ML22278A249

Rev. 21 to Updated Final Safety Analysis Report, Chapter 3, Section 3.4, Water Level (Flood) Design

ML22278A249
Person / Time
Site:	Limerick
Issue date:	09/19/2022
From:	Constellation Energy Generation
To:	Office of Nuclear Reactor Regulation
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LGS UFSAR 3.4 WATER LEVEL (FLOOD) DESIGN 3.4.1 FLOOD PROTECTION 3.4.1.1 Flood Protection Measures for Seismic Category I Structures Seismic Category I structures and the safety-related systems and components housed within them are listed in Table 3.4-1. As discussed in Section 2.4.2.2, the design basis flood level of the Schuylkill River at the site is 207 feet, including wave activity. The shortest horizontal distance from the contour at el 207' to the nearest safety-related structure is approximately 200 feet. Grade level is no lower than el 215' at any of the safety-related structures, and none of the safety-related structures has exterior openings below el 217'. Therefore, the safety-related structures are secure from Schuylkill River flooding and no special provisions for flood protection are necessary. Flooding from internal events is discussed in Section 3.6.1. The impact of local intense precipitation is discussed in Section 2.4.2.3.

Table 2.4-1 lists the access openings in safety-related structures. Two of these doors are below the design basis flood (el 207' MSL) but do not represent a flooding concern. The doors between the reactor enclosure and water pipe tunnel room 202 at el 201' are watertight. The doors between the control structure and the turbine enclosure at el 200' are watertight because the turbine enclosure is not watertight against the PMP as discussed in Section 2.4.2.3.

The failure of nonseismic Category I and nontornado protected tanks, vessels, and major pipes located outside buildings (Table 3.4-2) has also been evaluated and determined to not adversely affect safety-related structures, systems, and components as discussed below.

Tank Failure The location of tanks in the yard area is shown in Figure 3.8-58. Failure of the tanks on the west and south sides of the power plant complex (Table 3.4-2, items 1 through 5) will not cause potential flooding of safety-related structures, systems, and components. Any flooding due to a failure of these tanks will be contained within seismic Category IIA earth dikes, which will remain stable under both static and dynamic conditions. The design of the earth dikes is discussed in Sections 2.4.12 and 2.5.5.5.

The tanks on the north side of the power plant complex (Table 3.4-2, items 6 through 9) do not have seismically designed containments around them. Failure of these tanks could cause local flooding. This flooding would not adversely affect safety-related facilities for the following reasons:

a. Surface drainage in this area will drain water towards the Schuylkill River and Possum Hollow Run before it can reach the power plant complex.

b. Seismic Category I electrical cable and duct banks located in the vicinity of these tanks are adequate, as discussed below.

Even if the above dikes were to fail, there would be no impact on other safety-related structures, systems, or components due to site drainage.

CHAPTER 03 3.4-1 REV. 13, SEPTEMBER 2006

LGS UFSAR Failure of Cooling Tower Basin Wall The failure of the cooling tower basin wall (Table 3.4-2, items 10 & 11) would not adversely affect safety-related structures, systems, and components as discussed below.

The runoff pattern of water from the cooling tower basin wall failure would be similar to that caused by intense storm precipitation as discussed in Section 2.4.2.3 and shown in Figures 2.4-4, 2.4-5 and 2.4-6. Most of the flood water from the cooling tower basin would run away from the power plant complex. The worst case flood conditions for the power plant complex would be created by a failure of the south side of the Unit 1 cooling tower basin wall. For this case, a portion of the cooling tower basin water would flow towards the turbine enclosure. Although some limited turbine enclosure flooding may occur, there would be no impact on safety-related components.

The seismic Category I electrical cable and duct banks and valve pits located in the flow path of the water from the failed cooling tower basin are adequately protected as discussed below.

Failure of Circulating Water Conduit Failure of the conduit within the yard area between the cooling tower basin and the turbine enclosure (Table 3.4-2, item 12) will cause flooding of this area. Water from the damaged conduit will erode the soil cover and flood the yard.

The runoff pattern will be similar to that shown in Figure 2.4-4. The seismic Category I electrical cable and duct banks and valve pits, located in this area are adequate, as discussed below.

In the most severe case, all the water from the cooling tower basin could drain through the damaged conduit into the yard area between the cooling water pumphouse and turbine enclosure and cause flooding of the condenser pit. However, safety-related systems and components would not be damaged, as discussed in Section 10.4.1.3.3.

Failure of Major Yard Piping Failure of any of the pipes identified in Table 3.4-2, items 13 through 17, may cause local flooding.

However, the intensity and volume of water discharge from any of the pipes is less than that of the cooling water conduit failure discussed above and would not cause damage to any safety-related facilities. Soil erosion caused by failure of these pipes is discussed in Section 3.8.4.1.6.

Safety-related structures, systems, and components, including underground cables, will not be adversely affected by flooding or wetting caused by (1) design basis flood, (2) design basis precipitation or (3) failure of an outdoor tank or tank truck.

The safety-related structures, systems, and components located within the yard area are shown in Figure 3.8-58. Excavation for seismic Category I structures, pipelines, electrical duct banks, manholes, and underground diesel oil storage tanks are shown in Figure 2.5-37.

The design basis flood is not applicable to structures and yard facilities. The design basis flood level with respect to the Schuylkill River is 207 ft (Section 2.4.2.2), which is 10 feet lower than the lowest grade level entrance to any of the safety-related facilities.

CHAPTER 03 3.4-2 REV. 13, SEPTEMBER 2006

LGS UFSAR The protection of safety-related structures, systems, and components with respect to flooding caused by the design basis precipitation is discussed in Section 2.4.2.3. The runoff patterns for flood water caused by precipitation are shown in Figures 2.4-4 and 2.4-5. Below-grade parts of structures are protected from water intrusion as stated in Table 3.4-1.

The following seismic Category I yard facilities may be susceptible to flooding:

a. 5 Valve pits for RHRSW and ESW piping

b. 8 Valve pits for diesel generator fuel oil storage tanks The valve pits may be temporarily covered with water during an intense storm or major failure of an outdoor tank. However, the following protective features have been provided to resist resultant flooding of these facilities:

a. They are built as reinforced concrete boxes and are equipped with solid steel manhole covers with gaskets.

b. The tops of slabs of these facilities are elevated 3 to 12 inches above grade.

c. The RHRSW/ESW valve pits are equipped with drain pipes leading seepage water into normal waste drainage system.

d. The diesel generator fuel oil storage tank valve pits drain lines have been sealed off from the normal waste drainage system to minimize the potential for back flow from the downstream drainage pipe. Administrative controls shall assure that:

• These pits are periodically inspected (at least quarterly) at a frequency sufficient to assure that seepage water does not accumulate to a level at which it would enter the fuel oil tank or jeopardize safety related components in the pits.

• Accumulated seepage water is pumped out of the pits and out of the collection sumps of the concrete base surrounding the fuel oil storage tanks, as necessary.

• The pits are periodically inspected for material conditions adverse to pit flood resistance.

All electrical cables are designed to operate under water. Water absorption characteristics for all cables have been reviewed to confirm that even under flooded conditions, electrical cabling in manholes and duct banks will continue to operate properly. In addition, all electrical conduits that travel to electrical manholes outside the structures are sealed to prevent water from entering the structures through the electrical duct banks (Table 3.4-1).

The design features described above will also protect seismic Category I structures and yard facilities against internal flooding in case of failure of an outdoor tank or tank truck.

3.4.1.2 Permanent Dewatering System CHAPTER 03 3.4-3 REV. 13, SEPTEMBER 2006

LGS UFSAR All safety-related structures founded below the design basis groundwater levels can withstand the resulting hydrostatic loading. Safety-related systems and components not located within structures are founded above the maximum groundwater levels. Furthermore, the groundwater levels by safety-related structures, systems, and components are located below the top of rock. Therefore, a permanent dewatering system to protect safety-related structures, systems, and components is not required.

Groundwater does not affect the yard facilities because the design groundwater elevation (Section 2.4.13.5) is generally below yard facilities.

Rubberized flat dumbbell-type water stops have been provided at construction joints below the maximum expected groundwater level listed in Section 2.4.13.5 for all safety-related enclosures.

3.4.2 ANALYTICAL AND TEST PROCEDURES The safety-related structures listed in Table 3.4-1 are built of reinforced concrete with below-grade exterior walls that are a minimum of 1 foot thick for yard structures and 2 feet thick for the remaining listed structures. The roofs and parapets of all safety-related structures have been analyzed to verify their ability to withstand the static loading resulting from any water confined on the top of the structures due to local intense precipitation (discussed in Section 2.4.2.3).

The safety-related structures founded below the design basis groundwater levels (see Section 2.4.13.5) have been analyzed to verify their ability to withstand the resulting hydrostatic loading.

The spray pond pump structure and other safety-related structures at the spray pond have been designed to withstand the maximum water level in the spray pond.

CHAPTER 03 3.4-4 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.4-1 FLOOD LEVELS AT SAFETY-RELATED STRUCTURES SAFETY-RELATED SYSTEM ELEVATION OF DESIGN OR COMPONENTS LOWEST EXTERIOR FLOOD STRUCTURE HOUSED IN STRUCTURE ACCESS OPENING(1) ELEVATION

1. Reactor Primary containment, 217'-0" 207.0' Enclosure ECCS, miscellaneous safety-related systems and components

2. Diesel Diesel generator 217'-0" 207.0' Generator Enclosure

3. Spray Pond ESW system and RHRSW 268'-0" 254.9'(2) and Spray Pond system Pump Structure

4. Miscellaneous See Figure 3.8-58 See Figure 3.8-58 207.0' Yard Structures (1)

Penetrations below the lowest exterior access openings typically include underground electrical and piping penetrations. These penetrations are sealed water-tight.

(2)

The design basis flood level for the spray pond pumphouse in based on the design basis flood level of the spray pond (Section 2.4.8.2.1).

CHAPTER 03 3.4-5 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.4-2 YARD TANKS AND MAJOR PIPING (NONSEISMIC)

ITEM CAPACITY TYPE OF TORNADO NO. TANK OR PIPE DESCRIPTION OR FLOWS LOCATION CONTAINMENT PROTECTION 1 Unit 1 CST tank 200,000 gal. West of power plant complex Earth dikes None 2 Refueling Water 550,000 gal. West of power plant complex Earth dikes None 3 Unit 2 CST tank 200,000 gal. South of power plant complex Earth dikes None 4 Fuel oil storage tank 200,000 gal. South of power plant complex Earth dikes None 5 Deleted 6 Deleted 7 Deleted 8 Clarified water tank 200,000 gal. Water treatment plant north None None of power plant complex 9 Demineralized water tank 50,000 gal. Water treatment plant north None None of power plant complex 10 Cooling tower Basin 1 7x106 gal. North of power plant complex Reinforced None concrete walls 11 Cooling tower Basin 2 7x106 gal. North of power plant complex Reinforced None concrete walls 12 Cooling water pipes 8.0 ft diam. 450,000 gpm Between cooling tower and Underground Soil cover 4 pressure pipe each unit per unit turbine enclosure 13 36" makeup water pressure pipe 30,000 gpm From Schuylkill River to Underground Soil cover cooling tower 14 36" blowdown gravity pipe 18,000 gpm From cooling tower to Underground Soil cover Schuylkill River CHAPTER 03 3.4-6 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.4-2 (Cont'd) 15 36" makeup water pressure pipe 30,000 gpm From Perkiomen Creek to Underground Soil cover cooling tower 16 36" service water pressure pipe 35,000 gpm From cooling tower to Underground Soil cover turbine enclosure 17 12" fire loop pressure pipe 2,500 gpm Around plant complex Underground Soil cover CHAPTER 03 3.4-7 REV. 13, SEPTEMBER 2006

LGS UFSAR 3.5 MISSILE PROTECTION Where possible, the seismic Category I and safety-related structures, equipment, and systems are protected from missiles through basic station component arrangement so that the missile does not cause the failure of these structures, equipment, or systems. Where it is impossible to provide protection through plant layout, suitable physical barriers are provided to isolate the missile or to shield the critical system or component. Also, redundant seismic Category I components are suitably protected so that a single missile cannot simultaneously damage a critical system component and its backup system. Missile protection provides for either safe operation or shutdown during all operating conditions, operational transients, and postulated accident conditions. Table 3.2-1 provides a tabulation of safety-related structures, systems, and components, along with their applicable seismic category and quality group classification.

3.5.1 MISSILE SELECTION AND DESCRIPTION 3.5.1.1 Internally Generated Missiles (Outside Primary Containment)

There are three general sources of postulated missiles outside the primary containment:

a. Rotating component failure missiles

b. Pressurized component failure missiles

c. Gravitationally generated missiles 3.5.1.1.1 Rotating Component Failure Missiles The systems located outside the primary containment have been examined to identify and classify potential missiles. Redundant equipment is normally located in different areas of the plant or separated by walls so that a single missile from a rotating mass does not damage both redundant systems.

Catastrophic failure of rotating equipment such as pumps, fans, and compressors leading to the generation of missiles is not considered credible. Massive and rapid failure of these components is unlikely because of the conservative design, material characteristics, inspections, quality control during fabrication and erection, and prudent operation as applied to the particular component.

Additionally, the bases for considering it unlikely for rotating components, other than those identified in this section, to break through their casings and adversely impact safety-related equipment are the following:

A review of event reports on file at the Nuclear Safety Information Center, Oak Ridge National Laboratory, concerning failures of fans and missile generation indicated that a few fan failures have resulted in generation of missiles in safety-related areas of a nuclear facility. Small pump failures resulting in generation of missiles are considered more improbable than fan failures resulting in generation of missiles because pump casings are generally thicker than fan casings and pump speeds are generally lower than fan speeds. Even in the unlikely event that a rotating component does break through its casing, much of the component's kinetic energy would be dissipated in moving through the casing, thereby decreasing the probability of the component adversely CHAPTER 03 3.5-1 REV. 21, SEPTEMBER 2022

LGS UFSAR damaging a safety-related component. Therefore, generation of secondary missiles from the internally generated missiles described above is not considered credible. It is an even lower probability that a rotating component would adversely affect redundant safety-related systems because redundant equipment is generally located in different areas or separated by barriers.

Large pumps, such as the RHR and CS pumps, are installed so that the impeller section is surrounded by concrete; additionally, these pumps are located in separate compartments.

Therefore, they are not considered a missile hazard to other systems.

Missiles from HPCI or RCIC turbines would be contained by the walls of their compartments that form suitable barriers so that no other systems can be affected by a HPCI or RCIC turbine failure.

Other rotating equipment is not considered to constitute missile hazards because of its small size and/or the unlikelihood that it could move its rotating components through its housing. A review of the analyses of internally generated missiles performed for Palo Verde and San Onofre verified that postulated missiles from pumps and fans (e.g., a pump impeller or fan blade) typically do not have sufficient energy to penetrate the component casings. Because LGS uses pumps and fans that are generally designed and constructed in accordance with the same recognized industry codes and standards as those installed at Palo Verde and San Onofre, the results of the rigorous structural analyses conducted for those plants are indicative of the integrity of LGS equipment.

It was further verified that no internally generated missiles will cause loss of function of any system required for safe shutdown. The plant arrangement was reviewed to ensure that all essential systems or components are either remote from or separated by adequate barriers from potential missile sources.

In any case where a direct path exists between a potential missile source and equipment required for safe shutdown, potential missiles cannot impact more than a single component; therefore redundant equipment will be available to effect safe shutdown.

All HVAC fans located in safety-related enclosures are listed in Table 3.5-6, along with the reason why they are not credible missiles that could adversely affect redundant equipment needed for safe shutdown. For many fans, there are barriers that separate the potential fan blade missiles from essential systems. These barriers consist of walls, floors, and ceilings that totally enclose the potential missile and are sufficiently thick to prevent spalling. All redundant fans needed for safe shutdown are separated by adequate barriers. Class I seismic duct-work of a thickness substantially greater than normal duct-work is utilized in safety-related areas that have duct-work, which should preclude escape of a potential missile from connecting duct-work.

A review was also made of the fan installation drawings to ensure no inlet conditions to fans existed that could result in operating conditions that would impose cyclical fatigue on fan blades resulting in blade cracks. No such conditions exist.

The only fan blades that have the potential for damaging redundant components needed for safe shutdown, if they escape through their casing, are the drywell area unit cooler fans. A study was performed for these fans which demonstrated that the fan blades do not have sufficient kinetic energy to penetrate their casings. In addition, a complete physical inspection of all drywell unit cooler fan blades for blade angle, verification of proper torque on blade retaining bolts, and lock CHAPTER 03 3.5-2 REV. 21, SEPTEMBER 2022

LGS UFSAR wiring of the bolts by the manufacturer's representative was conducted to ensure that no blades would loosen during operation. This inspection is documented in the quality records.

3.5.1.1.2 Pressurized Component Failure Missiles The following are potential internal missiles from pressurized equipment:

a. Valve bonnets

b. Valve stems

c. Temperature detectors

d. Nuts and bolts Pressurized components in systems where service pressure exceeds 275 psig are evaluated as to their potential for becoming missiles.

Most valves of ANSI 900 psig rating and above and most valves of 600 psig rating, constructed in accordance with ASME Section III, are pressure seal bonnet type valves. For pressure seal bonnet valves, valve bonnets are prevented from becoming missiles by the retaining ring, which would have to fail in shear, 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 and hence the bonnets are not considered credible missiles.

The remaining valves of ANSI rating 900 psig and below are valves with bolted bonnets. Valve bonnets are prevented from becoming missiles by limiting stresses in the bonnet-to-body bolting material by requirements set forth in the ASME Section III and by designing flanges in accordance with applicable code requirements. Even if bolt failure were to occur, the likelihood of all bolts experiencing simultaneous complete severance failure is remote. The widespread use of valves with bolted bonnets and the low historical incidence of complete severance failure of bonnets confirm that bolted valve bonnets need not be considered as credible missiles.

Valve stems are not considered potential missiles if at least one feature in addition to the stem threads is included in their design to prevent ejection. Valves with back-seats are prevented from becoming missiles by this feature. In addition, air or motor-operated valve stems are effectively restrained by the valve operators.

Temperature or other detectors installed on piping or in wells are considered highly improbable missiles, since a complete and sudden failure of a circumferential weld is needed for a detector to become a missile.

Nuts, bolts, nut and bolt combinations, and nut and stud combinations have little stored energy and thus are of no concern as potential missiles.

In addition, high pressure gas cylinders and accumulators are not considered credible internally generated missiles impacting on redundant safety-related equipment for the following reasons:

CHAPTER 03 3.5-3 REV. 21, SEPTEMBER 2022

LGS UFSAR

a. All accumulators located in areas that contain safety-related equipment are listed in Table 3.5-3, along with the maximum pressure and temperature at which they operate. These accumulators have low stresses and operate in the "moderate energy" range. Therefore, any failures would not be of concern for missile generation.

b. Safety-related gas cylinders and accumulators are seismically supported, which ensures that they will not become gravity-generated missiles during design basis conditions. Nonsafety-related cylinders and accumulators are seismically restrained if they are located in the vicinity of safety-related components.

c. High pressure gas (H2, O2, N2, Kr) cylinders meet the manufacturing and test requirements of the DOT standards, 49CFR178.37, Specification 3AA and ICC standards. Vessel wall thickness requirements of both standards are comparable to ASME Section III, Class 3 requirements. Because of this highly conservative design, any failure of these cylinders that would cause them to become missiles is highly improbable.

d. Halon gas cylinders used in the fire protection system are located in the turbine enclosure and are isolated from safety-related equipment.

3.5.1.1.3 Gravitationally Generated Missiles Installed equipment and components in safety-related plant areas outside containment are designed and installed so that they would not present gravitational missile hazards to safety-related structures, systems, or components during or after an SSE. This is achieved for safety-related equipment by seismic Category I design and installation (in accordance with Regulatory Guide 1.29 criteria) and for nonsafety-related equipment by seismic Category IIA design and installation (Section 3.2). Nonpermanently installed equipment is either removed from the safety-related areas, adequately separated from safety-related components, or secured in place before reactor operation to ensure that it would not present a missile hazard.

The information requested in the generic letter dated December 22, 1980 regarding conformance to the criteria contained in NUREG-0612, "Control of Heavy Loads at Nuclear Power Plants" is provided in Reference 9.1-1.

3.5.1.2 Internally Generated Missiles (Inside Containment)

There are three general sources of postulated missiles inside the primary containment:

a. Rotating component failure missiles

b. Pressurized component failure missiles

c. Gravitationally generated missiles 3.5.1.2.1 Rotating Component Failure Missiles CHAPTER 03 3.5-4 REV. 21, SEPTEMBER 2022

LGS UFSAR The most substantial piece of NSSS rotating equipment is the recirculation pump and motor. This potential missile source is covered in detail in Reference 3.5-1.

It is concluded in Reference 3.5-1 that destructive pump overspeed can result in certain types of missiles. A careful examination of shaft and coupling failures shows that the fragments do not result in damage to the containment or to vital equipment.

a. Low energy missiles (kinetic energy less than 1,000 ft-lbs):

Low energy level missiles may be created at motor speeds of 300% of rated through failure of the end structure of the rotor. The structure consists of the retaining ring, the end ring, and the fans. Missiles potentially generated in this manner would strike the overhanging ends of the stator coils, the stator coil bracing, support structures, and two walls of 1/2 inch thick steel plate. Due to the ability of these structures to absorb energy, it is concluded that missiles would not escape this structure. It is at this point that frictional forces would tend to bring the overspeed sequence to a stop.

b. Medium energy missiles (kinetic energy less than 20,000 ft-lbs):

In the postulated event that the body of the rotor bursts, medium energy missiles could be created. The likelihood that these missiles would escape the motor is considered less than the likelihood of escape for the low energy missiles described above, due to the additional amount of material constraining missile escape, such as the stator coils and stator frame directly adjacent to the rotor.

c. The motor as a potential missile:

Since bolting is capable of carrying greater torque loads than the pump shaft, pump bolt failure is precluded. Since pump shaft failure decouples the rotor for the overspeed driving blowdown force, only those cases with peak torques less than that required to fail the pump shaft (five times rated) have the capability to drive the motor to overspeed. When missile generation probabilities are considered along with a discussion of the actual load-bearing capabilities of the system, it is evident that these considerations support the conclusion that it is unrealistic that the motor would become a missile.

It is concluded that the other rotating components inside the containment, such as fans, do not have sufficient energy to move the masses of their rotating parts through the housings in which they are contained and therefore are not considered missile hazards. Additional evaluation and discussion of the potential for rotating equipment generated missiles to impact safety-related equipment is provided in Section 3.5.1.1.1 above.

3.5.1.2.2 Pressurized Component Failure Missiles It is concluded that potential internal missiles inside the containment from pressurized components are not considered credible for the same reasons listed in Section 3.5.1.1.2.

3.5.1.2.3 Gravitationally Generated Missiles Installed equipment and components in the containment are designed so that they would not present gravitational missile hazards to safety-related structures, systems, or components during or CHAPTER 03 3.5-5 REV. 21, SEPTEMBER 2022

LGS UFSAR after an SSE. This is achieved for safety-related equipment by seismic Category I design and for nonsafety-related equipment by seismic Category IIA design, as discussed in Section 3.2.

Nonpermanently installed equipment is either removed from the containment or secured in place before reactor operation to ensure that it does not become dislodged and present a missile hazard.

3.5.1.3 Turbine Missiles The turbine inspection program which is being implemented in Section 10.2.3.6 uses a probabilistic approach for scheduling the inspection and replacement of the low pressure turbine rotors with shrunk-on discs. This turbine inspection program is based upon the missile probability analysis methodology in Reference 3.5-15 and the probabilistic approach established in NUREG 1048 (Reference 3.5-14). The approach in Reference 3.5-14 was found acceptable by the NRC for use in establishing maintenance and inspection schedules for specific turbine systems including the original Main Turbines installed at Limerick Generating Station. As a result of the replacement of the original Main Turbines, the missile probability analysis methodology in Reference 3.5-15 has been adopted to maintain the probabilistic approach to determine the inspection program for the installed turbine system. The overspeed protection system and turbine valve test interval was determined by References 3.5-16 and 3.5-17 in accordance with the missile probability analysis methodology in Reference 3.5-15.

The intent of the program is to ensure that the probability of a turbine generating a missile is maintained less than 1x10-5 per year. The program takes into account specific turbine wheel operating conditions, material properties, results of periodic in service inspections, and other factors. The program's determination of missile probability is based on the probabilities of individual parameters which may lead to the generation of the turbine missile. As a result, the program can facilitate evaluations of the effects of changes in any parameter. Table 3.5-9, Turbine System Reliability Criteria, has been extracted from table U.1 of Reference 3.5-14, for use with an unfavorably oriented turbine. The probability of unacceptable damage from turbine missile will be maintained at less than or equal to 1x10-7 per year.

Schedules for future inspection and replacement of low pressure turbine rotors with shrunk-on discs will be based on the probabilistic approach.

3.5.1.4 Missiles Generated by Natural Phenomena Only tornado-generated missiles have been considered. Missiles used in the design and assessment of structures and openings are listed in Table 3.5-4. All safety-related structures, systems, and components listed in Table 3.2-1 were reviewed for adequacy against tornado-generated missiles listed in Table 3.5-4. These safety-related structures, systems, and components are designed either to resist tornado missiles in accordance with Reference 3.5-6 or are protected by these tornado-resistant enclosures. The structures designed for these tornado missiles and the systems protected are listed in Table 3.3-2. Table 3.5-8 provides information on the characteristics of these barriers. Additionally, ESW and RHRSW systems yard piping is protected by burial and separation of redundant loops.

The exterior walls and roof thicknesses have been evaluated for the tornado-resistant enclosures listed in Table 3.3-2 and are capable of withstanding all of the missiles listed in Table 3.5-4. The 4000 psi strength concrete walls and roofs have minimum thicknesses of 24 inches and 18 inches, respectively (Table 3.5-8). This exceeds the minimum acceptable missile barrier thickness requirements specified in table 1 of SRP (NUREG-0800, July 1981) section 3.5.3.

CHAPTER 03 3.5-6 REV. 21, SEPTEMBER 2022

LGS UFSAR Where necessary to protect safety-related components, the doors in these enclosures have been designed to withstand the 1 inch steel rod, the utility pole, the 6 inch steel pipe, and the 12 inch pipe in addition to the original three design basis missiles. This design is in accordance with the SRP (NUREG-0800, July 1981).

The probability of any of the above tornado missiles penetrating any of the openings and structures (not specifically designed to be resistant to the above missiles) and adversely impacting safety-related components, such that safe shutdown is prevented, is considered low for the following reasons:

a. The total area of the nontornado-resistant features that have safety-related components located behind them is extremely small compared to the total area of the tornado-resistant portions of the enclosures.

b. To penetrate a nontornado-resistant feature and travel a sufficient distance to impact a safety-related component, a missile would need to strike the feature at a perpendicular angle.

c. Much of the missile's kinetic energy would be dissipated in breaking through the nontornado-resistant feature, which reduces the possibility of the missile adversely damaging a safety-related component even if it strikes one.

d. Redundant safety-related components are normally located in different areas of the plant or yard or are separated by walls so that a single tornado missile would not damage both redundant systems.

Evaluations have been carried out considering all seven tornado missiles listed in Table 3.5-4 and their penetration/impact on the openings and nontornado designed structures of the plant including manhole covers and valve pit roofs. Even though some singular safety-related components may be affected, damage would not occur to all redundant systems, and safe shutdown could be achieved even when assuming an additional single active failure.

LGS is in conformance with Regulatory Guide 1.117 regarding systems to be protected from tornado missiles except as discussed below where unacceptable damage to unprotected spray networks is not considered credible.

As described in Section 9.2.6, the ultimate heat sink at LGS is an excavated spray pond with a surface area of 9.9 acres. Four spray networks, each having 50% capacity for shutdown of two units, are provided.

Details of the spray pond excavation and finished grading are shown in Figures 3.8-55, 3.8-56, and 3.8-57. The general arrangement of the spray pond, spray networks, and spray pond pump structure is shown in Figure 9.2-6. The layout of the spray networks is shown in M-384.

As discussed in Section 3.5.1.4, all essential structures, systems, and components related to the ESW system, RHRSW system, and the UHS are protected from the effects of tornadoes and tornado missiles. Protection of the spray networks from tornado missiles is provided by location of CHAPTER 03 3.5-7 REV. 21, SEPTEMBER 2022

LGS UFSAR the network piping and sprays below the surrounding grade and by physical separation of the networks:

a. In all but the spillway area, the surrounding grade is in excess of el 260' while the top of the sprays are at el 258' and the spray network piping is between el 253'-5" and el 256'-8".

b. The closest branches of adjacent spray networks are separated by 65 ft.

c. The supply piping to adjacent networks is separated by 215 ft.

d. The networks are located at a minimum distance of 72 ft from the edge of the pond.

The use of elevational differences and physical separation to provide protection of the spray pond networks from tornado missiles is justified by the following considerations:

a. Only two spray networks are required for the safe shutdown of both units.

b. The only active failure that can compromise the operability of a spray network is failure of its supply valve (HV-012-032A, B, C or D). These valves may be manually operated to isolate damaged networks or to initiate the use of undamaged networks if their controls or motors are inoperable.

c. The physical arrangement of the spray networks precludes the possibility that large missiles can damage more than one spray network due to trajectory considerations.

Multiple missiles of sufficient energy and distribution to substantially damage multiple networks are unlikely. Network piping varies in size from the 30 inch diameter supply headers to the 2 inch diameter piping at the extreme ends of the distribution branches. Network piping wall thickness varies from 0.337-0.500 inch.

d. The loss of some sprays in a network does not result in substantial loss of heat removal capability for the entire network (each network contains 240 spray nozzles).

e. The design thermal performance of the spray pond is based on conservative design values of initial pond temperature and meteorology as described in Section 9.2.6.4.

For all expected conditions, the margin in thermal performance would be considerably greater than the 10% margin demonstrated under design conditions.

In fact, for average meteorological conditions, a single spray network is sufficient for the removal of the heat rejected from both units for at least a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period.

f. Interconnections are provided that allow the use of the cooling towers as a heat sink for ESW and RHRSW systems. Such operation may be initiated from the control room or locally by manual operation.

It is unlikely that tornado winds would compromise the heat removal capability of the spray pond networks, or the cooling towers, to the extent that safe shutdown of the units would be affected. The spray networks have been designed to withstand design basis tornado winds. While not specifically designed to withstand design CHAPTER 03 3.5-8 REV. 21, SEPTEMBER 2022

LGS UFSAR basis tornado winds, the cooling tower shell and supporting structure have been designed to withstand the following wind loading when either operating or dry:

Elevation Above Wind Velocity Grade (ft) (mph) 30 90 150 113 200 118 300 125 400 130 500 135 The cooling towers are expected to provide sufficient heat removal capability for the safe shutdown of the units even in the event that the tower fill is extensively damaged.

g. The loss of more than two spray networks and the coincident loss of the cooling towers due to tornado missiles is unlikely due to physical separation of the cooling towers and the spray pond. The cooling towers are located approximately 600 feet from the nearest portion of a spray network.

The likelihood of tornado winds and/or missiles affecting the safe shutdown capability of the cooling towers and spray networks at the same time is quite remote when the above described design factors are considered together with the variation in tornado intensity along its path length and width (Reference 3.5-2).

h. Tornado missiles are an insignificant contributor to plant risk because of the low frequency of occurrence of tornadoes in this region (EROL section 2.3.1.2.2) and the low likelihood of damaging missiles if one were to occur.

Even if the safe shutdown capability of the cooling towers and spray networks were compromised by tornado effects, use of the cooling tower basins and/or UHS in a "cooling pond type" mode would allow substantial time for spray network repair. A plant procedure governs such repair activities. This procedure will contain, at a minimum, the following elements:

1. Repair work on damaged spray networks will begin immediately, using materials, equipment, and personnel that have been verified to be available.

Procedure verification will be made each year.

2. On current loss of all UHS spray headers and all cooling tower cooling capacity, the spray pond will be operated as a closed cycle cooling pond until the temperature of the water reaches the design limit of 95F. In this mode, water will be returned to the pond via the winter bypass line to promote thermal mixing and minimize the likelihood or recirculation.

Under design basis conditions of initial pond temperature and meteorology, it would take approximately 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> for the pond to reach the 95F limit.

CHAPTER 03 3.5-9 REV. 21, SEPTEMBER 2022

LGS UFSAR Under average conditions, it would take approximately 10 hours1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br /> to reach this limit. Both numbers are for two-unit, full power operation. For single-unit operation, these times would be approximately 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> and 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br />, respectively. The heat rejection rate can be further reduced by depressurizing the reactor at a slower rate than 100oF/hr assumed in the design basis analysis.

3. When the pond reaches the design temperature limit, the sluice gates between the spray pond pumphouse wet wells and the spray pond will be closed. Water will be released from the cooling tower basins into the wet wells and pumped through the plant to service the required heat loads. The water will be returned to the spray pond and allowed to discharge over the blowdown weir and storm spillway.

The two cooling tower basins contain a total of 14 million gallons. If it is conservatively assumed that only one-half of this volume of water is available, there is sufficient water to provide makeup for the ESW and RHRSW pumps, operating in a once-through mode, for an additional 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />. In the unlikely event that the cooling tower basin walls have failed due to tornado missiles, the additional time of 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> would not be available. However, the spray pond PRA demonstrates that it is extremely improbable that the four spray pond networks would not be available.

4. Sufficient makeup water can be supplied to the cooling tower basins to sustain continuous operation in this mode for a number of sources as described in Item i.

i. The Schuylkill River makeup pumphouse is located approximately 1500 ft from the nearest cooling tower, making it unlikely that the pumphouse would be damaged by a tornado that would also compromise the spray pond networks and the cooling towers. The pumphouse is powered from the 2300 V plant services switchgear.

The switchgear can be fed using offsite power from either of the two plant substation via underground lines. The two substations are approximately 2000 ft apart, making it highly unlikely that both substation would be disabled by a tornado that would also compromise the spray pond networks and the cooling towers.

While an additional source of water is available from the pump station to provide the Perkiomen makeup supply located approximately 8 miles from the plant site, no reliance is being placed on this intake for the purpose of safety analysis nor the safety licensing basis for the facility.

If existing sources of makeup cannot be made available in a timely manner, makeup will be provided using available portable pumps of required size and capacity to pump water from the Schuylkill River to the spray pond pumphouse wet wells. The water would be pumped via a tie-in to the existing underground water pipeline from the Schuylkill River intake pumphouse to the cooling tower basins. It would flow via gravity to the pump pits. If a tie-in to the existing pipeline is not possible, the water would be pumped directly to the wetwell through temporary CHAPTER 03 3.5-10 REV. 21, SEPTEMBER 2022

LGS UFSAR lines. The portable pumps that would be used are either licensee owned or rental pumps. The required pumps will be verified to be available on an annual basis.

j. Plant EOPs address the various contingency actions available to the operators to deal with degraded UHS conditions. As indicated in the above discussions, substantial time is available for corrective operator actions. If UHS capability should be lost for such a period of time that conditions degraded considerably, the EOPs would direct the use of equipment that would achieve a safe stable state regardless of UHS capability.

3.5.1.5 Missiles Generated by Events Near the Site The safety-related structures, systems, and components were reviewed for adequacy against missiles externally generated by railroad explosions. The safety-related facilities are either designed to resist the externally generated missiles in accordance with Reference 3.5-6 or are protected by these missile-resistant barriers. The barriers designed to resist externally generated missiles and the corresponding systems and components protected by these missile barriers are listed in Table 3.5-7.

The nearest possible train explosion accident and its consequent missiles are considered to be the most severe missile-generating event that could occur near the site. The postulated missiles resulting from such an accident considered in the design of structures protecting safety-related systems are listed in Table 3.5-5. Missiles resulting from truck, industrial, and pipeline explosions would be less severe and therefore are not considered. As demonstrated in Section 2.2, there is no potential for missiles from ship or barge explosions or military installations. Descriptions of the railroad, its location relative to the plant, the railroad explosion, and explosions from other sources are given in Section 2.2.

3.5.1.6 Aircraft Hazards The safety-related structures, systems, and components were reviewed for adequacy against missiles externally generated by aircraft accidents. The safety-related facilities are either designed to resist the externally generated missiles in accordance with Reference 3.5-6 or are protected by these missile-resistant barriers. The barriers designed to resist externally generated missiles and the corresponding systems and components protected by these missile barriers are listed in Table 3.5-7.

Airports and aircraft activity in the vicinity of the site are discussed in Section 2.2.2. Based on the airport and aircraft information provided in Section 2.2.2, an analysis has been performed using the methodology of SRP section 3.5.1.6 to demonstrate that the probability of an aircraft accident causing damage to safety-related equipment or structures sufficient to result in radiological consequences that are a significant fraction of 10CFR50.67 exposure guidelines is lower than the acceptance criteria of SRP section 3.5.1.6.

3.5.1.6.1 Design of Safety-Related Structures Aircraft operating or capable of operating out of nearby airports are listed in Table 3.5-1. Based on the estimated maximum kinetic energy at impact and corresponding local damage to concrete elements, the Learjet is determined as the design aircraft for the design of safety-related structures.

CHAPTER 03 3.5-11 REV. 21, SEPTEMBER 2022

LGS UFSAR The characteristics of the Learjet are presented in Table 3.5-2, and the associated load curve acting on structures due to the impact is shown in Figure 3.5-1. The design aircraft is assumed to impact perpendicularly to the vertical surface of the buildings or to impact at a glide angle of 15° from horizontal on the roof of buildings. The reactor enclosure, control structure, and spray pond pump structure are designed to withstand the impact of the design aircraft without loss of structural integrity. Safety-related equipment and systems located therein are designed to remain functional.

The diesel generator enclosures are not designed to withstand aircraft impact.

3.5.1.6.2 Analysis Method The probability of damage to structures containing safety-related equipment (P) is calculated from the formula:

L M P= Nij Cij Aij (EQ. 3.5-1) i=1 j=1 where:

L = number of different flight paths that affect the site M = number of different types of aircraft that may constitute a hazard Nij = number of movements per year by aircraft j operating in flight path i Cij = probability of a crash per square mile per aircraft movement for aircraft j operating in flight path i Aij = effective plant area (square miles) for aircraft j operating in flight path i Values for the above parameters and related assumptions are discussed in the following sections.

3.5.1.6.2.1 Aircraft Movements (Nij)

In addition to the movements, type of aircraft, and other particulars of aircraft operation described in Section 2.2.2, the following considerations for each of the sources of aircraft movement apply.

Pottstown-Limerick Airport - The analysis assumes that 90% of the takeoffs (45% of the movements) are towards the west, in the direction of the site, and that 10% of the landings (5% of the movements) are from the west, in the direction of the site. This is based on the VFR and IFR information provided in Section 2.2.2. The flight restrictions discussed in Section 2.2.2 are accounted for in the analysis.

Based on the approach pattern for Pottstown-Limerick Airport, the holding area is at least 3.8 miles northeast of the runway, which would locate the holding area more than 5 miles northeast from the site. Due to the distance and direction, no additional movements involved with the holding pattern were considered in calculating the aircraft crash probability.

CHAPTER 03 3.5-12 REV. 21, SEPTEMBER 2022

LGS UFSAR Pottstown Municipal Airport - The proportions of single-engine and twin-engine aircraft using the airport are assumed to be 70% and 30%, respectively. Based on the information provided in Section 2.2.2, it is assumed that 50% of the takeoffs involve eastbound passes over the site, with the remainder to be north and west away from the site. Similarly, it is assumed that 50% of the landings involve northbound flight patterns over the site.

New Hanover Airport, Sunset Landing Strip and Perkiomen Valley Airport - The projected annual number of operations for these facilities located between 5 and 10 miles from the site is less than 500 D2. In accordance with SRP section 3.5.1.6, their contributions to the total probability are extremely small and are not considered in the analysis.

Federal Airways - The traffic data presented in Section 2.2.2 are employed in the analysis. It is conservatively assumed that the Boeing 727 is representative of traffic in these airways.

LGS Plant Site Heliport - The analysis is based on 156 landings and 156 takeoffs per year. The helicopter parameters used are given in Table 3.5-1. These parameters are bounded by those of the Learjet, which was the design basis aircraft for impact.

3.5.1.6.2.2 Aircraft Crash Probability (Cij)

Pottstown-Limerick and Pottstown Municipal Airports Aircraft crash probabilities from the SRP section 3.5.1.6, are used in this analysis. The SRP specifies the crash probability per square mile in an annular sector 60 on each side of the extended runway centerline as a function of the distance from the end of the runway.

Accordingly, the crash probability for U.S. general aviation at a distance of 1-2 miles of 1.5x10-7 per square mile per aircraft movement is used for Pottstown-Limerick Airport. This probability is also applied to the rotary-wing operations at that airport. For Pottstown Municipal Airport, the U.S.

general aviation probability for 4-5 miles, 1.2x10-8 per square mile per aircraft movement, is used.

The use of this value for Pottstown Municipal Airport is conservative because LGS lies outside the 60 sector of the extended runway centerline.

LGS Plant Site Heliport The crash probability for the helicopter traffic using the plant site heliport is based on NTSB and FAA nationwide statistics on the estimated number of helicopter flight hours for 1979 and the number of fatal helicopter crashes during takeoff and landing during that same year, and is equal to 4.88x10-7 per takeoff or landing.

Federal Airways For commercial aircraft in the Federal airways, a model for the probability that an aircraft crash will result at a distance (x) normal to its flight path is employed, which was developed by Solomon (References 3.5-7 and 3.5-8) using a negative exponential distribution. Solomon's model requires estimation of the deviation, which was determined by a 20 angle from the flight path. A deviation is estimated here by referencing Solomon's angle at 14,000 feet, and adding one mile for deviation in the airway. This gives the probability:

CHAPTER 03 3.5-13 REV. 21, SEPTEMBER 2022

LGS UFSAR f(x) = 1 e-x/1.97 (EQ. 3.5-2) 3.94 Where:

x = distance in miles The probability that an enroute crash will occur is about 0.45x10-9 per flight path mile, based upon U.S. commercial aviation performance in the period 1970-75 (Reference 3.5-9). Thus, the crash probability at the site per passage in each of the Federal airways can be estimated by the product of these two probabilities.

The resulting probabilities are as follows:

Probability Airway x (miles) (per square mile)

V29/147 1.3 5.9x10-11 Pottstown VOR 1.3 5.9x10-11 320 radial V210 8.0 2.2x10-12 V276 10.0 7.1x10-13 3.5.1.6.2.3 Critical Target Areas (Aij)

The aircraft crash probability estimates the chance that a given aircraft maneuver (passage, takeoff, or landing) would result in striking the ground within a specific unit area, without regard to the obstruction of surrounding objects. If the assumed flight ray of the aircraft in crashing should pass through an obstructing object, then the probability of crash into that obstructing object is the same as for the associated unit ground area.

