NL-17-0534, Joseph M. Farley Nuclear Plant, Updated Final Safety Analysis Report, Revision 27, Chapter 4 Through Chapter 5
Text
FNP-FSAR-4 TABLE 4.1-3 DESIGN LOADING CONDITIONS FOR REACTOR CORE COMPONENTS
REV 21 5/08 1. Fuel assembly weight
- 2. Fuel assembly spring forces
- 3. Internals weight
- 4. Control rod scram (equivalent static load)
- 5. Differential pressure
- 6. Spring preloads
- 7. Coolant flow forces (static)
- 8. Temperature gradients
- 9. Differences in thermal expansion
- a. Because of temperature differences b. Because of expansion of different materials
- 10. Interference between components
- 11. Vibration (mechanically or hydraulically induced)
- 12. One or more loops out of service
- 13. All operational transients listed in table 5.2-2
- 14. Pump overspeed
- 15. Seismic loads (operation basis earthquake and design basis earthquake)
- 16. Blowdown forces (due to cold and hot leg break)
FNP-FSAR-4 TABLE 4.2-1 MAXIMUM DEFLECTIONS SPECIFIED FOR REACTOR INTERNAL SUPPORT STRUCTURES
REV 21 5/08 Component Allowable Deflections (in.)
No Loss of Function Deflections (in.)
Upper Barrel radial inward 4.38 8.77 radial outward 0.5 1.0 Upper Package 0.1 0.15 Rod Cluster Guide tubes 1.0 1.75 REV 21 5/08 FUEL ASSEMBLY OUTLINE 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-1 (SHEET 1 OF 2)
REV 21 5/08 FUEL ASSEMBLY OUTLINE 17 X 17 VANTAGE 5 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-1 (SHEET 2 OF 2)
REV 21 5/08 PRE UNIT 2 CYCLE 3 AND UNIT 1 CYCLE 6 FUEL ASSEMBLY CROSS-SECTION 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 1 OF 8)
REV 21 5/08 UNIT 2 CYCLES 3-5, UNIT 1 CYCLES 6-8 FUEL ASSEMBLY CROSS-SECTION 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 2 OF 8)
REV 21 5/08 UNIT 2 CYCLE 6,UNIT 1 CYCLE 9 FUEL ASSEMBLY CROSS-SECTION 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 3 OF 8)
REV 21 5/08 UNIT 2 CYCLE 7, UNIT 1 CYCLE 10 FUEL ASSEMBLY CROSS-SECTION 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 4 OF 8)
REV 21 5/08 UNIT 2 CYCLE 8, UNIT 1 CYCLE 11 FUEL ASSEMBLY CROSS-SECTION 17 X 17 LOPAR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 5 OF 8)
REV 21 5/08 UNIT 2 CYCLE 9, 10, 11, 12, 13, UNIT 1 CYCLE 12, 13, 14, 15 FUEL ASSEMBLY 17 X 17 VANTAGE 5 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 6 OF 8)
REV 21 5/08 UNIT 1 CYCLE 16, 17 AND UNIT 2 CYCLE 14 FUEL ASSEMBLY OUTLINE 17 X 17 VANTAGE+ W/PROTECTIVE GRID ZIRLO FUEL RODS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 7 OF 8)
REV 21 5/08 UNIT 2 CYCLE 15 AND UNIT 1 CYCLE 18 AND AFTER FUEL ASSEMBLY OUTLINE 17 X 17 VANTAGE+ WITH LOW PRESSURE ZIRLO FUEL RODS, AND LOW FORCE HOLDDOWN SPRING JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-2 (SHEET 8 OF 8)
REV 21 5/08 PRE UNIT 2 CYCLE 3 AND UNIT 1 CYCLE 6 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 1 OF 8)
REV 21 5/08 UNIT 2 CYCLE 3, 4, 5 AND UNIT 1 CYCLE 6, 7, AND 8 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 2 OF 8)
REV 21 5/08 UNIT 2 CYCLE 6 AND UNIT 1 CYCLE 9 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 3 OF 8)
REV 21 5/08 UNIT 2 CYCLE 7, UNIT 1 CYCLE 10 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 4 OF 8)
REV 21 5/08 UNIT 2 CYCLE 8, UNIT 1 CYCLE 11 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 5 OF 8)
REV 21 5/08 UNIT 2 CYCLE 9, 10, 11, 12, 13 AND UNIT 1 CYCLE 12, 13, 14, 15 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 6 OF 8)
REV 21 5/08 UNIT 1 CYCLE 16, 17 AND UNIT 2 CYCLE 14 FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 7 OF 8)
REV 21 5/08 UNIT 2 CYCLE 15 AND UNIT 1 CYCLE 18 AND AFTER FUEL ROD SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-3 (SHEET 8 OF 8)
REV 21 5/08 GRID PLAN VIEW JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-4
REV 21 5/08 PRE UNIT 2 CYCLE 6 AND UNIT 1 CYCLE 9 TOP GRID TO NOZZLE ATTACHMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-5 (SHEET 1 OF 3)
REV 21 5/08 UNIT 2 CYCLE 6, 7, AND 8, UNIT 1 CYCLE, 9, 10, AND 11 TOP GRID TO NOZZLE ATTACHMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2.-5 (SHEET 2 OF 3)
REV 21 5/08 UNIT 2 CYCLE 9, UNIT 1 CYCLE 12 AND LATER TOP GRID-TO-NOZZLE ATTACHMENT DETAIL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-5 (SHEET 3 OF 3)
REV 21 5/08 ELEVATION VIEW, GRID-TO-THIMBLE ATTACHMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-6 (SHEET 1 OF 2)
REV 21 5/08 ELEVATION VIEW, GRID-TO-THIMBLE ATTACHMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-6 (SHEET 2 OF 2)
REV 21 5/08 GUIDE THIMBLE TO BOTTOM NOZZLE JOINT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-7
REV 21 5/08 TYPICAL CLAD AND PELLET DIMENSIONS AS A FUNCTION OF EXPOSURE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-8
REV 21 5/08 REPRESENTATIVE FUEL ROD INTERNAL PRESSURE AND LINEAR POWER DENSITY FOR THE LEAD BURNUP ROD AS A FUNCTION OF TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-9
REV 21 5/08 LOWER CORE SUPPORT ASSEMBLY (CORE BARREL ASSEMBLY)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-10
REV 21 5/08 UPPER CORE SUPPORT ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-11
REV 21 5/08 PLAN VIEW OF UPPER CORE SUPPORT STRUCTURE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-12
REV 21 5/08 FULL-LENGTH ROD CLUSTER CONTROL AND DRIVE ROD ASSEMBLY WITH INTERFACING COMPONENTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-13
REV 21 5/08 FULL-LENGTH ROD CLUSTER CONTROL ASSEMBLY OUTLINE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-14
REV 21 5/08 FULL LENGTH ABSORBER ROD JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-15
REV 21 5/08 BURNABLE ABSORBER ASSEMBLY (STANDARD BOROSILICATE GLASS)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-16 (SHEET 1 OF 3)
REV 21 5/08 UNIT 1 CYCLES 8 AND 9, UNIT 2 CYCLES 6 AND 7 BURNABLE ABSORBER ASSEMBLY (WET ANNULAR)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-16 (SHEET 2 OF 3)
REV 21 5/08 UNIT 1 CYCLE 10 UNIT 2 CYCLE 8 AND AFTER ABSORBER ASSEMBLY (WET ANNULAR)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-16 (SHEET 3 OF 3)
REV 21 5/08 BURNABLE ABSORBER ROD (STANDARD BOROSILICATE GLASS)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-17 (SHEET 1 OF 2)
REV 21 5/08 BURNABLE ABSORBER ROD (WET ANNULAR)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-17 (SHEET 2 OF 2)
REV 21 5/08 PRIMARY SOURCE ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-18
REV 21 5/08 SECONDARY SOURCE ASSEMBLY FOR UNIT 1 CYCLES 1 AND 2 ONLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-19A
REV 21 5/08 SECONDARY SOURCE ASSEMBLY FOR UNIT 1 CYCLES 2 TO 12 AND UNIT 2 CYCLES 1 TO 9 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-19B
REV 21 5/08 DOUBLE ENCAPSULATED SECONDARY SOURCE ASSEMBLY FOR UNIT 2 CYCLE 9 AND AFTER AND UNIT 1 CYCLE 12 AND AFTER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-19C
REV 21 5/08 THIMBLE PLUG ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-20 (SHEET 1 OF 2)
REV 21 5/08 STANDARDIZED THIMBLE PLUG ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-20 (SHEET 2 OF 2)
REV 21 5/08 FULL-LENGTH CONTROL ROD DRIVE MECHANISM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-21
REV 21 5/08 FULL-LENGTH CONTROL ROD DRIVE MECHANISM SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-22
REV 21 5/08 NOMINAL LATCH CLEARANCE AT MINIMUM AND MAXIMUM TEMPERATURE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-23
REV 21 5/08 CONTROL ROD DRIVE MECHANISM LATCH CLEARANCE THERMAL EFFECT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-24
REV 21 5/08 REMOVABLE ROD COMPARED TO STANDARD ROD JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-25
REV 21 5/08 REMOVABLE FUEL ROD ASSEMBLY OUTLINE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-26
REV 21 5/08 LOCATION OF REMOVABLE RODS WITHIN AN ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-27
REV 21 5/08 SCHEMATIC REPRESENTATION OF REACTOR CORE MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.2-28
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-2 (SHEET 1 OF 2)
NUCLEAR DESIGN PARAMETERS (FIRST CYCLE)
REV 21 5/08 Core Average Linear Power, kW/ft, including Densification Effects 5.20 Total Heat Flux Hot Channel Factor, F Q 2.32 Nuclear Enthalpy Rise Hot Channel 1.55 Factor, N H F Reactivity Coefficients Design Limits Best Estimate Doppler-only power, Coefficients, pcm/% power(b) (upper limit) -19.4 to -12.6 -12.2 to -8.1 Lower limit -10.2 to -6.7 -11.8 to -7.9 Doppler temperature coefficient (pcm/°F)(b) -2.9 to -1.4 -2.2 to -1.4 Moderator temperature coefficient (pcm/°F)(b) 0 -1 to -40 Boron coefficient (pcm/ppm)(b) -16 to -8
-13 to -9 Rodded moderator density (pcm/g/cm 3)(b) 0.43 x 10 5 .33 x 10 5 Delayed Neutron Fraction and Lifetime eff, BOL, (EOL) 0.0075, (0.0048) , BOL, (EOL)
µs 19.9 (18.1) Control Rods See table 4.3-3 Rod requirements Maximum bank worth, pcm(b) < 2300 Maximum ejected rod wo rth See chapter 15 Radial Factor (BOL to EOL)
Unrodded 1.37 to 1.25 D bank 1.58 to 1.42 D + C 1.63 to 1.42 D + C + B 1.80 to 1.55
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-2 (SHEET 2 OF 2)
Design Limits Best Estimate REV 21 5/08 Boron Concentrations Zero power, Keff = 1.00 cold rod cluster control assemblies out (1% uncertainty included)(a) 1429 Zero power, Keff = 1.00 hot rod cluster control assemblies out (1% uncertainty included)(a) 1419 Design basis refueling boron concentration 2000 Zero power, Keff 0.95, cold rod cluster control assemblies in (1% uncertainty included)(a) 1196 Full power, no xenon, Keff = 1.0, hot rod cluster control assemblies out 1195 Full power, equilibrium xenon, Keff = 1.0, hot rod cluster control assemblies out 906 Reduction with fuel burnup First cycle (ppm/GWd/Mtu)(c) ~60 Reload cycle (ppm/GWd/Mtu)(c) ~85
________________________
- a. Uncertainties are given in paragraph 4.3.3.3.
