition of load combinations and loads not defined in Note 1, refer to Table 3.8-1.2. 5. SRV and LOCA loads are considered negligible in the radwaste, control and diesel generator buildings, and shall not be used in the analysis and design of these buildings.

CPS/USAR CHAPTER 03 3.8-51 REV. 11, JANUARY 2005 TABLE 3.8-2 LOAD COMBINATIONS FOR STRUCTURAL STEEL LOAD FACTORS L SRV (g) (j) LOCA (g)(i)(j)

DESCRIPTION NO. D EHL SLL S C R o P o T o E E' W W t R a P a T a H' M Y r Y j Y m ADS 1V2P ALLV CH CO PS MV C ALLOWABLE STRESS 1 1.0 1.0 1.0 1.0 1.0 1.33 AISC Construction 2 1.0 1.0 1.0 1.0 1.0 AISC Test 3 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.33 AISC 4 1.0 1.0 1.0 1.0 1.0 1.0 1.0 AISC Normal 5 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 AISC 6 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 AISC 7 1.0 1.0 1.0 1.0 1.0 1.0 1.0 AISC Severe Environmental 8 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 AISC 9 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy 10 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy Abnormal 11 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy 12 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy 13 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy Extreme Environmental 14 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy 15 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy Abnormal/

Severe 16 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy 17 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy Abnormal/

Extreme 18 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.6 AISC£ .95 Fy NOTES: a) For construction combination, wind load for a 10-year recurrence interval shall be used. b) T a is based on a temperature corresponding to the pressure, P

a. c) Loads not applicable to a particular system may be deleted. d) If for any load combination, the effect of any load other than D reduces the load, it will be deleted from the combination. e) For E, E', W t, M & R a, the resultant effects for both horizontal and vertical force components shall also be determined by combining the individual effects by the square root of the sum of the squares. f) For loading combinations 2, 4, 5, 6, 7 and 8 use zero values for R o and T o with AISC allowables. These combinations shall also be checked using the values of R o and T o but the AISC allowables shall be increased by 33%. g) Only one load under each of these loadings shall be considered at one time. h) For load categories and load definitions refer to Table 3.8-1. S - Stability loads. Stability loads are ps uedo-static loads applied to a braced steel frame to assure suffic ient st rengt h and stiffness for column, beam, and girder stability. i) The loads due to a pool swell event are applied on the structural steel as pseudo static loads with dynamic load factors consistent with ductility ratios given in Standard Review Plan Section 3.5.3. j) SRV and LOCA loads are considered negligible in the radwaste, control and diesel generator buildings, and shall not be used in the analysis and design of these buildings.

CPS/USAR CHAPTER 03 3.8-52 REV. 11, JANUARY 2005 TABLE 3.8-3 (This table has been deleted)

CPS/USAR CHAPTER 03 3.8-53 REV. 11, JANUARY 2005 TABLE 3.8-4 LIST OF SPECIFICATIONS, CODES AND STANDARDS SPECI- FICATION REFERENCE NUMBER SPECIFICATION OR STANDARD DESIGNATION TITLE EDITION REMARKS 1 ACI 318-71 or 77 Supplement 1974 Building Code Requirements for Reinforced Concrete 1971 or 1977 Appendix "A" adopted for seismic design 2 ACI 301-72 Revision 1973 Specifications for Structural Concrete for Buildings 1973 3 ACI 307-68 ANSI A145.1-1968 Recommended Practice for Concrete Formwork 1968 4 ACI 305-72 ANSI A170.1-1972 Recommended Practice for Hot Weather Concreting 1972 5 ACI 211.1-74 Recommended Practice for Selecting Proportions for Normal Weight

Concrete 1974 6 ACI 349 Code Requirements for Nuclear Safety Related Structures 1976 and 1980 7 ACI 315-74 Manual of Standard Practice for Detailing Reinforced Concrete

Structures 1974 8 ACI 306-66 Recommended Practice for Cold Weather Concreting 1966 9 ACI 309-72 Recommended Practice for Consolidation of Concrete 1972 Title 69-56 10 (Deleted) 11 ACI 308-71 Recommended Practice for Curing Concrete 1971 Title 69-1 12 ACI 212 Guide for Use of Admixtures in Concrete ACI Journal Sept, 1971 Title 68-56 13 ACI 214-65 ANSI A146.1-1968 Recommended Practice for Evaluation of Compression Test Results of Field Concrete 1965 14 ACI 311-64 Recommended Practice for Concrete Inspection 1964 15 (Deleted) 16 ACI 304-73 Recommended Practice for Measuring, Mixing, Transporting and Placing

Concrete 1973 17 Report by ACI Committee 304 Placing Concrete by Pumping Methods ACI Journal May, 1971 Title 68-33 18 Report by ACI Committee 437 Subcommittee 1 Strength Evaluation of Existing Concrete Structures ACI Journal Nov, 1967 Title 64-61 19 (Deleted) 20 ACI-ASME-359 ASME Boiler & Pressure Vessel Code, Section III, Division 2, Concrete Reactor Vessels and Containments 1973 Issued for trial use and comment 21 AISC-69 or 78 Specification for the Design, Fabrication, and Erection of Structural

Steel for Buildings 1969 or 1978 CPS/USAR TABLE 3.8-4 (Cont'd) CHAPTER 03 3.8-54 REV. 11, JANUARY 2005 SPECI- FICATION REFERENCE NUMBER SPECIFICATION OR STANDARD DESIGNATION TITLE EDITION REMARKS 22 AISI Specification for the Design of Light Gage Cold-Framed Steel Structural Members 1968 or 1980 23 AWS D1.1 Structural Welding Code (with required visual inspec tion based upon VWAC, Revision 2) 1976 or 1977 See Note 1 24 AWS D12.1.61 Recommended Practice for Welding Reinforcing Steel, Metal Inserts and Connection in Reinforced Concrete

Construction 1961 25 ASME ASME Boiler and Pressure Vessel Code, Section III, Division 1, NE 1971 with Summer of 1973 Addenda For Containment Locks and Hatches 25a ASME ASME Boiler and Pressure Vessel Code, Section III, Division 1, NE 1974 with Summer of 1976 Addenda For Drywell Locks and

Hatches 25b ASME ASME Boiler and Pressure Vessel Code, Section III, Division 1, NF 1974 with Winter of 1975 Addenda For Reactor Pedestal 25c ASME ASME Boiler and Pressure Vessel Code, Section III, Division 1, ND 1977 For Fuel Pool Gates 25d ASME ASME Boiler and Pressure Vessel Code, Section III, Division 2 1977 For Fuel Pool Liners 26 (Deleted) 27 (Deleted) 28 ASTM Annual Books of ASTM Standards 29 (Deleted) 30 (Deleted) 31 UBC Uniform Building Code 1970 or 1979 32 (Deleted) 33 (Deleted) 34 (Deleted) 35 NRC Regulatory Guide 1.10 Mechanical Cadweld Splices in Reinforcing Bars of Concrete Containments Feb. 1, 1971 Withdrawn by the NRC 7/8/81 36 NRC Regulatory Guide 1.12 Instrumentation for Earthquakes Rev. 1, Apr. 1974 37 NRC Regulatory Guide 1.13 Spent Fuel Storage Facility Design Basis Rev. 1, Dec. 1975 38 NRC Regulatory Guide 1.15 Testing of Reinforcing Bars For Concrete Structures (revision 1) Dec. 28, 1972 Withdrawn by the NRC 7/8/81 39 NRC Regulatory Guide 1.18 Structural Acceptance Test for Concrete Primary Reactor Containments (Revision 1) Dec. 28, 1972 Withdrawn by the NRC 7/8/81 40 NRC Regulatory Guide 1.19 Nondestructive Examinations of Primary Containment Liner Welds (Revision 1) Aug. 11, 1972 Withdrawn by the NRC 7/8/81 41 NRC Regulatory Guide 1.26 Quality Group Classifications and

Standards Rev. 3, Feb. 1976 42 NRC Regulatory Guide 1.27 Ultimate Heat Sink Rev. 2, Jan. 1976 43 NRC Regulatory Guide 1.29 Seismic Design Classification Rev. 3, Sept. 1978

CPS/USAR TABLE 3.8-4 (Cont'd) CHAPTER 03 3.8-55 REV. 11, JANUARY 2005 SPECI- FICATION REFERENCE NUMBER SPECIFICATION OR STANDARD DESIGNATION TITLE EDITION REMARKS 44 NRC Regulatory Guide 1.31 Control of Stainless Steel Welding Rev. 3, Apr. 1976 45 (Deleted) 46 CRSI Manual of Standard Practice 1973 47 ANSI N45.2.5 Supplementary QA Requirements for Installation, Inspection and Testing of Structural Concrete and Structural Steel during Con-struction Phase of Nuclear Power Plants 1974 48 NRC Regulatory Guide 1.55 Concrete Placement in Category I

Structures Rev. 0, June 1973 Withdrawn by the NRC

7/8/81 49 NRC Regulatory Guide 1.57 Design Limits and Loading Combinations for Metal Primary Reactor Containment Systems and Components June 1973 50 NRC Regulatory Guide 1.136 Materials for Concrete Contain-ments (Article CC-2000 of the Code for Concrete Reactor Vessels and Containments)

Rev. 1, Oct. 1978 51 NRC Regulatory Guide 1.142 Safety Related Concrete Structures for Nuclear Power Plants (other than Reactor Vessels and Containments)

Rev. 0 April 1978 Explanatin of Abbreviations ACI - American Concrete Institute AISC - American Institute of Steel Construction AISI - American Iron and Steel Institute ANSI - American National Standards Institute API - American Petroleum Institute ASME - American Society of Mechanical Engineers ASTM - American Society for Testing and Materials AWS - American Welding Society CRSI - Concrete Reinforcing Steel Institute NEC - National Electric Code NRC - Nuclear Regulatory Commission UBC - Uniform Building Code VWAC - Visual Weld Acceptance Criteria NOTES: 1. Clarification to and deviation from portions of AWS D1.1 (and VWAC Revision 2 for visual inspection of welds made to the requirements of AWS D1.1) are made based on engineering evaluations.

CPS/USAR CHAPTER 03 3.8-56 REV. 11, JANUARY 2005 TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS CONTAINMENT MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MC-1 745 ft-0 in. 745 ft-0 in.

225° 18'-0" Equipment Hatch 1MC-2 741 ft-0 in. 741 ft-0 in.

78° 9'-10" Personnel Lock 1MC-3 832 ft-3 in. 832 ft-3 in.

60° 9'-10" Personnel Lock 1MC-4 764 ft-2-7/16 in. 180° -14 ft-0-1/2 in. 20 in. Fuel Transfer Tube 1MC-5 770 ft-9-1/6 in. 770 ft-8-11/16 in. 0

° + 10 ft-6 in. 24 in. Main Steam "C" 1MC-6 770 ft-9 in. 770 ft-8-5/8 in. 0

° +3 ft-6 in. 24 in. Main Steam "A" 1MC-7 770 ft-9 in. 770 ft-8-5/8 in. 0

° -3 ft-6 in. 24 in. Main Steam "D" 1MC-8 770 ft-9-1/16 in. 770 ft-8-11/16 in. 0

° -10 ft-6 in. 24 in. Main Steam "B" 1MC-9 763 ft-4-1/4 in. 763 ft-3-7/8 in. 0

° + 7 ft-0 in. 20 in. Feedwater "A" 1MC-10 763 ft-4-1/4 in. 763 ft-3-7/8 in. 0

° -7 ft-0 in. 20 in. Feedwater "B" 1MC-11 720 ft-0 in. 720 ft-0 in.

38° 20 in. RHR Pump Suction "A" 1MC-12 720 ft-0 in. 720 ft-0 in.

323° 20 in. RHR Pump Suction "B" 1MC-13 720 ft-0 in. 720 ft-0 in.

308° 20 in. RHR Pump Suction "C" 1MC-14 757 ft-6 in. 757 ft-6 in.

0° 18 in. RHR Shutdown Suction 1MC-15 764 ft-3 in. 764 ft-3 in.

63° 12 in. RHR LPCI "A" 1MC-16 748 ft-6 in. 748 ft-6 in.

335° 12 in. RHR LPCI "B" 1MC-17 748 ft-0 in. 748 ft-0 in.

300° 12 in. RHR LPCI "C" 1MC-18 740 ft-0 in. 740 ft-0 in.

94° 14 in. RHR Test to Supp. "A" 1MC-19 740 ft-0 in. 740 ft-0 in.

317° 14 in. RHR Test to Supp. "C" 1MC-20 752 ft-0 in. 752 ft-0 in.

275° 14 in. RHR Test to Supp. "B" 1MC-21 739 ft-0 in. 739 ft-0 in.

27° 2 in. RHR "A" P.R.V. 1MC-22 739 ft-0 in. 739 ft-0 in.

23° Spare 1MC-23 752 ft-0 in. 752 ft-0 in.

34° 2 in. RHR"A" P.R.V. 1MC-24 742 ft-0 in. 742 ft-0 in.

23° 12 in. RHR "A" P.R.V. 1MC-25 740 ft-0 in. 740 ft-0 in.

320° 2 in. RHR "B" P.R.V. (Pump Suction) 1MC-26 744 ft-9 in. 744 ft-9 in.

339° 12 in. RHR "B" P.R.V. (Heat Exchanger) 1MC-27 740 ft-0 in. 740 ft-0 in.

334° 1-1/2 in. RHR "B" P.R.V. (Shutdown Return) 1MC-28 720 ft-0 in. 720 ft-0 in.

355° 6 in. RCIC Pump Suction 1MC-29 740 ft-0 in. 740 ft-0 in.

303° 1-1/2 in. RHR "C" P.R.V. (Pump Suction) 1MC-30 740 ft-0 in. 740 ft-0 in.

296° 1-1/2 in. RHR "C" P.R.V. (Pump Discharge) 1MC-31 745 ft-0 in. 745 ft-0 in.

354° 6 in. RHR "B" P.R.V. (Crosstie to RCIC) 1MC-32 720 ft-0 in. 720 ft-0 in.

52° 20 in. LPCS Pump Suction 1MC-33 743 ft-0 in. 743 ft-0 in.

252° 12 in. HPCS Test to Supp. 1MC-34 720 ft-0 in. 720 ft-0 in.

66° 12 in. Suppression Pool Clean Up Suction 1MC-35 758 ft-0 in. 758 ft-0 in.

266° 10 in. HPCS Pump Discharge 1MC-36 752 ft-6 in. 752 ft-6 in.

100° 10 in. LPCS Pump Discharge 1MC-37 720 ft-0 in. 720 ft-0 in.

243° 20 in. HPCS Pump Suction 1MC-38 746 ft-0 in. 746 ft-0 in.

97° 4 in. LPCS P.R.V. (Pump Discharge) 1MC-39 740 ft-0 in. 740 ft-0 in.

357° Spare 1MC-40 739 ft-0 in. 739 ft-0 in.

30° 2 in. RCIC Min. Flow 1MC-41 739 ft-11-5/8 in. 740 ft-0 in.

9° 12 in. RCIC Turbine Steam EXH CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-57 REV. 11, JANUARY 2005 CONTAINMENT MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MC-42 756 ft-4-3/4 in. 756 ft-4-3/4 in. 0

° -11 ft-0 in. 4 in. RCIC Head Spray 1MC-43 758 ft-0-3/8 in. 757 ft-10-1/8 in. 0

° -5 ft-0 in. 8 in. RCIC Turbine Steam Supply 1MC-44 745 ft-0 in. 745 ft-0 in.

21° 3 in. RCIC Turbine Vacuum Breaker 1MC-45 756 ft-6-3/8 in. 756 ft-6 in.

0° -15 ft-0 in. 3 in. Main Steam Drain 1MC-46 756 ft-9 in. 756 ft-9 in.

95° 10 in. Component Cooling Water Supp. 1MC-47 756 ft-9 in. 756 ft-9 in.

98° 10 in. Component Cooling Water Return 1MC-48 758 ft-0 in. 758 ft-0 in.

121° 3 in. Shutdown Service Water Supply 1MC-49 763 ft-6 in. 763 ft-6 in.

124° 1 in. Breathing Air 1MC-50 790 ft-0 in. 790 ft-0 in.

95° 4 in Make-up Condensate Supply 1MC-51 790 ft-0 in. 790 ft-0 in.

98° Spare 1MC-52 758 ft-0 in. 758 ft-0 in.

127° 8 in. Fuel Pool Cooling & Cleanup 1MC-53 758 ft-0 in. 758 ft-0 in.

145° 10 in. Fuel Pool Cooling & Cleanup 1MC-54 762 ft-3 in. 762 ft-3 in.

0° + 2 ft-3 in. Spare 1MC-55 743 ft-0 in. 743 ft-0 in.

266° Spare 1MC-56 753 ft-0 in. 753 ft-0 in.

312° 10 in. F.P. Containment Standpipe 1MC-57 769 ft-11 in. 769 ft-11 in.

65° 3 in. Instrument Air 1MC-58 770 ft-0 in. 770 ft-0 in.

61° 1 in. Instrument Booster Air 1MC-59 770 ft-0 in. 770 ft-0 in.

68° 3 in. Service Air 1MC-60 756 ft-6 in. 756 ft-6 in.

0° + 11 ft-3 in. 6 in. RWCU Pump Supply 1MC-61 758 ft-3 in. 758 ft-3 in.

0° + 7 ft-11 in. 4 in. RWCU Pump Return 1MC-62 763 ft-6 in. 763 ft-6 in.

294° 2 in. Hydrogen Recombiner to Containment 1MC-63 758 ft-0 in. 758 ft-0 in.

187° 2 in. C.R.D. Pump Discharge 1MC-64 763 ft-8-7/8 in. 763 ft-8-7/8 in. 0

° + 15 ft-0 in. 4 in. RWCU to RHR Return 1MC-65 763 ft-8-7/8 in. 763 ft-8-7/8 in. 0

° -15 ft-0 in. 2 in. Radwaste Reprocessing & Disposal 1MC-66 752 ft-0 in. 752 ft-0 in.

75° Spare 1MC-67 743 ft-0 in. 743 ft-0 in.

245° 6 in. Containment Service Air (Cnmt. Press) 1MC-68 752 ft-0 in. 752 ft-0 in.

82° 1/2 & 3/4 in. Process Sampling 1MC-69 758 ft-0 in. 758 ft-0 in.

124° 3 in. Containment Equipment Drains 1MC-70 763 ft-9 in. 763 ft-9 in.

277° 3 in. Containment Floor Drains 1MC-71 752 ft-0 in. 752 ft-0 in.

58° 2 in. Hydrogen Recombiner from Contnt. 1MC-72 752 ft-0 in. 752 ft-0 in.

66° 2 in. Hydrogen Recombiner to Contnt. 1MC-73 745 ft-0 in. 745 ft-0 in.

195° Spare 1MC-74 762 ft-3 in. 762 ft-3 in.

0° -11 ft-9 in. 6 in. RT Decontamination 1MC-75 764 ft-0 in. 764 ft-0 in.

192° Spare 1MC-76 740 ft-0 in. 740 ft-0 in.

323° 1-1/4 in. RHR P.R.V. (Drain) 1MC-77 739 ft-11-5/8 in. 740 ft-0 in.

4° Spare 1MC-78 751 ft-0 in. 751 ft-0 in.

281° 4 in. Component Cooling Water Supply 1MC-79 748 ft-0 in. 748 ft-0 in.

259° 10 in. Suppression Pool Clean-Up Return 1MC-80 753 ft-0 in. 753 ft-0 in.

40° Spare 1MC-81 742 ft-0 in. 742 ft-0 in.

27° 6 in. Fire Protection 1MC-82 758 ft-0 in. 758 ft-0 in.

154° 10 in. Fire Protection CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-58 REV. 11, JANUARY 2005 CONTAINMENT MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MC-83 768ft-0 in. 768ft-0 in.

127° Spare 1MC-84 751ft-9 in. 751ft-9 in.

44° Spare 1MC-85 753ft-0 in. 753ft-0 in.

307° 6 in. Cycle Condensate 1MC-86 762ft-3 in. 762ft-3 in.

0° + 12ft-0 in. 4 in. RWCU to Condenser 1MC-87 753ft-0 in. 753ft-0 in.

48° 11/2 in. RHR "A" P.R.V. 1MC-88 743ft-0 in. 743ft-0 in.

335° 4 in. Component Cooling Water Return 1MC-89 765ft- 10-7/16 in.

765ft- 10-7/16 in.

25° 11/2 in. RHR Ht. Exch. Shell Vent 1MC-101 790ft-0 in. 790ft-0 in.

146° 36 in. Cont. Vent Air Supply 1MC-102 772ft-6 in. 772ft-6 in.

146° 36 in. Cont. Vent Air Purge & Exhaust 1MC-103 768ft-0 in. 768ft-0 in.

151° 10 in. Cont. Cooling Unit Chilled Water Supp. 1MC-104 768ft-0 in. 768ft-0 in.

145° 10 in. Cont. Cooling Unit Chilled Water Return 1MC-105 788ft-0 in. 788ft-0 in.

102° Spare 1MC-106 782ft-6 in. 782ft-6 in.

113° 12 in. Continuous Cnmt. Purge Air Exh.

1MC-107 769ft- 10-1/2 in.

769ft- 10-1/2 in. 205

° 10 in. Drywell Cooling Unit Chilled Water Supp.

1MC-108 769ft- 10 in. 769ft- 10 in. 202° 10 in. Drywell Cooling Unit Chiller Water Return 1MC-109 769ft- 10-1/4 in.

769ft- 10-1/4 in. 199

° 10 in. Drywell Cooling Unit Chilled Water Supp.

1MC-110 769ft- 9-3/4 in.

769ft- 9-3/4 in. 196

° 10 in. Drywell Cooling Unit Chiller Water Return 1MC-111 792ft-6 in. 792ft-6 in.

102° Spare 1MC-112 792ft-0 in. 792ft-0 in.

113° Spare 1MC-113 782ft-6 in. 782ft-6 in.

117° 12 in. Continuous Cnmt. Purge Air Supply 1MC-114 782ft-6 in. 782ft-6 in.

51° Spare 1MC-115 782ft-6 in. 782ft-6 in.

102° Spare 1MC-116 782ft-6 in. 782ft-6 in.

93° 2 in. Standby Liquid Control 1MC-150 782ft-6 in. 782ft-9 in.

62° 3/4 in. Cnmt. Pressure Monitors 1MC-151 782ft-6 in. 782ft-9 in.

65° 3/4 in. Cnmt. Pressure Monitors/ DW Pressure Recirculation 1MC-152 767ft-0 in. 767ft-3 in.

182° 3/4 in. Containment Monitoring 1MC-153 782ft-6 in. 782ft-9 in.

107° 3/4 in. Drywell Pressure 1MC-154 770ft-0 in. 770ft-3 in.

298° Spare 1MC-155 790ft-0 in. 790ft-3 in.

68° Spare 1MC-156 792ft-0 in. 792ft-3 in.

235° 3/4 in. Containment Pressure (SGTS Train A)/

DW Air Temp/Humidity Inst. Panel CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-59 REV. 11, JANUARY 2005 CONTAINMENT MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MC-157 792ft-0 in. 792ft-3 in.

248° 3/4 in. Suppression Pool Water Level Monitoring 1MC-158 782ft-6 in. 782ft-3 in.

60° Spare 1MC-159 763ft-6 in. 763ft-3 in.

120° Spare 1MC-160 767ft-0 in. 766ft-9 in.

190° 3/4 in. Containment Monitoring System 1MC-161 770ft-0 in. 769ft-9 in.

294° Spare 1MC-162 790ft-0 in. 790ft-3 in.

205° Spare 1MC-163 790ft-0 in. 790ft-3 in.

210° Spare 1MC-164 765ft-0 in. 764ft-9 in.

150° 3/4 in. Suppression Pool Makeup 1MC-165 792ft-6 in. 792ft-3 in.

128° 3/4 in. Containment Differential Pressure 1MC-166 789ft-0 in. 788ft-9 in.

291° 2 in. Hydrogen Recombiner from Containment 1MC-167 788ft-0 in. 788ft-3 in.

117° 3/4 in. Containment Pressure (SGTS Train B) 1MC-168 789ft-0 in. 789ft-3 in.

280° 3/4 in. Containment Differential Pressure 1MC-169 793ft-0 in. 792ft-9 in.

117° 3/4 in. Continuous Containment Purge Damper Control 1MC-170 790ft-0 in. 790ft-3 in.

65° Spare 1MC-171 767ft-0 in. 767ft-3 in.

185° 3/4 in. Suppression Pool Makeup 1MC-172 773ft- 1-1/8 in.

773ft- 4-15/16 in. 300

° 11/2 in. HR Ht. Exch. Shell Vent 1MC-173 788ft-0 in. 788ft-3 in.

242° -31/2 in 3/4 in. Containment Monitoring System 1MC-174 788ft-0 in. 788ft-3 in.

128° Spare 1MC-175 788ft-0 in. 788ft-3 in.

315° Spare 1MC-176 788ft-0 in. 788ft-3 in.

62° Spare 1MC-177 720ft-0 in. 720ft-0 in.

60° 11/4 in. Supp. Pool Water Level (RCIC) 1MC-178 745ft-0 in. 745ft-0 in.

55° Spare 1MC-179 720ft-0 in. 720ft-0 in.

150° 11/4 & 3/4 in. H.P. Core Spray System &

Suppression Pool Make-up 1MC-180 745ft-0 in. 745ft-0 in.

150° 11/4 in. H.P. Core Spray System 1MC-181 720ft-0 in. 720ft-0 in.

203° 11/4 in. Supp. Pool Water Level 1MC-182 742ft-6 in. 742ft-6 in.

203° Spare 1MC-183 720ft-0 in. 720ft-0 in.

260° 11/4 in. Supp. Pool Water Level 1MC-184 742ft-6 in. 742ft-6 in.

260° Spare 1MC-200 769ft-6 in. 769ft-6 in.

76° 3/4 in. Supp. Pool Water Level (RCIC) 1MC-201 769ft-6 in. 769ft-6 in.

79° Spare 1MC-202 769ft-6 in. 769ft-6 in.

82° Spare 1MC-203 782ft-6 in. 782ft-6 in.

298° 3/4 in. Containment Monitoring 1MC-204 782ft-6 in. 782ft-6 in.

295° 3 in. S/D Service Water 1MC-205 782ft-6 in. 782ft-6 in.

291° 3 in. S/D Service Water 1MC-206 788ft-0 in. 788ft-0 in.

245° 1 in. Instrument Air 1MC-207 745ft-0 in. 745ft-0 in.

190° Spare 1MC-208 768ft-0 in. 768ft-0 in.

120° 3 in. S/D Service Water Return 1MC-209 745ft-0 in. 745ft-0 in.

185° Spare 1MC-210 759ft-6 in. 759ft-6 in.

0° +4ft-6 in. 3/4 in. Post Accident Sample 1MC-211 756ft-6 in. 756ft-6in. 0

° +4ft-6 in. Spare CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-60 REV. 11, JANUARY 2005 CONTAINMENT ELECTRICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1P1B-1 773ft-0 in. 773ft-0 in. 42

°30' 18 in. Recirc. Pump 1A (P) 1P2B-1 773ft-0 in. 773ft-0 in. 317

°30' 18 in. Recirc. Pump 1B (P) 1K1E-2 769ft-0 in. 769ft-0 in. 37

°30' 12 in. Instrumentation 1 (K) 1K2B-1 769ft-0 in. 769ft-0 in. 302

°30' 12 in. Instrumentation (K) 1K1B-2 769ft-0 in. 769ft-0 in. 57

°30' 12 in. Instrumentation (K) 1K2B-2 769ft-0 in. 769ft-0 in. 322

°30' 12 in. Instrumentation (K) 1K1N 792ft-0 in. 792ft-0 in. 42

°30' 12 in. Neutron Monitoring System 1 (K) 1K2E-1 792ft-0 in. 792ft-0 in. 307

°30' 12 in. Instrumentation 2 (K) 1K1E 792ft-0 in. 792ft-0 in. 52

°30' 12 in. Instrumentation 1 (K) 1K4N 792ft-0 in. 792ft-0 in.

315° 12 in. Neutron Monitoring System 4 (K) 1C1E 794ft-0 in. 794ft-0 in.

40° 0' 12 in. Reactor Protection System 1 (C) 1C2E-1 794ft-0 in. 794ft-0 in.

305° 12 in. Reactor Protection System 2 (C) 1C3E 794ft-0 in. 794ft-0 in.

55° 0' 12 in. Reactor Protection System 3 (C) 1C4E 794ft-0 in. 794ft-0 in. 317

°30' 12 in. Reactor Protection System 4 (C) 1C1B-2 771ft-0 in. 771ft-0 in.

40° 0' 12 in. Information (C) 1C2B-1 771ft-0 in. 771ft-0 in.

310° 12 in. Information (C) 1C1B-1 771ft-0 in. 771ft-0 in.

50° 12 in. Information (C) 1C2B-2 771ft-0 in. 771ft-0 in. 320

° 0' 12 in. Information (C) 1K1B-1 769ft-0 in. 769ft-0 in. 52

°30' 12 in. Instrumentation (K) 1K2B-3 769ft-0 in. 769ft-0 in. 307

°30' 12 in. Instrumentation (K) 1K2N 794ft-6 in. 794ft-6 in.

240° 8 in. Neutron Monitoring System 2 (K) 1K3N 794ft-6 in. 794ft-6 in.

140° 8 in. Neutron Monitoring System 3 (K) 1P1E-1 796ft-0 in. 796ft-0 in. 42

°30' 12 in. Engineered Safety Feature 1 (P) 1P2E-1 796ft-0 in. 796ft-0 in. 307

°30' 12 in. Engineered Safety Feature 2 (P) 1P2E-3 796ft-0 in. 796ft-0 in. 52

°30' 12 in. Engineered Safety Feature 2 (P) 1P1B-4 773ft-0 in. 773ft-0 in. 47

°30' 12 in. Balance of Plant (P) 1P2B-4 773ft-0 in. 773ft-0 in. 312

°30' 12 in. Balance of Plant (P) 1SP-1 794ft-0 in. 794ft-0 in.

45° Spare 1C2E-2 794ft-0 in. 794ft-0 in.

310° 12 in. Engineered Safety Feature 2 (C) 1P1E-2 794ft-0 in. 794ft-0 in.

50° 12 in. Engineered Safety Feature 1 (P) 1K1E-1 792ft-0 in. 792ft-0 in. 37

°30' 12 in. Control Rod Indication 1 (K) 1K2E-2 792ft-0 in. 792ft-0 in. 302

°30' 12 in. Control Rod Indication 2 (K) 1K3E 792ft-0 in. 792ft-0 in. 57

°30' 12 in. Instrumentation 3 (K) 1K4E 792ft-0 in. 792ft-0 in.

320° 12 in. Instrumentation 4 (K) 1P1B-3 773ft-0 in. 773ft-0 in. 52

°30' 12 in. Balance of Plant (P) 1P2B-3 773ft-0 in. 773ft-0 in. 307

°30' 12 in. Balance of Plant (P) 1C1B-3 771ft-0 in. 771ft-0 in.

55° 12 in. Information (C) 1C2B-3 771ft-0 in. 771ft-0 in.

305° 12 in. Information (C) 1P2E-2 796ft-0 in. 796ft-0 in. 302

°30' 12 in. Engineered Safety Feature 2 (P) 1P1B-2 773ft-0 in. 773ft-0 in. 57

°30' 12 in. Balance of Plant (P) 1P2B-2 773ft-0 in. 773ft-0 in. 302

°30' 12 in. Balance of Plant (P)

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-61 REV. 11, JANUARY 2005 CONTAINMENT ELECTRICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1P1B-5 771ft-0 in. 771ft-0 in.

35° 12 in. Balance of Plant (P) 1SP-3 771ft-0 in. 771ft-0 in.

325° Spare 1K1B-3 775ft-9 in. 775ft-9 in.

192° 12 in. Testing Instrumentation (K) 1K1B-4 775ft-9 in. 775ft-9 in.

205° 12 in. Testing Instrumentation (K)

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-62 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-1 744ft-0 in. 744ft-0 in.

225° 16ft-0 in. Equipment Hatch 1MD-2 741ft-0 in. 741ft-0 in.

90° 10ft-0 in. Personnel Lock 1MD-3 90° 24 in. Manhole 1MD-4 786ft-0 in. 786ft-0 in.

60° +0ft-8 in. 3 in. Standby Liquid Conrol 1MD-5 771ft- 0-5/8 in. 771ft-0 in. 0

° +10ft-6 in. 24 in. Main Steam "C" 1MD-6 771ft-0-1/2 in.

770ft- 11-7/8 in. 0

° +3ft-6 in. 24 in. Main Steam "A" 1MD-7 771ft-0-1/2 in.

770ft- 11-7/8 in 0

° -3ft-6 in. 24 in. Main Steam "D" 1MD-8 771ft-0-5/8 in. 771ft-0 in.

0° -10ft-6 in. 24 in. Main Steam "B" 1MD-9 763ft- 7-1/8 in.

763ft- 6-1/2 in. 0

° +7ft-0 in. 20 in. Feedwater "A" 1MD-10 763ft- 7-1/8 in.

763ft- 6-1/2 in.

0° -7ft-0 in. 20 in. Feedwater "B" 1MD-11 744ft-0 in. 744ft-0 in.

145° 3/4 in. RR Pump Seal Purge "A" 1MD-12 744ft-0 in. 744ft-0 in.

325° 3/4 in. RR Pump Seal Purge "B" 1MD-13 763ft-0 in. 763ft-0 in.

119° 3/4 in. RR Process Sampling 1MD-14 757ft-6 in. 757ft-6 in.

0° 18 in. RHR Shutdown Suction 1MD-15 761ft-3-1/2 in. 761ft-3-1/2 in. 45° 12 in. RHR/LPCI "A" 1MD-16 764ft-0-1/2 in. 764ft-0-1/2 in. 210° 12 in. RHR/LPCI "B" 1MD-17 764ft-0-1/2 in. 764ft-0-1/2 in. 160° 12 in. RHR/LPCI "C" 1MD-18 731ft- 11-5/8 in. 762ft-11 in. 16

° 10 in. Main Steam RV Vents 1MD-19 731ft- 11-9/16 in.

726ft- 11-5/16 in. 37

° 10 in. Main Steam RV Vents 1MD-20 731ft- 11-3/4 in.

726ft- 11-9/16 in. 58

° 10 in. Main Steam RV Vents 1MD-21 731ft- 10-7/8 in.

726ft- 10-1/4 in. 79

° 10 in. Main Steam RV Vents 1MD-22 731ft- 10-7/8 in.

726ft- 10-3/16 in. 101

° 10 in. Main Steam RV Vents 1MD-23 731ft- 11-3/16 in.

726ft- 10-1/8 in. 122

° 10 in. Main Steam RV Vents 1MD-24 731ft- 11-5/16 in.

726ft- 10-13/16 in. 143

° 10 in. Main Steam RV Vents 1MD-25 731ft- 11-1/2 in 726ft- 10-1/2 in. 164

° 10 in. Main Steam RV Vents 1MD-26 732ft- 3/16 in. 726ft- 11-3/8 in. 185

° 10 in. Main Steam RV Vents 1MD-27 731ft- 11-1/4 in.

726ft- 10-7/16 in.

206° 10 in. Main Steam RV Vents 1MD-28 731ft- 10-5/8 in. 726ft-10 in. 238

° 10 in. Main Steam RV Vents 1MD-29 731ft- 11-1/16 in.

726ft- 10-13/16 in. 259

° 10 in. Main Steam RV Vents 1MD-30 731ft- 111/4 in. 726ft-10-5/8 in. 281

° 10 in. Main Steam RV Vents 1MD-31 731ft- 10-15/16 in. 726ft-10-9/16 in. 302

° 10 in. Main Steam RV Vents 1MD-32 731ft- 10-7/16 in.

726ft- 10-15/16 in. 323

° 10 in. Main Steam RV Vents 1MD-33 731ft- 9-3/4 in.

726ft- 10-3/4 in.

344° 10 in. Main Steam RV Vents 1MD-34 29' 8-5/8"* Drywell Head 1MD-35 769ft-51/4 in. 765ft-51/4 in.