The determination of target areas are based on the following considerations:

a. The skid area, associated with a ground strike ahead of a wall, followed by sliding into the wall. The controlling parameters for the target area of this mode are the width of the wall normal to the skid path (augmented by the wingspan of the striking aircraft), and the distance the aircraft will slide.

b. The roof or plan area (augmented by the wingspan of the striking aircraft), which may be projected to an equal ground area regardless of the angle of flight path slope.

c. The wall shadow area, (Bh) cot (A), which is the ground area projected by a wall of vertical dimension (h), width (B) normal to the flight path (and augmented by the CHAPTER 03 3.5-14 REV. 21, SEPTEMBER 2022

LGS UFSAR wingspan of the striking aircraft), with aircraft glide path making an angle (A) with the ground.

d. Shadowing of target structures by other structures.

Skid lengths are calculated from the equation provided in Reference 3.5-5:

Ls = 6.3x10-6 V2 (EQ. 3.5-3) k where:

Ls = skid length (miles) k = roughness factor (set equal to 2.5)

V = velocity (mph)

The glide angle is assumed to be 15 for fixed-wing aircraft. For rotary-wing aircraft, it is assumed to be 0 because of the poor aerodynamic characteristics of such aircraft. The skid length for rotary-wing aircraft is therefore considered to be zero feet.

The impact velocity for single-engine and twin-engine general aviation aircraft is estimated to be 110 mph. Therefore, from Equation 3.5-3, the skid length is 0.0305 miles or 161 feet. The impact velocity of a Boeing 727, assumed to be the average aircraft in the Federal airways, as discussed in Section 3.5.1.6.2.1, is taken as 290 mph, yielding a skid length of 0.212 miles or 1120 feet.

The analysis assumes that single-engine aircraft have a wingspan of 32 feet, twin-engine aircraft have a wingspan of 50 feet, and the wingspan of a Boeing 727 is 110 feet.

Pottstown-Limerick and Pottstown Municipal Airports As discussed in Section 3.5.1.6.1, the reactor enclosure, control structure, and spray pond pump structure are designed to withstand the impact of the largest aircraft capable of operating out of the nearby airports, the design basis Learjet, and are therefore not considered in the calculation of critical target areas.

The spray pond itself is not susceptible to aircraft damage because it is constructed completely in excavation, as discussed in Section 2.5. Aircraft damage to the spray pond spray network is not considered because the design provides for separation and redundancy so that the loss of a network would not affect the systems safety function.

The diesel generator enclosures for Units 1 and 2 each have an east-west dimension of 108 feet, a north-south dimension of 84.5 feet, and a height of 29 feet. The two structures are separated by 54 feet in the east-west direction.

Flight paths of concern for the diesel generator enclosures are those in the east, west, and northbound directions. There are no intervening structures for northbound aircraft. The critical CHAPTER 03 3.5-15 REV. 21, SEPTEMBER 2022

LGS UFSAR target area for single-engine aircraft is 3.51x10-3 square miles and 4.01x10-3 square miles for twin-engine aircraft.

For westbound traffic, the reactor enclosures and the administration building provide shadowing to the diesel generator enclosures. The reactor enclosures, located north of the diesel generator enclosures, reduce the effective target area for westbound aircraft to 84.5 feet plus one-half the aircraft's wingspan. The warehouse and shop area, located to the east of the diesel generator enclosures, effectively eliminates the skid area. The resultant critical target areas for single-engine aircraft is 1.36x10-3 square miles, and 1.49x10-3 square miles for twin-engine aircraft.

For eastbound traffic, shadowing to the diesel generator enclosures is provided by the reactor enclosures and the radwaste enclosures. The resultant critical target areas are 1.68x10-3 square miles for single-engine aircraft and 1.86x10-3 square miles for twin-engine aircraft.

For southbound traffic, the critical target area for diesel generators is zero due to the shadowing afforded by the reactor enclosure.

For rotary-wing aircraft, the critical target area of the diesel generator enclosure is 6.55x10-4 square miles.

LGS Plant Site Heliport The probability of a crashing helicopter impacting the diesel generator building is determined based on the known flight path of the helicopter traffic using the heliport, the closest distance this flight path comes to the diesel generator building, and the deviation from the flight path.

The possible deviation from the bounding traffic pattern for landing or takeoff path is represented by an exponential density function of the form:

f(x;) = e-x (EQ. 3.5-4) where:

f(x;)dx = the probability that the deviation is in the range (x, x+dx);

x = the perpendicular distance (deviation) from the flight path;

= the exponential distribution parameter.

The exponential distribution parameter, , is the inverse of the expected value of x. The expected value of x, x has been chosen as the mean distance which the helicopter can travel from an altitude (h). Therefore, x is given by:

x = h tan()

The angle is conservatively assumed to be the helicopter glide angle due to loss of a rotor. The expected altitude (h) of the helicopter on an approach or departure from the helipad is given by:

h = a tan()

CHAPTER 03 3.5-16 REV. 21, SEPTEMBER 2022

LGS UFSAR The angle , the approach angle, is conservatively assumed to be 15, and (a) is the distance from the helipad.

Using the above information, is given by:

= 1 a tan() tan()

The probability of impacting at a distance greater than (d) perpendicular to the flight path on one side of the flight path, is then given by the cumulative distribution function:

P(>d) = 0.5 e-ldl The probability of impact on the diesel generator enclosure is greatest when the helicopter is on the approach path 675 feet from the helipad and 1,100 feet laterally from the diesel generator enclosure. The average conditional probability of impacting the diesel generator enclosure from a crash for all three types of helicopters given in Table 3.5-1 is 7.66x10-4.

Federal Airways The reactor enclosure, control structure, and diesel generator enclosures are considered in the calculation of critical target areas. The spray pond pump structure represents only about 5% of the total target areas of safety-related structures, and its contribution to the overall probability is considered negligible.

The spray pond is not susceptible to aircraft damage because it is constructed completely in excavation, as discussed in Section 2.5. Aircraft damage to the spray pond spray network is not considered because the design provides for separation and redundancy so that the loss of a network would not affect the systems safety function.

Because the majority of airway traffic is northbound, the analysis considers the south faces of the diesel generator enclosures and the reactor enclosures as the target area. The calculated critical target area is 3.33x10-2 square miles.

3.5.1.6.3 Crash Probabilities Based on the given data and assumptions regarding the number of aircraft movements (Nij),

aircraft crash probability (Cij), and critical target areas for the plant (Aij), the probability of an aircraft striking to safety-related structures has been calculated.

For operations out of the Pottstown Municipal Airport, the probability of an aircraft striking the diesel generator enclosures is given by:

P = 1.2x10-8 * [1.73Ne + 1.40Nw + 3.66Nn]x10-3 yr-1 (EQ. 3.5-5) where:

Ne = number of eastbound movements over the site CHAPTER 03 3.5-17 REV. 21, SEPTEMBER 2022

LGS UFSAR Nw = number of westbound movements over the site Nn = number of northbound movements over the site and the target areas have been weighted to reflect the proportion of single and twin-engine traffic.

For Ne = Nn = 4000; Nw = 0, this yields a crash probability of 2.59x10-7 per year.

Similarly, for operations out of the Pottstown-Limerick Airport, the probability is given by:

P = 1.5x10-7 * [1.73Ne + 1.40Nw + 3.66Nn + 0.655Nr]x10-3 yr-1 (EQ. 3.5-6) where:

Nr = number of rotary-wing operations over the site from the Pottstown-Limerick Airport.

For Ne = 1500, Nw = 13500, Nn = 0; Nr = 0, this yields a crash probability of 3.2x10-6 per year.

For operations out of the LGS plant site heliport, the probability of a helicopter striking the diesel generator enclosure is the product of the number of takeoffs and landings per year, the probability of the helicopter crashing during takeoff or landing, and the probability that the crashing helicopter will impact the diesel generator building:

P = 312

• 4.88x10-7

• 7.66x10-4

= 1.17x10-7 per year The probability of an aircraft in one of the airways near the site striking the diesel generator enclosures, or the reactor enclosures is given by:

P = [(8395 + 20240)(590) + (46355)(22) + (4745)(7.1)]x10-13

• 0.0333

= 5.98x10-8 yr-1 (EQ. 3.5-7) 3.5.1.6.4 Probability of Aircraft Strike Resulting in Unacceptable Consequences CHAPTER 03 3.5-18 REV. 21, SEPTEMBER 2022

LGS UFSAR Given the aircraft impact resistant design of all safety-related structures except the diesel generator enclosures, an aircraft strike could jeopardize safe shutdown capability or have potential offsite dose consequences only under the following scenarios:

(1) The impacting aircraft is larger in size than the design basis aircraft described in Section 3.5.1.6.1, and the impact causes structural failure and/or damage to systems required to achieve or maintain safe shutdown, resulting in the possibility of offsite consequences in excess of 10CFR50.67 guidelines.

(2) The impacting aircraft strikes a diesel generator enclosure, damaging more than 2 diesel generators, thereby leaving that unit dependent on offsite power for achieving and maintaining shutdown. During the period required to return at least 2 diesel generators to service, a complete loss of offsite power occurs, which leaves the affected unit dependent on its Class 1E battery system to achieve and/or maintain safe shutdown. If offsite power is not restored before the batteries are exhausted, inadequate core cooling and offsite radiological consequences in excess of 10CFR50.67 guidelines could result.

The probabilities of scenarios (1) and (2) occurring are calculated as follows:

Scenario 1: Impact By Aircraft Larger Than Design Basis Aircraft This scenario is credible only for aircraft in Federal airways. It is conservatively assumed for this analysis that the impact of any aircraft using the Federal airways causes damage such that offsite radiological consequences in excess of 10CFR50.67 occur. The probability of such an impact causing unacceptable radiological consequences is therefore equal to the crash probability of an aircraft in one of the airways near the site, shown in Section 3.5.1.6.3 to be 5.98x10-8 per year.

Scenario 2: Diesel Generator Damage Caused By Impact Of Design Basis or Smaller Aircraft The probability that aircraft operating out of the Pottstown- Limerick Airport, the Pottstown Municipal Airport, and the LGS plant site heliport will impact the diesel generator enclosures is shown in Section 3.5.1.6.3 to be 3.2x10-6 per year, 2.59x10-7 per year, and 1.17x10-7 per year respectively, or a total probability, Pac = 3.58x10-6 per year.

Given that each unit's diesel generator enclosure is 108 ft wide, and that each of the 4 diesels per unit is located in a separate compartment inside that unit's diesel generator enclosure, it is highly unlikely that the impact of a typical twin-engine plane with a wingspan of 50 ft could disable more than 2 diesel generators.

In addition, only 2 of a unit's 4 diesel generators are required to achieve and maintain safe shutdown, assuming no additional failures. Despite the above, it is conservatively assumed for this analysis that any aircraft impacting a diesel generator enclosure disables all 4 diesel generators therein, leaving that unit dependent on offsite power for achieving and maintaining safe shutdown.

It is further conservatively assumed that immediate action would be taken to repair the diesel generators, and that at least 2 of the 4 would be restored to service 1 year following the aircraft impact.

CHAPTER 03 3.5-19 REV. 21, SEPTEMBER 2022

LGS UFSAR Using the methodology of appendix III of Reference 3.5-10, the probability, PLOOP, that a LOOP occurs (during the 1 year in which the affected unit would have no onsite ac backup power) is conservatively determined to be 0.2 per year, thus leaving that unit dependent on its Class 1E dc battery system to achieve or maintain shutdown.

It has been shown (Reference 3.5-11) that under loss of all ac power (station blackout) conditions, in a BWR with electrical and mechanical system designs similar to LGS, the batteries could power all systems required to achieve or maintain safe shutdown for as long as 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br /> before exhaustion. For this analysis, it is conservatively assumed that battery exhaustion occurs 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> after the LOOP occurs, after which inadequate core cooling and offsite radiological consequences in excess of 10CFR50.67 could occur.

Again using the methodology of appendix III of Reference 3.5-10, the probability, P4, that at least one offsite source of power is not returned to service within 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> following the LOOP is conservatively estimated to be 0.05 per LOOP event.

Based on the preceding information, the probability per year of an aircraft crash that disables one unit's diesel generators, leading to offsite radiological consequences in excess of 10CFR50.67 dose limits, can be calculated as follows:

P = PAC

• PLOOP

• P4

= (3.58x10-6 yr-1)(0.2 yr-1)(0.05) (EQ. 3.5-8)

= 3.58x10-8 per year.

Conclusion The total probability per year of an aircraft impact resulting in offsite radiological consequences in excess of 10CFR50.67 dose limits is the sum of the probabilities associated with Scenarios 1 and 2, or 9.56x10-8 per year. Realistically, this probability would be lower.

It is concluded that the acceptance criteria of SRP section 3.5.1.6 are met.

3.5.2 SYSTEMS TO BE PROTECTED Safety-related systems and structures are reviewed for missile protection and are listed in Table 3.2-1. Structures and barriers designed to provide for protection from external missiles are discussed in Section 3.5.1 and listed in Table 3.5-7, and their characteristics are listed in Table 3.5-8.

As discussed in Section 3.5.1, the only postulated internally generated missiles are those arising from failure of the HPCI and RCIC pump turbines. In this case, the pump-room walls are designed as barriers to isolate such missiles from other safety-related components.

The locations of safety-related equipment and the arrangement of the various compartments are shown in drawings M-110, M-111, M-112, M-113, M-114, M-115, M-116, M-117, M-118, M-119, M-120, M-121,M-122, M-123, M-124, M-125, M-126, M-127, M-128, M-129, M-130, M-131, M-132, CHAPTER 03 3.5-20 REV. 21, SEPTEMBER 2022

LGS UFSAR M-133, M-134, M-135, M-136, M-137, M-138, M-140, M-141, M-142, M-143, M-144, M-145, M-146, M-388, M-389, and M-390.

3.5.3 BARRIER DESIGN PROCEDURES Safety-related components necessary for safe shutdown and housed in the Category I structures are adequately protected against tornado missiles. The characteristics of structural wall barriers for Category I structures are shown in Table 3.5-8. Figure 3.5-2 shows typical reinforcement provided in the southernmost reactor enclosure concrete wall barrier.

The structures and barriers are designed to resist missile hazards in accordance with the procedures detailed in Reference 3.5-6.

The procedures include:

a. Prediction of local damage (penetration, perforation, and spalling) in the impact area, including estimation of the depth of penetration

b. Estimation of the barrier thickness required to prevent perforation

c. Prediction of the overall structural response of the barrier, and portions thereof, to missile impact Tornado missiles, as described in Table 3.5-4, were considered as deformable upon impact, and the structures were considered as rigid targets. Eighteen inch thick concrete, with a compressive strength of 4000 psi, was used in concrete walls and roofs subject to tornado missile impact. This satisfies the penetration formulae for elements subject to missiles as specified in Reference 3.5-6.

The ESW and RHRSW piping located within the yard area are installed underground with adequate cover for missile protection. A detailed assessment of soil cover for Category I yard piping shows a minimum depth of 4 feet, with most soil coverings exceeding 6 feet. The 4 foot depth was found to be adequate for tornado missiles (Table 3.5-4) in accordance with the criteria set forth in Reference 3.5-3.

The physical routing of the piping and typical profiles showing soil cover and installation details are shown in Figure 2.5-37. The cementitious backfill for pipe bedding is shown in detail 3 of Figure 2.5-37. This backfill was used predominantly for Category I piping to provide additional missile protection. The properties of cementitious backfill are defined in Section 2.5.4.5.4. Figure 2.5-37 shows the 4 foot minimum depth dimension of soil cover for Category I yard piping.

Doors are designed for tornado missiles unless internal secondary barriers are provided to protect safety-related components. Louvers are not designed for tornado missiles; however, internal secondary barriers are provided to protect safety-related components which are located behind the louvers. Figure 3.5-3 shows a typical secondary barrier for the south wall of the diesel generator enclosure (concrete labyrinth behind the door) used for tornado missile protection.

CHAPTER 03 3.5-21 REV. 21, SEPTEMBER 2022

LGS UFSAR 3.

5.4 REFERENCES

3.5-1 Letter, E.A. Hughes (GE) to R.C. DeYoung (NRC), "GE Recirculation Pump Potential Overspeed", (January 18, 1977).

3.5-2 NUREG/CR-2944, "Tornado Damage Risk Assessment", Reinhold & Ellingwood, Brookhaven National Lab, (September 1982).

3.5-3 "Depth Prediction for Earth-Penetrating Projectiles", Soil Mechanics and Foundation Division, ASCE, p. 6558, (May 1969).

3.5-4 "Probability of Missile Generation in General Electric Nuclear Turbines", GE Proprietary Report.

3.5-5 D.C. Gonyea, "An Analysis of the Energy of Hypothetical Wheel Missiles Escaping from Turbine Casings", GE Technical Information Series No. DF735L12, (February 1973).

3.5-6 "Design of Structures for Missile Impact", BC-TOP-9A, Revision 2, Bechtel Power Corporation, San Francisco, California, (September 1974).

3.5-7 Solomon, K.A., "Hazards Associated with Aircraft and Missiles", presented at American and Canadian Nuclear Society Meeting, Toronto, Canada, (June 1976).

3.5-8 Solomon, K.A., "Estimate of probability that an Aircraft will impact the PVNGS",

NUS-1416, NUS Corp., (June 1975).

3.5-9 National Air Transportation Safety Board, "Annual Review of Aircraft Accident Data", (Published 1972 and annually thereafter).

3.5-10 ASH-1400, "Reactor Safety Study: An Assessment of Accident Risks in U.S.

Commercial Nuclear Power Plants", NRC, (October 1975).

3.5-11 NUREG/CR-2182, Vol. I, "Station Blackout at Browns Ferry Unit One - Accident Sequence Analysis, Oak Ridge National Laboratory", (November 1981).

3.5-12 Barber, R.B., "Steel Road/Concrete Slab Impact Test (Experimental Simulation)",

Bechtel Power Corporation, (October 1973).

3.5-13 F.A. Vasallo, "Missile Impact Testing of Reinforced Concrete Panels", prepared for Bechtel Power Corporation, Calspan Corp., (January 1975).

3.5-14 "Safety Evaluation Report Relates to the Operation of Hope Creek Generating Station", NUREG-1048, Supplement 6, (July 1986).

CHAPTER 03 3.5-22 REV. 21, SEPTEMBER 2022

LGS UFSAR 3.5-15 Missile Analysis Report, Limerick Units 1 and 2," Siemens Energy, Inc., CT-27554, November 29, 2016 (including, as Appendix A: "Missile Probability Analysis Methodology for Limerick Generating Station, Units 1 and 2, with Siemens Retrofit Turbines Siemens Power Corporation Engineering Report - ER-9605, Revision 2, June 18, 1997 - SIEMENS PROPRIETARY) 3.5-16 "Westinghouse Limerick Generating Station Unit 1 & 2 Turbine Control System Upgrade Reliability Analysis and Electronic Overspeed Protection Fault Tree Analysis, Revision 4, November 2020."

3.5-17 LEAM-0017, Revision 1, Limerick Supplemental Turbine Missile Probability Analysis."

CHAPTER 03 3.5-23 REV. 21, SEPTEMBER 2022

LGS UFSAR Table 3.5-1 AIRCRAFT IMPACT DESIGN PARAMETERS GROSS WEIGHT APPROACH VELOCITY KINETIC ENERGY AIRCRAFT DESCRIPTION (lb) (knots) (ft/sec.) (K-ft)

Single reciprocating <5,000 110 186 2,686 engine Cessna (twin recip 7,250 111 188 3,979 engine)

Cessna jet (twin gas 10,500 115 194 6,136 turbine)

Learjet(1) (twin gas 13,500 139 235 11,577 turbine)

Navajo Chieftrain 7,000 100 169 3,104 Beech King Air-200 12,500 103 174 5,877 Cheyenne Turbo-Prop 11,200 115 194 6,545 Bell Long Ranger 4,150 50 85 466 (helicopter)

Aerospatiale Twin 5,291 50 85 594 Star (helicopter)

Bell 222 Model A 7,850 70 118 1,697 (helicopter) (takeoff)

(1) Largest aircraft capable of operating out of nearby airports; none are presently known to exist.

CHAPTER 03 3.5-24 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-2 CHARACTERISTICS OF LEARJET Resultant Load on Walls(2)

Assumed Mass Impact Peak Aircraft Weight (Kip- K.E. Momentum Duration Force Components(1) (lbs) sec2/ft) (Kip-ft) (Kip-sec) (sec) (Kips)

Fuselage 7,178 0.2229 6,155 52.3 0.10 846 Tip Tanks 1,358 0.0422 1,165 9.9 0.05 396 Wing 3,503 0.1088 3,005 25.6 0.10 512 Engines 1,158 0.0360 995 8.5 0.05 340 Empennage 305 0.0095 260 2.2 0.05 88 Total 13,500 0.4194 11,580 98.5 - (2)

(1)

All components are assumed traveling at 235 fps (139 knots) at impact.

(2)

Resultant load curves on structure due to the different components and total are shown in Figure 3.5-1.

CHAPTER 03 3.5-25 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-3 ACCUMULATORS LOCATED IN SAFETY-RELATED AREAS Maximum Maximum Pressure Temperature Description (psig) (F)

MSRV Accumulators 125 150 MSIV Accumulators 125 150 Diesel Air Start Reservoirs 250 120 Instrument Gas Receivers 125 135 CHAPTER 03 3.5-26 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-4 TORNADO-GENERATED MISSILE PARAMETERS IMPACT HORIZONTAL(4) KINETIC WEIGHT AREA VELOCITY ENERGY MISSILE (lb) (ft2) (mph)/(ft/sec) (ft-lb)

1. Wood plank (4"x12"x12') 200 0.333 300/440 6.01x105

2. Steel pipe (3" dia x 10',

schedule 40)(3) 78 0.067 144/211 5.39x104

3. Automobile(2)(3) 4000 20 72/106 6.98x105

4. Steel rod (1" dia x 3') 8 0.007 216/317 1.25x104

5. Utility pole (131/2" dia x 35',

not more than 30' 1490 1.266 144/211 1.03x106 above all grade elevations within 1/2 mile of the plant)

6. Steel pipe (6" dia x 15',

schedule 40)(1) 285 0.239 144/211 1.97x105

7. Steel pipe (12" dia x 15',

schedule 40)(1) 743 0.886 144/211 5.14x105 (1) The design basis for LGS included only missiles 1, 2, 3, 4 and 5. All safety-related structures and openings in structures have been assessed for the effects of missiles 6 and 7.

(2) LGS was originally designed for a postulated automobile missile not more than 25 ft above grade for all safety-related structures. All safety-related structures have been reassessed for the effect of the automobile at elevations up to 30 ft above all grade levels within 1/2 mile of the plant.

(3) LGS was originally designed for postulated missile velocities equal to 100 mph for the 3 in diameter steel pipe and 50 mph for the automobile. All safety-related structures have been reassessed for the revised velocities shown on the table.

(4) These missiles are considered to be capable of striking in all directions with vertical velocities equal to 80% of the horizontal velocities.

CHAPTER 03 3.5-27 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-5 RAILROAD ACCIDENT GENERATED MISSILE PARAMETERS WEIGHT INITIAL VELOCITY MISSILE (lb) (mph)

Part of tank car tank (112" dia x 62,000 200 (during ignition, traveling 50')(1) end on)

Steel pipe (3" dia x 10' 78 200 (tumbling) schedule 40)

Rail (3' long, 130 lb/yd) 130 200 (tumbling)

Type E coupler (7 in2 shank) 200 200 (tumbling)

Unfused bomb 750 100 (1)

These have been observed being propelled by burning gases at a very low trajectory not exceeding 30-35 feet above the ground. The elevation difference (77 feet) between the railroad and safety-related structures prevents impingement of this missile, and therefore it was not considered in the design.

CHAPTER 03 3.5-28 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-6 POTENTIAL HVAC FAN MISSILES Fan No. Description Area Elevation Notes OAV-163, SGTS exhaust fans 8 350 (2)

OBV-109 OAV-127, Control room emergency fresh 8 304 (1)

OBV-127 air supply fans OOV-126 Toilet room exhaust air fan 8 332 (2) 1AV-206, Reactor enclosure equipment 15,16, 313 (2) 1BV-206, compartment air exhaust fans 17,18 2AV-206, 2BV-206 OAV-132, Drywell purge exhaust fans 8 350 (2)

OBV-132 OAV-131, SGTS room exhaust fans 8 350 (2)

OBV-131 1AV-213, Reactor enclosure air 15,18 313 (2) 1BV-213, recirculation fans 2AV-213, 2BV-213 OAV-120, Auxiliary equipment room air 8 304 (2)

OBV-120 return/exhaust fans OAV-121, Control room ac return/exhaust 8 304 (2)

OBV-121 fans 1AV-512 Diesel generator enclosure DG Bldg (3) thru ventilation air exhaust fans 1HV-512 and 2AV-512 thru 2HV-512 CHAPTER 03 3.5-29 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-6 (Cont'd)

Fan No. Description Area Elevation Notes 1AV-201 Refueling floor air supply fans 15,18 313 (2) thru 1CV-201 and 2AV-201 thru 2CV-201 1AV-202 Reactor enclosure air supply fans 16,17 313 (2) thru 1CV-202 and 2AV-202 thru 2CV-202 1AV-204 Refueling floor air exhaust fans 16,17 313 (2) thru 1CV-204 and 2AV-204 thru 2CV-204 1AV-205 Reactor enclosure air exhaust fans 16,17 331 (2) thru 1CV-205 and 2AV-205 thru 2CV-205 OAV-124, Battery room air exhaust fans 8 304 (2)

OBV-124 OAV-114, Auxiliary equipment room 8 304 (2)

OBV-114 supply ac units OAV-116, Control room supply ac units 8 304 (2)

OBV-116 OAV-118, Emergency switchgear and 8 217 (2)

OBV-118 battery room ac units CHAPTER 03 3.5-30 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-6 (Cont'd)

Fan No. Description Area Elevation Notes 1AV-208, RCIC pump-room unit coolers 15,18 177 (3) 1BV-208, 2AV-208, 2BV-208 1AV-209, HPCI pump-room unit coolers 15,18 177 (3) 1BV-209, 2AV-209, 2BV-209 1AV-210 RHR pump-room unit coolers 15,16, 177 (3) thru 17,18 1HV-210 and 2AV-210 thru 2HV-210 1AV-211 CS pump-room unit coolers 11,12, 177 (3) thru 13,14 1HV-211 and 2AV-211 thru 2HV-211 1AV-212 Drywell area unit coolers Drywell 217 (4) thru 1HV-212 and 2AV-212 thru 2HV-212 OAV-140, SGTS room unit coolers 8 332 (3)

OBV-140 OAV-141, SGTS room access area unit 8 332 (3)

OBV-141 coolers OAV-543 Spray pond pumphouse air Spray (2)

Thru supply units pond ODV-543 00V900 CRD Decontamination 15 253 (3)

Room Unit Cooler CHAPTER 03 3.5-31 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-6 (Cont'd)

(1)

These potential missiles are installed in equipment that is only used during accident situations. They are not considered credible potential missiles because internal missile generation is not postulated to occur concurrently with other accidents.

(2)

These potential missiles are either remote from or separated by adequate barriers from all essential systems. Therefore, essential systems are protected from these potential missiles.

(3)

The potential missiles cannot impact on more than a single component. Therefore, redundant equipment will be available to effect safe shutdown.

(4)

The casing thickness is sufficient to prevent penetration by a loose blade. (During the September 24, 1982 Palo Verde event where the description indicated that a portion of a fan blade cut the liner plate, investigation revealed that the portion of blade escaped through a neoprene-coated fiberglass flexible connection between the fan discharge connection and the downstream duct. Pieces of blade struck the fan casing and the downstream duct. The duct was dented but no penetration of the fan casing or duct by the blade occurred. In the case of the LGS drywell unit coolers, the fan discharge connections are connected to a Class I Seismic Y duct fitting directly. Downstream of the fitting, stainless steel bellows with sliding metal flow liners are utilized rather than flexible fabric connections, thus precluding blade penetration.)

CHAPTER 03 3.5-32 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-7 MISSILE BARRIERS AND PROTECTED COMPONENTS FOR EXTERNALLY GENERATED MISSILES PROTECTED COMPONENTS MISSILE BARRIER Reactor coolant and other Primary containment safety-related equipment structure, reactor inside the containment enclosure walls, refueling cavity walls, internal structures and beams Control room, cable spreading Control structure room, switchgear, batteries, SGTS ECCS and other safety-related Reactor enclosure and equipment outside the primary internal walls containment Standby diesel generators Diesel generator enclosure(1)

Diesel fuel oil tank Earth; the tank is buried underground Spent fuel pool Reactor enclosure and fuel pool walls ESW pumps and RHRSW pumps Spray pond pump structure ESW and RHRSW yard piping Earth; piping is buried (1)

The diesel generator enclosure is not designed for design aircraft accident.

CHAPTER 03 3.5-33 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-8 CHARACTERISTICS OF EXTERNALLY GENERATED MISSILE BARRIERS MINIMUM MINIMUM CURING THICKNESS STRENGTH TIME BARRIER (in) (psi) (days)

Primary containment 74 4000 28 Reactor enclosure Walls 24 4000 28 Roofs 18 4000 28 Refueling cavity walls 48 4000 28 Spent fuel pool walls 54 4000 28 Control structure Walls 24 4000 28 Roof 18 4000 28 Diesel generator enclosure Walls 24 4000 28 Roof 24 4000 28 Spray pond pump structure Walls 24 4000 28 Roof 24 4000 28 (1)

On which strength is based CHAPTER 03 3.5-34 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.5-9 TURBINE SYSTEM RELIABILITY CRITERIA PROBABILITY, YR-1 UNFAVORABLY ORIENTED CRITERION TURBINE REQUIRED ACTION (A) P1 < 10-5 This is the general, minimum reliability requirement for loading the turbine and bringing the system on-line.

(B) 10-5 < P1 < 10-4 If this condition is reached during operation, the turbine may be kept in service until the next scheduled outage, at which time the licensee is to take action to reduce P1 to meet the appropriate A criterion (above) before returning the turbine to service.

(C) 10-4 < P1 < 10-3 If this condition is reached during operation, the turbine is to be isolated from the steam supply within 60 days, at which time the licensee is to take action to reduce P1 to meet the appropriate (A) criterion (above) before returning the turbine to service.

(D) 10-3 < P1 If this condition is reached at any time during operation, the turbine is to be isolated from the steam supply within 6 days, at which time the licensee is to take action to reduce P1 to meet the appropriate (A) criterion (above) before returning the turbine to service.

CHAPTER 03 3.5-35 REV. 13, SEPTEMBER 2006

LGS UFSAR 3.6 PROTECTION AGAINST DYNAMIC EFFECTS ASSOCIATED WITH THE POSTULATED RUPTURE OF PIPING This section describes the treatment of postulated ruptures in high energy and moderate energy piping located both inside and outside of the primary containment. The methods used to determine pipe rupture locations and to analyze the results of the ruptures, including jet thrust forces, jet impingement forces, piping dynamic responses, and compartment pressure-temperature transients, are described. Description is also provided for the design measures that have been implemented to ensure that the effects of any postulated rupture do not result in the loss of a required function that is necessary to mitigate the consequences of that pipe rupture.

The definitions of certain terms used within this section are provided in Section 3.6.3. The use of the term "low reactor level" in this section is generic rather than referring to a specific isolation signal set point.

NOTE: The computer program "COPDA", as described in Bechtel Topical Report BN-TOP-4 (Reference 3.6-1), was originally used to determine compartment pressure and temperature resulting from line breaks in Unit 1. The computer program "FLUD" (Reference 3.6.3), which replaced COPDA, was typically used for the same analysis in Unit 2. The mainframe program "FLUD" was later replaced in 1985 with a personal computer version "PCFLUD (Reference 3.6-12). The "PCFLUD" program had additional features added in 1993 (hot pipes, multiple HVAC fans, and "CONCOIL" room air cooler routine) and was renamed "CONCOIL-FLUD" or "CFLUD" (Reference 3.6-13). The supporting calculations for UFSAR Table 3.6-7 must be consulted for information regarding which specific program was used.

3.6.1 POSTULATED PIPING FAILURES IN FLUID SYSTEMS The failure of piping containing high energy or moderate energy has the potential of causing damage to surrounding structures, systems, and components. Depending on the fluid system involved and the rupture location, postulated piping failures can result in one or more of the effects described below. Essential systems and components are protected from these effects unless it can be demonstrated that their function is not impaired.

Pipe Whip Pipe whip is the unrestrained movement of a pipe due to the reaction force imposed on the pipe by fluid discharging from a rupture. Protection against pipe whip can be provided by intervening structural members between high energy piping and the essential systems and components, by providing pipe whip restraints on the high energy piping, or by locating essential systems and components sufficiently distant from high energy piping. Examples of typical pipe whip restraints and bumpers are shown in Figure 3.6-1.

Jet Impingement The blowdown of fluid from a rupture of a high energy pipe can exert forces on nearby equipment that could be high enough to cause damage to the equipment. Protection against jet impingement can be provided by installing jet impingement barriers to deflect the blowdown jet, or by locating essential systems and components a sufficient distance from high energy piping.

CHAPTER 03 3.6-1 REV. 18, SEPTEMBER 2016

LGS UFSAR Environmental Effects Pipe ruptures in high energy and moderate energy lines release fluid that can increase temperature, pressure, and humidity in the vicinity of the pipe rupture and also in remote areas that communicate with the local atmosphere. Essential systems and components may be exposed to abnormal conditions that have the potential of degrading the capability of the equipment to perform its function.

Piping systems whose rupture might generate hazardous environmental conditions are generally located in compartments that are capable of being isolated from other compartments containing essential systems and components. Isolation of compartments that enclose high energy lines is provided by maintaining normally closed access ways, automatically isolating ventilation duct-work, and sealing penetrations through walls and slabs. Compartments are designed to withstand the maximum internal pressurization developed as a result of a pipe rupture. Essential systems and components are either located in areas not affected by pipe ruptures or are qualified for operation under the maximum environmental conditions that they may be subjected to as a result of pipe ruptures.

Figures 3.6-43 through 3.6-45 are provided to show additional details of a typical pipe break analysis performed in accordance with SRP 3.6.1 and BTP ASB 3-1. The pipe break chosen is a break of the HPCI steam supply line in the HPCI compartment.

Water Spray Water released from pipe ruptures is a hazard to certain equipment, particularly electrical components. In most cases, spatial separation is adequate to prevent water spray from reaching essential systems and components. Essential systems and components that can potentially be subjected to water spray are either designed to operate when wetted or will be provided with barriers to deflect the spray.

Flooding Significant ruptures of fluid system piping may result in flooding in the vicinity of the rupture and in compartments through which the released fluid drains. The flooding rate and the total fluid volume released are computed based on the rupture configuration, the service of the fluid system, and the time required to isolate the rupture. The plant's drainage system handles minor releases of fluid with no adverse effects on essential systems and components.

Compartments that contain fluid systems with the potential for major releases of fluid are designed to contain the flooding and prevent significant leakage to adjoining compartments containing essential systems and components. Essential systems and components are either located in areas not subject to significant flooding or are designed to operate in a flooded environment.

3.6.1.1 Design Bases Pipe breaks are postulated to occur in high energy fluid system piping in accordance with the criteria stated in Section 3.6.2.1.1. Pipe cracks are postulated to occur in moderate energy fluid system piping in accordance with the criteria stated in Section 3.6.2.1.2. The effects of these piping ruptures require special consideration to ensure the following:

CHAPTER 03 3.6-2 REV. 18, SEPTEMBER 2016

LGS UFSAR

a. The ability to shut the reactor down safely is maintained.

b. Primary containment integrity is maintained.

c. Resultant offsite doses are below the values of 10CFR50.67.

In analyzing the effects of postulated piping ruptures, the following assumptions are made:

a Each break in high energy fluid system piping or crack in moderate energy fluid system piping is considered separately as a single postulated initial event occurring during normal plant conditions.

b. Offsite power is assumed to be unavailable if a trip of the turbine-generator or the RPS is a direct consequence of the postulated piping rupture.

c. A single active component failure is assumed to occur in systems used to mitigate the consequences of the postulated piping rupture and to shut the reactor down, except as noted in item d. below.

d. Where the postulated piping rupture is assumed to occur in one of the two redundant loops of either the RHR system or the RHRSW system, single active failures of components in the other loop of that system need not be assumed.

These two systems are dual-purpose moderate energy systems powered from both onsite and offsite sources and are designed, constructed, and inspected to standards appropriate for nuclear safety systems.

e. All available systems, including those actuated by operator actions, may be employed to mitigate the consequences of the postulated piping rupture. In judging the availability of systems, the postulated piping rupture and its direct consequences, and the assumed single active component failure and its direct consequences are considered. The feasibility of carrying out operator actions is judged on the availability of ample time and adequate access to equipment for the required actions.

f. An unrestrained whipping pipe is considered capable of causing breaks in impacted piping of smaller nominal pipe size and developing through-wall leakage cracks in impacted piping of equal or larger nominal pipe size with thinner wall thickness, except where experimental or analytical data demonstrates the capability to withstand the impact without failure. If such experimental or analytical data are used, additional documentation will be submitted to the NRC for review.

g. The effect of the jets resulting from each postulated pipe break in high energy fluid systems is reviewed and is evaluated only when such jets may have an effect on specific safety-related components that are required to effect a safe shutdown, maintain containment integrity, provide post-LOCA monitoring, or limit radioactive releases to site-allowable limits, as a result of a specific postulated pipe break event.

CHAPTER 03 3.6-3 REV. 18, SEPTEMBER 2016

LGS UFSAR The effects of jet impingement on the piping required to effect a safe shutdown, maintain containment integrity, provide post-LOCA monitoring, or limit radioactive releases to site-allowable limits, regardless of size or wall thickness, are also reviewed. Bounding cases are evaluated to ensure that acceptable stress levels in such piping are not exceeded.

3.6.1.2 Description A listing of high energy fluid system piping is provided in Table 3.6-1. All other piping in the plant that is pressurized above atmospheric pressure is considered to be moderate energy piping. The routing of piping within the reactor enclosure and the primary containment is shown in Figures 1.2-40 through 1.2-72.

For each pipe rupture location determined in accordance with the criteria of Section 3.6.2.1, an analysis is performed using the assumptions of Section 3.6.1.1 to verify that the consequences of the pipe rupture are acceptable. These analyses are summarized below for high energy and moderate energy fluid systems.

3.6.1.2.1 High Energy Fluid Systems 3.6.1.2.1.1 Reactor Recirculation System The two reactor recirculation loops are located entirely within the primary containment and are arranged on opposite sides of the reactor pedestal and reactor shield wall. Pipe whip restraints anchored in the reactor pedestal and reactor shield wall are provided for the recirculation loops and are arranged as shown in Figure 3.6-2. This system of restraints prevents unrestrained pipe whip resulting from a postulated rupture at any of the identified break locations. The restraints consist of two basic components: the frame attached to a support member and straps attached to the frame (two straps per frame). Either carbon steel cables or stainless steel bars are used as straps. The restraints on the 28-inch recirculation loop piping utilize stainless steel bars. The restraints on the 22-inch discharge header and the 12-inch risers utilize carbon steel cables. A schematic detail of a restraint is shown in Figure 3.6-3.

A review of the potential consequences of jet impingement resulting from recirculation loop ruptures has shown that breaks at the lower end of two of the 12-inch risers in each recirculation loop could result in impingement on the CRD withdraw piping. An evaluation of the consequences of damage to CRD withdraw piping resulting from jet impingement has shown that such damage would not prevent the associated control rods from being inserted into the reactor as necessary to affect a reactor shutdown. Electrical cabling associated with essential systems and components is routed to avoid jet impingement from postulated ruptures of recirculation loop piping.

Pipe Break Locations The postulated pipe break locations for the recirculation loop piping are shown in Figure 3.6-4. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-2. Blowdown thrust time histories for each break location are provided in Table 3.6-3. The recirculation piping design basis has been evaluated for the effects of power rerate and demonstrated to be adequate for the increases in pressure, temperature and flow due to power rerate. Details of the evaluation performed are documented in Reference 3.6-11.

CHAPTER 03 3.6-4 REV. 18, SEPTEMBER 2016

LGS UFSAR Compartment Pressure-Temperature Transient The pressure-temperature transient in the primary containment resulting from a complete circumferential rupture of one recirculation loop is discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability The sequence of events that would occur automatically to shut the reactor down and cool the core in the event of a recirculation loop rupture is discussed in Section 6.3.3. A combination of pipe whip restraints, jet impingement barriers, and separation by distance is used to ensure the availability of sufficient equipment to accomplish these functions.

3.6.1.2.1.2 Main Steam System The four 26-inch main steam lines are routed as shown in Figure 5.1-7 for the portion inside the primary containment, and in Figure 3.6-5 for the portion outside the primary containment. The A and B steam lines are connected to the east side of the reactor vessel and the C and D steam lines are connected to the west side of the vessel. All four steam lines penetrate the north side of the primary containment. The portion of the reactor enclosure through which the main steam lines are routed (between the primary containment and the turbine enclosure) is referred to as the main steam tunnel and is separated from other areas of the reactor enclosure by concrete walls and slabs. Only piping, valves, and associated instrumentation are located in the main steam tunnel.

Figure 3.6-6 shows an elevation view of the main steam tunnel.

The following features are incorporated into the design of the main steam line and nearby structures to mitigate the consequences of a main steam line break or to minimize the probability of its occurrence:

a. A venturi-type flow restrictor is located in each main steam line inside the primary containment. The flow restrictor reduces the rate of loss of reactor coolant from a main steam line break for break locations downstream of the restrictor. (The flow restrictors are described in Section 5.4.)

b. Each main steam line is provided with either three or four MSRVs that reduce the probability of breaks by protecting the steam line against overpressurization. (The MSRVs are described in Section 5.2.)

c. Each main steam line is provided with two fast-acting fail-safe MSIVs, one upstream and one downstream of the primary containment penetration. These valves close automatically upon receipt of signals indicating high steam flow or high temperature in the vicinity of the piping outside the primary containment (as well as upon receipt of other initiating signals), in order to terminate blowdown through breaks outside the primary containment. (The MSIVs are described in Section 5.4.)

d. Moment-limiting pipe whip restraints are located upstream of the inboard MSIVs and downstream of the outboard MSIVs in order to ensure the operability of these valves in the event of a main steam line break in the general vicinity of the valves.