- b. 1 pcm = (percent milli rho) = 10
-5 where is calculated from two statepoint values of Keff by ln (k 2/k 1). c. Gigawatt day (GWd) = 1000 megawatt day (1000 MWd). During the first cycle, fixed BA rods are present which significantly reduce the boron depletion rate compared to reload cycles.
]
FNP-FSAR-4 TABLE 4.3-3 REACTIVITY REQUIREMENTS FOR ROD CLUSTER CONTROL ASSEMBLIES
REV 21 5/08
[HISTORICAL]
[BOL EOL Reactivity Effects, Percent (First Cycle)
End of Life (Equilibrium Cycle)
Control requirements Fuel temperature (Doppler)(% ) 1.26 1.05 1.11 Moderator temperature (%) 0.23 1.07 1.20 Void (% ) 0.05 0.05 0.05 Redistribution (% ) 0.50 0.85 1.00 Rod Insertion Allowance (% ) 0.50 0.50 0.50 (1) Total control (% ) 2.54 3.52 3.86 Estimated rod cluster control assembly worth
(48 rods)
- a. All full-length assemblies inserted (% ) 9.88 9.57 8.50 b. All but one (highest worth) assemblies inserted (% ) 7.85 7.81 7.65 (2) Estimated rod cluster control assembly credit with 10-percent adjustment to accommodate uncertainties (3 to 10 percent) (% ) 7.06 7.03 6.88 Shutdown margin available (2-1) (% ) 4.52 3.51] 3.02 (a)
_________________ a. The design basis minimum shutdown is 1.77 percent.
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-4 AXIAL STABILITY INDEX PWR CORE WITH A 12-FT HEIGHT REV 21 5/08 Stability Index (hr-1) Burnup (MWD/T)
F z C B (ppm) Exp Calc 1550 1.34 1065 -0.041 -0.032
7700 1.27 700 -0.014 -0.006
5090(a) -0.0325 -0.0255 Radial Stability Index
2250(b) -0.068 0.07
__________________ a. 4-loop plant, 12-foot core in cycle 1, axial stability test.
- b. 4-loop plant, 12-foot core in cycle 1, radial (X-Y) stability test.
]
FNP-FSAR-4 TABLE 4.3-5 TYPICAL NEUTRON FLUX LEVELS (n/cm 2 -s) AT FULL POWER
REV 21 5/08 E > 1.0 Mev 5.53 Kev < E 01.0 Mev 0.625 ev E <5.53 Kev E < .625 ev (hardened spectrum)
Core center 6.51 x 10 13 1.12 x 10 14 8.50 x 10 13 3.00 x 10 13 Core outer radius at
midheight 3.23 x 10 13 5.74 x 10 13 4.63 x 10 13 8.60 x 10 12 Core top, on axis 1.53 x 10 13 2.42 x 10 13 2.10 x 10 13 1.63 x 10 12 Core bottom, on axis 2.36 x 10 13 3.94 x 10 13 3.50 x 10 13 1.46 x 10 13 Pressure vessel inner
wall, azimuthal peak, core midheight 2.77 x 10 10 5.75 x 10 10 6.03 x 10 10 8.38 x 10 10 FNP-FSAR-4 TABLE 4.3-6 COMPARISON OF MEASURED AND CALCULATED DOPPLER DEFECTS
REV 21 5/08 Plant Fuel Type Core Burnup (MWD/MTU) Measured (pcm) Calculated (pcm)(a) 1 Air and helium-filled 8460 1200 1210 2 Helium-filled 0 1130 1220 3 Helium-filled 0 1180 1220
_________________
- a. 2 1 5 k kln10pcmx=
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-7 (SHEET 1 OF 2)
BENCHMARK CRITICAL EXPERIMENT S (26,34,35) LEOPARD COMPARISONS REV 21 5/08 Description of Experiments(a) No. of Experiments LEOPARD K eff Using Experimental Bucklings UO 2 Al clad 14 1.0012 SS clad 19 0.9963 Borated H 2O 7 0.9989 Total 40 0.9985 U-Metal Al clad 43 0.9995 Unclad 20 0.9990 Total 61 0.9993 All above 101 0.9990
____________________ a. Reported in reference 25
.
FNP-FSAR-4 TABLE 4.3-7 (SHEET 2 OF 2)
AMPX - KENO COMPARISONS REV 21 5/08 General Description Enrichment w/o U235 Reflector Separating Material Characterizing Separation (cm)
- 1. UO 2 rod lattice 2.35 water water 11.92 2. UO 2 rod lattice 2.35 water water 8.39 3. UO 2 rod lattice 2.35 water water 6.39 4. UO 2 rod lattice 2.35 water water 4.46 5. UO 2 rod lattice 2.35 water Stainless steel 10.44 6. UO 2 rod lattice 2.35 water Stainless steel 11.47 7. UO 2 rod lattice 2.35 water Stainless steel 7.76 8. UO 2 rod lattice 2.35 water Stainless steel 7.42 9. UO 2 rod lattice 2.35 water boral 6.34 10. UO 2 rod lattice 2.35 water boral 9.03 11. UO 2 rod lattice 2.35 water boral 5.05 12. UO 2 rod lattice 4.29 water water 10.64 13. UO 2 rod lattice 4.29 water Stainless steel 9.76 14. UO 2 rod lattice 4.29 water Stainless steel 8.08 15. UO 2 rod lattice 4.29 water boral 6.72 16. U metal cylinders 93.2 bare air 15.43 17. U metal cylinders 93.2 paraffin air 23.84 18. U metal cylinders 93.2 bare air 19.97 19. U metal cylinders 93.2 paraffin air 36.47 20. U metal cylinders 93.2 bare air 13.74 21. U metal cylinders 93.2 paraffin air 23.48 22. U metal cylinders 93.2 bare plexiglas 15.74 23. U metal cylinders 93.2 paraffin plexiglas 24.43 24. U metal cylinders 93.2 bare plexiglas 21.74 25. U metal cylinders 93.2 paraffin plexiglas 27.94 26. U metal cylinders 93.2 bare steel 14.74 27. U metal cylinders 93.2 bare plexiglas, steel 16.67]
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-8 SAXTON CORE II ISOTOPICS ROD MY+, AXIAL ZONE 6
REV 21 5/08 Atom Ratio Measured (a) 2 Precision (%)
LEOPARD Calculation
U-234/U 4.65 x 10-5 +/-29 4.60 x 10-5 U-235/U 5.74 x 10-3 +/-0.9 5.73 x 10-3 U-236/U 3.55 x 10-4 +/-5.6 3.74 x 10-4 U-238/U 0.99386
+/-0.01 0.99385 Pu-238/Pu 1.32 x 10-3 +/-2.3 1.222 x 10
-3 Pu-239/Pu 0.73971
+/-0.03 0.74497 Pu-240/Pu 0.19302
+/-0.2 0.19102 Pu-241/Pu 6.014 x 10
-2 +/-0.3 5.74 x 10-2 Pu-242/Pu 5.81 x 10
-3 +/-0.9 5.38 x 10-3 Pu/U (b) 5.938 x 10
-2 +/-0.7 5.970 x 10
-2 Np-237/U-238 1.14 x 10-4 +/-15 0.86 x 10-4 Am-241/Pu-239 1.23 x 10
-2 +/-15 1.08 x 10-2 Cm-242/Pu-239 1.05 x 10
-4 +/-10 1.11 x 10-4 Cm-244/Pu-239 1.09 x 10
-4 +/-20 0.98 x 10-4
__________________ a. Reported in reference 37.
- b. Weight ratio.
]
FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-9 CRITICAL BORON CONCENTRATIONS (ppm),HZP, BOL REV 21 5/08 Plant Type Measured Calculated 2-Loop, 121 assemblies 10-foot core 1583 1589 2-Loop, 121 assemblies 12-foot core 1625 1624 2-Loop, 121 assemblies 12-foot core 1517 1517 3-Loop, 157 assemblies 12-foot core 1169 1161 3-Loop, 157 assemblies 12-foot core 1344 1319 4-Loop, 193 assemblies 12-foot core 1370 1355 4-Loop, 193 assemblies 12-foot core
] 1321 1309 FNP-FSAR-4
[HISTORICAL] [TABLE 4.3-10 COMPARISON OF MEASURED AND CALCULATED AG-IN-CD ROD WORTH
REV 21 5/08 2-Loop Plant, 121 Assemblies, 10-foot core Measured (pcm) Calculated (pcm)
Group B 1885 1893 Group A 1530 1649 Shutdown group 3050 2917 ESADA Critical(a), 0.69" Pitch, 2 w/o Pu0 2 , 8% Pu 240 , 9 Control Rods 6.21" rod separation 2250 2250 2.07" rod separation 4220 4160 1.38" rod separation 4100 4010
__________________
- a. Reported in reference 36.