270° 10 in. HPCS Pump Discharge 1MD-36 769ft-51/4 in. 769ft-51/4 in.

90° 10 in. LPCS Pump Discharge 1MD-37 764ft- 9-9/16 in.

769ft- 9-1/16 in. 76

°24' 7'-5" x 5'-0"** CRD Insert/Withdrawal Quad 1 1MD-38 764ft- 9-11/16 in.

769ft- 9-3/16 in. 101

°22' 5'-9" x 5'-0"** CRD Insert/Withdrawal Quad 3 1MD-39 764ft-9-9/16 in. 769ft- 9-1/16 in.

256°24' 7'-5" x 5'-0"** CRD Insert/Withdrawal Quad 2 1MD-40 764ft- 9-11/6 in.

769ft- 9-3/16 in. 281

°22' 5'-9" x 5'-0"** CRD Insert/Withdrawal Quad 4 1MD-41 748 ft-0 in. 748 ft-0 in.

298° 12 in Spare (Capped)

• Inside diameter of drywell head. **Dimentions of insert/withdrawal assemblies.

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-63 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-42 756ft- 4-3/4 in.

756ft- 4-3/4 in. 0

° -11ft-0 in. 4 in. RCIC Head Spray 1MD-43 758ft- 1-1/16 in.

758ft- 9/16 in. 0° -4ft- 9-7/16 in. 8 in. Steam to RHR/RCIC 1MD-44 764ft-0-1/2 in. 764ft-0-1/2 in. 204

° 12 in. Structural Integrity Test Cable 1MD-45 756ft-6 in. 756ft-6 in.

0° -15ft-0 in. 3 in. Main Steam Drain 1MD-46 744ft-3 in. 744ft-3 in.

54° 6 in. Component Cooling Water Supply 1MD-47 745ft-9 in. 745ft-9 in.

67° 6 in. Component Cooling Water Return 1MD-48 748ft-0 in. 748ft-0 in.

72° 3/4 in. Nuclear Boiler Instrumentation 1MD-49 747ft-0 in. 747ft-0 in.

54° 4 in. Spare (Capped) 1MD-50 790ft-0 in. 790ft-0 in.

70° 12 in. Spare (Capped) 1MD-51 792ft-8 in. 792ft-8 in.

260° 12 in. For Electrical Use 1MD-52 Not Used 1MD-53 775ft-9 in. 775ft-9 in.

121° 4 in. Drywell Cooling 1MD-54 790ft-0 in. 790ft-0 in.

90° 3 in. D. C. Welding Recpt. 1MD-55 764ft-0-1/2 in. 764ft-0-1/2 in. 154

° 1/2 & 1 in. Recirc. Pump 1A (Hyd. Line to Pwr. Unit) 1MD-56 748ft-0 in. 748ft-0 in.

285° 10 in. Spare (Capped) 1MD-57 764ft-0 in. 764ft-0 in.

45° 3 in. Instrument Air 1MD-58 759ft-0 in. 759ft-0 in.

52° 1 in. Instrument Booster Air 1MD-59 761ft-6 in. 761ft-6 in.

52° 3 in. Service Air 1MD-60 756ft-6 in. 756ft-6 in.

0° +11ft-3 in. 6 in. Reactor Water Clean-up Pump Suct. 1MD-61 748ft-0 in. 748ft-0 in.

170° 10 in. Spare (Capped) 1MD-62 748ft-0 in. 748ft-0 in.

330° 6 in. Rx Recirculation Pump Motor Feed 1MD-63 764ft-0 in. 764ft-0 in.

190° 2 in. Spare (Capped) 1MD-64 790ft-0 in. 790ft-0 in.

138° 18 in. High Range Radiation Monitor 1MD-65 748ft-0 in. 748ft-0 in.

177° 18 in. Spare (Capped) 1MD-66 748ft-0 in. 748ft-0 in.

150° 18 in. Spare (Capped) 1MD-67 758ft-0 in. 758ft-0 in.

182° 6 in. Spare (Capped) 1MD-68 748ft-0 in. 748ft-0 in.

160° 10 in. Spare (Capped) 1MD-69 752ft-6 in. 752ft-6 in.

167° 3 in. Drywell Equipment Drain 1MD-70 750ft-0 in. 750ft-0 in.

299° 3 in. Drywell Floor Drain 1MD-71 762ft-3 in. 762ft-3 in.

0° +2ft-3 in. 3 in. Spare (Capped) 1MD-72 790ft-0 in. 790ft-0 in.

130° 10 in. Drywell Vacuum Breaker 1MD-73 792ft-8 in. 792ft-8 in.

270° 8 in. Spare (Capped) 1MD-74 Not used 1MD-75 Not used 1MD-76 761ft-0 in. 761ft-0 in.

219° 2 in. Spare (Capped) 1MD-77 Not used 1MD-78 784 ft-8 in. 784ft-8 in.

309° 3/4 in. Reactor Pressure Vessel Level "D" 1MD-79 723ft-11 in. 723ft-11 in.

69° 6 in. Hydrogen Bubbler Pipe 1MD-80 723ft-11 in. 723ft-11 in.

249° 6 in. Hydrogen Bubbler Pipe 1MD-81 743ft-0 in. 743ft-8 in.

35° -5ft-0 in. 3/8 in. TIP System A 1MD-82 739ft-1 in. 739ft-9 in.

35° +5ft-0 in. 3/8 in. TIP System B CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-64 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-83 743 ft-0-1/8 in. 743 ft-8 in.

35° 3/8 in. TIP System C 1MD-84 739 ft-1-1/8 in. 739 ft-9 in.

35° 3/8 in. TIP System D 1MD-85 Not Used 1MD-86 739 ft-9 in. 739 ft-9 in.

35° +2 ft-6 in. 1/2 in. TIP System Purge Supply 1MD-87 776 ft-0 in. 776 ft-0 in.

44° 8 in. Spare (Capped) 1MD-88 761 ft-0 in. 761 ft-0 in.

230° 8 in. Spare (Capped) 1MD-89 762 ft-3 in. 762 ft-3 in.

0° -11 ft-9 in. 3/8 & 1/2 in. Post Accident Sample 1MD-90 790 ft-0 in. 790 ft-0 in.

80° 6 in. Spare (Capped) 1MD-91 790 ft-0 in. 790 ft-0 in.

120° 6 in. Spare (Capped) 1MD-92 792 ft-8 in. 792 ft-8 in.

280° 6 in. Spare (Capped) 1MD-93 784 ft-8 in. 784 ft-8 in. 320

°30' Spare (Capped) 1MD-94 764 ft-9 in. 764 ft-9 in.

314° 1/2 to 1 1/2 in. Recirc. Pump 1B - RHR 1MD-95 748 ft-0 in. 748 ft-0 in.

350° 2 in. Spare (Capped) 1MD-96 792 ft-8 in. 792 ft-8 in.

242° 1 in. Instrument Air 1MD-97 776 ft-0 in. 776 ft-0 in.

190° 4 in. Spare (Capped) 1MD-98 Not Used 1MD-99 748 ft-0 in. 748 ft-0 in.

0° 4 in. Spare (Capped) 1MD-100 749 ft-0 in. 749 ft-0 in.

265° 6 in. Drywell Pressure Test 1MD-101 793 ft-9-1/2 in. 793 ft-9-1/2 in. 313

° -5/8 in. 24 in. Drywell Purge Air Inlet 1MD-102 776 ft-11 in. 776 ft-11 in. 135

° +8 ft-0 in. 24 in. Drywell Purge Air Outlet 1MD-103 776 ft-11 in. 776 ft-11 in. 135

° +14 ft-8 in. 2 in. Spare (Capped) 1MD-104 791 ft-3 in. 791 ft-3 in. 304

°30' 24 in. High Range Radiation Detector 1MD-105 792 ft-0 in. 792 ft-0 in.

113° 24 in. Spare (Capped) 1MD-106 792 ft-0 in. 792 ft-0 in.

98° 1 in. Breathing Air 1MD-107 770 ft-0 in. 770 ft-0 in.

219° +11 ft-6 in. 10 in. Drywell Chilled Water Supply 1MD-108 770 ft-0 in. 770 ft-0 in.

219° +8 ft-6 in. 10 in. Drywell Chilled Water Return 1MD-109 771 ft-3 in. 771 ft-3 in.

200° 10 in. Drywell Chilled Water Supply 1MD-110 771 ft-3 in. 771 ft-3 in.

194° 10 in. Drywell Chilled Water Return 1MD-111 Located on Top of Drywell 6 in. From Drywell to H2 Mix. Fan "A" 1MD-112 Located on Top of Drywell 6 in. To Drywell from H2 Mix Fan "A" 1MD-113 785 ft-0 in. 785 ft-0 in.

218° +0 ft-2-1/2 in. 6 in. From Drywell to H2 Mix. Fan "B" 1MD-114 785 ft-0 in. 785 ft-0 in.

225° +0 ft- 2-5/16 in. 6 in. To drywell from H2 Mix. Fan "B" 1MD-115 787 ft-0 in. 787 ft-0 in.

120° 4 in. Spare (Capped)l 1MD-116 785 ft-0 in. 785 ft-0 in.

315° Spare (Capped) 1MD-117 790 ft-0 in. 790 ft-0 in.

45° 10 in. Drywell Vacuum Breaker 1MD-118 790 ft-0 in. 790 ft-0 in.

105° Spare (Capped) 1MD-119 764 ft-3 in. 764 ft-3 in.

215° 10 in. Drywell Vacuum Breaker 1MD-120 764 ft-0 in. 764 ft-0 in.

309° 10 in. Drywell Vacuum Breaker 1MD-121 Not Used 1MD-122 Not Used CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-65 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-123 764 ft-0 in. 764 ft-0 in.

170° 4 in. Spare (Capped) 1MD-124 764 ft-0 in. 764 ft-0 in.

180° 4 in. Fire Protection 1MD-125 766 ft-0 in. 766 ft-0 in.

47° 3 & 1-1/2 in. Cycled Condensate & RHR 1MD-126 763 ft-3 in. 763 ft-3 in.

270° +2 ft-0 in. 4 in. Comp. Cooling Water (Supply) 1MD-127 770 ft-0 in. 770 ft-0 in.

322° 4 in. Comp. Cooling Water (Return) 1MD-128 791'-3" 790'-10" 50

° 3/4 in. Reactor Pressure Level "A" 1MD-129 791'-3" 790'-10" 221

°45' 3/4 in. Reactor Pressure Level "B" 1MD-130 791'-3" 790'-10" 142

°30' 3/4 in. Reactor Pressure Level "C" 1MD-131 791'-3" 790'-10" 313

° 3/4 in. Reactor Pressure Level "D" 1MD-132 783'-3" 782'-10" 233

°30' 3/4 in. Reactor Pressure Level "B" 1MD-133 783'-3" 782'-10" 142

°30' 3/4 in. Reactor Pressure Level "C" 1MD-134 771'-3" 770'-10" 44

° 3/4 in. Reactor Pressure Level "A" 1MD-135 771'-3" 770'-10" 189

°30' 3/4 in. Reactor Pressure Level "B" 1MD-136 771'-3" 770'-10" 157

°15' 3/4 in. Reactor Pressure Level "C" 1MD-137 771'-2-3/4" 770'-9-3/4" 318

°30' 3/4 in. Reactor Pressure Level "D" 1MD-138 Not Used 1MD-139 Not Used 1MD-140 Not Used 1MD-141 748 ft-0 in. 748 ft-0 in.

20° 8 in. 1MD-142 748 ft-0 in. 748 ft-0 in.

292° 8 in. Spare (Capped) Spare (Capped) 1MD-143 Not Used 1MD-144 Not Used 1MD-145 Not Used 1MD-146 1MD-147 748 ft-0 in. 748 ft-0 in.

30° 12 in. Not Used Spare (Capped) 1MD-148 748 ft-0 in. 748 ft-0 in.

270° 12 in. Inservice Inspection Cables 1MD-150 784 ft-11 in. 784 ft-6 in.

135° 3/4 in. Instrumentation 1MD-151 778 ft-6 in. 778 ft-1 in.

115° 3/4 in. Spare (Capped) 1MD-152 765 ft-6 in. 765 ft-1 in.

65° +17 ft-3 in. 3/4 in. Instrumentation 1MD-153 782 ft-4 in. 781 ft-11 in. 222

° 3/4 in. Spare (Capped) 1MD-154 775 ft-11 in. 775 ft-6 in. 197

°30' 3/4 in. Containment Monitoring System 1MD-155 763 ft-4 in. 762 ft-11 in. 167

° +17 ft-3 in. 3/4 in. Instrumentation 1MD-156 781 ft-11 in. 781 ft-6 in. 320

°30' 3/4 in. Spare (Capped) 1MD-157 774 ft-11 in. 774 ft-0 in.

314° 3/4 in. Spare (Capped) 1MD-158 763 ft-4 in. 762 ft-11 in. 288

° +17 ft-3 in. 3/4 in. Spare (Capped) 1MD-159 784 ft-11 in. 784 ft-6 in.

40° 3/4 in. Reactor Press. Vessel Level "A" 1MD-160 778 ft-6 in. 778 ft-1 in.

40° 3/4 in. Reactor Press. Vessel Level "A" 1MD-161 765 ft-6 in. 765 ft-1 in.

35° +17 ft-3 in. 3/4 in. Main Steam "A"/"C" & R.C.I.C. Steam 1MD-162 Not Used 1MD-163 744 ft-0 in. 744 ft-5 in.

173° 3/4 in. Drywell Pressure 1MD-164 744 ft-0 in. 744 ft-5 in.

179° 3/4 in. Recirc. Pump "A" Flow/Leak Detection (1E31-N764)

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-66 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-165 750 ft-0 in. 750 ft-5 in.

188° 3/4 in. Instrumentation 1MD-166 744ft-0 in. 743ft-7 in.

269° 3/4 in. Spare (Capped) 1MD-167 768ft-0 in. 768ft-6 in.

280° +17ft-3 in. 3/4 in. Instrumentation 1MD-168 765ft-0 in. 765ft-6 in.

28° +17ft-3 in. 3/4 in. Instrumentation 1MD-169 765ft-0 in. 764ft-6 in.

302° -17ft-3 in. 3/4 in. HPCS Leak Detection 1MD-170 744ft-0 in. 744ft-5 in.

45° 3/4 in. Instrumentation 1MD-171 744ft-0 in. 744ft-5 in.

75° 3/4 in. Reactor Water Cleanup Flow 1MD-172 Not used 1MD-173 744ft-0 in. 743ft-7 in.

155° 3/4 in. Instrumentation 1MD-174 771ft-0 in. 770ft-6 in.

272° -17ft-3 in. 1-1/2 in. Spare (Capped) 1MD-175 747ft-0 in. 747ft-5 in.

264° 3/4 in. Instrumentation 1MD-176 785ft-6 in. 785ft-1 in.

256° 3/4 in. Instrumentation 1MD-177 747ft-0 in. 747ft-5 in.

188° 3/4 in. Recirc. Pump "A" Flow 1MD-178 744ft-0 in. 744ft-5 in.

290° 3/4 in. Drywell Pressure 1MD-179 764ft-6 in. 764ft-0 in.

272° -17ft-3 in. 3/4 in. Instrumentation 1MD-180 744ft-0 in. 744ft-5 in.

295° 3/4 in. Recirc. Pump "B" Flow 1MD-181 747ft-0 in. 747ft-5 in.

155° 3/4 in. Recirc. Pump "A" #1 & 2 Seal Cavity 1MD-182 765ft-0 in. 764ft-6 in.

60° +17ft-3 in. 3/4 in. RI/RH Leak Detection System 1MD-183 764ft-0 in. 763ft-6 in.

152° -17ft-3 in. 3/4 in. Instrumentation 1MD-184 764ft-0 in. 763ft-6 in.

158° -17ft-3 in. 3/4 in. Spare (Capped) 1MD-185 765ft-0 in. 764ft-6 in.

236° +17ft-3 in. 3/4 in. Main Steam Line "D" 1MD-186 765ft-ft-0 in. 764ft-6 in. 274

° +17ft-3 in. 3/4 in. Main Steam Flow 1MD-187 747ft-0 in. 746ft-7 in.

145° 3/4 in. Instrumentation 1MD-188 749ft-0 in. 749ft-6 in.

295° 3/4 in. Recirc. Pump "B" #1 & 2 Seal Cavity 1MD-189 744ft-0 in. 744ft-5 in.

276° 3/4 in. Spare (Capped) 1MD-190 771ft-0 in. 771ft-6 in.

173° -17ft-3 in. 3/4 in. Spare (Capped) 1MD-191 771ft-0 in. 771ft-6 in.

212° -17ft-3 in. 3/4 in. Spare (Capped) 1MD-192 761ft-6 in. 761ft-0 in.

91° 3/4 in. Instrumentation 1MD-193 771ft-ft-0 in. 771ft-6 in. 134

° -17ft-3 in. 3/4 in. Spare (Capped) 1MD-194 771ft-0 in. 771ft-6 in.

167° -17ft-3 in. 3/4 in. Containment Monitoring System 1MD-195 768ft-0 in. 768ft-6 in.

173° -17ft-3 in. 3/4 in. Spare (Capped) 1MD-196 771ft-0 in. 770ft-6 in.

180° -17ft-3 in. 1 in. Spare (Capped) 1MD-197 765ft-0 in. 764ft-6 in.

120° +17ft-3 in. 3/4 in. Spare (Capped) 1MD-198 751ft-0 in. 750ft-7 in.

155° 3/4 in. Spare (Capped) 1MD-199 771ft-0 in. 770ft-6 in.

287° 1 in. Spare (Capped) 1MD-200 765ft-ft-0 in. 764ft-6 in. 208

° +17ft-3 in. 3/4 in. Main Steam (Loop "A") - RHR (Loop "C") 1MD-201 750ft-0 in. 749ft-6 in.

262° 3/4 in. Instrumentation 1MD-202 Not used 1MD-203 747ft-0 in. 747ft-7 in.

340° 3/4 in. Spare (Capped) 1MD-204 747ft-5 in. 747ft-0 in.

355° 3/4 in. Instrumentation 1MD-205 768ft-0 in. 768ft-6 in.

270° +17ft-3 in. 3/4 in. Drywell Pressure 1MD-206 761ft-6 in. 761ft-0 in.

57° +17ft-3 in. 3/4 in. Spare (Capped)

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-67 REV. 11, JANUARY 2005 DRYWELL MECHANICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1MD-207 750ft-0 in. 749ft-6 in.

75° 3/4 in. Spare (Capped) 1MD-208 777ft-0 in. 777ft-6 in.

165° 3/4 in. Spare (Capped) 1MD-209 774ft-0 in. 774ft-6 in.

155° 3/4 in. Spare (Capped) 1MD-210 777ft-0 in. 777ft-6 in.

155° 3/4 in. Spare (Capped) 1MD-211 777ft-0 in. 777ft-6 in.

170° 3/4 in. Spare (Capped)

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-68 REV. 11, JANUARY 2005 DRYWELL ELECTRICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1ED-1 791ft- 10-23/32 in.

791ft- 10-23/32 in.

62.5° 6 in. Neutron Monitoring 1K1N 1ED-2 791ft- 11-1/32 in 791ft- 11-1/32 in.

64° 6 in. Neutron Monitoring 1K1N 1ED-3 793ft-3 in. 793ft-3 in.

62.5° 4 in. Instrumentation 1K1E 1ED-4 793ft-3 in. 793ft-3 in.

63.5° 4 in. Instrumentation 1K1E 1ED-5 775ft-0 in. 775ft-0 in.

36° 6 in. Medium Voltage 6900 V Recirc. Pump 1P1B 1ED-6 775ft-0 in. 775ft-0 in.

37.5° 6 in. Medium Voltage 6900 V Recirc. Pump 1P1B 1ED-7 774ft-0 in. 774ft-0 in.

44° 4 in. Instrumentation 1K1B 1ED-8 774ft-0 in. 774ft-0 in.

45° 4 in. Instrumentation 1K1B 1ED-9 794ft-9 in. 794ft-9 in.

62.5° 4 in. Control 1C1E 1ED-10 794ft-9 in. 794ft-9 in.

63.5° 4 in. Control 1C1E 1ED-11 796ft-3 in. 796ft-3 in.

62.5° 4 in. L.V. Power 1P1E 1ED-12 796ft-3 in. 796ft-3 in.

63.5° 4 in. L.V. Power 1P1E 1ED-13 796ft-3 in. 796ft-3 in.

64.5° 4 in. L.V. Power 1P1E 1ED-14 774ft-0 in. 774ft-0 in.

52° 4 in. L.V. Power 1P1B 1ED-15 774ft-0 in. 774ft-0 in.

53° 4 in. L.V. Power 1P1B 1ED-16 774ft-0 in. 774ft-0 in.

54° 4 in. L.V. Power 1P1B 1ED-17 774ft-0 in. 774ft-0 in.

48° 4 in. Control 1C1B 1ED-18 774ft-0 in. 774ft-0 in.

49° 4 in. Control 1C1B 1ED-19 774ft-9 in. 774ft-9 in.

48° 4 in. Control 1C1B 1ED-20 774ft-9 in. 774ft-9 in.

49° 4 in. Control 1C1B 1ED-21 773ft-9 in. 773ft-9 in.

135° -11ft- 4-7/32 in. 4 in. Control 1C3E 1ED-22 773ft-9 in. 773ft-9 in.

135° -11ft- 11-19/32 in. 4 in. Control 1C3E 1ED-23 773ft-9 in. 773 ft-9 in.

135° -12ft- 6-15/16 in. 4 in. Control 1C3E 1ED-24 792ft-6 in. 792ft-6 in.

157° -13ft- 1-1/16 in. 6 in. Neutron Monitoring 1K3N 1ED-25 792ft-6 in. 792ft-6 in.

157° -12ft- 3-1/16 in. 6 in. Neutron Monitoring 1K3N 1ED-27 773ft-9 in. 773ft-9 in.

135° -13ft- 1-19/32 in. 4 in. Control 1C3E 1ED-28 773ft-9 in. 773ft-9 in.

135° -13ft- 9-1/8 in. 4 in. Control 1C3E 1ED-29 773ft-9 in. 773ft-9 in.

203° 6 in. Instrumentation 1K2E 1ED-30 773ft-9 in. 773ft-9 in.

205° 6 in. Instrumentation 1K2E 1ED-31 776ft-3 in. 776ft-3 in.

203° 4 in. Control 1C2E 1ED-32 776ft-3 in. 776ft-3 in.

204° 4 in. Control 1C2E 1ED-33 776ft-3 in. 776ft-3 in.

205° 4 in. Control 1C2E 1ED-34 774ft- 2-3/8 in.

774ft- 2-3/8 in.

242° 4 in. L.V. Power 1P2E 1ED-35 774ft- 2-5/32 in.

774ft- 2-5/32 in.

243° 4 in. L.V. Power 1P2E 1ED-36 774ft- 2-3/16 in.

774ft- 2-3/16 in.

244° 4 in. L.V. Power 1P2E 1ED-37 773ft-9 in. 773ft-9 in. 300.5

° 4 in. Instrumentation 1K2B 1ED-38 773ft-9 in. 773ft-9 in. 301.5

° 4 in. Instrumentation 1K2B 1ED-39 773 ft-9 in. 773ft-9 in. 302.5

° 4 in. Instrumentation 1K2B 1ED-40 773ft-9 in. 773ft-9 in. 303.5

° 4 in. Instrumentation 1K2B 1ED-41 773ft-9 in. 773ft-9 in. 304.5

° 4 in. Instrumentation 1K2B 1ED-43 774ft-0 in. 774ft-0 in. 310.5

° 4 in. Control 1C2B CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-69 REV. 11, JANUARY 2005 DRYWELL ELECTRICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1ED-44 774ft-0 in. 774ft-0 in. 311.5

° 4 in. Control 1C2B 1ED-45 774ft-0 in. 774ft-0 in. 312.5

° 4 in. Control 1C2B 1ED-46 774ft-0 in. 774ft-0 in. 306.5

° 4 in. L.V. Power 1P2B 1ED-47 774ft-0 in. 774ft-0 in. 307.5

° 4 in. L.V. Power 1P2B 1ED-48 774ft-0 in. 774ft-0 in. 308.5

° 4 in. L.V. Power 1P2B 1ED-49 773ft-9 in. 773ft-9 in. 296.5

° 6 in. Spare 1ED-50 773ft-9 in. 773ft-9 in.

298° 6 in. Control 1C2E 1ED-51 774ft-0 in. 774ft-0 in. 315.5

° 6 in. Neutron Monitoring 1K4N 1ED-52 774ft-0 in. 774ft-0 in.

317° 6 in. Neutron Monitoring 1K4N 1ED-53 774ft-0 in. 774ft-0 in. 318.5

° 4 in. Instrumentation 1K4E 1ED-54 774ft-0 in. 774ft-0 in.

320° 4 in. Instrumentation 1K4E 1ED-55 774ft-0 in. 774ft-0 in.

322° 4 in. Control 1C4E 1ED-56 774ft-0 in. 774ft-0 in.

323° 4 in. Control 1C4E 1ED-57 774ft-0 in. 774ft-0 in.

325° 4 in. Control 1C4E 1ED-58 773ft-9 in. 773ft-9 in.

135° -16ft- 6-3/32 in. 4 in. Instrumentation 1K3E 1ED-59 773ft-9 in. 773ft-9 in.

135° -15ft- 8-3/8 in. 4 in. Instrumentation 1K3E 1ED-60 774ft-9 in. 774ft-9 in.

203° 4 in. Control 1C2E 1ED-61 774ft-9 in. 774ft-9 in.

205° 4 in. Neutron Monitoring 1K2N 1ED-64 792ft-0 in. 792ft-0 in. 103.5

° 6 in. Control Rod Drive Position Indication 1K1E 1ED-65 792ft-0 in. 792ft-0 in.

105° 6 in. Control Rod Drive Position Indication 1K1E 1ED-66 792ft-0 in. 792ft-0 in. 106.5

° 6 in. Control Rod Drive Position Indication 1K1E 1ED-67 792ft-0 in. 792ft-0 in.

108° 6 in. Control Rod Drive Position Indication 1K1E 1ED-68 794ft-3 in. 794ft-3 in.

303° 6 in. Control Rod Drive Position Indication 1K2E 1ED-69 794ft-3 in. 794ft-3 in. 304.5

° 6 in. Control Rod Drive Position Indication 1K2E 1ED-70 794ft-3 in. 794ft-3 in.

306° 6 in. Control Rod Drive Position Indication 1K2E 1ED-71 794ft-3 in. 794ft-3 in. 307.5

° 6 in. Control Rod Drive Position Indication 1K2E 1LD-1 751ft-0 in. 751ft-0 in.

47° 3 in. Ltg./Comm.

1LD-2 751ft-0 in. 751ft-0 in.

48° 3 in. Ltg./Comm.

1LD-3 751ft-0in. 751ft-0 in. 49

° 3 in. Ltg./Comm. 1LD-4 750ft-0 in. 750ft-0 in.

273° 3 in. Ltg./Comm.

1LD-5 750ft-0 in. 750ft-0 in.

274° 3 in. Ltg./Comm.

1LD-6 750ft-0 in. 750ft-0 in.

275° 3 in. Ltg./Comm.

1LD-7 775ft-0 in. 775ft-0 in.

57° 3 in. Ltg./Comm.

1LD-8 775ft-0 in. 775ft-0 in.

58° 3 in. Ltg./Comm.

1LD-9 743ft-0 in. 743ft-0 in.

5° 3 in. Ltg./Comm.

1LD-10 774ft-0 in. 774ft-0 in.

57° 3 in. Ltg./Comm.

1LD-11 774ft-0 in. 774ft-0 in.

58° 3 in. Ltg./Comm.

1LD-12 739ft-3 in. 739ft-3 in.

142° 3 in. Ltg./Comm.

1LD-13 742ft-6 in. 742ft-6 in.

178° 3 in. Ltg./Comm.

1LD-14 784ft-0 in. 784ft-0 in.

58° 3 in. Ltg./Comm.

1LD-15 774ft-0 in. 774ft-0 in.

279° 3 in. Ltg./Comm.

CPS/USAR TABLE 3.8-5 CONTAINMENT AND DRYWELL PENETRATIONS (ContInued) CHAPTER 03 3.8-70 REV. 11, JANUARY 2005 DRYWELL ELECTRICAL PENETRATIONS No. Centerline Elevation Inside Centerline Elevation Outside Azimuth Distance Parallel to Azimuth Nom. Pipe Size Description 1LD-16 774ft-0 in. 774ft-0 in.

280° 3 in. Ltg./Comm.

1LD-17 774ft-0 in. 774ft-0 in.

281° 3 in. Ltg./Comm.

1LD-18 774ft-0 in. 774ft-0 in.

282° 3 in. Ltg./Comm.

CPS/USAR CHAPTER 03 3.8-71 REV. 11, JANUARY 2005 TABLE 3.8-6 PREDICTED DEFORMATION OF THE CONTAINMENT DURING THE PRESSURE TEST*

ITEM LOCATION METER NO.** DEFLECTION (IN) Radial deflec-tion of cylin-

der wall 10 feet above base

48 feet above base

Midheight of cylinder 123 feet above base

Dome springline

Equipment hatch RC1-RC6 RC7-RC12 RC13-RC18 DC19-DC21 DC22-DC24 RHC25-RHC36 0.16 0.43 0.43 0.86***

0.72***

0.43 Vertical de-

flection of

dome with reference to

cylinder base Dome springline

1st intermediate point (at El. 909 ft)

2nd intermediate point (at El. 920 ft)

Dome apex VC1-VC6 VC7 VC8 VC9 0.25 0.42

0.45

0.45

• Test pressure is 17.25 psig.

• • The locations of the deflection meters are shown on Drawing S27-1401.

• • • The value given is the diametrical deflection of the dome springline.

CPS/USAR CHAPTER 03 3.8-72 REV. 11, JANUARY 2005 TABLE 3.8-7.1 PREDICTED DEFORMATION OF THE DRYWELL FOR THE PRESSURE TEST*

ITEM LOCATION METER NO.** DEFLECTION (in.) 10 ft. above base RD1 - RD3 0.105 Radial deflection 48 ft. above base RD4 - RD6 0.254 of drywell wall Top of drywell RD7 - RD9 0.098

{Equipment hatch region RHD10 - RHD21 0.238 Base to top of drywell VD1 - VD3 0.250 Vertical deflection

{Top of shield wall

to top of drywell VD4 - VD6 0.800

• The test pressure is 30.0 psig. ** The locations of the deflection meters are shown on Drawing S27-1401.

CPS/USAR CHAPTER 03 3.8-73 REV. 11, JANUARY 2005 TABLE 3.8-7.2 PREDICTED STRAINS OF THE DRYWELL FOR THE PRESSURE TEST*

STRAIN ITEM LOCATION GAUGE NO.** INSIDE FACE OUTSIDE FACE Hoop strain of 10 feet above base SD1 - SD2 0.00027 0.00027 drywell wall 48 feet above base SD3 - SD4 0.00059 0.00062 Vertical strain of 10 feet above base SD1 - SD2 0.00007 0.00001 drywell wall 48 feet above base SD3 - SD4 0.00002 0.00004

• The test pressure is 30.0 psig. ** The locations of the strain gauges are shown in Drawing S27-1401.

CPS/USAR CHAPTER 03 3.8-74 REV. 11, JANUARY 2005 TABLE 3.8-8 MATERIAL PROPERTIES FOR CONTAINMENT ULTIMATE CAPACITY STUDY Material Specified Strength Average Tested Strength Standard Deviation Concrete 4,000 psi 6,086 psi

1. 11 rein. steel 60 ksi 68.6 ksi 3.3 ksi #14 rein. steel 60 ksi 73.3 ksi 2.9 ksi #18 rein. steel 60 ksi 71.1 ksi 2.6 ksi Carbon steel liner (1/4 in.) 32 ksi 48.4 ksi 3.3 ksi Stainless steel liner 30 ksi 43.0 ksi 3.7 ksi (1/4 in. and 1/2 in.)

CPS/USAR CHAPTER 03 3.8-75 REV. 11, JANUARY 2005 TABLE 3.8-9 CODES USED FOR DESIGN AND CONSTRUCTION OF STRUCTURAL ITEMS INSIDE CONTAINMENT 1,3 DESIGN MATERIAL FABRICATION ERECTION ANI STAMPING ITEM SPECIFICATION CONTROL 2. WELDING EXAMINATION WELDING EXAMINATION 1. CONTAINMENT

a. Liner backed by concrete ASME, (proposed) Section III, Division 2 (1973) ASME, Section II (1971,Sum.'73) CC-2500, ASME, Sec. III, Div. 2 (Proposed-1973) CC-4500, ASME, Sec. III, Div. 2 (Proposed-1973) and ASME, Section IX (1971, Summer '73) NRC Regulatory Guide 1.19 Rev. 1 and NE-5000, ASME, Sec. III, Div. 1 (1971, Sum. '73) CC-4500,ASME, Section III, Div. 2 (Proposed - 1973) and ASME, Section IX (1971, Summer '73) NRC Regulatory Guide 1.19, Rev. 1 and NE-5000, ASME, Sec. III, Div. 1 (1971, Sum. '73)

No No b. Pipe Penetration Sleeves ASME, Section III Division 1 (1974 with Summer '74 Addenda), Sub- section NE ASME, Section II (1971, Sum.

'73) NE-2000 ASME, Sec. III, Div. 1 (1971, Sum. '73) ASME, Section IX (1971, Summer '73)

NE-2000, ASME, Sec. III, Div. 1 (1971, Sum. '73) NE-4000, ASME, Sec. III, Div. 1 (1971, Summer '73) NE-5000, ASME, Sec. III, Div. 1 (1971, Summer '73) Yes No c. Personnel Locks ASME, Section III Division 1 (1971 with Summer '73 Addenda), Sub-section NE ASME, Section II

(1971, Summer

'73) NA-4000 &

NE-2000, ASME, Sec. III, Div. 1 (1971, Sum. '73) NE-4000, ASME, Sec. III, Div. 1 (1971, Sum. '73)

and ASME, Section IX (1971, Sum. '73) NE-5000, ASME, Sec. III, Div. 1 (1971, Sum. '73) NE-4000, ASME, Sec. III, Div. 1 (1971, Sum. '73)

and ASME, Section IX (1971, Sum. '73) NE-5000, ASME, Sec. III, Div. 1 (1971, Sum. '73) Yes Yes d. Equipment Hatch ASME, Section III Division 1 (1971 with Summer '73 Addenda), Sub-section NE ASME, Section II

(1971, Summer

'73) NA-4000 &

NE-2000, ASME, Sec. III, Div. 1 (1971, Sum. '73) NE-4000, ASME, Sec. III, Div. 1 (1971, Sum. '73)

and ASME, Section IX (1971, Sum. '73) NE-5000, ASME, Sec. III, Div. 1 (1971, Sum. '73) NE-4000, ASME, Sec. III, Div. 1 (1971, Sum. '73)

and ASME, Section IX (1971, Sum. '73)

NE-5000, ASME Sec. III, Div. 1 (1971, Sum. '73) Yes No 2. DRYWELL a. Suppression Pool Liner (Backed by Concrete)

ASME, (Proposed)

Section III, Division 2 (1973) ASME, Section II (1974, Winter

'75) Equivalent to NB-2000 & NA-4000, ASME, Sec.