The main steam lines are provided with pipe whip restraints inside the primary containment and in the main steam tunnel. Typical restraints inside the primary containment are shown in Figure CHAPTER 03 3.6-5 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6-1. Figure 3.6-6 shows the locations of the restraints in the main steam tunnel. As shown in Figure 3.6-7, the first two restraints downstream of the outboard MSIVs span between the east and west walls of the tunnel and restrain all four steam lines. Corbels extending out from the north wall of the main steam tunnel limit the possible upward movement of the upper elbow of each steam line in the tunnel. Additionally, the vertical portion of the steam line run in the tunnel is restrained against the north wall of the tunnel at two separate locations.

After entering the Unit 1 turbine enclosure from the main steam tunnel, the main steam lines are routed along the west side of the control structure before turning westward to run the length of the turbine enclosure. This arrangement is shown in Figure 3.6-5. In order to prevent a main steam line break in this portion of piping from causing pipe whip impact on the control structure wall, bumpers are provided between the control structure and the elbows of the two steam lines closest to the control structure (26" EBB-103 and 26" EBB-104). The Unit 2 arrangement is similar, but opposite hand.

In reviewing the potential consequences of jet impingement resulting from main steam line breaks, it was determined that breaks at the elbow at the lower end of the near-vertical piping run inside the primary containment could result in impingement on the CRD withdraw piping. An evaluation of the consequences of damage to CRD withdraw piping resulting from jet impingement has shown that such damage would not prevent the associated control rods from being inserted into the reactor core as necessary to affect a reactor shutdown. Electrical cabling associated with essential systems and components is routed to avoid jet impingement from postulated breaks of main steam piping.

Pipe Break Locations The postulated pipe break locations for the main steam piping, and also the pipe whip restraint locations, are shown in Figures 3.6-8 and 3.6-9. The calculated stress levels and usage factors, and the postulated break types, for the main steam piping inside the primary containment are listed in Table 3.6-4. Breaks for the main steam piping outside containment are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments. The main steam piping basis has been evaluated for the effects of power rerate and demonstrated to be adequate for the increases in pressure, temperature and flow due to power rerate. Details of the evaluation performed are documented in Reference 3.6-11.

Compartment Pressure-Temperature Transients The pressure-temperature transient in the primary containment resulting from a complete circumferential break of one main steam line is discussed in Section 6.2.1.

Protection against overpressurization of the main steam tunnel in the event of a main steam line break in the tunnel is provided by two sets of blowout panels. One set of blowout panels is located in the north wall of the tunnel and vents to the turbine enclosure. The second set of blowout panels is located in the vent stack that leads upward from the tunnel and discharges to the atmosphere above the top of the reactor enclosure. This vent stack can be seen in drawings M-123 and M-138.

A pressure-temperature transient analysis for the case of a main steam line break in the tunnel was performed using the analytical technique described in BN-TOP-4 (Reference 3.6-1) and the blowdown data provided in Table 3.6-6 and Figure 3.6-10. The flow model is shown in Figure 3.6-11 and the results of the analysis are listed in Table 3.6-7. The main steam tunnel is designed CHAPTER 03 3.6-6 REV. 18, SEPTEMBER 2016

LGS UFSAR to withstand the maximum pressure developed, and the MSIVs are qualified to operate under environmental conditions more severe than those calculated to occur.

Protection against overpressurization of the Unit 1 turbine enclosure as a result of a main steam line break within its boundaries is provided by blowout panels along the north exterior wall of the turbine enclosure. As shown in drawings M-111 and 112, these blowout panels are located between column lines 12 and 18 and vent the main condenser area of the turbine enclosure directly to the atmosphere. The main steam lines are routed in areas that communicate directly with the main condenser area for the entire length of their run within the turbine enclosure. The Unit 2 arrangement is similar.

A pressure-temperature transient analysis for the case of a main steam line break in the turbine enclosure was performed using the analytical technique described in Reference 3.6-1 and with the blowdown data provided in Table 3.6-6. The flow model is shown in Figure 3.6-12 and the results of the analysis are listed in Table 3.6-7. The turbine enclosure is designed to withstand the maximum pressure developed.

The secondary containment structure is capable of withstanding the effects of a high energy pipe rupture occurring inside the secondary containment without loss of integrity. For all pipe breaks except those in the RWCU system, the time-dependent mass and energy release rates are calculated using the Moody model including friction effects. Blowdown for the main steam line breaks is determined in accordance with the assumptions in Reference 3.6-14. A double-ended guillotine break is assumed, and both forward and reverse flows are included for all cases except for HPCI and RCIC line breaks in the HPCI and RCIC compartments, respectively. The postulated pipe breaks in the HPCI and RCIC compartments are at the upstream side of a normally closed isolation valve, therefore only forward flow is considered.

Mass and energy release rates for pipe breaks in the RWCU system are calculated with RELAP4/MOD5 using the Henry-Fauske model option for subcooled conditions and the Moody critical flow model option for saturated conditions.

Verification of Reactor Shutdown Capability Breakage of a main steam line inside the primary containment would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

For a main steam line break outside the primary containment, the MSIVs will be closed automatically because of high steam flow, low reactor water level, or high temperature in the vicinity of the main steam lines. Closure of the MSIVs will cut off steam flow to the feedwater pump turbines, causing the pumps to coast down and stop. The reactor will be scrammed by low reactor water level or closure of the MSIVs. After isolation of the reactor vessel, the RCS pressure will increase until the setpoint of the MSRVs is reached. Steam will then be automatically discharged to the suppression pool to limit the pressure rise.

Low reactor water level will initiate operation of the HPCI and RCIC systems to maintain reactor water level. If both the HPCI and RCIC systems are unavailable, the ADS will be automatically initiated to depressurize the reactor vessel so that the LPCI and core spray systems can inject CHAPTER 03 3.6-7 REV. 18, SEPTEMBER 2016

LGS UFSAR water into the vessel. These latter two systems are initiated automatically by low reactor water level, and will provide sufficient flow to restore reactor water level and cool the core.

After reactor water level has been restored, and the RCS pressure and temperature have decreased sufficiently, the shutdown cooling mode of the RHR system can be initiated to bring the reactor to cold shutdown.

A combination of pipe whip restraints, jet impingement barriers, and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of a main steam line break in either the primary containment or the main steam tunnel.

Essential systems and components located in these areas are qualified to operate under the environmental conditions resulting from a break. Since no essential systems and components are located in the turbine enclosure, no special provisions are necessary to provide protection for equipment in this area from the effects of pipe breaks. Blowout panels are provided to prevent overpressurization of the turbine enclosure, and bumpers are provided to prevent a whipping main steam line from impacting the control structure.

3.6.1.2.1.3 Feedwater System The discharge lines from the three feedwater pumps are routed into a common mixing header in the turbine enclosure. From this header, two parallel 24-inch feedwater lines enter the main steam tunnel and then penetrate the primary containment. Inside the primary containment, the two lines diverge to form symmetrical headers on opposite sides of the reactor vessel. Each header splits into three 12-inch risers that attach to the reactor vessel nozzles. The routing of the feedwater lines in the main steam tunnel and primary containment is shown in drawings M-234, M-305 and Figure 3.6-6.

Each feedwater containment penetration is provided with three check valves as containment isolation valves, one inside the drywell and two in the main steam tunnel. In the event of a feedwater line break outside the primary containment, these check valves will, as described below, close to prevent backflow from the reactor vessel. Thus, flow from the break would be from the feedwater pump side only.

The operability of the first outboard containment isolation valve (i.e. closest to the drywell penetration) is ensured in the event of a feedwater line break by providing moment-limiting whip restraints designed for this purpose. The operability of the check valve inside containment and the second outboard check valve need not be protected from overstress due to pipe whip, since only two isolation valves are necessary to provide adequate redundancy to ensure the containment isolation function, and it is not possible for one pipe break to affect both the inboard and second outboard check valves. More detailed discussion of feedwater check valve operability is provided in Section 3.9.3.2b.2.

The feedwater lines are provided with pipe whip restraints inside the primary containment, in the main steam tunnel, and in the turbine enclosure immediately outside the main steam tunnel.

Typical restraints inside the primary containment are shown in Figure 3.6-1. As shown in Figure 3.6-13, the restraint in the tunnel spans between the east and west walls of the tunnel and restrains both feedwater lines. Corbels extending out from the north wall of the main steam tunnel limit the possible upward movement of the upper elbows of the feedwater startup recirculation lines that connect to the feedwater lines in the main steam tunnel. Feedwater line restraints located in the turbine enclosure and which are attached to the control structure and reactor enclosure walls are CHAPTER 03 3.6-8 REV. 18, SEPTEMBER 2016

LGS UFSAR shown in Figure 3.6-14. Bumpers on the control structure walls protect the control structure from a whipping feedwater line.

Electrical cabling associated with other essential systems and components is routed to avoid jet impingement from postulated breaks of feedwater piping.

Pipe Break Locations The postulated pipe break locations for the feedwater piping, and also the pipe whip restraint locations, are shown in Figures 3.6-15 and 3.6-16. The calculated stress levels and usage factors, and the postulated break types, for the feedwater piping inside the primary containment are listed in Table 3.6-8. Breaks for the feedwater piping outside containment are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Compartment Pressure-Temperature Transients The pressure-temperature transient in the primary containment resulting from a break of any of the sizes of feedwater lines in the drywell is exceeded in severity by the transients resulting from recirculation loop and main steam line breaks, which are discussed in Section 6.2.1. The pressure-temperature transient resulting from a feedwater line break in the main steam tunnel or turbine enclosure is exceeded in severity by the transient resulting from a main steam line break, which is discussed in Section 3.6.1.2.1.2.

Verification of Reactor Shutdown Capability Breakage of a feedwater line inside the primary containment would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

For a feedwater line break outside the primary containment, differential pressure across the containment isolation check valves in the reverse direction will cause these valves to close rapidly, isolating the reactor vessel from the break. The loss of feedwater flow will cause the reactor water level to drop, initiating a reactor scram. Water level will continue to drop because of steam generation from decay heat, causing closure of the MSIVs as well as initiation of the RCIC and HPCI systems. Once the reactor has been scrammed and the RCS isolated, the sequence of events is similar to that for a main steam line break outside the primary containment.

A combination of pipe whip restraints, jet impingement barriers, and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of a feedwater line break in either the drywell or the main steam tunnel. Since no essential systems and components are located in the turbine enclosure, no special provisions are necessary to provide protection for equipment in this area from the effects of pipe breaks. Bumpers are provided in the turbine enclosure to protect the control structure from the impact of a whipping feedwater line.

3.6.1.2.1.4 Condensate System The condensate system is located entirely within the turbine enclosure. No pipe whip restraints are provided for the condensate piping.

CHAPTER 03 3.6-9 REV. 18, SEPTEMBER 2016

LGS UFSAR Pipe Break Locations Since the condensate system consists of non-nuclear class piping, breaks are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Compartment Pressure-Temperature Transients Since the normal fluid temperature in the condensate system is less than 135F, no significant pressure-temperature transient would result from condensate line breaks.

Verification of Reactor Shutdown Capability In the event of a condensate line break, the feedwater pumps would trip because of low suction pressure. The resulting loss of feedwater flow would result in closure of the containment isolation check valves, preventing reactor blowdown through the break. The sequence of events that would occur from this point on is the same as for breakage of a feedwater line outside the primary containment.

Since no essential systems and components are located in the turbine enclosure, no special provisions are necessary to provide protection for equipment in this area from the effects of pipe breaks. The flooding effects of a condensate line break are exceeded by the effects of a circulating water line expansion joint rupture in the turbine enclosure, a discussion of which is contained in Sections 10.4.1 and 10.4.5.

3.6.1.2.1.5 Reactor Water Cleanup System The RWCU system takes suction from reactor recirculation loop B via the RHR shutdown cooling suction line inside primary containment. The 6-inch RWCU suction line is routed up to el 297'-3", at which point it penetrates the primary containment. The RWCU piping is then routed through the various RWCU equipment compartments at el 283' and el 313' in the reactor enclosure, including the containment penetration compartment, three RWCU pump compartments, the regenerative heat exchanger compartment, two nonregenerative heat exchanger compartments, two RWCU holding pump compartments, and two RWCU filter/demineralizer compartments. The RWCU return piping is routed from the containment penetration compartment directly into the main steam tunnel, where the piping splits into two 4-inch lines. One line connects to the A feedwater line (DBB-103) and the other line connects to the RCIC injection line immediately upstream of its connection to the B feedwater line (DBB-104). Both of these connections to the feedwater lines are located between the two outboard containment isolation valves in the feedwater lines. Another RWCU line, the 4-inch RWCU alternate return line, has a branch connection to the RWCU return line at a point just upstream of the location where the return line splits into the two lines leading to the feedwater lines. From this branch connection, the RWCU alternate return line is routed downward to a containment penetration at elevation 261 feet, which is also in the main steam tunnel. The routing of this RWCU piping is shown in drawings M-210, M-211, M-225, M-231, M-232, M-235, M-296, M-302, M-303, M-306, M-322, and M-323.

The RWCU suction line is provided with two fast-acting isolation valves, one upstream and one downstream of the primary containment penetration. These valves close automatically upon receipt of signals indicating high differential flow (RWCU suction vs. return) or high temperature in CHAPTER 03 3.6-10 REV. 18, SEPTEMBER 2016

LGS UFSAR the vicinity of the piping outside the drywell (as well as upon receipt of other initiating signals), in order to terminate blowdown through breaks outside the drywell. In order to ensure the operability of these valves in the event of an RWCU suction line break, moment-limiting pipe whip restraints are located upstream of the inboard containment isolation valve and downstream of the outboard containment isolation valve. Whip restraints are also located on the portion of the RWCU suction line within the drywell and on the portion of the return line within the main steam tunnel.

In reviewing the potential consequences of jet impingement resulting from RWCU line breaks, it was determined that breaks near the connection of the RWCU return lines to the RCIC injection and feedwater line in the main steam tunnel could result in impingement on the outboard MSIV operators. A steel plate barrier is provided to protect the MSIV operators from this potential source of jet impingement. It was also determined that RWCU line breaks in the RWCU containment penetration compartment could result in impingement on the RWCU outboard containment isolation valve. A steel plate barrier is provided to protect the isolation valve from this potential source of jet impingement.

Pipe Break Locations The postulated pipe break locations for the RWCU pump suction piping inside the primary containment, and also the pipe whip restraint locations, are shown in Figure 3.6-17. The calculated stress levels and usage factors, and the postulated break types, for this portion of the RWCU piping are listed in Table 3.6-10. Piping isometrics for the RWCU pump suction piping in the RWCU isolation valve compartment and for the RWCU return piping in the main steam tunnel are shown in Figures 3.6-39 and 3.6-18, respectively. Breaks in the latter portions of the RWCU piping are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Compartment Pressure-Temperature Transients The pressure-temperature transient in the primary containment resulting from a break in the portion of the RWCU suction line in the drywell is exceeded in severity by the transients resulting from recirculation loop breaks, main steam line breaks, and small steam leaks, which are discussed in Section 6.2.1.

Protection against overpressurization of the RWCU equipment compartments in the reactor enclosure as a result of RWCU line breaks in these areas is provided by interconnecting steam venting paths between the various compartments and by blowout panels leading to the outside atmosphere. The two nonregenerative heat exchanger compartments are vented into the regenerative heat exchanger compartment via a short steam venting plenum. The regenerative heat exchanger compartment is vented in turn to the adjacent containment penetration compartment. Each of the three pump rooms is vented individually to the containment penetration compartment. The containment penetration compartment is vented, via a blowout panel, to an exhaust stack located on the west side of the reactor enclosure for Unit 1 and on the east side of the reactor enclosure for Unit 2. The stack is open to the atmosphere at its upper and lower ends.

The venting pathway from the containment penetration compartment to the exhaust stack is shown in drawings M-215 and M-316.

CHAPTER 03 3.6-11 REV. 18, SEPTEMBER 2016

LGS UFSAR Pressure-temperature transient analyses for the cases involving RWCU line breaks in the RWCU equipment compartments were performed using the analytical technique described in Reference 3.6-1 and with the blowdown data provided in Table 3.6-6. These blowdown data were developed using Reference 3.6-2. The flow model for breaks in the RWCU equipment compartments is shown in Figure 3.6-19, and the results of the analyses are listed in Table 3.6-7. No analyses are performed for the RWCU holding pump compartments and filter/demineralizer compartments since the piping in these areas contains fluid at a temperature of 150F or less, so that no significant pressure-temperature transient results from a pipe break.

The RWCU equipment compartments are designed to withstand the maximum pressure developed due to a pipe break. The outboard containment isolation valves for the RWCU system and other systems located in the containment penetration compartment are qualified to operate under environmental conditions more severe than those calculated to occur due to pipe break.

Verification of Reactor Shutdown Capability Breakage of the RWCU suction line inside the drywell would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

For an RWCU line break outside the drywell, the RWCU containment isolation valves will be closed automatically due to high flow in the RWCU suction line or high temperature in the RWCU equipment compartments. Backflow from the feedwater lines into the RWCU return line will be prevented by closure of the check valve in the return line. If the break occurs in the RWCU piping within the main steam tunnel, the MSIVs will close automatically due to high temperature in the tunnel. Once MSIV isolation has occurred, the sequence of events is similar to that for a main steam line break outside the primary containment. If the break occurs in the RWCU equipment compartments, no reactor scram or MSIV closure will occur, due to the isolation of the RWCU system and rapid termination of the blowdown. After an RWCU isolation, there is sufficient capability to shutdown the reactor. The operator will investigate the cause of the RWCU isolation and take appropriate actions, in accordance with the plant Technical Specifications and approved operating procedures. The Technical Specifications and/or the operating procedures, which include off normal and EOPs, may, based on plant conditions, require the operator to initiate a normal shutdown of the reactor.

A combination of pipe whip restraints, jet impingement barriers, and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of an RWCU line break occurring in the drywell, main steam tunnel, or RWCU equipment compartments. Essential systems and components located in these areas are qualified to operate under the environmental conditions resulting from the break. Among the RWCU equipment compartments, only the containment penetration compartment contains safety-related equipment:

the primary containment purge line, two of the four LPCI injection lines, and the RWCU outboard containment isolation valve. The containment isolation valves in the purge line and the LPCI injection lines are normally closed during reactor operation and are not required to operate after a RWCU line break outside primary containment; therefore they require no protection. The RWCU outboard containment isolation valve is protected from jet impingement by a steel plate barrier, and the cabling to the valve is routed so as to avoid jet impingement.

CHAPTER 03 3.6-12 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6.1.2.1.6 Reactor Vessel Drain The reactor vessel drain consists of 2-inch piping originating at a nozzle on the reactor vessel bottom head. From this location inside the reactor vessel pedestal, the piping increases to 21/2 inches in diameter, penetrates the pedestal into the general drywell area, increases to 4 inches in diameter, and is routed upward to its point of connection to the 6-inch RWCU suction line inside the drywell.

Pipe Break Locations The postulated pipe break locations for the reactor vessel drain line, and also the pipe whip restraint locations, are shown in Figure 3.6-20. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-12.

Compartment Pressure-Temperature Transients Since the reactor vessel drain line is located entirely within the drywell, breakage of this line would have no effect on plant areas outside the primary containment. The pressure-temperature transient in the primary containment resulting from a reactor vessel drain line break is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability Breakage of the reactor vessel drain line would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3. A combination of pipe whip restraints and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of a reactor vessel drain line break.

3.6.1.2.1.7 HPCI Steam Supply Line The HPCI steam supply piping has a nominal diameter of 10 inches for the portion inside the drywell and 12 inches for the portion outside the drywell. The supply line connects to main steam line C inside the drywell. From its connection to the main steam line, the HPCI steam supply line is routed downward and then horizontally along the drywell wall. The line then penetrates the drywell at el 244'-8", entering the isolation valve compartment located at floor el 217' in the reactor enclosure. The steam supply line penetrates the floor of the isolation valve compartment and enters the HPCI pump compartment located at el 177'. The routing of this line is shown in drawings M-213, M-225, M-226, M-227, M-228, M-295, M-297, M-296, M-318, M-319, M-229, and M-320. During normal reactor operation, the line is pressurized from main steam line C up to the HPCI turbine steam supply valve (HV-55-F001).

The HPCI steam supply line is provided with two fast-acting isolation valves, one upstream and one downstream of the primary containment penetration. These valves close automatically upon receipt of signals indicating high steam flow or high temperature in the vicinity of the piping outside the drywell (as well as upon receipt of other initiating signals), in order to terminate blowdown through breaks outside the drywell.

CHAPTER 03 3.6-13 REV. 18, SEPTEMBER 2016

LGS UFSAR Moment-limiting pipe whip restraints are located upstream of the inboard containment isolation valve and downstream of the outboard containment isolation valve in order to ensure the operability of these valves in the event of a break in the HPCI steam supply line near the valves.

Whip restraints are also located on the HPCI steam supply line and in the isolation valve compartment at el 217' in the reactor enclosure. A typical restraint inside the drywell is shown in Figure 3.6-1.

Pipe Break Locations The postulated pipe break locations for the HPCI steam supply line, and also the pipe whip restraint locations, are shown in Figure 3.6-21 for the portion of the line inside the drywell and in Figure 3.6-22 for the portion of the line outside the drywell. The calculated stress levels and usage factors, and the postulated break types, are listed in Tables 3.6-13 and 3.6-14.

Compartment Pressure-Temperature Transients The pressure-temperature transient in the primary containment resulting from a break in the portion of the HPCI steam supply line in the drywell is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Protection against overpressurization of the HPCI pump compartment and the isolation valve compartment at el 217' in the reactor enclosure as a result of HPCI steam supply line breaks in these areas is provided by steam venting paths and blowout panels leading to the outside atmosphere. The isolation valve compartment is vented to the atmosphere via blowout panels located on the south side of the reactor enclosure, as shown in drawings M-118, M-123 and M-138. The HPCI pump compartment is vented to the isolation valve compartment via hinged, metal plate blowout panels located in the floor at el 217'. These latter panels lift to relieve pressurization in the HPCI pump compartment but do not allow pressurization in the isolation valve compartment to result in steam flow in the reverse direction, i.e., down into the HPCI pump compartment.

Pressure-temperature transient analyses for the cases involving HPCI steam supply line breaks in the HPCI pump compartment and the HPCI piping area were performed using the analytical technique described in Reference 3.6-1 and the blowdown data provided in Table 3.6-6. The flow model for breaks in these two compartments is shown in Figure 3.6-23, and the results of the analyses are listed in Table 3.6-7. Figures 3.6-43 through 3.6-45 provide additional details of a typical pipe break analysis. The case of an HPCI steam supply line break in the isolation valve compartment was analyzed using the CFLUD computer program (Reference 3.6-13) and the blowdown data provided in Table 3.6-6. The flow model for this case is shown in Figure 3.6-24, and the results of the analysis are listed in Table 3.6-7.

The HPCI pump compartment and the isolation valve compartment are designed to withstand the maximum pressures developed due to pipe break. All equipment located in the isolation valve compartment that is required to operate following breakage of the HPCI steam supply line is qualified to operate under environmental conditions more severe than those calculated to occur due to pipe break. No equipment located in the HPCI pump compartment is required to operate following breakage of the HPCI steam supply line.

CHAPTER 03 3.6-14 REV. 18, SEPTEMBER 2016

LGS UFSAR Verification of Reactor Shutdown Capability Breakage of the HPCI steam supply line inside the drywell would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

For an HPCI steam supply line break outside the drywell, the steam supply line containment isolation valves will be closed automatically, terminating reactor vessel blowdown. No reactor scram will occur, due to the isolation of the HPCI steam line and rapid termination of the blowdown.

After a HPCI isolation, there is sufficient capability to shutdown the reactor. The operator will investigate the cause of the HPCI isolation and take the appropriate actions, in accordance with the plant Technical Specifications and approved operating procedures, to assure that the plant is in safe condition. The Technical Specifications and/or the operating procedures, which include off normal and EOPs, may, based on plant conditions, require the operator to initiate a normal shutdown of the reactor.

A combination of pipe whip restraints and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of an HPCI steam supply line break occurring in the drywell, the isolation valve compartment, or the HPCI pump compartment. Essential systems and components located in these areas are qualified to operate under the environmental conditions resulting from the break. Electrical cabling associated with essential systems and components is routed so as to avoid jet impingement from the postulated break.

No pipe whip restraints are provided for the portion of the HPCI steam supply line located within the HPCI pump compartment. Since no essential systems and components (other than the HPCI system itself) are located in the HPCI pump compartment, such protection is not necessary.

3.6.1.2.1.8 RCIC Steam Supply Line The RCIC steam supply line has a nominal diameter of 4 inches for the portion inside the drywell and 6 inches for the portion outside the drywell. The supply line connects to main steam line B inside the drywell. From its connection to the main steam line, the RCIC steam supply line is routed generally downward and then horizontally along the drywell wall. The line then penetrates the drywell at el 243'-6", entering the isolation valve compartment located at floor el 217' in the reactor enclosure. The steam supply line penetrates the floor of the isolation valve compartment and enters the RCIC pump compartment located at el 177'. The routing of this line is shown in drawings M-213, M-215, M-225, M-226, M-227, M-229, M-239, M-295, M-296, M-297, M-310, M-316, M-318, and M-320. During normal reactor operation, the line is pressurized from main steam line B up to the RCIC turbine steam supply valve (HV-50-F045).

The RCIC steam supply line is provided with two fast-acting isolation valves, one upstream and one downstream of the primary containment penetration. These valves close automatically upon receipt of signals indicating high steam flow or high temperature in the vicinity of the piping outside the drywell (as well as upon receipt of other initiating signals), in order to terminate blowdown through breaks outside the drywell.

Moment-limiting pipe whip restraints are located upstream of the inboard containment isolation valve and downstream of the outboard containment isolation valve in order to ensure the CHAPTER 03 3.6-15 REV. 18, SEPTEMBER 2016

LGS UFSAR operability of these valves in the event of a break in the RCIC steam supply line near the valves.

Whip restraints are also located on the RCIC steam supply line inside the drywell and in the isolation valve compartment at el 217' in the reactor enclosure. A typical restraint inside the drywell is shown in Figure 3.6-1.

Pipe Break Locations The postulated pipe break locations for the RCIC steam supply line, and also the pipe whip restraint locations, are shown in Figure 3.6-25 for the portion of the line inside the drywell and in Figure 3.6-26 for the portion of the line outside the drywell. The calculated stress levels and usage factors, and the postulated break types, are listed in Tables 3.6-15 and 3.6-16.

Compartment Pressure-Temperature Transients The pressure-temperature transient in the primary containment resulting from a break in the portion of the RCIC steam supply line in the drywell is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Protection against overpressurization of the RCIC pump compartment and the isolation valve compartment at el 217' in the reactor enclosure as a result of RCIC steam supply line breaks in these areas is provided by steam venting paths and blowout panels leading to the outside atmosphere. The isolation valve compartment is vented to the atmosphere via blowout panels located on the south side of the reactor enclosure, as shown in drawings M-118, M-123 and M-138. The RCIC pump compartment is vented to the isolation valve compartment via hinged, metal plate blowout panels located in the floor at el 217'. These latter panels lift to relieve pressurization in the RCIC pump compartment but do not allow pressurization in the isolation valve compartment to result in steam flow in the reverse direction, i.e., down into the RCIC pump compartment.

Pressure-temperature transient analyses for the cases involving RCIC steam supply line breaks in the RCIC pump compartment and the RCIC upper pipe tunnel were performed using the analytical technique described in Reference 3.6-1 and the blowdown data provided in Table 3.6-6. The flow model for breaks in these two compartments is shown in Figure 3.6-27, and the results of the analyses are listed in Table 3.6-7.

The RCIC pump compartment is designed to withstand the maximum pressure developed due to pipe break. No equipment located in the RCIC pump compartment is required to operate following breakage of the RCIC steam supply line. The pressure-temperature transient in the isolation valve compartment at el 217' resulting from a break in the portion of the RCIC steam supply line within that compartment is exceeded in severity by the transient resulting from HPCI steam supply line breakage in the same compartment. All equipment located in the isolation valve compartment that is required to operate following breakage of the RCIC steam supply line is qualified to operate under environmental conditions more severe than those calculated to occur due to breakage of the HPCI steam supply line.

Verification of Reactor Shutdown Capability Breakage of the RCIC steam supply line inside the drywell would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

CHAPTER 03 3.6-16 REV. 18, SEPTEMBER 2016

LGS UFSAR For an RCIC steam supply line break outside the drywell, the steam supply line containment isolation valves will be closed automatically, terminating reactor vessel blowdown. No reactor scram will occur, due to the isolation of the RCIC steam line and rapid termination of the blowdown.

After a RCIC isolation, there is sufficient capability to shutdown the reactor. The operator will investigate the cause of the RCIC isolation and take appropriate actions, in accordance with the plant Technical Specifications and approved operating procedures, to assure that the plant is in a safe condition. The Technical Specifications and/or the operating procedures, which include off normal and EOPs, may, based on plant conditions, require the operator to initiate a normal shutdown of the reactor.

A combination of pipe whip restraints and separation by distance or intervening structure is used to ensure the availability of essential systems and components in the event of an RCIC steam supply line break occurring in the drywell, the isolation valve compartment, or the RCIC pump compartment. Essential systems and components located in these areas are qualified to operate under the environmental conditions resulting from the break. Electrical cabling associated with essential systems and components is routed so as to avoid jet impingement from the postulated break.

No pipe whip restraints are provided for the portion of the RCIC steam supply line located within the RCIC pump compartment. Since no essential systems and components are located within the RCIC pump compartment, such protection is not necessary.

3.6.1.2.1.9 Main Steam Drain Lines The main steam drainage piping connects to the four main steam lines both inside and outside the drywell. Inside the drywell, 2-inch drain lines that connect to each of the main steam lines are headered together into a single 3-inch line, which then penetrates the drywell wall. This 3-inch drain header is provided with two containment isolation valves, one upstream and one downstream of the containment penetration, both of which are normally closed. Outside the drywell (in the main steam tunnel), 2-inch drain lines that connect to each of the main steam lines are headered together into a single 3-inch line, which then connects to the drain header from the drain lines inside the drywell. Downstream of the connection of the two drain headers, the common line splits into two 3-inch lines again, one routed to the main condenser and one routed to the liquid waste management system. These latter two lines are both provided with normally closed valves located within the main steam tunnel. The routing of the main steam drain lines is shown in drawings M-217, M-226, M-297, and M-326.

Pipe Break Locations The postulated pipe break locations for the main steam drain lines inside the drywell are shown in Figure 3.6-28. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-17. Breaks in the main steam drain lines within the main steam tunnel are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Compartment Pressure-Temperature Transients The pressure transient in the primary containment resulting from a break in a main steam drain line within the drywell is exceeded in severity by transients resulting from recirculation loop breaks and main steam line breaks. The temperature transient in the primary containment resulting from a CHAPTER 03 3.6-17 REV. 18, SEPTEMBER 2016

LGS UFSAR main steam drain line break is exceeded in severity by the transient resulting from a small steam leak. These design basis transients are discussed in Section 6.2.1.

The pressure-temperature transient in the main steam tunnel resulting from a main steam drain line break within the tunnel is exceeded in severity by the transient resulting from a main steam line break.

Verification of Reactor Shutdown Capability Breakage of a main steam drain line inside the drywell would result in an unisolable blowdown of steam from the reactor vessel through the broken line. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

For the case of a main steam drain line break inside the main steam tunnel, the resultant temperature rise in the tunnel would cause the MSIVs and the MSL drain isolation valves, if open to close automatically, thereby terminating steam blowdown through the break. Once the MSIVs and MSL drain isolation valves have been closed, the sequence of events is similar to that for a main steam line break outside the drywell.

Separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of a main steam drain line break in either the drywell or the main steam tunnel.

3.6.1.2.1.10 RPV Head Vent Line The RPV head vent line is located entirely within the drywell. From its connection point to a 4-inch flanged nozzle on the RPV head, the line reduces to 2 inches in diameter and then is routed generally downward to a penetration through the containment seal plate. From this point, the line continues on downward to its connection with 26-inch main steam line C.

Pipe Break Locations The postulated pipe break locations for the RPV head vent line, and also the pipe whip restraint locations, are shown in Figure 3.6-30. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-19.

Compartment Pressure-Temperature Transients Since the RPV head vent line is located entirely within the drywell, breakage of this line has no effect on plant areas outside the primary containment. The pressure transient in the primary containment resulting from a break in the RPV head vent line is exceeded in severity by transients resulting from recirculation loop breaks and main steam line breaks. The temperature transient in the primary containment resulting from a break in the RPV head vent line is exceeded in severity by the transient resulting from a small steam leak. These design basis transients are discussed in Section 6.2.1.

CHAPTER 03 3.6-18 REV. 18, SEPTEMBER 2016

LGS UFSAR Verification of Reactor Shutdown Capability Breakage of the RPV head vent line would result in an unisolable blowdown of steam into the drywell. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

Separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of an RPV head vent line break.

3.6.1.2.1.11 Standby Liquid Control System Injection Line The discharge lines from the SLCS injection pumps penetrate the drywell through two separate penetrations and are headered together inside the drywell to form a single 2-inch line. From the vicinity of the drywell penetrations (between el 265' and el 271'), this line is routed horizontally and upward to its connection to core spray injection line B (12" DCA-319). In addition to the containment isolation valves, the SLCS injection line is provided with a check valve near the connection to the core spray line. Only that portion of the SLCS injection line between the core spray line and the inboard check valve is considered high energy during periods when the reactor is pressurized.

Pipe Break Locations The postulated pipe break locations for the SLCS injection line are shown in Figure 3.6-31. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-20.

Compartment Pressure-Temperature Transients Since the high energy portion of the SLCS injection line is located entirely within the drywell, breakage of this line would have no effect on plant areas outside the primary containment. The pressure-temperature transient in the primary containment resulting from a break in the SLCS injection line is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability Breakage of the SLCS injection line between the RPV and the first check valve in that line would result in an unisolable blowdown from the reactor vessel into the drywell. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

Separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of an SLCS injection line break.

3.6.1.2.1.12 RHR Shutdown Cooling Suction Line The RHR shutdown cooling suction line is a 20-inch line connected to recirculation loop B in the drywell. The line is routed downward and then horizontally to its containment penetration at el 244'-8". The line is provided with two containment isolation valves (one inboard and one outboard CHAPTER 03 3.6-19 REV. 18, SEPTEMBER 2016

LGS UFSAR of the penetration), both of which are normally closed. Thus, only that portion of the line between the recirculation loop and the inboard containment isolation valve is considered high energy. The routing of the line is shown in drawings M-213, M-217, M-225, M-295, M-296, and M-326.

The RHR shutdown cooling suction line is provided with pipe whip restraints on the portion of the line inside the drywell. Examples of these restraints are shown in Figure 3.6-1.

Pipe Break Locations The postulated pipe break locations for the RHR shutdown cooling suction line, and also the pipe whip restraint locations, are shown in Figure 3.6-32. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-21.

Compartment Pressure-Temperature Transients Since the high energy portion of the RHR shutdown cooling suction line is located entirely within the drywell, breakage of the line would have no effect on plant areas outside the primary containment. The pressure-temperature transient in the primary containment resulting from a break in the RHR shutdown cooling suction line is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability Breakage of the RHR shutdown cooling suction line would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

A combination of pipe whip restraints and separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of a break in the RHR shutdown cooling suction line.

3.6.1.2.1.13 RHR Shutdown Cooling Return Line One 12-inch RHR shutdown cooling return line is associated with RHR loop A and a second return line is associated with RHR loop B. Since the two return lines are routed symmetrically on opposite sides of the drywell, the following discussion applies to both lines. The RHR shutdown cooling return line, from the discharge of the associated RHR pump and heat exchanger, penetrates the south side of the drywell at el 244'-8". The line is then routed horizontally following the curvature of the drywell wall until a point due west (for loop A; due east for loop B) of the reactor vessel centerline is reached. From this point, the line is routed upward to el 265'-4" and then horizontally to its connection with the discharge riser of the reactor recirculation loop. The routing of this piping is shown in drawings M-213, M-217, M-225, M-295, M-296, and M-326.

The RHR shutdown cooling return line is provided with two containment isolation valves: a normally closed globe valve outside the drywell and a check valve inside the drywell. Only that portion of the line between the reactor recirculation line and the inboard check valve is considered high energy during periods when the reactor is pressurized.

The RHR shutdown cooling return line is provided with one pipe whip restraint, located at the upper elbow of the vertical portion of the line. A detail of this restraint is shown in Figure 3.6-1.

CHAPTER 03 3.6-20 REV. 18, SEPTEMBER 2016

LGS UFSAR Pipe Break Locations The postulated pipe break locations for the RHR shutdown cooling return line, and also the pipe whip restraint locations, are shown in Figure 3.6-32. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-21.

Compartment Pressure-Temperature Transients Since the high energy portion of the RHR shutdown cooling return line is located entirely within the drywell, breakage of the line would have no effect on plant areas outside the primary containment.

The pressure-temperature transient in the primary containment resulting from a break in the RHR shutdown cooling return line is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability Breakage of the RHR shutdown cooling return line would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

A combination of pipe whip restraints and separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of a break in the RHR shutdown cooling return line.

3.6.1.2.1.14 LPCI Injection Line There are four 12-inch LPCI injection lines, one associated with each of the four RHR pumps. The four lines are routed symmetrically inside the drywell, with the A and C injection lines entering the west side of the drywell and the B and D lines entering the east side of the drywell. The following discussion applies to all four lines. The LPCI injection line penetrates the drywell at el 285'-2" and is routed up to el 297'-3" inches where it connects to the reactor vessel nozzle. The routing of this piping is shown in drawings M-215, M-217, M-234, M-235, M-305, M-306, M-316, and M-326.

The LPCI injection line is provided with two containment isolation valves: a normally closed gate valve outside the drywell and a check valve inside the drywell. Only that portion of the line between the reactor vessel nozzle and the inboard check valve is considered high energy during periods when the reactor is pressurized.

The LPCI injection line is restrained to prevent pipe whip inside the drywell. A typical restraint is shown in Figure 3.6-1.

Pipe Break Locations The postulated pipe break locations for the LPCI injection line, and also the pipe whip restraint locations, are shown in Figure 3.6-33. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-22 CHAPTER 03 3.6-21 REV. 18, SEPTEMBER 2016

LGS UFSAR Compartment Pressure-Temperature Transients Since the high energy portion of the LPCI injection line is located entirely within the drywell, breakage of the line would have no effect on plant areas outside the primary containment. The pressure-temperature transient in the primary containment resulting from a break in the LPCI injection line is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

Verification of Reactor Shutdown Capability Breakage of the LPCI injection line would result in an unisolable blowdown of the reactor vessel.

The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

A combination of pipe whip restraints and separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of a break in the LPCI injection line.

3.6.1.2.1.15 Core Spray Injection Line There are two core spray injection lines, one associated with core spray pumps A and C and one associated with core spray pumps B and D. Since the two lines are routed symmetrically within the drywell, the following discussion applies to both lines. The core spray injection line penetrates the north side of the drywell at el 297'-3" and is routed up to el 306'-7" before connecting to the RPV nozzle. The routing of the piping is shown in drawings M-217, M-235, M-306, and M-326.

The core spray injection line is provided with containment isolation valves both inside and outside the drywell, the inboard valve being a check valve. Only that portion of the line between the reactor vessel nozzle and the inboard check valve is considered high energy during periods when the reactor is pressurized.

The core spray injection line is restrained to prevent pipe whip inside the drywell. Typical restraints are shown in Figure 3.6-1.

Pipe Break Locations The postulated pipe break locations for the core spray injection line, and also the pipe whip restraint locations, are shown in Figure 3.6-34. The calculated stress levels and usage factors, and the postulated break types, are listed in Table 3.6-23.

Compartment Pressure-Temperature Transients Since the high energy portion of the core spray injection line is located entirely within the drywell, breakage of the line would have no effect on plant areas outside the primary containment. The pressure-temperature transient in the primary containment resulting from a break in the core spray injection line is exceeded in severity by the transients resulting from recirculation loop breaks and main steam line breaks, which are discussed in Section 6.2.1.

CHAPTER 03 3.6-22 REV. 18, SEPTEMBER 2016

LGS UFSAR Verification of Reactor Shutdown Capability Breakage of the core spray injection line would result in an unisolable blowdown of the reactor vessel. The sequence of events that would occur automatically to shut the reactor down and cool the core is discussed in Section 6.3.3.

A combination of pipe whip restraints and separation by distance and intervening structure is used to ensure the availability of essential systems and components in the event of a break in the core spray injection line.

3.6.1.2.1.16 Control Rod Drive Hydraulic The 2 CRD water pumps are located in the turbine enclosure. The high energy discharge pipes from the 2 pumps are headered together, and a single 2-inch pipe is routed from the turbine enclosure into the reactor enclosure at el 201'. This 2-inch line is then routed upward to the CRD hydraulic system master control station located at el 253' of the reactor enclosure. From the master control station, a 2-inch cooling water header and a 2-inch charging water header are routed to the groups of HCUs on the west side and east side of the drywell. Containment isolation valves are provided in each header between the master control station and the HCUs. The piping between the isolation valves has been upgraded to be equivalent to ASME III, Class 2.

Pipe Break Locations Since the CRD pump discharge line, the cooling water header, and the charging water header originally consisted entirely of non-nuclear class piping, breaks are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments. The original analysis bounds the effects of the addition of containment isolation valves in the headers.

Compartment Pressure-Temperature Transients Since the normal fluid temperature in the CRD hydraulic system is less than 120F, no significant pressure-temperature transient would result from postulated breaks.

Verification of Reactor Shutdown Capability Loss of water pressure due to a break in the CRD pump discharge line, cooling water header, or charging water header will not prevent the control rods from being inserted into the reactor core. At reactor pressures of 450 psig or higher, reactor pressure alone is sufficient to fully insert the control rods. At lower reactor pressures, the scram accumulators assist in supplying the energy necessary to insert the control rods.

3.6.1.2.1.17 Auxiliary Steam Line From the auxiliary boiler, auxiliary steam is distributed via an 8-inch header to the various steam-consuming components in the turbine enclosure and the radwaste enclosure. This auxiliary steam header passes through the offgas pipe tunnel north of the control structure at el 187'. The auxiliary steam system also provides steam to the RCIC and HPCI systems for testing purposes.

CHAPTER 03 3.6-23 REV. 18, SEPTEMBER 2016

LGS UFSAR Pipe Break Locations Since the auxiliary steam line consists of non-nuclear class piping, breaks are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Compartment Pressure-Temperature Transients A pressure-temperature transient analysis for the case of an auxiliary steam line break in a safety-related area was not performed because the piping does not pass through any safety related areas with the exception of the test lines to HPCI and RCIC. However, a pressure-temperature transient analysis was performed for the auxiliary steam line break in the offgas pipe tunnel to ensure control structure integrity. The HPCI and RCIC test lines are normally isolated from the reactor enclosure by closed valves. The lines are only energized in the reactor enclosure for short periods (i.e., less than 1% of the time) and are therefore considered as moderate energy lines.