]
FNP-FSAR-4 REV 21 5/08
TABLE 4.3-11 COMPARISON OF MEASURED AND CALCULATED MODERATOR COEFFICIENTS AT HZP, BOL
Plant Type/ Control Bank Configuration Measured iso (a) (pcm/°F) Calculated iso (a) (pcm/°F) 3-loop, 157 assemblies, 12-foot core D at 160 steps - 0.50 - 0.50 D in, C at 190 steps - 3.01 - 2.75 D in, C at 28 steps - 7.67 - 7.02 B, C, and D in - 5.16 - 4.45 2-loop, 121 assemblies, 12-foot core D at 180 steps
+ 0.85 + 1.02 D in, C at 180 steps - 2.40 - 1.90 C and D in, B at 165 steps - 4.40 - 5.58 B, C, and D in, A at 174 steps - 8.70 - 8.12 4-loop, 193 assemblies, 12-foot core ARO - 0.52 - 1.2 D in - 4.35 - 5.7 D and C in - 8.59 -10.0 D, C, and B in
-10.14 -10.55 D, C, B, and A in
-14.63 -14.45
_________________ a. Isothermal coefficients, which include the Doppler effect in the fuel.
FT/k kln10 1 2 5 iso°=
FNP-FSAR-4 REV 21 5/08 TABLE 4.3-12 95/95 K eff FOR SPENT FUEL RACK STORAGE CONFIGURATIONS Configuration Nominal Enrichment w/o U-235 No Soluble Boron 95/95 K eff Soluble Boron Credit 95/95 K eff All Cell 2.15 0.99201 0.93741 2-out-of-4 Checkerboard 5.0 0.94285 0.N/A* Burned/Fresh Checkerboard 1.6/3.9 0.99415 0.94025
_________________ *No soluble boron credit is necessary for the 2-out-of-4 checkerboard to maintain K eff < 0.95.
FNP-FSAR-4 REV 22 8/09 TABLE 4.3-13 95/95 K eff FOR SPENT FUEL CASK LOADING OPERATIONS Configuration Nominal Enrichment w/o U-235 No Soluble Boron 95/95 K eff Soluble Boron Credit 95/95 K eff Cask Storage 2.09 0.970 0.945
REV 21 5/08 FUEL LOADING ARRANGEMENT FOR INITIAL CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-1 (SHEET 1 OF 2)
REV 21 5/08 TYPICAL RELOAD FUEL LOADING ARRANGEMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-1 (SHEET 2 OF 2)
REV 21 5/08 PRODUCTION AND CONSUMPTION OF HIGHER ISOTOPES, TYPICAL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-2
REV 21 5/08 BORON CONCENTRATION VERSUS CYCLE BURNUP WITH BURNABLE ABSORBERS, TYPICAL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-3
REV 21 5/08 TYPICAL DISCRETE BURNABLE ABSORBER ROD ARRANGEMENTS WITHIN AN ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-4 (SHEET 1 OF 2)
REV 21 5/08 TYPICAL IFBA ARRANGEMENT WITHIN AN ASSEMBLY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-4 (SHEET 2 OF 2)
REV 21 5/08
[UNIT 1 CYCLE 1 BURNABLE ABSORBER LOADING PATTERN JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-5
]
REV 21 5/08
[UNIT 2 CYCLE 1 BURNABLE ABSORBER LOADING PATTERN JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-6 (SHEET 1 OF 3)
]
REV 21 5/08 TYPICAL DISCRETE BURNABLE ABSORBER LOADING PATTERN JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-6 (SHEET 2 OF 3)
REV 21 5/08 TYPICAL IFBA LOADING PATTERN JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-6 (SHEET 3 OF 3)
REV 21 5/08
[CYCLE 1 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, UNRODDED CORE, HOT FULL POWER, NO XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-7 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, UNRODDED CORE, HOT FULL POWER, NO XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-7 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-8 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-8 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, BANK D AT INSERTION LIMIT, HOT FULL POWER, E QUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-9 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE, BANK D AT INSERTION LIMIT, HOT FULL POWER, EQUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-9 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 NORMALIZED POWER DENSITY DISTRIBUTION NEAR MIDDLE OF LIFE, UNRODDED CORE, HOT FULL POWER, E QUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-10 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD NORMALIZED POWER DENSITY DISTRIBUTION NEAR MIDDLE OF LIFE, UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-10 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 NORMALIZED POWER DENSITY DISTRIBUITON NEAR END OF LIFE, UNRODDED CORE, HOT FULL POWER, E QUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-11 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD NORMALIZED POWER DENSITY DISTRIBUTION NEAR END OF LIFE, UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-11 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 RODWISE POWER DISTRIBUTION IN A TYPICAL ASSEMBLY (ASSEMBLY G-9) NEAR BEGINNING OF LIFE, HOT FULL POWER, EQUILIBRIUM XENON, UNRODDED CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-12 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD RODWISE POWER DISTRIBUTION IN A TYPICAL ASSEMBLY (ASSEMBLY E-10) NEAR BEGINNING OF LIFE, HOT FULL POWER, EQUILIBRIUM XENON, UNRODDED CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-12 (SHEET 2 OF 2)
REV 21 5/08
[CYCLE 1 RODWISE POWER DISTRIBUTION IN A TYPICAL ASSEMBLY (ASSEMBLY G-9) NEAR END OF LIFE, HOT FULL POWER, E QUILIBRIUM XENON, UNRODDED CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-13 (SHEET 1 OF 2)
]
REV 21 5/08 TYPICAL RELOAD RODWISE POWER DISTRIBUTION IN TYPICAL ASSEMBLY (ASSEMBLY E-10) NEAR END OF LIFE, HOT FULL POWER EQUILIBRIUM XENON, UNRODDED CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-13 (SHEET 2 OF 2)
REV 21 5/08 TYPICAL HFP AXIAL POWER SHAPE OCCURRING AT BEGINNING OF LIFE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-14
REV 21 5/08 TYPICAL HFP AXIAL POWER SHAPE OCCURRING AT MIDDLE OF LIFE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-15
REV 21 5/08 TYPICAL HFP AXIAL POWER SHAPE OCCURRING AT END OF LIFE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-16
REV 21 5/08 A TYPICAL COMPARISON OF ASSEMBLY AXIAL POWER DISTRIBUTION WITH CORE AVERAGE AXIAL DISTRIBUTION BANK "D" SLIGHTLY INSERTED JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-17
REV 21 5/08 FLOW CHART FOR DETERMINING SPIKE MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-18
REV 21 5/08 PREDICTED POWER SPIKE DUE TO SINGLE NONFLATTENED GAP IN THE ADJACENT FUEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-19
REV 21 5/08 POWER SPIKE FACTOR AS A FUNCTION OF AXIAL POSITION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-20
REV 21 5/08 MAXIMUM F QX POWER VERSUS AXIAL HEIGHT DURING NORMAL OPERATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-21
REV 21 5/08 PEAK POWER DURING CONTROL ROD MALFUNCTION OVERPOWER TRANSIENTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-22
REV 21 5/08 PEAK POWER DURING BORATION/DILUTION OVERPOWER TRANSIENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-23
REV 21 5/08 COMPARISON BETWEEN CALCULATED AND MEASURED RELATIVE FUEL ASSEMBLY POWER DISTRIBUTION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-24
REV 21 5/08 COMPARISON OF CALCULATED AND MEASURED AXIAL SHAPE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-25
REV 21 5/08 MEASURED VALUES OF F Q FOR FULL POWER ROD CONFIGURATIONS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-26
REV 21 5/08
[DOPPLER TEMPERATURE COEFFI CENT AT BOL AND EOL VERSUS TEFF FOR CYCLE 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-27
]
REV 21 5/08
[DOPPLER ONLY POWER COEF FICIENT VERSUS POWER LEVEL AT BOL AND EOL CYCLE 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-28
]
REV 21 5/08
[DOPPLER ONLY POWER DEFECT VERSUS PERCENT POWER, BOL AND EOL CYCLE I JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-29
]
REV 21 5/08
[MODERATOR TEMPERATURE COEFFICIENT - BOL, CYCLE 1, NO RODS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-30
]
REV 21 5/08
[MODERATOR TEMPERATURE COEFFICIENT EOL, CYCLE 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-31
]
REV 21 5/08
[MODERATOR TEMPERATURE COEFFICIENT AS A FUNCTION OF BORON CONCENTRATION - BOL CYCLE 1, NO RODS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-32
]
REV 21 5/08
[HOT FULL POWER MODERATOR TEMPERATURE COEFFICIENT DU RING CYCLE 1 FOR THE CRITICAL BORON CONCENTRATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-33
]
REV 21 5/08
[TOTAL POWER COEFFICIENT VERSUS PERCENT POWER FOR BOL AND EOL, CYCLE 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-34
]
REV 21 5/08
[TOTAL POWER DEFECT BOL, EOL, CYCLE 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-35
]
REV 21 5/08 ROD CLUSTER CONTROL ASSEMBLY PATTERN JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-36
REV 21 5/08 ACCIDENTAL SIMULTANEOUS WITHDRAWAL OF TWO CONTROL BANKS EOL, HZP BANKS D AND B MOVING IN THE SAME PLANE, TYPICAL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-37
REV 21 5/08 DESIGN - TRIP CURVE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-38
REV 21 5/08 NORMALIZED ROD WORTH VERSUS ROD INSERTION ALL RODS BUT ONE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-39
REV 21 5/08
[AXIAL OFFSET VERSUS TIME, PWR CORE WITH A 12-FT HEIGHT AND 121 ASSEMBLIES JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-40
]
REV 21 5/08
[XY XENON TEST THERMOCOUPLE RESPONSE QUADRANT TILT DIFFERENCE VERSUS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-41
]
REV 21 5/08
[CALCULATED AND MEAS URE DOPPLER DEFECT AND COEFFICIENTS AT BOL TWO-LOOP PLANT, 121 ASSEMBLIES, 12-FOOT CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-42
]
REV 21 5/08
[COMPARISON OF CALCULATED AND MEASURED BORON CONCENTRATION FOR 2-LOOP PLANT, 121 ASSEMBLIES, 12-FT CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-43
]
REV 21 5/08
[COMPARISON OF CALCULATED AND MEASURED C B 2-LOOP WITH 121 ASSEMBLIES, 12-FOOT CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-44
]
REV 21 5/08
[COMPARISON OF CALCULATED AND MEASURED C B IN 3-LOOP PLANT, 157 ASSEMBLIES, 12-FOOT CORE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-45
]
REV 21 5/08 FARLEY MINIMUM IFBA REQUIREMENTS FOR FRESH ASSEMBLY IN BURNED/FRESH CHECKERBOARD STORAGE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.3-46
FNP-FSAR-4 4.4-1 REV 27 4/17 The overall objective of the thermal and hydraulic design of the reactor core is to provide adequate heat transfer which is compatible with the heat generation distribution in the core, such that heat removal by the reactor coolant system (RCS) or the emergency core cooling system (ECCS), when applicable, assures that the following requirements are met:
A. Fuel damage (a) is not expected during normal operation and operational transients (Condition I) or any transient conditions arising from faults of moderate frequency (Condition II). It is not possible, however, to preclude a very small number of rod failures. These will be within the capability of the plant cleanup system and are consistent with the plant design bases.