III, Div. 1 (1974 Edition, Summer '74) Equivalent to NB-4000, ASME, Sec. III, Div. 1 (1974, Summer '74) and ASME Section IX (1974 Edition, Winter 1975)

or AWS D1.1, Rev. 1-76 NB-5000, ASME, Sec. III, Div. 1 (1974, Win. '75)

and ASME, Section V (1974 Edition, Winter '75) NB-4000, ASME, Sec. III, Div. 1 (1974, Summer '74)

and ASME, Section IX (1974, Summer '74)

or AWS D1.1, Rev. 1-'76 NB-5000, ASME, Sec. III, Div. 1 (1974, Summer '75)

and ASME, Section V (1974, Summer

'75) No No CPS/USAR TABLE 3.8-9 CODES USED FOR DESIGN AND CONSTRUCTION OF STRUCTURAL ITEMS INSIDE CONTAINMENT 1,3 (Continued) CHAPTER 03 3.8-76 REV. 11, JANUARY 2005 DESIGN MATERIAL FABRICATION ERECTION ANI STAMPING ITEM SPECIFICATION CONTROL 2. WELDING EXAMINATION WELDING EXAMINATION b. Form Plate AISC 1969 or 1978 ASTM Certified Material Test Reports Only ASME, Sec. IX (1974, Win. '75)

OR AWS D1.1, Rev 1-'76 Not Applicable ASME, Section IX (1974 Edition, Summer '74)

OR AWS D1.1-Rev. 1, 1976 NB-5000, ASME, Sec. III, Div. 1 (1974 Edition, Summer '75)

and ASME, Section V (1974, Summer '75)

No No c. Pipe Penetration Sleeves ASME, Section III, Div. 1 (1974, Summer 1974), Subsection NF ASME, Section II (1974, Sum. '74)

NF-2000& NA-4000, ASME, Section III, Div. 1 (1974, Sum. '74) ASME, Sec. III (1974, Sum. '74)

AND ASME, Section IX (1974, Sum. '74) NF-2000, ASME, Section III, Division 1 (1974, Sum. '74) ASME, Section IX (1974, Sum. '74)

OR AWS D1.1, Rev. 1-'76 NB-5000, ASME, Sec. III, Div. 1 (1974, Sum. '75) and ASME, Section V (1974, Sum. '75)

No No d. Personnel Lock & Hatch, Drywell

Head ASME, Section III, Division 1, Sub- section NE, (1974 Edition, Summer

'76 Addenda) ASME, Section II

(1974, Summer

'76) NE-2000 & NA-4000, ASME, Sec. III, Div. 1 (1974, Sum. '76)

NE-4000, ASME Sec. III, Div. 1 (1974, Sum. '76) and ASME, Section IX (1974, Sum. '76)

NE-5000, ASME Sec. III, Div. 1 (1974, Sum. '76) and ASME, Sec. V (1974, Sum. '76) NE-4000, ASME, Sec. III, Div. 1 (1974, Sum. '76 and ASME, Sec. IX (1974, Sum. '76)

or AWS D1.1, Rev. 1-'76 NB-5000, ASME, Sec. III, Div. 1 (1974, Sum. '75) and ASME, Section V (1974, Sum. '76) Yes No e. Refueling Bellows ASME, Section III, Division 1, Sub-section NE, (1977 Edition, Winter

'77 Addenda) ASME, Sec. II, (1977, Winter

'77) NE-2000 & NA- 4000, ASME, Sec. III, Div. 1 (1977, Winter

'77) NE-4000, ASME, Sec. III, Div. 1 (1977, Win. '77)

and ASME, Section IX (1977, Win. '77) NE-5000, ASME, Sec. III, Div. 1 (1977, Win. '77)

and ASME, Section V (1977, Win. '77) NE-4000, ASME, Sec. III, Div. 1 (1974, Sum. '74)

and ASME, Section IX (1974, Sum. '74) NE-5000, ASME, Sec. III, Div. 1 (1974, Sum. '74)

and ASME, Section V (1974, Sum. '74)

No No 3. REACTOR PEDESTAL ASME, Section III, Division 1, Sub- section NF (1974 Edition, Winter

'75 Addenda) ASME, Sec. II (1974, Winter

'75) NA-4000, NF- 2000, ASME, Section III, Div. 1 (1974, Win. '75) NF-4000, ASME, Sec. III, Div. 1 (1974, Win. '75)

and ASME, Section IX (1974, Winter '75) NF-5000, ASME, Sec. III, Div. 1 (1974, Win. '75) and ASME, Section V (1974, Win. '75) NF-4000, ASME, Sec. III, Div. 1 (1974, Win. '75)

and ASME Sec. IX (1974, Win. '75) NF-5000, ASME, Sec. III, Div. 1 (1974, Win. '75)

and ASME, Section V (1974, Win. '75)

No No CPS/USAR TABLE 3.8-9 CODES USED FOR DESIGN AND CONSTRUCTION OF STRUCTURAL ITEMS INSIDE CONTAINMENT 1,3 (Continued) CHAPTER 03 3.8-77 REV. 11, JANUARY 2005 DESIGN MATERIAL FABRICATION ERECTION ANI STAMPING ITEM SPECIFICATION CONTROL 2. WELDING EXAMINATION WELDING EXAMINATION 4. WEIR WALL LINER ASME, (Proposed) Section III, Division 2 (1973) ASME, Sec. II, (1974, Win. '75) Equivalent to NB-2000 & NA-4000, ASME, Sec. III, Div. 1 (1974, Sum. '74) ASME, Section IX (1974, Win. '75)

OR AWS D1.1, Rev. 1-

1976 Equivalent to CC-5500, ASME, Sec. III, Div. 2 (Proposed, 1973) and NB-5000, ASME, Sec. III, Div.1 (1974, Sum. '74) ASME, Section IX (1974, Sum. '74)

OR AWS D1.1, Rev. 1-

1976 Equivalent to CC-5500, ASME, Sec. III, Div. 2 (Proposed, 1973) and NB-5000, ASME, Sec. III, Div. 1 (1974, Sum. '75) and ASME Sec. V (1974, Sum. '75)

No No 5. REACTOR SHIELD WALL AISC 1969 or

1978 ASTM Equivalent to NB-2500 & NA-4000, ASME, Sec. III, Div. 1 (1974, Win. '75) ASME, Section IX (1974, Winter '75)

OR AWS D1.1, Rev. 1-

1976 NB-5000, ASME Sec. III, Div.1 (1974, Win. '75) and ASME, Section V (1974, Win. '75) ASME, Section IX (1974, Sum. '74)

OR AWS D1.1, Rev. 1-

1976 NB-5000, ASME, Sec. III, Div. 1 (1974, Sum. '75) and ASME, Section V (1974, Sum. '75)

No No 6. CONTAINMENT POOL a. Stainless Steel Liner ASME, Section III Division 2 (1977) ASME, Sec. II (1977 Edition) Equivalent to NB-2000 & NA-4000, ASME, Sec. III, Div. 1 (1977) CC-4000, ASME, Sec. III, Div. 2

(1977) and ASME, Section IX

(1977) Equivalent to CC-5500, ASME, (1977) and ASME, Section V

(1977) CC-4000, ASME, Sec. III, Div. 2 and ASME, Section IX (1977) Equivalent to CC-5500, ASME, (1977) and ASME, Section V

(1977) No No b. Pool Gates ASME, Section III Division 1 (1977)

Subsection ND ASME, Sec. II (1977 Edition) ND-2000 & NA 4000, ASME, Section III, Division 1 (1977) ND-4000, ASME, Sec. III, Div. 1

(1977) and ASME, Section IX

(1977) Equivalent to CC-5500, ASME, Sec. III, Div. 2 (1977) and ASME, Section V

(1977) ND-4000, ASME, Sec. III, Div. 1

(1977) and ASME Sec. IX

(1977) Equivalent to CC-5500, ASME, Sec. III, Div. 2 (1977) and ASME, Section V

(1977) No No 7. PIPE WHIP RESTRAINTS AISC (1969 or

1978) ASTM NF-2000 & NA- 4000, ASME, Sec. III, Div. 1 (1977, Win. '78) NF-4000, ASME, Sec. III, Div. 1 (1974, Win. '75) and ASME, Section IX (1974, Win. '75) Equivalent to NF-5000, ASME, Sec. III, Div. 1 (1974, Win. '75) and ASME, Sec. V (1974, Win. '75) NF-4000, ASME, Sec. III, Div.1 (1974, Winter '75) and ASME, Section IX (1974, Winter '75) Equivalent to NF-5000, ASME, Sec. III, Div. 1 (1974, Winter '75) and ASME, Section V (1974, Win. '75)

No No CPS/USAR TABLE 3.8-9 CODES USED FOR DESIGN AND CONSTRUCTION OF STRUCTURAL ITEMS INSIDE CONTAINMENT 1,3 (Continued) CHAPTER 03 3.8-78 REV. 11, JANUARY 2005 DESIGN MATERIAL FABRICATION ERECTION ANI STAMPING ITEM SPECIFICATION CONTROL 2. WELDING EXAMINATION WELDING EXAMINATION 8. CONCRETE AND REBAR a. Concrete (Includes containment wall, drywell wall, weir wall, contain-ment pool walls &

refueling floor slab)

ASME, (Proposed) Section III, Division 2 (1973) ACI Codes and Standards (Refer to Table 3.8-4) ANSI N45.2.5, Draft 3, Rev. 1 Jan. 1974 Not Applicable Not Applicable Not Applicable Not Applicable No No b. Rebar (Includes same items as above)

ASME, (Proposed) Section III, Division 2 (1973)

ASTM A615 (Refer to Table 3.8-4) Certified Material Test Reports Only CRSI (Refer to Table 3.8-4) CRSI (Refer to Table 3.8-4) NRC Regulatory Guide 1.10, Rev. 1 and CC 4300, ASME Sec.

III, Div. 2 (1975) NRC Regulatory Guide 1.10, Rev. 1 and CC 4300, ASME ec. III, Div. 2 (1975)

No No NOTES: 1. References to the ASME Code paragraphs made throughout this table are for technical requirements only; administrative requirements, such as preparation of Design Specification, Design Report, Code Data Report, ANI involvement or stamping, are not included unless noted in last two columns. 2. Material Control includes requirements for Certified Material Test Reports, impacts, identification & traceability, unless otherwise indicated. 3. Table contents may include non-code reference information.

CPS/USAR ATTACHMENT A3.8 CONTAINMENT DESIGN LOADS

CPS/USAR CHAPTER 03 A3.8-i REV. 11, JANUARY 2005 ATTACHMENT A3.8 - CONTAINMENT DESIGN LOADS TABLE OF CONTENTS PAGE A3.

8.1 INTRODUCTION

A3.8-1 A3.8.2 DEVELOPMENT OF SRV LOADS A3.8-1 A3.8.2.1 Description of the Phenomena A3.8-2 A3.8.2.2 Analytical Models for Rams Head Loads A3.8-2 A3.8.2.2.1 Bubble Dynamic Model A3.8-3 A3.8.2.2.2 Method of Images Model A3.8-4 A3.8.2.2.3 Vent Clearing Model A3.8-7 A3.8.2.2.4 Assumptions Used In the Implementation of the Analytical Model A3.8-7 A3.8.2.2.5 Method of Implementation A3.8-7 A3.8.2.2.6 Selection of Discharge Cases A3.8-9 A3.8.2.2.6.1 Symmetric Discharge Case A3.8-9 A3.8.2.2.6.2 Asymmetric Discharge Case A3.8-9 A3.8.2.3 Quencher Design Loads on Suppression Pool Boundaries A3.8-10 A3.8.3 DEVELOPMENT OF LOCA LOADS A3.8-10 A3.8.3.1 Description of LOCA (DBA) A3.8-10 A3.8.3.2 Description of LOCA (IBA) A3.8-14 A3.8.3.3 Description of LOCA (SBA) A3.8-14 A3.8.3.4 Design Loads for LOCA Events A3.8-15 A3.8.3.4.1 Water Jet Loads A3.8-16 A3.8.3.4.2 LOCA Air Bubble Loads A3.8-16 A3.8.3.4.2.1 LOCA Air Bubble Loads on the Weir Annulus and Vent System A3.8-16 A3.8.3.4.2.2 LOCA Air Bubble Loads on the Exterior of the Drywell Wall A3.8-16 A3.8.3.4.2.3 LOCA Air Bubble Loads on the Containment A3.8-17 A3.8.3.4.2.4 LOCA Air Bubble Loads on the Basemat A3.8-17 A3.8.3.4.3 Pool Swell Drag and Impact Loads A3.8-17 A3.8.3.4.4 Fallback Loads A3.8-18 A3.8.3.4.5 Froth Impingement Loads on the HCU Floor A3.8-18 A3.8.3.4.6 Condensation Oscillation Loads A3.8-18 A3.8.3.4.7 Chugging Loads A3.8-19 A3,8.3.4.7.1 Chugging Loads Applied to Top Vent A3.8-20 A3.8.3.4.7.2 Pool Boundary Chugging Loads A3.8-21 A3.8.3.4.7.3 Cyclic Temperature Due to Chugging A3.8-22 A3.8.3.4.7.4 Suppression Pool Thermal Stratification A3.8-22 A3.8.4 DEFINITION OF OTHER LOADS A3.8-22 A3.8.5 LOAD COMBINATIONS A3.8-22 CPS/USAR TABLE OF CONTENTS (CONT'D)

CHAPTER 03 A3.8-ii REV. 11, JANUARY 2005 PAGE A3.8.6 ANALYSIS METHODS A3.8-23 A3.8.6.1 General A3.8-23 A3.8.6.1.1 Analysis for SRV Loading A3.8-23 A3.8.6.1.1.1 Rams Head Discharge A3.8-24 A3.8.6.1.1.1.1 Loading A3.8-24 A3.8.6.1.1.1.2 Model for Analysis A3.8-24 A3.8.6.1.1.1.3 Method of Analysis A3.8-24 A3.8.6.1.1.2 Quencher Discharge A3.8-25 A3.8.6.1.1.2.1 Loading A3.8-25 A3.8.6.1.1.2.2 Model for Analysis A3.8-25 A3.8.6.1.1.2.3 Method of Analysis A3.8-25 A3.8.6.2 Analysis for LOCA Loads A3.8-26 A3.8.6.2.1 Annulus Pressurization A3.8-26 A3.8.6.2.2 LOCA Bubble A3.8-26 A3.8.6.2.2.1 Loading A3.8-26 A3.8.6.2.2.2 Model for Analysis A3.8-26 A3.8.6.2.2.3 Method of Analysis A3.8-26 A3.8.6.2.3 Froth Impingement A3.8-27 A3.8.6.2.3.1 Loading A3.8-27 A3.8.6.2.3.2 Model for Analysis A3.8-27 A3.8.6.2.3.3 Method of Analysis A3.8-27 A3.8.6.2.4 Condensation Oscillation A3.8-27 A3.8.6.2.4.1 Loading A3.8-27 A3.8.6.2.4.2 Model for Analysis A3.8-27 A3.8.6.2.4.3 Method of Analysis A3.8-27 A3.8.6.2.5 Chugging A3.8-28 A3.8.6.2.5.1 Loading A3.8-28 A3.8.6.2.5.2 Model for Analysis A3.8-28 A3.8.6.2.5.3 Method of Analysis A3.8-28 A3.

8.7 REFERENCES

A3.8-28 CPS/USAR ATTACHMENT A3.8 - CONTAINMENT DESIGN LOADS CHAPTER 03 A3.8-iii REV. 11, JANUARY 2005 LIST OF TABLES NUMBER TITLE PAGE A3.8-1 Bubble Dynamics Equations A3.8-31 A3.8-2 Extreme Values of Coefficients for the Symmetric Discharge Case A3.8-33 A3.8-3 Extreme Calculated Pressures for the Symmetric Discharge Case A3.8-34 A3.8-4 Extreme Values of Fourier Coefficients for the Asymmetric Discharge Case A3.8-35 A3.8-5 Extreme Calculated Pressures for the Asymmetric Discharge Case A3.8-36 A3.8-6 Chugging Loads A3.8-37 A3.8-7 Soil Strain Versus Modulus A3.8-38

CPS/USAR ATTACHMENT A3.8 - CONTAINMENT DESIGN LOADS LIST OF FIGURES CHAPTER 03 A3.8-iv REV. 11, JANUARY 2005 NUMBER TITLE A3.8-1 Array of Imaginary Sources and Sinks for Method of Images Model of Suppression Pool A3.8-2 Plan of Clinton Suppression Pool Showing the Vents Active in the Symmetric Loading Case A3.8-3 Plan of Clinton Suppression Pool Showing the Vents Active in the Asymmetric Loading Case A3.8-4 Cross Section of Suppression Pool A3.8-5 Symmetric Wall Loading - Zone 4 - Normalized Average Pressure A3.8-6 Symmetric Wall Loading - Zone 4 - Normalized 1st Cosine Harmonic A3.8-7 Symmetric Wall Loading - Zone 4 - Normalized 2nd Cosine Harmonic A3.8-8 Symmetric Wall Loading - Zone 4 - Normalized 3rd Cosine Harmonic A3.8-9 Symmetric Wall Loading - Zone 4 - Normalized 4th Cosine Harmonic A3.8-10 Asymmetric Discharge Wall Loading - Zone 4 - Normalized Average Pressure A3.8-11 Asymmetric Discharge Wall Loading - Zone 4 - Normalized 1st Cosine Harmonic A3.8-12 Asymmetric Discharge Wall Loading - Zone 4 - Normalized 2nd Cosine Harmonic A3.8-13 Asymmetric Discharge Wall Loading - Zone 4 - Normalized 3rd Cosine Harmonic A3.8-14 Asymmetric Discharge Wall Loading - Zone 4 - Normalized 4th Cosine Harmonic A3.8-15 Loss-of-Coolant Accident Chronology (DBA)

A3.8-16 Pressure Distribution on Suppression Pool Wetted Surface A3.8-17 Dynamic Loads Associated with Initial Bubble Formation in the Pool A3.8-18 Loads at HCU Floor Elevation Due to Pool Swell Froth Impact and Two-Phase Flow A3.8-18a NRC Acceptance Criteria For Froth Impact: Peak Amplitude of Pressure Pulse A3.8-19 Condensation Oscillation Forcing Function on the Drywell Wall O.D. Adjacent to the Top Vent A3.8-20 Condensation Oscillation Load Spatial Distribution on the Drywell Wall, Containment Wall, and Basemat A3.8-21 Peak Pressure Pulse Main in Top Vent During Chugging A3.8-22 Peak Force Pulse Main in Top Vent During Chugging A3.8-23 Average Force Pulse Train in Top Vent During Chugging A3.8-24 Average Pressure Pulse Train in Top Vent During Chugging A3.8-25 Typical Pressure Time History for Weir Annulus During Chugging A3.8-26 Underpressure Distribution on the Weir Wall and Drywell I.D. Wall During Chugging A3.8-27 Peak Pressure Pulse Train on the Weir Wall and Drywell I.D. Wall During Chugging CPS/USAR ATTACHMENT A3.8 - CONTAINMENT DESIGN LOADS LIST OF FIGURES CHAPTER 03 A3.8-v REV. 11, JANUARY 2005 A3.8-28 Mean Pressure Pulse Train on the Weir Wall and Drywell I.D. Wall During Chugging A3.8-29 Normalized Weir Annulus Pressure Pulse Attenuation A3.8-30 Typical Pressure Time-History on the Pool Boundary During Chugging A3.8-31 Suppression Pool Chugging Normalized Peak Underpressure Attenuation A3.8-32 Suppression Pool Chugging Normalized Spike Attenuation A3.8-33 Suppression Pool Chugging Spike Duration "d" as a Function of Location in the Pool A3.8-34 Suppression Pool Chugging Normalized Peak Post Chug Oscillations A3.8-35 Circumferential Underpressure Amplitude Attenuation A3.8-36 Circumferential Post Chug Oscillation Amplitude Attenuation A3.8-37 Suppression Pool Chugging Normalized Mean Underpressure and Post Chug Oscillation Attenuation A3.8-37a Suppression Pool Chugging Normalized Mean Underpressure Attenuation A3.8-38 Drywell Top Vent Cyclic Temperature Profile and Area of Application During Chugging A3.8-39 Drywell Top Vent Cyclic Temperature Profile During Chugging A3.8-40 Suppression Pool Temperature Profile for Large Breaks A3.8-41 SRV Quencher All Valve Vertical Response Spectra for Containment Wall, Elevation 712'-0" A3.8-42 SRV Quencher All Valve Vertical Response Spectra for Drywell Wall Elevation 712'-0" A3.8-43 SRV Quencher All Valve Vertical Response Spectra for Pedestal Elevation 724'-1-3/4" A3.8-44 LOCA Bubble Horizontal Response Spectra for Containment, Elevation 712'-0" A3.8-45 LOCA Bubble Vertical Response Spectra for Drywell Wall Elevation 712'-0' A3.8-46 LOCA Bubble Horizontal Vertical Response Spectra for RPV, Elevation 753'-

3 3/8" A3.8-47 DELETED A3.8-48 Finite Element Model (Used for Quencher and LOCA Analysis)

CPS/USAR CHAPTER 03 A3.8-1 REV. 11, JANUARY 2005 CONTAINMENT DESIGN LOADS A3.

8.1 INTRODUCTION

The methodologies used to determine the design-basis loads for structural components for the Clinton Power Station (CPS) are presented herein. The methodology for each loading phenomenon is discussed in detail and the critical load combinations and acceptance criteria are identified. Finally, the system analysis method is discussed. Section A3.8.2 discusses the methodology for determining safety-relief valve (SRV) actuation loads as they apply to the suppression pool boundaries and submerged structures.

Section A3.8.3 discusses the loads resulting from a loss-of-coolant accident (LOCA). Each LOCA-related phenomenon is identified and the loading methodology is presented. Section A3.8.4 identifies the remaining loads that act on the structure i.e., normal, seismic, thermal, etc. Section A3.8.5 identifies the load combinations and acceptance criteria, and Section A3.8.6 presents the analysis methods used to determine structural responses and the acceptance of the structural design. A3.8.2 DEVELOPMENT OF SRV LOADS The structural design basis for the Clinton Power Station was based on the assumption that rams head discharge devices would be installed on the SRV discharge lines. This assumption was consistent with the technical understanding of the phenomena and licensing at the time the construction permit was issued (see Paragraph 5, pp.6-9 of Reference 1.) Since that time, General Electric (GE) has determined that the quencher discharge device is a desirable alternative to the rams head device in that it substantially reduces the suppression pool boundary loads resulting from air-clearing phenomena during SRV discharge and minimizes thermal effects in the suppression pool. Further, GE has specified the quencher device for the Standard BWR/6-238 design and recommended the device for BWR/6-Mark III application (see Attachment A of Reference 2). The quencher device has been incorporated into the CPS design. Therefore, this section presents the methodologies used for determining the structural design-basis loads for both the devices. The former are presented to show compliance with the licensing commitments described in Reference 1, and the latter are presented to demonstrate the compatibility of the design with the quencher discharge device loads. The design-basis SRV loads on BOP piping and equipment are presented in Attachment A3.9.

A3.8.2.1 Description of the Phenomena Prior to the actuation of a pressure relief valve, the piping between the SRV discharge and the suppression pool water surface is full of air at drywell pressure and temperature. The discharge piping terminates at the quencher in the suppression pool, with the water level inside the pipe at the same level as the water at the pool surface. When a relief valve lifts, the effluent reactor steam causes a rapid pressure buildup in the discharge pipe. This results in a rapid compression of the column of air in the discharge pipe and acceleration of the water in the submerged portion of the piping and the expulsion of the water through the line. The pressure in the pipe builds to a peak as the last of the water is expelled. The compressed cushion of air between the water slug and the steam exits through CPS/USAR CHAPTER 03 A3.8-2 REV. 11, JANUARY 2005 the quencher, forming a number of clouds of small bubbles which begin to expand to the lower pool pressure. These bubbles continue to expand, displacing the pool water and propagating a pressure disturbance throughout the suppression pool. When the gas pressure reaches equilibrium with the local hydrostatic pressure, the transient would cease were it not for the inertia of the accelerated water mass. The inertia of the water drives the gas system past the point of equilibrium, and a negative pressure (with respect to local hydrostatic pressure) results within the bubble. The negative pressure in the bubble decelerates the water mass and reverses its motion in an attempt to reach equilibrium. Again, the inertia of the water drives the system past the equilibrium pressure, and the process repeats in a cyclic manner. The dynamics of the air-water system are manifest in pressure oscillations (similar to that of a springmass system) arising from the expansion and contraction of the bubble coupled with the inertial effects of the moving water mass. The oscillations are repeated with an identifiable frequency until the bubble reaches the pool surface. The magnitude of the pressure disturbance in the suppression pool decreases with increasing distance from the point of discharge, resulting in a damped oscillatory load at every point on the structures and pool boundaries below the pool surface. A3.8.2.2 Analytical Models for Rams Head Loads This section discusses the analytical models utilized to determine the bubble dynamics and wall loading data for discharge of an SRV through a rams head discharge device. The vent-clearing portion of the analysis is described in Reference 3. It is the goal of this section to provide the basic background necessary to understand the analytical procedures and assumptions utilized in the design-basis determination. These conclusions support the analytical models discussed in the following sections. A3.8.2.2.1 Bubble Dynamics Model The analytical model for the bubble dynamics is described in Subsections A3.8.3.1 and A3.8.3.4 and in Section 4.0 of Reference 3. The differential equations and initial conditions for bubble pressure, radius, and depth are summarized in Table A3.8-1. The analytical model includes the following features and assumptions: a. Bubble formation efficiency, , accounts for energy lost by heat transfer to the pool; consequently, this lost energy is not present in the bubble, as explained in Reference 3. A value of 0.1 was used. b. The environmental pressure (P) includes the hydrostatic pressure in the pool which varies as the bubble rises. The reference pressure (Pref) was assumed to be 14.7 psi above the pool surface. The oscillatory and vertical motions of the bubble are dependent upon the time-dependent description of the bubble depth, Z. Since this analysis includes the variation of bubble pressure with elevation, it is not advisable to assume that these effects are negligible as in Reference 3. c. The vertical bubble motion description includes the effects of the bubble's virtual mass, drag, buoyancy, and preferential upward expansion. A drag coefficient (C D) of 2.5 is used. The bubble radial displacement term (Rt) is included only when the bubble is expanding (R

> 0), representing preferential upward CPS/USAR CHAPTER 03 A3.8-3 REV. 11, JANUARY 2005 expansion of the bubble. The bubble is assumed to remain at the same initial depth until its formation is completed. Since both drag force and bubble mass are included in this analysis, it is not necessary to assume that these effects are negligible as in Reference 3. d. The initial bubble pressure (P) is calculated as a stagnation pressure based on vent line discharge pressure, density, and velocity. Reference 3 uses the static pressure rather than the stagnation pressure as the initial bubble pressure.

Since the stagnation pressure is larger than the static pressure, a higher (more conservative) initial bubble pressure is used. The discharge velocity is assumed to be sonic. The rams head geometry is taken into account in calculating discharge pressure and density by dividing by 2, assuming that two bubbles are formed. This results from the assumption that sonic velocity is established immediately upon clearing at both the end of the vent line and the rams head

outlets. e. The time required to discharge air (t d) is based on the initial mass of air in the vent line and mass flow rate (m) based on discharge density and velocity. At this time (t d), the first term in the pressure differential equation (Equation A3.8-8 in Table A3.8-1) is dropped and the bubble is allowed to rise. f. The bubble dynamics calculation is terminated when the bubble surface reaches the pool surface. A3.8.2.2.2 Method of Images Model The method of images is a classical technique applied in many scientific and engineering disciplines. Hydrodynamic images have been utilized to describe the flow of an ideal fluid for over a century (see References 4 and 5). Usually, however, the fluid has been assumed to be infinite or semi-infinite in extent (see References 6 and 7). In order to describe the pressure field behavior in a finite fluid (such as a suppression pool), a particular interpretation of the method of hydrodynamic images was developed (Reference 3). Good agreement between the application of the method of images and a series of experiments has been reported (References 8 and 9). An experiment was designed and performed in order that the application of the method of hydrodynamic images to the fluid in a suppression pool could be verified. Verification of the principle of superposition with respect to the pressure field resulting from discharging of multiple SRV lines was also substantiated. As many of the normal operating conditions as possible were simulated in this scaled experiment. Based on the results of these experimental investigations, the method of hydrodynamic images is judged to be well suited for the suppression pool analyses. Reference 3 presents the analytical models and equations that describe the bubble dynamics and the pool response. The development utilizes basic hydrodynamic potential flow theory (References 10 through 13), which assumes an ideal fluid. The bubble dynamics description is embodied in the Rayleigh equation, and the pool response description is derived from the generalized Bernoulli equation: )t(f t V 2 1 L P 2=++ (A3.8-1)

CPS/USAR CHAPTER 03 A3.8-4 REV. 11, JANUARY 2005 where: P = fluid pressure; L = fluid density; V = fluid velocity;

= force potential; = velocity potential; and f(t) = arbitrary function of time. The velocity potential, , satisfies the LaPlace equation:

0 2= (A3.8-2) and it must also satisfy the initial and boundary conditions. The result is substituted into the generalized Bernoulli equation (A3.8-1), and an expression for the pressure distribution in the pool as a function of space and time is obtained. The boundary conditions which are imposed on the potential function are as follows: a. There is no flow across the rigid boundaries. b. The pressure at the free surface is constant and unaffected by the bubble.

c. At the bubble surface, the pressure is given as a function of time by the bubble dynamics equation. The first two boundary conditions are satisfied through the use of the method of images. If image sources and sinks are placed as depicted in Figure A3.8-1, the effects of the rigid boundaries and the free surface can be described. The rigid boundaries are simulated by flow sources, and the free surface is simulated by sinks. The sources and sinks all have the same

uniform strength. The potential function must be harmonic to satisfy the LaPlace equation. For a simple point source, the following form is appropriate (Reference 12).

r M= (A3.8-3) where: M = source or sink strength, and r = radial distance from the point source. Since the sum of harmonic functions is also harmonic, the effect of all the simple point sources and sinks in the image array can be included through a simple summation. Thus, the local pressure (potential) in the suppression pool can be described by:

CPS/USAR CHAPTER 03 A3.8-5 REV. 11, JANUARY 2005 ==+/-++/-+=i i n1i bi b n0i wi 2 bbb r 1)R1(r 1gc2 RL)PP(RPP (A3.8-4) where: L = liquid density; P = local pressure; P = environmental pressure; P b = bubble pressure; R b = bubble radius; b R = time rate of change of bubble radius; r bi = vector distance from the point of interest to the center of i th source or sink; r wi = distance from the point on the boundary to the i th image; i=0 for the bubble; n = total number of real and imaginary sources and sinks; and

+/-1 = (+) depicts a source; (-) depicts a sink. It should be noted that equation A3.8-4 is more general than the equivalent expression (Equation 35) given in Reference 3. Equation A3.8-4 applies during the entire bubble oscillation and not just at the extremes where R = 0. Hence, Equation A3.8-4 represents the total bubble pressure (i.e., the static and dynamic contributions are both included). The third boundary condition is that the pressure at the bubble surface is equal to the internal pressure of the bubble. This procedure accounts for the bubble dynamics, bubble rise, effects of the rigid and free boundaries, and pressure at the interface between the bubble and the suppression pool fluid. The method was applied to the Mark III containment cylindrical geometry in all three dimensions, hence a general solution is embodied. The effects of the convex curvature of the drywell wall and concave curvature of the containment wall (relative to suppression pool fluid) are included in this general solution. Thus, the parametrics include the effects of pool geometry, bubble characteristics, and bubble location. A3.8.2.2.3 Vent Clearing Model The analytical model for the vent clearing transient model is fully described in Reference 3. This model has been extended to include the effect of friction on the water slug in the discharge line and implemented in the design calculations.

CPS/USAR CHAPTER 03 A3.8-6 REV. 11, JANUARY 2005 A3.8.2.2.4 Assumptions Used in the Implementation of the Analytical Model In order to implement the analytical models described above, certain assumptions were required. These assumptions relate to assuring that conservative results are produced for use in design calculations. These assumptions are: a. The actual plant geometry and piping configurations are used in the analysis. b. The spring setpoints, Section III of the ASME Code, Article NB-7000, and additional conservative factors, to accommodate manufacturing tolerances, were used to determine the maximum individual SRV flow rates for use in the analysis. c. The method of images was employed to account for the finite pool boundaries and size. d. The loadings are determined by combining the effects of each bubble pair (two bubbles per rams head) for a given SRV discharge line using the SRSS method.

Then the effects of each bubble pair were linearly summed. e. The bubbles were located on the axis of the discharge device 4 feet from the exit plane of the device. A3.8.2.2.5 Method of Implementation In order that the loads are properly and accurately defined at the boundaries of the suppression pool as concisely as possible, the analytical models are implemented as follows: a. The suppression pool vertical and horizontal boundaries are divided into three zones of equal length and the loading function is evaluated as a pressure at the midpoint of each zone for each boundary. (See Figure A3.8-4). b. The pressure is determined as described above at azimuthal increments of 2

° around the pool boundary. c. The loading functions, i.e., the pressure distribution on each zone, are then represented as a Fourier Series. Thus, from Reference 14: ++==nsinBncosA(2 Ao)(P n n1n) where: ,...2,1,0ndncos)(P 1 A n== and ,...2,1,0ndnsin)(P 1 B n==

CPS/USAR CHAPTER 03 A3.8-7 REV. 11, JANUARY 2005 Here, P() = the calculated pressure distribution on each zone, A n , B n = the Fourier cosine and sine coefficients as defined above, and = azimuth with the origin at the centerline of the RPV. Each coefficient is then normalized by the maximum absolute value of each function over all nine zones. Therefore: ,...2,1,0ndncos)(P A 1 a n n== and ...3,2,1,0ndnsin)(P B 1 b n n== and )nsinbncosa(2 a)(p n n1n o++==. where: p() is the normalized Fourier representation of the pressure function acting on each zone at any instant, A n* and B n* are the maximum absolute values of the Fourier coefficients for all zones for all time, and a n and b n are as defined above. This process is repeated at each time step (typically 0.002 second) throughout the transient.

A3.8.2.2.6 Selection of Discharge Cases The discharge cases selected for design purposes are those which give the maximum vertical and horizontal response for the containment and also the maximum overturning moment. These criteria are met by two discharge cases, the symmetric discharge case and the asymmetric case. These are discussed in the following. A3.8.2.2.6.1 Symmetric Discharge Case The symmetric discharge case is defined as the simultaneous actuation of all 16 SRV's.

A unique conservative bubble pressure time history is produced for each discharge line considering the valve setpoint, discharge line air mass, clearing time, discharge pressure, etc. These unique bubble dynamics are then considered for valves actuating simultaneously in the methodology described above. Figure A3.8-2 shows the distribution of the discharge devices in the suppression pool. Figure A3.8-4 shows a cross section of the pool which indicates the zones over which the boundary pressure time histories are produced.

CPS/USAR CHAPTER 03 A3.8-8 REV. 11, JANUARY 2005 Table A3.8-2 indicates the maximum and minimum values for the zonal Fourier coefficients (A n , B n) for the first four harmonics. Table A3.8-3 summarizes the minimum and maximum point (local) pressures calculated for each zone. Figures A3.8-5 through A3.8-9 show representative values of the normalized Fourier cosine coefficients (a n) as they vary in time throughout the transient. A3.8.2.2.6.2 Asymmetric Discharge Case The asymmetric discharge case is defined as the subsequent actuation of the low setpoint valve together with the simultaneous failure (open) of the valve connected to an adjacent discharge line. The bubble dynamics are produced as discussed above, with the exception that the discharge line undergoing the subsequent discharge is considered to have an elevated water level due to the assumption that the previous transient has not died out. The loading functions are otherwise prepared as discussed above. Figure A3.8-3 shows the distribution of the discharge devices in the suppression pool, and the zonal definitions are shown in Figure A3.8-4 as before. Table A3.8-4 contains the zonal extrema for the Fourier coefficients (A n , B n), and Table A3.8-5 shows the zonal extrema for the calculated point (local) pressures.