Verification of Reactor Shutdown Capability Breakage of the auxiliary steam line would have no effect on operation of the reactor. Except for the test lines for HPCI and RCIC, auxiliary steam lines are not located in areas containing safety related equipment. In the case of the HPCI and RCIC test lines, they are considered to be moderate energy lines based on the short periods in which they are energized. As moderate energy lines, they are acceptable since all rooms containing safety related equipment are qualified for a moderate energy line break.

3.6.1.2.1.18 Plant Heating Steam Piping From the auxiliary boiler, plant heating steam is distributed to various steam-consuming components in the turbine enclosure and the diesel generator compartments. Additionally, heating steam is provided to the refueling floor and the railroad air-lock in the reactor enclosure and in the diesel corridor outside of the reactor enclosure, however, none of these areas contain safety related equipment necessary to shut down the reactor or maintain primary containment integrity.

Pipe Break Locations Because the plant heating steam piping consists of non-nuclear class piping, breaks are postulated to occur at each location of potential high stress, such as pipe fittings, valves, and welded attachments.

Verification of Reactor Shutdown Capability Breakage of the plant heating steam piping would have no effect on operation of the reactor because reactor systems are not located in areas through which the heating steam piping is routed. The only areas of the plant through which heating steam piping is routed that contain safety-related systems are the diesel generator compartments, the refuel floor, the diesel corridor, and the railroad air lock. In addition to the diesel generators, the safeguard MCCs for the following valves are located within diesel generator enclosures:

a. ESW to RHRSW valves (HV-11-11A, 11B, 15A, 15B; drawing M-11).

b. ESW to TECW heat exchanger valves (HV-11-105, 107, 205, 207; Figure 9.2-2).

CHAPTER 03 3.6-24 REV. 18, SEPTEMBER 2016

LGS UFSAR

c. TECW heat exchanger to RHRSW valves (HV-12-110, 210; drawing M-12).

d. CST to ECCS (HV-55-124, 224; drawing M-55).

The MCCs for these valves are distributed throughout the diesel generator cells such that the loss of the MCCs in any one cell would not compromise the safety function of the systems. Thus, an HELB in one cell would not affect the safe shutdown capability of the plant.

A rupture of the plant steam heating piping in one cell would not affect the diesel generators located in other cells. Thus, a failure of this system would result in loss of only the one diesel. The safe shutdown capability of the plant would therefore not be compromised because only three of the four diesels are required in the event of a LOOP.

Because the loss of components in any diesel generator compartment from a high or moderate energy line break would not precipitate a trip of the turbine-generator or a trip of the RPS, offsite power is assumed to remain available in the analysis of the effects of postulated piping failures in these systems (BTP ASB 3.1, paragraph 3.b.1).

There are no essential systems or components as defined in Branch Technical Position SPLB 3-1 (Protection against postulated piping failures in fluid systems outside containment) on the refueling floor, the diesel corridor or the railroad airlock.

3.6.1.2.2 Moderate Energy Fluid Systems Each room, compartment, or area containing components essential for safe shutdown has been evaluated for the effects of postulated ruptures in moderate energy piping.

Safe shutdown components were evaluated for operability in the postulated water spray and/or flooding environment. Safe shutdown components in the reactor enclosure, the control structure, the diesel generator enclosure, the spray pond pumphouse, the main steam tunnel, and the RHRSW pipe tunnel were considered.

Those components not designed for operability in a water spray and/or a flooding environment were assumed to fail. Electrical components that are needed for safe shutdown, but that would fail under the water spray/flooding conditions, were protected from water spray and flooding conditions.

The synergistic effects of an independent single active failure, in addition to the effects of the pipe break, were evaluated and were found acceptable subject to any limitations discussed below. For some systems (such as RHR), a single active failure in the redundant loop of the same system has been excluded in accordance with Section 3.6.1.1.

The crack sizes postulated, and the nominal pipe sizes in which moderate energy pipe cracks are postulated to occur, are discussed in Section 3.6.2.1.3. Additional criteria used in the moderate energy fluid systems analysis are discussed in Section 3.6.1.1.

3.6.1.2.2.1 Primary Containment All equipment within the primary containment that must operate during or after a LOCA is qualified for the appropriate environmental conditions, as described in Section 3.11. Wetting associated CHAPTER 03 3.6-25 REV. 18, SEPTEMBER 2016

LGS UFSAR with the postulated failure of any moderate energy pipe is within the bounds of that qualification.

Consequently, no item-by-item discussion of this less severe event is required to verify that capability of safely shutting down the plant.

3.6.1.2.2.2 Reactor Enclosure Compartments or areas on each elevation that contain safe shutdown electrical equipment were evaluated. Piping and mechanical drawings for each elevation of the reactor enclosure are provided in Section 1.2. All safety-related equipment whose operability in a water spray or flooding environment must be ensured were either protected from the consequences of the pipe break or are designed to remain operable in that environment.

3.6.1.2.2.3 Control Structure Compartments or areas that contain safety-related equipment on each elevation were evaluated beginning with el 200', the lowest elevation of the control structure. All safety-related equipment whose operability in a water spray or flooding environment must be ensured were either protected from the consequences of the pipe break or were designed to operate in that environment. Piping and mechanical drawings for each elevation of the control structure are provided in Section 1.2.

3.6.1.2.2.4 Diesel Generator Compartments Moderate energy piping systems that are routed through the diesel generator compartments include:

a. ESW to the diesel generator heat exchangers

b. diesel fuel oil supply lines

c. diesel generator starting air

d. demineralized water from the jacket water expansion tank to the diesel generator

e. lube oil from the storage tank to the diesel generator

f. fire protection system (normally dry)

Each of the moderate energy piping systems is integral to the operation or protection of the diesel generator located in the same compartment as the piping. Thus, a failure of one or more of these systems would result in loss of only the one diesel. The safe shutdown capability of the plant would therefore not be compromised because only three of the four diesels are required in the event of a LOOP.

In addition to the diesel generators, a safeguard MCC is located within each diesel generator enclosure. The impact of a pipe break in the diesel generator enclosure on these MCCs is discussed in Section 3.6.1.2.1.18.

CHAPTER 03 3.6-26 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6.1.2.2.5 Spray Pond Pumphouse The spray pond pumphouse contains two separate ESW and RHRSW pump areas, two RHRSW and ESW pipe-ways, two RHRSW valve compartments, and several other compartments.

Each pump-room contains two RHRSW pump motors and two ESW pump motors. The pumps are common to both Unit 1 and Unit 2.

During normal plant operation, the ESW and RHRSW piping is depressurized except during surveillance testing of ESW, RHRSW, and the diesel generators; and during operation for reactor shutdown cooling, suppression pool cooling, and spray pond cooling/chemistry control. The frequency of the surveillance tests is in accordance with the Technical Specifications, and the frequency of operation for other purposes varies with plant conditions.

A moderate energy line break in these lines during normal plant operation is unlikely. However, if a line break should occur, water spray could impinge on the pump motors in the room. In this event, up to two ESW and two RHRSW pumps could be disabled. The loss of the two ESW pumps is of no concern because shutdown cooling loads would be provided by the service water system.

Because nonsafeguard heat removal systems are available, the RHRSW system is not critical for safe shutdown of the reactors.

The RHRSW valve compartments contain valves that are essential for the safe operation of the ESW system. However, as discussed above, shutdown cooling loads would be provided by the service water system because offsite power would be assumed to be available.

A pipe break in the wet pit, the ESW and RHRSW pipe-way, or the access hatch area would not impact safe shutdown capability because these compartments do not contain components that are essential for safe shutdown.

Spray pond pumphouse flooding is limited to the compartments adjoining the pump area. The pump area itself cannot flood because water would flow through open grating into the spray pond.

Flooding in the adjacent compartments could damage at most a single mechanical division.

Because it is not necessary to postulate a single active failure in the redundant RHRSW system, and because flooding can damage only one mechanical division of RHRSW, safe shutdown capability will not be impaired.

3.6.1.2.2.6 Main Steam Tunnel Safe shutdown components in the main steam tunnel include the outboard MSIVs.

The MSIVs are qualified to operate in environments more severe than that caused by moderate energy water spray. The valves are located well above the maximum water flood level. Therefore, the occurrence of a moderate energy pipe crack in this compartment would have no effect on safe shutdown capability.

CHAPTER 03 3.6-27 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6.1.2.2.7 RHRSW Pipe Tunnel A moderate energy pipe break in the RHRSW pipe tunnel would not impact safe shutdown capability.

The only safe shutdown components on this elevation are two RHRSW valves. Each of these valves are qualified to be submersible and would withstand postulated water spray and flooding conditions.

3.6.1.3 Safety Evaluation The analyses of postulated pipe ruptures summarized in Section 3.6.1.2 verify that the consequences of any single rupture of fluid system piping in the plant will not prevent safe shutdown of the reactor from being achieved.

The offsite radiological consequences of piping ruptures are enveloped by a reactor recirculation system break for all breaks inside primary containment, and by main steam system and feedwater system breaks for all breaks outside primary containment. The radiological consequences of these breaks are presented in Sections 15.6.5, 15.6.4, and 15.6.6, respectively.

Special consideration has been given to protecting the control room and other areas of the control structure containing essential systems and components from the effects of postulated pipe ruptures. Exterior walls of the control structure above the slab at el 217' are designed as steam-tight in those areas where the walls could be subject to steam pressurization resulting from rupture of high energy fluid system piping outside the control structure. HVAC ducts penetrating these portions of the control structure walls are equipped with back pressure dampers, and other types of penetrations through the walls are designed as steam-tight.

As described in Sections 3.6.1.2.1.2 and 3.6.1.2.1.3, those portions of the main steam and feedwater lines routed near the control structure are provided with pipe whip restraints and bumpers to prevent a postulated rupture of these lines from causing unacceptable damage to the control structure walls.

3.6.2 DETERMINATION OF PIPE FAILURE LOCATIONS AND DYNAMIC EFFECTS ASSOCIATED WITH POSTULATED PIPING FAILURES Information concerning break and crack location criteria and methods of analysis is presented in this section. The location criteria and methods of analysis are needed to evaluate the dynamic effects associated with postulated ruptures of high energy and moderate energy piping inside and outside the primary containment.

3.6.2.1 Criteria Used to Determine Pipe Break and Crack Locations and Their Configurations 3.6.2.1.1 Break Locations in High Energy Fluid System Piping 3.6.2.1.1.1 Piping in Containment Penetration Areas High energy pipes penetrating the primary containment are provided with moment-limiting restraints that are located reasonably close to the containment isolation valves and are designed to CHAPTER 03 3.6-28 REV. 18, SEPTEMBER 2016

LGS UFSAR withstand the loadings resulting from a pipe break either inboard of the inboard isolation valve or outboard of the outboard isolation valve so that neither isolation valve operability nor leak-tight integrity of the containment penetration would be impaired as a result of such pipe breaks.

Terminal ends of high energy piping that penetrate the containment are considered to originate beyond the containment isolation valve and its moment-limiting restraint, both inboard and outboard.

Breaks are not postulated in these portions of high energy piping in containment penetration areas provided that the following design stress and fatigue limits are satisfied:

For ASME Section III, Class 1 Piping

a. The stress intensity range Sn, calculated for normal and upset conditions by equation (10) of paragraph NB-3653, does not exceed 2.4 Sm, and the cumulative usage factor associated with normal, upset, and testing conditions is less than 0.1, or

b. The stress intensity range Sn, calculated for normal and upset conditions by equation (10) of paragraph NB-3653 exceeds 2.4 Sm but does not exceed 3.0 Sm and the cumulative usage factor associated with normal, upset, and testing conditions is less than 0.1, or

c. The stress intensity range Sn, calculated for normal and upset conditions by equation (10), exceeds 3.0 Sm, but the stress intensity ranges computed by equations (12) and (13) of paragraph NB-3653 are less than 2.4 Sm and the cumulative usage factor associated with normal, upset, and testing conditions is less than 0.1

d. Breaks are always postulated whenever the usage factor exceeds 0.1 regardless of stress.

e. The loading resulting from a postulated pipe break beyond these portions of the piping does not cause the stress as calculated by equation (9) in paragraph NB-3652 to exceed 2.25 Sm, except for the portion of piping between the isolation valve and the adjacent restraints protecting the operability of the valve. For this latter portion of piping, higher stresses are permitted provided that a plastic-hinge is not formed and the operability of the isolation valve is ensured.

After the as-built analysis is completed, the licensee will conduct a comparison against the SRP criteria (NUREG-0800) (BTP MEB 3-1) and demonstrate no additional breaks for the following situations:

a. The stress calculated by equation (10) is between 2.4 Sm and 3.0 Sm and

b. The ASME Code version of 1979 Summer Addenda or later is used for analysis.

CHAPTER 03 3.6-29 REV. 18, SEPTEMBER 2016

LGS UFSAR For ASME Section III, Class 2 and 3 Piping

a. The maximum stress ranges as calculated by the sum of equations (9) and (10) in paragraph NC-3652, considering normal and upset plant conditions, does not exceed 0.8(1.2 Sh + SA).

b. The maximum stress, as calculated by equation (9) in paragraph NC-3652, under the loadings resulting from a postulated rupture of fluid system piping beyond these portions of piping does not exceed 1.8 Sh, except for the portion of the piping between the isolation valve and the adjacent restraints protecting the operability of the valve. For this latter portion of the piping, higher stress is permitted, provided that a plastic-hinge is not formed and the operability of the isolation valve is assured.

In addition to these stress and fatigue criteria, high energy piping in containment penetration areas must meet the following requirements:

a. Welded pipe support attachments are avoided to eliminate stress concentrations.

b. The number of circumferential and longitudinal pipe welds and branch connections is minimized.

c. The length of the piping run is minimized, consistent with requirements to keep stress levels low and provide access for inservice inspection.

d. The design at points of pipe fixity (such as pipe anchors or welded connections at containment penetrations) does not require welding directly to the outer surface of the piping (flued, integrally forged pipe fittings are acceptable), except where such welds are 100% volumetrically examinable in service and a detailed stress analysis is performed to demonstrate compliance with the limits of the stress and fatigue criteria stated above.

e. In accordance with the inservice inspection plan, the inservice examination completed during each inspection interval will provide either a 100% volumetric examination of circumferential pipe welds within these portions of piping or those examinations required by the Risk Informed Inservice Inspection (RISI) Program as applied to these portions of piping also known as the Break Exclusion Region (BER). Inservice inspection of the RCPB is discussed in Section 5.2.4.

f. For piping constructed in accordance with ANSI B31.1, all welds will be fully radiographed.

After the as-built analysis is completed, the licensee will conduct a comparison against the SRP criteria (NUREG-0800) (BTP MEB 3-1) and demonstrate no additional breaks for the following situations:

a. The stress calculated by equation (10) is between 2.4 Sm and 3.0 Sm and

b. The ASME Code version of 1979 Summer Addenda or later is used for analysis.

CHAPTER 03 3.6-30 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6.2.1.1.2 Recirculation System Piping Pipe breaks in the Unit 1 recirculation system are postulated to occur at the following locations:

a. Terminal ends of a piping run or branch run.

b. At intermediate locations between terminal ends where the maximum stress range between any two load sets (including the zero load set), as calculated according to ASME Section III, subarticle NB-3600, for upset plant conditions and an independent OBE event, meets the following requirements:

1. The stress range, as calculated using equation (12) or (13), exceeds 2.4 Sm and the cumulative usage factor associated with normal, upset, and testing conditions is less than 0.1.

2. The stress range calculated using equation (10) exceeds 2.4 Sm but is less than 3.0 Sm, and the cumulative usage factor exceeds 0.1.

3. The stress range calculated using equation (10) exceeds 3.0 Sm, and the cumulative usage factor exceeds 0.1.

4. Breaks are always postulated whenever the usage factor exceeds 0.1 regardless of stress.

5. If two or more intermediate break locations cannot be determined by stress or usage factor limits, two arbitrary intermediate break locations are selected on a reasonable basis. This basis includes consideration of fitting locations and/or highest stress or usage factor locations. Where more than two such intermediate locations are possible using the application of the above reasonable basis, those two locations possessing the greatest damage potential are used. A break at each end of a fitting can be classified as two discrete break locations when the stress analysis is sufficiently detailed to differentiate stresses at each postulated break.

After the as-built analysis is completed, the licensee will conduct a comparison against the SRP criteria (NUREG-0800) (BTP MEB 3-1) and demonstrate no additional breaks for the following situations:

a. The stress calculated by equation (10) is between 2.4 Sm and 3.0 Sm, and

b. The ASME Code version of 1979 Summer Addenda or later is used for analysis.

For Unit 2, break locations are similarly postulated except that the two arbitrary intermediate break locations of b.4 above need not be postulated provided the following criteria are met:

a. The piping system is adequately resistant to IGSCC and is not susceptible to unanticipated water hammer thermal transient events.

CHAPTER 03 3.6-31 REV. 18, SEPTEMBER 2016

LGS UFSAR

b. The piping system is included in the startup testing program for piping steady-state vibrations.

c. The elimination of the break location does not change equipment environmental qualification or structural design criteria.

3.6.2.1.1.3 Class 1 Piping (Other Than Recirculation System Piping and Piping in Containment Penetration Areas)

Breaks in Class 1 piping (ASME Section III) for Unit 1 are postulated to occur at the following locations:

a. At terminal ends of piping runs or branch runs.

b. At intermediate locations between terminal ends, as determined by one of the two following criteria:

1. At each location of potential high stress such as pipe fittings (elbows, tees, reducers, etc) valves, and welded attachments.

2. At each location where, for normal and upset load conditions, none of the following stress and fatigue limits are met:

(a) The stress intensity range Sn, calculated by equation (10) of paragraph NB-3653, does not exceed 2.4 Sm and the cumulative usage factor associated with normal, upset, and testing conditions is less than 0.1.

(b) The stress intensity range Sn, as calculated by equation (10) of paragraph NB-3653, exceeds 2.4 Sm but is less than 3.0 Sm, and the cumulative usage factor is less than 0.1.

(c) The stress intensity range Sn exceeds 3.0 Sm, but the stresses computed by equations (12) and (13) of paragraph NB-3653 are less than 2.4 Sm, and the cumulative usage factor is less than 0.1.

(d) Breaks are always postulated whenever the usage factor exceeds 0.1 regardless of stress.

3. When the above stress and fatigue criteria result in less than two intermediate break locations, a minimum of two locations are chosen based on highest stress, as calculated by equation (10) of paragraph NB-3653.

The two locations are separated by a change of direction of the force resulting from pipe break. If the piping run has no more than one change of direction, a minimum of one intermediate break location is chosen. Any given fitting is considered as a single break location regardless of the number of breaks postulated at different locations on that fitting.

CHAPTER 03 3.6-32 REV. 18, SEPTEMBER 2016

LGS UFSAR Intermediate pipe break locations are initially based on committed design piping stress calculations in accordance with the above criteria. As a result of piping reanalysis, the highest stress locations may be shifted. An initially determined pipe break location will not be changed as a consequence, however, unless one of the following conditions exists:

(a) Reanalysis shows that maximum stress range or cumulative usage factor at another location not only exceeds that for the initial pipe break location but also exceeds the above pipe break criteria. In addition, the break at the new location results in more serious consequences to safety-related systems than the initial break location.

(b) Significant changes are made in the routing, size, or wall thickness of the pipe after the initial pipe break determination.

4. After the as-built analysis is completed, the licensee will conduct a comparison against the SRP criteria (NUREG-0800) (BTP MEB 3-1) and demonstrate no additional breaks for the following situations:

(a) The stress calculated by equation (10) is between 2.4 Sm and 3.0 Sm and (b) The ASME Code version of 1979 Summer Addenda or later is used for analysis.

For Unit 2, break locations are similarly postulated except that arbitrary intermediate breaks from b.3 above need not be postulated provided the following criteria are met:

a. The piping system is adequately resistant to IGSCC and is not susceptible to unanticipated water hammer thermal transient events.

b. The piping system is included in the startup testing program for piping steady-state vibrations.

c. The elimination of the break location does not change equipment environmental qualification or structural design criteria.

The following Unit 2 class 1 piping systems meet all of the above criteria:

• Main Steam Inside Containment

• Feedwater Inside Containment

• RWCU Inside Containment

• Reactor Vessel Drain

• HPCI Steam Supply Inside Containment CHAPTER 03 3.6-33 REV. 18, SEPTEMBER 2016

LGS UFSAR

• RCIC Steam Supply Inside Containment

• SLCS

• RHR Shutdown Cooling Suction

• RHR Shutdown Cooling Return

• LPCI Injection

• Core Spray Injection 3.6.2.1.1.4 Class 2 and 3 Piping (Other Than Recirculation System Piping and Piping in Containment Penetration Areas)

Breaks in Class 2 and 3 piping (ASME Section III) for Unit 1 are postulated to occur at the following locations:

a. At terminal ends of piping runs or branch runs.

b. At intermediate locations between terminal ends, as determined by one of the three following criteria:

1. At each location of potential high stress, such as pipe fittings (elbows, tees, reducers, etc), valves, and welded attachments.

2. At each location where the maximum stress range, as calculated by the sum of equations (9) and (10) of paragraph NC-3652, considering normal and upset plant conditions, exceeds 0.8(1.2 Sh + SA).

3. When the above stress and fatigue criteria result in less than two intermediate break locations, a minimum of two locations are chosen based on highest stress, as calculated by the sum of equations (9) and (10) of paragraph NC-3652. The two locations are separated by a change of direction of the force resulting from pipe break. If the piping run has no more than one change of direction, a minimum of one intermediate break location is chosen. Any given fitting is considered as a single break location regardless of the number of breaks postulated at different locations on that fitting.

Intermediate pipe break locations are initially based on committed design piping stress calculations in accordance with the above criteria. As a result of piping reanalysis, the highest stress locations may be shifted. An initially determined pipe break location will not be changed as a consequence, however, unless one of the following conditions exists:

(a) Reanalysis shows that maximum stress range or cumulative usage factor at another location not only exceeds that for the initial pipe CHAPTER 03 3.6-34 REV. 18, SEPTEMBER 2016

LGS UFSAR break location but also exceeds the above pipe break criteria. In addition, the break at the new location results in more serious consequences to safety-related systems than the initial break location.

(b) Significant changes are made in the routing, size, or wall thickness of the pipe after the initial pipe break determination.

4. At each extreme of the piping run adjacent to basic protective structures, when the piping system contains no fitting, valve, or welded attachment.

For Unit 2, break locations are similarly postulated except that the arbitrary intermediate breaks from b.3 above need not be postulated provided the following criteria are met:

a. The piping system is adequately resistant to IGSCC and is not susceptible to unanticipated water hammer thermal transient events.

b. The piping system is included in the startup testing program for piping steady-state vibrations.

c. The elimination of the break location does not change equipment environmental qualification or structural design criteria.

The following Unit 2 Class 2, 3 piping systems meet all of the above criteria:

• RWCU Outside Containment

• HPCI Steam Supply Outside Containment

• RCIC Steam Supply Outside Containment

• Main Steam Outside Containment 3.6.2.1.1.5 Non-Nuclear Class Piping Breaks in non-nuclear class piping are postulated to occur at the following locations:

a. At terminal ends of piping runs or branch runs.

b. At each intermediate location of potential high stress, such as pipe fittings (elbows, tees, reducers, etc), valves, and welded attachments.

Alternatively, the break locations for non-nuclear class piping can be selected according to the same criteria used for Class 2 and 3 piping, provided that all necessary analyses are made.

These breaks in nonseismic Category I piping have been postulated at those locations that would result in the maximum amount of damage and the safety-related systems and components have adequate protection from these piping breaks as discussed below.

CHAPTER 03 3.6-35 REV. 18, SEPTEMBER 2016

LGS UFSAR The only high energy seismic Category II piping in the control structure is the portion of the steam supply line to the offgas recombiner preheater that is located within the recombiner compartments.

The only high energy seismic Category II piping in the reactor enclosure is the RWCU piping inside the RWCU filter/demineralizer compartments and the RWCU holding pump compartments. In both cases, the walls of the compartments are capable of withstanding the pipe whip and jet impingement forces and the compartment pressurization that could result from breaks at the most adverse locations. Because there are no safety-related components located in these compartments, safety-related components will not be affected by high energy pipe breaks within these compartments.

The diesel generator enclosures have high energy seismic Category IIA piping in the form of heating steam supply lines (11/2" JBD-330/430). A rupture of such a line could create maximum temperatures of less than 340F and pressures less than 15 psia (0.3 psig) due to the large "bird grill" vent openings. As only one diesel generator enclosure is affected at any one time, the assumed loss of one diesel generator is acceptable.

Other safety-related structures, such as the spray pond pumphouse, do not contain high energy seismic Category II piping.

Safety-related components are protected from the effects of high energy pipe breaks in nonsafety-related structures by the walls that separate the safety-related structures from the nonsafety-related structures. To the extent necessary to prevent unacceptable damage to safety-related components, these walls are designed to withstand the pipe whip and jet impingement forces and compartment pressurization that could result from breaks at the most adverse locations in high energy seismic Category II piping within the nonsafety-related structures.

3.6.2.1.2 Crack Locations in Moderate Energy Fluid System Piping Through-wall leakage cracks are postulated to occur in moderate energy piping located in areas containing essential systems and components. Cracks are postulated to occur at terminal ends of piping runs or branch runs, and at intermediate locations selected in accordance with either of the two following criteria:

a. At each location of potential high stress, such as pipe fittings (elbows, tees, reducers, etc), valves, and welded attachments

b. For Class 1 piping (ASME Section III), at locations where the maximum range of stress intensity as calculated by equation (10) of paragraph NB-3653 exceeds 1.2 Sm, and for Class 2 or 3 piping (ASME Section III) or non-nuclear piping, at locations where the maximum stress range as calculated by the sum of equations (9) and (10) of paragraph NC-3652 exceeds 0.4(1.2 Sh + SA).

The above criteria notwithstanding, cracks are not postulated in those portions of moderate energy piping located in the following areas:

a. Areas in which high energy pipe breaks are postulated, provided that moderate energy piping cracks would not result in more severe environmental conditions than the high energy pipe breaks.

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b. Between containment isolation valves, provided that:

1. The piping meets the requirements of subarticle NE- 1120 of the ASME B&PV Code

2. The maximum range of stress intensity for Class 1 piping (ASME Section III) as calculated by equation (10) of paragraph NB-3653 does not exceed 1.2 Sm, and the maximum stress range for Class 2 and 3 (ASME Section III) or non-nuclear piping as calculated by the sum of equations (9) and (10) of paragraph NC-3652 does not exceed 0.4(1.2 Sh + SA).

3.6.2.1.3 Types of Breaks and Cracks in Fluid System Piping Circumferential Breaks A circumferential break is assumed to result in (a) severance of a high energy pipe on a plane perpendicular to the pipe axis, and (b) separation amounting to at least a one diameter lateral displacement of the ruptured piping ends unless physically limited by piping restraints, structural members, or piping stiffness. 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.

All analyses to determine the blowdown (reaction) force on the segment of piping that contains a circumferential break are based on unobstructed discharge from 100% of the cross-sectional area of the pipe. This is consistent with the assumption of a one diameter lateral displacement of the ruptured piping ends. The only pipe break analyses that have involved lateral displacements of less than one pipe diameter are analyses concerning jet impingement forces. In certain cases where pipe whip restraints are located on both sides of a postulated circumferential break, and the design of the restraints prevents the two ends of the break from achieving a one diameter displacement, credit is taken for one end of the broken pipe causing partial blockage of the fluid being discharged from the opposite side of the break. Similarly, in certain cases where one side of the break is an RPV nozzle safe-end and the other side of the break is restrained from achieving a one diameter displacement relative to the nozzle, credit is taken for partial blockage of the fluid discharging from the restraint pipe end. This methodology can result in a reduction of the jet impingement force on potential impingement targets.

Circumferential breaks are postulated in high energy fluid system piping of nominal pipe size greater than 1-inch, at the locations determined by the criteria listed in Section 3.6.2.1.1, except where it can be shown that the maximum stress is in the circumferential direction and is at least 1.5 times the longitudinal stress, in which case only a longitudinal break is postulated.

Longitudinal Breaks A longitudinal break is assumed to result in an axial split parallel to the pipe axis, without causing pipe severance. The break opening area is assumed to be equal to the effective cross- sectional flow area of the pipe at the break location. The split is assumed to be oriented (but not concurrently) at two diametrically opposed points on the piping circumference so that the jet reaction force causes out-of-plane bending of the piping configuration. Piping movement is assumed to occur in the direction of the jet reaction unless limited by piping restraints, structural members, or piping stiffness.

CHAPTER 03 3.6-37 REV. 18, SEPTEMBER 2016

LGS UFSAR Longitudinal breaks are postulated in high energy fluid system piping of nominal pipe sizes of 4 inches and larger, at the locations determined by the criteria listed in Section 3.6.2.1.1, with the following exceptions. Longitudinal breaks are not postulated:

a. At terminal ends

b. At intermediate break locations chosen to satisfy the criterion for a minimum number of break locations

c. At locations where it can be shown that the maximum stress is in the longitudinal direction and is at least 1.5 times the circumferential stress, in which case only circumferential breaks need to be postulated.

Through-Wall Leakage Cracks Through-wall leakage cracks are postulated to occur in moderate energy fluid system piping exceeding a nominal pipe size of 1 inch, at the locations determined by the criteria listed in Section 3.6.2.1.2. A crack is assumed to occur at any orientation about the circumference of a pipe. Fluid flow from a crack is based on a circular opening with an area equal to that of a rectangle one-half pipe diameter in length and one-half pipe wall thickness in width.

3.6.2.2 Analytical Models to Define Forcing Functions and Response Models (Recirculation System Only) 3.6.2.2.1 Analytical Methods to Define Blowdown Forcing Functions The rupture of a pressurized pipe causes the flow characteristics of the system to change, creating reaction forces that can dynamically excite the piping system. The reaction forces are a function of time and space and depend upon the fluid state within the pipe prior to rupture, break flow area, frictional losses, plant system characteristics, piping system, and other factors. The methods used to calculate the reaction forces for recirculation system piping are presented below.

The criteria that are used for calculation of fluid blowdown forcing functions include:

a. 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 experimentally determined thrust coefficient. Limited pipe displacement at the break location, line restrictions, flow limiters, positive pump-controlled flow, and the absence of energy reservoirs may be taken into account, as applicable, in the reduction of jet discharge.

b. All breaks are assumed to attain full pipe break area instantaneously, i.e., a rise time not exceeding one millisecond is used for the initial pulse.

Blowdown forcing functions are determined by either of two methods as described below.

CHAPTER 03 3.6-38 REV. 18, SEPTEMBER 2016

LGS UFSAR Moody Model The predicated blowdown forces on pipes fed by a pressure vessel can be described by transient and steady-state forcing functions. The forcing functions used are based on methods described in Reference 3.6-4. These are simply described as follows:

a. 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).

b. 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 occur at the break end, changes in direction of piping, 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.

c. 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 (Po) 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 PoA).

d. Time histories of transient pressure, flow rate, and other thermodynamic properties of the fluid can be used to calculate the blowdown force on the pipe using the following equation:

F= [(P-Pa) + u2 ] A gc (EQ. 3.6-1) where:

F = blowdown force Pa = pressure at exit plane P = ambient pressure u = velocity at exit plant

= density at exit plane A = area of break gc = Newton's constant CHAPTER 03 3.6-39 REV. 18, SEPTEMBER 2016

LGS UFSAR

e. 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 accounting for these effects is presented in Reference 3.6-4. For subcooled water, a reduction from the theoretical maximum of 2.0 PoA is found through the use of Bernoulli's equation and standard equations such as Darcy's equation, which account for friction.

RELAP3 The computer code RELAP3 (Reference 3.6-5) 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:

T G E2VE

= PE Poo + (EQ. 3.6-2)

AE Gc R = - T

• AE AE (EQ. 3.6-3) where:

T = thrust per unit break area (lbf/ft2)

AE PE = exit pressure (lbf/ft2)

Poo = receiver pressure (lbf/ft2)

GE = exit mass flux (lbm/sec-ft2)

VE = exit specific volume (ft3/lbm) gc = Newton's constant (32.174 ft-lbm/lbf-sec2)

R = reaction force on the pipe (lbf) 3.6.2.2.2 Pipe Whip Dynamic Response Analyses The prediction of time-dependent and steady-thrust reaction loads caused by blowdown of subcooled, saturated, and two-phase fluid from a ruptured pipe is used in design and evaluation of dynamic effects of pipe breaks. A detailed discussion of the analytical methods employed to compute these blowdown loads for recirculation system piping is given in Section 3.6.2.2.1.

Analytical methods used to account for this loading are discussed below.

The criteria used for performing the pipe whip dynamic response analyses for recirculation system piping include:

a. A pipe whip analysis is performed for each postulated pipe break. However, a given analysis can be used for more than one postulated break location if the CHAPTER 03 3.6-40 REV. 18, SEPTEMBER 2016

LGS UFSAR 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 that 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 is assumed 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 absorber. Limits of strain are imposed that are similar to strain levels allowed in restraint plastic members. 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 to show that the consequences do not result in direct damage to any essential system or component.

f. Components such as vessel safe ends and valves that 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 limits imposed by the ASME B&PV Code for essential components under faulted loading.

The pipe whip analysis was performed using the PDA computer program (Reference 3.6-6). 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. A static nonlinear cantilever beam analysis is used for these locations to obtain the relationship between the pipe bending moment and the deflection (or rotation). Similar to the 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 pipe motion are formulated using an energy consideration, and the equations are numerically integrated in small time steps to yield time history information of the deformed pipe.

Considerable testing and analyses have demonstrated that potential rebound does not cause unacceptable increases in restraint deformation following the first quarter cycle loading for the GE restraint design and piping system experiencing blowdown thrust forces. Generic tests were performed on a 12-inch pipe size restraint with two primary loading configurations that represent the typical conditions during the postulated pipe rupture. Any other loading condition results in a combination of these two extremes. These loading configurations are:

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a. Load applied perpendicular to the restraint frame base against the cable

b. Load applied parallel to the base against one side of the frame.

The mass/inertia and stiffness properties of the recirculation system in the pipe dynamic analysis (PDA model) are represented as described below. A generic representation of the pipe in any given analysis is shown in Figure 3.6-41. If the stiffness of the piping segment located between A and B is such that:

a. The slope of BD at B = 0, then in the analysis, the pipe is treated as built-in at B.

b. The slope of BD at B =/ 0 (considerably different), then in the analysis, the pipe is considered to have a fixed, simple support (pinned end) at B.

To analyze the pipe with both ends supported (Figure 3.6-42) with the above computer model, two simplifications were made in the PDA program. First, an equivalent point mass is assumed at D instead of pipe length DE. The inertia characteristics of this mass rotating around point B are calculated to be identical to those of pipe length DE rotating around point E. Secondly, an equivalent resisting force is calculated for any deflection for the case of a built-in end from the bending moment-angular deflection relationship for pipe length DE. This equivalent force is subtracted from the applied thrust-force when calculating the net energy. The new model resulting from these simplifications is shown in Figure 3.6-42. The PDA computer program is further described in Section 3.9.1.2.2.6.

A comprehensive verification has been performed to demonstrate the conservatisms inherent in the PDA pipe whip computer program and the analytical methods utilized. This is described in Reference 3.6-7. Part of this verification program included an independent analysis of the recirculation system piping for the 1969 Standard Plant Design by Nuclear Services Corporation, under contract to GE. The recirculation system piping was chosen for study due to its complex piping arrangement and assorted pipe sizes. The analysis included elastic-plastic pipe properties, elastic-plastic restraint properties, and gaps between the restraint and pipe as documented in Reference 3.6-7. The piping/restraint system geometry and properties and fluid blowdown forces were the same in both analyses. However, a linear approximation was made by Nuclear Services Corporation for the restraint load-deflection curve supplied by GE. This approximation is demonstrated in Figure 3.6-36. 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 Nuclear Services Corporation 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 3.6-24.

A comparison of the Nuclear Services Corporation analysis with the PDA analysis, as presented in Table 3.6-25 and Figure 3.6-37, shows that PDA predicts higher loads in 15 of the 18 restraints analyzed. This is due to the Nuclear Services Corporation 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 Nuclear Services Corporation 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 CHAPTER 03 3.6-42 REV. 18, SEPTEMBER 2016

LGS UFSAR performance, and pipe movement predictions within the meaningful design requirements for these low probability postulated accidents.

3.6.2.3 Analytical Models to Define Forcing Functions and Response Models (Systems Other Than Recirculation System)

Analyses to determine the jet impingement effects and the piping and restraint displacements resulting from a pipe break are performed in accordance with BN-TOP-2 (Reference 3.6-8).

Analysis of jet thrust forces are described in section 2.2 of BN-TOP-2. Fluid jet impingement forces are discussed in section 2.3 of BN-TOP-2. Impulsive loading and impact combined with impulsive loading are described in sections 3.2 and 3.3, respectively, of BN-TOP-2.

Alternatively, nonlinear time history dynamic analyses are performed. The forcing function used in piping dynamic analysis is obtained using either Reference 3.6-2 or 3.6-9. A typical forcing function and the piping system model used for the dynamic response analysis is provided in Figure 3.6-38. Pipe restraint rebound effects are also considered in this analysis.

A specific case for which time history dynamic analyses are used is the analysis of pipe breaks near containment isolation valves whose operability following the break must be ensured. For each break postulated to occur near a containment isolation valve in a high energy line, a dynamic analysis is performed using the appropriate pipe break forcing function in order to determine the stresses in the pipe and the loads on the moment-limiting restraints near the valve. The dynamic analysis verifies that the stress in the pipe at the isolation valve is maintained below the yield strength of material, in order to ensure valve operability. Since the section modulus of the valve is much greater than that of the pipe, the stress in the valve body is kept below the yield strength of the valve. Therefore, deformation in the valve body as a result of nearby pipe breaks is severely limited and remains in the elastic range so that binding of the valve internals cannot occur.

The criteria for dynamic analyses when used in verifying valve operability are as follows:

a. An analysis of the piping system is performed for the most severe of the postulated longitudinal and circumferential breaks at the break locations determined in accordance with the criteria of Section 3.6.2.1.

b. The loading condition of a piping system prior to a postulated rupture in terms of internal pressure, temperature, and stress state is that condition associated with reactor operation at 100% power.

The basis of selecting 100% power as the loading condition of a piping system prior to rupture is as follows:

1. Pipe rupture analysis state-of-the-art involves several conservative steps and assumptions in all phases of break design (e.g., probability of break, the postulated speed of break propagation, the structural material properties, and the structural stability characteristics of pipe break restraint structures).

2. For those portions of piping systems that are normally pressurized during normal plant operation at power mode, the thermodynamic states in the piping systems are those of full (100%) thermal power.

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3. There is a much higher probability for scheduled plant operation at 100%

power (or less) than at higher ratings.

The combined effects of these considerations result in designs sufficiently capable of sustaining breaks at higher power levels.

c. For a circumferential break, pipe break dynamic transient force analyses are performed only for that end (or ends) of the pipe or branch that is connected to a contained fluid energy reservoir having sufficient capacity to develop a jet stream.

d. Dynamic analytic methods used for calculating the piping and piping/restraint system response to the pipe break forces adequately account for the effects of the following:

1. Translational masses (and rotational masses for major components) and stiffness properties of the piping system, restraint system, major components, and support walls

2. Transient forcing function(s) acting on the piping system

3. Elastic and inelastic deformation of piping and/or restraint

4. The design clearance between the pipe and the restraint

e. A 10% increase in minimum specified design yield strength (Sy) is used to account for strain rate effects in dynamic analyses.

The criteria for dynamic analyses when used in verifying the design of pipe whip restraints other than valve operability restraints are as follows:

a. A design analysis of the pipe whip restraint system is performed for each postulated longitudinal and circumferential break at the break locations determined in accordance with Section 3.6.2.1.

b. The loading condition of a piping system prior to a postulated break in terms of internal pressure, temperature, and stress state is that condition associated with reactor operation at 100% power.

c. For a circumferential break, pipe break dynamic transient forcing function calculations are performed only for that end (or ends) of the pipe or branch that is connected to a contained fluid energy reservoir having sufficient capacity to develop a jet stream.

d. An energy balance is established between the work done by the pipe break force on the pipe during the first quarter cycle of movement and the sum of the strain energy temporarily stored as elastic strain energy and that energy dissipated as plastic strain energy in members of the whip restraint.

CHAPTER 03 3.6-44 REV. 18, SEPTEMBER 2016

LGS UFSAR

e. In the event that the energy balance design analysis criteria are not satisfied, a dynamic analysis of the piping and piping restraint system is performed. This analysis is described by the same criteria described for valve operability dynamic analysis.

f. Dynamic analytical methods used for designing the pipe whip restraints adequately account for the effects of the following:

1. Transient forcing functions acting on the piping system.

2. Design clearance between the pipe and the restraint in the direction of pipe whip motion 3. Elastic stiffness and yield load capacity of the restraint under load applied in the direction of the pipe motion

4. Impact energy-absorbing capacity of crushable material used in the restraint is equivalent to 100% of the available capacity

5. Ductility capacity of the restraint loaded under impact beyond its yield load corresponds to not more than 50% of the ultimate uniform strain of the material.

3.6.2.4 Dynamic Analysis Methods to Verify Integrity and Operability (Recirculation System Only)

Pipe whip restraints, as differentiated from piping supports, are designed to function and carry load for an extremely low probability gross failure in a piping system carrying high energy fluid. The piping integrity does not usually depend on the piping whip restraints for any loading combination.

When the 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 pipe whip restraints (i.e., those devices that serve only to control the movement of a ruptured pipe following gross failure) will be subjected to once in a lifetime loading. For the purpose of design, the pipe break event is considered to be a faulted plant condition and the pipe, its restraints, and the structure to which the restraint is attached, are analyzed and designed accordingly.

As described in Section 3.6.1.2.1.1, the pipe whip restraints used for the recirculation system consist of straps (either carbon steel wire ropes or stainless steel bars) attached to a steel frame.

The analytical methods used in the design of these restraints are similar to those used for the Fermi Unit 2 and Duane Arnold plants. They have, however, been improved by incorporation of the latest force-deflection data available for wire rope and by using the GE PDA code for the dynamic analysis. Load capacities for the restraint frames were developed by using the SAP code (a finite-element structural analysis program), and were confirmed by a test series using slowly applied loading methods to determine restraint load-deflection data in the tangential direction (parallel to the restraint base). The results of this test program are presented in Reference 3.6-10.