B. The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged, (a) although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.
C. The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry following transients arising from Condition IV events.
In order to satisfy the above requirements the following design bases have been established for the thermal and hydraulic design of the reactor core.
Basis There will be at least a 95-percent probability that DNB will not occur on the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Condition I and II events) at a 95-percent confidence level.
Discussion The design method employed to meet the DNB design basis for the VANTAGE 5 and LOPAR fuel assemblies is the revised thermal design procedure (RTDP), reference 2. With the RTDP methodology, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes, and DNB correlation predictions are considered statistically to obtain DNB uncertainty factors. Based on the DNB uncertainty factors, RTDP
_________________ a. Fuel damage as used here is defined as penetration of the fission product barrier (i.e. the fuel rod clad).
FNP-FSAR-4 4.4-2 REV 27 4/17 design limit DNBR values are determined such that there is at least a 95-percent probability at a 95-percent confidence level that DNB will not occur on the most limiting fuel rod during normal operation and operational transients and during transient conditions arising from faults of moderate frequency (Condition I and II events as defined in ANSI N18.2).
Uncertainties in the plant operating parameters (pressurizer pressure, primary coolant temperature, reactor power, and reactor coolant system flow) have been evaluated for the Farley Units 1 and 2 for RTD bypass loops eliminated (references 3 and 4). In the departure from nucleate boiling ratio (DNBR) analyses with RTDP, a set of plant operating parameter uncertainties was used as bounding for operation with RTD bypass loops eliminated. Only the random portion of the plant operating parameter uncertainties is included in the statistical combination. Instrumentation bias is treated as a direct DNBR penalty. Since the parameter uncertainties are considered in determining the RTDP design limit DNBR values, the plant safety analyses are performed using input parameters at their nominal values.
The RTDP design limit DNBR values are 1.24 and 1.23 for the typical and thimble cells, respectively, for VANTAGE 5 fuel and 1.25 and 1.24 for the typical and thimble cells, respectively, for LOPAR fuel.
The design limit DNBR values are used as a basis for the technical specifications and for
consideration of the applicability of items requiring NRC approval as defined in 10 CFR 50.59.
To maintain DNBR margin to offset DNB penalties such as those due to fuel rod bow (paragraph 4.4.2.2.7) and transition core (paragraph 4.4.2.2.8), the safety analyses were performed to DNBR limits higher than the design limit DNBR values. The difference between the design limit DNBRs and the safety analysis limit DNBRs results in available DNBR margin. The net DNBR margin, after consideration of all penalties, is available for operating and design flexibility.
The option of thimble plug removal has been included in all of the DNBR analyses performed for the VANTAGE 5 and LOPAR fuel. The primary impact of thimble plug removal on the thermal-hydraulic analysis is an increase in the core bypass flow. Bypass flow is assumed to be ineffective for core heat removal. The increased bypass flow is included in all of the flow and DNBR values presented in table 4.4-1.
Operation with thimble plugs in place reduces the core bypass flow through the fuel assembly thimble tubes. The reduction in core bypass flow for operation with the thimble plugs in place is a DNBR benefit. The increased margin associated with the use of a full complement of thimble plugs can be used to offset DNBR penalties.
The standard thermal design procedure (STDP) is used for those analyses where RTDP is not applicable. In the STDP method, the parameters used in analysis are treated in a conservative way from a DNBR standpoint. The parameter uncertainties are applied directly to the plant
safety analyses input values to give the lowe st minimum DNBR. The DNBR limit for STDP is the appropriate DNB correlation limit increased by sufficient margin to offset the applicable DNBR penalties.
FNP-FSAR-4 4.4-3 REV 27 4/17 Discussion
By preventing departure from nucleate boiling, adequate heat transfer is assured between the fuel cladding and the reactor coolant, thereby preventing fuel damage as a result of inadequate cooling. Maximum fuel rod surface temperature is not a design basis, as it will be within a few degrees of coolant temperature during operation in the nucleate boiling region. Limits provided by the nuclear control and protection systems are such that this design basis will be met for transients associated with Condition II events, including overpower transients. There is an additional large DNBR margin at rated power operation and during normal operating transients.
Basis During modes of operation associated with Condition I and Condition II events, the maximum fuel temperature shall be less than the melting temperature of UO
- 2. The UO 2 melting temperature for at least 95 percent of the peak kW/ft fuel rods will not be exceeded at the 95-percent confidence level. Melting temperature of UO 2 is taken as 5080°F (1) unirradiated and reducing 58°F per 10,000 MWd/Mtu. By precluding UO 2 melting, the fuel geometry is preserved and possible adverse effects of molten UO 2 are eliminated.
To preclude center melting, and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700°F has been selected as the overpower limit. This provides sufficient margin for uncertainties in the thermal evaluations as described in paragraph 4.4.2.10.1.
Discussion Fuel rod thermal evaluations are performed at rated power, maximum overpower, and during transients at various burnups. These analyses assure that this design basis, as well as the fuel integrity design bases given in section 4.2, are met. They also provide input for the evaluation of Condition III and IV faults given in chapter 15.
Basis A minimum of 92.9 percent of the thermal flowrate will pass through the fuel rod region of the core and will be effective for fuel rod cooling. Coolant flow through the thimble tubes, as well as the leakage from the core barrel baffle region into the core, are not considered effective for heat removal.
Discussion Core cooling evaluations are based on the thermal flowrate (minimum flow) entering the reactor vessel. A maximum of 7.1 percent of this value is allotted as bypass flow. This includes rod FNP-FSAR-4 4.4-4 REV 27 4/17 cluster control (RCC) guide thimble cooling flow, head cooling flow, baffle leakage, and leakage
to the vessel outlet nozzle.
The maximum bypass flow fraction of 7.1 percent assumes no plugging devices, burnable absorbers, or secondary source rods in the RCC guide thimble tubes which do not contain RCC
rods.
Basis Modes of operation associated with Condition I and II events shall not lead to hydrodynamic instability.
The above design basis, together with the fuel clad and fuel assembly design bases given in paragraph 4.2.1.1, are sufficiently comprehensive so that additional limits are not required.
Fuel rod diametral gap characteristics, moderator-coolant flow velocity and distribution, and
moderator void are not inherently limiting. Each of these parameters is incorporated into the thermal and hydraulic models used to ensure that the above mentioned design criteria are met.
For instance, the fuel rod diametral gap characteristics change with time (see paragraph 4.2.1.3.1) and the fuel rod integrity is evaluated on that basis. The effect of the moderator flow velocity and distribution (see paragraph 4.4.2.3) and moderator void distribution (see paragraph 4.4.2.5) are included in the core thermal (THINC) evaluation and thus affect the design bases.
Meeting the fuel clad integrity criteria covers possible effects of clad temperature limitations. As noted in paragraph 4.2.1.3.1, the fuel rod conditions change with time. A single clad temperature limit for Condition I or Condition II events is not appropriate, since it would of necessity be overly conservative. A clad temperature limit is applied to the loss-of-coolant accident (LOCA) (subsection 15.4.1), and locked rotor accident.
Table 4.4-1 provides a comparison of the design parameters for the 17 x 17 LOPAR fuel and
the VANTAGE 5 fuel.
FNP-FSAR-4 4.4-5 REV 27 4/17 Consistent with the thermal hydraulic design bases described in subsection 4.4.1, the following discussion pertains mainly to fuel pellet temperature evaluation. A discussion of fuel clad integrity is presented in paragraph 4.2.1.3.1.
The thermal hydraulic design ensures that the maximum fuel temperature is below the melting point of UO 2 (melting point of 5080°F (1) unirradiated and reducing by 58°F per 10,000 MWd/Mtu). (To preclude center melting, and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700°F has been selected as the overpower limit. This provides sufficient margin for uncertainties in the thermal evaluations as described in paragraph 4.4.2.10.1.) The temperature distribution within the fuel pellet is predominantly a function of the local power density and the UO 2 thermal conductivity. However, the computation of radial fuel temperature distributions combines crud, oxide, clad, gap, and pellet conductances. The factors that influence these conductances, such as gap size (or contact pressure), internal gas pressure, gas composition, pellet density, and radial power distribution within the pellet, etc., have been combined into a semi-empirical thermal model (see paragraph 4.2.1.3.1) which includes a model for time dependent fuel densification as given in references 5, and 100 for this section. This thermal model enables the determination of these factors and their net effects on temperature profiles. The temperature predictions have been compared to inpile fuel temperature measurements(6-12,13) and melt radius data(14,15) with good results.
Effect of Fuel Densification on Fuel Rod Temperatures
Fuel densification results in fuel pellet shrinkage. This affects the fuel temperatures in the
following ways:
A. Pellet radial shrinkage increases the pellet diametral gap, which results in increased thermal resistance of the gap and, thus, higher fuel temperatures (see paragraph 4.2.1.3.1).
B. Pellet axial shrinkage may produce pellet-to-pellet gaps resulting in local power spikes and, thus, higher total heat flux hot channel factor, F Q, and local fuel temperatures. Application of a local power spike factor is no longer necessary for Westinghouse fuel designs, as described in paragraph 4.3.2.2.5.
C. Pellet axial shrinkage will result in a fuel stack height reduction and an increase in the linear power generation rate (kW/ft) for a constant core power level. Using the methods described in references 5 and 16, the increase in linear power for the fuel rod specifications, listed in table 4.3-1, is 0.2 percent. This value remains applicable for reference 100.