Figures A3.8-10 through A3.8-14 show representative values of the normalized Fourier cosine coefficients (a n), indicating their variance with time throughout the analysis. A3.8.2.3 Quencher Design Loads on Suppression Pool Boundaries The methodology and procedure for determining design boundary loads with a quencher discharge device are discussed in Section 2.2 of Attachment A3.9. A3.8.3 DEVELOPMENT OF LOCA LOADS During a LOCA, the structures forming the containment system and other structures within the containment experience dynamic phenomena. This section provides numerical information on the dynamic loads that these phenomena impose on the CPS containment system structures. A3.8.3.1 Description of LOCA (DBA)

Figure A3.8-15 shows the events occurring during DBA and the potential loading conditions associated with these events (Reference 18). If the hypothetical guillotine break is postulated to occur within the annulus between the RPV and the biological shield wall the shield annulus will experience a rapid pressurization. During this short-lived transient the annulus is pressurized before the break flow can eventually find its way into the drywell through the various penetrations and openings in the shield wall and the opening at the top of the shield wall. As the drywell pressure increases, the water initially standing in the vent system discharges into the pool, clearing the vents. During this vent-clearing process, the water leaving the horizontal CPS/USAR CHAPTER 03 A3.8-9 REV. 11, JANUARY 2005 vents forms jets in the suppression pool and causes water jet impingement loads on the structures within the suppression pool and on the containment wall opposite the vents. During the vent-clearing transient, the drywell is subjected to a pressure differential, and the weir wall experiences a vent-clearing reaction force. Immediately following vent clearing, an air and steam bubble forms at the exit of the vents. The bubble pressure initially is conservatively assumed equal to the current drywell pressure (drywell pressure at vent clearing is calculated as 17.56 psig, however, the conservative value of 20.1 psig is the design basis peak pressure), which transmits a pressure wave through the suppression pool and results in loading on the suppression pool boundaries and on equipment located in the suppression pool. As air and steam flow from the drywell becomes established in the vent system, the initial vent exit bubble expands to suppression pool hydrostatic pressure. Tests at the GE Large Scale Pressure Suppression Test Facility (PSTF) (Reference 17) show that the steam fraction of the flow is condensed, but that continued injection of drywell air and expansion of the air bubble results in a rise in the suppression pool surface.

During the early stages of this process, the pool swells in a bulk mode (i.e., a slug of solid water is accelerated upward by the air). During this phase of pool swell, structures close to the pool surface will experience impact loads as the rising pool surface reaches the lower surface of the structure. In addition to these initial impact loads, these same structures will experience drag loads as water flows past them. Equipment in the suppression pool located above the elevation of the bottom vent will also experience drag

loads. Data from PSTF air tests indicates that after the pool surface has risen approximately 1.6 times the initial submergence of the top vent, the water ligament thickness has decreased to 2 feet or less when the ligament begins to break up, and the impact loads are significantly reduced. This phase is referred to as incipient breakthrough. Ligament thickness continues to decrease until complete breakthrough is reached and the air bubble can vent to the containment free space. The breakthrough process results in formation

of an air/water froth. Continued injection of drywell air into the suppression pool results in froth pool swell. This froth swell impinges on structures it encounters, but the two-phase nature of the fluid results in loads that are much smaller than the impact loads associated with bulk pool swell. When the froth reaches the bottom elevation of the open portion of the floors on which the hydraulic control units for the control rod drives are located, approximately 22 feet above pool level, the froth encounters a flow restriction; at this elevation, there is approximately 35% open area. The froth pool swell experiences a two-phase pressure drop as it flows through the available open areas. This pressure differential represents a load on both the floor structures and the adjacent containment and drywell. The result is a discontinuous pressure loading at this elevation. The pool swell impact and impingement target data presented in Attachment A3.9 are applicable to small structures (less than 20 inches). This restriction on the application of the impact test data is necessary, since the basic tests involved targets with a width of 20 inches. For this size target, only the suppression pool water in the immediate vicinity of the target has to be redirected by the impact impulse, thus the impact loads are not dependent upon the pool CPS/USAR CHAPTER 03 A3.8-10 REV. 11, JANUARY 2005 swell water ligament thickness. Attachment A3.9 discusses application of PSTF impact data to small structures. For floors that are expansive enough to decelerate a large sector of the pool rather than a small region of the pool in the vicinity of the target, the impulsive loading on the floor is dependent upon the momentum of the entire slug and is related to slug thickness. As drywell air flow through the horizontal vent system decreases and the air/water suppression pool mixture experiences gravity-induced phase separation, pool upward movement stops, and the "fallback" process starts. During this process, floors and other flat structures experience downward loading. No pressure increase on the containment wall has been observed experimentally. The pool swell transient associated with drywell air venting to the pool typically lasts 3 to 5 seconds. Following this, there is a long period of high steam flow rate through the vent system; data indicates that this steam will be entirely condensed in a region adjacent to the vent exits. For the DBA reactor blowdown, steam condensation lasts for a period of approximately 1 minute. Structural loadings during the steam condensation phase of the accident are relatively small and are included in the containment loading specification. As the reactor blowdown proceeds, the primary system is depleted of high-energy fluid inventory, and the steam flow rate to the vent system decreases. This reduced steam flow rate leads to a reduction in the drywell/containment pressure differential, which in turn results in a sequential re-covering of the horizontal vents. Re-covering of a particular vent row occurs when the vent stagnation differential pressure corresponds to the suppression pool hydrostatic pressure at the row of vents. Toward the end of the reactor blowdown, the top row of vents is capable of condensing the reduced blowdown flow, and the two lower rows will be totally re-covered. As the blowdown steam flow decreases to very low values, the water in the top row of vents starts to oscillate back and froth, causing what has become known as vent "chugging." This action results in dynamic loads on the top vents and on the weir wall opposite the upper row of vents and on the drywell and containment. Since this phenomenon is steam mass flux dependent (the chugging threshold appears to be in the range of 10 lb/sec/ft 2), it is present for all break sizes. For smaller breaks, it is the only mode of condensation that the vent system will experience. Shortly after a DBA, the emergency core cooling system (ECCS) pumps automatically start up and pump condensate and/or suppression pool water into the reactor pressure vessel. This water cascades into the drywell from the break. The time at which this occurs depends upon break size and location. Because the drywell is full of steam at the time of vessel flooding, the sudden introduction of cool water causes steam condensation and drywell depressurization. When the drywell pressure falls below the contai nment pressure, the drywell vacuum relief valve open and air from the containment enters the drywell to equalize the drywell and containment pressure. During this drywell depressurization transient, there is a period of negative pressure on the drywell structure, and a conservative negative load condition is therefore specified for the drywell design. While the drywell pressure is below the containment pressure the containment atmosphere flows from the wetwell through the vacuum breakers to the drywell and the suppression pool water flows through the LOCA vents into the drywell. The suppression pool water spills over the weir wall and induces drag loads on piping, equipment, and structures below the top of the weir wall. Impact and drag velocities are computed by methods similar to those described in Reference 21. Impact and drag loads resulting from the upward movement CPS/USAR CHAPTER 03 A3.8-11 REV. 12, JANUARY 2007 of the water in the weir annulus are computed by the methods described in Appendix L of Reference 21. The downward drag loads due to weir swell are computed by the same method as the pool swell fall back loads with the velocity being computed as the free fall velocity from the maximum calculated weir swell height (740'-9"). Suppression pool water is continuously recirculated through the core by the ECCS pumps, resulting in a slow heatup of the suppression pool. To control suppression pool temperature, the RHR system is put into service. After several hours, the heat exchangers control and limit the suppression pool temperature increase. It is conservatively calculated that the pool will reach a peak temperature of 184.5

°F. The increase in air and water vapor pressure at these temperatures results in a pressure loading of the containment, as discussed in Chapter 6. The post-DBA containment heatup and pressurizat ion transient is terminated when the RHR heat exchangers reduce the pool temperature and containment pressure to nominal values.

A3.8.3.2 Description of LOCA (IBA)

An intermediate size break is defined as a break that is less than the DBA but is of sufficient magnitude to automatically depressurize the primary system due to loss of fluid and/or automatic initiation of the ECCS systems (Reference 18). In practice, this means liquid breaks greater than 0.05 ft 2 and steam breaks greater than 0.4 ft 2 as determined by analysis. In general, the magnitude of dynamic loading conditions associated with a loss-of-coolant accident decreases with decreasing break size. However, the intermediate break is examined because the automatic depressurization system (ADS) is involved and simultaneous actuation of the multiple SRV's committed to this system introduces containment system loads, as discussed in Section A3.8.6. A3.8.3.3 Description of LOCA (SBA)

Small breaks are defined as breaks not large enough to depressurize the reactor automatically (Reference 18). The dynamic loads produced by this class of accident are small. However, there are certain conditions associated with smaller reactor system breaks that must be considered during the design process. Specifically, the drywell and weir wall must be designed for the thermal loading conditions that can be generated by a small steam break (SBA). For a definition of the design conditions, the following sequence of events is postulated. With the reactor and containment operating at maximum normal conditions, a small break is postulated to occur, allowing blowdown of reactor steam to the drywell. The resulting drywell pressure increase leads to a high drywell pressure signal that scrams the reactor and activates the containment isolation system. Drywell pressure continues to increase at a rate dependent on the size of the assumed steam leak. A pressure increase to approximately 3 psig depresses the water level in the weir annulus until the level reaches the top of the upper row of vents. At this time, air and steam enter the suppression pool. Steam is condensed, and the air passes to the containment free space. The latter results in gradual pressurization of the containment at a rate dependent upon the air carryover. Entrainment of the drywell air in the steam flow through the vents results in all the drywell air being carried over to the containment. At this time, containment pressurization ceases. The drywell is now full of steam and has a positive pressure differential sufficient to keep the weir annulus water level depressed to the top vents and chugging can occur. Reactor blowdown steam continues to be condensed in the suppression pool.

CPS/USAR CHAPTER 03 A3.8-12 REV. 11, JANUARY 2005 The thermodynamic process associated with blowdown of primary system fluid is one of constant enthalpy. If the primary system break is below the normal RPV water level, blowdown flow consists of reactor water. Upon depressurizing from reactor pressure to drywell pressure, approximately one-third of this water flashes to steam, two-thirds remains as liquid, and both phases will be in a saturated condition at drywell pressure. Thus, if the drywell is at atmospheric pressure, the steam-and-liquid blowdown will have a temperature of 212

° F. If the primary system rupture is located so that the blowdown flow is reactor steam, the resultant temperature in the drywell is significantly higher than the saturated temperature associated with liquid blowdown. This is because a constant enthalpy decompression of high-pressure saturated steam results in a superheat condition. For example, decompression of 1,000-psia saturated steam to atmospheric pressure results in 298

° F superheated steam (86

° F of superheat). Reactor operators are alerted to the SBA incident by the leak detection system, or high drywell-pressure signal, and reactor scram. For the purpose of evaluating the duration of the superheat condition in the drywell, it is assumed that operator response to the small break is to

depressurize the reactor using selected relief valves and to activate the RHR system to control the suppression pool temperature. (This conserva tively assumes that the main condenser is not available and the operators must use the suppression pool for an energy sink.) Reactor cooldown is assumed to be started 30 minutes after the break and to proceed at a rate of 100

° F/hr. This results in a reactor cooldown to 212

° F or less in approximately 3 to 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />, after which time the blowdown flow rate is terminated. It should be noted that the end-of-blowdown chugging phenomenon discussed in Subsection A3.8.3.4 also will occur during a small-break accident and will last the duration of reactor depressurization. A3.8.3.4 Design Loads for LOCA Events The design pressures and temperatures for the drywell which result from the various LOCA phenomena and assumptions are discussed in detail in Subsection 6.2.1.1.3. These analyses are not repeated here, but the consequences of the transients are incorporated in the design as discussed in Subsection A3.8.5, Attachments B3.8 and A3.9, and Sections 3.9 and 3.10. Also, the effects of a pipe break within the shield annulus are discussed in detail in Subsection A3.8.6.2 and accounted for in the design. The effects of pool swell impact and drag loads on piping and equipment located in the containment are quantified and incorporated into the design as discussed in Attachment A3.9. The discussions that follow quantify the effects of the LOCA transient that act directly on the containment structure. The incorporation of these effects into the structural design is discussed in Sections A3.8.5 and A3.8.6. A3.8.3.4.1 Water Jet Loads During the vent clearing transient, the weir wall, the weir side of the drywell wall and the LOCA vents will experience drag loads due to the suppression pool water being forced through the weir annulus and the LOCA vents by the rising drywell atmospheric pressure. These loads are negligible when compared with other loads (i.e., condensation oscillation and chugging) that act on these surfaces later in the transient.

CPS/USAR CHAPTER 03 A3.8-13 REV. 11, JANUARY 2005 The water jets emitted from the LOCA vents during the vent-clearing transient impinge on the containment wall. However, the loading imposed on the containment wall is insignificant when compared with the load caused by the air bubble loads that apply later in the transient, as discussed in Reference 18. A3.8.3.4.2 LOCA Air Bubble Loads A3.8.3.4.2.1 LOCA Air Bubble Loads on the Weir Annulus and Vent System Once the flow of air, steam and water droplets has been established in the vent system, there will be a static pressure reduction in the weir annulus and vent system that leads to a uniform outward pressure of approximately 10 psid on the weir wall. This loading was calculated with the flow model described in Reference 15, and for design purposes is assumed to exist for the first 30 seconds of the blowdown (Reference 18). A3.8.3.4.2.2 LOCA Air Bubble Loads on the Exterior of the Drywell Wall During the vent clearing process, the drywell pressure reaches a peak pressure of 20.1 psig. During the subsequent vent flow phase of the blowdown, the peak pressure differential does not exceed the 20.1-psid value even when it is assumed that pool swell results in some two-phase flow reaching the containment annulus restriction at the HCU floor. Interaction between the pool swell water mass and the limited number of structures at or near the pool surface does not significantly affect the drywell pressure. During bubble formation, the outside of the drywell wall will be subject to varying pressures. A bounding range of 0 to 20.1 psid is specified on those sections of the drywell wall below the suppression pool surface. The basis for this specification is the knowledge that the minimum pressure increase is 0 psi, and that the maximum bubble pressure can never exceed the peak drywell pressure of 20.1 psid, as discussed in Reference 18. As the water rises to a maximum of 18 feet above the nominal pool surface, the pressure decreases linerally from 20.1 psid to 0 psid, as shown in Figures A3.8-16 and A3.8-17. A3.8.3.4.2.3 LOCA Air Bubble Loads on the Containment The magnitude of the pressure increase on the containment wall following vent clearing is dependent upon the rate at which the drywell air bubble accelerates the suppression pool water. Circumferential variations in the air flow rate may occur due to drywell air/steam mixture variations, but these variations result in negligible variations in the containment bubble pressure

load. The large-scale PSTF test data are the basis for specifying the maximum asymmetic load of 10 psid. A maximum increase of 10 psid on the containment wall was observed in the PSTF at the Mark III drywell peak calculated pressure of 36.5 psia. Thus, use of a 10-psid asymmetric pressure condition applied in a worst-case distribution is a bounding specification for the symmetric load on the containment wall, as discussed in Reference 18. The symmetric loading specification is based on the same test results and is specified as a symmetric pressure load of 10 psid applied over the containment surface below the suppression pool surface (Reference 18). Above the suppression pool surface, the pressure is attenuated linearly to 0 psid over 18 feet, as shown in Figure A3.8-16. The time history for both the symmetric and asymmetric loading function if given in Figure A3.8-17.

CPS/USAR CHAPTER 03 A3.8-14 REV. 11, JANUARY 2005 A3.8.3.4.2.4 LOCA Air Bubble Loads on the Basemat The outer half of the basemat will experience a 10-psid bulk pressure load associated with the air bubble formation discussed above. The inner half of the basemat will experience a pressure rise that is assumed to increase from 10 psid at the center of the pool to 20.1 psid at the drywell wall. The increase in pressure over the inner 50% of the pool width is assumed to be linear, as discussed in Reference 18 and as shown in Figure A3.8-16. The time history for the transient is shown in Figure A3.8-17. A3.8.3.4.3 Pool Swell Drag and Impact Loads The loads imposed on the structure and equipment within the pool swell zone are discussed in Attachment A3.9. For pool swell loading, Clinton piping systems above the suppression pool are typically restrained in the vertical direction utilizing pinned-pinned rigid supports. The concern is related to lateral instability of piping systems under steady-state pool swell/froth drag loading. A review was made of all affected piping systems with significant runs of horizontal piping. This review has shown that the affected piping is rigidly restrained in the lateral direction of seismic/pool dynamic response spectrum loading. The lateral rigid restraints are adequately designed and spaced to eliminate potential instability problems. The NRC concern has therefore been considered in the Clinton design. (Q&R 210.03) A3.8.3.4.4 Fallback Loads The drag loads due to pool swell fallback on the drywell, containment and basemat are negligible compared with the initial pool swell drag and bubble loads. Fallback drag loads on piping and equipment are specified in Attachment A3.9. A3.8.3.4.5 Froth Impingement Loads on the HCU Floor The HCU floor is approximately 23.5 feet above the suppression pool and is approximately 11 feet above the elevation at which breakthrough is expected to occur. Froth will reach the HCU floor approximately 0.5 second after the top vent clears and will generate both impingement loads on the structure and a load due to flow pressure differential as it passes through the restricted annulus area at this elevation.

This will result in vertical loads on the drywell wall, an upward loading on the HCU floor and an outward pressure loading on the containment wall. The time history for the froth impingement impact and drag loads is shown in Figure A3.8-18. The froth impingement load definition can be

determined from Figure A3.8-18a. The HCU floor has been assessed for this load.

Following the completion of the design of the structures and components at the HCU floor elevation, Reference 22 was received which gave the NRC acceptance criteria for the froth impact load. The design of the Clinton Power Station is judged to be adequate for these new loads. A3.8.3.4.6 Condensation Oscillation Loads Following the initial LOCA pool swell transient, during which the drywell air is vented to the containment free space, there will exist a high steam mass flow through the top vents, and CPS/USAR CHAPTER 03 A3.8-15 REV. 11, JANUARY 2005 condensation oscillation will occur. Vent mass fluxes up to 25 lb/sec-ft 2 will occur as the result of a main steam or recirculation line break. For a more detailed discussion and for a discussion of the test data on which the following methodology is based, see Reference 18. The condensation oscillation forcing function used for design is defined as a summation of four harmonically related sine waves developed from a regression analysis of the data obtained in test series 5807 of Reference 16. The results are:

2)t(A)(P= { 0.80 sin [2 f (t) ] (A3.8-5) + 0.30 sin [4 f (t) ] + 0.15 sin [6 f (t) ] + 0.20 sin [8 f (t) ] }

where: P() = pressure amplitude for a cycle beginning at time t and ending at t+Tp (psid), A(t) = peak to peak amplitude variation with time (psid) = 5.5[3.395 - 0.106t + 1.15 log t - 7.987(log t) 2 + 7.688 (log t) 3 - 1.344 (log t) 4 , (A3.8-6) f(t) = fundamental frequency variation with time (sec-:)

= 0.8[2.495-0.225t - 0.742 log t (A3.8-7) + 10.514(log t) 2 - 9.271(log t) 3 + 3.208(log t) 4], t = time (sec) from initiation of LOCA, 3t30, = time (sec) from the beginning of each cycle, 0 Tp, and Tp = f(t)-1 (sec). P() from Equation A3.8-5 has been calculated for four cycles and is shown in Figure A3.8-19. The spatial distribution of the forcing function amplitude over the wetted surface of the suppression pool is shown in Figure A3.8-20. The amplitudes shown in Figure A3.8-20 are normalized to 1.0 at the centerline of the top vent. A3.8.3.4.7 Chugging Loads In addition to the bulk drywell pressure fluctuations, pressure pulses have been observed when steam bubbles collapse in the vents, in a process known as chugging. The dominant pressure response to the top vent during chugging is that of a pulse train, with the peak amplitude of the CPS/USAR CHAPTER 03 A3.8-16 REV. 11, JANUARY 2005 pulses varying randomly from chug to chug. The pulse train associated with a chug consists of a sequence of four pulses with exponentially decreasing amplitude. The dominant pressure response to chugging in the suppression pool is characterized by a pre-chug underpressure, chug (pressure spike), and a post-chug oscillation. The chugging phenomenon as observed in PSTF tests has a random amplitude and frequency. Although it is expected that chugging will occur randomly among the vents, synchronous chugging in all top vents is conservatively assumed. Each vent is expected to be periodically subjected to the peak observed pressure spike. The pool boundary load definitions consider that the chugging loads transmitted to the drywell wall, base mat, weir wall, and containment are the result of several vents chugging simultaneously at different amplitudes. A detailed discussion of the data on which the following load definitions are based is presented in Reference 18. A3.8.3.4.7.1 Chugging Loads Applied to Top Vent Within the top vent, the peak pressure pulse train shown in Figure A3.8-21 is applied for local or independent evaluation of vents. Although some variation is observed in the pressure distribution from the top to the bottom of the vent, it is conservatively assumed that during the chugging event the entire top vent wall is simultaneously exposed to spatially uniform pressure pulses. Because some net imbalance in the pressure distribution gives rise to a vertical load, the peak force pulse train shown in Figure A3.8-22 is applied vertically upward over the projected vent area concurrently with the peak pressure pulse train shown in Figure A3.8-21 to evaluate local effects on the vents. For global effect, the average force pulse train shown in Figure A3.8-23 is applied vertically over the projected area of all top vents concurrently with the average pressure pulse train within the vent shown in Figure A3.8-24. The underpressure preceding the pressure pulse train within the top vent is very small compared with the peak (spike) over-pressure. The mean measured pressure in the vent (results from tests) was -9 psid with a standard deviation of

+/- 3 psid. On this basis, the specified design value is -15 psid. A bounding underpressure of -19.5 psid was calculated for

inside vent surface. The pressure pulses generated inside the top vents during chugging propagate toward the weir annulus. The dominant pressure response in the weir annulus during chugging is characterized by a pre-chug underpressure followed by a pressure pulse train as shown in Figure A3.8-25.

The load applied to the weir annulus (weir wall and drywell wall) is described by a prechug underpressure, defined as a half sine wave as shown in Figure A3.8-26, followed by a pressure pulse train as shown in Figures A3.8-27 and A3.8-28. For local load considerations, the peak amplitudes are applied. For global considerations, the mean amplitudes are applied. Vertical attenuation of the weir underpressure is small; for design evaluation, no attenuation is assumed. For the pressure pulse train, the attenuation on the weir wall and the drywell ID wall in the vertical direction is shown in Figure A3.8-29. For distribution in the circumferential direction, the local loading for the pre-chug underpressure attenuates as shown in Figure A3.8-29. For the global loads, there is no attenuation in the circumferential direction.

CPS/USAR CHAPTER 03 A3.8-17 REV. 11, JANUARY 2005 A3.8.3.4.7.2 Pool Boundary Chugging Loads The chugging load applied to the pool boundary (drywell, base mat and containment) is described by the typical forcing function shown in Figure A3.8-30. The forcing function consists of a prechug underpressure defined as a half sine wave, a triangular pulse (pressure spike) loading characterized by a time duration "d" and a post-chug oscillation described by a damped sinusoid. The impulse is at its maximum magnitude and duration near the top vent on the drywell wall due to the localized nature of the phenomena. The amplitude of the pre-chug underpressure and the post-chug oscillation are also maximum at this location.

For local load considerations on the pool boundary: Pre-chug underpressure: peak amplitude - Table A3.8-6; and distribution - Figure A3.8-31. Pulse (spike): peak amplitude - Table A3.8-6; distribution - Figure A3.8-32; and duration - Figure A3.8-33. Post-chug oscillation: peak amplitude - Table A3.8-6; and distribution - Figure A3.8-34: For distribution in the horizontal (circumferential) direction, the pre-chug underpressure attenuates on the drywell, base mat and containment, as shown in Figure A3.8-35. The pulse attenuation is the same as the lower portion of the vertical attenuation shown in Figure A3.8-32, except that the peak is at the vent centerline, and the post-chug oscillation attenuates on the drywell, base mat and containment, as shown in Figure A3.8-36. The profiles in Figures A3.8-35 and A3.8-36 represent the peak observed value at one vent, with the other vents chugging at the mean value.

For global load considerations on the pool boundary: Pre-chug underpressure: mean amplitude - Table A3.8-6; and distribution - Figure A3.8-37a. Pulse (spike): mean amplitude - Table A3.8-6; distribution - Figure A3.8-32; and duration - Figure A3.8-33.

CPS/USAR CHAPTER 03 A3.8-18 REV. 11, JANUARY 2005 Post-chug oscillation: mean amplitude - Table A3.8-6; and distribution - Figure A3.8-37. There is no horizontal attenuation for this loading. A3.8.3.4.7.3 Cyclic Temperature Due to Chugging The chugging phenomenon includes the continual evacuation and reflooding of the vents and periodic collapse of steam bubbles on the drywell wall. The temperature transient and area of application are shown in Figures A3.8-38 and A3.8-39. A3.8.3.4.7.4 Suppression Pool Thermal Stratification During the period of steam condensation in the suppression pool, the pool water in the immediate vicinity of the vents is heated. For the Mark III configuration, most of the condensing steam mass and energy are released to the pool through the top vents. The hot water rises by natural convection, and the cold water is displaced toward the bottom of the pool. The vertical temperature gradient resulting from this effect is known as thermal stratification. The short-term thermal stratification for the large break accident used in the containment evaluation is shown in

Figure A3.8-40. A3.8.4 DEFINITION OF OTHER LOADS Loads other than the SRV and LOCA loads used in the structural design of the containment structure are discussed in Chapter 3 of the F SAR. Variation of suppression pool temperature with depth is accounted for in the design of the containment. A3.8.5 LOAD COMBINATIONS Load combinations used in the structural design of the containment structure are discussed in Section 3.8. A3.8.6 ANALYSIS METHODS A3.8.6.1 General Dynamic structural analysis for the SRV and LOCA loads is discussed in Subsections A3.8.6.1.1 and A3.8.6.2. A brief description of each of the loads precedes the discussion of the

analysis. Structural responses to pool dynamic loads are most pronounced in the region where the loads occur, namely, the suppression pool. The suppression pool region is bounded by axisymmetric structures such as the containment wall, drywell wall, and the basemat. Therefore, it is appropriate and efficient to use axisymmetric thin shell of revolution structural responses can be accurately determined. The presence of adjacent structures is accounted for by including in the analysis model the adjacent structures that will have significant response to pool dynamic loads. These structures are represented by axisymmetric plates and shells having masses and stiffnesses equivalent to CPS/USAR CHAPTER 03 A3.8-19 REV. 11, JANUARY 2005 their actual configuration. The formulation of these axisymmetric finite elements does not deviate from the general finite element approach. Unlike the seismic phenomenon where the loading represents ground motion, the pool dynamic loads represent conditions of forced vibration which is relatively localized and the structural responses attenuate away with distance from the source of the loading. In addition, the free boundary conditions at the outer edge of the modeled base slab conservatively ignore any stiffening effect of the base slab extension to the other unit. Therefore, it is not necessary to model the two unit complex. (Q&R 220.43) Typical response of the structure is shown in Figures A3.8-41 through A3.8-46.

A3.8.6.1.1 Analysis of SRV Loading SRV loads for both the rams head and quencher discharge devices have been considered in the CPS structural analysis. The preliminary design of the containment was based on the conservative rams head load. It was later decided that the quencher device would be installed at CPS, since it imposed a reduced loading. The containment structure originally designed for the rams head loads also is assessed for the quencher loads. The analysis for the rams head load is presented in Section A3.8.6.1.1.1. The analysis for the quencher load is described in Section A3.8.6.1.1.2. A3.8.6.1.1.1 Rams Head Discharge A3.8.6.1.1.1.1 Loading The rams head SRV loads are described in detail in Section A3.8.2. Both a symmetric and an asymmetric load case are considered. The peak positive pressure for the symmetric rams head load is 83.5 psid and the peak negative pressure is 21.6 psid. The peak positive and negative pressures for the asymmetric load case are 70.0 psid and 11.5 psid, respectively. The frequency of the loads varies approximately from 5 hertz to 13 hertz, and the load has a time duration of approximately 0.75 seconds. A3.8.6.1.1.1.2 Model for Analysis The plant structure and underlying soil, were analyzed for the rams head discharge using an axisymmetric finite element model. The soil-structure model is similar to the model used for the quencher analysis shown in Figure A3.8-48. Appropriate material properties are used to describe the various structural and soil model components. The structural components were represented by thinshell finite elements. Solid finite elements were used to model the soil from bedrock up to the basemat. To consider soil effects beyond the finite element model boundary, an artificial viscous boundary (nonreflecting boundary) is used for horizontal motion. The viscous boundary provides the capability to model an infinitely long system utilizing a model of finite size. The vertical motion at the side boundary of the modeled soil media is assumed to be unrestrained. At the bottom interface between silt and bedrock, fixed boundary conditions are used. More information concerning the viscous boundary can be found in Appendix C - DYNAX. (Q&R 220.42)

CPS/USAR CHAPTER 03 A3.8-20 REV. 11, JANUARY 2005 Soil properties are strain dependent and are given in Table 2.5-48. In order to obtain appropriate elastic modulus corresponding to the strain, an iterative procedure utilizing an axisymmetric DYNAX model is carried out. Iteration is continued until the assumed strain levels and the actual strain level converge. The final values of strains and the corresponding soil moduli used in the analysis are shown in Table 3.8-7. (Q&R 220.41) A3.8.6.1.1.1.3 Method of Analysis DYNAX, a validated Sargent & Lundy (S&L) proprietary computer program, was used in the dynamic analysis of the rams head loading. DYNAX is a finite element program for the static or dynamic analysis of axisymmetric shells and solids. A description of the program along with validation information is presented in Appendix C. The analysis for the rams head loading was performed using white noise transfer functions. Structural responses were calculated in the frequency domain and then transferred to the time domain using the S&L proprietary computer program FAST. A description of the program along with validation information is presented in

Appendix C. A3.8.6.1.1.2 Quencher Discharge A3.8.6.1.1.2.1 Loading The quencher SRV loadings are described in detail in Section A3.9.2.2 of Attachment A3.9. Fifty-nine trial load cases for the quencher discharge were supplied by GE. Out of the 59 load cases, 11 were randomly selected for analysis to represent a nonexceedance probability of at least 84/84. Three cases of quencher SRV actuations were considered in this analysis: All Valve, automatic depressurization system (ADS), and single valve - subsequent actuation. A3.8.6.1.1.2.2 Model for Analysis The analysis for the quencher discharge loading was done using an axisymmetric model. Figure A3.8-48 shows the soil-structure model. Thin-shell finite elements were used to model the structures. The axisymmetric thin-shell finite element models of the RPV are supplied by the General Electric Company. The soil is modeled using solid elements. Fluid finite elements described in Reference 19 are used to simulate the suppression pool water. A3.8.6.1.1.2.3 Method of Analysis The DYNAX computer program was used in the analysis for the quencher loads. The analysis was performed using white noise transfer functions and the FAST Fourier Transform of a single bubble loading (Reference 20). Responses for the multiple bubbles were obtained by appropriate superposition of various single bubble responses. The RSG computer program was used to obtain a response spectrum from the acceleration time history for each of the 11 randomly selected load cases. The response spectra for equipment and subsystem design were then obtained by enveloping the response spectra for the 11 loadings. RSG, a validated S&L proprietary computer program, is described in

Appendix C.

CPS/USAR CHAPTER 03 A3.8-21 REV. 11, JANUARY 2005 A3.8.6.2 Analysis for LOCA Loads The LOCA loads included for structural analysis were those due to annulus pressurization, LOCA bubble, froth impingement, condensation oscillation, and chugging. Descriptions of the analysis methods for each of the loads are presented in Subsections A3.8.6.2.1 through A3.8.6.2.5, respectively. A3.8.6.2.1 Annulus Pressurization Annulus pressurization loads were considered in the shield wall analysis and design. A discussion regarding annulus pressurization can be found in Subsection 6.2.1.2.3.2. A3.8.6.2.2 LOCA Bubble A3.8.6.2.2.1 Loading The LOCA bubble loading is described in detail in Section A3.8.3 of this attachment.

A3.8.6.2.2.2 Model for Analysis The finite-element model described in Subsection A3.8.6.1.1.2.2 (Figure A3.8-48) was used to analyze the LOCA bubble load effects. A3.8.6.2.2.3 Method of Analysis The DYNAX computer program was used to obtain the dynamic structural responses due to the LOCA bubble load time histories. Displacements, accelerations, forces and moments of the structures were obtained using a direct numerical integration algorithm to solve the governing differential equations. The LOCA bubble pressure loading was applied to the model as equivalent concentrated nodal force time histories. These time histories were applied on the suppression pool basemat ad on the pool walls to a height 18 feet above the pool surface. The nodal force time histories represent the meriodinal distribution of the load on the drywell and containment walls and the radial distribution of the pressure on the basemat. Fourier harmonics were used to account for the circumferential distribution of the load. The resultant acceleration response time histories at critical locations were used to obtain response spectra for equipment and sub-system design. RSG was used to calculate the acceleration response spectra. The asymmetric LOCA bubble load was applied over 180

° of the suppression pool boundary, including basemat and pool walls. Responses due to this asymmetric loading are considered only for local containment structural evaluation. A3.8.6.2.3 Froth Impingement A3.8.6.2.3.1 Loading The froth impingement loading is described in detail in Section A3.8.3 of this attachment. The peak froth impingement pressure is 15 psid. The load time history is shown in Figure A3.8-18.

CPS/USAR CHAPTER 03 A3.8-22 REV. 11, JANUARY 2005 A3.8.6.2.3.2 Model for Analysis The finite-element model described in Subsection A3.8.6.1.1.2.2 (Figure A3.8-48) was used to analyze the froth impingement load effects. A3.8.6.2.3.3 Method of Analysis The method of dynamic analysis for the froth impingement load is similar to that described in Subsection A3.8.6.2.2.3. The impingement pressures are applied as nodal load time histories to the drywell floor. Structural responses are obtained using the DYNAX computer program. Acceleration response spectra are obtained at critical locations for equipment and subsystem design using the RSG

computer program. A3.8.6.2.4 Condensation Oscillation A3.8.6.2.4.1 Loading The condensation oscillation loading is described in detail in Section A3.8.3 of this attachment. The peak pressure specified for design is

+/- 7 psid, as shown in Figure A3.8-19. The frequency of the condensation oscillation load varies, as shown in Equation A3.8-7. A3.8.6.2.4.2 Model for Analysis The finite-element model described in Subsection A3.8.6.1.1.2.2 (Figure A3.8-48) was used to analyze the condensation oscillation load effects. A3.8.6.2.4.3 Method of Analysis The dynamic analysis for the condensation oscillation load is similar to that described in Subsection A3.8.6.2.2.3. The DYNAX computer program was used to obtain structural responses for the axisymmetric condensation oscillation pressures. The load was applied to the suppression pool boundary as nodal force time histories. The RSG computer program was used to obtain acceleration response spectra for equipment and subsystem design at critical locations.

A3.8.6.2.5 Chugging A3.8.6.2.5.1 Loading The chugging loading is described in detail in Section A3.8.3 of this attachment. Symmetric and asymmetric chugging loads are defined by GE. The time histories applied at the top vent on the drywell wall consist of a vertical force pulse train made up of triangularly shaped spikes. The peak pressure pulse train occurs within the top vent and has a magnitude of 540 psid. Refer to

Figure A3.8-21. The time histories applied in the weir annulus area consist of two phases: a sinusoidal pre-chug underpressure, and a pressure pulse train made up of triangularly shaped spikes. The peak pressure pulse train applied to the weir annulus is 43 psid. Refer to Figure A3.8-27.

CPS/USAR CHAPTER 03 A3.8-23 REV. 11, JANUARY 2005 The time histories applied in the suppression pool area consist of three phases: a sinusoidal pre-chug underpressure, a triangularly shaped pulse (pressure spike), and a post-chug oscillation defined as an exponentially decaying sine function. The peak spike pressure (i.e.,

the peak time history pressure) occurs on the drywell wall and has a magnitude of 100 psid. Refer to Figure A3.8-30 and Table A3.8-6. A3.8.6.2.5.2 Model for Analysis The finite-element model described in Subsection A3.8.6.1.1.2.2. (Figure A3.8-48) was used to analyze the chugging load effects. A3.8.6.2.5.3 Method of Analysis The method of analysis for the symmetric and asymmetric chugging load is similar to that described in Subsection A3.8.6.2.2.3 Loads were applied to the containment wall, drywell wall, and base mat as nodal force time histories. Responses due to the symmetric and asymmetric chugging loads were obtained using the DYNAX computer program. Acceleration response spectra due to symmetric chugging loads were obtained at critical locations for equipment and subsystem design using the RSG computer program. Responses due to asymmetric chugging loads are considered only for local containment structural evaluation.