The specific design objectives for the restraints are:

a. The restraints shall in no way increase the RCPB stresses by their presence during any normal mode of reactor operation or condition CHAPTER 03 3.6-45 REV. 18, SEPTEMBER 2016

LGS UFSAR

b. The restraint system shall function to stop the movement of pipe failure (gross loss of piping integrity) without allowing damage to critical components or missile development

c. The restraints shall provide minimum hinderance to inservice inspection of the process piping.

For the purposes of design, the pipe whip restraints are designed for the following dynamic loads:

a. Blowdown thrust of the pipe section that impacts the restraint

b. Dynamic inertia loads of the moving pipe section that is accelerated by the blowdown thrust and subsequent impact on the restraint

c. Design characteristics of the pipe whip restraints are included and verified by the pipe whip dynamic analysis described in Section 3.6.2.2.2

d. Since the pipe whip restraints are not contacted during normal plant operation, the postulated pipe rupture event is the only design loading condition.

As previously described, the recirculation loop pipe whip restraints are composed of two parts, the straps and the restraint frame. Both parts of the restraining device function as load-carrying members, and will deflect under load. The load configurations for a restraint are shown in Figure 3.6-3. The components of the restraints are categorized as Type I and II, as described below:

a. Type I - radial load-carrying members: These members, consisting of cables or bars, will absorb energy when loaded in the direction perpendicular to the restraint base by elastic and plastic deformations (Figure 3.6-3).

b. Type II - tangential load-carry members: These members, consisting of restraint frames, will absorb energy when loaded in the direction parallel to the base by plastic deformation (Figure 3.6-3).

Each of these components is constructed of a different material in order to fulfill different design objectives. The design requirements and design limits for each component are therefore different.

They are specified as below:

a. Type I - Straps

1. For carbon steel wire ropes, the maximum acceptable load is 90% of the load-carrying capacity of the cable in the restraint configuration. This limit takes into consideration the efficiency reduction experienced when a cable is wrapped around a pipe. This means that the design load is limited to about 75% of the minimum certified load-carrying capacity of the cable in tension.

2. For stainless steel bars, the design limit base was 50% of the minimum uniform ultimate tensile elongation.

CHAPTER 03 3.6-46 REV. 18, SEPTEMBER 2016

LGS UFSAR

b. Type 2 - Restraint Frames Design limits for the ASTM A36 restraint frames are as follows:

1. Design Load The load-bearing member is primarily a cantilever beam with an extra support (the diagonal plate) at approximately mid-span. At loads approaching the plastic moment capability of the beam, the plastic-hinge forms at the section determined by an elastic structural analysis. The maximum design load and the ultimate load are calculated based on plastic moment capability of this section, with the diagonal plate stressed uniformly at the minimum ultimate stress of 58,000 psi as specified for ASTM A36 material.

2. Design Deflection The design and ultimate deflection are calculated assuming the beam remains straight and rotates about a point on the upper surface of the beam.

The maximum design deflection at the load point is calculated assuming the diagonal plate undergoes 10% elongation. This corresponds to 50% of the minimum ultimate elongation of 20%, as specified for ASTM A36 material.

The ultimate deflection of the beam is based on 20% ultimate elongation of the diagonal plate.

3.6.2.5 Dynamic Analysis Methods to Verify Integrity and Operability (Systems Other Than Recirculation System The pipe whip restraints provided for protection from high energy pipe breaks are of two basic types: independent restraints and operability restraints. Independent restraints are provided solely to protect nearby structures and equipment from damage due to whipping pipes, and are designed so that a gap is maintained between the pipe and the restraint during normal plant conditions.

Operability restraints are provided near primary containment isolation valves whose operability is required following a break of the pipe in which they are installed. These operability restraints are designed to limit the stress in the piping near the valve to below the yield strength of the material in order to ensure operability of the valve. To accomplish this function, it is necessary to minimize the gap between the pipe and the restraint so that contact will occur during normal plant conditions.

The following high energy piping systems in the drywell are not provided with pipe whip restraints:

a. Reactor vessel drain line (4" DCA-101)

b. Main steam drain lines (2" and 3" DBA-105)

c. RPV head vent line (2" DBA-108)

d. SLCS injection line (2" DCA-112)

CHAPTER 03 3.6-47 REV. 18, SEPTEMBER 2016

LGS UFSAR Each of these lines has been evaluated, using the guidance of Regulatory Guide 1.46, to verify that in the event of a pipe break, damage to structures, systems, or components needed for safe shutdown would not occur. Therefore, the ability to shut the reactor down safely is maintained if a break should occur in any of these lines.

3.6.2.5.1 Design Loading Combinations The design loading combinations applied in the design of pipe whip restraints are categorized with respect to the plant operating conditions which are identified as normal, upset, emergency, and faulted as described in Section 3.9.3.1.1. Pipe break is considered as a faulted plant condition.

3.6.2.5.2 Design Stress Limits Operability Restraints When restraints for piping are designed so that contact between pipe and restraint will occur during normal plant conditions, the design loading combinations for normal, upset, emergency, and faulted conditions are applicable. In evaluating the supports and restraints for ASME Section III Class 1, 2, and 3, the design stress limits applied in evaluating loading combinations for normal, upset, emergency, and faulted (except for pipe rupture) conditions are those given in Tables 3.9-12 and 3.9-16. After rupture of the supported pipe occurs, the piping system is no longer within the jurisdiction of ASME Section III because the pressure boundary has been breached. The restraints are evaluated for pipe rupture loads as described in Section 3.6.2.3.

Pipe restraints are included on main steam, feedwater, RWCU, HPCI, and RCIC system to ensure the operability of the containment isolation valves during a postulated break event that touch the pipe during normal operation. The restraints are modeled in the thermal and dynamic analysis as active (1/16 inch or less gaps) during all loading conditions.

The operability of the isolation valves protected by operability restraints is assured by limiting the pipe break dynamic stress in the adjacent pipe. Stresses at the junction of this component with the pipe are limited to the dynamic yield strength of the pipe material (1.1 Sy). Between the containment penetration inboard/outboard isolation valves, pipe dynamic stress is limited to be less than 2.25 Sm.

Independent Restraints When restraints are designed solely to control movement following a postulated pipe rupture and to function independently of the normal support system, only the design pipe rupture loads are applicable.

To ensure that restraints function independently of the normal support system, the motions of the intact pipe due to all normal and upset plant conditions and the vibratory motion of the SSE are calculated and used to specify a minimum clearance between the pipe and the restraint. Wherever possible, gaps between pipes and restraints are maximized to avoid possible contact during plant operation. Where a particular location requires minimizing a gap, special features are provided to permit adjustment of the gap size during hot functional testing.

Independent restraints are evaluated for the pipe rupture loads as described in Section 3.6.2.3.

CHAPTER 03 3.6-48 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6.2.6 Guard Pipe Assembly Design Criteria Guard pipe assemblies are not used in this plant.

3.6.3 DEFINITIONS Certain terms used in Sections 3.6.1 and 3.6.2 have specified meanings as described below:

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

High Energy Fluid Systems - Fluid systems that, during normal plant conditions, are either in operation or maintained pressurized under conditions where either or both of the following are met:

a. Maximum operating temperature exceeds 200F

b. Maximum operating pressure exceeds 275 psig 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 are met:

a. Maximum operating temperature is 200F or less

b. Maximum operating pressure is 275 psig or less A system that qualifies as a high energy fluid system for only short periods and qualifies as moderate energy fluid system for the majority of the time is classified as a moderate energy fluid system provided that the total time the system operates within high energy pressure/temperature conditions is less than either of the following:

a. 2% of the time that the system operates as a moderate energy fluid system

b. 1% of the normal operating life span of the plant Normal Plant Conditions - Plant operating conditions during reactor startup, operation at power, hot standby, or reactor cooldown to cold shutdown condition.

Upset Plant 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.

Sh and SA - Allowable stresses at maximum (hot) temperature and allowable stress range for thermal expansion, respectively, as defined in ASME Section III, Article NC-3600.

Sm - Design stress intensity as defined in ASME Section III, Article NB-3600.

CHAPTER 03 3.6-49 REV. 18, SEPTEMBER 2016

LGS UFSAR Sn - Primary plus secondary stress intensity range for normal and upset conditions as defined in ASME Section III, paragraph NB-3653.

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 part of the single failure.

Terminal Ends - Extremities of piping runs that connect to structures, components (e.g., vessels, pumps, valves), or pipe anchors that act as rigid constraints to piping thermal expansion. A branch connection to a main piping run is a terminal end of the branch run, except when all three of the following conditions are in effect:

a. The nominal size of the branch run is at least half that of the main run

b. The intersection is not rigidly constrained to the building structure

c. The branch run and main run are included together in the same piping stress analysis model For piping in containment penetration areas, terminal ends are selected at points located immediately beyond the required moment-limiting restraints inside and outside containment. In piping runs which are maintained pressurized during normal plant conditions for only a portion of the run (i.e., up to the first normally closed valve), a terminal end of such runs is the piping connection to this closed valve.

3.

6.4 REFERENCES

3.6-1 "Subcompartment Pressure Analyses", BN-TOP-4, Rev. 0, Bechtel Power Corporation, San Francisco, California, (July 1976).

3.6-2 RELAP4/MOD5, "Computer Program for Transient Thermal-Hydraulic Analysis of Nuclear Reactors and Related Systems," ANCR-NUREG-1335, (September 1976).

3.6-3 FLUD, "Thermofluid Dynamics for a System of Interconnected Compartments,"

NE017, version 3, (July 18, 1980).

3.6-4 "System Criteria and Applications for Protection Against the Dynamic Effects of Pipe Break", GE Specification No. 22A2625.

3.6-5 RELAP3, "A Computer Program for Reactor Blowdown Analysis", IN-1321, Reactor Technology TID-4500, (June 1970).

3.6-6 "PDA - Pipe Dynamic Analysis Program for Pipe Rupture Movement" (proprietary filing), GE Report NEDE-10813.

3.6-7 "Final Report Pipe Rupture Analysis of Recirculation System for 1969 Standard Plant Design", Nuclear Services Corporation Report No. GEN-02-02.

CHAPTER 03 3.6-50 REV. 18, SEPTEMBER 2016

LGS UFSAR 3.6-8 "Design for Pipe Break Effects", BN-TOP-2, Rev. 2, Bechtel Power Corporation, San Francisco, California, (May 1974).

3.6-9 F.J. Moody, "Fluid Reaction and Impingement Loads", ASCE Specialty Conference on Structural Design of Nuclear Plant Facilities, Vol 1, pp. 219-262, (December 1973).

3.6-10 "Recirculation System Pipe Whip Restraint for the BWR 4, 218 and 251, Mark I and Mark II Product Line Plant", GE Design.

3.6-11 "Limerick Atomic Power Stations Units 1 & 2 Power Rerate Evaluation of Main Steam

& Recirculation Piping System, "GE Report GE-NE-123-E014-0193, Rev. 0 (July 1993) (part of N-00E-177-00003).

3.6-12 PCFLUD, "Thermofluid Dynamics for a System of Interconnected Compartments,"

MAP-120, Version 4.0 (July 15, 1992).

3.6-13 CONCOIL-FLUD (CFLUD), "Thermofluid Dynamics for a System of Interconnected Compartments," Version 1.0 (December 1, 1993).

3.6-14 GEH 0000-0158-9651-NP, Revision 0, October 2013, LGS Main Steam Isolation Valve Response Time Testing Analysis (SDOC G-080-VC-00489).

CHAPTER 03 3.6-51 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-1 HIGH ENERGY FLUID SYSTEM PIPING FLUID SYSTEM EXTENT OF HIGH ENERGY PIPING Reactor recirculation From reactor vessel suction nozzle to recirculation pump to reactor vessel discharge nozzles (Drawing M-43)

Main steam From reactor vessel nozzles to turbine stop valves (Drawings M-01 and M-41)

Feedwater From condensate filter/demineralizers through feedwater heaters and feedwater pumps to reactor vessel nozzles (Drawings M-06 and M-41)

Condensate From condensate pump discharge through steam jet air ejector condenser, steam packing exhauster, and condensate filter/ demineralizers (Drawings M-05 and M-16)

RWCU From shutdown cooling suction line through RWCU pumps, regenerative and non-regenerative heat exchangers, and cleanup filter/demineralizers to feedwater lines (Drawings M-43, M-44 and M-45)

Reactor vessel drain From reactor bottom head nozzle to RWCU line inside primary containment (Drawings M-43 and M-44)

HPCI steam supply From main steam line C to HPCI turbine steam supply valve (Drawing M-55)

RCIC steam supply From main steam line B to RCIC turbine steam supply valve (Drawings M-49 and M-50)

Main steam drain line From main steam lines inside drywell to inboard containment isolation valve; from main steam lines outside drywell to outboard containment isolation valve and drain line isolation valves (Drawing M-41)

CHAPTER 03 3.6-52 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-1 (Cont'd)

RPV head vent line From reactor vessel head nozzle to main steam line C (Drawing M-41)

SLCS From reactor vessel nozzle to first upstream check valve (Drawing M-42 and M-48)

RHR shutdown cooling From reactor recirculation loop to inboard suction containment isolation valve (Drawing M-51)

RHR shutdown cooling From reactor recirculation loop to first return upstream check valve (Drawing M-51)

LPCI injection From reactor vessel nozzle to first upstream check valve (Drawing M-51)

Core spray injection From reactor vessel nozzle to first upstream check valve (Drawing M-52)

CRD hydraulic From drive water pump to master control station to HCUs (Drawing M-46)

Auxiliary steam From auxiliary boiler to various steam-consuming components (Drawing M-21)

Plant heating steam From auxiliary boiler to various steam-consuming components CHAPTER 03 3.6-53 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-2 RECIRCULATION PIPING SYSTEM STRESS LEVELS AND PIPE BREAK DATA (5)(6)

(UNIT 1)

STRESS RATIOS PER ASME CODE EQUATIONS Basis for Node Eq.(10) Eq.(12) Eq.(13) Usage Break Break Point(1) 3 Sm (2) 3 Sm (2) 3 Sm (2) Factor Type(3) Selection(4)

A. RECIRCULATION LOOP A 001 0.630 0.079 0.606 0.00 C TE 500 1.225 0.447 0.847 0.06 L MBL 220 1.209 0.136 0.957 0.05 C&L MBL 236 0.734 0.088 0.646 0.00 C TE 200 1.39 0.098 0.993 0.15 C&L MBL 216 0.821 0.199 0.677 0.00 C TE 165 1.849 0.198 0.831 0.12 C&L MBL 296 0.786 0.099 0.698 0.00 C TE 240 1.532 0.312 0.948 0.29 C&L MBL 256 0.745 0.152 0.635 0.00 C TE 260 1.222 0.265 0.784 0.05 C&L MBL 276 0.727 0.218 0.560 0.00 C TE B. RECIRCULATION LOOP B 001 0.575 0.069 0.512 0.00 C TE 16 1.590 0.281 0.741 0.16 L MBL 800 1.179 0.401 0.826 0.04 L MBL 220 1.175 0.112 0.945 0.04 C&L MBL 236 0.708 0.074 0.642 0.00 C TE 200 1.355 0.077 1.006 0.13 C&L MBL 216 0.792 0.209 0.642 0.00 C TE 165 1.838 0.203 0.832 0.11 C&L MBL 296 0.817 0.093 0.739 0.00 C TE 240 1.499 0.326 0.894 0.25 C&L MBL 256 0.771 0.176 0.648 0.00 C TE 260 1.235 0.274 0.799 0.06 C&L MBL 276 0.793 0.223 0.585 0.00 C TE CHAPTER 03 3.6-54 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-2 (Cont'd)

RECIRCULATION PIPING SYSTEM STRESS LEVELS AND PIPE BREAK DATA (UNIT 2)

STRESS RATIOS PER ASME CODE EQUATIONS Basis for Node Eq.(10) Eq.(12) Eq.(13) Usage Break Break Point(1) 3 Sm (2) 3 Sm (2) 3 Sm (2) Factor Type(3) Selection(4)

A. RECIRCULATION LOOP A 001 0.40 0.08 0.37 0.00 C TE 236 0.79 0.40 0.40 0.00 C TE 216 0.79 0.37 0.44 0.00 C TE 296 0.94 0.51 0.45 0.00 C TE 256 1.18 0.79 0.43 0.02 C TE 276 0.73 0.34 0.40 0.00 C TE B. RECIRCULATION LOOP B 001 0.43 0.06 0.41 0.00 C TE 236 0.57 0.07 0.52 0.00 C TE 216 0.72 0.19 0.54 0.00 C TE 296 0.63 0.09 0.57 0.00 C TE 256 0.70 0.16 0.54 0.00 C TE 276 0.70 0.20 0.53 0.00 C TE (1) Locations of the nodes are shown in Figure 3.6-4 (2) Sm: Design stress intensity as defined in ASME Section III, Article NB-3600 (Section 3.6.3)

(3) Break types are indicated as follows:

C - Circumferential L - Longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - Terminal end MBL - Intermediate break locations selected to satisfy the requirement for a minimum number of break locations.

(5) The recirculation piping design basis has been evaluated for the effects of power rerate and demonstrated to be adequate for the increases in pressure, temperature and flow due to power rerate. Detail of the evaluation performed are documented in Reference 3.6-11.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-55 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-3 REACTOR RECIRCULATION PIPING SYSTEM FLUID BLOWDOWN THRUST TIME HISTORIES (UNIT 1)

(Refer to Diagram A below)

Initial Intermediate Steady-State Time Duration Time to Reach Node Point of Break Side of Force Fo Force Fint Force Fss sec) (sec)

Break Location(1) Type Break (kips) (kips) (kips) of Fo(T1) Steady-State (T2) 001 CRCMF 540.00 476.80 179.80 0.00186 0.08006 16 (loop B only) LONG 540.00 564.30 621.54 0.00064 0.0032 220 CRCMF PUMP 373.48 291.98 330.53 0.00145 0.0210 236 CRCMF PUMP 134.42 101.48 118.96 0.00143 0.02098 200 CRCMF PUMP 373.48 281.78 330.53 0.00145 0.021 216 CRCMF PUMP 134.42 101.48 118.96 0.00143 0.02098 296 CRCMF PUMP 134.42 101.48 118.96 0.00143 0.02098 240 CRCMF PUMP 373.48 281.98 330.53 0.00145 0.021 256 CRCMF PUMP 134.42 101.48 118.96 0.00143 0.02098 260 CRCMF PUMP 373.48 281.98 330.53 0.00145 0.021 276 CRCMF PUMP 134.42 101.48 118.96 0.00143 0.02098 220 CRCMF VESSEL 373.48 373.48 373.48 0.0038 0.0368 200 CRCMF VESSEL 373.48 373.48 373.48 0.0038 0.0368 165 CRCMF VESSEL 373.48 373.48 373.48 0.0038 0.0368 240 CRCMF VESSEL 373.48 373.48 373.48 0.0038 0.0368 260 CRCMF VESSEL 373.48 373.48 373.48 0.0038 0.0368 CHAPTER 03 3.6-56 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-3 (Cont'd)

REACTOR RECIRCULATION PIPING SYSTEM FLUID BLOWDOWN THRUST TIME HISTORIES (UNIT 2)

(Refer to Diagram A below)

Initial Intermediate Steady-State Time Duration Time to Reach Node Point of Break Side of Force Fo Force Fint Force Fss sec) (sec)

Break Location(1) Type Break (kips) (kips) (kips) of Fo(T1) Steady-State (T2) 001 C 540.00 476.80 179.80 0.00186 0.08006 236 C PUMP 134.42 101.48 118.96 0.00143 0.02098 216 C PUMP 134.42 101.48 118.96 0.00143 0.02098 296 C PUMP 134.42 101.48 118.96 0.00143 0.02098 256 C PUMP 134.42 101.48 118.96 0.00143 0.02098 276 C PUMP 134.42 101.48 118.96 0.00143 0.02098 Fo (1)

SEE FIGURE 3.6-4 FOR IDENTIFICATION OF POSTULATED BREAK LOCATIONS.

Fint Fss T1 T2 Diagram A CHAPTER 03 3.6-57 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-4 MAIN STEAM PIPING STRESS LEVELS AND PIPE BREAK DATA (7)(8)

(UNIT 1)

Stress Cumulative Pipe Break Basis for Node Node By Eq. 10 Usage Stress Limit Break Break Point(1) Type(2) (ksi)(6) Factor 2.4 Sm(ksi) Type(3) Selection(4)

LINE A 1 BW 25.94 0.0 42.21 C TE 3 CURB 51.45 0.14 42.21 C&L SFL 3 CURE 50.18 0.20 42.21 C&L SFL 5 SWE (5) (5) 42.21 C TE (TYP for 3 PSVs) 6 FLA (5) (5) 42.21 C TE (TYP for 3 PSVs) 24 CUR 39.03 0.01 42.21 C MBL 40 BW 34.72 0.01 42.21 C TE LINE B 1 BW 24.81 0.0 42.21 C TE 3 CURB 48.56 0.11 42.21 C&L SFL 3 CURE 45.87 0.19 42.21 C&L SFL 5 SWE (5) (5) 42.21 C TE (TYP for 4 PSVs) 6 FLA (5) (5) 42.21 C TE (TYP for 4 PSVs) 23 CUR 36.19 0.01 42.21 C MBL 46 BW 33.09 0.01 42.21 C TE LINE C 2 BW 24.03 0.0 42.21 C TE 3 CURB 46.48 0.10 42.21 C&L SFL 4 CURE 44.56 0.17 42.21 C&L SFL 5 SWE (5) (5) 42.21 C TE (TYP for 3 PSVs) 6 FLA (5) (5) 42.21 C TE (TYP for 3 PSVs) 26 CUR 36.73 0.01 42.21 C MBL 48 BW 34.07 0.01 42.21 C TE CHAPTER 03 3.6-58 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-4 (Cont'd)

(UNIT 1)

Stress Cumulative Pipe Break Basis for Node Node By Eq. 10 Usage Stress Limit Break Break Point(1) Type(2) (ksi)(6) Factor 2.4 Sm(ksi) Type(3) Selection(4)

LINE D 1 BW 33.21 0.01 42.21 C TE 3 CURB 50.28 0.11 42.21 C&L SFL 3 CURE 48.04 0.13 42.21 C&L SFL 5 SWE (5) (5) 42.21 C TE (TYP for 4 PSVs) 6 FLA (5) (5) 42.21 C TE (TYP for 4 PSVs) 31 CUR 39.09 0.01 42.21 C MBL 47 BW 34.13 0.01 42.21 C TE CHAPTER 03 3.6-59 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-4 (Cont'd)

MAIN STEAM PIPE STRESS LEVELS AND PIPE BREAK DATA (UNIT 2)

Stress Cumulative Pipe Break Basis for Node Node By Eq. 10 Usage Stress Limit Break Break Point(1) Type(2) (ksi)(6) Factor 2.4 Sm(ksi) Type(3) Selection(4)

LINE A 1 BW 32.44 0.007 42.48 C TE 5 SWE 41.77 0.251 42.48 C TE (TYP for 3 PSVs) 6 FLA (5) (5) 42.48 C TE (TYP for 3 PSVs) 40 BW 32.30 0.008 42.48 C TE LINE B 1 BW 32.00 0.007 42.48 C TE 5 SWE 37.58 0.218 42.48 C TE (TYP for 4 PSVs) 6 FLA (5) (5) 42.48 C TE (TYPE for 4 PSVs) 46 BW 32.60 0.009 42.48 C TE LINE C 2 BW 29.77 0.007 42.48 C TE 5 SWE 41.65 0.294 42.48 C TE (TYP for 3 PSVs) 6 FLA (5) (5) 42.48 C TE (TYP for 3 PSVs) 48 BW 33.78 0.010 42.48 C TE LINE D 1 BW 29.95 0.006 42.48 C TE 5 SWE 40.32 0.126 42.48 C TE (TYP for 4 PSVs) 6 FLA (5) (5) 42.48 C TE (TYP for 4 PSVs) 47 BW 32.18 0.007 42.48 C TE CHAPTER 03 3.6-60 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-4 (Cont'd)

(1)

Locations of nodes are shown in Figure 3.6-8.

(2)

Node types are designated as follows:

BW - Butt-welding tee CUR - Elbow CURB - Elbow beginning CURE - Elbow ending SWE - Sweepolet FLA - Flange (3)

Break types are indicated as follows:

C - Circumferential L - Longitudinal (4)

Symbols used to denote the basis for break selection are as follows:

TE - Terminal end SFL - The stress and fatigue limits established in Section 3.6.2.1.1.3 are not met.

MBL - Intermediate break location selected to satisfy the criteria for a minimum number of break locations.

(5)

Stress values not calculated; terminal end break assumed.

(6)

Refer to Section 3.6.2.1.1.3.

(7)

The main steam piping design basis has been evaluated for the effects of power rerate and demonstrated to be adequate for the increase in pressure, temperature and flow due to power rerate. Details of the evaluation performed are documented in Reference 3.6-11.

(8)

The information posted in this table was used for original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-61 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-5 INTENTIONALLY LEFT BLANK CHAPTER 03 3.6-62 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-6 BLOWDOWN DATA FOR HIGH ENERGY PIPE BREAKS OUTSIDE PRIMARY CONTAINMENT BLOWDOWN (2) ISOLATION VALVE CLOSURE TIME AFTER MASS VALVE SIGNAL TOTAL BREAK FLOW RATE ENTHALPY ISOLATION CLOSING TIME DELAY TIME INTERVAL HIGH ENERGY LINE (sec) (lb/sec) (Btu/lb)(3) VALVES (sec) (sec) (sec) .

Main steam line 0.00 13,356 1192.4 HV41-F022A,B,C&D 5.0 1.0 6.0 (26 inch EBB-101, EBB-102, 0.076 8,543 1192.4 HV41-F028A,B,C&D 5.0 1.0 6.0 EBB-103, or EBB-104) 0.16 9,177 1192.4 1.0 15,972 1192.4 1.001 19,398 595.0 5.0 19,398 595.0 6.0 0 595.0 RWCU suction line 0.0 2,820 525.3 HV44-F001 10.0 max. 2.0 12.0 max.

(6 inch DCC-103) 0.1 2,319 524.7 HV44-F004 10.0 max. 2.0 12.0 max.

0.2 1,818 513.0 HV44-F039(1) - - -

0.3 1,486 509.8 0.4 1,200 506.3 0.5 1,027 503.7 0.6 905 501.3 0.7 791 497.4 0.8 723 495.0 0.9 701 491.0 1.0 638 485.1 1.26 563 485.0 6.0 563 485.0 9.0 0 485.0 RWCU pump discharge line

• 0.00 1410

• 526.4 HV44-F001 10.0 max. 2.0 12.0 max.

(4 inch DCC-101) 0.02 912

• 526.4 HV44-F004 10.0 max. 2.0 12.0 max.

0.03 623

• 526.4 HV44-F039(1) - - -

0.11 473

• 526.4 0.25 688

• 526.4 0.37 583

• 526.4 2.00 531

• 526.4 15.00 273

• 526.4

• Mass flow rate scale from values for 3 break to 4 break by using the ratio of the break areas as a scaling factor (A4/A3) = 1.74.

CHAPTER 03 3.6-63 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-6 (Cont'd)

BLOWDOWN (2) ISOLATION VALVE CLOSURE TIME AFTER MASS VALVE SIGNAL TOTAL BREAK FLOW RATE ENTHALPY ISOLATION CLOSING TIME DELAY TIME INTERVAL HIGH ENERGY LINE (sec) (lb/sec) (Btu/lb)(3) VALVES (sec) (sec) (sec) .

RWCU pump discharge line 0.0 1,164 526.4 HV44-F001 10.0 max. 2.0 12.0 max.

at inlet to regenerative 0.2 947 526.4 HV44-F004 10.0 max. 2.0 12.0 max.

heat exchanger 0.3 904 526.4 HV44-F039(1) - - -

(4 inch DCC-101) 0.35 1,051 526.4 1.0 682 526.4 1.5 503 526.4 2.0 471 526.4 16.0 223 526.4 RWCU pump discharge line 0.00 1,164 203.9 HV44-F001 10.0 max. 2.0 12.0 max.

at inlet to nonregenerative 0.34 1,056 203.9 HV44-F004 10.0 max. 2.0 12.0 max.

heat exchanger(4 inch DCC-102) 0.59 868 203.9 HV44-F039(1) - - -

0.59 868 125.2 1.00 677 131.4 3.72 464 149.4 3.72 464 316.2 16.00 464 316.2 16.00 224 91.0 20.00 224 91.0 HPCI steam supply line 0.0 1,470 1192.4 HV55-F002 12.0 1.0 13.0 at turbine inlet valve 0.24 1,045 1192.4 HV55-F003 12.0 1.0 13.0 (12 inch EBB-108) 0.36 280 1192.4 13.0 280 1192.4 14.0 0 1192.4 HPCI steam supply line 0.0 2,940 1192.4 HV55-F002 12.0 1.0 13.0 in piping area 0.11 1,958 1192.4 HV55-F003 12.0 1.0 13.0 (12 inch EBB-108) 0.14 1,594 1192.4 0.22 266 1192.4 13.0 266 1192.4 14.0 0 1192.4 CHAPTER 03 3.6-64 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-6 (Cont'd)

BLOWDOWN (2) ISOLATION VALVE CLOSURE TIME AFTER MASS VALVE SIGNAL TOTAL BREAK FLOW RATE ENTHALPY ISOLATION CLOSING TIME DELAY TIME INTERVAL HIGH ENERGY LINE (sec) (lb/sec) (Btu/lb)(3) VALVES (sec) (sec) (sec) .

HPCI steam supply line 0.0 2,940 1192.4 HV55-F002 12.0 1.0 13.0 in isolation valve compartment 0.135 1,272 1192.4 HV55-F003 12.0 1.0 13.0 (12 inch EBB-108) 0.23 902 1192.4 0.475 328 1192.4 13.0 328 1192.4 14.0 0 1192.4 RCIC steam supply line 0.0 380 1192.4 HV49-F007 7.2 1.0 8.2 at turbine inlet valve 0.311 168 1192.4 HV49-F008 7.2 1.0 8.2 (6 inch EBB-109) 0.43 40 1192.4 7.2 40 1192.4 8.2 0 1192.4 RCIC steam supply line 0.0 760 1192.4 HV49-F007 7.2 1.0 8.2 in upper pipe tunnel 0.13 382 1192.4 HV49-F008 7.2 1.0 8.2 (6 inch EBB-109) 0.26 85.4 1192.4 0.302 42 1192.4 7.2 42 1192.4 8.2 0 1192.4 (1)

Valve closure time is not applicable for HV44-F039 since it is a check valve. This valve prevents backflow of water from the feedwater lines into the RWCU equipment compartments in the event of a break.

(2)

The blowdown table is based on original power level. Environmental effects from blowdown are addressed based 3527 MWt conditions in Table 3.6-7 and 3.6-9, which bounds the operation at MUR power level of 3515 MWt.

(3)

Enthalpy of system is conservatively assumed for saturated steam at 1000 psig (0% moisture carryover) based on original design conditions.

CHAPTER 03 3.6-65 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 PRESSURE-TEMPERATURE TRANSIENT ANALYSIS RESULTS FOR (6)

HIGH ENERGY PIPE BREAKS OUTSIDE PRIMARY CONTAINMENT (Unit 1)

TIME(5) TIME (5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F) (sec)

A. Main Steam Line Break in Main Steam Tunnel

1. Main steam tunnel 10.18 3.0 321 0.60

2. Main steam tunnel 8.18 3.1 324 0.62 vent stack (lower-region)

3. Main steam tunnel 5.32 3.1 325 0.64 vent stack (mid-region)

4. Main steam tunnel 2.93 3.2 319 0.74 vent stack (upper-region)

5. Main steam tunnel security -- -- -- --

plenum (U2 Only)

6. Main condenser area 0.56 0.23 179 5.8

7. Steam venting plenum 0.61 0.23 182 5.5 B. Main Steam Line Break in Main Condenser Area

6. Main condenser area 2.33 4.2 208 4.95

7. Steam venting plenum 2.33 4.3 208 5.16 C. RWCU Suction Line Break in Penetration Room

6. Nonregenerative heat 3.02 0.65 129(4) 0.51 exchanger room "A"

7. Nonregenerative heat 3.02 0.64 128(4) 0.51 exchanger room "B"

9. Regenerative heat 2.92 0.65 127(4) 0.65 exchanger room

10. RWCU pump-room 2.91 0.38 105(4) 0.80

13. RWCU penetration room 2.92 0.40 202(4) 7.12 CHAPTER 03 3.6-66 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 1)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F) (sec)

D. RWCU Pump Discharge Line Break in Pump-Room

6. Nonregenerative heat 0.44(4) 0.78 115(4) 0.78 exchanger room "A"

7. Nonregenerative heat 0.45(4) 0.76 115(4) 0.76 exchanger room "B"

9. Regenerative heat 0.44(4) 0.74 112(4) 0.29 exchanger room

10. RWCU pump-room (B, C only) 1.96(4) 0.33 222 3.07 10A.RWCU pump room (A only) 6.2 0.33 229 3.07

13. RWCU penetration room 0.44(4) 0.87 206(4) 15.00 E. RWCU Pump Discharge Line Break in Regenerative Heat Exchanger Room

6. Nonregenerative heat 2.45(4) 0.92 113(4) 0.11 exchanger room "A"

7. Nonregenerative heat 2.45(4) 0.92 112(4) 0.11 exchanger room "B"

9. Regenerative heat 2.42(4) 1.01 221 6.65 exchanger room

10. RWCU pump-room 1.78(4) 1.13 113(4) 0.43

13. RWCU penetration room 1.77(4) 1.16 215 15.95 F. RWCU Pump Discharge Line Break in Nonregenerative Heat Exchanger Room "A"

6. Nonregenerative heat 1.56(4) 10.18 221 16.00 exchanger room "A"

7. Nonregenerative heat 1.06(4) 10.82 111(4) 0.10 exchanger room "B"

9. Regenerative heat 0.50(4) 15.44 208(4) 16.13 exchanger room

10. RWCU pump-room 0.36(4) 16.23 107(4) 0.38

13. RWCU penetration room 0.36(4) 16.20 183(4) 16.43 CHAPTER 03 3.6-67 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 1)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F) (sec)

G. RWCU Pump Discharge Line Break in Nonregenerative Heat Exchanger Room "B"

6. Nonregenerative heat 1.07(4) 10.73 111(4) 0.10 exchanger room "A"

7. Nonregenerative heat 2.03(4) 10.03 217(4) 16.00 exchanger room "B"

9. Regenerative heat 0.50(4) 15.55 208(4) 16.17 exchanger room

10. RWCU pump-room 0.36(4) 16.20 107(4) 0.40

13. RWCU penetration room 0.36(4) 16.23 183(4) 16.49 H. HPCI Steam Supply Line Break in HPCI Pump-Room

17. HPCI pump-room 2.94 0.24 307 15.16

18. HPCI piping area 2.22 0.25 308 15.20

21. Isolation valve 1.03 0.33 241 14.00 compartment

22. Steam venting tunnel 0.61 0.33 238 15.99 I. HPCI Steam Supply Line Break in HPCI Piping Area

17. HPCI pump-room 2.54 0.14 264 15.51

18. HPCI piping area 6.64 0.13 295 0.13

21. Isolation valve 1.52 0.21 241 15.71 compartment

22. Steam venting tunnel 0.91 0.21 235 12.88 J. HPCI Steam Supply Line Break in Isolation Valve Compartment

21. Isolation valve 1.51 0.18 273 16.55 compartment

22. Steam venting tunnel 0.98 0.18 270 15.93 CHAPTER 03 3.6-68 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 1)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F) (sec)

K. RCIC Steam Supply Line Break in RCIC Pump-Room

19. RCIC pump-room 2.94 0.27 229 11.17

20. RCIC upper pipe tunnel 2.56 0.29 218 11.15

21. Isolation valve 0.50 0.38 129 13.99 compartment

22. Steam venting tunnel 0.50 0.38 128 0.38 L. RCIC Steam Supply Line Break in RCIC Upper Pipe Tunnel

19. RCIC pump-room 2.68 0.16 153 0.16

20. RCIC upper pipe tunnel 5.77 0.03 306 7.04

21. Isolation valve 0.50 0.22 142 13.99 compartment

22. Steam venting tunnel 0.50 0.22 135 11.12 CHAPTER 03 3.6-69 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

PRESSURE-TEMPERATURE TRANSIENT ANALYSIS RESULTS FOR HIGH ENERGY PIPE BREAKS OUTSIDE PRIMARY CONTAINMENT (Unit 2)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F)(3) (sec)

A. Main Steam Line Break in Main Steam Tunnel

1. Main steam tunnel 11.39 3.1 320 0.60

2. Main steam tunnel 9.78 3.2 325 0.62 vent stack (lower-region)

3. Main steam vent stack 7.81 3.2 325 0.67 (security plenum)

4. Main steam tunnel 2.72 3.4 319 0.75 vent stack (upper-region)

5. Main steam tunnel 10.30 3.2 320 0.60 security plenum

6. Main condenser area 0.5 0.23 182 5.9

7. Steam venting plenum 0.54 0.23 187 5.6 B. Main Steam Line Break in Main Condenser Area

6. Main condenser area 2.33 4.2 208 5.0

7. Steam venting plenum 2.33 4.3 208 5.2 C. RWCU Suction Line Break in Penetration Room

6. Nonregenerative heat 3.02 0.65 129(4) 0.51 exchanger room "A"

7. Nonregenerative heat 3.02 0.64 128(4) 0.51 exchanger room "B"

9. Regenerative heat 2.92 0.65 127(4) 0.65 exchanger room

10. RWCU pump-room 2.91 0.38 105(4) 0.80

13. RWCU penetration room 2.92 0.40 202(4) 7.12 CHAPTER 03 3.6-70 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 2)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F)(3) (sec)

D. RWCU Pump Discharge Line Break in Pump-Room

6. Nonregenerative heat 0.44(4) 0.78 115(4) 0.78 exchanger room "A"

7. Nonregenerative heat 0.45(4) 0.76 115(4) 0.76 exchanger room "B"

9. Regenerative heat 0.44(4) 0.74 112(4) 0.29 exchanger room

10. RWCU pump-room(B, C only) 1.96(4) 0.33 222 3.07 10A.RWCU pump-room (A only) 6.2 0.33 229 3.07

13. RWCU penetration room 0.44(4) 0.87 206(4) 15.00 E. RWCU Pump Discharge Line Break in Regenerative Heat Exchanger Room

6. Nonregenerative heat 2.4(4) 0.92 113(4) 0.11 exchanger room "A"

7. Nonregenerative heat 2.45(4) 0.92 11(4) 0.11 exchanger room "B"

9. Regenerative heat 2.42(4) 1.01 221 6.65 exchanger room

10. RWCU pump-room 1.78(4) 1.13 113(4) 0.43

13. RWCU penetration room 1.77(4) 1.16 215 15.95 F. RWCU Pump Discharge Line Break in Nonregenerative Heat Exchanger Room "A"

6. Nonregenerative heat 1.56(4) 10.18 221 16.00 exchanger room "A"

7. Nonregenerative heat 1.06(4) 10.82 111(4) 0.10 exchanger room "B"

9. Regenerative heat 0.50(4) 15.44 208(4) 16.13 exchanger room

10. RWCU pump-room 0.36(4) 16.23 107(4) 0.38

13. RWCU penetration room 0.36(4) 16.20 183(4) 16.43 CHAPTER 03 3.6-71 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 2)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F)(3) (sec)

G. RWCU Pump Discharge Line Break in Nonregenerative Heat Exchanger Room "B"

6. Nonregenerative heat 1.11(4) 10.73 111(4) 0.10 exchanger room "A"

7. Nonregenerative heat 2.11(4) 10.03 214(4) 16.00 exchanger room "B"

9. Regenerative heat 0.50(4) 15.55 206(4) 16.17 exchanger room

10. RWCU pump-room 0.38(4) 16.20 107(4) 0.40

13. RWCU penetration room 0.38(4) 16.23 184(4) 16.49 H. HPCI Steam Supply Line Break in HPCI Pump-Room

17. HPCI pump-room 2 50 0.25 299 15.17

18. HPCI piping area 2.03 0.25 298 15.20

21. Isolation valve 1.04 0.34 241 17.00 compartment

22. Steam venting tunnel 0.68 0.35 239 16.03

23. Security plenum 0.51 0.35 238 15.98 I. HPCI Steam Supply Line Break in HPCI Piping Area

17. HPCI pump-room 2.79 0.14 243 0.14

18. HPCI piping area 2.79 0.14 299 4.74

21. Isolation valve 1.39 0.22 271 17.00 compartment

22. Steam venting tunnel 0.97 0.22 260 16.04

23. Security plenum 0.72 0.22 260 15.97 CHAPTER 03 3.6-72 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-7 (Cont'd)

(Unit 2)

TIME(5) TIME(5)

PEAK(2) AFTER PEAK AFTER PRESSURE BREAK TEMPERATURE BREAK COMPARTMENT(1) (psig) (sec) (F)(3) (sec)

J. HPCI Steam Supply Line Break in Isolation Valve Compartment

21. Isolation valve 1.60 0.20 271 16.99 compartment

22. Steam venting tunnel 1.21 0.20 270 16.06

23. Security plenum 0.83 0.20 270 15.97 K. RCIC Steam Supply Line Break in RCIC Pump-Room

19. RCIC pump-room 1.84 0.27 228 11.21

20. RCIC upper pipe tunnel 1.47 0.29 218 11.19

21. Isolation valve 0.50 0.38 129 14.00 compartment

22. Steam venting tunnel 0.50 0.38 128 0.38

23. Security plenum 0.50 0.38 128 0.38 L. RCIC Steam Supply Line Break in RCIC Upper Pipe Tunnel

19. RCIC pump-room 2.00 0.16 152 0.16

20. RCIC upper pipe tunnel 2.47 0.03 295 6.97

21. Isolation valve 0.50 0.22 142 14.00 compartment

22. Steam venting tunnel 0.50 0.22 137 12.31

23. Security plenum 0.50 0.22 135 11.10 (1) Compartment numbers used in this table correspond to the compartment numbers used in the flow models (Figures 3.6-11, 3.6-12, 3.6-19, 3.6-23, 3.6-24, and 3.6-27).

(2) The compartment design pressures and the pressure-temperature transient analysis results are in Table 3.6-9.

(3) For Unit 2, design bulk temperatures may be less. Note also that temperatures in this area due to breaks elsewhere may be bounding.

(4) The value shown is based on original power level. It is bounded by another break in this compartment where the value shown is for the 3527 MWt power level.

(5) Time shown is based on original power level and was not recalculated for rerate since it is not used for any design basis evaluations.

(6) Values shown unless noted are based on power level of 3527 MWt. The values (except the main steam line valves) were established based on the original values and a multiplier. The multiplier was calculated based on a maximum pressure increase associated with a power level of 3527 MWt and its impact on the blowdown and subsequent impact to subcompartment pressures and temperatures.