As described in reference 5, fuel rod thermal evaluations (fuel centerline, average, and surface temperatures) are determined throughout the fuel rod lifetime with consideration of time-dependent densification. Maximum fuel average and surface temperatures, shown in figure 4.4-1 as a function of linear power density (kW/ft),
are peak values attained during the fuel lifetime. Figure 4.4-2 presents the peak FNP-FSAR-4 4.4-6 REV 27 4/17 value of fuel centerline temperature versus the linear power density that is attained during the fuel lifetime.
The maximum pellet temperatures at the hot spot during full-power steady state, and at the maximum overpower trip point are shown in table 4.1-1. The principal factors which are employed in the determination of the fuel temperature are discussed below.
The thermal conductivity of uranium dioxide was evaluated from data reported by Howard, et al.(17); Lucks, et al.
(18); Danial, et al.(19); Feith (20); Vogt, et at.(21); Nishijima, et al.
(22); Wheeler, et al.(23); Godfrey, et al.
(24); Stora, et al.(25); Bush (26); Asamoto, et al.
(27); Kruger (28); and Gyllander (29).
At the higher temperatures, thermal conductivity is best obtained utilizing the integral conductivity to melt which can be determined with more certainty. From an examination of the data, it has been concluded that the best estimate for the value of 2800°C Kdt is 93 W/cm. This conclusion is based on the integral values reported by Gyllander (29), Lyons, et al.
(30), Coplin, et al.
(31), Duncan(14), Bain (32), and Stora (33). The design curve for the thermal conductivity is shown in figure 4.4-3. The section of the curve at temperatures between 0°C and 1300°C is in excellent agreement with the recommendation of the IAEA panel.
(34) The section of the curve above 1300°C is derived for an integral value of 93 W/cm.(14,29,33)
Thermal conductivity for UO 2 at 95-percent theoretical density can be presented best by the following equation:
313T10775.8T238.08.11 1 K (4.4-1) where C cm W K CT An accurate description of the radial power distribution as a function of burnup is needed in determining the power level for incipient fuel melting and other important performance parameters such as pellet thermal expansion, fuel swelling, and fission gas release rates.
FNP-FSAR-4 4.4-7 REV 27 4/17 This information on radial power distributions in UO 2 fuel rods is determined with the neutron transport theory code, LASER. The LASER code has been validated by comparing the code predictions on radial burnup and isotopic distributions with measured radial microdrill
data.(35,36) A "radial power depression factor," f, is determined using radial power distributions predicted by LASER. The factor f enters into the determination of the pellet centerline temperature, T C, relative to the pellet surface temperature, T S, through the expression:
c s T T 4f'qdT)T(k (4.4-2) where k (T) = he thermal conductivity for UO 2 with a uniform density distribution.
q' = the linear power generation rate.
The temperature drop across the pellet-clad gap is a function of the gap size and the thermal conductivity of the gas in the gap. The gap conductance model is selected such that when
combined with the UO 2 thermal conductivity model, the calculated fuel centerline temperatures reflect the inpile temperature measurements. A more detailed discussion of the gap conductance model is presented in references 5, and 100.
The fuel rod surface heat transfer coefficients during subcooled forced convection and nucleate boiling are presented in paragraph 4.4.2.8.1
The outer surface of the fuel rod at the hot spot operates at a temperature of approximately 660°F for steady-state operation at rated power throughout core life because of the onset of nucleate boiling. Initially (beginning of life), this temperature is that of the clad metal outer surface.
During operation over the life of the core, the buildup of oxides and crud on the fuel rod surface causes the clad surface temperature to increase. Allowance is made in the fuel
center melt evaluation for this temperature rise. Since the thermal hydraulic design basis limits DNB, adequate heat transfer is provided between the fuel cladding and the reactor coolant so that the core thermal output is not limited by considerations of the clad
temperature.
FNP-FSAR-4 4.4-8 REV 27 4/17 The total heat flux hot channel factor, F Q, is defined by the ratio of the maximum to core average heat flux. As presented in table 4.3-2 and discussed in paragraph 4.3.2.2.1, F Q for normal operation is 2.32 for LOPAR fuel and 2. 50 for VANTAGE 5 fuel. This results in a peak local power of 12.63 kW/ft for LOPAR fuel and 13.61 kW/ft for VANTAGE 5 fuel. As described in paragraph 4.3.2.2.6, the peak linear power for determination of protection setpoints is less than 22.4 kW/ft.
The centerline temperature at this kW/ft must be below the UO 2 melt temperature over the lifetime of the rod, including allowances for uncertainties. The melt temperature of unirradiated UO 2 is 5080°F (1) and decreases by 58°F per 10,000 MWd/Mtu. The most limiting centerline temperatures at a given local power occur at beginning of life. From figure 4.4-2, the centerline temperature at 22.4 kW/ft is below that required to produce melting.
Fuel centerline temperature at rated 100% power and at the maximum overpower trip point is presented in table 4.1-1.
The phenomenon of fuel rod bowing (37) must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNBR and/or margin obtained from measured plant operating parameters (such as N H F or core flow), which are less limiting than those required by the plant safety analysis, can be used to offset the effect of rod bow.
For the safety analysis of the Farley units, sufficient DNBR margin was maintained (paragraph 4.4.1.1) to accommodate the full and low flow rod bow DNBR penalties which are based on the methodology in reference 38. The rod bow DNBR penalties that are applicable to LOPAR fuel assembly analyses using the WRB-1 DNB correlation and to VANTAGE 5 fuel assembly analyses using the WRB-2 DNB correlation were determined using the
methodology in reference 38.
The maximum rod bow penalties ( 2% DNBR) accounted for in the design safety analysis are based on an assembly average burnup of 24,000 MWd/Mtu. At burnups greater than 24,000 MWd/Mtu, credit is taken for the effect of N H F burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory, and no additional rod bow penalty is required (reference 39).
In the upper spans of the VANTAGE 5 fuel assembly, additional restraint is provided with the intermediate flow mixer (IFM) grids such that the grid-to-grid spacing in those spans with IFM grids is approximately 10 in. compared to approximately 20 in. in the other spans. Using the NRC approved scaling factor results in predicted c hannel closure in the limiting 10-in. spans of less than 50% closure; therefore, no rod bow DNBR penalty is required in the 10-in. spans in the VANTAGE 5 safety analyses.
FNP-FSAR-4 4.4-9 REV 27 4/17 The LOPAR and VANTAGE 5 designs have been show n to be hydraulically compatible in reference 40.
The Westinghouse transition core DNB methodology is given in references 41, 42, and 43. Using this methodology, transition cores are analyzed as if the entire core consisted of one assembly type (full LOPAR or full VANTAGE 5). The resultant DNBRs are then reduced by the appropriate transition core penalty.
The VANTAGE 5 fuel assembly has a higher mixing vane grid loss coefficient relative to the LOPAR mixing vane grid loss coefficient. In addition, the VANTAGE 5 fuel assembly has IFM grids located in spans between mixing vane grids, where no grid exists in the LOPAR assembly. The higher loss coefficients and the additional grids introduce localized flow redistribution from the VANTAGE 5 fuel assembly into the LOPAR assembly at the axial zones near the mixing vane grid and the IFM grid position in a transition core. Between the grids, the tendency for velocity equalization in parallel open channels causes flow to return to the VANTAGE 5 fuel assembly. The localized flow redistribution described above actually benefits the LOPAR assembly. This benefit more than offsets the slight mass flow bias due to velocity equalization at nongridded locations. Thus, the analysis for a full core of LOPAR is appropriate for that fuel type in a transition core. There is no transition core DNBR penalty for the LOPAR fuel.
The transition core penalty is a function of the number of VANTAGE 5 fuel assemblies in the core based on the methodology of reference 44. Modifications to the magnitude of the DNBR transition core penalty for a VANTAGE 5/LOPAR transition are given in reference 45.
Sufficient DNBR margin is maintained in the VANTAGE 5 safety analysis to completely offset this transition core penalty.
The minimum DNBRs for the rated power, design overpower, and anticipated transient conditions are given in table 4.4-1. The minimum DNBR in the limiting flow channel is typically downstream of the peak heat flux location (hot spot) because of the increased downstream enthalpy rise.
DNBRs are calculated by using the correlation and definitions described in paragraphs 4.4.2.3.1 and 4.4.2.3.2. The THINC-IV computer code (discussed in paragraph 4.4.3.4.1) is used to determine the flow distribution in the core and the local conditions in the hot channel for use in the DNB correlation. The use of hot channel factors is discussed in paragraph 4.4.3.2.1 (nuclear hot channel factors) and in paragraph 4.4.2.3.4 (engineering hot channel factors).
FNP-FSAR-4 4.4-10 REV 27 4/17 The primary DNB correlation used for the analysis of the 17 x 17 LOPAR fuel is the WRB-1 correlation (reference 46). The primary DNB correlation used for the analysis of the VANTAGE 5 fuel is the WRB-2 correlation (reference 40).
The WRB-1 correlation was developed based exclusively on the large bank of mixing vane grid rod bundle critical heat flux (CHF) data (over 1100 points) that Westinghouse has collected.
The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over a wide range of variables than the previous correlation used in design.
This correlation accounts directly for both typical and thimble cold wall cell effects, uniform and nonuniform heat flux profiles, and variations in rod heated length and in grid spacing.
The applicable range of parameters for the WRB-1 correlation is as follows:
Pressure 1440 P < 2490 psia Local Mass Velocity 0.9 G loc/10 6 3.7 lb/ft 2-h Local Quality
-0.2 X loc 0.3 Heated Length, Inlet to CHF Location L h 14 ft Grid Spacing 13 g sp 32 in. Equivalent Hydraulic Diameter 0.37 d e 0.60 in. Equivalent Heated Hydraulic Diameter 0.46 d h 0.59 in.
Figure 4.4-5, sheet 1 shows measured CHF plotted against predicted critical heat flux using the WRB-1 correlation.
A correlation limit DNBR of 1.17 for the WRB-1 correlation has been approved by the NRC for
17 x 17 LOPAR fuel.
The WRB-2 DNB correlation was developed to take credit for the VANTAGE 5 intermediate flow mixer (IFM) grid design. A limit of 1.17 is also applicable for the WRB-2 correlation. Figure 4.4-5, sheet 2 shows measured CHF plotted against predicted CHF using the WRB-2 correlation.