A3.

8.7 REFERENCES

1. "Supplement No. 1 to the Safety Evaluation Report by the Office of Nuclear Reactor Regulation, U.S. Nuclear Regulatory Commission in the Matter of Illinois Power Company, Clinton Power Station, Units 1 and 2, Docket Nos. 50-461 and 50-462,"

NUREG-75-013, December 1975. 2. "Information Report - Mark III Containment Dyanmic Loading Conditions", Appendix B, General Electric Company Report, NED0-11314-08, submitted as Amendment 37 to 238-NI-GESSAR, December 1975, and as revised through Amendment 43. 3. "Safety Relief Valve Discharge Analytical Models", by P. Valandani, General Electric Company Report NEDE-20942-P, May 1975. 4. Hydrodynamics , by H. Lamb, 6th Edition, Cambridge Press, 1945.

5. Theoretical Hydrodynamics, L. M. Milne-Thomson, 5th Edition, MacMillan, 1968.

6. Underwater Explosions, by R. H. Cole, Dover Press, 1965. 7. "On the Best Location of a Mine Near the Sea Bed," by M. Friedman and B. Shiffman, NDRC, AM37.14, May 1944, (Unclassified). 8. "Safety-Relief Valve Discharge Analytical Models," by P. Valandani, General Electric Company Report, NED0-20942, May 1975. 9. "Model Verification Test for Safety/Relief Valve Air Clearing Loads," by H. Choe, General Electric Company Report, NEDM-13423, July 1975.

CPS/USAR CHAPTER 03 A3.8-24 REV. 11, JANUARY 2005

10. Theoretical Acoustics, by P. M. Moore and Ingard, McGrawHill, 1968.

11. Methods of Theoretical Physics, Volumes 1 and II, by P. M. Moore and Fishback, McGraw-Hill, 1953.

12. Advanced Mechanics of Fluids, edited by H. Rouse, John Wiley, 1959. 13. "Hydrodynamic Images", by L. M. Milne-Thomson, Proceedings of the Cambridge Philosophical Society, Vol. 36, 1940.

14. Handbook of Applied Mathematics, edited by C. E. Pearson, Van Nostrand Rheinhold Co., 1974. 15. "The General Electric Mark III Pressure Suppression Containment System Analytical Model", by W. J. Bilanin, General Electric Company Report, NED0-20533, June 1974, and Supplement 1, August 1975. 16. "Mark III Confirmatory Test Program - 1/ 3 Scale Condensation and Stratification Phenomena-Test Series 5807", General Electric Company Report, NEDE-21596-P, March 1977. 17. "Mark III Confirmatory Test Program Phase 1 - Large Scale Demonstration Tests," General Electric Company Report NEDM-13377, October 1974. 18. General Electric Co., Appendix 3B of 238-NI-GESSAR(22A7000) Revision 2, submitted for GESSAR docket STN 50-447, dated June 12, 1981, application for final design approval. 19. Transmission of Shock Waves into Submerged Fluid Filled Vessels", by A. J. Kalinowski, ASME Conference on FSI Phenomena in Pressure Vessel and Piping Systems, TRP-

TB-026, 1977. 20. A. K. Singh et al., "Response Using Dynamic Influence Coefficients," Seventh Conference on Electronic Computation, ASCE, St. Louis, Mo., August 1979. 21. "Containment Loads," Appendix 3B to 238-NI-GESSAR II, Docket No. STN 50-447, General Electric Report 22A7000, Rev. 2 dated June 12, 1981. 22. "Draft Technical Evaluation Report on Mark III LOCA-Related Hydrodynamic Load Definition - Generic Technical Activity B-10". A report by the Containment Systems Branch, Division of Systems Integration, U. S. Nuclear Regulatory Commission, transmitted by Mr. T. P. Speis to Mr. H. Pfefferlen of General Electric Company on

October 8, 1982.

CPS/USAR CHAPTER 03 A3.8-25 REV. 11, JANUARY 2005 TABLE A3.8-1 BUBBLE DYNAMICS EQUATIONS Differential Equations

=R RPm PR4 nk3PP b d d 3 (3.8A-8) R/R 2 3 g)PP(R 2 L c= (3.8A-9) where: P = P ref + L Z. For (t > td): vm/FZ= where gmgR 3 4ZZRC 2 1 b 3 L 2DL+=, and (3.8A-10) 3 Lbv R b 2mm+=. (3.8A-11)

Then dtzzz t t t12 2 1+=, and (3.8A-12)

CPS/USAR TABLE A3.8-1 (Cont'd) CHAPTER 03 A3.8-26 REV. 11, JANUARY 2005

)tt(Rdtzzz12 t t12 t 2 1+=

• last term included for 0R> only (3.8A-13) Initial Conditions (t = 0):

8 VR;2 DR;PgkV;g4 V 2 P P L d d 2 d 2ddd===+= (3.8A-15) )tt(V 2Rm;)tt0(0z;LZ d d d 2 b d s=== (2.8A-16) where: as 2 b)LL(D8/m= for a rams head, and (3.8A-17) bbdm/mt=. (3.8A-18)

CPS/USAR CHAPTER 03 A3.8-27 REV. 11, JANUARY 2005 TABLE A3.8-2 EXTREME VALUES OF COEFFICIENTS FOR SYMMETRIC DISCHARGE CASE ZONE A 0 B 1 A 1 B 2 A 2 B 3 A 3 B 4 A 4 1. MAX 10.930 2.204 4.111 0.432 2.027 1.571 1.572 0.432 0.541 MIN -3.470 -1.671 -3.266 -1.824 -3.408 -1.694 -1.460 -0.942 -0.726 2. MAX 26.238 3.862 8.157 0.935 3.542 3.719 3.020 0.902 1.247 MIN -7.666 -3.879 -5.614 -4.622 -7.892 -4.367 -2.642 -1.456 -1.651 3. MAX 26.667 3.573 7.930 0.911 3.412 3.399 2.703 0.802 1.061 MIN -7.865 -4.037 -5.179 -4.663 -7.580 -4.055 -2.431 -1.124 -1.386 4. MAX 25.796 3.431 7.645 0.871 3.259 3.205 2.547 0.748 0.985 MIN -7.615 -3.906 -4.973 -4.442 -7.237 -3.840 -2.292 -1.033 -1.254 5. MAX 21.230 2.828 6.227 0.688 2.583 2.446 1.947 0.544 0.722 MIN -6.254 -3.156 -4.100 -3.491 -5.759 -2.907 -1.751 -0.773 -0.944 6. MAX 17.511 2.362 5.072 0.539 2.029 1.823 1.454 0.375 0.505 MIN -5.146 -2.545 -3.433 -2.713 -4.514 -2.142 -1.306 -0.558 -0.662 7. MAX 16.611 2.253 4.820 0.508 1.915 1.711 1.365 0.348 0.471 MIN -4.879 -2.405 -3.274 -2.554 -4.261 -2.005 -1.226 -0.525 -0.618 8. MAX 12.742 1.781 3.763 0.394 1.489 1.349 1.078 0.277 0.379 MIN -3.735 -1.836 -2.587 -1.967 -3.314 -1.565 -0.968 -0.435 -0.499 9. MAX 4.863 0.705 1.465 0.151 0.577 0.526 0.432 0.108 0.150 MIN -1.422 -0.696 -1.024 -0.751 -1.279 -0.603 -0.378 -0.183 -0.198 NORM (psi) 26.667 4.037 8.157 4.663 7.592 4.267 3.020 1.456 1.651 NORM (ksf) 3.5400 0.5513 1.1746 0.67115 1.1364 0.6144 0.4349 0.2097 6.2377

CPS/USAR CHAPTER 03 A3.8-28 REV. 11, JANUARY 2005 TABLE A3.8-3 EXTREME CALCULATED PRESSURES FOR SYMMETRIC DISCHARGE CASE*

ZONE P-WALL** MAX (psid)

P-WALL** MIN (psid) 1 17.20 -4.752 2 39.42 -9.383 3 36.46 -8.850 4 34.88 -8.483 5 28.91 -6.875 6 22.82 -5.570 7 21.62 -5.276 8 16.66 -4.055 9 6.382 -1.548

• The extreme values do not necessarily occur at the same point in space or time for the several zones. ** This is the pressure due to air bubble only, i.e., it does not include hydrostatic pressure.

CPS/USAR CHAPTER 03 A3.8-29 REV. 11, JANUARY 2005 TABLE A3.8-4 EXTREME VALUES OF FOURIER COEFFICIENTS FOR ASYMMETRIC DISCHARGE CASE ZONE A 0 B 1 A 1 B 2 A 2 B 3 A 3 B 4 A 4 1. MAX 1.004 1.821 0.361 1.377 0.543 0.888 0.517 0.479 0.363 MIN -0.436 -0.734 -0.969 -0.422 -1.303 -0.567 -0.991 -0.874 -0.935 2. MAX 2.405 4.342 0.784 3.241 1.125 2.032 1.003 1.014 0.655 MIN -0.962 -1.592 -2.318 -0.876 -3.077 -1.252 -2.270 -1.892 -2.047 3. MAX 2.441 4.339 0.791 3.104 1.089 1.830 0.915 0.849 0.555 MIN -0.985 -1.606 -2.322 -0.847 -2.956 -1.132 -2.052 -1.588 -1.711 4. MAX 2.361 4.182 0.763 2.962 1.041 1.723 0.863 0.784 0.515 MIN -0.953 -1.549 -2.239 -0.810 -2.822 -1.067 -1.933 -1.471 -1.585 5. MAX 1.944 3.408 0.621 2.350 0.824 1.319 0.659 0.577 0.378 MIN -0.783 -1.260 -1.824 -0.641 -2.238 -0.816 -1.479 -1.082 -1.166 6. MAX 1.604 2.777 0.505 1.850 0.647 0.987 0.491 0.406 0.264 MIN -0.645 1.024 1.486 0.503 1.760 0.610 1.105 0.759 0.819 7. MAX 1.521 2.631 0.478 1.747 0.610 0.927 0.461 0.379 0.247 MIN -0.611 -0.970 -1.408 -0.475 -1.662 -0.573 -1.038 -0.709 -0.765 8. MAX 1.167 2.026 0.367 1.359 0.474 0.734 0.363 0.306 0.199 MIN -0.468 -0.745 -1.083 -0.369 -1.292 -0.453 -0.821 -0.572 -0.619 9. MAX 0.446 0.776 0.140 0.525 0.182 0.287 0.142 0.122 0.079 MIN -0.178 -0.285 -0.415 -0.142 -0.499 -0.177 -0.321 -0.228 -0.246 NORM (psid) 2.441 4.342 2.322 3.241 3.077 2.032 2.270 1.892 2.047 NORM (ksf) 0.3515 0.6252 0.3344 0.4667 0.4431 0.2926 0.3269 0.2724 0.2948

CPS/USAR CHAPTER 03 A3.8-30 REV. 11, JANUARY 2005 TABLE A3.8-5 EXTREME CALCULATED PRESSURES FOR THE ASYMMETRIC DISCHARGE CASE*

ZONE P-WALL** MAX (psid)

P-WALL** MIN (psid) 1 9.379 -4.202 2 20.390 -7.611 3 7.120 -6.551 4 16.120 -6.158 5 12.580 -4.756 6 9.789 -3.610 7 9.232 -3.398 8 7.214 -2.663 9 2.796 -1.033

• The extreme values do not necessarily occur at the same point in space or time for the several zones. ** This is the pressure due to air bubble only, i.e., it does not include hydrostatic pressure.

CPS/USAR CHAPTER 03 A3.8-31 REV. 11, JANUARY 2005 TABLE A3.8-6 CHUGGING LOADS*

LOCATION PRE-CHUG UNDERPRESSUREAND DURATION PULSE (SPIKE) AND DURATION "d" POST-CHUG OSCILLATION AND FREQUENCY PEAK (A) MEAN (A) PEAK MEAN PEAK (B) MEAN (B) Drywell Wall -5.8 psid -2.65 psid 100 psid 24 psid

+/-6.50 psid

+/-2.2 psid 125 ms 125 ms 8 ms 8 ms 10-12 hertz 10-12 hertz Containment -1.3 psid -1.0 psid 3 psid 0.7 psid

+/-1.7 psid +/-1.00 psid 125 ms 125 ms 2 ms 2 ms 10-12 hertz 10-12 hertz Base Mat -1.8 to

-1.3 psid

-1.34 to -1.0 psid 10 to 3 psid 2.4 to 0.7 psid +/-2.1 to +/-1.7 psid +/-1.29 to +/-1.0 psid 125 ms 125 ms 4 to 2 ms 4 to 2 ms 10-12 hertz 10-12 hertz

• Source: Reference 18 CPS/USAR CHAPTER 03 A3.8-32 REV. 11, JANUARY 2005 TABLE A3.8-7 (Q&R 220.41) SOIL STRAIN VERSUS MODULUS SOIL LAYER*

FROM TOP STRAIN MODULUS OF ELASTICITY E (ksf) 1 0.1034x10

-3 62.5x10 2 2 0.1106x10

-3 62.5x10 2 3 1.1134x10

-3 62.5x10 2 4 0.1410x10

-3 62.5x10 2 5 0.1928x10

-4 56.0x10 3 6 0.1871x10

-4 57.0x10 3 7 0.1814x10

-4 58.0x10 3 8 0.1814x10

-4 58.0x10 3 9 0.1571x10

-4 60.0x10 3 10 0.1393x10

-4 60.0x10 3 11 0.1215x10

-4 66.0x10 3 12 0.9900x10

-5 71.0x10 3 13 0.7684x10

-5 80.0x10 3 14 0.7475x10

-5 80.0x10 3 15 0.7265x10

-5 80.0x10 3 16 0.1085x10

-4 38.1x10 3 17 0.8488x10

-5 58.0x10 3 18 0.7230x10

-5 60.0x10 3 19 0.6651x10

-5 62.5x10 3 20 0.6071x10

-5 65.0x10 3 21 0.7474x10

-5 60.0x10 3 Bed Rock

• For soil layer see Figure A3.8-48.

CPS/USAR ATTACHMENT B3.8 CONTAINMENT STRUCTURAL DESIGN ASSESSMENT

CPS/USAR CHAPTER 03 B3.8-i REV. 11, JANUARY 2005 ATTACHMENT B3.8 - CONTAINMENT STRUCTURAL DESIGN ASSESSMENT TABLE OF CONTENTS PAGE B3.

8.1 INTRODUCTION

B3.8-1 B3.8.2 STRUCTURAL RESPONSE DUE TO SRV AND LOCA LOADS B3.8-1 B3.8.3 DESIGN ASSESSMENT B3.8-2 B3.

8.4 CONCLUSION

S B3.8-2 CPS/USAR CHAPTER 03 B3.8-ii REV. 11, JANUARY 2005 ATTACHMENT B3.8 - CONTAINMENT STRUCTURAL DESIGN ASSESSMENT LIST OF TABLES NUMBER TITLE PAGE B3.8-1 Design Assessment Stresses for Loads without Temperature B3.8-3 B3.8-2 Design Assessment Stresses for Loads with Temperature B3.8-4

CPS/USAR CHAPTER 03 B3.8-iii REV. 11, JANUARY 2005 ATTACHMENT B3.8 - CONTAINMENT STRUCTURAL DESIGN ASSESSMENT LIST OF FIGURES NUMBER TITLE B3.8-1 Force Plots - Containment Wall SRV - All Valve B3.8-2 Force Plots - Containment Wall LOCA Bubble B3.8-3 Force Plots - Containment Wall LOCA - Froth Impingement B3.8-4 Force Plots - Containment Wall LOCA - Condensation Oscillation B3.8-5 Force Plots - Containment Wall LOCA - Chugging B3.8-6 Force Plots - Drywell SRV - All Valve B3.8-7 Force Plots - Drywell LOCA Bubble B3.8-8 Force Plots - Drywell LOCA - Froth Impingement B3.8-9 Force Plots - Drywell LOCA - Condensation Oscillation B3.8-10 Force Plots - Drywell LOCA - Chugging B3.8-11 Locations of Design Assessment Sections

CPS/USAR CHAPTER 03 B3.8-1 REV. 11, JANUARY 2005 CONTAINMENT STRUCTURAL DESIGN ASSESSMENT B3.

8.1 INTRODUCTION

The functions and capabilities of the containment structure, the loads and load combinations used in the containment structural design, and the applicable design codes and regulatory guides have been described in Section 3.8 and Attachment A3.8. Structural responses due to normal, pressurization, and earthquake loads have been included in Section 3.8. This attachment presents responses of the containment structure due to the safety/relief valve (SRV) and loss-of-coolant accident (LOCA) suppression pool dynamic loads, and an assessment of the containment structure for the final design loads. Section B3.8.2 provides design responses of the containment wall and dome and the drywell structures due to the SRV and LOCA loads. Section B3.8.3 includes the assessment of the structural capacity with respect to the critical load combinations for the final design loads. B3.8.2 STRUCTURAL RESPONSE DUE TO SRV AND LOCA LOADS Responses of the containment wall and dome and the drywell for loads other than SRV and LOCA have been presented in Section 3.8. Containment wall and dome force responses due to dead load, suppression pool hydrostatic load, accident pressure load, and safe-shutdown earthquake are shown in Figure 3.8-17. Similar responses for the drywell are shown in Figure

3.8-33. Typical force response plots of the containment wall and dome due to the SRV and LOCA loads are shown in Figures B3.8-1 through B3.8-5. Typical drywell force response plots are presented in Figures B3.8-6 through B3.8-10. It should be noted that these plots represent the envelope of the maximum positive and the envelope of the maximum negative values that occur along the height of the structure and that the positive and negative values at a particular elevation do not occur concurrently. Response plots for the SRV loading are for the all valve case and for reasons of simplicity are plotted for only one of the eleven trials which were analyzed (refer to Subsection A3.8.6.1.1.2). For the LOCA loads, the symmetric loading case is used for the force plots. In the assessment of the structural capacity, however, various trials of the SRV loads and symmetric and asymmetric cases of the LOCA loads are enveloped. B3.8.3 DESIGN ASSESSMENT Results of the design assessment for the containment structure are provided in Tables B3.8-1 and B3.8-2. Six representative locations on the basemat, containment wall, containment dome, and drywell are considered as shown in Figure B3.8-11. Concrete and reinforcing stresses due to membrane and flexural forces are calculated for both meridional and circumferential sections. Sections are evaluated for the loads and load combinations shown in Table 3.8-1.1. The critical stresses for the service and factored load categories are tabulated for two conditions, loads without temperature (Table B3.8-1) and loads with temperature (Table B3.8-2).

CPS/USAR CHAPTER 03 B3.8-2 REV. 11, JANUARY 2005 The stresses due to combined bending and axial loads with and without temperature loads are calculated using the Sargent and Lundy proprietary computer program TEMCO. The program is described in Appendix C. B3.

8.4 CONCLUSION

As shown in Tables B3.8-1 and B3.8-2, the stresses are all within the allowables for the final design loads and therefore the design is adequate.

CPS/USAR CHAPTER 03 B3.8-3 REV. 11, JANUARY 2005 TABLE B3.8-1 DESIGN ASSESSMENT STRESSES FOR LOADS WITHOUT TEMPERATURE SERVICE CONDITION FACTORED CONDITION REINFORCING TENSION(ksi) CONCRETE COMPRESSION(psi) REINFORCING TENSION(ksi) CONCRETE COMPRESSION(psi)

SECTION (1) MERIDIONAL HOOPMERIDIONAL HOOP MERIDIONAL HOOP MERIDIONAL HOOP 1 30 (4) 17 (4) 840 (4) 690 (4) 43 (19) 20 (19) 1150 (19) 810 (19) 2 6 (3) 26 (3) 510 (4) 220 (4) 22 (19) 37 (13) 630 (19) 250 (11) 3 19 (3) 23 (3) 160 (4) 140 (4) 27 (6) 32 (6) 190 (11) 170 (11) 4 18 (4) 11 (4) 1260 (4) 570 (4) 46 (19) 23 (19) 1620 (11) 870 (19) 5 19 (4) 24 (3) 740 (4) 370 (3) 43 (26) 42 (25) 1110 (26) 670 (25) 6 30 (4) 21 (4) 650 (4) 710 (4) 36 (19) 30 (26) 950 (19) 920 (19)

Allowable Stresses 30 ksi 1800 psi 54 ksi 3000 psi (1) See Figure B3.8-11 for Section locations. (2) Numbers in parentheses identify the load combinations corresponding to the stress values given. Refer to Table 3.8-1.1 an d following note (3) for load combination information. (3) Load combinations 24, 25, and 26 are load combinations 6, 13, and 19 of Table 3.8-1.1 respectively, except with design pres sure of 30 psi and without SRV loads. Load combinations 27 and 28 are load combinations 13 and 19 except with long term temperature and without accident pressure or compartment pressure, no LOCA or SRV loads.

CPS/USAR CHAPTER 03 B3.8-4 REV. 11, JANUARY 2005 TABLE B3.8-2 DESIGN ASSESSMENT STRESSES FOR LOADS WITH TEMPERATURE SERVICE CONDITION FACTORED CONDITION REINFORCING TENSION(ksi) CONCRETE COMPRESSION(psi) REINFORCING TENSION(ksi) CONCRETE COMPRESSION(psi)

SECTION (1) MERIDIONAL HOOP MERIDIONAL HOOP MERIDIONAL HOOP MERIDIONAL HOOP 1 29 (4) 19 (4) 780 (4) 780 (4) 41 (19) 22 (19) 1090 (19) 890 (19) 2 11 (3) 29 (3) 810 (4) 410 (4) 35 (19) 35 (13) 1250 (19) 630 (19) 3 17 (3) 17 (3) 790 (4) 830 (4) 31 (6) 27 (19) 1130 (19) 1110 (19) 4 19 (4) 8 (4) 1790 (4) 1260 (4)37 (11) 17 (11) 2110 (11) 1440 (11) 5 30 (4) 28 (4) 1030 (4) 330 (4) 54 (26) 55 (26) 1630 (19) 1310 (25) 6 39 (4) 33 (4) 1110 (4) 1240 (4)42 (28) 43 (28) 1520 (19) 1580 (19)

Allowable Stresses 40 ksi 1800 psi 60 ksi 3000 psi (1) See Figure B3.8-11 for Section locations. (2) Numbers in parentheses identify the load combinations corresponding to the stress values given. Refer to Table 3.8-1.1 and following note (3) for load combination information. (3) Load combinations 24, 25, and 26 are load combinations 6, 13, and 19 of Table 3.8-1.1 respectively, except with design pre ssure of 30 psi and without SRV loads. Load combinations 27 and 28 are load combinations 13 and 19 except with long term temperature and without accident pressure or compartment pressure, no LOCA or SRV loads.

CPS/USAR ATTACHMENT C3.8 EVALUATION OF SAFETY-RELATED MASONRY WALLS

CPS/USAR CHAPTER 03 C3.8-1 REV. 11, JANUARY 2005 EVALUATION OF SAFETY-RELATED MASONRY WALLS C3.8.1 Criteria Used for Design of Category 1 Masonry Walls at Clinton Station The loads and load combinations used for the design of Category 1 masonry walls for the Clinton Station are shown in Table C3.8-1. The material properties for the masonry walls are shown in Table C3.8-2. The basic allowable stresses for unreinforced masonry walls are those given in UBC 1979 for inspected workmanship and are shown in Table C3.8-3. The seismic damping values for masonry walls are 2% for normal/severe load combinations and 4% for extreme load combinations. C3.8.2 Design and Analysis Considerations The seismic analysis for the safety related concrete masonry walls is in accordance with the requirements of the Clinton Station USAR, Section 3.7. The assumptions and modeling techniques consider proper boundary conditions, cracking of concrete, if any, and the dynamic behavior of the masonry walls. The analysis of the safety related masonry walls consider both in-plane and out-of-plane loads, and interstory drift effects are considered. There are no concrete masonry shear walls at the Clinton Station, and there are no safety related concrete masonry walls which are subject to accident pipe reaction (Yr), jet impingement (Yj), or missile impact (Ym).

CPS/USAR CHAPTER 03 C3.8-2 REV. 11, JANUARY 2005 Table C3.8-1 Load Combinations for Category 1 Concrete Masonry SRV* LOCA - Pool Dynamics* Allowable Stresses Load Category D L E E 1 ALL 1V2P ADS PS CH CO MYC Masonry Steel 1.0 1.0 Normal 1.0 1.0 1.0 AISC Sovern Environmental 1.0 1.0 1.0 1.0 AISC 1.0 1.0 1.0 1.0 1.0 1.0 (UBC Allowable)

Abnormal 1.0 1.0 1.0 1.0 1.6 x AISC .95 Fy Extreme Environmental 1.0 1.0 1.0 1.0 1.6 x AISC .95 Fy 1.0 1.0 1.0 1.0 1.0 1.0 1.0 Abnormal/Severe Environmental 1.0 1.0 1.0 1.0 1.0 1.6 x AISC .95 Fy 1.0 1.0 1.0 1.0 1.0 1.0 1.0 Abnormal/Extreme Environmental 1.0 1.0 1.0 1.0 1.0 1.6 x AISC .95 Fy *Only one load under each of these loadings shall be considered at one time. Load Symbols are defined as follows: D = Dead load of masonry wall including atachment loads L = Live load E = Operating Basis Earthquake (OBE)

E 1 = Safe Shutdown Earthquake (SSE)

SRV 1V2P = Safety/Relief Valve (SRV) discharge loading due to discharge of one Safety/Relief Valve subsequent activation.

SRV ADS = SRV loading due to seven (ADS) Safety/Relief Valve discharges SRV ALLV = SRV loadng due to 16 (ALL) Safety/Relief Valve discharges LOCA MVC = LOCA loading due to main vent clearing LOCA PS = LOCA loading due to pool swell LOCA CO = LOCA loading due to condensation oscillation LOCA CH = LOCA loading due to chugging CPS/USAR CHAPTER 03 C3.8-3 REV. 11, JANUARY 2005 Table C3.8-2 MASONRY MATERIAL PROPERTIES

1. Hollow Concrete Masonry Blocks ASTM C90, Grade N-I Minimum ultimate compressive strength (fc) shall be 1000 psi based upon the gross cross-sectional area of the blocks 2. Solid Concrete Masonry Blocks ASTM C145, Grade N-I Minimum ultimate compressive strength (fc) shall be 1800 psi based upon the gross cross-section area of the blocks. 3. Grouted Concrete Masonry Blocks Hollow Blocks: ASTM C90, N-I Grout : ASTM C476 4. Mortar, Type M* ASTM C270 with minimum compressive strength m o = 2,500 psi 5. Masonry Compressive Strength, f m 1,350 psi, Type M mortar 6. Reinforcement for Concrete Masonry Truss or ladder type F y = 65ksi, ASTM A82 *For the Clinton Station Type M mortar shall be used for all masonry construction.

CPS/USAR CHAPTER 03 C3.8-4 REV. 11, JANUARY 2005 Table C3.8-3 ALLOWABLE STRESSES/STRAINS FOR CONCRETE MASONRY INSPECTED WORKMANSHIP UNREINFORCED MASONRY (Table Values are based on UBC 1979)

Allowable Stresses (PSI) Normal/Severe Environ. Load Comb. Extreme Envison. Load Comb Type M Mortar Type M Mortar Over Stress Factor Used Ult Shr Strn Type of Stress Type of Unit (c) m o = 2500 PSI m o = 2500 PSI 1. Compression Flexural and Axial F m & F a H S G 170 175 225 425 438 562 2.5 2.5 2.5 2. Bearing Under Conc. Loads H S G 255 262 337 637 655 842 2.5 2.5 2.5 Tension in Flexure

1) Normal to bed Joint F t1 H S G 12 (a) 12 12 12 (a) 12 12 1.0 1.0 1.0 3. 2) Parallel to bed Jonts F t11 H S G 24 24 50 36 36 75 1.5 1.5 1.5 4. Shear Out of Plane Loads and In-Plane Loads H S G 12 12 25 12 12 25 1.0 1.0 1.0 (b) .001 (a) This value is allowed for tensile stresses due to inplane loads for existing walls only. Use F t1 = 0 psi for new walls. F t1 = 12 psi is also allowed for wall strip around opening spanning vertically. (b) Shear strain is equal to /h where is relative horiz. floor displacement and h is the height of wall. (c) H = Hollow Units G = Grouted Units S = Solid Units

CPS/USAR CHAPTER 03 3.10-1 REV. 11, JANUARY 2005 3.10 SEISMIC QUALIFICATION OF SEISMIC CATEGORY I INSTRUMENTATION AND ELECTRICAL EQUIPMENT Seismic qualification includes the equipment qualification due to applicable hydrodynamic and LOCA loads which are addressed in Section A3.9. This section addresses the dynamic qualification of all Class lE electrical equipment, instrumentation and their supports. Section 3.9 addresses the dynamic qualification of safety-related mechanical equipment and Section 3.11 addresses the environmental qualification of Class lE electrical equipment and instrumentation. All Class lE electrical equipment and instrumentation are designed to withstand, without functional impairment, the effects of the Safe Shutdown Earthquake (SSE) defined in Subsection of 3.7.1 and the hydrodynamic loads discussed in Section A3.9.

The requirements of IEEE-344 and Regulatory Guide 1.100 are met for equipment identified in this section.

Per Section 6.1.1 of IEEE Standard 344-75, electrical equipment must be tested on a shake table with mounting and configuration similar to actual service, unless adequate justification can be made to extend the qualification to an untested orientation or configuration. Safety-related switchgears in Div. 1, 2 and 3 Auxiliary Power Systems were seismically qualified with breakers in racked-in/racked-up/connected configurations to meet the requirements of IEEE-344-75. However, to address certain breaker configurations other than the original qualification (which may exist during on-line maintenance activities of the switchgear breakers), an evaluation has been performed and has been determined that there is no adverse impact on the intended safety function of the affected switchgear and other devices (such as relays, instruments) in adjacent cubicles in service, provided the duration of these seismically unanalyzed conditions is for a limited period of time. The switchgears addressed in this evaluation are safety-related Div. 1 and 2, 4160V/6900V switchgear, Div. 3 HPCS switchgear and 480V Unit Substation switchgear in the Unit Auxiliary

power systems. The identification of all Class lE BOP and NSSS electrical equipment and instrumentation utilized in various systems of the plant is provided in Nuclear Station Engineering Standard MS-02.00. The information provided in MS-02.00, Maintenance of Equipment Qualification Program Manual is as follows: a. Equipment Number and Name: Provides the specific number and name of the equipment. This provides a correlation between other documents and drawings. b. Equipment Manufacture/Model: Identifies the manufacturer or vendor of the equipment and equipment model number. c. Qualification Package: List the specific document package which demonstrates qualification. Equipment functional times, environmental zones, and equipment categories can be obtained from the respective environmental qualification document package referenced in MS-02.00.

CPS/USAR CHAPTER 03 3.10-2 REV. 11, JANUARY 2005 3.10.1 Seismic Qualification Criteria 3.10.1.1 BOP and NSSS Compliance with IEEE-344 BOP and NSSS Class lE electrical equipment and instrumentation is qualified to meet the requirements of IEEE-344. The dynamic qualification and its documentation is verified to show that the equipment performs its function during and after the SSE event. This event is combined with any applicable hydrodynamic event as discussed in Section A3.9. Analysis, testing, or a combination of test and analysis is used to qualify the equipment. The method of qualification is identified in the qualification document package referenced in MS-02.00. An analysis or test or a combination of test and analysis is used for qualifying Class lE electrical equipment. Analytical methods when used are consistent with Section 5.0 of IEEE-344. Testing is primarily used for qualifying Class lE electrical active equipment. The equipment tested is in compliance with Section 6.0 of IEEE-344. When the equipment could not be practically qualified by methods using analysis or testing alone because of its size and/or complexity, a method of combined analysis and testing was used. 3.10.2 Methods and Procedures for Qualifying Electrical Equipment and Instrumentation 3.10.2.1 BOP Equipment All Class lE electrical equipment and instrumentation are identified in Nuclear Station Engineering Standard MS-02.00. These components are designed to withstand the effects of the SSE, hydrodynamic and LOCA loads as applicable, without f unctional impairment. For each unique piece of instrumentation or electrical equipment a dynamic qualification report has been prepared by the equipment vendor in accordance with the requirements of the equipment procurement specifications. The load combinations used for BOP equipment qualification are those delineated in Tables A3.9-6 and A3.9-7. The equipment is qualified in the entire frequency range of interest to include the higher frequencies (greater than 33 Hz) for the hydrodynamic and LOCA events. For those components qualified by testing, the requirements of IEEE-344 are verified or adequate justification is provided that the component does meet these requirements. Since several components/devices are used at different locations for the same application or for a different application, the approach taken for qualification is to test the component for the bounding or worst case location and for the different applications. For example, a relay in one system may have as its safety function to deenergize and open its contact within a certain time, while in another system it may be required to energize and close its contacts. In such a situation, the relay would be tested in both modes under the worst case dynamic conditions to assure operability. To the extent practical, dynamic qualification tests of equipment were performed while the equipment was subjected to normal operating loads. Where it was demonstrated by prior testing or analysis that operating loads such as pressure, torque, flow, voltage, current, and thermal expansion did not cause significant stress loads within the equipment, or where such operating loads are not significant to equipment operability, operation under such loads during testing was not required. The equipment was monitored and evaluated during and after the test CPS/USAR CHAPTER 03 3.10-3 REV. 11, JANUARY 2005 for malfunction or failure and, upon completion of the test, was thoroughly inspected for damage. The results of such tests were documented and reviewed for compliance with IEEE-

344. Testing was used for the seismic qualification of equipment auxiliary components, such as relays, switches, and instruments necessary for proper operation. As far as possible, these components were tested and qualified with the equipment mounted in a manner similar to the field mounting condition. The input motion was applied simultaneously to the vertical axis and one principal horizontal axis, or when single axis tests were performed, adequate justification was provided. The maximum input motion acceleration was equal to or in excess of the maximum dynamic acceleration at the equipment mounting. The qualification document package referenced in Nuclear Station Engineering Standard MS-02.00 gives specific details on methods, results, and analysis. In some cases where it was found that different pieces of equipment have similar characteristics, the test program was based upon testing of prototype equipment. The test

reports furnished by the equipment supplier we re reviewed for compliance with IEEE-344. 3.10.2.2 NSSS Equipment All Class lE NSSS instrumentation and electrical equipment are identified in Nuclear Station Engineering Standard MS-02.00. These components are designed to withstand the effects of the SSE, hydrodynamic and LOCA loads as applicable, without f unctional impairment. For each unique piece of instrumentation or electrical equipment, a dynamic qualification report has been prepared and assessed to verify compliance with IEEE-344 and the loading combination as defined in Table 3.9-2. The qualification shall also be verified in the entire frequency range of interest to include the higher frequencies (greater than 33 Hz) of the hydrodynamic and LOCA events, or adequate justification is provided that the component does meet the requirements. The dynamic loading criterion used in the design and subsequent qualification of all Class lE instrumentation and electrical equipment supplied by GE is as follows: The Class lE equipment shall be capable of performing all safety-related functions during (1) normal plant operation, (2) anticipated transients, (3) design basis accidents, and (4) postaccident operation while being subjected to, and after the cessation of, the accelerations resulting from the SSE or hydrodynamic loads at the point of attachment of the equipment to the building or supporting structure. The criteria for each of the devices used in the Class lE systems depend on the use in a given system; for example, a relay in one system may have as its safety function to deenergize and open its contacts within a certain time, while in another system it must energize and close its contacts. Since GE supplies many devices for many applications, the approach taken was to test the device in the worst case configuration in which it might be used. In this way, the capability of protective action initiation and the proper operation of safety-failure circuits is

assured. From the basic input ground motion data, a series of response curves at various building elevations are developed after the building layout is completed. Standard requirement levels that meet or exceed the maximum expected unique plant information is included in the purchase CPS/USAR CHAPTER 03 3.10-4 REV. 11, JANUARY 2005 specifications for Seismic Category I equipment. Suppliers of equipment such as batteries and racks, instrument racks, control consoles, etc., are required to submit test data, operating experience and/or calculations to substantiate that their components, systems, etc. will not suffer loss of function during or after dynamic loadings. The magnitude and frequency of the loadings which each component will experience are determined by its specific location within the plant. All Class lE electrical equipment and instrumentation are evaluated for their capability to perform the safety function under the combinations of seismic and hydrodynamic vibration loadings shown in Table 3.9-2. 3.10.2.3 NSSS Testing Procedures for Qualifying lE Electrical Equipment and Instrumentation (Excluding Motors and Valve-Mounted Equipment)

The test procedure required that the devices be mounted on the table of the vibration machine in a manner similar to that in which it is installed. The device was tested in the operating states in which it is to be used when performing its Class lE functions and these states were monitored before, during, and after the test to assure proper function and absence of spurious function. In the case of relays, both energized and deenergized states and normally open and normally closed contact configurations were tested if the relay is used in those configurations for its Class lE functions. The first step was to search for resonances in each device. This was done since resonances cause amplification of the input vibration and is the most likely cause of malfunction. The resonance search was usually run at low acceleration levels (0.2g) in order to avoid damaging the test sample in case a severe resonance was encountered. The resonance search was generally run up to 60 Hz to account for high frequency effects. If the device was large the vibrations were monitored by accelerometers placed at critical locations from which accelerations were measured and compared with the input acceleration level at the table to

determine resonance. The method used for qualification is a dynamic excitation with a single sinusoidal frequency with peak acceleration amplitude at several discrete frequencies. The vibratory excitation was applied in three orthogonal axes individually. Additionally, after conducting the frequency scan and resonance determination, the devices were tested to determine their malfunction limit. This test was a necessary adjunct to the assembly test as explained below. The malfunction limit test was run at each resonant frequency as determined by the resonance search. In this test, the acceleration level was gradually increased until either the device malfunctioned or the limit of the vibration machine was reached. If no resonances were detected (as was usually the case), the device was considered to be rigid, and the malfunction limit was therefore independent of frequency. To achieve maximum acceleration from the vibration machine, rigid devices were malfunction-tested at the upper test frequency of 60 Hz, since that allowed the maximum acceleration to be obtained from deflection-limited machines. The above procedures were required of purchased devices as well as those supplied by GE. Vendor test results were reviewed and if unacceptable, the tests were repeated either by GE or the vendor. If the vendor tests were adequate, the device was considered qualified to the limits of the test.