CHAPTER 03 3.6-73 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-8 FEEDWATER PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION INSIDE PRIMARY CONTAINMENT) (5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 10 TTJ 77.05 0.0603 42.21 C TE 45 SWP 91.10 0.5213 42.21 C TE 55 TEE 104.58 0.4841 42.21 C&L SFL 70 EL 69.14 0.1374 42.21 C&L SFL 75 TTJ 76.74 0.6192 42.21 C TE 100 TEE 100.79 0.3651 42.21 C&L SFL 110 EL 72.19 0.1337 42.21 C&L SFL 115 TTJ 79.85 0.6136 42.21 C TE 170 EL 73.22 0.1316 42.21 C&L SFL 180 TTJ 78.87 0.6103 42.21 C TE 197 TTJ 74.17 0.0575 42.21 C TE UNIT 2 10 TTJ 43.06 0.0531 42.21 C TE 45 SWP 70.45 0.2216 42.21 C TE 55 TEE 87.89 0.3960 42.21 C&L SFL 115 TTJ 56.98 0.2482 42.21 C TE 180 TTJ 61.18 0.2388 42.21 C TE 197 TTJ 68.59 0.1942 42.21 C TE 75 TTJ 59.97 0.2699 42.21 C TE 100 TEE 102.14 0.8011 42.21 C&L SFL (1) Locations of the nodes listed in this table are shown in Figure 3.6-15.

(2) Node types are designated as follows:

TEE - butt-welding tee TTJ - tapered transition joint EL - elbow SWP - sweepolet (3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) The values provided above are based upon original licensed power conditions. These CUF values are conservatively based derived by using Sm for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-74 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-9 COMPARTMENT DESIGN PRESSURES AND PRESSURE-TEMPERATURE TRANSIENT ANALYSIS RESULTS UNIT 1 UNIT 2 (2)

MAX TRANSIENT MAX TRANSIENT(2)

DESIGN PRESSURE(1) PEAK PRESSURE DESIGN PRESSURE(1) PEAK PRESSURE COMPARTMENT PSIG PSIG PSIG PSIG .

Main Steam Tunnel 10.2 10.18 10.47 11.39 Main Steam Tunnel Vent Stack (Lower-Region) 8.2 8.18 8.99 9.78 Main Steam Tunnel Vent Stack (Mid-Region) 5.7 5.32 -- --

Main Steam Tunnel Vent Stack (Security Plenum) -- -- 6.97 7.81 Main Steam Tunnel Vent Stack (Upper-Region) 5.7 2.93 5.7 2.72 MST Security Plenum -- -- 2.14 2.33 Main Condenser Area 2.14 2.33 2.14 2.33 Steam Vent Plenum 2.14 2.33 10.47 10.30 HPCI Pump 2.7 2.94 2.7 2.79 HPCI Piping Area 6.1 6.64 6.1 2.79 RCIC Pump 2.8 2.94 2.8 2.00 RCIC Upper Piping Area 5.4 5.77 5.4 2.47 RHR, HPCI, RCIC Isolation Valve 1.41 1.52 1.47 1.60 HPCI, RCIC Steam Venting Tunnel 0.9 0.98 1.11 1.21 Security Plenum -- -- 1.11 0.83 Non-Regenerative Heat Exchanger 2.9 3.02 2.9 3.02 Non-Regenerative Heat Exchanger 2.9 3.02 2.9 3.02 Regenerative Heat Exchanger 2.9 2.92 2.9 2.92 RWCU Pump (A) 2.8 6.2 2.8 6.2 RWCU Pump (B and C) 2.8 2.91 2.8 2.91 RWCU Penetration 2.8 2.92 2.8 2.92 (1)

In each compartment, the maximum pre-rerate peak transient pressure was used as a basis to establish the compartment design pressure. This is considered to be appropriate because of the conservatism in the analytical models used to calculate the mass and energy release rates.

(2)

Maximum transient peak pressures are for power rerate conditions (3527 mwt). Although some peak pressures exceed listed design pressures, there are sufficient margin and conservatism in the existing design to accommodate the rerate condition. These pressures have been evaluated to be acceptable for each compartment.

CHAPTER 03 3.6-75 REV. 18, SEPTEMBER 2016

LGS UFSAR Table 3.6-10 RWCU PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION INSIDE PRIMARY CONTAINMENT)(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ.10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 38 BW 28.735 0.0002 32.784 C TE 77 TTJ 53.039 0.2145 32.784 C&L SFL 79 TTJ 52.2 0.1883 32.784 C&L SFL 82 TTJ 51.379 0.1666 32.784 C&L SFL 90 TTJ 50.828 0.1497 32.784 C&L SFL 103 TTJ 54.446 0.3216 32.784 C&L SFL 115 SWP 45.461 0.0098 32.784 C TE UNIT 2 39 STR 11.19 0.0 32.784 C TE 103 TTJ 55.18 0.2643 32.784 C&L SFL 115 SWP 45.36 0.0597 32.784 C TE 47 TTJ 48.56 0.1348 32.784 C&L SFL 53 TTJ 48.72 0.1394 32.784 C&L SFL (1) Locations of the nodes listed in this table are shown in Figure 3.6-17.

(2) Node types are designated as follows:

TTJ - tapered transition joint BW - butt weld SWP - sweepolet STR - straight pipe (3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-76 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-11 INTENTIONALLY LEFT BLANK CHAPTER 03 3.6-77 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-12 REACTOR VESSEL DRAIN PIPING STRESS LEVELS AND PIPE BREAK DATA(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 553 RED 60.009 0.1068 33.936 C&L SFL 643 TTJ 65.95 0.5875 33.936 C&L SFL 645 TTJ 65.636 0.5678 33.936 C&L SFL 670 WLD 49.61 0.0871 33.936 C&L SFL 675 TTJ 56.538 0.1465 33.936 C&L SFL 685 TTJ 56.273 0.1358 33.936 C&L SFL 690 WLD 46.38 0.0573 33.936 C&L SFL 720 RED 80.543 0.6964 33.936 C&L SFL 725 TEE 61.08 0.2194 33.936 C SFL 790 RED 57.937 0.1171 32.784 C SFL 847 DMJ 54.883 0.1320 32.784 C SFL 860 TTJ 39.605 0.0161 42.210 C TE 866 RED 64.49 0.2543 32.784 C SFL 875 TTJ 29.37 0.00 32.784 C TE 917 TTJ 41.877 0.0237 32.784 C SFL 925 TTJ 43.832 0.0311 32.784 C SFL 935 TTJ 43.356 0.0289 32.784 C SFL UNIT 2 553 RED 55.85 0.0979 32.78 C&L SFL 643 TTJ 48.49 0.0555 32.78 C&L SFL 645 TTJ 48.29 0.0527 32.78 C&L SFL 670 WLD 45.62 0.0639 32.78 C&L SFL 675 TTJ 42.75 0.0141 32.78 C&L SFL 685 TTJ 42.83 0.0121 32.78 C&L SFL 720 RED 65.23 0.7513 32.78 C&L SFL 725 TEE 59.26 0.1430 32.78 C SFL 860 TTJ 50.96 0.0535 42.210 C TE 875 TTJ 50.74 0.1630 32.784 C TE (1) Locations of the nodes listed in this table are shown in Figure 3.6-20.

(2) Node types are designated as follows:

TEE - butt-welding tee TTJ - tapered transition joint RED - reducer DMJ - dissimilar metal joint WLD - weldolet CHAPTER 03 3.6-78 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-12 (Cont'd)

(3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5)

The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress Maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-79 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-13 HPCI STEAM SUPPLY PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION INSIDE PRIMARY CONTAINMENT)(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 197 TTJ 47.5 0.0084 42.2 C TE 200 EL 48.4 0.0018 42.2 C MBL 225 EL 58.4 0.0061 42.2 C MBL 358 STR 21.1 0.0002 42.2 C TE UNIT 2 197 TTJ 53.86 0.0289 42.2 C TE 358 STR 19.16 0.0004 42.2 C TE (1)

Locations of the nodes listed in this table are shown in Figure 3.6-21.

(2)

Node types are designated as follows:

TTJ - tapered transition joint EL - elbow STR - straight pipe (3)

Break types are indicated as follows:

C - circumferential L - longitudinal (4)

Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5)

The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximu (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 6% for the rerate condition.

(6)

The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-80 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-14 HPCI STEAM SUPPLY PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION OUTSIDE PRIMARY CONTAINMENT)(5)(6)

PIPE BREAK STRESS (ksi) STRESS LIMIT BASIS FOR NODE NODE 0.8(1.2Sh+SA) BREAK BREAK POINT(1) TYPE(2) EQ. 9 EQ. 10 TOTAL (ksi) TYPE(3) SELECTION(4)

UNIT 1 45 BW 6.87 18.19 25.06 32.4 C TE 55 TEE 7.52 15.59 23.11 32.4 C MBL 65 EL 7.50 10.83 18.33 32.4 C MBL 85 BW 7.02 5.49 12.51 32.4 C TE UNIT 2 95 ST 9.62 12.39 22.01 32.4 C TE 150 EL 10.72 11.03 21.75 32.4 C TE (1)

Locations of the nodes listed in this table are shown in Figure 3.6-22.

(2)

Node type are designated as follows:

TTJ - tapered transition joint TEE - butt-welding tee EL - elbow BW - butt weld ST - straight pipe (3)

Break types are indicated as follows:

C - circumferential L - longitudinal (4)

Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected to satisfy the requirement for a minimum number of break locations (5)

The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 6% for the rerate condition.

(6)

The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-81 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-15 RCIC STEAM SUPPLY PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION INSIDE PRIMARY CONTAINMENT)(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ.10 Usage 2.4 S m Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 250 EL 35.62 0.0003 42.2 C TE 320 EL 41.7 0.0008 42.2 C MBL 350 EL 47.4 0.0013 42.2 C MBL 391 TTJ 68.5 0.29 42.2 C TE UNIT 2 250 EL 12.05 0.0000 42.2 C TE 391 TTJ 62.60 0.1201 42.2 C TE (1)

Location of the nodes listed in this table are shown in Figure 3.6-25.

(2)

Node types are designated as follows:

TTJ - tapered transition joint EL - elbow RED - reducer (3)

Break types are indicated as follows:

C - circumferential L - longitudinal (4)

Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5)

The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 6% for the rerate condition.

(6)

The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-82 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-16 RCIC STEAM SUPPLY PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION OUTSIDE PRIMARY CONTAINMENT)(5),(6)

PIPE BREAK STRESS (ksi) STRESS LIMIT BASIS FOR NODE NODE 0.8(1.2Sh+SA) BREAK BREAK POINT(1) TYPE(2) EQ. 9 EQ. 10 TOTAL (ksi) TYPE(3) SELECTION(4)

UNIT 1 40 BW 7.62 1.60 9.22 32.4 C TE 100 EL 8.14 18.63 26.77 32.4 C MBL 105 EL 7.66 16.84 24.50 32.4 C MBL 117 STR 7.40 1.69 9.08 32.4 C TE UNIT 2 30 TTJ 6706 927 7633 32400 C TE 105 EL 6712 2618 9330 32400 C TE (1)

Locations of the nodes listed in this table are shown in Figure 3.6-26.

(2)

Node types are designated as follows:

TTJ - tapered transition joint EL - elbow BW - butt weld STR - straight pipe (3)

Break types are indicated as follows:

C - circumferential L - longitudinal (4)

Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected to satisfy the requirement for a minimum number of break locations (5)

The values provided above are based upon original licensed power conditions. These cumulative usage factor (CUF) values are conservatively derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased up to 6% for the rerate conditions.

(6)

The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-83 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-17 MAIN STEAM DRAINAGE PIPING STRESS LEVELS AND PIPE BREAK DATA (PORTION INSIDE PRIMARY CONTAINMENT)(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 10 SW 37.8 0.0196 42.2 C TE 30 SW 39.12 0.0168 42.2 C MBL 51 TEE 41.0 0.0075 42.2 C MBL 65 SW 28.4 0.0143 42.2 C MBL 90 SW 28.5 0.0124 42.2 C TE 110 TTJ 13.24 0.0 42.2 C TE 195 TEE 40.8 0.0014 42.2 C MBL 220 SW 28.2 0.0141 42.2 C MBL 240 SW 26.3 0.0113 42.2 C TE 276 SW 39.5 0.0171 42.2 C MBL 295 SW 37.64 0.0196 42.2 C TE UNIT 2 10 SW 53.58 0.0402 42.2 C TE 90 SW 38.49 0.0174 42.2 C TE 110 TTJ 18.32 0.0006 42.2 C TE 240 SW 38.93 0.0320 42.2 C TE 295 SW 57.16 0.0582 42.2 C TE (1) Locations of the nodes listed in this table are shown in Figure 3.6-28.

(2) Node types are designated as follows:

TEE - butt-welding tee TTJ - tapered transition joint SW - socket weld (3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-84 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-18 INTENTIONALLY LEFT BLANK CHAPTER 03 3.6-85 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-19 RPV HEAD VENT PIPING STRESS LEVELS AND PIPE BREAK DATA(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi ) Type(3) Selection(4)

UNIT 1 5 TTJ 47.61 0.0541 42.21 C TE 10 CUR 46.95 0.0012 42.21 C MBL 28 RED 54.22 0.3862 42.21 C SFL 30 DMJ 59.49 0.1037 42.21 C SFL 33 TTJ 28.92 0.0 42.21 C TE 500 TEE 44.85 0.0157 42.21 C MBL 620 SW 25.50 0.0672 42.21 C TE 835 SW 54.69 0.0621 42.21 C MBL 920 SW 51.97 0.0232 42.21 C TE 292 SW 74.85 0.5044 42.21 C SFL UNIT 2 5 TTJ 32.80 0.0032 42.21 C TE 292 SW 74.85 0.5044 42.21 C SFL 40 TTJ 21.09 0.00 33.94 C TE 620 SW 19.03 0.0017 42.21 C TE 910 SW 68.56 0.2355 42.21 C TE (1) Locations of the nodes are shown in Figure 3.6-30.

(2) Node types are designated as follows:

TTJ - Tapered transition joint CUR - Butt-weld elbow RED - Reducer DMJ - Dissimilar metal joint TEE - Butt-welding tee SW - Socket weld (3) Break types are indicated as follows:

C - Circumferential (4) Symbols used to denote the basis for break selection are as follows:

TE - Terminal end MBL - Intermediate break location selected to satisfy the requirement for a minimum number of break locations.

SFL - The stress and fatigue limits established in Section 3.6.2.1.1.3 are not met.

(5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-86 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-20 SLCS INJECTION PIPING STRESS LEVELS AND PIPE BREAK DATA(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 5 SW 57.25 0.1478 33.8 C TE 15 SW 58.67 0.1784 33.8 C SFL 20 SW 59.43 0.1887 33.8 C SFL 35 SW 60.53 0.2049 33.8 C SFL 40 SW 61.15 0.2208 33.8 C SFL 42 SW 61.36 0.2285 33.8 C SFL 43 SW 60.96 0.2204 33.8 C SFL 70 SW 64.29 0.3164 33.8 C SFL 75 SW 64.10 0.3139 33.8 C SFL 76 SW 63.46 0.2940 33.8 C TE UNIT 2 7 SW 40.31 0.0455 32.44 C TE 76 SW 46.69 0.0559 32.44 C TE (1) Locations of the nodes are shown in Figure 3.6-31.

(2) Node type are designated as follows:

SW - Socket weld (3) Break types are indicated as follows:

C - Circumferential (4) Symbols used to denote the basis for break selection are as follows:

TE - Terminal end break SFL - The stress and fatigue limits established in Section 3.6.2.1.1.3 are not met.

(5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-87 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-21 RHR SHUTDOWN COOLING PIPING STRESS LEVELS AND PIPE BREAK DATA(7)(8)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 107 BW 37.999 0.0193 32.79 C TE 109 EL 60.727 0.3363 32.79 C&L SFL 125 TTJ 60.924 0.976 32.79 C&L SFL 142 TTJ 58.386 0.831 32.79 C&L SFL 150 TTJ 60.330 0.8259 32.79 C TE 288 TTJ 38.46 0.0242 32.79 C TE 291 TTJ 37.659 0.0238 32.79 C MBL 293 TTJ 40.264 0.0258 32.79 C MBL 297 BW 32.920 0.0463 32.79 C TE 325 TTJ 39.737 0.0269 32.79 C TE 335 TTJ 41.565 0.0314 32.79 C MBL 345 TTJ 45.696 0.0676 32.79 C MBL 350 BW 34.127 0.0515 32.79 C TE UNIT 2 107 TTJ 33.59 0.0141 32.79 C TE 109(5) EL 51.20 0.0415 32.79 C&L SFL 125 TTJ 54.48 0.3439 32.79 C&L SFL 142 TTJ 53.48 0.3065 32.79 C&L SFL 150 TTJ 56.73 0.4447 32.79 C TE 288 TTJ 35.98 0.0231 32.79 C TE 297 TTJ 37.48 0.0396 32.79 C TE 325 TTJ 38.16 0.0219 32.79 C TE 350 BW 32.51 0.0319 32.79 C TE 293(6) TTJ 46.84 0.0527 32.79 C&L SFL (1) Locations of the nodes listed in this table are shown in Figure 3.6-32.

(2) Node types are designated as follows:

TTJ - tapered transition joint EL - elbow BW - butt weld CHAPTER 03 3.6-88 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-21 (Cont'd)

(3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) EQ.13 = 38.37 ksi (6) EQ.13 = 38.13 ksi (7) The values provided above are based upon original licensed power conditions. These CUF values are conservatively based derived by using Sm for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(8) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-89 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-22 LPCI INJECTION PIPING STRESS LEVELS AND PIPE BREAK DATA(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 65 TTJ 30.87 0.0008 32.78 C T 67 DMJ 56.97 0.12 32.78 C&L SFL 90 EL 117.84 0.81 32.78 C&L SFL 100 TTK 157.16 0.64 32.78 C&L SFL 140 EL 156.47 0.80 32.78 C&L SFL 145 STR 109.32 0.44 32.78 C&L SFL 150 TTJ 154.95 0.62 32.78 C TE UNIT 2 65 TTJ 37.48 0.0012 32.78 C TE 90 EL 117.84 0.4709 32.78 C&L SFL 100 TTJ 143.83 0.6160 32.78 C&L SFL 140 EL 114.41 0.4580 32.78 C&L SFL 150 TTJ 140.38 0.6493 32.78 C TE (1) Locations of the nodes listed in this table are shown in Figure 3.6-33.

(2) Node types are designated as follows:

TTJ - tapered transition joint EL - elbow DMJ - dissimilar metal joint STR - straight pipe (3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-90 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-23 CORE SPRAY INJECTION PIPING STRESS LEVELS AND PIPE BREAK DATA(5)(6)

Pipe Break Stress Cumula- Stress By tive Limit Basis for Node Node EQ. 10 Usage 2.4 Sm Break Break Point(1) Type(2) (ksi) Factor (ksi) Type(3) Selection(4)

UNIT 1 25 TTJ 30.19 0.0016 32.78 C TE 30 DMJ 66.87 0.1365 32.78 C&L SFL 63 STR 106.87 0.3789 32.78 C&L SFL 65 EL 124.58 0.7034 32.78 C&L SFL 72 RED 108.52 0.7573 32.78 C&L SFL 73 STR 109.03 0.4141 32.78 C&L SFL 75 TTJ 117.54 0.4955 32.78 C TE UNIT 2 25 TTJ 30.49 0.0037 32.78 C TE 75 TTJ 56.29 0.0733 32.78 C TE (1) Locations of the nodes listed in this table are shown in Figure 3.6-34.

(2) Node types are designated as follows:

TTJ - tapered transition joint EL - elbow RED - reducer DMJ - dissimilar metal joint STR - straight pipe (3) Break types are indicated as follows:

C - circumferential L - longitudinal (4) Symbols used to denote the basis for break selection are as follows:

TE - terminal end SFL - the stress and fatigue limits established in Section 3.6.2.1.1.3 are not met MBL - intermediate break location selected in order to satisfy the requirement for a minimum number of break locations (5) The values provided above are based upon original licensed power conditions. These Cumulative Usage Factor (CUF) values are conservatively based derived by using Stress maximum (Sm) for design temperatures instead of the load pair operating temperatures. For rerate power conditions, using the more realistic approach, the CUF values are found to be equal to or less than those in the table. No new pipe break is required for rerate power conditions. The maximum equation 10 stresses may be increased by up to 2% for the rerate condition.

(6) The information posted in this table was used for the original System piping analysis. Refer to System ASME III piping analysis for current information.

CHAPTER 03 3.6-91 REV. 15, SEPTEMBER 2010

LGS UFSAR Table 3.6-24 RESTRAINT DATA USED IN VERIFICATION OF RECIRCULATION SYSTEM PIPE WHIP RESTRAINT DESIGN(1)

PIPE INITIAL EFFECTIVE TOTAL SIZE REST LOAD LIMIT, CLEARANCE CLEARANCE CLEARANCE (in) DIRECTION C2(2) N(2) RESTRAINT(2) (in) (in) (in) 12 0 27,733 0.24 6.129 4 1.941 5.941 12 90 14,795 0.401 9.063 4 12.247 16.247 16 0 109,265 0.24 6.278 4 1.934 5.934 6 90 62,599 0.377 8.978 4 12.187 16.187 24 0 102,228 0.24 8.222 4 1.984 5.984 24 90 55,531 0.375 11.972 4 13.685 17.685 24 38(3) 109,888 0.24 5.588 4 5.698 9.698 24 52(3) 109,835 0.24 5.473 4 8.462 12.462 (1)

The restraint data listed applies to one bar of a restraint.

(2)

F = C2 ( restraint)N where F is the resistance force for one bar of a restraint and where ( restraint) = ( pipe) - (total clearance)

(3)

Applies to restraint RCR 3 only.

CHAPTER 03 3.6-92 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.6-25 COMPARISON OF PDA AND NSC CODES FRACTION OF DESIGN RESTRAINT RESTRAINT PIPE BREAK RESTRAINT NO. OF BARS LOAD (kips) DEFLECTION(in) DEFLECTION (%) DEFLECTION (in)

DESIGNATION(1) DESIGNATION(1) PDA NSC PDA NSC PDA NSC PDA NSC PDA NSC RC1J RCR1 5 5 803.2 788.3 6.57 7.926 79.93 96.4 17.72 15.58 RC2LL RCR1 5 5 766.4 458.4 14.99 7.495 125 62.6 35.83 24.52 RC3LL RCR2 6 6 747.0 639.7 2.27 3.73 27.65 45.35 17.16 20.11 RC3LL RCR2 6 6 796.6 780.3 10.22 10.54 57.8 59.6 41.48 43.0 RC4LL RCR3 5 5 846.0 838.4 7.64 8.05 92.95 97.98 18.87 16.43 RC4LL RCR3 8 8 1319.0 1073.9 5.43 4.62 99.23 76.85 23.28 17.25 RC4CV RCR3 8 8 1260.7 1275.0 4.49 5.58 80.37 99.89 22.56 18.73 RC6AV RCR3 8 8 928.5 722.5 1.22 1.77 22.46 31.7 23.68 95.39 RC7J RCR7 6 6 953.3 801.6 6.28 5.76 76.4 70.12 16.46 21.63 RC8LL RCR6 4 4 599.0 0 8.28 0 112.46 0 26.76 6 6 895.0 0 8.16 0 110.76 0 29.316 8.39 RC9CV RCR6 4 4 575.8 520.16 4.16 5.53 50.63 67.33 13.2 14.56 RC9LL RCR8 6 6 830.2 546.8 11.408 6.815 95.29 56.9 36.612 26.24 RC11A RCR8 6 6 818.3 493.6 10.98 5.99 91.72 50.07 31.404 23.71 RC13 RCR10 4 4 668.4 478.4 5.87 3.66 93.5 58.39 13.37 10.44 RC16 RCR11 4 4 687.4 518.4 6.59 4.38 105 69.86 15.37 10.22 RC14CV RCR20 8 8 285.0 309.6 2.83 5.88 46.3 95.92 15.45 13.96 RC14LL RCR20 8 8 116.3 129.9 0.96 3.36 10.5 37.1 22.13 23.56 (1)

Break designations and restraint designations are shown on Figure 3.6-37.

CHAPTER 03 3.6-93 REV. 13, SEPTEMBER 2006

LGS UFSAR 3.7 SEISMIC DESIGN All systems and equipment of the NSSS are defined as either seismic Category I or nonseismic Category I. The requirements for seismic Category I qualification are given in Section 3.2, along with a list of systems, components, and equipment which are so categorized. Seismic design requirements for nonseismic Category I items are also defined in Section 3.2. These items are classified as either Category II or Category IIA.

All systems, components, and equipment related to plant safety are designed to withstand SSE and OBE.

The SSE is that earthquake which is based upon an evaluation of the maximum earthquake potential considering the regional and local geology, and seismology and specific characteristics of local subsurface material. It is that earthquake which produces the maximum vibratory ground motion for which seismic Category I systems and components are designed to remain functional.

These systems and components are those necessary to ensure:

a. The integrity of the RCPB.

b. The capability to shut down the reactor and maintain it in a safe shutdown condition.

c. The capability to prevent or mitigate the radiological consequences of accidents which could result in potential offsite exposures, comparable to the dose limits of 10 CFR 50.67.

The OBE is that earthquake which, considering the regional and local geology, and seismology and specific characteristics of local subsurface material, could reasonably be expected to affect the plant site during the operating life of the plant. It is that earthquake which produces the vibratory ground motion for which those features of the nuclear power plant necessary for continued operation, without undue risk to the health and safety of the public, are designed to remain functional.

In addition to seismic loads, Category I structures, systems and components are also reviewed for hydrodynamic loads as described in Appendix 3A.

3.7.1 SEISMIC INPUT 3.7.1.1 Design Response Spectra The site design response spectra used are shown in Figures 3.7-1 and 3.7-2. These spectra are for the horizontal components of the OBE and the SSE respectively. The response spectra for the OBE are scaled or normalized to a maximum horizontal ground acceleration of 71/2% of gravity.

The response spectra for the SSE are normalized to a maximum horizontal ground acceleration of 15% of gravity. The values for the vertical component of the design response spectra are 2/3 of the horizontal design response spectra described above. The response spectra are based on data developed from records of previous earthquake activity and represent an envelope of motion expected at a sound rock site from a nearby earthquake (Section 2.5.2).

Regulatory Guide 1.60 (December, 1973) "Design Response Spectra for Seismic Design of Nuclear Power Plants" was not used for the development of the spectra, because LGS Units 1 and 2 were docketed for construction permit review in March 1970, and the spectra were finalized in CHAPTER 03 3.7-1 REV. 19, SEPTEMBER 2018

LGS UFSAR 1973. Further, a letter dated December 21, 1973 from J.M. Hendrie (NRC) to R.M. Collins (Bechtel) states that Regulatory Guide 1.60 is applicable only to the plants docketed for construction permit review after April 1, 1973.

3.7.1.2 Design Time History A synthetic time history of motion was generated by modifying the actual records of the 1952 Taft earthquake according to the techniques proposed in Reference 3.7-1. This synthetic time history was then further modified to develop the time history shown in Figure 3.7-3, which corresponds to the design response spectra. The duration of the time history is 15 seconds. Figures 3.7-4 through 3.7-9 show a comparison of the time history response spectra and the design response spectra for 0.5%, 1%, 2%, 3%, 5%, and 7% damping values.

The spectra are computed at the frequency values as given in table 5-1 of Reference 3.7-2.

3.7.1.3 Critical Damping Values 3.7.1.3.1 Critical Damping Values (NSSS)

The damping values indicated in Table 3.7-1 are used in the response analysis of various structures and systems, and in preparation of floor response spectra used as forcing inputs for piping and equipment analysis or testing. It can be seen that the values given in Table 3.7-1 are somewhat less than those given in Regulatory Guide 1.61 (October 1973), "Damping Values for Seismic Design of Nuclear Power Plants". The calculated responses are, therefore, conservative.

Alternative critical damping values for piping may be used as described in Section 3.7.1.3.3.

3.7.1.3.2 Critical Damping Values (Non-NSSS)

Critical damping values expressed as a percentage of critical damping and used for seismic Category I structures, equipment, and piping for both the OBE and SSE are given in Table 3.7-2.

Alternative critical damping values for piping may be used as described in Section 3.7.1.3.3.

Regulatory Guide 1.61 is not used as a design basis as discussed in Section 1.8, except as discussed in Section 3.7.1.3.3. However, all the values shown in Table 3.7-2 are equivalent to or more conservative than those in Regulatory Guide 1.61 with the exception of the SSE value for welded steel structures. The damping value of 5% (PSAR table C.2.1) is based on information given in Reference 3.7-6. The 5% value has been used, with appropriate design margins, because the stress levels for SSE conditions are allowed to approach the yield point.

3.7.1.3.3 Alternative Critical Damping Values and Spectral Peak Broadening for Piping (NSSS and Non-NSSS)

Alternative critical damping values, as provided in ASME B&PV Code,Section III, Division 1 Code Case N-411 may be used. When used, the provisions listed below are applied unless otherwise approved by References 3.7-12 and 3.7-13. References 3.7-12 and 3.7-13 approved the use of Independent Support Motion (ISM) response spectra methodology in conjunction with the damping values specified in ASME Code Case N-411, as the basis for eliminating snubbers from 31 piping systems on Unit 1 and Unit 2.

CHAPTER 03 3.7-2 REV. 19, SEPTEMBER 2018

LGS UFSAR

• The code case damping is applied only to uniform (or envelope) response spectra loading analysis for seismic and seismiclike building filtered hydrodynamic loads and the annulus pressurization loading.

• 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.

• Modal and direction combination of the three earthquake directions are combined in accordance with Regulatory Guide 1.92.

• 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.

• Assurance that the predicted piping displacements are such that adequate clearance exists with respect to adjacent components and equipment.

• Line mounted equipment is designed to withstand the increased pipe motion.

• The code case damping is not used for analyzing the dynamic response of piping systems incorporating supports designed to dissipate energy by yielding.

• The code case damping is not applied to piping analytical models that incorporate equipment with natural frequencies below 20 hertz.

• The code case damping is not applied to piping in which stress-corrosion cracking has occurred unless a case specific evaluation is reviewed by the NRC.

When time history or independent support motion response spectrum analysis is utilized, the following critical damping values are applied:

a. Time history analysis with 0.5% damping for OBE, 1% damping for SSE, or Regulatory Guide 1.61 damping for hydrodynamic loads.

b. Independent support motion response spectra analysis with Regulatory Guide 1.61 damping.

ASME Section III, Division I Code Case N-397, "Alternative Rules to the Spectral Broadening Procedures of N-1226.3 for Classes 1, 2 and 3 Piping,Section III, Division 1" may be selectively used on a load case basis. This code case is applicable to all spectral analysis load cases and all methodologies except it was not applied to a spectral analysis load case that utilized the independent support motion analysis methodology.

3.7.1.4 Supporting Media for Seismic Category I Structures All seismic Category I structures are supported on sound rock or concrete backfill bearing on sound rock, except for some yard facilities such as valve pits and portions of electrical duct banks and underground piping which are supported on natural soil or fill (Section 2.5.4.10).

For the dynamic analysis of the rock-founded structures, soil-structure interaction is considered to be negligible due to the high stiffness of the rock. However, the floor response spectra developed CHAPTER 03 3.7-3 REV. 19, SEPTEMBER 2018

LGS UFSAR for the reactor enclosure and the containment for equipment analysis are based on a model that considered the flexibility of the supporting medium. The modulus of elasticity, the shear-wave velocity, and the density of the supporting medium used in the analysis are 3.0x106 psi, 6000 fps, and 150 lbs/ft3, respectively. Additional soil-structure interaction studies were performed for the containment structure and reactor and control enclosures to assess the sensitivity of the structural response to variations in the design basis rock modulus. Modal analyses have demonstrated that for a +/-50% of the design elastic modulus range, variations in structural frequency do not exceed 10% for predominant modes. These results indicate that a reduction in rock modulus to 1.5x106 psi would not produce significant effects on the structural response. It is therefore concluded that the average dynamic elastic modulus value of 3.0x106 psi is adequate for design.

3.7.2 SEISMIC SYSTEM ANALYSIS Seismic Category I structures and systems, and components of the NSSS that fall under the category of a seismic system, are discussed here. Seismic systems are analyzed for both OBE and SSE.

3.7.2.1 Seismic Analysis Method 3.7.2.1.1 Seismic Analysis Methods (NSSS)

Analysis of seismic Category I NSSS systems and components was accomplished, where applicable, using the response spectrum or time history approach. Either approach utilizes the natural period, mode shapes, and appropriate damping factors of the particular system. Certain pieces of equipment having very high natural frequencies may be analyzed statically. In some cases, dynamic testing of equipment may be used for seismic qualification.

A time history analysis involves the solution of the equations of the dynamic equilibrium (Section 3.7.2.1.1.1) by means of the methods discussed in Section 3.7.2.1.1.2. In this case, the duration of motion is of sufficient length to ensure that the maximum values of response are obtained.

A response spectrum analysis involves the solution of the equations of motion (Section 3.7.2.1.1.1) by the method discussed in Section 3.7.2.1.1.3.

3.7.2.1.1.1 The Equations of Dynamic Equilibrium Assuming that the force due to damping is proportional to velocity, the dynamic equilibrium equations for a lumped-mass, distributed stiffness system are expressed in matrix form as:

[M]{u(t)} + [C]{u(t)} + [K]{u(t)} = {P(t)} (EQ. 3.7-1) where:

u(t) = time-dependent displacement of nonsupport points relative to the supports u(t) = time-dependent velocity of nonsupport points relative to the supports u(t) = time-dependent acceleration of nonsupport points relative to the supports CHAPTER 03 3.7-4 REV. 19, SEPTEMBER 2018

LGS UFSAR

[M] = diagonal matrix of lumped masses

[C] = damping matrix

[K] = stiffness matrix P(t) = time-dependent inertial forces acting at nonsupport points 3.7.2.1.1.2 Solution of the Equations of Motion by Mode - Superposition The first technique used for the solution of the equations of motion is the method of mode-superposition.

The set of homogeneous equations represented by the undamped free vibration of the system is:

[M]{u(t)} + [K]{u(t)} = {0} (EQ. 3.7-2)

Since the free oscillations are assumed to be harmonic the displacements can be written as

{u(t)} = {} e (EQ. 3.7-3) where:

{} = column matrix of the amplitude of displacement {u}

= circular frequency of oscillation t = time Substituting Equation 3.7-3 and its derivatives into Equation 3.7-2 and noting that (e) is not necessarily zero for all values of (t) yields

[-2[M] + [K)){} = {0} (EQ. 3.7-4)

Equation 3.7-4 is the classical algebraic eigenvalue problem, wherein the eigenvalues are the frequencies of vibrations, wi, and the eigenvectors are the mode shapes, {}i.

For each frequency here is a corresponding solution vector {}i. It can be shown that the mode shape vectors are orthogonal with respect to the weighting matrix [K] in the n-dimensional vector space.

The mode shape vectors are also orthogonal with respect to the mass matrix [M].

The orthogonality of the mode shapes is used to effect a coordinate transformation of the displacements, velocities, and accelerations, so that the response in each mode is independent of the response of the system in any other mode. Thus, the problem becomes one of solving (n) independent differential equations rather than (n) simultaneous differential equations; and, since the system is linear, the principle of superposition holds, and the total response of the system oscillating simultaneously in (n) modes is determined by direct addition of the responses in the individual modes.

CHAPTER 03 3.7-5 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.7.2.1.1.3 Analysis by Response Spectrum As an alternative to the step-by-step mode-superposition method described in Section 3.7.2.1.1.2, the response spectrum method may be used. The response spectrum method is based on the fact that the modal responses can be expressed as a set of integral equations, rather than a set of differential equations. The advantage of this form of solution is that for a given ground motion, the only variables under the integral are the damping factor and the frequency. Thus, for a specified damping factor, it is possible to construct a curve which gives a maximum value of the integral as a function of frequency. This curve is called a response spectrum for the particular input motion and the specified damping factor. The integral has units of velocity, consequently the maximum of the integral is called the spectral velocity.

Using the calculated natural frequencies of vibration of the system, the maximum values of the modal responses are determined directly from the appropriate response spectrum. The modal maxima are then combined as discussed in Section 3.7.3.7.1.

The total seismic structural response is predicted by combining the response calculated from the two horizontal and the vertical analyses. When the response spectrum method is used, the methods for combining the loads from the three analyses is based on the method described in Section 3.7.2.6.

3.7.2.1.1.4 Support Displacements in Multisupported Structure The methods described in Section 3.7.2.1.1.5 are used, where applicable, to account for the effects of relative anchor motion in the case of a multisupported structure.

3.7.2.1.1.5 Dynamic Analysis of Seismic Category I Systems and Components The time history technique and the response spectrum technique were used as applicable for the dynamic analysis of seismic Category I NSSS systems and components which are sensitive to dynamic seismic events.

a. Dynamic Analysis of Piping Systems Each pipeline is idealized as a mathematical model consisting of lumped masses connected by elastic members. The stiffness matrix for 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 change 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 the response spectrum method of analysis. When the piping system is being anchored and supported at points with different excitations, the response spectrum analysis is performed using a response spectrum which envelopes all the response spectra of anchors and supports. Alternatively, except as restricted by Section 3.7.1.3.3, the multiple response spectra/independent support motion method of analysis may be used where distinct time histories or response spectra are applied to all piping system attachment points.

The relative modal displacement between anchors is determined from the dynamic analysis of the primary structures. The results of the relative anchor point CHAPTER 03 3.7-6 REV. 19, SEPTEMBER 2018

LGS UFSAR displacement per mode are used for a static analysis to determine the secondary stresses due to relative anchor point displacements. The modal stresses are combined by the SRSS method to obtain the resultant secondary stresses.

b. Dynamic Analysis of Equipment Equipment is idealized as a mathematical model consisting of lumped masses connected by elastic members or springs. The Table 3.9-6 series documents calculated and allowable loads for various loading combinations, including seismic loads, applied to major NSSS systems and components.

The response spectrum analysis is performed using the enveloped response spectrum of all attachment points. Alternatively, for equipment supported at two or more points located at different elevations in the same primary structure, the response spectrum for the most severe single support can be applied uniformly to all support points. As a second alternative, if appropriate, the response spectra at the elevation near the center of gravity of the equipment may be taken as the design spectra.

The relative modal displacement between supports is determined from the dynamic analysis of the structure. The relative support point displacements per mode are used for a static analysis to determine the secondary stresses due to support displacements. The modal stresses are combined by the SRSS method to obtain the resultant secondary stresses. Further details are given below.

c. Differential Seismic Movement of Interconnected Components The procedure for considering differential displacements for equipment anchored and supported at points with different displacement excitations is as follows:

The relative displacements between the supporting points induces additional stresses in the equipment supported at these points. These stresses can be evaluated by performing a static analysis where each of the supporting points is displaced a prescribed amount. From the dynamic analysis of the complete structure, the time history of displacement at each supporting point is available.

These displacements are used to calculate stresses by determining the peak modal responses. The stresses thus obtained for each natural mode are then superimposed for all modal displacements of the structure by the SRSS method.

In the static calculation of the stresses due to relative displacements in the response spectrum method, the maximum value of the modal displacement is used.

Therefore, the mathematical model of the equipment is subjected to the maximum displacement at its supporting points obtained from the modal displacements. This procedure is repeated for the significant modes (modes contributing most to the total displacement response at the supporting point) of the structure. The total stresses due to relative displacement are obtained by combining the modal results using the SRSS method. Since the maximum displacements for different modes do not occur at the same time, the SRSS method is a realistic and practical method.

When a component is covered by the ASME B&PV Code, then the stresses due to relative displacement as obtained above are treated as secondary stresses.

CHAPTER 03 3.7-7 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.7.2.1.1.6 Seismic Qualification by Testing For certain seismic Category I equipment and components where dynamic testing is necessary to ensure functional integrity, test performance data and results reflect the following:

a. Performance data of equipment which, under the specified conditions, has been subjected to dynamic loads equal to or greater than those to be experienced under the specified seismic conditions

b. Test data from previously tested comparable equipment which, under similar conditions, has been subjected to dynamic loads equal to or greater than those specified

c. Actual testing of equipment in accordance with one of the methods described in Sections 3.9 and 3.10 Alternate test procedures that satisfy the requirements of these criteria are allowed, subject to review by the responsible engineer.

3.7.2.1.2 Seismic Analysis Methods (Non-NSSS)

In the analysis of seismic Category I structures, two distinct objectives must be satisfied:

a. Development of in-structure seismic response characteristics, where necessary, for use in the analysis and design of seismic Category I systems, equipment, and components

b. Determination of seismic force distribution within the various structures resulting from the design criteria free-field seismic input, for use in the design of seismic Category I structures Two analytical procedures were employed to determine the seismic responses to the Category I structures. In general, a modal response spectrum analysis was used to compute the in-structure seismic responses, including nodal accelerations, nodal displacements, and member forces.

Alternatively, a time history analysis procedure was used to generate the in-structure seismic responses discussed above. In addition, time history analysis was used to generate all floor response spectra. The mathematical idealization of the structural characteristics of the various seismic Category I structures was accomplished by a lumped-parameter beam-stick model. The general analytical methods and modeling techniques used in these analyses are in accordance with Reference 3.7-2. The seismic design criteria input is defined in terms of the OBE and SSE design response spectra (Section 3.7.1.1), the synthetic time history (Section 3.7.1.2), and the soil-structure interaction parameters (Section 3.7.2.4) used for development of floor response spectra for equipment assessment. Refer to Figures 3.7-10 through 3.7-19 for either a pictorial representation or an actual sketch of the mathematical models used. A complete description of the formulation of the mathematical models and their use is provided in Section 3.7.2.3.2.

3.7.2.2 Natural Frequencies and Response Loads The natural frequencies of the primary containment, the reactor enclosure, and the control structure below 33 cps are shown in Tables 3.7-5 and 3.7-6 respectively. The significant mode CHAPTER 03 3.7-8 REV. 19, SEPTEMBER 2018

LGS UFSAR shapes of the containment and the reactor enclosure and control structure are shown on Figures 3.7-20 through 3.7-30. The mode shapes for the primary containment are for the horizontal and vertical directions. The reactor enclosure and control structure mode shapes are for each of the three principal directions: east-west, north-south, and vertical.

Tables 3.7-7 through 3.7-16 show the response (i.e., displacements, accelerations, shear forces, bending moments, and axial forces) of the primary containment and the reactor enclosure and control structure for both OBE and SSE.

Response spectra at critical locations are shown on Figures 3.7-31 through 3.7-40. The curves are shown for each of the principal directions at the damping values shown (Section 3.7.2.15 for further discussion of damping values). Figures 3.7-41 and 3.7-42 represent the response spectra of the refueling area using soil-structure interaction.