Use of this correlation has been conservatively modified to utilize a penalty above a certain high quality threshold within the approved ranges (reference 101).
The applicable range of parameters for the WRB-2 correlation is as follows:
Pressure 1440 P 2490 psia Local Mass Velocity 0.9 G loc/10 6 3.7 lb/ft 2-h Local Quality
-0.1 X loc 0.3 Heated Length, Inlet to CHF Location L h 14 ft Grid Spacing 10 g sp 26 in. Equivalent Hydraulic Diameter 0.33 d e 0.5101 in. Equivalent Heated Hydraulic Diameter 0.45 d h 0.66 in.
FNP-FSAR-4 4.4-11 REV 27 4/17 The W-3 DNB correlation (references 47 and 48) is used for both fuel types where the primary DNBR correlations are not applicable. The WRB-1 and WRB-2 correlations were developed based on mixing vane data and, therefore, are only applicable in the heated rod spans above the first mixing vane grid. The W-3 correlation, which does not take credit for mixing vane grids, is used to calculate DNBR value in the heated region below the first mixing vane grid. In addition, the W-3 correlation is applied in the analysis of accident conditions where the system pressure is below the range of the primary correlations. For system pressures in the range of 500 to 1000 psia, the W-3 correlation limit is 1.45 (reference 49). For system pressures greater than 1000 psia, the W-3 correlation limit is 1.30. A cold wall factor (CWF) (reference 50) is applied to the W-3 DNB correlation to account for the presence of the unheated thimble surfaces.
The DNB heat flux ratio (DNBR) as applied to typical cells (flow cells with all walls heated) and thimble cells (flow cells with heated and unheated walls) is defined as:
loc"qN,DNB"q DNBR (4.4-4) where FEU,DNB"qN,DNB"q (4.4-5) EU,DNB"q is the uniform DNB heat flux as predicted by the WRB-1 DNB correlation, WRB-2 DNB correlation, or the W-3 DNB correlation (typical cell only).
F is the flux shape factor to account for nonuniform axial heat flux distributions (reference
- 51) with the "C" term modified as in reference 48.
loc"q is the actual local heat flux.
The DNBR as applied to the W-3 DNB correlation when a cold wall is present is as follows:
loc"qCW,N,DNB"q DNBR where FCWFD,EU,DNB"qCW,N,DNB"q h
FNP-FSAR-4 4.4-12 REV 27 4/17 hD,EU,DNB"q is the uniform DNB heat flux as predicted by the W-3 cold wall DNB correlation (reference 48) when not all flow cell walls are heated (thimble cold wall cell).
107.014.00535.0 6X78.1Dh509.8 1000 P0619.0 10 G732.4e372.1376.1Ru0.1)52reference(CWF and Dh De1Ru The rate of heat exchange by mixing between flow channels is proportional to the difference in the local mean fluid enthalpy of the respective channels, the local fluid density, and flow velocity. The proportionalism is expressed by the dimensionless thermal diffusion coefficient, TDC, which is defined as:
pVa W TDC (4.4-12) where:
W = flow exchange rate per unit length, lbm/ft-s.
P = fluid density, lbm/ft
- 3. V = fluid velocity, ft/s.
a = lateral flow area between channels per-unit-length, ft 2/ft.
The application of the TDC in the THINC analysis for determining the overall mixing effect or heat exchange rate is presented in reference 53.
Westinghouse has also sponsored and directed mixing tests at Columbia University.
(54) These series of tests using the "R" mixing vane grid design on 13-, 26-, and 32-in. grid spacings were conducted in pressurized water loops at Reynolds numbers similar to that of a PWR core under the following single- and two-phase (subcooled boiling) flow conditions:
Pressure 1500 to 2400 psia Inlet Temperature 332°F to 642°F Mass Velocity 1.0 to 3.5 x 10 6 lb/h ft 2 FNP-FSAR-4 4.4-13 REV 27 4/17 Reynolds Number 1.34 to 7.45 x 10 5
Bulk Outlet Quality
-52.1 to -13.5 percent TDC is determined by comparing the THINC code predictions with the measured subchannel exit temperatures. Data for 26-in. axial grid spacing are presented in figure 4.4-6 where the thermal diffusion coefficient is plotted versus the Reynolds number. TDC is found to be independent of the Reynolds number, mass velocity, pressure, and quality over the ranges tested. The two-phase data (local, subcooled boiling) fell within the scatter of the single-phase
data.
The effect of two-phase flow on the value of TDC has been demonstrated by Cadek, (54) Rowe and Angle,(55, 56) and Gonzalez-Santalo and Griffith.
(57) In the subcooled boiling region the values of TDC were indistinguishable from the single-phase values. In the quality region, Rowe and Angle show that in the case with rod spacing similar to that in PWR reactor core geometry, the value of TDC increased with quality to a point and then decreased, but never below the
single-phase value.
Gonzalez-Santalo and Griffith showed that the mixing coefficient increased as the void fraction increased.
The data from these tests on the "R" grid showed that a design TDC value of 0.038 (for 26-in.
grid spacing) can be used in determining the effect of coolant mixing in the THINC analysis.
A mixing test program similar to the one described above was conducted at Columbia University for the 17 x 17 geometry and mixing vane grids on 26-in. spacing.
(58) The mean value of TDC obtained from these tests was 0.051, and all data were well above the current design value of
0.038. Since the actual grid spacing of 17 x 17 LOPAR fuel is approximately 20 in., additional margin is available for this design, as the value of TDC increases as grid spacing decreases.(54)
The inclusion of three IFM grids in the upper spans of the VANTAGE 5 fuel assembly results in a grid spacing of approximately 10 in. Per reference 40, a design TDC value of 0.038 was chosen as a conservatively low value for use in VANTAGE 5 to determine the effect of coolant mixing in the core thermal performance analysis.
The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quantities. The heat flux hot channel factor considers the local maximum linear heat generation rate at a point (the "hot spot"), and the enthalpy rise hot channel factor involves the maximum integrated value along a channel (the "hot channel").
Each of the total hot channel factors is the product of a nuclear hot channel factor (see table 4.3-2 and paragraph 4.4.3.2) describing the neutron power distribution and an engineering hot
channel factor, which allows for variations in flow conditions and fabrication tolerances. The engineering hot channel factors are made up of subfactors which account for the influence of FNP-FSAR-4 4.4-14 REV 27 4/17 the variations of fuel pellet diameter, density, enr ichment, and eccentricity; inlet flow distribution; flow redistribution; and flow mixing.
Heat Flux Engineering Hot Channel Factor, E Q F The heat flux engineering hot channel factor is used to evaluate the maximum linear heat generation rate in the core. This subfactor is determined by statistically combining the fabrication variations for fuel pellet diameter, density, and enrichment and has a value of 1.03 at the 95-percent probability level with 95-percent confidence. As shown in reference 59, no DNB penalty need be taken for the short, relatively low-intensity heat flux spikes caused by variations in the above parameters, as well as fuel pellet eccentricity and fuel rod diameter variation.
Enthalpy Rise Engineering Hot Channel Factor, E H F The effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise is directly considered in the THINC core thermal subchannel analysis (paragraph 4.4.3.4.1) under any reactor operating condition. The items considered contributing to the enthalpy rise engineering hot channel factor are discussed below:
A. Pellet diameter, density, and enrichment:
Variations in pellet diameter, density, and enrichment are considered statistically in establishing the limit DNBRs (paragraph 4.4.1.1) for the RTDP (reference 2) employed in this application. Uncertainties in these variables are determined from sampling manufacturing data.
B. Inlet Flow Maldistribution:
The consideration of inlet flow maldistribution in core thermal performances is discussed in paragraph 4.4.3.1.2. A design basis of 5-percent reduction in coolant flow to the hot assembly is used in the THINC IV analysis.
C. Flow Redistribution:
The flow redistribution accounts for the reduction in flow in the hot channel because of the local or bulk boiling. The effect of the nonuniform power distribution is inherently considered in the THINC analysis for every operating condition which is evaluated.
D. Flow Mixing:
The subchannel mixing model incorporated in the THINC Code and used in reactor design is based on experimental data (60) discussed in paragraph 4.4.3.4.1. The mixing vanes incorporated in the spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly, as well as between adjacent assemblies. This mixing reduces the enthalpy rise in the hot channel resulting from local power peaking or unfavorable mechanical tolerances.
FNP-FSAR-4 4.4-15 REV 27 4/17
Significant quadrant power tilts are not anticipated during normal operation since this phenomenon is caused by some asymmetric perturbation. A dropped or misaligned RCCA could cause changes in hot channel factors. Ho wever, these events are analyzed separately in chapter 15. This discussion will be confined to flux tilts caused by x-y xenon transients, inlet temperature mismatches, enrichment variations within tolerances, and so forth.
The design value of the enthalpy rise hot channel factor N H F, which includes an 8-percent uncertainty (as discussed in paragraph 4.3.2.2.7), is assumed to be sufficiently conservative that flux tilts up to, and including, the alarm point (see technical specifications) will not result in values of N H F greater than that assumed in this submittal. The design value of F Q does not include a specific allowance for quadrant flux tilts.
The calculated core average and the hot subchannel maximum and average void fractions are presented in table 4.4-2 for operation at full power. The void fraction distribution in the core at various radial and axial locations is presented in reference 61. The void models used in the THINC-IV computer code are described in paragraph 4.4.2.8.3.
Assembly average coolant mass velocity and enthalpy at various radial and axial core locations are given in figures 4.4-7 through 4.4-9. Coolant enthalpy rise and flow distributions are shown for the 4-ft elevation (1/3 of core height) in figure 4.4-7 and 8-ft elevation (2/3 of core height) in
figure 4.4-8, and at the core exit in figure 4.4-9. These distributions are representative of a Westinghouse 3-loop plant. The THINC code analysis for this case utilized a uniform core inlet enthalpy and inlet flow distribution.