CPS/USAR CHAPTER 03 3.10-5 REV. 11, JANUARY 2005 3.10.2.4 Qualification of Valve Mounted Equipment - NSSS The piping analyses establish the response spectra, power spectral density function or time history characteristics, and horizontal and vertical accelerations for the pipe-mounted equipment. Class lE motor-operated valve actuators were qualified per IEEE-382. The safety/relief valves, including the electrical components mounted on the valve, are subjected to a dynamic test. This test is described in Subsection 3.9.3.2.1.5.2. 3.10.2.5 Qualification of NSSS Motors Seismic qualification of the ECCS motors is discussed in Subsection 3.9.2.2.1.6.7 in conjunction with the ECCS pump and motor assembly. Seismic qualification of the Standby Liquid Control pump (SLC) motor is discussed in Subsection 3.9.2.2.1.6.10 in conjunction with the SLC pump motor assembly. 3.10.3 Methods and Procedure of Analysis of Testing of Supports of Electrical Equipment and Instrumentation 3.10.3.1 BOP Seismic Category I Electrical Equipment and Instrument Supports 3.10.3.1.1 Battery Racks, Instrument Racks, Control Consoles, Cabinets and Panels Response spectra curves at the appropriate locations, consisting of the response due to seismic, hydrodynamic and LOCA loads (where applicable) have been supplied to the equipment vendor. Methods used for qualification are analysis, testing, or a combination thereof in accordance with the procedures in Subsection 3.10.2.1. The qualification results for BOP components is presented in the qualification document package referenced in Nuclear Station Engineering

Standard MS-02.00. 3.10.3.2 NSSS Dynamic Analysis, Test Procedures, and Restraint Measures 3.10.3.2.1 Instrument Racks, Control Consoles, Cabinets, and Panels Class lE electrical equipment supplied by GE is used in many systems on many different plants under widely varying dynamic loading requirements. The seismic qualification tests were generally performed at all frequencies from 5 to 60 Hz. The actual qualification range was generally 1 to 60 Hz. However, since test facility capability was sometimes limited, the lower frequency tested was 5 Hz. A combination of test and analysis was used to assure that all component resonances were determined.

Some GE-supplied Class lE devices were qualified by analysis only. Analysis was used for passive mechanical devices and was sometimes used in combination with testing for larger assemblies containing Class lE devices. For instance, a test might have been run to determine if there were natural frequencies in the equipment within the critical loading frequency range. If the equipment was determined to be free of natural frequencies within the critical frequency range, then it was assumed to be rigid and a static analysis was performed as shown in Attachment A3.10. If it had natural frequencies in the critical frequency range, then calculations of transmissibility and responses to varying input accelerations were determined to see if Class CPS/USAR CHAPTER 03 3.10-6 REV. 11, JANUARY 2005 lE devices mounted in the assembly would operate without malfunctioning. A sample analysis is shown in Attachment B3.10. In general, the testing of Class lE equipment was accomplished using the procedure described in the following paragraph. Assemblies (i.e., control panels) containing devices which have had dynamic malfunction limits established were tested by mounting the assembly on a vibration machine in the field-mounted configuration as far as practical. Whenever exceptions to this were identified, additional justification was provided. A low-level resonance search was then conducted. As with the devices, the assemblies were tested in the three major orthogonal axes. The resonance search was run in the same manner as described previously for devices. If resonances were present, the transmissibility between the input and the location of each Class lE device was determined by measuring the accelerations at each device location and calculating the amplification between it and the input. From the transmissibilities the response at any Class lE device location for any given input was determined analytically. (It was conservatively assumed that the transmissibilities were linear as a function of acceleration even though they actually decrease as acceleration is increased.) If the device input accelerations were determined to be below their malfunction limits, then the assembly was considered a rigid body with a transmissibility equal to one so that a device mounted on it would be limited directly by the assembly input acceleration. There are basically three generic panel types. One or more of each type was tested using the above procedures.

Figures 3.10-1 through 3.10-3 illustrate the three basic panel types and show typical accelerometer locations. The results of the dynamic tests on the Class lE panels supplied by GE are presented in the qualification document package referenced in Nuclear Station Engineering Standard MS-02.00. The full acceleration level tests described above demonstrated that most of the panel types had more than adequate mechanical strength and that a given panel design acceptability was just a function of its amplification factor and the malfunction levels of the devices mounted in it.

Subsequent panels were, therefore, tested at lower acceleration levels and the transmissibilities measured to the various devices as described above. By dividing the malfunction levels of the devices by the panel transmissibility, between the device and the panel input, the panel seismic qualification level could be determined. Several high-level tests have been run on selected generic panel designs to assure conservatism in using the transmissibility analysis described. In cases where the supports or panels are not separable from the components being qualified, an integrated method of testing is performed with the component mounting fastened to the test table in a manner identical to the actual installation. Thus a qualification procedure in accordance with that outlined in Subsection 3.10.3.2.1 is used. 3.10.3.3 Design of Cable Trays, C able Tray Supports, and Conduit Supports 3.10.3.3.1 General All safety-related cable trays, their hangers, conduits, and their supports are designed to meet the requirements of Seismic Category I electrical components.

CPS/USAR CHAPTER 03 3.10-7 REV. 11, JANUARY 2005 3.10.3.3.2 Loads

a. Dead load D: includes dead weight of the cables, cable trays, conduits and self weight of the hanger. b. Live Load L: a live load of 200 pounds is considered for the design of cable trays and cable tray handers, for the construction loading case only. c. E: operating-basis earthquake or safe shutdown earthquake, whichever is larger. d. Safety-relief valve discharge load.

SRV ALL = SRV loading due to 16 (all) safety/relief valve discharge.

SRV 1V2P = SRV loading due to one safety/relief valve subsequent actuation.

SRV ADS = SRV loading due to seven (ADS) safety/relief valve discharge. e. LOCA dynamic response loads P

d. MVC = LOCA loading due to main vent clearing.

PS = LOCA loading due to pool swell. CO = LOCA loading due to condensation oscillation. CH = LOCA loading due to chugging.

AP = LOCA loading due to annulus pressurization and associated pipe breaks. 3.10.3.3.3 Load Combinations and Design Limits

a. Cable Tray Supports D 1.0 x allowable D* + L 1.33 x allowable but not to exceed 0.95 F y (D* tray weight only without cables)

++++++++++++++++++++PSEDAPEDMVCED COSRVED CHSRVED COSRVED CHSRVEDSRVED ADS ADSP2V1P2V1 ALL 1.6 x allowable, but a minimum factor of safety of 1.05 will be maintained against yield

CPS/USAR CHAPTER 03 3.10-8 REV. 11, JANUARY 2005 b. Conduit Supports

D 1.0 x allowable

+++++++++++++++++++++++++APSRVED PSSRVED MVCSRVED CHSRVED COSRVED COSRVED CHSRVED SRVEDSRVEDP2V1P2V1P2V1 ADS ADSP2V1P2V1 ALLP2V1 1.6 x allowable, but less than or equal to 0.95F y 3.10.3.3.4 Procedure for Analysis and Design The dynamic analysis and design of the cable tray hangers is performed using computer programs PIPSYS and SEISHANG. Both the progr ams utilize a response spectrum method of analysis. Different dynamic loads are combined by the square root of the sum of the squares method, with the exception of the condensation oscillation load, which is combined by absolute sum. The stresses and reactions from the different directional excitations are combined by the square root of the sum of the squares method. The equivalent static analysis of the cable trays is performed using the computer program SEISHANG and peak of the response spectrum.

Conduit supports are also designed using equivalent static approach and peak of the response spectrum. The PIPSYS program performs a multimode analysis. Stresses and reactions from all significant modes are combined using methods in compliance with NRC Regulatory Guide 1.92.

The SEISHANG program performs a single mode analysis. Seven percent damping was used

in the analysis for SSE. The slenderness ratio for compression members shall be as follows:

Member Type Maximum Slenderness

Ratio (k1/r) Compression members (verticals, diagonals and longitudinal braces) in floor and wall mounted supports (i.e., compression

system supports) 200 Compression members (verticals, diagonals and longitudinal braces) in ceiling mounted supports (i.e., tension system

supports) 300 A detailed discussion of these programs is given in Appendix C.

CPS/USAR CHAPTER 03 3.10-9 REV. 11, JANUARY 2005 3.10.3.3.5 Applicable Codes, Standards and Specifications

a. AISI "Specification for Design of Cold-formed Steel Structural Members," 1968 Edition and 1980 Edition. b. AISC "Specification for the Design Fabrication and Erection of Structural Steel for Buildings," (1969 or 1978). c. AWS D1.3, "Structural Welding Code - Sheet Steel," (1978 Edition). d. AWS D1.1, "Structural Welding Code - Steel." Clarifications and exceptions to AWS D1.1 and D1.3 are made based on engineering evaluations. 3.10.3.3.6 Instrument Tubing Supports The dynamic design and analysis of instrument tubing supports are in accordance with the simplified dynamic analysis discussed in Subsection 3.7.3.8.6.

CPS/USAR ATTACHMENT A3.10 SAMPLE SEISMIC STATIC ANALYSIS

CPS/USAR CHAPTER 03 A3.10-1 REV. 11, JANUARY 2005 ATTACHMENT A3.10 SAMPLE SEISMIC STATIC ANALYSIS PART I Part I presents a set of curves from which static seismic analysis of standard enclosures can be quickly performed. A standard enclosure is any enclosure listed in the Enclosure Standards Manual. The enclosures are assumed to be floor mounted, using all mounting holes with 5/8 inch steel bolts or studs each having an effective area of 0.2256 in2. Using an elastic limit of one half the ultimate strength, the bolts are assumed to have a maximum safe tension stress and maximum safe shear stress of 28,000 PSI and 21,000 PSI, respectively. The curves are based on a design basis earthquake having a horizontal acceleration of 1.5G and a vertical acceleration of 0.5G. It is assumed that each enclosure is mounted alone and not coupled directly to any other enclosure. The static analysis consists of determining the maximum allowable safe weight of the enclosure and its components for which the mounting bolt stresses are not exceeded. The curves of Figure A3.10-1 have been derived for this purpose. To use the curves given in Figure A3.10-1, first determine from Table A3.10-1 the curve designation of the enclosure being considered. Next, using the corresponding curve in Figure A3.10-1, determine the maximum safe weight per bolt for a given height of the center of gravity. The maximum safe enclosure weight is then determined by multiplying the weight per bolt by the total number of enclosure mounting bolts. Comparison with the actual weight of the enclosure and its components then indicates whether or not the mounting bolt stresses are exceeded. If the comparison shows that the maximum safe weight per bolt is exceeded, steps should be taken to increase the effective bolt area by welding the enclosure to its mounting, increasing the number of mounting bolts, adding top braces to a wall, or using another appropriate method to ensure safe operation during seismic disturbance.

CPS/USAR CHAPTER 03 A3.10-2 REV. 11, JANUARY 2005 TABLE A3.10-1 STANDARD ENCLOSURES Curve Enclosure Width Depth Mode of Failure C1 Instrument Rack 24" 24" Instrument Rack 24" 30" Vertical Board 24" 24" Side to Side Vertical Board 24" 30" Benchboard 24" 48" Benchhoard 24" 54" C2 Instrument Rack 30" 30" Instrument Rack 30" 24" Instrument Rack 48" 24" Instrument Rack 60" 24" Front to Back Instrument Rack 72" 24" OR Instrument Rack 96" 24" Back to Front Vertical Board 36" 24" Vertical Board 48" 24" Vertical Board 60" 24" Vertical Board 72" 24" Vertical Board 96" 24" C3 Instrument Rack 48" 30" Instrument Rack 60" 30" Instrument Rack 72" 30" Instrument Rack 96" 30" Vertical Board 36" 30" Front to Back Vertical Board 48" 30" OR Vertical Board 60" 30" Back to Front Vertical Board 72" 30" Vertical Board 96" 30" C4 Console 96" 42" Back to Front C5 Benchboard 48" 54" Side to Side Benchboard 48" 48" C6 Benchboard 72" 48" Benchboard 96" 48" Front to Back Console 96" 48" C7 Benchboard 72" 54" Back to Front Benchboard 96" 54" CPS/USAR CHAPTER 03 A3.10-3 REV. 11, JANUARY 2005 PART II Part II presents the necessary assumptions and equations for the calculation of the maximum normal and shear stresses in the mounting bolts of any enclosure under seismic disturbance.

The following assumptions and conventions are made: a. The enclosure under consideration is assumed to be a rigid body in equilibrium with respect to its mounting. b. The forces on the enclosure due to seismic accelerations are assumed to act through the enclosure's center of gravity. c. The enclosure is assumed to have a known weight W as well as a known center of gravity located at X, Y, Z with respect to a right-handed coordinate system. d. The right-handed coordinate system is arbitrarily assumed to be located at the front left-hand lower corner of the enclosure with the positive X-axis to the right along the front edge, the positive Y-axis toward the back of the enclosure, and the positive Z-axis toward the top of the enclosure. e. The stresses on the enclosure mounting bolts are assumed to be greatest when the horizontal component of the floor acceleration is perpendicular to a side of the enclosure and the vertical component of the acceleration is downward. f. It is assumed that the enclosure tends to rotate about an axis parallel to either the X-axis or the Y-axis, dependent upon the direction of the horizontal acceleration. The location of the axis of rotation is dependent upon the mounting configuration of the enclosure. g. There is assumed to be no friction between the enclosure and its mounting. h. The horizontal shear force due to the horizontal component of the acceleration is assumed to be distributed equally among the mounting bolts. i. All mounting bolts are assumed to be identical. The following procedure outlines the equations involved in determining the mounting bolt stresses. From the geometric configuration of the mounting bolts it is found that the tension forces in the bolts are related by

'F d d F j j i i= (1) where Fi and Fj are the tension forces acting on the i-th and the j-th bolts, respectively, and di and dj are the perpendicular distances of the i-th and the j-th bolts, respectively, from the axis about which the enclosure tends to rotate. When the enclosure is mounted directly to the floor, the axis of rotation will be an edge of the enclosure. For other mounting configuration, care must be exercised in determining the axis.

CPS/USAR CHAPTER 03 A3.10-4 REV. 11, JANUARY 2005 Summing moments about the enclosure's axis of rotation, the equation relating the unknown bolt tension forces to known quantities is found to be

[],L)12A(Z1AWdF...dFdFNN2211+*=+++ (2) where N is the number of mounting bolts, A1 and A2 are the relative magnitudes of the horizontal and vertical components of the floor acceleration, respectively, and L is the perpendicular distance between the line of action of the vertical acceleration through the center of gravity and the axis about which the enclosure tends to rotate. Substituting (1) into (2), the j-th tension force is

[]N 2 2 2 2 2 j jd...ddL)12A(Z1AWd F++++**= (3) The other tension forces are determined using Equation (1).

The tension stress T is related to the tension force by A F T= (4) Where A is the effective cross-sectional area of a mounting bolt.

Summing forces in the direction of the horizontal force acting upon the enclosure and making use of assumptions 7 and 8, the shear stress on the i-th bolt is AN1AW S**= (5) Due to the combined tension and shear stresses, the maximum tension stress, (Ti) , and the maximum shear stress, (S ) present in the i-th bolt are 2 i 2 iSi 2 T i 2 TmaxT++= (6) and 2 i 2 i)S(i 2 TmaxS+= (7) For a detailed derivation of Equations (6) and (7), the reader is directed to Strength of Materials, by Ferdinand L. Singer, Chapter 9, Section 6. To apply the above equations to determine the maximum tension and shear stresses, the following is required: Center of Gravity X, Y, Z Inches Horizontal Seismic Acceleration A1 - G Vertical Seismic Acceleration A2 - G CPS/USAR CHAPTER 03 A3.10-5 REV. 11, JANUARY 2005 Number of Bolts N Area Each Bolt A Square Inches Bolts distance from Axis of d 1 , d 2...d N Inches Rotation PROCEDURE: a. Determine the axis about which the cabinet tends to rotate for a given floor motion. b. Determine, using Equation (3), the tension force acting on the j-th mounting bolts (arbitrarily choose one). c. Determine the tension forces acting on the remaining mounting bolts from application of Equation (1). d. Calculate the tension stress acting on each bolt using Equation (4) and the results of Step 3. e. Calculate the horizontal shear stress from Equation (5). f. Determine the maximum tension stresses using Equation (6) and the results of Steps 4 and 5. g. Determine the maximum shear stresses using Equation (7) and the results of Steps 4 and 5. h. Compare these maximum stresses and allowable stresses of one half the ultimate strength (in PSI) for the bolt material.

CPS/USAR CHAPTER 03 A3.10-6 REV. 11, JANUARY 2005 PART III I. PURPOSE The purpose of Part III is to document a static seismic analysis which was performed to verify that the mounting bolts of the standard cabinets are capable of withstanding seismic environment.

II. SCOPE The scope of this report is limited to the static analysis of the mounting bolt stresses of five (5) standard cabinets. The standard cabinets are: a. Area Radiation Monitor, 236x400 (911) b. TIP Control, 236x401 (913) c. Start-up Neutron Monitor, 236x402 (936) d. Power Range Monitor, 236x403 (937)

e. Rod Position Information System, 236x404 (927)

III. DISCUSSION The Seismic Design Guide, 225A4582, was used in conducting the static seismic analysis. Each cabinet was assumed to be floor mounted using 5/8" bolts in all mounting holes. The maximum safe tension stress and maximum safe shear stress was assumed to have a horizontal acceleration of 1.5G and a vertical 1 acceleration of 0.5G. The weight of each cabinet was estimated using the weight of each major component listed in the parts lists for each cabinet. The height of the center of gravity of each cabinet was calculated using the weight and center of gravity of each of the major

components. The following data sheets include the necessary information for determining the factor of safety for each cabinet.

1 Equal to one-half the ultimate strength as given in Machinery's Handbook, Fourteenth Edition.

CPS/USAR CHAPTER 03 A3.10-7 REV. 11, JANUARY 2005 SEISMIC DESIGN VERIFICATION DATA SHEET Cabinet Name, Area Radiation Monitor (MPLH12P605) Applied Horizontal Acceleration 1.5 G Applied Vertical Acceleration 0.5 G Tension Stress (Maximum Safe) 28,000 PSI Weight of Cabinet 675 Lbs.

Number of Mounting Bolts 4 Height of Center of Gravity 48 Inches Maximum Allowable Weight Per Bolts 830 Lbs/Bolt (From Curve No. Cl on Page 8 of Seismic Design Guide, 225A4582)

Maximum Allowable Cabinet Weight 830 Lbs/Bolt

• 4 Bolts =

3,320 Lbs.

Maximum Allowable Weight Factor of Safety = = 4.9 Weight Cabinet Names: TIP Control, (H12P607) Applied Horizontal Acceleration 1.5 G Applied Vertical Acceleration 0.5 G Tension Stress (Maximum Safe) 28,000 PSI Shear Stress (Maximum Safe) 21,000 PSI Weight of Cabinet 755 Lbs. Number of Mounting Bolts 8 Height of Center of Gravity 50 Inches Maximum Allowable Weight Per Bolt 1,110 Lbs.

(From Curve No. C3 on Page 8 of Seismic Design Guide, 225A4582) Maximum Allowable Cabinet Weight 1,110 Lbs/Bolt

• 8 Bolts = 8,880 Lbs.

Maximum Allowable Weight Factor of Safety = =11.7 Weight Cabinet Name: Start-Up Neutron Monitor, (H12P633) Applied Horizontal Acceleration 1.5 G Applied Vertical Acceleration 0.5 G Tension Stress (Maximum Safe) 28,000 PSI Shear Stress (Maximum Safe) 21,000 PSI Weight of Cabinet 1,910 Lbs.

Number of Mounting Bolts 12 Height of Center of Gravity 50 Inches Maximum Allowable Weight Per Bolt 1,110 Lbs/Bolt (From Curve No. C3 on Page 8 of Seismic Design Guide, 225A4582)

CPS/USAR SEISMIC DESIGN VERIFICATION DATA SHEET (Continued) CHAPTER 03 A3.10-8 REV. 11, JANUARY 2005 Maximum Allowable Cabinet Weight 1,110 Lbs/Bolt

• 12 Bolts =

13,320 Lbs Maximum Allowable Weight Factor of Safety = = 11.9 Weight Cabinet Name: Power Range Monitor, (328x105) (H12P608) Applied Horizontal Acceleration 1.5 G Applied Vertical Acceleration 0.5 G Tension Stress (Maximum Safe) 28,000 PSI Shear Stress (Maximum Safe) 21,000 PSI Weight of Cabinet 4,345 Lbs.

Number of Mounting Bolts 40 Height of Center of Gravity 46 Inches Maximum Allowable Weight Per Bolt 1,210 Lbs/Bolt (From Curve No. C3 on Page 8 of Seismic Design Guide, 225A4582) Maximum Allowable Cabinet Weight 1,210 Lbs/Bolt

• 40 Bolts = 48,400 Lbs.

Maximum Allowable Weight Factor of Safety = 11.1 Weight Cabinet Name: Rod Position Information System, (H12P615) Applied Horizontal Acceleration 1.5 G Applied Vertical Acceleration 0.5 G Tension Stress (Maximum Safe) 28,000 PSI Shear Stress (Maximum Safe) 21,000 PSI Weight of Cabinet 2,500 Lbs.

Number of Mounting Bolts 20 Height of Center of Gravity 45 Inches Maximum Allowable Weight Per Bolt 1,225 Lbs/Bolt (From Curve No. C3 on Page 8 of Seismic Design Guide, 225A4582)

Maximum Allowable Cabinet Weight 1,225 Lbs/Bolt

• 20 Bolts = 24,500 Lbs.

Maximum Allowable Weight Factor of Safety = 9.8 Weight CPS/USAR CHAPTER 03 A3.10-9 REV. 11, JANUARY 2005 IV. CONCLUSION Review of the Factor of Safety of each standard cabinet indicates that the mounting bolts of each cabinet are capable of withstanding seismic disturbance as specified in the Seismic Design Guide.

CPS/USAR ATTACHMENT B3.10 SAMPLE PANEL FREQUENCY ANALYSIS

CPS/USAR CHAPTER 03 B3.10-1 REV. 11, JANUARY 2005 The method of analysis used to determine the resonant frequency of the panel is as follows: a. Calculate the moment of inertia of the corner post structure. b. First assume a simplified structure and calculate the frequency using the expression: k/w/13.3w/k)2/)g((w/kg2/1(f=== =/13.3f Where:

f = frequency g = 386 in./sec 2 k = spring rate #/in.

w = weight # = deflection = w/k weight distribution is assumed to be uniform. c. Additional structural components are added and the moment and frequency recalculated. The calculated resonant frequency of 7.4 Hz for the panel and 5.9 Hz for the benchboard was obtained using only the corner posts and the top. The addition of skin (3/8-in. steel) and 2-in. x 1/4-in. steel stiffeners will raise the frequency further. This proves that resonances cannot exist in the unstable region below 5 Hertz. FIRST APPROXIMATION For First Approximation lump the 4 corner posts together and assume the panel is a cantilever beam fixed on one end and uniformly loaded (see Figure B3.10-1).

The natural frequency is 2.6 Hz so we will have to use more of the structure.

SECOND APPROXIMATION For a second approximation, consider two 0.18" x 30" barriers in addition to the corner posts. The plan view of the panel is shown in Figure B3.10-2. In the X direction just one barrier will raise the frequency to 30 Hz. Use 4 inches of the back panel for each of the two barriers (see Figure B3.10-3) and the natural frequency in the Y direction becomes 4 Hz. The deflection equation used so far is very conservative: it assumes that the 4 corner posts are lumped together and that the structure can deflect like a simple cantilever beam. Actually the corners are separated by an angle frame which is stiffer than the corner posts. This will force the structure to deflect as shown in Figure B3.10-4.

CPS/USAR CHAPTER 03 B3.10-2 REV. 11, JANUARY 2005 In the simulated model we are not conservative (if we used all of the members) but we are very close. The reason we are not quite correct is because the stiff top frame will deflect slightly as shown below. The calculated frequency is 7.4 Hz which is above the necessary 5 Hz. The benchboard H11 P601 which weighs 4000 pounds, the calculated natural frequency is 5.9 Hz which is still above the 5 Hz test frequency minimum. NOTE: This neglects the barriers, the end and front panels, top plate, the stiffening of the lower part of the structure due to the bench board geometry, and all other members of the structure.

CPS/USAR CHAPTER 03 3.11-1 REV. 11, JANUARY 2005 3.11 ENVIRONMENTAL QUALIFICATION OF MECHANICAL AND ELECTRICAL EQUIPMENT 3.11.1 Introduction Environmental equipment qualification efforts for the Clinton Power Station Unit 1 began with the issuance of original equipment procurement specifications. These documents contained requirements to ensure General Design Criteria (10 CFR 50 Appendix A) 1, 2, 4, and 23 were satisfied and included IEEE 323 and 344. Since the issuance of IE Bulletin 79-01B, NUREG-0588 (Reference 2) and the Commission Memorandum and Order (CLI-80-21) of May 23, 1980 (Reference 1), an effort was initiated to compare the Clinton Environmental Qualification (CEQ) program against the requirements as stated in these documents. This included recalculation or verification of environmental parameters (radiation, temperature, pressure, humidity) to ensure consistency with guidelines contained in NUREG-0588, Rev. 1, Category 1 requirements. Also, a complete electrical systems analysis was performed to revalidate the listing of electrical equipment and components required to satisfy six safety goals for plant operation and

shutdown. These are: a. Safe shutdown b. Containment isolation

c. Core coverage

d. Residual heat removal

e. Containment integrity

f. Effluent control The objective of this study was to establish a comprehensive list of Class lE electrical equipment in harsh plant zones that require qualification to support these safety goals. The environmental qualification program consists of essentially three phases: a. The assessment and evaluation phase that was performed for equipment already qualified to specified environmental condition s prior to issuance of NUREG-0588. Furthermore, this appraisal verifies the actual basis of qualification for the equipment against the NUREG-0588 criteria. This phase is discussed further in Subsection 3.11.5. b. The ongoing qualification phase that uses the environmental parameters of Tables 3.11-5 through 3.11-8, 3.11-11 through 3.11-14, and 3.11-21 through 3.11-25, which are based on NUREG-0588. This phase is discussed further in

Subsection 3.11.6. c. The requalification phase includes the following options:

CPS/USAR CHAPTER 03 3.11-2 REV. 11, JANUARY 2005

1. Relocation

2. Reanalysis

3. Retest

4. Replacement of equipment that does not adequately qualify to NUREG-0588 requirements with a qualified replacement. This phase is discussed further in Subsection 3.11.7.

3.11.2 Definitions

a. Harsh environmental zone - An area in the plant where one or more environmental parameters change significantly due to postulated loss of coolant accident (LOCA) or high-energy line breaks (HELB's). b. Mild environmental zone - Those areas where gradual changes in environmental parameters may occur but these changes do not exceed the normal operating ranges specified for equipment therein. As an example, the loss of offsite electrical power (LOOP) could result in a loss of ventilating equipment and change the normal environment, resulting in a mild environment. c. Engineered safety feature (ESF) sy stems - These systems are provided to mitigate the consequences of design basis accidents and are discussed in detail

in Chapter 6. d. High energy lines - Those lines at pressures above 275 psig and/or temperature above 200° F. Refer to Subsection 3.6.1.1.1.b. 3.11.3 Safety Systems and Supporting Equipment In the physical layout of the CPS plant equipment, specific attention was given to the location of Class lE equipment in the containment building. As far as practical, redundant safety-related cooling is provided for essential control panels, auxiliary equipment panels, cable spreading areas, and essential switchgear. Furthermore, all Class lE electrical system logics, interlocks, controllers, indicators, recorders, relays, etc. that comprise the control circuitry are located in the mild environmental zones wherever feasible. The engineered safety feature systems are the first line systems required to achieve or maintain safe reactor shutdown, containment isolation, reactor core cooling, containment heat removal, core residual heat removal, and prevention of significant release of radioactive materials to the environment. The functional aspects of the systems that are required to support these safety functions are described separately in the appropriate sections of the USAR for each system and support equipment. All Class lE electrical equipment which is located in a harsh environment, as defined in the scope of 10 CFR 50.49 (Reference 12), is included in List 1 of Engineering Standard MS-02.00, Maintenance of Equipment Qualification Program Manual (Reference 13).

CPS/USAR CHAPTER 03 3.11-3 REV. 11, JANUARY 2005 3.11.4 NUREG-0588 Parameters Considered in Qualification Phase

a. Temperature - This parameter for each plant environmental zone is discussed in Subsection 3.11.9, and the discrete values for each zone are provided in Table 3.11-5. b. Radiation - This parameter for each plant environmental zone is discussed in Subsection 3.11.9, and the discrete values for each zone are provided in Table

3.11-5. c. Aging - Aging effects on equipment are considered in the qualification program. This includes electrical and mechanical cycling for equipment wherever appropriate. For most equipment, the Arrhenius methodology is used for determining the accelerated thermal aging requirements, and the aging acceleration rate is defined for each component in the specific environmental qualification program. Aging is addressed in greater detail in specific environmental qualification (EQ) binders prepared for the equipment and available in the Clinton Power Station Central File. d. Seismic and dynamic - For lE electrical equipment, the dynamic qualification is discussed in Section 3.10. e. Chemical environment - Since demineralized water is used for all safety systems, chemical spray is not a concern for Clinton Environmental Qualification.

f. Pressure - This parameter for each plant environmental zone is discussed in Subsection 3.11.9, and the discrete values for each zone are provided in Table

3.11-5. g. Humidity - This parameter for each plant environmental zone is discussed in Subsection 3.11.9, and the discrete values for each zone are provided in Table

3.11-5. h. Submergence - This parameter for each plant environmental zone is discussed in Subsection 3.11.9, and the discrete values for components are addressed in the applicable EQ binders.

i. Synergism - This effect has been addressed in the EQ binders for those materials affected by this phenomenon.

j. Dust - There are general administrative housekeeping procedures to maintain plant cleanliness at acceptable levels, including a maintenance and replacement schedule for HVAC filter units.

k. Margins - For equipment that is type tested, appropriate margins have been applied to the service conditions in conformance with Subsection 6.3.1.5 of IEEE 323 (Reference 3). This is addressed in more detail in the EQ binders. A one-hour margin has been applied when operability times are less than 10 hours1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br />. In cases where a deviation from this is taken for specific equipment, a justification is provided in the applicable EQ binders.

CPS/USAR CHAPTER 03 3.11-4 REV. 11, JANUARY 2005 l. Test sequence - For equipment that is type tested, the sequence of testing is in conformance with Subsection 6.3.2 of Reference 3 unless otherwise noted and justified in the detailed EQ binders. m. Periodic surveillance and maintenance - Qualified life is established based on the qualification of the most limiting material identified in the equipment. A maintenance and surveillance schedule is established for the equipment. This is addressed further in Subsection 3.11.11 and in the individual qualification

packages. n. Containment spray - This parameter for environmental zones which are subject to containment spray is discussed in Subsection 3.11.9.2.2 and is addressed further in the individual EQ binders as applicable. The operation of the containment spray system is addressed in Subsection 6.2.2. Environmental Zones H-1, H-15, H-26, H-37, and H-44 are subject to containment spray. 3.11.5 Assessment and Evaluation Phase The assessment phase includes an identification of all equipment important to safety in terms of its function during normal and abnormal environments. If the equipment is used in several systems throughout the plant, the most demanding safety function and the most severe environment is the condition used to assess its qualification. Each piece of BOP and NSSS lE electrical equipment is identified specifically by manufacturer, model number and type. Additionally, for all active equipment, the time required to perform its respective essential safety function is identified. In this phase of the work, the plant areas are divided into environmental zones that are determined by various plant events. The zone classifications are harsh and mild environmental zones, and are described in Subsection 3.11.9. The applicable normal, abnormal, and accident plant conditions for each harsh zone are shown in Table 3.11-5. The assessment and evaluation are based on the parameters identified or referenced in Table 3.11-5. Further, in this phase a detailed appraisal of equipment's capability to perform its safety function in an accident environment associated with LOCA or HELB was performed based on available qualification test reports. 3.11.5.1 BOP Assessment and Evaluation Phase For each unique piece of BOP Class lE electrical equipment located within a harsh environmental zone, an equipment qualification report was submitted by the equipment vendor (or testing laboratory) in accordance with the requirements of the equipment technical specifications. The assessment and evaluation of this equipment is based on the environmental parameters listed in Table 3.11-5. All BOP Class lE equipment is included in List 1 of MS-02.00. This list also provides equipment number, manufacturer, type/model number and EQ binder number.

CPS/USAR CHAPTER 03 3.11-5 REV. 11, JANUARY 2005 For all environmentally qualified electrical equipment listed in List 1 a detailed EQ checklist (which forms a part of the EQ binder) in compliance with NUREG-0588 is prepared and sent to the Clinton Power Station Central File (see Subsection 3.11.12). 3.11.5.2 NSSS Assessment and Evaluation Phase All NSSS Class lE equipment is included in List 1 of MS-02.00. This list also provides equipment number, manufacturer, model number and EQ binder number.