3.7.2.3 Procedures Used for Modeling 3.7.2.3.1 Procedures Used for Modeling (NSSS) 3.7.2.3.1.1 Modeling Techniques for Seismic Category I Systems and Components An important step in the seismic analysis of seismic Category I NSSS systems and components is the procedure used for modeling. The techniques currently being used are represented by lumped masses and a set of spring-dashpots idealizing both the inertial and stiffness properties of the system. The details of the mathematical models are determined by the complexity of the actual structures and the information required for the analysis.

3.7.2.3.1.2 Modeling of Reactor Pressure Vessel and Internals The seismic loads on the RPV and internals are based on a dynamic analysis of an entire RPV enclosure complex, with the appropriate forcing function applied at ground level. For this analysis, the models shown in Figure 3.7-19 and the mathematical model of the enclosure are coupled together.

This mathematical model consists of lumped masses connected by elastic (linear) members. The stiffness properties of the model are determined using the elastic properties of the structural components. This includes the effects of both bending and shear. In order to facilitate hydrodynamic mass calculations, several mass points (fuel, shroud, vessel) are selected at the same elevation. The various lengths of CRD housings are grouped into the two representative lengths as shown in Figure 3.7-19. 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 loads.

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 the stiffnesses of light components such as jet pumps, incore guide tubes and housings, spargers, and their supply headers. This is done to reduce the complexity of the dynamic model. To find seismic responses of these components, the floor response spectra generated from the system analysis are used.

The presence of 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 CHAPTER 03 3.7-9 REV. 19, SEPTEMBER 2018

LGS UFSAR introduction of a hydrodynamic mass matrix, which serves 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 3.7-5. The seismic model of the RPV and internals has two generalized coordinates in the horizontal directions for each mass point considered in the analysis. The remaining generalized vertical coordinate is excluded because the vertical mode frequencies of RPV and internals are well above the significant horizontal mode frequencies. A separate vertical analysis is performed. The two rotational coordinates about each node point are excluded because the moment contribution from rotary inertia 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 shroud support plate is loaded in its own plane during a seismic event and hence is extremely stiff, and, therefore, may 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.

3.7.2.3.2 Procedures Used for Modeling (Non-NSSS)

The mathematical models used for the analysis of all major Category I structures are fixed base, lumped-mass, elastic spring models, except for the primary containment, the reactor enclosure and control structure, where the lumped-mass seismic model considers a rigid mass resting on soil springs. The soil spring constants are determined from formulae given in table 3-2 of Reference 3.7-2. Two separate and independent analyses are done for the containment structure. The section properties for these two analyses are based on uncracked and cracked concrete sections, respectively. The same models are used both for the response spectrum and time history analyses. The mathematical models of the primary containment are shown in Figures 3.7-10 and 3.7-11, and those of the reactor enclosure and control structure are shown in Figures 3.7-14 and 3.7-15. The additional mathematical models used for development of floor response spectra, using soil-structure interaction for these structures, are shown in Figures 3.7-12, 3.7-13, 3.7-16, and 3.7-17.

For all models, the masses are located at elevations of mass concentrations, such as floors and roofs. However, in the case of the containment, a structure of continuous mass distribution, masses are lumped at approximate maximum intervals of 15 feet along the containment shell and reactor pedestal. These methods of mass distribution are in accordance with the procedures of section 3.2 of Reference 3.7-2. All equipment, components, and piping systems are lumped into the supporting structure except for the reactor vessel, which is incorporated into the lumped-parameter analysis of the containment structure to account for potential coupling. The detailed analysis of the NSSS is performed using a decoupled model as discussed in Section 3.7.2.3.1.

A more refined RPV model is being used in the DAR (Appendix 3A) to evaluate the anticipated effects of hydrodynamic loads in combination with seismic loads.

3.7.2.4 Soil-Structure Interaction Since the seismic Category I structures are founded on competent bedrock, a soil spring approach to characterize soil-structure interaction is not used in the dynamic analysis. A simplified lumped-mass method using a fixed base model is used. However, for a more refined analysis of containment and reactor enclosure, the underlying foundation medium is considered to interact with the structure. The equivalent soil spring constant and damping coefficient are computed in accordance with the formulae of table 3-2 of Reference 3.7-2, and the analysis carried out by the CHAPTER 03 3.7-10 REV. 19, SEPTEMBER 2018

LGS UFSAR methods discussed in appendix D of Reference 3.7-2. The resulting structure-foundation interaction coefficients are listed in Table 3.7-17.

3.7.2.5 Development of Floor Response Spectra 3.7.2.5.1 Floor Response Spectra (NSSS)

See Section 3.7.3.6.1.

3.7.2.5.2 Floor Response Spectra (Non-NSSS)

The time history method of analysis was used to develop the floor response spectra. A discussion of the technique of finding the nodal time history and then producing the spectrum may be found in Sections 4.2 and 5.2 of Ref 3.7-2.

3.7.2.6 Three Components of Earthquake Motion 3.7.2.6.1 NSSS See Section 3.7.3.6.1.

3.7.2.6.2 Non-NSSS The response spectrum method was used in seismic analysis of structures. Independent analyses are performed for the vertical and two horizontal (east-west and north-south) directions. Regulatory Guide 1.92, "Combining Responses and Spatial Components in Seismic Response Analysis" states that the design response value is obtained by taking the SRSS of the maximum co-directional responses caused by each of the three components of earthquake motion at a particular point of the structure. However, for design purposes, the response value used is the maximum value obtained by adding the response due to the vertical earthquake with the larger value of the response due to one of the horizontal earthquakes by the absolute sum method. A discussion of the adequacy of the conservatism provided by this two-component absolute summation technique as compared to the Regulatory Guide 1.92 requirement of a three-component SRSS procedure is provided below.

The general conditions for which the absolute summation of two resultants method is conservative may be demonstrated by considering the following:

RHmin = fRHmax where 0 f 1 where f = RHmin/RHmax and RV = CRHmax 0 C < where C = RV/RHmax where:

RHmax = Larger of the two seismic co-directional responses due to either of the horizontal earthquake components CHAPTER 03 3.7-11 REV. 19, SEPTEMBER 2018

LGS UFSAR RHmin = Smaller seismic co-directional responses due to the other orthogonal horizontal earthquake component RV = Seismic co-directional response due to vertical earthquake f = ratio of RHmin to RHmax C = ratio of RV to RHmax.

Therefore SRSS of three components can be expressed as follows:

(R H max)2 + (R H min)2 = (R v)2 (R Hmax)2 + f2 (R Hmax)2 + C2 (R Hmax)2

( 1 + f2 + C2 R H max ) (EQ. 3.7-5)

The absolute sum of two components can be expressed as follows:

RHmax + RV RHmax + CRHmax

( (1 + C)2 ) RHmax

( 1 + 2C + C2 ) RHmax (EQ. 3.7-6)

Equation 3.7-6 (absolute sum of two components) will be greater than or equal to Equation 3.7-5 (three-component SRSS) when:

1 + 2C + C2 1 + f2 + C2 or 1 + 2C + C2 1 + f2 + C2 or 2C f2 or C 1/2 f2 where 0 f = RHmin 1 (EQ. 3.7-7 )

RHmax This relationship is shown in Figure 3.7-46.

The relative conservatism of the two-component absolute and the three-component SRSS summation techniques may be illustrated by considering the ratio of Equations 3.7-6 and 3.7-5.

This ratio will be defined as y. Then:

CHAPTER 03 3.7-12 REV. 19, SEPTEMBER 2018

LGS UFSAR y = ABS of EQ. 3.7-6 =

( 1 + 2C + C ) R2 H

max

( 1+f +C ) R 2 2 H

max SRSS of EQ. 3.7-5 1

1 + 2C + C2 2 y = (EQ. 3.7-8) 1 + f2 + C2 where:

0 f = R H min 1, as before R H max This relationship is shown in Figure 3.7-47.

As shown above, the LGS two-component absolute summation method is conservative when the co-directional response due to the vertical excitation is equal to or greater than one-half the higher of the two horizontal responses (C 1/2), regardless of the relationship between the two horizontal responses.

The minimum possible ratio between the two-component absolute summation method and the three-component SRSS procedure is equal to 0.707. This would occur only when the response due to the vertical excitation is zero (C = 0) and the two horizontal responses resulting from the two horizontal excitations are equal (f = 1). However, this case is unlikely to occur, and any other relationship of the various response components would produce a ratio larger than 0.707. The ratio between the two procedures, as shown in Figure 3.7-47 would be greater than one in most cases.

To further demonstrate the relationship of the seismic response components on the LGS structures, an evaluation has been performed for selected critical structural elements within the structures. Details of the evaluation are provided in the following paragraphs.

a. Containment Exterior Shell For the structural design of the containment shell, consideration of two horizontal components is not necessary due to the axisymmetric nature of the shell. The maximum resulting loads from two horizontal earthquake components would not be coincident and would occur 90 apart on the circumference of the shell.

Furthermore, when the resultant force from one horizontal component is maximum at a given location, the resultant force from the orthogonal horizontal component would be zero. This relationship corresponds to f = 0 in Figures 3.7-46 and 3.7-47.

Therefore, the two-component absolute summation technique would produce a more conservative design.

b. Reactor Enclosure and Control Structure CHAPTER 03 3.7-13 REV. 19, SEPTEMBER 2018

LGS UFSAR Stress evaluations were performed for critical locations in the reactor enclosure and control structure. The northeast control structure and southwest reactor enclosure corner model locations are selected because of their sensitivity to large orthogonal responses due to out-of-plane seismic motion. The results obtained from the SSE seismic analysis are combined by the two-component ABS and three-component SRSS methods and are compared in Table 3.7-20. For 14 out of 20 wall locations evaluated, the seismic stress response comparison shows that the two-component ABS method is more conservative than the three-component SRSS method. In general, the two-component ABS method and the three-component SRSS method produce comparable results. Although the differences between the combined stress from the two methods are small, the values for some locations suggest that the two-component ABS method may be conservative when the seismic load is considered to act alone.

There are 6 wall locations where the two-component ABS method is less conservative. The critical wall location is the southwest corner of the reactor enclosure, at el 177', where the differential between ABS stresses and SRSS stresses equals 14% (Table 3.7-20). (This location is considered critical not only because of the difference between the ABS and SRSS stresses but also because of the level of the stresses.) However, when the seismic stresses are combined with the stresses due to other design loads, the effect of the differential between ABS stresses and SRSS stresses is reduced in proportion with the ratio of seismic wall load to total wall load. In addition, the critical wall (wall line "D" at el 177') is loaded to only 52% of its capacity (based on combined axial and bending stresses due to all design loads as shown in Figure 3A-432).

Therefore, when considering the reduced effect of ABS vs. SRSS in conjunction with the ample excess capacity of the wall, the use of the ABS method for calculating seismic stresses provides adequate conservatism when considered in combination with other loads in the design of the reactor enclosure and control structure.

3.7.2.7 Combination of Modal Responses 3.7.2.7.1 Combination of Modal Responses (NSSS)

See Section 3.7.3.7.1.

3.7.2.7.2 Combination of Modal Responses (Non-NSSS)

The modal responses (i.e., shears, moments, deflections, accelerations, and inertia forces) are combined by either the sum of the ABS method, or by the SRSS method with consideration of closely spaced modes. Two consecutive modes are defined as closely spaced when their frequencies differ from each other by 10% or less of the lower frequency. When the SRSS method is used, Regulatory Guide 1.92 shall be adopted for the combination of modal responses.

3.7.2.8 Interaction of Non-Category I Structures with Seismic Category I Structures The turbine enclosure, auxiliary boiler enclosure, and the administration building are the only major non-Category I structures adjacent to seismic Category I structures. These non-Category I CHAPTER 03 3.7-14 REV. 19, SEPTEMBER 2018

LGS UFSAR structures are designed for seismic loading in accordance with the UBC (Reference 3.7-3). In addition, the turbine enclosure was dynamically analyzed to ensure the capacity to withstand a SSE without collapsing on or impairing the integrity of the adjacent reactor and control structures.

Similarly, the other non-Category I structures were analytically evaluated to ensure that they will not collapse on or otherwise impair the integrity of adjacent seismic Category I structures when subjected to the design seismic loads.

Structural separations have been provided to ensure that interaction between Category I and non-Category I structures does not occur. The minimum separation gap between the buildings is twice the relative displacement for the design seismic loads.

3.7.2.9 Effects of Parameter Variations on Floor Response Spectra 3.7.2.9.1 Effects of Parameter Variations on Floor Response Spectra (NSSS)

To account for potential variations in the primary structure frequencies due to uncertainties in material properties of the soil and structure, soil-structure interaction techniques, approximation in damping, and approximation in dynamic modeling, the computed floor response spectra are peak-broadened by +/-15%. This is consistent with the requirements of Regulatory Guide 1.122, although this regulatory guide is not the design basis requirement for the LGS construction permit.

3.7.2.9.2 Effects of Parameter Variations on Floor Response Spectra (Non-NSSS)

To account for variations in the structural frequencies owing to uncertainties in the material properties of the structure and to approximations in the modeling techniques used in the seismic analysis, the computed floor response spectra are smoothed, and peaks associated with each of the structural frequencies are broadened. In lieu of making a parametric study considering changes in the material properties and other variables, the spectrum is broadened on either side of the peak value by 15% of the frequency at which the peaks occur.

3.7.2.10 Use of Constant Vertical Static Factors Vertical seismic system multimass dynamic models are used to obtain vertical response loads for the seismic design of seismic Category I structures. Therefore, constant vertical static factors are not used to account for vertical response to earthquakes for the seismic design of Category I structures.

3.7.2.11 Methods Used to Account for Torsional Effects Torsional effects for the reactor enclosure, diesel generator enclosure, spray pond pumphouse, and radwaste enclosure are accounted for as follows:

A static analysis is performed to account for torsion on these structures. The eccentricity is determined using the distance between the center of mass and the center of rigidity of the individual structure. The inertial force from the response spectrum analysis is applied at the center of mass. The resulting torsional moment is equal to the inertial force times the eccentricity. The shear forces due to the torsional moment are then distributed to the walls. The total seismic shear in a given structural element is equal to the sum of the absolute values of the shear due to translation and torsion. Torsional effects are negligible for the containment because of the symmetry of the structure about the vertical center line.

CHAPTER 03 3.7-15 REV. 19, SEPTEMBER 2018

LGS UFSAR Dynamic analyses have been performed to verify the adequacy of the static-equivalent method described above. These dynamic analyses were performed for the reactor enclosure, diesel generator enclosure and spray pond pumphouse where structural eccentricities are larger than 5%

of the larger building plan dimension.

Inertia forces, obtained from dynamic analysis, are multiplied by the mass eccentricity (distance between center of mass and center of rigidity) to obtain a torsional moment, which is distributed to structural walls for design assessment. This equivalent static method was found to produce conservative design torsional moments; more so than results obtained from a 3-D dynamic analysis which included structural eccentricities. The dynamic analysis was performed as follows:

a. 3-D horizontal stick models were constructed for the reactor enclosure, diesel generator enclosure, and spray pond pumphouse with structural masses at calculated eccentricities (Figures 3.7-48, 3.7-49, and 3.7-50). The direction that has the larger mass eccentricity was selected for dynamic response analysis.

b. Modal analysis was performed to determine the vibration frequencies of the structures with eccentricity accounted for. The vibration frequencies with eccentricity were compared to the vibration frequencies with zero eccentricity (Tables 3.7-21 through 3.7-23).

c. Time history analyses were performed for the east-west models of the spray pond pumphouse, diesel generator enclosure, and reactor enclosure.

Torsional moments and shear forces at the base slab were compared with the moments and shear forces of the original design obtained using the equivalent static approach (Tables 3.7-24 through 3.7-27) .

Torsional moment and shear data comparisons show that the original design values using the static approach are conservative.

3.7.2.12 Comparison of Responses A comparison of the results of modal design response spectrum analysis and modal time history analysis for the containment is shown in Figures 3.7-4 through 3.7-9. The responses are calculated for both OBE and SSE. The uncracked containment model as described in Section 3.7.2.3.2 is used for the analyses. The responses due to synthetic time history input are lower than those due to response spectrum, because the sum of the absolute modal responses are used in the latter analysis, whereas the phase differences of the modal responses are considered in the time history analysis.

3.7.2.13 Methods for Seismic Analysis of Dams Dams are not provided on LGS.

3.7.2.14 Determination of Seismic Category I Structure Overturning Moments The overturning moment for seismic Category I structures is the absolute sum of the moments at the base of each stick of the mathematical model. For each stick, the moment at the base is CHAPTER 03 3.7-16 REV. 19, SEPTEMBER 2018

LGS UFSAR determined by combining the modal overturning moments. The modal moments are combined by the methods described in Section 3.7.2.7.

The components of the earthquake motion used are the same as those discussed in Section 3.7.2.6. Section 3.8.5 discusses the factor of safety against overturning for several loadings, including seismic loads.

3.7.2.15 Analysis Procedure for Damping 3.7.2.15.1 Analysis Procedure for Damping (NSSS)

In a linear dynamic analysis, the procedure to be utilized to properly account for damping in different elements of a coupled system model is as follows:

a. The values for structural damping of the various structural elements of the model are first specified. Each value is referred to as the damping ratio (Bj) of a particular component which contributes to the complete stiffness of the system.

b. Perform a modal analysis of the linear system model. This will result in a modal matrix () normalized so that:

iT Ki = Wi2 where:

K = the stiffness matrix Wi = circular natural frequency of mode i iT = the transpose () which is a column vector of ()

corresponding to the mode shape of mode i Matrix () contains all translational and rotational coordinates.

c. Using the strain energy of the individual components as a weighting function, the following equation can be derived to obtain a suitable damping ratio (Bi) for the ith mode.

N

[ i j Kj i]

T (EQ. 3.7-9)

Bi j=1 Wi2 where:

N = Total number of structural elements Mode shape for mode i ( i as transpose)

T i =

i = Percent damping associated with element j CHAPTER 03 3.7-17 REV. 19, SEPTEMBER 2018

LGS UFSAR Kj = Stiffness matrix of element j Wi = Circular natural frequency of mode i 3.7.2.15.2 Analysis Procedure for Damping (Non-NSSS)

The structures consist of reinforced concrete and welded/bolted structural steel. Damping values for these materials are shown in Table 3.7-2. However, in the seismic analysis of structures supported on rock, damping values of 2% and 5% are used for OBE and SSE respectively for welded/bolted structural steel, as well as for reinforced concrete. Therefore, the analysis by composite modal damping is not necessary.

For the time history analysis of containment and reactor enclosure where the soil-structure interaction is taken into account, the composite modal damping technique was used as described in appendix D of Reference 3.7-2.

3.7.3 SEISMIC SUBSYSTEM ANALYSIS This section discusses the seismic analysis of subsystems, i.e., equipment, piping, Class 1E cable trays, and supports for seismic Category I HVAC ducts and cable trays.

3.7.3.1 Seismic Analysis Methods 3.7.3.1.1 Equipment Seismic qualification of equipment was performed by using one of the following methods:

a. Analysis

b. Dynamic testing

c. Combination of analysis and dynamic testing 3.7.3.1.1.1 Analysis For the purpose of analysis, equipment is idealized as a system of lumped masses and springs, for which frequencies and mode shapes are determined for vibration in the vertical direction and two orthogonal horizontal directions. For each direction of vibration, the spectral acceleration per mode is obtained from the appropriate response spectrum curve (i.e., corresponding to the location of the equipment) at the natural frequency of the equipment. Seismic loading, in terms of inertia forces, moments and shears, is determined for each direction using the response spectrum technique, summing the absolute values per mode. If the orientation of the equipment is not designated on the equipment location drawing, the horizontal seismic loading is taken as the maximum loading (worst case) obtained, using each horizontal direction of vibration and the appropriate horizontal response spectrum curve(s). If the frequencies of all equipment modes (determined by either analysis or testing) are above the frequency of the appropriate response spectrum curve at which the acceleration is constant in the rigid (high frequency) range, the seismic loading consists of static loading corresponding to that acceleration level. If the equipment damping is unknown, the response spectrum curve for 0.5% damping is used to arrive at a conservative seismic loading.

CHAPTER 03 3.7-18 REV. 19, SEPTEMBER 2018

LGS UFSAR The damping value used for the OBE is increased for the SSE, where sufficient justification is established.

3.7.3.1.1.2 Dynamic Testing In lieu of performing dynamic analysis, seismic adequacy may be established by providing dynamic test or previous dynamic environmental (performance) data which demonstrate that the equipment meets the seismic design criteria. The data include at least one of the following:

a. Recent test data acquired from dynamic tests of equipment

b. Dynamic test data from previously tested comparable equipment

c. Performance data from equipment which, during normal operating conditions, have been subjected to dynamic loads equal to or greater than those defined in Section 3.7.3.1.1.1 Typical test methods used are as follows:

a. Single-frequency sine beat test

b. Single-frequency dwell test

c. Multifrequency test 3.7.3.1.1.3 Combination of Analysis and Dynamic Testing Certain equipment was qualified by a combination of analysis and dynamic testing. Experimental methods are used to aid in the formulation of the mathematical model for the equipment. Mode shapes and frequencies are determined experimentally and incorporated in the mathematical model of the equipment. The model is then subsequently analyzed by the procedure described in Section 3.7.3.1.1.1.

3.7.3.1.2 Piping Systems Reference 3.7-4 describes the methods used for seismic analysis of piping systems. Reference 3.7-4 is followed on LGS with the following exceptions:

In seismic analysis the modal responses are combined by SRSS, and lower damping values than specified in Reference 3.7-4 are used. Alternative analytical methods and damping values may be utilized as described in Section 3.7.1.3.3.

See Section 3.7.3.7.2.

3.7.3.1.3 Class 1E Cable Trays The cable trays are seismically qualified by the capacity evaluation method which consists of the following:

CHAPTER 03 3.7-19 REV. 19, SEPTEMBER 2018

LGS UFSAR

a. Calculation of the fundamental frequency of the cable tray based on the tray properties obtained from static tests

b. Seismic load computation based upon the tray frequency and the design spectra

c. Calculation of the tray allowable capacity

d. Evaluation of the tray capacity by interaction formula 3.7.3.1.4 Supports for Seismic Category I HVAC Ducts The supports for HVAC ducts are analyzed by the response spectrum method (Reference 3.7-2).

3.7.3.1.5 Supports for Seismic Category I Electrical Raceway Systems This section defines the procedures used for the design of the supports of electrical raceway systems, i.e., cable tray, conduit, and wireway gutter systems, subject to the seismic and other applicable loads. The raceway support system usually consists of raceways, horizontal and vertical support members, and lateral and longitudinal bracing members.

3.7.3.1.5.1 Loading Combinations The adequacy of raceway systems to withstand seismic and other applicable static loads is determined according to the loading combinations and allowable responses given in Table 3.7-3.

Load combinations and allowable stresses including the effects of hydrodynamic loads due to LOCA and SRV discharge are given in Table 3A-21.

3.7.3.1.5.2 Analytical Techniques Either of two methods of analysis is used. Method 1 is a simplified method of analysis which determine the fundamental frequency of braced supports using two-dimensional analysis.

Frequencies are determined in each of three principal directions. Then loads are determined by taking the spectral accelerations times the weight; and stresses are determined from static analysis. All members and connections are checked using stress criteria.

Method 2 uses a three-dimensional computer analysis and includes springs to represent joint stiffnesses. Response spectrum analyses are done to determine stresses and deformations. The number of stress cycles is determined by multiplying the time of maximum earthquake motion by the natural frequency of the system.

The allowable number of cycles is taken from Reference 3.7-8 for the joint rotations calculated.

Only overhead connections are checked for fatigue because the test results (Reference 3.7-8, p.

7-19) demonstrate that failures occur only in overhead connections.

The basis for the design criteria and analysis Method 2 is the "Cable Tray and Conduit Raceway Test Program" (References 3.7-7 through 3.7-10).

3.7.3.1.5.3 Damping CHAPTER 03 3.7-20 REV. 19, SEPTEMBER 2018

LGS UFSAR Damping of 10% of the critical is used for the design of cable tray support systems; 7% damping for conduit and wireway gutter trapeze-type support systems; 5% damping for conduit and wireway gutter nontrapeze-type support systems. The recommended damping values for cable tray and conduit systems, developed from the test program, are shown in Figure 3.7-45. Wireway gutters were not tested: however, the manner in which they are constructed (with more bolted connections and more cables than conduit) provides more damping mechanisms than are present in conduit systems so that using the same damping value as in conduit systems is conservative.

3.7.3.1.5.4 Operating Basis Earthquake The OBE is considered in the load combinations only for the overhead connections which are checked for fatigue. The OBE stresses are not checked during design for two reasons: first, raceway systems do not fail in a brittle or catastrophic mode as demonstrated by the test program in which such failures did not occur and the electrical systems were able to continue to function in all cases. Thus, there is no need to limit the OBE stresses to the low levels usually used to preclude such failures. Second, the OBE stresses will always be less than the SSE stresses as demonstrated below.

Based on Figure 3.7-45, a comparison of response spectra for corresponding damping values specified in Section 3.7.3.1.5.3 demonstrates that for all response spectra the OBE acceleration values are less than the corresponding SSE acceleration values (References 3.7-8 and 3.7-10).

Thus, the OBE acceleration response and stresses are below the SSE acceleration response and stresses.

3.7.3.2 Determination of Number of Earthquake Cycles 3.7.3.2.1 Determination of Number of Earthquake Cycles (NSSS) 3.7.3.2.1.1 NSSS Piping Fifty peak OBE cycles are postulated for fatigue evaluation.

3.7.3.2.1.2 Other NSSS Equipment and Components To evaluate the number of cycles which exist within a given earthquake, a typical BWR enclosure/reactor dynamic model was excited by three different recorded time histories: May 18, 1940, El Centro NS component 29.4 sec; 1952, Taft N 69o W component, 30 sec; and March 1957, Golden Gate S 80o E component, 13.2 seconds. The modal response is truncated so that the response of three different frequency bandwidths could be studied: 0-10 Hz; 10-20 Hz; and 20-50 Hz. This is 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 3.7-18 was formed. A comparison has shown that the LGS design basis response spectrum is bounded by the spectra of the three earthquakes (Golden Gate, Taft, and El Centro) in the GE base study (Figure 3.7-43).

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 was found in the following manner:

CHAPTER 03 3.7-21 REV. 19, SEPTEMBER 2018

LGS UFSAR

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 3.7-18 according to the frequency range within which the fundamental frequency lies.

c. For fatigue evaluation, 0.5% (0.005) of these cycles are conservatively assumed to be at the peak load and 4.5% (0.045) at or above three quarter peak. The remainder of the cycles have negligible contribution to fatigue usage.

The SSE 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 the evaluation is not required by ASME Section III.

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 of one-tenth of the SSE intensity, and one earthquake of 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, ten peak OBE cycles are postulated for fatigue evaluation.

Table 3.7-19 shows the calculated number of fatigue cycles and the number of fatigue cycles used in design.

3.7.3.2.2 Determination of Number of Earthquake Cycles (Non-NSSS)

In general, the design of the equipment is not fatigue controlled. For equipment qualified by analysis, fatigue is not a controlling factor because the equipment is designed to remain below 90% of the yield strength of the material for the extreme loading condition. The number of stress cycles considered is 60 (5 OBE and 1 SSE events at 10 cycles each). Based on ASME Section III, appendix I criteria (figure I-9-1), this number of cycles will not result in a reduction of allowable stresses.

Any fatigue effects in tested equipment are accounted for by the duration of the test.

Consequently, the number of cycles of the earthquake is considered.

In order to conduct a fatigue evaluation for nuclear Class I piping, the number of cycles for a given load set is obtained. This is done by considering ten maximum stress cycles per earthquake and 5 OBEs and 1 SSE to occur within the life of the plant. The results of the fatigue calculations for the most limiting BWR/4 component are shown in Table 3.7-4.

3.7.3.3 Procedures Used for Modeling 3.7.3.3.1 Procedures Used for Modeling (NSSS) 3.7.3.3.1.1 Modeling of Piping Systems CHAPTER 03 3.7-22 REV. 19, SEPTEMBER 2018

LGS UFSAR The continuous piping system is modeled as an assemblage of beams. The mass of each beam is lumped at the nodes connected by weightless elastic members representing the physical properties of each segment. The pipe lengths between mass points are no greater than the length which would have a natural frequency of 33 Hz, when calculated as a simply supported beam. 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 centers of gravity, with respect to the center line of the pipe, is included in the analytical model. Where the torsional effect is found to cause pipe stresses less than 500 psi, this effect is neglected.

3.7.3.3.1.2 Modeling of 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. A mode is considered significant if the corresponding natural frequency is less than 33 Hz, and the stresses calculated from this mode 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, because the natural frequencies of the equipment are generally in the higher frequency range of the floor spectra.

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 higher than the frequency at which the floor spectra peak occurs. If such is not the case, the model is adjusted to give more conservative results.

3.7.3.3.1.3 Field Location of Supports and Restraints The field location of seismic supports and restraints for seismic Category I piping and piping systems components was selected to satisfy the following two conditions:

a. The location selected must furnish the required response to control strain within allowable limits.

b. Adequate structure strength for attachment of the components must be available.

The final location of seismic supports and restraints for seismic Category I piping, piping system components, and equipment (including the placement of snubbers), was checked against the CHAPTER 03 3.7-23 REV. 19, SEPTEMBER 2018

LGS UFSAR drawings and instructions issued by the engineer. An additional examination of these supports and restraining devices was made to ensure that the location and characteristics of these supports and restraining devices are consistent with the dynamic and static analyses of the systems.

3.7.3.3.2 Procedures Used for Modeling (Non-NSSS)

Mathematical models which describe the mass and the stiffness properties of the equipment are used. The models define the dynamic behavior of the equipment within the frequency range of interest. The boundary conditions are modeled to reflect the actual mounting conditions. The equipment is represented by lumped-mass models. Massless elastic members are used to connect the masses.

Supports for HVAC ducts are modeled as two-dimensional, lumped- mass, plane frame models.

The masses are lumped at the center of the ducts. The electrical raceway support system analytical techniques are discussed in Section 3.7.3.1.5.2. Equivalent structural properties for cable tray analysis were determined from the load-deflection tests (Reference 3.7-11).

Sections 2 and 3 of Reference 3.7-4 discuss the techniques and procedures used to model piping other than the buried type.

3.7.3.4 Basis of Selection of Frequencies 3.7.3.4.1 Basis of Selection of Frequencies (NSSS)

All frequencies in the range of 0.25-33 Hz are considered in the analysis and testing of structures, systems, and components. These frequencies cover the natural frequencies of most of the components and structures under consideration. If the fundamental frequency of a component is 33 Hz, it is treated as rigid and analyzed accordingly. Frequencies less than 0.25 Hz are not considered, as they represent very flexible structures and are not encountered in this plant.

The frequency range of between 0.25 Hz and 33 Hz covers the range of the broad-band response spectrum used in the design.

3.7.3.4.2 Basis of Selection of Frequencies (Non-NSSS)

The natural frequencies of components are calculated. Only those modes which have natural frequencies <33 Hz are considered in the dynamic analysis. If a component has a frequency 33 Hz, it is considered as rigid. If the natural frequency of the component falls within the broadened peak of the response spectrum curve, then it is designed to take the applied load.

3.7.3.5 Use of Equivalent Static Load Method of Analysis (Non-NSSS)

The equivalent static load method is used when the natural frequency of the equipment is not determined. If the equipment can be adequately represented by a single-degree-of-freedom system, then the applied inertia load is equal to the weight of the equipment times the peak value of the response spectrum curve. Seismic acceleration coefficients for multidegree of freedom systems, which may be in the resonance region of the amplified response spectra curves, are increased by 50% to account conservatively for the increased modal participation.

Appendix D of Reference 3.7-4 discusses the use of equivalent static load method of analysis as applicable to piping.

CHAPTER 03 3.7-24 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.7.3.6 Three Components of Earthquake Motion 3.7.3.6.1 Three Components of Earthquake Motion (NSSS)

The simultaneous use of three components of earthquake motion was not a design basis requirement of the construction permit for the LGS plant. However, the NSSS systems and components are evaluated to the requirement of Regulatory Guide 1.92.

a. Response Spectrum Method Response spectra generated by GE are developed considering three components of earthquake motion. The individual responses in each orthogonal direction are combined by SRSS of the colinear contribution due to the three directions of earthquake motion. These are used to predict the total response at each frequency.

b. Time history Method When the time history method of analysis is used, one of the following options is used to obtain the peak value of any particular response of interest:

1. When maximum colinear contributions due to the three directions of earthquake motion are calculated separately, the total response is obtained as the SRSS combination of the colinear values.

2. When colinear time history responses from each of the three components of the earthquake motion are calculated individually by the step-by-step method and then combined algebraically at each time step, the maximum response is obtained as the peak value from the combined time solution.

3. When a response at each time step is calculated directly based on the simultaneous application of the three earthquake components, the maximum response is determined by scanning the combined time history solution.

The components of earthquake motion must be statistically independent for Options 2 and 3. Also, the time history method precludes the need to consider closely spaced modes.

3.7.3.6.2 Three Components of Earthquake Motion (Non-NSSS)

For equipment, cable trays, and supports for cable trays and HVAC ducts, the three spatial components of the earthquake are considered in the same manner as for structures (Section 3.7.2.6).

The criteria used for combining the results of horizontal and vertical seismic responses for piping systems are described in section 5.1 of Reference 3.7-4.

3.7.3.7 Combination of Modal Responses 3.7.3.7.1 Combination of Modal Responses (NSSS)

CHAPTER 03 3.7-25 REV. 19, SEPTEMBER 2018

LGS UFSAR All piping and equipment analyzed or supplied by GE are evaluated to the requirements of Regulatory Guide 1.92.

When the response spectra method of modal analysis is used, all modes except the closely spaced modes (i.e., the difference between any two natural frequencies is equal to or less than 10%) are combined by the SRSS as described in Section 3.7.3.7.1.a. Closely spaced modes are combined by the double sum method with absolute sign as described in Section 3.7.3.7.1.b.

In the time history method of dynamic analysis, the vector sum at every time step is used to calculate the combined response. The use of the time history method precludes the need to consider modal spacing.

a. SRSS The SRSS method is defined mathematically as:

n 1/2 R = (Ri)2 l (EQ. 3.7-10) i=1 where:

R = Combined response Ri = Response in the (i) mode n = Number of modes considered in the analysis

b. Procedure of Combining Closely Spaced Modal Response This method is defined mathematically as:

1/2 N N l R = RkRs Eks l (EQ. 3.7-11 )

k=1 s=1 l where:

R = representative maximum values of a particular response of a given element to a given component of excitation Rk = peak value of the response of the element due to the kth mode N = number of significant modes considered in the modal response combination Rs = peak value of the response of the element attributed to the sth mode CHAPTER 03 3.7-26 REV. 19, SEPTEMBER 2018

LGS UFSAR also:

2 1 1 Wk - Ws l Eks = 1 + l (EQ. 3.7-12) 1 1 wk + s ws l k

in which:

1 2 wk = wk (1 - )1/2 and k

1 k = k + 2 tdwk where:

wk = modal frequency k = damping ratio in the kth mode td = duration of the earthquake.

3.7.3.7.2 Combination of Modal Responses (Non-NSSS)

The modal responses of equipment are combined by the SRSS method. The absolute values of two closely spaced modes are added first before combining with the other modes by the SRSS method. Two consecutive modes are defined as closely spaced when their frequencies differ from each other by 10% or less.

Procedures given in Regulatory Guide 1.92 for combining modal responses, when closely spaced modes are present, are not complied with in the seismic response spectra analysis for piping. All modal responses are combined by SRSS in the response spectra method of modal analysis for seismic loading (OBE and SSE). Seismic response spectra used in the piping analysis corresponds to conservative damping values of 1/2% for OBE and 1% for SSE. Alternative analytical methods and damping values may be utilized as described in Section 3.7.1.3.3.

3.7.3.8 Analytical Procedure for Piping 3.7.3.8.1 Analytical Procedure for Piping (NSSS)

The analytical procedures for piping analysis have been described in Section 3.7.2.1.1.5.a.

Methods to include differential piping support movements at different support points are also described there.

3.7.3.8.2 Analytical Procedure for Piping (Non-NSSS)

CHAPTER 03 3.7-27 REV. 19, SEPTEMBER 2018

LGS UFSAR The design criteria and the analytical procedures applicable to piping systems are as described in section 2 of Reference 3.7-4. Alternatively, except as restricted by Section 3.7.1.3.3, the multiple response spectra/independent support motion method of analysis may be used where distinct response spectra are applied to the piping system attachment points. The methods used to consider differential piping support movements at different support points are described in section 4 of Reference 3.7-4.

3.7.3.9 Multiple Supported Equipment Components with Distinct Inputs 3.7.3.9.1 Multiple Supported Equipment Components with Distinct Inputs (NSSS)

The procedure and criteria for analysis has been described in Section 3.7.2.1.1.5.b.

3.7.3.9.2 Multiple Supported Equipment Components with Distinct Inputs (Non-NSSS)

For cable trays and HVAC ducts whose supports have two or more distinct inputs, a response spectrum curve envelopes the curves at all support locations. When a piping system whose supports have two or more distinct inputs, a response spectrum curve may be used for analysis that envelopes the curves at all locations. Alternatively, except as restricted by Section 3.7.1.3.3, the multiple response spectra/independent support motion method of analysis may be used where distinct response spectra are applied to the piping system attachment points. Section 4 of Reference 3.7-4 discusses the methods used for the analysis of multiple supported piping due to differential seismic anchor movement.

3.7.3.10 Use of Constant Vertical Static Factors 3.7.3.10.1 Use of Constant Vertical Static Factors (NSSS)

Constant vertical static factors are not used for NSSS components.

3.7.3.10.2 Use of Constant Vertical Static Factors (Non-NSSS)

Constant vertical static factors are not used in the seismic design of subsystems.

3.7.3.11 Torsional Effects of Eccentric Masses 3.7.3.11.1 Torsional Effects of Eccentric Masses (NSSS)

Torsional effects of eccentric masses is discussed in Section 3.7.3.3.1.1.

3.7.3.11.2 Torsional Effects of Eccentric Masses (Non-NSSS)

The torsional effects of valves and other eccentric masses are considered in the seismic analysis of piping by the techniques discussed in section 3.2 of Reference 3.7-4.

CHAPTER 03 3.7-28 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.7.3.12 Buried Seismic Category I Piping Systems and Tunnels (Non-NSSS)

Buried seismic Category I piping is analyzed and designed for seismic effects in accordance with section 6.0 of Reference 3.7-2. For portions of buried piping in soil, the soil response along the pipeline is determined in accordance with the procedure in Section 2.5.4.7.1.

The diesel oil storage tank structures and the ESW/RHRSW pipe tunnel (located beneath the diesel generator enclosure) are seismic Category I and are analyzed using an equivalent static load method.

Other tunnels at LGS are nonseismic Category I.

3.7.3.13 Interaction of other Piping with Seismic category I Piping 3.7.3.13.1 Interaction of other Piping with seismic Category I Piping (NSSS)

When other piping is attached to seismic Category I piping, the other piping is analytically simulated in a manner that does not significantly degrade the accuracy of the seismic Category I piping analysis. Furthermore, the other piping is designed to withstand the SSE without failing in a manner that would cause the seismic Category I piping to fail.

3.7.3.13.2 Interaction of Other Piping with seismic Category I Piping (Non-NSSS)

The techniques used to consider the interaction of seismic Category I piping with non-Category I piping are as follows:

Seismic boundary anchors are designed for the combined loads generated from both sides of a boundary anchor. The loads from the seismic Category I side are actual calculated loads, and the loads from the nonseismic Category I side are determined by one of the following:

a. The actual calculated seismic loads if the nonseismic side piping is dynamically analyzed for seismic events

b. The actual calculated loads if the nonseismic side piping is designed to a conservative simplified seismic design criteria (e.g., by simplified span methods such as those used for designed of small piping)

c. The loads determined by the plastic capability of the piping.

3.7.3.14 Seismic Analysis for Reactor Internals (NSSS)

The modeling of RPV internals is discussed in section 3.7.2.3.1.2. The damping values are given in Table 3.7-1.

3.7.3.15 Analysis Procedures for Damping 3.7.3.15.1 Analysis Procedures for Damping (NSSS)

Analysis procedures for damping are discussed in section 3.7.2.15.1.

CHAPTER 03 3.7-29 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.7.3.15.2 Analysis Procedure for Damping (Non-NSSS)

If the equipment damping is unknown, the response spectrum curve for 0.5% damping is used to arrive at a conservative seismic loading. The damping values used for the OBE are increased for the SSE, where sufficient justification is established.

3.7.4 SEISMIC INSTRUMENTATION 3.7.4.1 Comparison with Seismic Monitoring System Regulatory Guides The seismic monitoring system instrumentation was designed to comply with the intent of Regulatory Guide 1.12 (Rev. 2), Regulatory Guide 1.166 and Regulatory Guide 1.167.

3.7.4.1.1 Comparison to Regulatory Guide 1.12, Rev. 2 Regulatory Guide 1.12 describes the seismic instrumentation that is acceptable to the NRC for satisfying the requirements of Part 20 and Appendix S to Part 50.

Limerick is in compliance with the guidance with this Reg Guide and takes no exceptions.

3.7.4.1.2 Comparison to Regulatory Guide 1.166 Regulatory Guide 1.166 describes the pre-earthquake planning and post earthquake responses to a seismic event. The Reg Guide states that the response spectrum and Cumulative Absolute Velocity (CA V) be calculated to determine OBE exceedance. Limerick is in compliance with the guidance with this Reg Guide and takes no exceptions.

3.7.4.1.3 Comparison to Regulatory Guide 1.167 Regulatory Guide 1.167 describes the guidance on the restart of a nuclear reactor after a seismic event. Limerick is in compliance with the guidance with this Reg Guide and takes no exceptions.

3.7.4.2 Location and Description of Instrumentation The following instrumentation is provided for Unit 1 only, as essentially the same response is expected at Unit 2.

a. Six triaxial time history Strong Motion Recorders with built in sensors.

b. One response spectrum analyzer (software run on local laptop computer).

c. A system control panel which provides seismic event visual and audible annunciators.

d. Analysis software for monitoring of equipment and determination of exceedance of QBE based on CAV and response spectra.

CHAPTER 03 3.7-30 REV. 19, SEPTEMBER 2018

LGS UFSAR All instrument characteristics meet the requirements of Regulatory Guide 1.12, Rev. 2, Section 4.0.

The operability of the seismic monitoring instrumentation ensures that sufficient capability is available to promptly determine the magnitude of a seismic event and evaluate the response of those features important to safety. This capability is required to permit comparison of the measured response to that used in the design basis for the unit.