The analytical model and experimental data used to calculate the pressure drops shown in table 4.4-1 are described in paragraph 4.4.2.8. The core pressure drop includes the fuel assembly, lower core plate, and upper core plate pressure drops. The full power operation pressure drop values shown in table 4.4-1 are the unrecoverable pressure drops across the vessel, including the inlet and outlet nozzles, and across the core. These pressure drops are based on the best-estimate flow for actual plant operating conditions as described in subsection 5.5.1. This FNP-FSAR-4 4.4-16 REV 27 4/17 subsection also defines and describes the thermal design flow (minimum flow) which is the basis for reactor core thermal performance and the mechanical design flow (maximum flow) which is used in the mechanical design of the reactor vessel internals and fuel assemblies.
Since the best-estimate flow is that flow which is most likely to exist in an operating plant, the calculated core pressure drops in table 4.4-1 are based on this best-estimate flow rather than the thermal design flow.
Uncertainties associated with the core pressure drop values are discussed in paragraph 4.4.2.10.2.
The fuel assembly holddown springs (figure 4.2-2) are designed to keep the fuel assemblies in contact with the lower core plate under all Condition I and II events with the exception of the
turbine overspeed transient associated with a loss of external load. The holddown springs are designed to tolerate the possibility of an over-deflection associated with fuel assembly liftoff for this case and provide contact between the fuel assembly and the lower core plate following this transient. More adverse flow conditions occur during a LOCA. These conditions are presented in subsection 15.4.1.
Hydraulic loads at normal operating conditions are calculated based on the mechanical design flow, which is described in section 5.1, and accounting for the minimum core bypass flow based on manufacturing tolerances. Core hydraulic loads at cold-plant startup conditions are also based on this flow, but are adjusted to account fo r the coolant density difference. Conservative core hydraulic loads for a pump overspeed transient, which create flowrates 20 percent greater than the mechanical design flow, are evaluated to be greater than twice the fuel assembly weight.
The hydraulic verification tests for the LOPAR fuel assembly and the VANTAGE 5 fuel assembly are discussed in references 62 and 40, respectively.
Forced convection heat transfer coefficients are obtained from the familiar Dittus-Boelter correlation(63), with the properties evaluated at bulk fluid conditions:
K CGD023.0 K hD4.0 p8.0 e e (4.4-12) where h = heat transfer coefficient, Btu/h-ft 2-°F.
FNP-FSAR-4 4.4-17 REV 27 4/17 D e = equivalent diameter, ft.
K = thermal conductivity, Btu/h-ft-°F.
G = mass velocity, lb/h-ft
- 2. = dynamic viscosity, lb/ft-h.
C p = heat capacity, Btu/lb-°F.
This correlation has been shown to be conservative (64) for rod bundle geometries with pitch-to-diameter ratios in the range used by PWRs. The onset of nucleate boiling occurs when the clad wall temperature reaches the amount of superheat predicted by Thom's (65) correlation. After this occurrence, the outer clad wall temperature is determined by:
5.0 sat"q 1260 Pexp072.0T where Tsat = wall superheat, T w - T sat'. "q = wall heat flux, Btu/h-ft
- 2.
p = pressure, psia.
T w = outer clad wall temperature, °F.
T sat = saturation temperature of coolant at P, °F.
FNP-FSAR-4 4.4-18 REV 27 4/17 Unrecoverable pressure losses occur as a result of viscous drag (friction) and/or geometry changes (form) in the fluid flow path. The flow field is assumed to be incompressible, turbulent, single-phase water. These assumptions apply to the core and vessel pressure drop calculation for the purpose of establishing the primary loop flowrate. Two-phase considerations are neglected in the vessel pressure drop evaluation because the core-average void is negligible (paragraph 4.4.2.5 and table 4.4-2). Two-phase flow considerations in the core thermal subchannel analyses are considered and the models are discussed in paragraph 4.4.3.1.3.
Core and vessel pressure losses are calculated by equations of the form:
1442gVD LFKc 2 e L (4.4-14) where:
P L = unrecoverable pressure drop, lb f/in 2. = fluid density, lb/ft
- 3. L = length, ft.
D e = equivalent diameter, ft.
V = fluid velocity, ft/s.
g c = 32.174, 2 f mslbftlb K = form loss coefficient, dimensionless.
F = friction loss coefficient, dimensionless.
Fluid density is assumed to be constant at the appropriate value for each component in the core and vessel. Because of the complex core and vessel flow geometry, precise analytical values for the form and friction loss coefficients are not available. Therefore, experimental values for these coefficients are obtained from geometrically similar models.
Values are quoted in table 4.4-1 for unrecoverable pressure loss across the reactor vessel, including the inlet and outlet nozzles, and across the core. The results of full-scale tests of core components and fuel assemblies were utilized in developing the core pressure loss characteristic. The pressure drop for the vessel was obtained by combining the core loss with correlation of 1/7th-scale model hydraulic test data on a number of vessels(66, 67) and form loss relationships.
(68) Moody (69) curves were used to obtain the single-phase friction factors.
FNP-FSAR-4 4.4-19 REV 27 4/17 Tests of the primary coolant loop flowrates will be made (paragraph 4.4.4.1) prior to initial criticality to verify that the flowrates used in the design, which were determined in part from the pressure losses calculated by the m ethod described here, are conservative.
There are three separate void regions considered in flow boiling in a PWR as illustrated in figure 4.4-10. They are the wall void region (no bubble detachment), the subcooled boiling region (bubble detachment), and the bulk boiling region.
In the wall void region, the point where local boiling begins is determined when the clad temperature reaches the amount of superheat predicted by Thom's (65) correlation (discussed in paragraph 4.4.2.8.1). The void fraction in this region is calculated using Maurer's (70) relationship. The bubble detachment point, where the superheated bubbles break away from the wall, is determined by using Griffith's (71) relationship.
The void fraction in the subcooled boiling region (that is after the detachment point) is calculated from the Bowring (72) correlation. This correlation predicts the void fraction from the detachment point to the bulk boiling region.
The void fraction in the bulk boiling region is predicted by using homogeneous flow theory and assuming no slip. The void fraction in this region is therefore a function only of the thermodynamic quality.
DNB core safety limits are generated as a function of coolant temperature, pressure, core
power, and axial power imbalance. Steady-state operation within these safety limits ensures that the DNB design basis is met. Figure 15.1-1 shows the DNBR limit lines and the resulting overtemperature delta T trip lines (which become part of the technical specifications), plotted as T vs. Tavg for various pressures.
This system provides adequate protection against anticipated operational transients that are slow with respect to fluid transport delays in the primary system. In addition, for fast transients, e.g., uncontrolled rod bank withdrawal at power incident (subsection 15.2.2), specific protection functions are provided as described in section 7.2, and the uses of these protection functions are described in chapter 15. (See table 15.1-3.)
The thermal response of the fuel rod is discussed in paragraph 4.4.3.7.
FNP-FSAR-4 4.4-20 REV 27 4/17
As discussed in paragraph 4.4.2.2, the fuel temperature is a function of crud, oxide, clad, gap, and pellet conductances. Uncertainties in the fuel temperature calculation are essentially of two types: fabrication uncertainties, such as variations in the pellet and clad dimensions and the pellet density; and model uncertainties, such as variations in the pellet conductivity and the gap conductance. These uncertainties have been qualified by comparison of the thermal model to the in-pile thermocouple measurements, (6-12) by out-of-pile measurements of the fuel and clad properties,(17-28) and by measurements of the fuel and clad dimensions during fabrication. The resulting uncertainties are then used in all evaluations involving the fuel temperature. The effect of densification on fuel temperature uncertainties is also included in the calculation of the total
uncertainty.
In addition to the temperature uncertainty descr ibed above, the measurement uncertainty in determining the local power and the effect of density and enrichment variations on the local power are considered in establishing the heat flux hot channel factor. These uncertainties are described in paragraph 4.3.2.2.1.
Reactor trip setpoints, as specified in the technical specifications, include allowance for instrument and measurement uncertainties, such as calorimetric error, instrument drift and channel reproducibility, temperature measurement uncertainties, noise, and heat capacity variations.
Uncertainty in determining the cladding temperature results from uncertainties in the crud and oxide thicknesses. Because of the excellent heat transfer between the surface of the rod and the coolant, the film temperature drop does not appreciably contribute to the uncertainty.
Core and vessel pressure drops based on the best-estimate flow, described in section 5.1, are quoted in table 4.4-1. The uncertainties quoted are based on the uncertainties in both the test results and the analytical extension of these values to the reactor application. A major use of the core and vessel pressure drops is to determine the primary system coolant flowrates. In addition, as discussed in paragraph 4.4.4.1, tests on the primary system prior to initial criticality will be made to verify that a conservative primary system coolant-flowrate has been used in the design and analyses of Farley Nuclear Plant.
The effects of uncertainties in the inlet flow maldistribution criteria used in the core thermal analyses are discussed in paragraph 4.4.3.1.2.
FNP-FSAR-4 4.4-21 REV 27 4/17 The uncertainty in the DNB correlation (paragraph 4.4.2.3) can be written as a statement on the probability of not being in DNB based on the statistics of the DNB data. This is discussed in paragraph 4.4.2.3.2.
The uncertainties in the DNBRs calculated by THINC analysis (see paragraph 4.4.3.4.1) because of uncertainties in the nuclear peaking factors are accounted for by applying conservatively high values of the nuclear peaking factors and including measurement error allowances in the statistical evaluation of the limit DNBR (paragraph 4.4.1.1) using the RTDP (reference 2).
In addition, conservative values for the engineering hot channel factors are used as discussed in paragraph 4.4.2.3.4.
The results of a sensitivity study (61) with THINC-IV show that the minimum DNBR in the hot channel is relatively insensitive to variations in the core-wide radial power distribution (for the same value of FH). The ability of the THINC-IV computer code to accurately predict flow and enthalpy distributions in rod bundles is discussed in paragraph 4.4.3.4.1 and in reference 73. Studies have been
performed (61) to determine the sensitivity of the minimum DNBR in the hot channel to the void fraction correlation (see also paragraph 4.4.2.8.3); the inlet velocity and exit pressure distributions, assumed as boundary conditions for the analysis; and the grid pressure loss coefficients. The results of these studies show that the minimum DNBR in the hot channel is relatively insensitive to variations in these parameters. The range of variations considered in these studies covered the range of possible variations in these parameters.
The uncertainties associated with loop flowrates are discussed in section 5.1. A thermal design flow is defined for use in core thermal performance evaluations which accounts for both prediction and measurement uncertainties. In addition, another 7.1 percent of the thermal design flow is assumed to be ineffective for core heat removal capability because it bypasses the core through the various available vessel flow-paths described in paragraph 4.4.3.1.1.