For each unique piece of NSSS Class lE electrical equipment located within a harsh zone, an equipment qualification report was submitted by the equipment vendor (or testing laboratory) in accordance with the requirements of the equipment technical specifications. For all environmentally qualified electrical equipment listed in List 1, a detailed EQ checklist (which forms a part of the EQ binder) in compliance with NUREG-0588 is prepared and sent to the Clinton Power Station Central File (see Subsection 3.11.12). 3.11.6 Status of Ongoing Qualification Efforts The ongoing qualification phase uses the environmental zone parameters of Table 3.11-5 for both BOP and NSSS equipment qualification. 3.11.7 Requalification Phase 3.11.7.1 Equipment in Harsh Environments Instances in which Class lE equipment did not meet the requirements of NUREG-0588 were resolved by implementing one of the following options: a. Relocation of the equipment from a "harsh" to a less severe "harsh" or a "mild" environment provided the system logic or other design functions were not affected. b. Replacement of the equipment with upgraded equipment. c. Protection of the equipment such that a more realistic set of environmental conditions could be achieved for the specific location. d. Retesting the equipment using appropriate environmental parameters based on its unique location. 3.11.7.2 Equipment in Mild Environments Since equipment in mild environments is not exposed to extreme environmental parameters during or following a design basis event, requalification to levels described for harsh environment equipment is not required per Reference 12. 3.11.8 Electrical Equipment Tabulation and Format List 1 of MS-02.00 includes the Class lE equipment for all environmental zones in alphabetical order. The information presented in this list for electrical equipment is as follows:

CPS/USAR CHAPTER 03 3.11-6 REV. 12, JANUARY 2007 a. Equipment number: Provides the specific plant numbers of the equipment (for ease of reference and correlation with other documents and drawings) and the generic name (type) of the equipment. b. Equipment manufacturer: Identifies the manufacturer of the equipment. c. Type/Model number: Provides the equipment catalog number. d. EQ binder number: Provides a reference to the environmental qualification binder for the equipment. 3.11.9 Plant Environmental Zones The plant areas containing Class lE equipment are divided into two zones based on the environmental conditions that are expected to occur as a result of various plant events. These zone classifications are termed harsh and mild environmental zones. The mild environmental zone is discussed in Subsection 3.11.10. The harsh environmental zones are discussed in the subsections below.

The environmental parameters for the abnormal and normal service conditions represent conservative selections chosen to bound the real conditions that may occur in these zones. A more refined or more detailed analysis has been performed for specific equipment to establish more realistic and representative environmental parameters than the values specified for the specific environmental zone. In those cases where unique calculations are prepared, they have been made part of the environmental qualification and part of the environmental records. 3.11.9.1 Harsh Environmental Zones Due to LOCA, HELB or LOOP The definition of the environmental conditions within the harsh environmental zones is based on results of analyses of postulated accidents. The postulated accidents are loss-of-coolant (LOCA), high energy line break (HELB), and a loss-of-offsite power (LOOP).

The following sections summarize the basis for the definition of the harsh environmental zones. LOCA, HELB and LOOP were investigated and the bounding conditions presented. Where possible, several plant areas have been grouped into a single zone with an H-x designation that bounds the environmental conditions in each of the individual areas. The various environmental zones are represented on plant general arrangement drawings (Drawing M01-1600 Sheets 6 through 21). The environmental conditions for each zone pertaining to pressure, temperature, relative humidity, duration and submergence are defined in Subsections 3.11.9.3 through 3.11.9.59 and Table 3.11-5. The "Submergence or Spray" section applies to the containment and drywell for design considerations resulting from suppression pool dynamic events such as suppression pool swell or weir swell, and from containment spray. The radiation environment is defined in Subsection 3.11.9.2 and Table 3.11-5. Flood protection for the Clinton Power Station is discussed in Subsection D3.6.4.

CPS/USAR CHAPTER 03 3.11-7 REV. 12, JANUARY 2007 3.11.9.2 Radiation and Containment Spray 3.11.9.2.1 Radiation The design basis accidents addressed in the determination of the radiation environment are the loss of coolant accident (LOCA), the fuel handling accident (FHA) and the high energy line break accident (HELB). The LOCA produces the most severe radiation environment, and as such is used as the design basis accident where it is applicable. The environmental conditions produced by the HELB accident last for a relatively short time, such that the associated radiation environment, which is expressed in terms of total integrated radiation dose, is not significant.

The FHA also produces less severe radiation environment compared to LOCA, but is used as the design basis accident for those components which are required to survive a FHA, but not the LOCA. The radiation environment after a design-basis accident is determined based upon the assumptions provided in NUREG-0588. The source terms are calculated using the reactor data provided in Table 12.2-1 and the RACER Code (Reference 2 of Subsection 12.3.5). It is assumed that the containment leaks into the gas-control boundary at the design-basis leak rate, as specified in Subsection 6.2.6, and that the gas-control boundary is evacuated by the standby gas treatment system (SGTS). The exhaust rate and filter efficiencies of the SGTS are listed in Subsection 6.5.1. The effects of post LOCA recirculation fluids have been included in the determination of radiation environment in accordance with NUREG-0737, Section II.B.2. A harsh environment attributed to radiation is defined as an environment with a total integrated dose value greater than 1 x 10 4 rad (C), in which the major dose contribution is from a postaccident condition. The Radiation Qualification Dose values to be utilized for testing of equipment in each environmental zone are listed in Table 3.11-5. The dose values are the sum of gamma radiation dose values and beta radiation dose values. The conservatively calculated total dose value for each zone is less than, or equal to, the value specified as the radiation qualification dose. The calculated dose is the sum of the integrated dose for 40 years of normal operation plus one year of radiation exposure in a postaccident condition. 3.11.9.2.2 Containment Spray Equipment which is located in the containment and outside of the drywell, except equipment located in cubicles, will be environmentally qualified for the containment spray requirements. The following spray test requirements are used to simulate the containment spray system: Initiation of spray - 10 minutes after LOCA Spray duration - one hour Spray loading - one gpm per square foot Water chemistry - demineralized water 3.11.9.3 Environmental Zone H-1 Zone H-1 is the suppression pool in the containment building. This zone is identified as area

C.1.1 in Drawing M01-1600-6. For this zone, the bounding environmental conditions result from postulated design-basis LOCA events or from the normal and abnormal operation of the main steam safety/relief valves. The conditions arising from safety/relief valve operation are not CPS/USAR CHAPTER 03 3.11-8 REV. 12, JANUARY 2007 discussed here. For a full discussion of these phenomena, including the short-term LOCA pool swell, condensation oscillation, and chugging, refer to Sections A3.8 and A3.9.

a. Pressure The determination of the design pressure is discussed in Section 6.2. The pressure range is given in Table 3.11-5.

b. Temperature The temperature profile is given in Figures 6.2-3 (Curve 2) or 6.2-12 (Curve 2) and 6.2-7a (Curve 3) and 6.2-7b (Curve 3). The determination of the design temperature is discussed in Section 6.2. c. Relative humidity Since this zone is normally flooded, the relative humidity for this zone cannot be defined. d. Duration The time histories for the spectrum of LOCA events are given in Section 6.2 and Figures 6.2-2 (wetwell), 6.2-3 (Curve 2), 6.2-6a (Curve 2), 6.2-6b (Curve 2), 6.2-7a (Curve 3), 6.2-7b (Curve 3), 6.2-11 (wetwell), and 6.2-12 (Curve 2). The temperature and pressure at the end of these curves is conservatively assumed to persist up to 100 days. e. Submergence or Spray This zone is flooded under normal and accident conditions. 3.11.9.4 Environmental Zone H-2 Zone H-2 is the lower elevation of the drywell. This zone is denoted as area C.1.2 in Drawing M01-1600-6. The bounding environmental conditions in this zone result from design-basis LOCA events.

a. Pressure The pressure environment for this zone is defined as the design basis pressure transient for the drywell. The determination of this design basis is discussed in Section 6.2. Table 3.11-6 shows the envelope of the LOCA pressure conditions addressed in Section 6.2.

b. Temperature The temperature environment for this zone is defined as the design basis temperature transient for the drywell. The determination of the design basis is discussed in Section 6.2, Table 3.11-6 shows the envelope of the LOCA temperature conditions addressed in Section 6.2.

CPS/USAR CHAPTER 03 3.11-9 REV. 12, JANUARY 2007 c. Relative humidity Table 3.11-6 gives the relative humidity envelope as a function of time for the environmental qualification of equipment. The environment is conservatively assumed to be all steam for the first 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> following an accident and at 100%

relative humidity thereafter.

d. Duration The time histories for the spectrum of LOCA events are shown in Section 6.2 and Figures 6.2-2 (drywell), 6.2-3 (Curve 1), 6.2-6a (Curve 1), 6.2-6b (Curve 1), 6.2-7a (Curve 1), 6.2-7b (Curve 1), 6.2-11 (drywell), and 6.2-12 (Curve 1). The time history given in Table 3.11-6 envelops all these events plus the small HELB. e. Submergence or Spray This zone is normally dry and is provided with sumps and drains to maintain this condition during normal operation. Following a design basis LOCA, the drywell depressurization will cause water to be drawn into the drywell through the LOCA vents. The nature of this transient is such that the drywell could be flooded to the top of the weir wall (see Reference 10). 3.11.9.5 Environmental Zone H-3 Zone H-3 is the portion of the drywell inside the pedestal under the RPV. This zone is designated area C.1.3 in Drawing M01-1600-6. The bounding environmental conditions for this zone result from design-basis LOCA events.

a. Pressure The pressure environment for this zone is defined by the spectrum of LOCA events discussed in Section 6.2. The large HELB and the small HELB are, in this environmental zone, a subset of the spectrum of LOCA events. The bounding pressure transient in this environmental zone is a composite of the large HELB and small HELB pressure transients given in Table 3.11-6.

b. Temperature The temperature environment for this zone is defined by the spectrum of LOCA events discussed in Section 6.2. The bounding temperature transient is given in Table 3.11-6 as the small HELB conditions. The small HELB in Table 3.11-6 is a subset of the spectrum of LOCA events and its temperature transient bounds that of all other line breaks in this environmental zone. c. Relative humidity The relative humidity for this zone is conservatively assumed to be 100% except for the time period immediately following the accident when the atmosphere is considered to be all steam. See Table 3.11-6.

CPS/USAR CHAPTER 03 3.11-10 REV. 12, JANUARY 2007

d. Duration The time histories for the spectrum of LOCA events are enveloped by the time histories presented in Table 3.11-6. These time histories are the results of GE analysis per Reference 7 and enveloped per Reference 8. e. Submergence or Spray Same as Section e for Zone H-2. 3.11.9.6 Environmental Zone H-4 Zone H-4 is an area of the fuel building. This zone is identified as area F.2.1 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a LOCA.

a. Pressure The bounding pressure environment for this zone is defined by the design parameters for the standby gas tr eatment system and the secondary containment. The accident condition analysis is discussed in Subsection 6.2.3 and Table 3.11-14 gives the envelope for the pressure conditions following an

accident.

b. Temperature The bounding temperature environment for this zone is defined by the design parameters for the standby gas tr eatment system and the secondary containment. The accident condition analysis is discussed in Subsection 6.2.3 and Table 3.11-14 gives the envelope for the temperature conditions following an

accident. c. Relative humidity The relative humidity following a LOCA does not exceed 96% is given in Reference 16.

d. Duration The envelope of the transient events is given in Table 3.11-14. This envelope is developed from the analysis discussed in Subsection 6.2.3. 3.11.9.7 Environmental Zone H-5 This zone is composed of the majority of the fuel building. The zone is identified as areas F.1.1, F.1.2, F.1.3, F.1.4, F.1.5, F.1.6, F.1.9, F.1.11, F.2.2, F.2.3, F.2.4, F.2.5, F.2.6, F.2.7, and F.3.1 in Drawing M01-1600 Sheets 6 through 9. The bounding environmental conditions result from a LOCA. a. Pressure Same as Zone H-4.

CPS/USAR CHAPTER 03 3.11-11 REV. 12, JANUARY 2007

b. Temperature Same as Zone H-4. c. Relative humidity Same as Zone H-4.

d. Duration Same as Zone H-4. 3.11.9.8 Environmental Zone H-6 Zone H-6 is the HPCS pump cubicle in the fuel building. This zone is denoted as area F.1.7 in Drawing M01-1600-6. For equipment qualification purposes environmental conditions are based on the bounding environmental conditions as discussed below.

a. Pressure The pressure environment is defined by Reference 7 which presents a Mark III generic analysis of the accident transients and the envelope of the analysis of Reference 8. This envelope bounds the results of analysis discussed in Subsection 6.2.3. Table 3.11-7 presents the envelope for the LOCA events and Table 3.11-8 presents the envelope for the HELB accidents.

b. Temperature The temperature environment is defined by the Mark III generic analysis performed by GE (Reference 7) and the envelope of the analysis of Reference 8.

Table 3.11-7 presents the envelope for the LOCA events and Table 3.11-8 presents the envelope for HELB accidents. c. Relative humidity Table 3.11-7 shows the relative humidity envelope determined by the generic Mark III analysis. Table 3.11-8 shows the relative humidity envelope for the postulated HELB events.

d. Duration The time histories for the LOCA and HELB events are given in Tables 3.11-7 and

3.11-8. 3.11.9.9 Environmental Zone H-7 Zone H-7 is composed of the fuel building floor drain tank cubicles, the fuel building tunnel, and the fuel pool heat exchanger cubicle. These areas are identified as F.1.8, F.1.10, F.1.12, and F.2.8 in Drawing M01-1600 Sheets 6 and 7. The bounding environmental conditions in these areas result from a LOCA.

CPS/USAR CHAPTER 03 3.11-12 REV. 12, JANUARY 2007

a. Pressure Same as Zone H-4.

b. Temperature Same as Zone H-4. c. Relative humidity Same as Zone H-4.

d. Duration Same as Zone H-4. 3.11.9.10 Environmental Zone H-8 Zone H-8 is composed of general areas of the auxiliary building. These areas are identified as A.1.11, A.1.13 and A.1.14 in Drawing M01-1600-6. The bounding environmental conditions for this area result from a LOOP.

a. Pressure The areas in this zone are outside the secondary containment and thus are not subject to any pressure effects due to a LOCA. The pressure variations which would occur in this area during a LOOP transient are given in Table 3.11-23.

b. Temperature The areas in this zone are outside the secondary containment and thus are not subject to the temperature effects due to a LOCA. The bounding temperature is determined as the result of a postulated LOOP. The extreme expected temperatures are cited in Table 3.11-23. c. Relative humidity The extreme values of relative humidity for this area are given in Table 3.11-23.

d. Duration The only transients of significance relate to the temperature. The transient is assumed to last for the duration of the LOOP. 3.11.9.11 Environmental Zone H-9 Zone H-9 is composed of the access aisle and floor drain pump cubicle denoted areas A.1.1 and A.1.2 in Drawing M01-1600-6. The bounding environmental conditions for this area result from a HELB in the pipe chase or a LOCA.

CPS/USAR CHAPTER 03 3.11-13 REV. 12, JANUARY 2007

a. Pressure The pressure conditions resulting from a LOCA are mitigated by the standby gas treatment system and held below 1 inch of water as discussed in Subsection

6.2.3. The pressure conditions resulting from a HELB (in this area or in adjacent areas) or LOCA are given in Table 3.11-8.

b. Temperature The temperature conditions resulting from a LOCA are determined in Subsection 6.2.3 and the values are given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve

8). The temperature conditions resulting from a HELB in these areas or in adjacent areas are tabulated in Table 3.11-8. c. Relative humidity The relative humidity for this zone is given in Table 3.11-8 for the spectrum of

breaks. d. Duration The LOCA pressure transient is inconsequential. The LOCA temperature transient is given as the "ECCS equipment rooms" and the "RHR-C pump room" shown in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 8), respectively. The temperature at the end of the curve shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. The HELB pressure conditions are of minor importance. The peak values given in Table 3.11-8 are assumed to persist for the duration of the accident. 3.11.9.12 Environmental Zone H-10 Zone H-10 is the floor drain tank cubicle denoted as area A.1.3 in Drawing M01-1600-6. The bounding environmental conditions in this zone result from a HELB in the adjacent pipe chase or a LOCA.

a. Pressure Same as Zone H-9.

b. Temperature Same as Zone H-9. c. Relative humidity Same as Zone H-9.

CPS/USAR CHAPTER 03 3.11-14 REV. 12, JANUARY 2007

d. Duration Same as Zone H-9. 3.11.9.13 Environmental Zone H-11 Zone H-11 is composed of the RHR-C and LPCS pump cubicles designated areas A.1.4 and A.1.10 in Drawing M01-1600-6. The bounding environmental conditions for this zone are determined by a LOCA or by a HELB in the individual cubicles.

a. Pressure The pressure conditions from a LOCA are mitigated by the standby gas treatment system and held below 1 inch of water as discussed in Subsection

6.2.3. The pressure transients following a HELB (in the adjacent cubicles) or LOCA are detailed in Table 3.11-8.

b. Temperature The temperature transients following HELB's in adjacent areas are given in Table

3.11-8. The temperature transient following a LOCA is determined in Subsection 6.2.3 and the values are given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4 and Curve 8). c. Relative humidity The range of relative humidities expected following the various accidents is given in Table 3.11-8.

d. Duration The LOCA pressure transient is inconsequential. The LOCA temperature transient is given as the "ECCS equipment rooms" shown in Figure 6.2-138 (Curve 2) and "RHR-C pump room" and "LPCS pump room" shown in Figure 6.2-142 (Curve 4 and Curve 8). The temperature at the end of the curve shown in Figure 6.2-142 is conservatively assumed to persist for up to 100 days. The pressure and temperature time-history of a HELB in the adjacent cubicles is given in Table 3.11-8. The pressures cited are assumed to persist for the duration of the transient. 3.11.9.14 Environmental Zone H-12 Zone H-12 is composed of the RHR-A, RHR-B and RCIC pump and heat exchanger cubicles shown as areas A.1.6, A.1.7 and A.1.9 in Drawing M01-1600-6 and areas A.1.6 and A.1.9 in Drawing M01-1600 Sheets 7 through 9. The bounding environmental parameters are determined by a LOCA or a HELB in the individual cubicles or the adjacent RHR-C pump cubicle.

CPS/USAR CHAPTER 03 3.11-15 REV. 12, JANUARY 2007

a. Pressure The pressure transients for the LOCA, HELB in an adjoining cubicle, and HELB in the adjacent RHRC pump cubicle are shown in Table 3.11-8. The peak conditions for the break in the individual cubicles is determined by the analysis of Reference 9. The remainder of the pressure transients are from the analysis in References 7 and 8.

b. Temperature The temperature transients following HELB's in adjacent areas are given in Table

3.11-8. The temperature transient following a LOCA is determined in Subsection 6.2.3 and the values are given in Table 3.11-8. c. Relative humidity The relative humidity for the HELB in the individual cubicles is not considered, rather an all steam environment is specified for conservativeness. Table 3.11-8 presents the bounding relative humidity for the other transients discussed in Section a above.

d. Duration The peak values determined for the HELB in the adjacent cubicles are specified in Table 3.11-8 are conservatively assumed to persist for the duration of the accident. The time-histories for the remainder of the transients are also shown in Table 3.11-8. 3.11.9.15 Environmental Zone H-13 Zone H-13 is the RCIC instrument panel room shown as area A.1.8 in Drawing M01-1600-6. The bounding environmental conditions for this area result from the LOCA, a HELB in the RCIC pump cubicle, or a HELB in the RHR-A, RHR-B or RHR-C pump cubicles.

a. Pressure The pressure transients for the LOCA, HELB in the RCIC cubicles, HELB in an adjoining cubicle, and HELB in the adjacent RHR-C pump cubicles are shown in Table 3.11-8. The peak conditions for the break in the individual cubicles are determined by the analysis of Reference 9. The remainder of the pressure transients are from the analysis in References 7 and 8.

b. Temperature Same as Zone H-12 for area A.1.7. c. Relative humidity Same as Zone H-12.

CPS/USAR CHAPTER 03 3.11-16 REV. 12, JANUARY 2007

d. Duration Same as Zone H-12. 3.11.9.16 Environmental Zone H-14 Zone H-14 is comprised of stairways and general areas of the auxiliary building that are denoted as areas A.2.11, A.2.20, A.2.22, and A.2.23 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a simultaneous LOCA and LOOP.

a. Pressure The areas under consideration in this zone are outside of the secondary gas control boundary and thus do not experience any pressure effects from a LOCA. Also since no high energy lines are located in these areas the pressure transient is inconsequential. (See Table 3.11-23.)

b. Temperature The peak temperatures in this area will occur following the simultaneous LOCA and LOOP. The peak calculated temperatures for these zones are listed in Table

3.11-23. c. Relative humidity The range of relative humidity following a LOOP is given in Table 3.11-23.

d. Duration The pressure consequences of a LOCA or LOOP are inconsequential. LOCA temperature conditions do not have a direct effect on the areas in this zone.

Post-LOCA boundary conditions are used in computing the SGTS and secondary containment area bounding environmental temperatures. The transient is assumed to last for the duration of the LOCA. 3.11.9.17 Environmental Zone H-15 Zone H-15 is the wetwell area above the suppression pool in the containment. This zone is identified as area C.2.1 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a spectrum of LOCA events consisting of a large HELB or a small HELB as described in Table 3.11-11.

a. Pressure The pressure environment following a LOCA is based on the GE Mark III generic accident analysis of Reference 7. The CPS-unique analyses are discussed in Section 6.2. The generic analyses as enveloped in Reference 8 bound the CPS-unique analysis. The results of the generic analyses are given in Table 3.11-11.

CPS/USAR CHAPTER 03 3.11-17 REV. 12, JANUARY 2007

b. Temperature The temperature environment following a LOCA is based on the GE Mark III generic accident analysis of Reference 7. The CPS-unique analyses are discussed in Section 6.2. The generic analyses as enveloped in Reference 8 bound the CPS-unique analysis. The results of the generic analyses are given in Table 3.11-11. c. Relative humidity The relative humidity environment following a LOCA is based on the GE Mark III generic accident analysis of Reference 7. The CPS-unique analyses are discussed in Section 6.2. The generic analyses as enveloped in Reference 8 bound the CPS unique analysis. The results of the generic analyses are given in Table 3.11-11.

d. Duration Table 3.11-11 gives the time-histories following design basis LOCA events. e. Submergence or Spray Equipment in this area may be submerged or exposed to a spray from pool swell or operation or the containment spray. The pool swell phenomenon is described in Section A3.8. 3.11.9.18 Environmental Zone H-16 Zone H-16 is the drywell. This zone is denoted as area C.2.2 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from design-basis LOCA events.

a. Pressure Same as Zone H-2.

b. Temperature Same as Zone H-2. c. Relative humidity Same as Zone H-2.

d. Duration Same as Zone H-2. e. Submergence or Spray The portion of the drywell below floor Elevation 740 feet 9 inches will be subject to the flooding and/or spray effects of weir swell. The zone of influence of the CPS/USAR CHAPTER 03 3.11-18 REV. 12, JANUARY 2007 weir swell phenomena is shown in Reference 10. 3.11.9.19 Environmental Zone H-17 Zone H-17 is composed of auxiliary building access areas A.2.1, A.3.1, A.3.5, A.3.8, A.3.9, A.4.1, A.4.5, A.4.10, and A.4.11 in Drawing M01-1600 Sheets 7 through 9.. The environmental conditions for this zone result from a LOCA or a HELB in ECCS pump cubicles.

a. Pressure The pressure transients resulting from a LOCA are mitigated by the standby gas treatment system and held below 1 inch of water as discussed in Subsection 6.2.3. The pressure transient resulting from a HELB is documented in Reference

16. The peak pressures are shown in Table 3.11-5.

b. Temperature The temperature transient during a LOCA is not specifically evaluated as a part of the analysis discussed in Subsection 6.2.3. The temperature transient is chosen conservatively to be that of the LPCS pump cubicle as shown in Figures

6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature following a HELB in the adjacent ECCS pump cubicles is shown in Table 3.11-5. c. Relative humidity The relative humidity is estimated conservatively to be 100% for all accident conditions.

d. Duration The LOCA pressure transient is inconsequential. The LOCA temperature transients are taken conservatively to be that of the LPCS pump cubicle as given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature at the end of Curve 4 shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. The conditions resulting from a HELB in adjacent ECCS pump cubicles are taken conservatively to be the values stated in Table 3.11-5 for the duration of the transient. 3.11.9.20 Environmental Zone H-18 Zone H-18 is the RCIC pipe tunnel that is designated as area A.2.3 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a LOCA or a HELB in the main steam tunnel in the auxiliary building.

CPS/USAR CHAPTER 03 3.11-19 REV. 12, JANUARY 2007

a. Pressure The pressure transient for the LOCA is mitigated by the standby gas treatment system and held below 1 inch of water as discussed in Subsection 6.2.3. The pressure transient for a HELB in the adjacent main steam tunnel is determined in Reference 11. The peak pressure determined by this analysis, 8.2 psig, is assumed to persist for the duration of the transient.

b. Temperature The temperature transient for the LOCA conditions is calculated as discussed in

Subsection 6.2.3. The temperature for the HELB condition is shown in the bounding temperature curves provided in Reference 16. The temperature transient following a LOCA is controlled by the operation of the SGTS and is given in Figures 6.2-138 (Curve 3) and 6.2-141 (Curve 1). c. Relative humidity For the LOCA condition the relative humidity is assumed conservatively to be 100%. For the HELB condition the environment is considered to be all steam.

d. Duration The pressure transient for the LOCA condition is inconsequential. The temperature transient is given in Figures 6.2-138 (Curve 3) and 6.2-141 (Curve 1). The temperature at the end of Curve 1 shown in Figure 6.2-141 is conservatively assumed to persist up to 100 days.

The temperature time-history for the HELB condition is analyzed or extrapolated for 100 days. 3.11.9.21 Environmental Zone H-19 This zone is the personnel hatch access area denoted as area A.2.5 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a HELB in the adjacent RWCU pump cubicles or a LOCA.

a. Pressure The pressure transient resulting from a LOCA is mitigated by the standby gas treatment system and the peak pressure is held below 1 inch of water as discussed in Subsection 6.2.3. The pressure transients resulting from a HELB in the adjacent RWCU pump cubicles are determined from Reference 9. The peak pressure value appears in Table 3.11-5.

CPS/USAR CHAPTER 03 3.11-20 REV. 12, JANUARY 2007

b. Temperature The temperature transient during a LOCA is not specifically evaluated as part of the analysis discussed in Subsection 6.2.3. The temperature is conservatively chosen to be that of the adjacent LPCS pump cubicle as shown in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature following a HELB in the adjacent RWCU pump cubicles is provided in Reference 16. c. Relative humidity The relative humidity is estimated conservatively to be 100% for all accident conditions.

d. Duration The LOCA pressure transient is inconsequential. The LOCA temperature transient is taken conservatively to be that of the LPCS pump cubicle as given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature at the end of Curve 4 shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. The conditions resulting from a HELB in the adjacent RWCU pump cubicles are taken conservatively to be the values stated in Table 3.11-5 for the duration of the transient. 3.11.9.22 Environmental Zone H-20 Zone H-20 is the auxiliary building pipe tunnel designated area A.2.6 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from a RWCU line break in the tunnel or from a LOCA.

a. Pressure The pressure transient following a LOCA is mitigated by the standby gas treatment system and held below 1 inch of water as discussed in Section 6.2.3. The pressure transient following a HELB in the tunnel is taken from Reference 9. The peak pressure value is stated in Table 3.11-5.

b. Temperature The temperature transient following a LOCA is not specifically evaluated.

Conservatively, the temperature transient in the LPCS pump cubicle is assumed. This transient is given in figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature following a HELB is provided in Reference 16.

CPS/USAR CHAPTER 03 3.11-21 REV. 12, JANUARY 2007 c. Relative humidity The relative humidity is assumed conservatively to be 100% following a LOCA. The environment is considered as all steam following the HELB.

d. Duration The pressure transient following a LOCA is inconsequential as discussed in Subsection 6.2.3. The temperature time-history is assumed conservatively to be that of the LPCS cubicle as shown in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature at the end of Curve 4 shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. The peak values for a HELB stated in Table 3.11-5 are assumed to persist for the duration of the accident. 3.11.9.23 Environmental Zone H-21 Zone H-21 is composed of storage rooms and a stairway designated areas A.2.8, A.2.15, and A.2.21 in Drawing M01-1600-7. The bounding environmental conditions result from a LOOP.

a. Pressure The pressure transient following a LOOP is provided in Table 3.11-23.

b. Temperature The temperature transients following a LOCA or LOOP are provided in

Reference 16. c. Relative humidity The relative humidity following a LOOP is provided in Table 3.11-23.

d. Duration Time-histories for the LOOP temperatures are presented in Reference 16. 3.11.9.24 Environmental Zone H-22 Zone H-22 is composed of the air locks designated as areas A.2.9 and A.2.19 in Drawing M01-1600-7. The bounding environmental conditions result from a HELB in adjacent RHR or RWCU pump cubicles or from a LOCA.

a. Pressure The pressure transient following a LOCA is mitigated by the standby gas treatment system as discussed in Subsection 6.2.3. The pressure transient following a HELB results in the peak pressures stated in Table 3.11-5. Those pressures are based on the analysis given in Reference 9.

CPS/USAR CHAPTER 03 3.11-22 REV. 12, JANUARY 2007

b. Temperature The temperature transient following a LOCA is not specifically evaluated for these areas. The temperature transients are assumed conservatively to be that of the adjacent ECCS pump cubicles as determined in Subsection 6.2.3. The temperature following a HELB is shown in Reference 16. c. Relative humidity The relative humidity is assumed conservatively to be 100% following any accident.

d. Duration The pressure transient following a LOCA is inconsequential. The temperature transient is assumed conservatively to be that of the adjacent ECCS pump cubicles as given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature at the end of Curve 4 shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. 3.11.9.25 Environmental Zone H-23 Zone H-23 is the MSIV room access area designated as A.2.12 in Drawing M01-1600-7. The bounding environmental conditions in this zone result from a LOCA or a HELB in the adjacent

RCIC pipe tunnel.

a. Pressure The pressure transient following a LOCA is mitigated by the standby gas treatment system. The peak pressure is held below 1 inch of water as stated in

Subsection 6.2.3. The peak pressure following a HELB in the adjacent RCIC pipe tunnel is stated in Table 3.11-5. This pressure is determined from the analysis of Reference 9.

b. Temperature The temperature transient following a LOCA is discussed in Subsection 6.2.3.

Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4) present the time-dependent

temperature profiles. The temperature following a HELB in the adjacent RCIC cubicle is shown in Reference 16. c. Relative humidity The relative humidity following any accident is assumed to be 100%.

CPS/USAR CHAPTER 03 3.11-23 REV. 12, JANUARY 2007

d. Duration The pressure transient following a LOCA is inconsequential. The temperature transient is given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperature at the end of Curve 4 shown in Figure 6.2-142 is conservatively assumed to persist up to 100 days. The accident conditions stated in Table 3.11-5 are presumed to persist for the duration of the HELB accident. 3.11.9.26 Environmental Zone H-24 Zone H-24 is composed of the MSIV cubicles shown as areas A.2.13 and A.2.14 in Drawing M01-1600-7. The bounding environmental conditions for these cubicles result from a LOCA, LOOP, Loss of HVAC, or a HELB in the adjacent RCIC pipe tunnel.

a. Pressure Same as Zone H-23 during LOCA and HELB. The value for LOOP is the same as that for LOCA. The zone pressure is within its normal pressure range during a Loss-of-HVAC (see Table 3.11-5).

b. Temperature Same as Zone H-23 during LOCA and HELB. The value for LOOP is the same as that for LOCA. The peak temperatures for the cubicles in Zone H-24 are

given in Table 3.11-5. c. Relative Humidity Same as Zone H-23 during LOCA and HELB. The value for LOOP is the same as that for LOCA. The range of relative humidity for the cubicles in Zone H-24 are given in Table 3.11-5.

d. Duration Same as Zone H-23 during LOCA and HELB. The value for LOOP is the same as that for LOCA. Environmental conditions due to Loss-of-HVAC are assumed to persist for 7 days. 3.11.9.27 Environmental Zone H-25 Zone H-25 is composed of the RWCU pump cubicles shown as areas A.2.16, A.2.17, and A.2.18 in Drawing M01-1600-7. The bounding environmental conditions in this zone result from a LOCA or from a HELB within the cubicles.

a. Pressure The pressure following a LOCA is mitigated by the standby gas treatment system. The pressure is held below 1 inch of water as discussed in Subsection 6.2.3.

CPS/USAR CHAPTER 03 3.11-24 REV. 12, JANUARY 2007 The pressure following a HELB is documented in Reference 16. The peak pressures determined by this analysis are given in Table 3.11-5.

b. Temperature The temperature for the LOCA condition was not specifically evaluated for these cubicles. It is assumed conservatively that the temperature transient for the LPCS pump cubicle applies to these cubicles. The temperature following a HELB does not exceed 212

°F except for a short-term spike and is provided in Reference 16. c. Relative humidity The relative humidity is conservatively assumed to be 100% following a LOCA. Following a HELB, the relative humidity is assumed to be 100%.

d. Duration The pressure transient following a LOCA is inconsequential as discussed in Subsection 6.2.3. The temperature transients following a LOCA are conservatively assumed to be the same as those for the LPCS pump cubicle given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The temperatures at the end of the curves shown in Figure 6.2-142 are conservatively assumed to persist up to 100 days. The conditions stated in Table 3.11-5 for the HELB are presumed to persist for the duration of the accident. 3.11.9.28 Environmental Zone H-26 Zone H-26 is the containment above the HCU floor elevation. These areas are designated C.3.1, C.4.1, and C.5.1 in Drawing M01-1600 Sheets 8, 9, and 10. The bounding environmental conditions result from design-basis LOCA events.

a. Pressure Same as Zone H-15.

b. Temperature Same as Zone H-15. c. Relative humidity Same as Zone H-15.

d. Duration Same as Zone H-15.

CPS/USAR CHAPTER 03 3.11-25 REV. 12, JANUARY 2007 e. Submergence or Spray Same as Zone H-15 without pool swell. 3.11.9.29 Environmental Zone H-27 Zone H-27 is the drywell proper at the core midplane. Area C.3.2 in Drawing M01-1600-8 corresponds to this zone.

a. Pressure Same as Zone H-2.

b. Temperature Same as Zone H-2. c. Relative humidity Same as Zone H-2.

d. Duration Same as Zone H-2. 3.11.9.30 Environmental Zone H-28 Zone H-28 is the annular area between the RPV and the biological shield wall noted as areas C.3.3 and C.4.3 in Drawing M01-1600 Sheets 8 and 9. The bounding environmental conditions in this zone result from a LOCA inside the annular area that defines the zone or a small line break HELB (LOCA) in the drywell per Zone H-2.

a. Pressure The pressure transient for the HELB (LOCA) events in the drywell is the same as for Zone H-27 as given in Table 3.11-6. The pressure transients following a recirculation or feedwater line break inside the annulus are stated in terms of the peak pressures in Table 3.11-13. The analysis by which these pressures were determined is discussed in Subsection

6.2.1.2. b. Temperature The temperature transient following the HELB (LOCA) events in the drywell is the same as for Zone H27 as given in Table 3.11-6. The temperature transients following the HELB (LOCA) events inside the annulus are conservatively assumed to be the saturation temperatures for the peak pressures determined in Subsection 6.2.1.2. These temperatures are tabulated in Table 3.11-13.

CPS/USAR CHAPTER 03 3.11-26 REV. 12, JANUARY 2007 c. Relative humidity The relative humidity following a HELB (LOCA) in the drywell is given in Table 3.11-6. The relative humidity following a HELB (LOCA) in the RPV annulus is

given in Table 3.11-13.

d. Duration The time-histories for HELB (LOCA) events in the drywell are given in Table 3.11-6. The peak conditions for the HELB (LOCA) events inside the annulus as given in Table 3.11-13 are assumed to persist for the duration of the transients as discussed in Subsection 6.2.1.2. 3.11.9.31 Environmental Zone H-29 Zone H-29 is the main steam pipe tunnel in the containment. This is designated area C.3.4 in Drawing M01-1600-8. The bounding environmental conditions for this zone result from a LOCA or a RWCU line break in the tunnel. Main steam and feedwater line breaks are not considered because these lines are enclosed within guard pipes.

a. Pressure The pressure transient following a LOCA is discussed in Section 6.2. The bounding environmental conditions following a LOCA are specified from the generic Mark III evaluation of Reference 7 and the envelopes of Reference 8. The pressure transient following a HELB is discussed in Subsection 6.2.1.2.3.3.

b. Temperature The temperature transient following a LOCA is discussed in Section 6.2. The bounding environmental conditions following a LOCA are specified from the generic analysis of Reference 7 and the envelopes of Reference 8. The temperature transient following a HELB does not exceed 220

°F and is shown in Reference 16. c. Relative humidity The atmosphere in the zone is conservatively assumed to have a relative humidity of 100% following a LOCA and to be all steam following a HELB. These values are given in Tables 3.11-5 and 3.11-11. The atmosphere in this zone is conservatively assumed to be all steam following a HELB (see Table 3.11-5).

d. Duration The time-histories for the LOCA events are given in Table 3.11-11. The pressure and temperature time histories for the HELB are analyzed for 100 days or until conditions are restored to normal.