3.7.4.2.1 Triaxial Time History Strong Motion Recorders Triaxial time history Strong Motion Recorders (SMR) contain solid state seismic sensors and they measure and record the time varying acceleration at the sensor location. This data is stored locally and then sent to the main control panel when requested by the system for conversion and comparison with reference information. Each SMR contains a three axis solid state accelerometer mounted inside the SMR. All SMR's have their principal axes oriented identically, with one horizontal axis parallel to the major horizontal axis assumed in the seismic analysis. The SMR's are located as shown in table 3. 7-28.

3.7.4.2.2 Remote Triaxial Time History Strong Motion Recorders There are two self contained Strong Motion Recorders (SMR) located in the Spary Pond Pump House. The Spray Pond Pump House is the Alternate Category 1 building being monitored per Reg Guide 1.12. The first SMR is located at the building foundation at elevation 237' and the second SMR is located at elevation 268'. These two SMR's are not connected to the seismic control panel in the Main Control Room (MCR). These SMR are accessible by a dedicated laptop for download of the time history seismic response.

3.7.4.2.3 Triaxial Seismic Switch (DELETED) 3.7.4.2.4 Response Spectrum and CAV Analyzer The response spectrum and CAV analyzer is software that resides in the Network Control Center (NCC) and Earthquake Analysis PC which is located in the seismic control panel in the MCR. This software is used for calculation of CAV based on the recorded seismic response to an earthquake, and to compare the recorded data with the site design spectra for each of the connected SMRs location. It determines if the response spectra or CAV has exceeded the OBE criteria for signaling the local annunciator to alarm.

3.7.4.2.5 System Control Panel A panel located in the control room (OO-C693) houses the monitoring, and spectrum analysis units which are used in conjunction with the SMR sensors to produce a time history and frequency-amplitude record of the seismic event. The panel also contains display and alarm equipment associated with the NCC, Earthquake Analysis PC, and the system power supply units.

3.7.4.3 Control Room Operator Notification Activation of the seismic monitoring system causes an audible and visual annunciation in the control room alerting plant operators in the MGR to the seismic monitoring system panel.

CHAPTER 03 3.7-31 REV. 19, SEPTEMBER 2018

LGS UFSAR Should the system or seismic SMRs develop errors or equipment failures, the system will activate an audible and visual Trouble alarm at the Seismic Control panel.

Should the system sense a seismic event from the Free Field or the reactor building foundation sensor, of a large enough magnitude, the system will activate an audible and visual System Triggered alarm at the Seismic Control panel.

Should the system sense and analyze a seismic event that exceeds the OBE limits, the system will activate an audible and visual OBE Exceeded alarm at the Seismic Control panel.

The spray pond pump house SMRs are not interconnected, and have only local indication of error, data, run, and power. The locally recorded data from these two SMRs can be downloaded and analyzed.

3.7.4.4 Comparison of Measured and Predicted Responses Initial determination of the seismic event level is performed immediately after the event by comparing the measured response spectra from the Free Field with the calculated OBE and SSE response spectra for the corresponding location. An outline of the order of actions to be taken after a seismic event is provided in Figure 3.7-44.

3.7.4.5 Controls Applicability: The seismic monitoring instrumentation shall be operable at all times. This instrumentation is considered operable when it is capable of performing its specified functions and when all necessary attendant instrumentation, controls, electrical power, cooling or seal water, lubrication or other auxiliary equipment that are required for the system, subsystem, train, component, or device to perform its functions are also capable of performing their related support functions.

Actions: For requirements when one or more seismic monitoring instruments are inoperable, reference the TRM actions.

3.7.4.6 Surveillance Requirements Each of the seismic monitoring instruments shall be demonstrated operable by the performance of the channel check, channel functional test and channel calibration at the frequencies shown and defined in Table 3.7-29.

Each of the seismic monitoring instruments which is accessible during power operation and which is actuated during a seismic event greater than or equal to 0.01g, and which does not self-reset, shall be restored to operable status within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> and a channel calibration performed within 5 days following the seismic event. Data shall be retrieved from the actuated instruments and analyzed to determine the magnitude of the vibratory ground motion. A Special Report shall be prepared and submitted to the Nuclear Regulatory Commission pursuant to Specification 6.9.2 of the Technical Specifications within ten days describing the magnitude, frequency spectrum and resultant effect upon unit features important to safety.

CHAPTER 03 3.7-32 REV. 19, SEPTEMBER 2018

LGS UFSAR 3.

7.5 REFERENCES

3.7-1 N.C, Tsai, "Spectrum Compatible Motions for Design Purposes", Journal of Engineering Mechanics Division, ASCE, Vol, 98, No. EM2, Proc, Paper 8807, pp.

345-356 (April 1972) .

3.7-2 "Seismic Analyses of Structures and Equipment for Nuclear Power Plants",

BC-TOP-4A, Rev. 3, Bechtel Power Corporation, San Francisco, California, (November 1974).

3.7-3 "Uniform Building Code", by International Conference of Building Officials, Whittier, California, 1970 Edition.

3.7-4 "Seismic Analysis of Piping Systems" BP-TOP-1 Rev. 3. Bechtel Power Corporation, San Francisco, California, (January 1976).

3.7-5 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).

3.7-6 N.M. Newmark, "Design Criteria for Nuclear Reactors Subject to Earthquake Hazards", Proc IAEA Panel on Aseismic Design and Testing of Nuclear Facilities, Japan Earthquake Engineering Promotion Society, Tokyo, Japan, (1967).

3.7-7 "Development of Analysis and Design Techniques from Dynamic Testing of Electrical Raceway Support Systems", Technical Report, Bechtel Power Corporation, (July 1979).

3.7-8 Cable Tray and Conduit Raceway Seismic Test Program- Release 4", Test Report

1. 1053-21.1-4, Volumes 1 and 2, ANCO Engineers, Inc., (December 15, 1978).

3.7-9 P.Y. Hatago and G.S. Reimer, "Dynamic Testing of Electrical Raceway Support Systems for Economical Nuclear Power Plant Installations", presented at the IEEE-PES, (February 4-9, 1979).

3.7-10 "Cable Tray and Conduit Raceway Seismic Test Program- Release 4", Addendum to Test Report #1053-21.1-4, Volume 3, ANCO Engineers, Inc., (May 1980).

3.7-11 "Cable Trays Seismic Qualification Report for the Limerick Generating Station Units 1 and 2", Specification 8031-E-49, P-W Industries, Inc., (September 18, 1975).

3.7-12 "Safety Evaluation Report Relating to the Use of ASME Code Case N-411 for Limerick Generating Station, Unit 2", R.M. Bernero (NRC), (March 27, 1987).

3.7-13 NRC Letter to PECO, dated April 10, 1991, Snubber Reduction Program, Limerick Generating Station, Unit 1 (TAC No. 80069).

CHAPTER 03 3.7-33 REV. 19, SEPTEMBER 2018

LGS UFSAR Table 3.7-1 CRITICAL DAMPING VALUES FOR NSSS MATERIALS(1,2)

CRITICAL DAMPING (%)

ITEM OBE CONDITION SSE CONDITION Reinforced concrete structures 2.0 5.0 Welded structural assemblies 1.0 2.0 (equipment and supports)

Bolted or riveted structural 2.0 3.0 assemblies Vital piping systems 0.5 1.0 Drywell (coupled) 2.0 5.0 RPV support skirt, shroud head, 2.0 2.0 separator, and guide tubes CRD housings 3.5 3.5 Fuel 7.0 7.0 Steel frame structures 2.0 3.0 (1)

Other values may be used if they are indicated to be reliable by experiment or study.

(2)

Alternative critical damping values for piping systems may be used as described in Section 3.7.1.3.3.

CHAPTER 03 3.7-34 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-2 CRITICAL DAMPING VALUES FOR NON-NSSS MATERIALS(1)

CRITICAL DAMPING (%)

ITEM OBE CONDITION SSE CONDITION Equipment 0.5 1 Piping systems 0.5 1 Welded steel structures 2 5 Bolted steel structures 3 7 Reinforced concrete structures 2 5 (except as noted below for specific loading conditions in primary containment)

Primary containment (for those 3 7 loading conditions in which DBA and seismic loadings are combined)

(1)

Alternative critical damping values for piping systems may be used as described in Section 3.7.1.3.3 CHAPTER 03 3.7-35 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-3 LOAD COMBINATIONS AND ALLOWABLE RESPONSES FOR ELECTRICAL RACEWAY SYSTEM Equation Condition Load Combination(1) Allowable Response(1 )

1 Normal D+L Fs(4) 2(5) Normal/ D+E (2,4)

Severe 3 Abnormal/ D+E' (2,3,4)

Extreme (1)

For notations, see Tables 3.8-2 and 3.8-9.

(2)

The following equation is applicable for bending in overhead connections:

5nEQ + nEQ 1.0 NOBE NSSE where:

nEQ = Total number of load/stress cycles per earthquake.

NOBE = Allowable number of load/stress cycles per OBE event.

NSSE = Allowable number of load/stress cycles per SSE event.

(3)

The following criteria are used for checking the members. In no case shall the allowable stress exceed 0.90 Fy in bending, 0.85 Fy in axial tension or compression, and 0.50 Fy in shear. Where the design is governed by requirements of stability (local or lateral buckling),

the actual stress shall not exceed 1.5 Fs.

(4)

Allowable shear and normal loads in connections are determined from the manufacturers' data or from code allowable stresses, whichever is applicable. The allowable values are increased 50% for load combination equation 3.

(5)

Equation 2 applies only to connections for fatigue considerations.

CHAPTER 03 3.7-36 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-4 RESULTS OF FATIGUE CALCULATIONS FOR THE MOST LIMITING BWR/4 COMPONENT BWR/4 RPV FEEDWATER NOZZLE(1)

Loading Fatigue Usage 10 OBE Cycles 0.006 All Others(2) 0.967 Total 0.973 (1)

The most limiting calculation for the BWR/4 product line.

(2)

All other fatigue contributions due to SRV, thermal operating transients, etc.

CHAPTER 03 3.7-37 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-5 NATURAL FREQUENCIES OF PRIMARY CONTAINMENT BELOW 33 Hz FREQUENCY (Hz)

MODE HORIZONTAL VERTICAL NO. UNCRACKED CRACKED UNCRACKED CRACKED 1 4.496 2.878 10.620 7.731 2 8.545 7.589 17.573 12.857 3 15.393 9.864 - 19.809 4 17.831 14.836 - 26.954 5 27.389 17.221 - -

6 29.135 22.085 - -

7 - 23.735 - -

8 - 28.448 - -

9 - 31.175 - -

CHAPTER 03 3.7-38 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-6 NATURAL FREQUENCIES OF THE REACTOR ENCLOSURE AND CONTROL STRUCTURE BELOW 33 Hz FREQUENCY (Hz)

MODE NO. VERTICAL N-S E-W 1 1.993 2.495 3.235 2 3.007 7.757 10.515 3 3.272 11.531 15.108 4 3.888 16.119 22.063 5 4.043 21.310 30.057 6 4.154 28.090 -

7 4.262 - -

8 4.314 - -

9 4.546 - -

10 4.575 - -

11 4.630 - -

12 4.862 - -

13 5.088 - -

14 5.959 - -

15 7.949 - -

16 9.339 - -

17 9.975 - -

18 10.069 - -

19 10.112 - -

20 10.329 - -

21 10.618 - -

22 10.661 - -

23 10.814 - -

24 11.053 - -

25 11.333 - -

26 12.553 - -

27 12.998 - -

28 13.254 - -

29 13.328 - -

30 13.527 - -

31 13.731 - -

32 13.882 - -

33 13.923 - -

34 14.896 - -

35 17.918 - -

36 22.877 - -

37 24.077 - -

38 25.244 - -

39 30.395 - -

CHAPTER 03 3.7-39 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-7 PRIMARY CONTAINMENT SEISMIC INERTIA FORCES, DISPLACEMENTS, AND ACCELERATIONS (HORIZONTAL DIRECTION)

INERTIA FORCES DISPLACEMENTS ACCELERATIONS (kips) (10-2 ft) (g)

UNCRACKED CRACKED UNCRACKED CRACKED UNCRACKED CRACKED CONDITION CONDITION CONDITION CONDITION CONDITION CONDITION MASS POINT OBE SSE OBE SSE OBE SSE OBE SSE OBE SSE OBE SSE 1 141 230 179 269 1.738 2.663 4.330 6.125 0.735 1.193 0.928 1.390 2 669 1083 816 1216 1.650 2.525 4.038 5.750 0.685 1.108 0.835 1.245 3 1346 2165 1481 2169 1.513 2.313 3.625 5.163 0.610 0.980 0.670 0.981 4 1321 2098 1330 1913 1.338 2.050 3.113 4.425 0.509 0.808 0.511 0.736 5 1114 1749 1254 1854 1.189 1.813 2.713 3.850 0.428 0.671 0.481 0.771 6 971 1520 1105 1658 1.054 1.613 2.325 3.313 0.361 0.566 0.411 0.618 7 739 1151 913 1360 0.941 1.438 2.050 2.925 0.308 0.480 0.380 0.566 8 733 1160 868 1275 0.848 1.300 1.813 2.575 0.306 0.488 0.363 0.533 9 649 1038 771 1135 0.773 1.184 1.613 2.300 0.300 0.480 0.358 0.526 10 2251 3631 2624 3850 0.700 1.074 1.425 2.038 0.291 0.470 0.340 0.498 11 993 1614 1171 1730 0.551 0.849 1.061 1.513 0.258 0.419 0.305 0.450 12 1153 1895 1463 2226 0.405 0.625 0.713 1.019 0.219 0.360 0.279 0.424 13 1271 2156 1356 2110 0.263 0.409 0.378 0.543 0.186 0.315 0.199 0.309 14 1976 3431 1509 2410 0.160 0.251 0.148 0.215 0.150 0.260 0.114 0.183 15 200 333 179 264 1.713 2.788 3.850 5.488 0.714 1.186 0.636 0.941 16 326 536 281 413 1.588 2.425 3.238 4.600 0.573 0.940 0.493 0.724 17 293 473 290 421 1.313 2.013 2.663 3.788 0.469 0.758 0.465 0.675 18 234 375 251 364 1.159 1.775 2.363 3.363 0.409 0.655 0.439 0.635 19 233 376 240 353 0.915 1.400 1.838 2.613 0.346 0.563 0.359 0.526 20 188 306 238 366 0.534 0.823 0.989 1.413 0.266 0.435 0.339 0.523 21 280 469 355 565 0.379 0.586 0.615 0.885 0.244 0.409 0.310 0.493 22 384 661 411 664 0.244 0.380 0.310 0.449 0.200 0.345 0.215 0.346 23 560 876 445 644 3.325 5.075 5.425 7.713 1.249 2.113 1.073 1.550 24 866 1334 755 1081 2.713 4.150 4.588 6.525 0.991 1.526 0.864 1.238 25 896 1413 793 1143 2.150 3.288 3.788 5.388 0.816 1.286 0.721 1.040 26 764 1223 720 1045 1.634 2.513 3.063 4.350 0.628 1.005 0.591 0.859 27 - - - - 1.413 2.163 2.738 3.900 - - - -

28 - - - - 1.159 1.775 2.363 3.363 - - - -

29 - - - - 1.011 1.550 2.088 2.963 - - - -

30 - - - - 0.709 1.088 1.400 2.000 - - - -

31 - - - - 0.160 0.251 0.148 0.215 - - - -

32 - - - - 3.213 4.900 5.275 7.488 - - - -

33 - - - - 1.663 2.563 4.100 5.828 - - - -

34 - - - - 1.216 1.863 2.450 3.482 - - - -

CHAPTER 03 3.7-40 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-8 PRIMARY CONTAINMENT SEISMIC DISPLACEMENTS AND ACCELERATIONS (VERTICAL DIRECTION)

DISPLACEMENTS ACCELERATIONS (10-2 ft) (g)

UNCRACKED CRACKED UNCRACKED CRACKED CONDITION CONDITION CONDITION CONDITION MASS POINT OBE SSE OBE SSE OBE SSE OBE SSE 1 0.133 0.201 0.386 0.551 0.205 0.318 0.395 0.591 2 0.133 0.200 0.386 0.551 0.204 0.316 0.394 0.589 3 0.133 0.200 0.376 0.535 0.201 0.314 0.368 0.543 4 0.131 0.198 0.353 0.500 0.198 0.305 0.321 0.466 5 0.129 0.195 0.329 0.466 0.193 0.298 0.285 0.411 6 0.128 0.191 0.304 0.433 0.188 0.289 0.258 0.373 7 0.125 0.189 0.284 0.404 0.184 0.283 0.244 0.355 8 0.123 0.185 0.264 0.375 0.180 0.275 0.233 0.343 9 0.121 0.183 0.245 0.349 0.175 0.269 0.221 0.326 10 0.119 0.179 0.226 0.321 0.171 0.261 0.206 0.308 11 0.114 0.171 0.190 0.271 0.165 0.251 0.178 0.264 12 0.109 0.164 0.155 0.220 0.159 0.243 0.151 0.224 13 0.103 0.155 0.116 0.168 0.150 0.231 0.129 0.198 14 0.098 0.148 0.088 0.125 0.145 0.223 0.109 0.166 15 0.169 0.266 0.171 0.243 0.260 0.403 0.238 0.345 16 0.168 0.255 0.170 0.241 0.256 0.398 0.235 0.341 17 0.165 0.250 0.168 0.238 0.249 0.385 0.228 0.329 18 0.163 0.248 0.166 0.235 0.248 0.375 0.223 0.320 19 0.155 0.235 0.155 0.220 0.234 0.360 0.200 0.288 20 0.146 0.223 0.144 0.203 0.221 0.343 0.180 0.259 21 0.138 0.203 0.128 0.180 0.201 0.310 0.154 0.223 22 0.121 0.183 0.113 0.160 0.179 0.275 0.138 0.203 23 0.108 0.161 0.099 0.140 0.155 0.238 0.121 0.183 24 0.184 0.280 0.186 0.264 0.293 0.459 0.275 0.405 25 0.183 0.278 0.185 0.261 0.289 0.451 0.270 0.398 26 0.179 0.271 0.181 0.256 0.278 0.433 0.259 0.378 27 0.174 0.264 0.176 0.250 0.266 0.413 0.245 0.356 28 0.159 0.241 0.161 0.228 - - - -

29 0.170 0.258 0.173 0.248 - - - -

30 0.165 0.250 0.168 0.238 - - - -

CHAPTER 03 3.7-41 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-9 PRIMARY CONTAINMENT SHEAR FORCES AND MOMENTS SHEAR FORCES MOMENTS (kips) (103k-ft)

UNCRACKED CRACKED UNCRACKED CRACKED CONDITION CONDITION CONDITION CONDITION JOINT MEMBER NO. NO. OBE SSE OBE SSE OBE SSE OBE SSE 0 0 0 0 1 141 230 179 269 33 0.9 1.5 1.1 1.8 34 628 966 966 796 2 2.0 3.1 2.3 3.3 2 1,098 1,746 1,274 1,876 3 16.3 25.9 18.8 27.6 3 3,454 5,438 3,475 5,060 4 70.1 110.5 73.3 106.9 4 4,261 6,744 4,408 6,364 5 125.5 198.1 130.5 189.6 5 5,179 8,183 4,974 7,133 6 184.3 291.1 189.1 273.3 6 6,095 9,584 5,366 7,684 7 237.9 375.9 239.8 344.6 7 6,749 10,561 6,061 8,735 8 300.9 474.6 291.1 416.8 8 7,285 11,345 6,614 9,558 9 359.1 565.3 338.9 477.1 9 7,715 11,995 6,944 10,038 10 418.6/ 657.5/ 374.4/ 534.1/

420.5 654.6 371.6 530.4 10 9,278 14,263 7,835 11,308 11 546.4 850.9 453.4 646.0 11 10,019 15,438 8,716 12,538 12 680.4 1052.8 543.9 776.0 12 11,085 17,163 9,455 13,500 13 807.0 1260.0 641.5 914.6 13 12,153 18,900 10,693 15,400 14 919.0 1411.3 708.4 1008.5 15 0.0 0.0 0.0 0.0 14 1,041 1,618 1,055 1,516 16 20.5 31.9 20.8 29.9 15 814 1,285 813 1,173 CHAPTER 03 3.7-42 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-9 (Cont'd)

SHEAR FORCES MOMENTS (kips) (103k-ft)

UNCRACKED CRACKED UNCRACKED CRACKED CONDITION CONDITION CONDITION CONDITION JOINT MEMBER NO. NO. OBE SSE OBE SSE OBE SSE OBE SSE 17 36.6 57.1 36.9 53.0 16 616 995 569 829 18 41.8/ 65.4/ 41.9/ 60.1/

14.5 29.0 23.8 34.6 17 1,565 2,440 1,394 2,000 29 26.8 41.8 30.9 44.6 18 1,565 2,440 1,394 2,000 19 1,754 2,743 1,621 2,329 30 55.0/ 84.6/ 47.6/ 74.3/

46.9 72.3 37.4 68.3 20 431 700 1,353 1,978 20 42.5 65.4 55.9 80.0 21 485 765 1,441 2,076 21 39.0 60.0 65.1 93.5 22 685 1,085 1,630 2,351 22 34.6 53.0 78.5 112.0 23 1,041 1,688 2,041 3,014 31 44.4 68.5 96.1 137.5 23 0.0 0.0 0.0 0.0 25 560 876 445 644 32 1.8 2.8 1.4 2.0 26 159 268 139 209 24 2.4 4.1 2.0 3.1 27 425 680 338 490 25 9.9 16.1 8.0 11.9 28 1,168 1,805 996 1,426 26 26.6 41.5 20.9 30.1 29 1,889 2,950 1,676 2,408 27 37.4 57.5 31.5 45.1 30 1,899 2,950 1,676 2,408 34 49.4 76.1 42.1 60.3 31 1,899 2,950 1,676 2,408 28 52.6 81.1 45.0 64.4 10 1.9 2.9 2.8 3.9 32 968 1,484 708 1,009 30 8.1 12.4 4.3 6.3 33 95,713 146,875 79,175 112,675 CHAPTER 03 3.7-43 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-10 PRIMARY CONTAINMENT AXIAL FORCES AXIAL FORCE (kips)

MEMBER NO. UNCRACKED CONDITION CRACKED CONDITION OBE SSE OBE SSE 1 40 61 76 114 2 239 370 461 690 3 684 1063 1273 1889 4 1196 1854 2086 3063 5 1698 2626 2793 4064 6 2196 3393 3426 4953 7 2635 4064 3909 5615 8 3060 4711 4326 6186 9 3430 5273 4706 6735 10 4245 6513 5368 7680 11 4865 7450 5810 8304 12 5443 8319 6286 9003 13 5900 9003 6528 9345 14 73 113 66 96 15 219 340 240 291 16 374 579 341 495 17 1491 2310 1376 1998 18 1491 2310 1376 1998 19 1641 2539 1506 2181 20 2123 3269 1918 2755 21 2256 3474 2015 2889 22 2381 3661 2093 2944 23 2511 3854 2154 3071 24 118 185 110 164 25 331 519 311 459 26 429 670 401 590 27 988 1538 919 1341 28 988 1538 919 1341 29 988 1538 919 1341 CHAPTER 03 3.7-44 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-11 REACTOR ENCLOSURE AND CONTROL STRUCTURE E-W INERTIA FORCES, DISPLACEMENTS, AND ACCELERATIONS INERTIA FORCES DISPLACEMENTS ACCELERATIONS (k) (10-2 ft) (g)

MASS POINT OBE SSE OBE SSE OBE SSE 1 8,840 14,393 0.412 0.637 0.184 0.300 2 6,681 10,574 0.736 1.140 0.250 0.396 3 10,605 16,776 0.957 1.470 0.290 0.459 4 1,446 2,269 1.260 1.940 0.314 0.493 5 11,955 18,851 1.450 2.240 0.341 0.538 6 1,660 2,600 1.650 2.540 0.332 0.520 7 11,794 18,564 1.830 2.810 0.350 0.551 8 1,718 2,689 2.060 3.180 0.336 0.526 9 21,634 33,907 2.160 3.330 0.334 0.524 10 2,729 4,250 2.290 3.520 0.338 0.527 11 25,796 40,472 2.440 3.760 0.395 0.620 12 17,795 27,783 3.020 4.650 0.658 1.027 CHAPTER 03 3.7-45 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-12 REACTOR ENCLOSURE AND CONTROL STRUCTURE N-S INERTIA FORCES, DISPLACEMENTS, AND ACCELERATIONS INERTIA FORCES DISPLACEMENTS ACCELERATIONS (k) (10-2 ft) (g)

MASS POINT OBE SSE OBE SSE OBE SSE 1 8,354 13,645 0.419 0.647 0.174 0.284 2 6,791 10,735 0.932 1.440 0.302 0.475 3 11,026 17,361 1.300 2.000 0.302 0.475 4 1,475 2,311 1.800 2.760 0.320 0.502 5 12,320 19,355 2.120 3.270 0.351 0.552 6 1,809 2,841 2.470 3.800 0.361 0.568 7 12,469 19,661 2.770 4.270 0.370 0.584 8 1,892 2,945 3.210 4.940 0.370 0.576 9 26,060 40,829 3.410 5.250 0.403 0.631 10 2,996 4,638 3.740 5.750 0.371 0.575 11 29,068 45,167 4.140 6.370 0.445 0.692 12 20,473 31,545 5.510 8.470 0.757 1.166 CHAPTER 03 3.7-46 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-13 REACTOR ENCLOSURE AND CONTROL STRUCTURE VERTICAL INERTIA FORCES, DISPLACEMENTS, AND ACCELERATIONS INERTIA FORCES DISPLACEMENTS ACCELERATIONS (k) (10-2 ft) (g)

MASS POINT OBE SSE OBE SSE OBE SSE 1 6,320 10,050 0.232 0.359 0.133 0.211 2 2,961 4,682 0.264 0.409 0.147 0.232 3 2,765 4,355 0.286 0.443 0.154 0.243 4 534 838 0.317 0.491 0.166 0.260 5 3,315 5,199 0.336 0.521 0.172 0.271 6 446 697 0.356 0.552 0.178 0.278 7 2,804 4,383 0.373 0.578 0.184 0.287 8 632 990 0.399 0.618 0.196 0.308 9 2,943 4,616 0.409 0.634 0.202 0.316 10 928 1,457 0.426 0.660 0.211 0.332 11 4,760 7,486 0.439 0.681 0.219 0.345 12 3,242 5,167 0.462 0.717 0.248 0.396 13 675 1,077 0.272 0.422 0.149 0.238 14 687 1,098 0.316 0.489 0.164 0.262 15 1,035 1,649 0.541 0.838 0.238 0.380 16 606 952 0.751 1.161 0.287 0.451 17 2,204 3,460 0.988 1.529 0.448 0.704 18 73 114 1.109 1.715 0.503 0.787 19 38 60 1.228 1.898 0.589 0.918 20 3,933 6,153 0.985 1.524 0.378 0.591 21 2,911 4,581 1.249 1.932 0.496 0.781 22 4,743 7,407 0.835 1.292 0.298 0.464 23 4,759 7,431 0.840 1.299 0.297 0.465 24 2,917 4,589 1.255 1.941 0.497 0.782 25 3,948 6,176 0.989 1.530 0.379 0.594 26 2,899 4,476 3.560 5.495 0.836 1.291 27 6,994 10,800 2.693 4.157 0.689 1.064 28 1,925 2,982 0.645 0.997 0.550 0.852 29 4,077 6,334 0.585 0.906 0.471 0.732 30 2,980 4,611 0.787 1.216 0.534 0.827 31 2,906 4,498 0.749 1.158 0.387 0.599 32 1,001 1,548 0.956 1.477 0.551 0.852 33 2,312 3,573 1.112 1.717 0.653 1.009 34 3,908 6,057 1.226 1.895 0.774 1.154 35 717 1,120 1.332 2.062 0.725 1.132 36 1,363 2,106 1.900 2.934 0.703 1.087 37 2,120 3,305 1.406 2.175 0.696 1.085 38 488 756 0.995 1.538 0.334 0.517 39 909 1,419 0.910 1.408 0.351 0.548 40 1,212 1,875 0.777 1.201 0.677 1.047 41 956 1,477 2.231 3.445 0.276 0.427 42 1,283 1,981 2.488 3.842 0.649 1.001 43 967 1,494 1.981 3.058 0.468 0.723 CHAPTER 03 3.7-47 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-13 (Cont'd)

INERTIA FORCES DISPLACEMENTS ACCELERATIONS (k) (10-2 ft) (g)

MASS POINT OBE SSE OBE SSE OBE SSE 44 3,060 4,723 5.016 7.743 1.062 1.639 45 1,594 2,460 2.356 3.637 0.633 0.978 46 1,680 2,593 3.501 5.405 1.028 1.587 47 3,144 4,854 3.105 4.794 0.712 1.098 48 893 1,379 2.489 3.844 0.375 0.580 49 490 759 0.999 1.544 0.335 0.519 50 899 1,403 0.905 1.401 0.347 0.542 51 2,122 3,308 1.404 2.171 0.697 1.086 52 713 1,113 1.333 2.063 0.721 1.126 53 2,510 4,020 3.994 6.418 0.215 0.344 54 455 720 0.842 1.317 0.958 1.516 CHAPTER 03 3.7-48 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-14 REACTOR ENCLOSURE AND CONTROL STRUCTURE E-W SHEAR FORCES AND MOMENTS SHEAR FORCE MOMENTS (102 k) (103 k-ft)

JOINT MEMBER NO. NO. OBE SSE OBE SSE 1 11,040 17,000 1 834 1,288 2 9,486 14,600 2 779 1,201 3 8,439 12,990 3 706 1,091 4 7,032 10,830 4 694 1,073 5 6,146 9,474 5 659 1,018 6 5,162 7,966 6 655 1,012 7 4,306 6,654 7 608 941 8 3,098 4,797 8 596 922 9 2,589 4,013 9 423 656 10 1,794 2,781 10 400 621 11 1,032 1,611 11 178 278 12 0 0 CHAPTER 03 3.7-49 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-15 REACTOR ENCLOSURE AND CONTROL STRUCTURE N-S SHEAR FORCES AND MOMENTS SHEAR FORCE MOMENTS (102 k) (103 k-ft)

JOINT MEMBER NO. NO. OBE SSE OBE SSE 10,430 16,060 1 832 1,283 2 9,106 14,020 2 780 1,204 3 8,176 12,580 3 719 1,111 4 6,835 10,520 4 712 1,100 5 6,009 9,255 5 640 988 6 5,041 7,766 6 628 969 7 4,198 6,469 7 592 915 8 3,130 4,823 8 584 902 9 2,699 4,164 9 448 693 10 1,916 2,951 10 423 655 11 1,187 1,830 11 205 316 12 0 0 CHAPTER 03 3.7-50 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-16 REACTOR ENCLOSURE AND CONTROL STRUCTURE AXIAL FORCES AXIAL FORCE (k)

MASS POINT OBE SSE 1 62,283 96,586 2 47,396 73,527 3 44,400 68,961 4 40,645 63,210 5 39,736 61,810 6 35,719 55,599 7 34,779 54,144 8 31,475 49,013 9 30,699 47,814 10 20,903 32,677 11 18,163 24,838 12 5,162 8,273 13 13,581 21,050 14 12,796 19,810 15 12,484 19,305 16 11,121 17,201 17 6,483 10,053 18 4,009 6,203 19 3,454 5,354 CHAPTER 03 3.7-51 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-17 STRUCTURE-FOUNDATION INTERACTION COEFFICIENTS EQUIVALENT SPRING EQUIVALENT DAMPING STRUCTURE MOTION CONSTANT COEFFICIENT Primary Translational 4.15x107 k/ft 2.01x105 k-sec/ft containment Rocking 8.12x1010 k-ft/rad 7.82x107 k-ft-sec/rad Vertical 4.87x107 k/ft 3.48x105 k-sec/ft Reactor enclosure Translational: E-W 8.17x107 k/ft 8.98x105 k-sec/ft and control N-S 8.55x107 k/ft 8.79x105 k-sec/ft structure Rocking: E-W 2.04x1012 k-ft/rad 9.23x109 k-ft-sec/rad N-S 6.22x1011 k-ft/rad 2.63x109 k-ft-sec/rad Vertical 9.90x107 k/ft 1.74x106 k-sec/ft CHAPTER 03 3.7-52 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-18 NUMBER OF DYNAMIC RESPONSE CYCLES EXPECTED DURING A SEISMIC EVENT FOR NSSS SYSTEMS AND COMPONENTS FREQUENCY BAND WIDTH (Hz) 0-10 10-20 20-50 Total number of seismic cycles 168 359 643 Number of seismic cycles (0.5% of total) between 75%

and 100% of peak load 0.8 1.8 3.2 Number of seismic cycles (4.5% of total) between 50%

and 75% of peak load 7.5 16.2 28.9 CHAPTER 03 3.7-53 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-19 FATIGUE EVALUATION DUE TO SEISMIC LOAD CALCULATED NO. OF DESIGN NO. OF PEAK CYCLES AT PEAK STRESS CYCLES PER COMPONENT STRESS OBE

1. Reactor Pressure Vessel

a. Vessel See Table 3.7-18 10

b. Shroud support See Table 3.7-18 10

c. Skirt See Table 3.7-18 10

2. Seismic Category I Piping

a. Recirculation lines (1) 10

b. Steam lines (1) 10 (1)

Design number of peak stress cycles used in analysis CHAPTER 03 3.7-54 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-20 REACTOR ENCLOSURE AND CONTROL STRUCTURE - SEISMIC RESPONSE COMPARISON OF 2-COMPONENT ABS VERSUS 3-COMPONENT SRSS METHODS Combined Axial and Bending Stresses on Concrete Total Seismic Stress (KSF)

Wall Section Due to: (Axial and Bending Stress on Concrete Wall Section)

SSE Vertical SSE N-S SSE E-W Excitation Excitation Excitation (P) (Mc) (Mc) 3 Component SRSS 2 Component ABS Elevation (A) (KSF) I N-S (KSF) I E-W (KSF) (Regulatory Guide 1.92) (Section 3.7.2.6)  % Difference Reactor Enclosure Wall - Southwest Corner 352 12.27 23.4 10.2 28.3 35.7 +26 333 0.92 18.6 13.4 22.9 19.5 -15 313 0.88 26.4 19.5 32.8 27.3 -17 304 13.85 26.5 22.1 37.2 40.4 +9 283 14.19 37.1 30.6 50.1 51.3 +2 269 14.56 43.4 33.0 56.4 58.0 +3 253 14.92 51.7 39.7 66.9 66.6 -

239 16.11 59.4 45.7 76.7 75.5 -2 217 16.44 71.4 55.2 91.7 87.8 -4 201 15.18 57.2 55.9 81.4 72.4 -11 177 14.35 63.1 64.2 91.2 78.6 -14 Control Structure - Northeast Corner 352 12.27 25.4 3.9 28.5 37.7 +32 333 0.92 20.1 5.1 20.8 21.0 +1 313 0.88 29.1 7.5 30.1 30.0 -

304 13.85 29.8 8.5 33.9 43.7 +29 283 14.19 41.8 11.7 45.7 56.0 +23 269 14.56 45.6 12.6 49.5 60.2 +22 253 14.92 54.3 15.2 58.3 69.2 +19 239 16.11 61.9 17.5 66.3 78.0 +18 217 16.44 74.4 21.2 79.1 90.8 +15 201 15.18 74.7 21.4 79.2 89.9 +14 177 14.35 73.7 24.6 79.0 88.1 +12 CHAPTER 03 3.7-55 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-21 SPRAY POND PUMPHOUSE(1)

FREQUENCIES WITH AND WITHOUT ECCENTRICITIES GLOBAL DIRECTION: E-W VIBRATION FREQUENCIES (CPS)

MODE 3-D MODEL ORIGINAL MODEL NO. WITH ECCENTRICITY WITHOUT ECCENTRICITY 1 17.8 18.2 2 41.8 42.3 (1)

Figure 3.7-48 CHAPTER 03 3.7-56 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-22 DIESEL GENERATOR ENCLOSURE(1)

FREQUENCIES WITH AND WITHOUT ECCENTRICITIES GLOBAL DIRECTION: E-W VIBRATION FREQUENCIES (CPS)

ORIGINAL MODEL 3-D MODEL MODE WITHOUT ECCENTRICITY WITH ECCENTRICITY 1 9.33 8.84 2 22.59 20.13 3 48.05 37.50 (1)

Figure 3.7-49 CHAPTER 03 3.7-57 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-23 REACTOR ENCLOSURE(1)

FREQUENCIES WITH AND WITHOUT ECCENTRICITIES GLOBAL DIRECTION: E-W VIBRATION FREQUENCIES (CPS)

MODE 3-D MODEL ORIGINAL MODEL NO. WITH ECCENTRICITY WITHOUT ECCENTRICITY 1 3.738 3.769 2 11.397 11.863 3 16.587 17.944 4 21.685 23.007 5 24.906 30.083 (1)

Figure 3.7-50 CHAPTER 03 3.7-58 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-24 SPRAY POND PUMPHOUSE COMPARISON OF TORSIONAL MOMENTS FROM ORIGINAL DESIGN WITH VALUES OBTAINED FROM 3-D MODEL WITH ECCENTRICITY TORSIONAL MOMENT (MT, K-FT)

DIRECTION EARTHQUAKE ORIGINAL DESIGN 3-D STICK MODEL E-W OBE 15,370 7,724 DBE 27,620 14,230 CHAPTER 03 3.7-59 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-25 DIESEL GENERATOR ENCLOSURE COMPARISON OF TORSIONAL MOMENTS FROM ORIGINAL DESIGN WITH VALUES OBTAINED FROM 3-D MODEL WITH ECCENTRICITY TORSIONAL MOMENT (MT, K-FT)

DIRECTION EARTHQUAKE ORIGINAL DESIGN 3-D STICK MODEL E-W OBE 96,804 74,940 DBE 149,988 101,200 CHAPTER 03 3.7-60 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-26 REACTOR ENCLOSURE COMPARISON OF TORSIONAL MOMENTS FROM ORIGINAL DESIGN WITH VALUES OBTAINED FROM 3-D MODEL WITH ECCENTRICITY TORSIONAL MOMENT (MT, K-FT)

DIRECTION EARTHQUAKE ORIGINAL DESIGN 3-D STICK MODEL E-W DBE 3.5X106 2.36X106 (SSE)

CHAPTER 03 3.7-61 REV. 13, SEPTEMBER 2006

LGS UFSAR Table 3.7-27 REACTOR ENCLOSURE COMPARISON OF DESIGN SHEARS AND SHEARS OBTAINED FROM 3-D MODEL WITH ECCENTRICITY SEISMIC EVENT - DBE WALLS AT EL 177 FT (GROUND FLOOR)

SHEAR FROM 3-D MODEL WITH DESIGN SHEAR ECCENTRICITY WALL(1) (KIPS) (KIPS)

Line J 44,171 37,800 Line D 75,385 64,193 Line ML 26,359 22,099 (1)

Figure 3.7-50 CHAPTER 03 3.7-62 REV. 13, SEPTEMBER 2006

LGS UFSAR TABLE 3.7-28 SEISMIC MONITORING INSTRUMENTATION MINIMUM MEASUREMENT INSTRUMENTS INSTRUMENTATION AND SENSOR LOCATIONS RANGE OPERABLE

1. Triaxial Time-History Accelerometers (T/As)

a. Sensors /Recorders

1) XE-VA-131/XR-VA-131 Free Field +/- 4 g 1 (East of PPC)

2) XE-VA-132/XR-VA-132 Reactor Enclosure +/- 4 g 1 Foundation (Loc. 111-R11-177)

3) XE-VA-133/XR-VA-133 Reactor Enclosure +/- 4 g 1 Elevation "A" (Loc. 304-R 11-217)

4) XE-VA-134/XR-VA-134 Reactor Enclosure +/- 4 g 1 Elevation "B" (Loc. 506-R15-283)

5) XE-VA-135/XR-VA-135 Foundation of an +/- 4 g 1 Independent Seismic Category I Structure (Spray Pond Pump House, El 237')

6) XE-VA-136/XR-VA-136 Independent +/- 4 g 1 Seismic Cat. 1 Structure Elevation (Spray Pond Pump House, El 268')

2. Network Control Center/ 0.1 - 100 Hz 1*

Earthquake Analysis PC

• With reactor control room indication and annunciation. Receives signal from Strong Motion Recorders.

CHAPTER 03 3.7-63 REV. 19, SEPTEMBER 2018

LGS UFSAR TABLE 3.7-29 SEISMIC MONITORING INSTRUMENTATION SURVEILLANCE REOUIREMENTS CHANNEL**

CHANNEL* FUNCTIONAL CHANNEL***

INSTRUMENTATION AND SENSOR LOCATIONS CHECK TEST CALIBRATION

1. Triaxial Time-History Accelerometers (T/As)

a. Sensors /Recorders

1) XE-VA-131/XR-VA-131 Free M SA R Field (East of PPC)

2) XE-VA-132/XR-VA-132 Reactor M SA R Enclosure Foundation (Loc. 111-R 11-177)

3) XE-VA-133/XR-VA-133 Reactor M SA R Enclosure Elevation "A" (Loc. 304-R11-217)

4) XE-VA-134/XR-VA-134 Reactor M SA R Enclosure Elevation "B" (Loc. 506-R 15-283)

5) XE-VA-135/XR-VA-135 Foundation M SA R of an Independent Seismic Category 1 Structure (Spray Pond Pump House, El 237')

6) XE-VA-136/XR-VA-136 Independent M SA R Seismic Category I Structure (Spray Pond Pump House, El 268')

2. Network Control Center/

Earthquake Analysis PC M SA R

• Channel Check - The qualitative verification of the functional status of the instrument. This check is an "in situ" test and may be the same as a channel functional test.

• • Channel Functional Test- (Secondary Calibration). The determination without adjustment that an instrument, sensor, or system responds to a known input of such character that it will verify the instrument, sensor, or system is functioning in a manner that can be calibrated.

• • • Channel Calibration - (Primary Calibration). The determination and, if required, adjustment of an instrument, sensor, or system such that it responds within a specific range and accuracy to an acceleration, velocity, or displacement input, as applicable, or responds to an acceptable physical constant.

CHAPTER 03 3.7-64 REV. 19, SEPTEMBER 2018

LGS UFSAR TABLE 3.7-29 (Contd)

SEISMIC MONITORING INSTRUMENTATION SURVEILLANCE REOUIREMENTS Surveillance Frequency Notation Notation Frequency M At least once per 31 days SA At least once per 184 days R (Refueling Interval) At least once per 24 months (731 days)

CHAPTER 03 3.7-65 REV. 19, SEPTEMBER 2018