CPS/USAR CHAPTER 03 3.11-27 REV. 12, JANUARY 2007 3.11.9.32 Environmental Zone H-30 Zone H-30 is the main steam tunnel in the auxiliary building. This zone is denoted as areas A.3.3 and A.4.3 in Drawing M01-1600 Sheets 8 and 9. The bounding environmental conditions in this zone result from a LOCA or a HELB in the tunnel.

a. Pressure For LOCA, same as Zone H-18. The pressure transient for a HELB in the main steam tunnel is determined in Reference 11. The peak pressure of 13.8 psig determined by this analysis is assumed to persist for the duration of the transient.

b. Temperature Same as Zone H-18. c. Relative humidity Same as Zone H-18.

d. Duration Same as Zone H-18. 3.11.9.33 Environmental Zone H-31 Zone H-31 is composed of the areas in the auxiliary building switchgear rooms that are subject to elevated radiation after a LOCA. The bounding environmental conditions in these areas, shown on Drawing M01-1600 Sheets 8 and 9 as A.3.6, A.3.7, A.4.6, and A.4.7, result from normal operating conditions for these areas.

a. Pressure The areas in this zone are outside of the primary containment and thus receive no pressure effects due to a LOCA.

b. Temperature The areas in this zone are outside of primary containment and thus receive no temperature effects due to a LOCA. c Relative Humidity The areas in this zone are outside of primary containment and thus experience no change in relative humidity due to a LOCA.

d. Duration This is not applicable. The radiation, which is the only parameter that changes after an accident, is expressed as total integrated dose.

CPS/USAR CHAPTER 03 3.11-28 REV. 12, JANUARY 2007 3.11.9.34 Environmental Zone H-32 Zone H-32 is the drywell proper above a floor elevation of 778 feet. This zone is designated C.4.2 in Drawing M01-1600-9.

a. Pressure Same as Zone H-2.

b. Temperature Same as Zone H-2. c. Relative humidity Same as Zone H-2.

d. Duration Same as Zone H-2. 3.11.9.35 Environmental Zone H-33 Zone H-33 is the pipe tunnel designated as area C.4.4 in Drawing M01-1600-9. The bounding environmental conditions for this tunnel result from a LOCA and/or a HELB inside the tunnel.

a. Pressure The LOCA pressure conditions are determined from the generic analysis of Reference 7 and the envelopes of Reference 8. The bounding pressure transients are given in Table 3.11-11. The pressure in area C.4.4 due to a HELB in the RWCU system remains nearly constant (0 psig) throughout the transient.

b. Temperature The LOCA temperature conditions are determined from the generic analysis of Reference 7 and the envelopes of Reference 8. The bounding temperature transients are given in Table 3.11-11. The temperature following a HELB does not increase significantly from the normal temperature. c. Relative humidity The relative humidity following a LOCA/HELB is assumed conservatively to be 100%.

CPS/USAR CHAPTER 03 3.11-29 REV. 12, JANUARY 2007

d. Duration The time-histories for the LOCA events are provided in Table 3.11-11. The temperature time-history for the HELB event is provided analyzed for 100 days. 3.11.9.36 Environmental Zone H-34 Zone H-34 is the RWCU backwash receiving tank cubicle shown as area C.4.5 in Drawing M01-1600-9. a. Pressure Same as Zone H-33, except refer to Table 3.11-12 instead of 3.11-11.

b. Temperature Same as Zone H-33, except refer to Table 3.11-12 instead of 3.11-11. c. Relative humidity Same as Zone H-33, except refer to Table 3.11-12 instead of 3.11-11.

d. Duration Same as Zone H-33, except refer to Table 3.11-12 instead of 3.11-11. 3.11.9.37 Environmental Zone H-35 Zone H-35 is composed of the filter/demineralizer recirculating pump cubicles and the pipe tunnel designated as areas C.4.7 and C.5.8 in Drawing M01-1600 Sheets 9 and 10. The bounding environmental conditions in this zone result from a LOCA or a HELB.

a. Pressure The bounding environmental conditions in area C.4.7 are the same as those for area C.4.5 in Zone H-34. The response of area C.5.8 to a LOCA is the same as that of area C.4.5 in Zone H-34. The pressure in area C.5.8 due to a HELB in the RWCU valve room is assumed to remain at its peak value of 4 psig throughout the transient.

b. Temperature The response of area C.4.7 to a LOCA or a HELB is the same as that of area C.4.5 in Zone H-34. The response of area C.5.8 to a LOCA is also the same as the response of area C.4.5 in Zone H-34. The temperature in area C.5.8 due to a HELB in the RWCU valve room is shown in Reference 16. c. Relative humidity The relative humidity following either a LOCA or a HELB is assumed conservatively to be 100%.

CPS/USAR CHAPTER 03 3.11-30 REV. 12, JANUARY 2007

d. Duration The time histories for the LOCA events are provided in Table 3.11-12. The temperature for the HELB event are provided analyzed for 100 days. e. Submergence or Spray Flooding is discussed in Subsection D3.6.4. 3.11.9.38 Environmental Zone H-36 Zone H-36 is the RWCU crossover pipe tunnel shown as area C.5.11 in Drawing M01-1600-10. The bounding environmental conditions for this zone result from a spectrum of LOCA events in the drywell consisting of the large HELB or the small HELB as described in Table 3.11-12 or a HELB within the pipe tunnel.

a. Pressure The pressure transient following a LOCA is defined by the generic evaluation in Reference 7 and the envelopes of Reference 8. These values are shown in Table 3.11-12. The analysis for the pressure following a HELB is discussed in Section 6.2.1.2.3.6.

b. Temperature The temperature transient following a LO CA is defined by the generic evaluation in Reference 7 and the envelopes of Reference 8. The values are given in Table 3.11-12. The HELB conditions are determined as the saturation temperature for the pressure determined above. c. Relative humidity The relative humidity is provided in Table 3.11-5.

d. Duration The time history for the LOCA events are given in Table 3.11-12. 3.11.9.39 Environmental Zone H-37 Zone H-37 is the containment building general area above an Elevation of 856 feet, 0 inches.

Zone H-37, area C.7.1, is shown on Drawing M01-1600-11.

a. Pressure Same as Zone H-15.

CPS/USAR CHAPTER 03 3.11-31 REV. 12, JANUARY 2007

b. Temperature Same as Zone H-15. The upper region of this zone is subject to a steam environment of approximately 220

° F due to thermal stratification in the dome. This condition is conservatively assumed to persist for 100 days. c. Relative humidity Same as Zone H-15.

d. Duration Same as Zone H-15. In addition, the upper region of this zone is subject to a steam environment of approximately 220

° F due to thermal stratification in the dome. This condition is conservatively assumed to persist for 100 days. e. Submergence or Spray Equipment located in this zone is subject to the effects of the containment spray system operation. 3.11.9.40 Environmental Zone H-38 Zone H-38 is that portion of the gas control boundary that surrounds the containment structure

above floor Elevation 801 feet, 9 inches shown as area C.5.2 in Drawing M01-1600 Sheets 10 and 11. The bounding environmental conditions in this zone result from a LOCA.

a. Pressure The pressure environment for this zone is maintained below 1 inch of water by the standby gas treatments system following a LOCA. The analysis of this zone is discussed in Subsection 6.2.3.

b. Temperature The temperature environment for this zone is not explicitly evaluated as part of the analysis discussed in Subsection 6.2.3. The temperature for this zone is conservatively chosen to be that of the fuel building following a LOCA. This environment is tabulated in Table 3.11-14. c. Relative humidity The relative humidity is given in Table 3.11-14.

d. Duration The pressure transient following a LOCA is inconsequential. The temperature transient is conservatively assumed to be that of the fuel building as shown in Table 3.11-14.

CPS/USAR CHAPTER 03 3.11-32 REV. 12, JANUARY 2007 3.11.9.41 Environmental Zone H-39 Zone H-39 consists of the RWCU regenerative and nonregenerative heat exchanger cubicles shown as areas C.5.3 and C.5.4 in Drawing M01-1600 Sheets 9 and 10. The bounding environmental conditions for this zone result from a LOCA or a HELB in the cubicles.

a. Pressure The pressure transient following a LOCA is determined from the generic analysis of Reference 7 and the envelopes of Reference 8. This transient is shown in Table 3.11-12. The HELB analysis is discussed in Subsection 6.2.1.2.

b. Temperature The temperature transient following a LO CA is determined from the generic analysis of Reference 7 and the envelopes of Reference 8. This transient is given in Table 3.11-12. The temperature for the HELB is assumed to be the saturation temperature for the pressure described in Section 6.2.1.2. c. Relative humidity The relative humidity is assumed conservatively to be 100% following a LOCA. Following a HELB the environment is assumed to be all steam.

d. Duration The accident transient conditions are analyzed or extrapolated for 100 days. 3.11.9.42 Environmental Zone H-40 Zone H-40 is the combustible gas control equipment cubicles shown as areas C.4.8 in Drawing M01-1600-9 and C.5.5 in Drawing M01-1600-10. The bounding environmental conditions for this zone result from a LOCA.

a. Pressure Table 3.11-5 (Table 3.11-22) gives the pressure values.

b. Temperature Table 3.11-5 (Table 3.11-22) gives the temperature values. c. Relative humidity Table 3.11-5 (Table 3.11-22) gives the relative humidity values.

CPS/USAR CHAPTER 03 3.11-33 REV. 12, JANUARY 2007

d. Duration Table 3.11-5 (Table 3.11-22) gives the duration. 3.11.9.43 Environmental Zone H-41 Zone H-41 is the filter/demineralizer holding pump cubicle shown as area C.5.6 in Drawing M01-1600-10. The bounding conditions in this zone result from a LOCA or a HELB within the cubicle. a. Pressure Same as Zone H-36, except refer to Table 3.11-11 instead of Table 3.11-12, and refer to Section 6.2.1.2.3.7 instead of Section 6.2.1.2.3.6.

b. Temperature Same as Zone H-36, except refer to Table 3.11-11 instead of Table 3.11-12, and use the saturation temperature for pressure in Section 6.2.1.2.3.7 instead of Section 6.2.1.2.3.6. c. Relative humidity Same as Zone H-36.

d. Duration Same as Zone H-36, except refer to Table 3.11-11 instead of Table 3.11-12.

Conditions following a HELB are analyzed or extrapolated for 100 days. 3.11.9.44 Environmental Zone H-42 Zone H-42 is the filter/demineralizer valve room shown as area C.5.7 in Drawing M01-1600-10. The bounding environmental conditions for this zone result from a LOCA or a HELB in the cubicles.

a. Pressure Same as Zone H-36, except refer to Section 6.2.1.2.3.9 instead of Section

6.2.1.2.3.6.

b. Temperature Same as Zone H-36, except use the saturation temperature for pressure described in Section 6.2.1.2.3.9. c. Relative humidity Relative humidity following LOCA and HELB is provided in Table 3.11-5.

CPS/USAR CHAPTER 03 3.11-34 REV. 12, JANUARY 2007

d. Duration Same as Zone H-36. 3.11.9.45 Environmental Zone H-43 Zone H-43 is composed of the filter/demineralizer cubicles shown as areas C.5.9 and C.5.10 in Drawing M01-1600-10. The bounding environmental conditions for this zone result from a LOCA or a HELB within the cubicles.

a. Pressure The pressure-time histories for areas C.5.9 and C.5.10 for the LOCA events are provided in Table 3.11-12. The pressure-time histories for areas C.5.9 and C.5.10 due to the HELB events are discussed in Section 6.2.1.2.3.8.

b. Temperature The response of areas C.5.9 and C.5.10 to the LOCA events is shown in Table 3.11-12. The temperature response of these areas to the HELB event is shown

in Reference 16. c. Relative humidity The relative humidity is provided in Table 3.11-5.

d. Duration The time histories for the LOCA events are provided in Table 3.11-12. The time histories for the HELB events are provided in Reference 16. The conditions that exist at the end of these curves are assumed to persist for 100 days. 3.11.9.46 Environmental Zone H-44 Zone H-44 is the containment general area at the refueling floor. This zone is identified as area C.6.1 in Drawing M01-1600-11. The bounding environmental conditions for this zone result from design basis LOCA events.

a. Pressure Same as Zone H-15.

b. Temperature Same as Zone H-15. c. Relative humidity Same as Zone H-15.

CPS/USAR CHAPTER 03 3.11-35 REV. 12, JANUARY 2007

d. Duration Same as Zone H-15. 3.11.9.47 Environmental Zone H-45 Zone H-45 is composed of the RWCU heat exchanger valve cubicles shown as areas C.5.12 and C.5.13 in Drawing M01-1600 Sheets 9 and 10. The bounding environmental conditions for these cubicles result from a LOCA or a HELB in the cubicles.

a. Pressure Same as Zone H-36, except the analysis is discussed in Sections 6.2.1.2.3.5.

b. Temperature Same as Zone H-36, except refer to Section 6.2.1.2.3.5 to determine the pressure to be used to determine saturation temperatures. c. Relative humidity Relative humidity is provided in Table 3.11-5.

d. Duration The time history for the LOCA events are given in Table 3.11-12. The pressure and temperature following a HELB event are analyzed or extrapolated for 100 days. 3.11.9.48 Environmental Zone H-46 Zone H-46 is composed of the hydrogen recombiner cubicles in the control building. These are shown as areas D.1.8 and D.1.9 in Drawing M01-1600-12. The bounding environmental conditions for this zone result from the operation of the hydrogen recombiners following a

LOCA. a. Pressure Table 3.11-5 gives the envelope for the pressure conditions following a LOCA.

b. Temperature Table 3.11-5 gives the envelope for the temperature conditions following a

LOCA. c. Relative humidity Table 3.11-5 gives the relative humidity following a LOCA.

CPS/USAR CHAPTER 03 3.11-36 REV. 12, JANUARY 2007

d. Duration The LOCA conditions stated in Table 3.11-5 are applied for 100 days following the accident. 3.11.9.49 Environmental Zone H-47 Zone H-47 is composed of the standby gas treatment system filter train cubicles shown as areas D.2.8 and D.2.9 in Drawing M01-1600-13. The bounding environmental conditions in this zone result from the operation of the standby gas treatment system after a LOCA.

a. Pressure Table 3.11-5 gives the envelope for the pressure conditions following a LOCA.

b. Temperature Table 3.11-5 gives the envelope for the temperature conditions following a

LOCA. c. Relative humidity Table 3.11-5 gives the relative humidity following a LOCA.

d. Duration The LOCA conditions stated in Table 3.11-5 are applied for 100 days following the accident. 3.11.9.50 Environmental Zone H-48 This environmental zone is not used. 3.11.9.51 Environmental Zone H-49 This environmental zone is not used.

3.11.9.52 Environmental Zone H-50 This environmental zone is not used.

3.11.9.53 Environmental Zone H-51 Zone H-51 is the control building general area shown as D.2.10 in Drawing M01-1600-13. This area experiences an elevated radiation level following LOCA but the LOCA produces no other change in the environmental parameters. Bounding environmental conditions are produced by a LOOP transient in the area.

a. Pressure There is no significant pressure transient in the area following a LOOP.

CPS/USAR CHAPTER 03 3.11-37 REV. 12, JANUARY 2007

b. Temperature Summer and winter transients were calculated to provide a temperature envelope for these areas. The extreme temperature for the summer condition was reached at the end of the 100-day transient following a LOOP. For the winter transient there was a slight increase in the temperature at the end of 100 days. The initial condition in the area represents the minimum temperature condition. Temperature extremes are presented in Table 3.11-5. c. Relative Humidity The high normal relative humidity and the low normal relative humidity were used, respectively, as initial conditions for the winter and summer transients to provide a bounding relative humidity envelope. Humidity extremes are presented in Table 3.11-5.

d. Duration Time histories for the summer and winter LOOP temperatures are analyzed for

100 days. 3.11.9.54 Environmental Zone H-52 Zone H-52 is the radwaste pipe tunnel in the auxiliary building. This zone is identified as area A.2.2 in Drawing M01-1600-7. The bounding environmental conditions result from a LOCA or HELB within the tunnel.

a. Pressure The pressure transient following a LOCA is mitigated by the standby gas treatment system and held below 1 inch of water. See Subsection 6.2.3 for the supporting analysis. The pressure following a HELB within the sub-compartment results in the peak pressure stated in Table 3.11-5. This pressure is based on the analysis in

Reference 9.

b. Temperature The determination of the temperature for this area following a LOCA has not been specifically determined as part of the analysis presented in Subsection 6.2.3. However, the use of the temperature curves for the LPCS pump room profiles that appear in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4) are conservative for this zone. The temperature following a HELB is assumed conservatively to be the saturation temperature for the peak pressure given in Table 3.11-5.

CPS/USAR CHAPTER 03 3.11-38 REV. 12, JANUARY 2007 c. Relative humidity The relative humidity following LOCA does not rise significantly above the normal maximum. The environment following a HELB is assumed to be all steam.

d. Duration The pressure transient following a LOCA is inconsequential. The temperature transient is given in Figures 6.2-138 (Curve 2) and 6.2-142 (Curve 4). The conditions given at the end of Curve 4 in Figure 6.2-142 are conservatively applied over the period from 30 to 100 days. 3.11.9.55 Environmental Zone H-53 Zone H-53 is the area C.2.3 in Drawing M01-1600-7. The bounding environmental conditions for this zone result from design-basis LOCA events.

a. Pressure Same as Zone H-16.

b. Temperature Same as Zone H-16. c. Relative humidity Same as Zone H-16.

d. Duration Same as Zone H-16. e. Submergence or Spray Same as Zone H-16. 3.11.9.56 Environmental Zone H-54 Zone H-54 is the cubicle in the turbine building containing the turbine lube oil reservoir. This zone is denoted as area T.3.1 in Drawing M01-1600-21. The bounding environmental conditions in this zone results from a HELB involving a high energy instrument line.

a. Pressure This zone will experience a negligible pressure transient due to a HELB in a main steam instrument line. The peak pressure is given in Table 3.11-5.

CPS/USAR CHAPTER 03 3.11-39 REV. 12, JANUARY 2007

b. Temperature The peak temperature is conservatively chosen as that resulting from an isenthalpic expansion from reactor pressure to the cubicle pressure. The peak temperature is specified in Table 3.11-5. c. Relative humidity The relative humidity for this zone is given in Table 3.11-5.

d. Duration The cubicle would stay at peak temperature and humidity conditions for 100 days or until isolation of the break is achieved. 3.11.9.57 Environmental Zone H-55 Zone H-55 is the head cavity which is above, and separated from the drywell by the refueling bulkhead. The other boundaries are formed by the drywell head and the RPV head. There are ventilation paths between the head cavity and the rest of the drywell. This zone is denoted as

area C.5.14 in Drawing M01-1600-10. The bounding environmental temperature and humidity conditions result from a small HELB inside the head cavity. A large HELB outside the head cavity provides the bounding pressure condition during the initial portion of the transient.

a. Pressure This zone pressurizes to 30 psig in 1.5 seconds following a large HELB outside the head cavity and remains at this level during the first 40 seconds of the transient. During the next 5 seconds the pressure drops to 15 psig. After the first 45 seconds of the transient the bounding pressure environment is that for a small HELB inside the head cavity as given in Table 3.11-24.

b. Temperature The bounding environmental temperature condition is that for a small HELB inside the head cavity. In 1 second after the break a peak temperature of 340

°F is reached. The temperature transient is given in Table 3.11-24. c. Relative Humidity The bounding humidity environmental condition is that for a small HELB inside the head cavity in which an all steam environment is specified for the first 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> of the transient followed by 100% relative humidity for the rest of the 100 day

duration.

d. Duration Durations are as specified in Table 3.11-24 for the bounding line breaks as described above.

CPS/USAR CHAPTER 03 3.11-40 REV. 12, JANUARY 2007 e. Submergence or Spray The refueling bulkhead is designed to allow flooding of the head area during refueling. All vents have watertight caps which are closed to allow flooding of the

area. 3.11.9.58 Environmental Zone H-56 Zone H-56 denotes one general area in the auxiliary building outside secondary containment.

The area borders on RHR pump rooms A and B and is designated A.2.7 in Drawing M01-1600-

7. The bounding environmental conditions are produced by a LOOP transient in this zone.

a. Pressure The pressure is not likely to deviate significantly from the normal range as shown in Table 3.11-23.

b. Temperature Summer and winter transients were calculated to provide a temperature envelope for the area. Temperature extremes were reached at the end of the 100 day duration considered following a LOOP. These temperatures are presented in Table 3.11-5. c. Relative Humidity The high normal relative humidity and the low normal relative humidity were used respectively, as initial conditions for the winter and summer transients, to provide a bounding relative humidity envelope. Humidity extremes are presented in Table 3.11-5.

d. Duration The transient is assumed to last for the duration of the LOOP. 3.11.9.59 Environmental Zone H-57 Zone H-57 is the "control building general area" shown as D.2.11 and D.2.12 in Drawing M01-1600-13. This area experiences an elevated radiation level following LOCA but the LOCA produces no other change in the environmental parameters. Bounding environmental conditions are produced by a LOOP transient in the area.

a. Pressure Same as Zone H-51.

b. Temperature Same as Zone H-51.

CPS/USAR CHAPTER 03 3.11-41 REV. 12, JANUARY 2007 c. Relative Humidity Same as Zone H-51.

d. Duration Same as Zone H-51. 3.11.9.60 Environmental Zone H-58 Zone H-58 denotes one general area in the auxiliary building outside secondary containment. The area is a corridor adjacent to the RCIC instrument panel room and is designated are A.1.12 in Drawing M01-1600-6. The bounding environmental conditions are produced by a LOOP transient in this zone.

a. Pressure The pressure is not likely to deviate significantly from the normal range as shown in Table 3.11-23.

b. Temperature Summer and winter transients were calculated to provide a temperature envelope for this area. Temperature extremes were reached at the end of the 100 day duration considered following a LOOP. These temperatures are presented in Table 3.11-5. c. Relative Humidity The high normal relative humidity and the low normal relative humidity were used respectively, as initial conditions for the winter and summer transients, to provide a bounding relative humidity envelope. Humidity extremes are presented in Table 3.11-5.

d. Duration The transient is assumed to last for the duration of the LOOP. 3.11.9.61 Enviromental Zone H-59 Revision Zone H-59 denotes one general area in the auxiliary building outside secondary containment. The area borders on RHR pump room A and is designated A.1.15 in Drawing M01-1600-6. The bounding environmental conditions are produced by a LOOP transient in this zone. a. Pressure The pressure is not likely to deviate significantly from the normal range as shown in Table 3.11-23.

CPS/USAR CHAPTER 03 3.11-42 REV. 12, JANUARY 2007 b Temperature Summer and winter transients were calculated to provide a temperature envelope for this area. Temperature extremes were reached at the end of the 100 day duration considered following a LOOP. These temperatures are presented in Table 3.11-5. c. Relative Humidity The high normal relative humidity and the low normal relative humidity were used respectively, as initial conditions for the winter and summer transients, to provide a bounding relative humidity envelope. Humidity extremes are presented in Table 3.11-5.

d. Duration The transient is assumed to last for the duration of the LOOP. 3.11.9.62 Enviromental Zone H-60 Zone H-60 is composed of one general area of the auxiliary building. This area is identifieed as A.1.16 in Drawing M01-1600-6. The bounding environmental conditions for this area result from a LOOP. a. Pressure The area in this zone is outside the secondary containment and thus is not subject to any pressure effects due to a LOCA. The pressure variations which would occur in this area during a LOOP transient are given in Table 3.11-23.

b. Temperature The area in this zone is outside the secondary containment and thus is not subject to the temperature effects due to a LOCA. The bounding temperature is determined as the result of a postulated LOOP. The extreme expected temperatures are cited in Table 3.11-23. c. Relative Humidity The extreme values of relative humidity for this area are given in Table 3.11-23.

d. Duration The only transients of significance relate to the temperature. These are assumed to last for the duration of the LOOP.

CPS/USAR CHAPTER 03 3.11-43 REV. 11, JANUARY 2005 3.11.10 Mild Environmental Zones 3.11.10.1 Temperature, Pressure, and Relative Humidity The normal service conditions of the plant represent those conditions expected to occur during Planned Operation, with the normal HVAC systems in operation. The upper end of the temperature range represents the maximum value expected with all heat-generating equipment in the respective zones operating during a design-basis summer day. The lower end of the temperature range represents the minimum value expected during a design-basis winter day with equipment (except lighting) not operating and normal HVAC systems operating. 3.11.10.2 Mild Environment Radiation The radiation environment for normal reactor operation is determined on the basis of the reactor operating at full power, fuel leaking at the design-basis level, and assuming that all the systems at the station are operating normally at full capacity. The normal radiation environment is reported as the integrated radiation dose that can be absorbed by components over the 40-year lifetime of the station. The upper limit of the total integrated dose is 10 4 Rad (C) for mild environments. 3.11.11 Program for Continuing Qualification As a part of the engineering review of each equipment qualification report, the qualified life of each equipment or component is identified, along with any maintenance requirements for maintaining this qualified life. This information is recorded on separate sheets in the total qualification package, and made available to the plant operational maintenance staff at the time of system turnover. These replacement and maintenance requirements will be integrated into the Clinton Power Station preventive maintenance program described in existing plant procedures. This preventive maintenance program includes the following features: a. A computerized data base that includes performance frequency, assigned responsibility and applicable procedures, instructions, and requirements. b. Periodic computer-generated schedules are produced and distributed to assigned personnel to perform maintenance checks. c. Scheduled activities are performed and documented on schedules, data sheets, or other maintenance forms prescribed in procedures. d. Completed schedules are used to update and provide feedback to the data base for rescheduling activities. e. All documentation is retained for trend analysis and maintenance planning. f. A manual scheduling and tracking system is provided as a backup to the computer system. Maintenance actions resulting from Maintenance and/or Surveillance Testing are noted and result in the generation of a work document (either directly or subsequent to the generation of a CPS/USAR CHAPTER 03 3.11-44 REV. 14, JANUARY 2011 CPS Condition Report). Each work document generated for safety-related equipment is analyzed to determine if the required maintenance action was due to age-related degradation as a result of environmental conditions such as temperature, pressure, radiation, humidity, etc. If a maintenance action is the result of age-related degradation, the Preventive Maintenance and/or Surveillance Testing schedules are reevaluated and modified, as necessary, to prevent future age-related degradation problems. 3.11.12 Central File Description The Central File maintained at the Clinton plant site shall consist of the following items for Environmental Qualification of Electrical/Mechanical Equipment: a. A copy of USAR Section 3.11 including all tables and figures. b. System Description. c. Environmental qualification (EQ) binders for the equipment contained in List 1 of MS-02.00. Typically this would contain:

1. Index Page

2. Issue Summary Sheet 3. Checklist for NUREG-0588 requirements (Tab A & Tab B). 4. Maintenance and Surveillance Requirements (filed separately as MS-02.00). 5. Qualification Test Report (Tab F).

6. Analysis, Calculation, or Justification to support qualification (Tab C).

7. Equipment Identification (Tab D). 8. All supporting documents including excerpts from appropriate references (if required) (Tab G). 9. System Components Evaluation Worksheets (SCEW) (Tab H). d. Purchase order records pertaining to the Safety-Related equipment. e. All other references pertaining to the equipment such as drawings, etc. that are not listed in items a through d above. 3.11.13 TMI Items Requiring Environmental Qualification TMI items requiring environmental qualification such as area radiation monitors, miscellaneous recorders and transducers, etc., were added to List 1 of MS-02.00 as the systems were designed and the equipment purchased. Requirements resulting from the TMI-2 accident are addressed in Appendix D.

CPS/USAR CHAPTER 03 3.11-45 REV. 11, JANUARY 2005 3.11.14 Mechanical Equipment Qualification 3.11.14.1 Introduction Mechanical equipment which is environmentally qualified includes all safety-related active equipment located in harsh environmental zones. Other safety-related active mechanical equipment located in mild environmental zones and passive mechanical equipment is considered qualified by its construction as required by national codes, standards and NRC regulations. Mechanical equipment which is environmentally qualified is listed in List 1 of MS-02.00 as follows: a. Equipment Number: Provides the plant tag numbers of the equipment for reference and correlation with other documents and drawings.

b. Manufacturer: Identifies the manufacturer or vendor of the equipment.

c. Type/Model: Identifies the equipment catalog number. Components of the mechanical equipment listed in List 1 of MS-02.00 contain predominately metallic and some non-metallic materials. Since the effects of temperature, humidity and radiation are relatively insignificant for metallic components the environmental qualification is based only on their non-metallic materials. Examples of components upon which qualification is based are: - Gaskets - Grease - O-Rings - Packing

- Diaphragms - Seals

- Oil - Hydraulic Fluids - Lubricants The evaluation of equipment for qualification is done with the aid of a checklist which address the following items: 1. Equipment name, number, v endor/manufacturer and model number.

2. Vendor's equipment qualification report number, reference and status.

3. Equipment description, including location. 4. References to standards and documents, where applicable, if requirements are met by the qualification. 5. Identifies method of qualification; (i.e. Operating Experience, Analysis, Certificate of Compliance, Type Testing, or a combination thereof).

CPS/USAR CHAPTER 03 3.11-46 REV. 11, JANUARY 2005 3.11.14.2 Qualification Procedure For mechanical equipment listed in List 1 of MS-02.00 a materials list of organic components is generated from drawings, catalogues or the equipment vendor's instruction manuals.

Subsequently, an evaluation of the environmental impact on these organic components is made as follows: a. Certificate of Compliance Research procurement specifications to determine if environmental parameters for qualification of organic material were included in the equipment vendor's scope of work. If these parameters were specified, qualification is achieved through acquisition of a certificate of compliance from the equipment vendor. In some cases, the vendor's qualification report may also be reviewed to verify qualification.

b. Analysis When environmental parameters are not contained in the procurement specification, qualification will be established by analysis of those organic materials whose degradation could affect the ability of the equipment to perform its safety related function. Such analysis will demonstrate similarity between qualified and unqualified materials and their function, or show material suitability through degradation analysis such as Arrhenius calculations or other suitable methods. The resulting analysis will be made part of the qualification documentation. The feasibility of this approach depends on the type of components (organics) and their composition. c. Operating Experience Qualification of mechanical equipment using operating experience is used as a basis for environmental qualification should a) and b) not be feasible. This evaluation is done using similar equipment with a successful operating history in a service environment equal to or more severe than the environment for the equipment in question. The validity of operating experience as a means of qualification is determined from the type and amount of available supporting documentation, the service conditions and equipment performance. As this approach qualifies the equipment for normal environments, additional material degradation analysis is done to qualify the equipment for the Design Basis Events. This information is documented in the environmental

qualification documentation. d. Type Testing If environmental qualification of the item is not in the vendor's scope of work and b) and c) are not viable, the equipment is qualified by limited type testing of the components (organic) in question. A test specification is prepared either uniquely for Clinton or in a shared program with other utilities. The specification includes specific environmental qualification requirements and profiles for the equipment being qualified. Qualification results are reviewed and comments resolved prior to completing the equipment

qualification package.

CPS/USAR CHAPTER 03 3.11-47 REV. 11, JANUARY 2005 3.11.14.3 Central File Description The qualification document is maintained in the Central File (filed by equipment qualification package) at the Clinton Plant Site and consist of the following items: a. Individual component qualification documentation, (i.e. checklist and attachments) which identify conclusions of evaluation, equipment life, and maintenance requirements. b. Verification and quality assurance records associated with equipment qualification.

3.11.15 References

1. "Commission Memorandum and Order," CL1-80-21, May 23, 1980.

2. "Interim Staff Position on Environmental Qualification of Safety-Related Electrical Equipment," NUREG-0588 Revision 1, July, 1981. 3. "IEEE Standard for Qualifying lE Equipment for Nuclear Power Generating Stations," IEEE-323. 4. "IEEE Recommended Practices for Seismic Qualification of Class lE Equipment for Nuclear Power Generating Stations," IEEE-344. 5. "IEEE Standard for Design Qualification of Safety Systems Equipment Used in Nuclear Power Generating Stations," IEEE-627. 6. U.S. NRC letter to operating license applicants, "Qualification of Safety-Related Electrical Equipment," February 21, 1980. 7. "Containment Loads Appendix 3B to the 238 Nuclear Island GE Standard Safety Analysis Report - Submitted Amendment 2 to Application FDA review - STN 50-447," GE Document 22A7000, Revision 2. 8. "BWR Equipment Environmental Interface Data," GE Document 22A6926AA. 9. "High Energy Line Breaks Outside Containment," Sargent & Lundy Nuclear Safeguards and Licensing Division, Calculation 3C10-0477-001. 10. "Assessment of Weir Wall Annulus Flow During Drywell Depressurization," Sargent & Lundy Nuclear Safeguards and Licensing Division, Calculation 3C10-1280-001. 11. "Pressure and Temperature Effects of Postulated High Energy Line Breaks Outside of Containment," Sargent & Lundy Nuclear Safeguards and Licensing Division, Calculation 3C10-1274-001. 12. 10 CFR 50.49, Environmental Qualification of Electrical Equipment Important to Safety for Nuclear Power Plants. 13. Nuclear Station Engineering Standard MS-02.00, Maintenance of Equipment Qualification Program Manual.

CPS/USAR CHAPTER 03 3.11-48 REV. 11, JANUARY 2005 14. "Secondary Containment Functional Design," Sargent & Lundy Calculation 3C10-1079-001. 15. "Secondary Containment Subcompartment Parameters For Environmental Qualification of Equipment," Sargent & Lundy Calculation 3C10-0182-004. 16. "Equipment Environmental Design Conditions," Design Criteria Document.

CPS/USAR CHAPTER 03 3.11-49 REV. 11, JANUARY 2005 The following tables have been removed from the USAR. Table 3.11-1 Environmental Qualification List - BOP Electrical Equipment Table 3.11-2 Environmental Qualification List - NSSS Electrical Equipment Table 3.11-3 Qualification Table for BOP Components Table 3.11-4 Qualification Table for NSSS Components Table 3.11-5 Environmental Zone Summary Table*

Table 3.11-6 Abnormal Conditions in Drywell*

Table 3.11-7 Abnormal Conditions in ECCS Pump Rooms*

Table 3.11-8 Environmental Conditions for ECCS Subareas in Secondary Gas Control Boundary* Table 3.11-9 Deleted during FSAR Development Table 3.11-10 Deleted during FSAR Development Table 3.11-11 Abnormal Conditions in Containment*

Table 3.11-12 Abnormal Conditions in Cnmtainment Equipment Cubicles*

Table 3.11-13 Abnormal Conditions Around Shield Wall & RPV*

Table 3.11-14 Abnormal Conditions in Fuel Building* Table 3.11-15 Deleted during FSAR Development Table 3.11-16 Deleted during FSAR Development Table 3.11-17 Deleted during FSAR Development Table 3.11-18 Deleted during FSAR Development Table 3.11-19 Deleted during FSAR Development Table 3.11-20 Environmental Qualification List of Active NSSS & BOP Mechanical Equipment Table 3.11-21 Time-Temperature Profile for Area F.3.1*

Table 3.11-22 Abnormal Conditions in CGCS Cubicles Table 3.11-23 Environmental Zone Max & Min Temperature, Pressure, & Relative Humidity for Transient Conditions* Table 3.11-24 Abnormal Conditions for Area Above RPV Head and Below Drywell Head* Table 3.11-25 Abnormal Conditions in Drywell Under RPV*

• Information related to environmental conditions may be located in DC-ME-09-CP, Equipment Environmental Design Conditions Design Criteria Document.