As discussed in paragraph 4.4.2.7.2, hydraulic loads on the fuel assembly are evaluated for a pump overspeed transient which creates flowrates 20 percent greater than the mechanical design flow. The mechanical design flow as stated in section 5.1 is greater than the best estimate or most likely flowrate value for the actual plant operating condition.
FNP-FSAR-4 4.4-22 REV 27 4/17 The value of the mixing coefficient, TDC, used in THINC analyses for this application is 0.038 for LOPAR fuel and VANTAGE 5 fuel.
The results of the mixing tests done on 17 x 17 LOPAR geometry, as discussed in paragraph 4.4.2.3.3, had a mean value of TDC of 0.059 and standard deviation of = 0.007. Hence the current design value of TDC is almost three standard deviations below the mean for 26-in. grid spacing.
Plant configuration data for the thermal hydraulic and fluid systems external to the core are provided in the appropriate chapters 5, 6, and 9. Implementation of the emergency core cooling system is discussed in chapter 15. Some specific areas of interest are the following:
A. Total coolant flowrates for the reactor coolant system and each loop are provided in table 5.1-1. Flowrates employed in the evaluation of the core are presented in section 4.4.
B. Total RCS volume, including pressurizer and surge line and RCS liquid volume (including pressurizer water at steady-state power conditions), are given in table 5.1-1.
C. The flowpath length through each volume may be calculated from physical data provided in the above-referenced sections.
D. The height of fluid in components of the RCS may be determined from the physical data presented in section 5.5. The components of the RCS are water filled during power operation, with the pressurizer being approximately 60-percent water filled.
E. The elevation of components of the RCS relative to the reactor containment are shown in figures 1.2-6 and 1.2-7. Components of the ECCS are to be located in a manner which meets the criteria for NPSH described in section 6.3, and provide the minimum emergency flow as discussed in sections 15.3 and 15.4.
F. Line lengths and sizes for the safety injection system are determined in a manner which guarantees a total system resistance which provides, as a minimum, the fluid delivery rates assumed in the safety analyses described in chapter 15.
G. The minimum flow areas for components of the RCS are presented in section 5.5, Component and Subsystem Design.
H. The steady-state pressure and temperature distributions through the RCS are presented in table 5.1-1.
FNP-FSAR-4 4.4-23 REV 27 4/17
The following flowpaths for core bypass flow are considered:
A. Flow through the spray nozzles into the upper head for head cooling purposes.
B. Flow entering into the RCC guide thimbles to cool the core component rods.
C. Leakage flow from the vessel inlet nozzle directly to the vessel outlet nozzle through the gap between the vessel and the barrel.
D. Flow introduced between the baffle and the barrel for the purpose of cooling these components and not considered available for core cooling.
E. Flow entering into the core from the barrel baffle region through the gaps between the baffle plates.
The above contributions are evaluated to confirm that the design basis value of 7.1-percent core bypass flow is met. This design bypass value is also used in the evaluation of the core pressure drops quoted in table 4.4-1 and the determination of reactor flowrates in section 5.1. Flow model test results for the flowpath through the reactor are discussed in paragraph 4.4.2.8.2.
Data have been considered from several 1/7-scale hydraulic reactor model tests(66)(67)(74) in arriving at the core inlet flow maldistribution criteria to be used in the THINC analyses (see paragraph 4.4.3.4.1). THINC I analyses made using these data have indicated that a conservative design basis is to consider a 5-percent reduction in the flow to the hot assembly.(53) The same design basis of 5-percent reduction to the hot assembly inlet is used in the THINC-IV analyses.
The experimental error estimated in the inlet velocity distribution has been considered as outlined in reference 61, where the sensitivity of changes in inlet velocity distributions to hot channel thermal performance is shown to be small. Studies(61) made with the THINC-IV model show that it is adequate to use the 5-percent reduction in inlet flow to the hot assembly for a loop out of service, based on the experimental data in references 66 and 67.
The effect of the total flowrate on the inlet velocity distribution was studied in the experiments of reference 66. As was expected, on the basis of the theoretical analysis, no significant variation could be found in inlet velocity distribution with reduced flowrate.
FNP-FSAR-4 4.4-24 REV 27 4/17 Two empirical friction factor correlations are used in the THINC-IV computer code (described in paragraph 4.4.3.4.1).
The friction factor in the axial direction, parallel to the fuel rod axis, is evaluated using the Novendstern-Sandberg correlation.(75) This correlation consists of the following:
A. For isothermal conditions, this correlation uses the Moody(69) friction factor, including surface roughness effects.
B. Under single-phase heating conditions, a factor is applied based on the values of the coolant density and viscosity at the temperature of the heated surface and at the bulk coolant temperature.
C. Under two-phase flow conditions, the homogeneous flow model proposed by Owens (76) is used with a modification to account for a mass velocity and heat flux effect.
The flow in the lateral directions, normal to the fuel rod axis, views the reactor core as a large tube bank. Thus, the lateral friction factor proposed by Idel'chick(68) is applicable. This correlation is of the form 2.0ReAF L L (4.4-15) where: A = is a function of the rod pitch and diameter as given in reference 68.
L Re = is the lateral Reynolds number based on the rod diameter.
Extensive comparisons of THINC-IV predictions using these correlations to experimental data are given in reference 73 and verify the applicability of these correlations in PWR design.
The core power distribution, which is largely established at beginning of life by fuel enrichment, loading pattern, and core power level, is also a function of variables such as control rod worth and position and fuel depletion throughout lifetime. Radial power distributions in various planes of the core are often illustrated for general interest. However, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater importance for DNB analyses. These radial power distributions, characterized by N H F(defined in paragraph 4.3.2.2.2), as well as axial heat flux profiles, are discussed in the following two sections.
FNP-FSAR-4 4.4-25 REV 27 4/17 N H F Given the local power density q' (kW/ft) at point x, y, z in a core with N fuel rods and height H, rodsall 0 0 N Hdz)z,y,x(qH N 1dz)zo,yo,xo(qHMaxpowerrod averagepowerrodhot F The way in which N H F is used in the DNB calculation is important. It is obvious that the location of minimum DNBR will depend on the axial profile and the value of DNBR will depend on the enthalpy rise to that point. Basically, the maximum value of the rod integral is used to identify the most likely rod for minimum DNBR. An axial power profile is obtained which, when normalized to the design value of N H F, recreates the axial heat flux along the limiting rod. The surrounding rods are assumed to have the same axial profile with rod average powers which are typical of distributions found in hot assemblies. In this manner, worst-case axial profiles can be combined with worst-case radial distributions for reference DNB calculations.
It should be noted again that N H F, is an integral and is used as such in the DNB calculations.
Local heat fluxes are obtained by using hot channel and adjacent channel explicit power shapes which take into account variations in horizontal power shapes throughout the core. The sensitivity of the THINC-IV analysis to radial power shapes is discussed in reference 61. For operation at a fraction P of full power, the design N H F, used is given by:
fuelLOPARfor)]P1(3.01[30.1F N H fuel5 VANTAGEfor)]P1(3.01[70.1F N H It should be noted that the maximum value of the analysis of record N H F for both Unit 1 and Unit 2 is 1.70 as indicated above. However, the maximum N H F in the Unit 2 Technical Specifications is 1.65, pending a Technical Specification change submittal for increasing the Unit 2 N H F from 1.65 to 1.70.
The permitted relaxation of N H F is included in the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits, (77) thus allowing greater flexibility in the nuclear design.
As discussed in paragraph 4.3.2.2, the axial heat flux distribution can vary as a result of rod motion, power change, or because of spatial xenon transients which may occur in the axial direction. Consequently, it is necessary to measure the axial power imbalance by means of the FNP-FSAR-4 4.4-26 REV 27 4/17 ex-core nuclear detectors (as discussed in paragraph 4.3.2.2.7) and protect the core from excessive axial power imbalance. The reactor trip system provides automatic reduction of the trip setpoint in the overtemperature T channels on excessive axial power imbalance; that is, when an extremely large axial offset corresponds to an axial shape which could lead to a DNBR which is less than that calculated for the reference DNB design axial shape.
The reference DNB design axial shape used in the automatic reduction of the overtemperature T setpoint is a chopped cosine shape with a peak-to-average of 1.55.
A general summary of the steady-state thermal hydraulic design parameters including thermal output, flowrates, etc., is provided in table 4.4-1 for all loops in operation.
As stated in subsection 4.4.1, the design bases of the application are to prevent departure from nucleate boiling and to prevent fuel melting for Condition I and II events. The protective systems described in chapter 7 (Instrumentation and Controls) are designed to meet these bases. The response of the core to Condition II transients is given in section 15.
The objective of reactor core thermal design is to determine the maximum heat removal capability in all flow subchannels and show that the core safety limits, as presented in technical specifications, are not exceeded while compounding engineering and nuclear effects. The thermal design takes into account local variations in dimensions, power generation, flow redistribution, and mixing.
THINC-IV is a realistic three-dimensional matrix model developed to account for hydraulic and nuclear effects on the enthalpy rise in the core.(61, 73) The behavior of the hot assembly is determined by superimposing the power distribution among the assemblies on the inlet flow distribution while allowing for flow mixing and flow distribution between assemblies. The average flow and enthalpy in the hottest assembly is obtained from the core-wide, assembly-by-assembly analysis. The local variations in power, fuel rod and pellet fabrication, and mixing within the hottest assembly are then superimposed on the average conditions of the hottest assembly in order to determine the conditions in the hot channels.
Steady-State Analysis The THINC-IV computer program as approved by the NRC (78,79) is used to determine coolant density, mass velocity, enthalpy, vapor void, static pressure, and DNBR distributions along parallel flow channels within a reactor core under all expected operating conditions. The THINC-IV code is described in detail in references 61, 73, and 78, including models and FNP-FSAR-4 4.4-27 REV 27 4/17 correlations used. In addition, a discussion on experimental verification of THINC-IV is given in reference 73. The core region being studied is considered to be made up of a number of contiguous elements in a rectangular array extending the full length of the core. An element may represent any region of the core, from a single assembly to a subchannel.
The momentum and energy exchange between elements in the array are described by the equations for the conservation of energy and mass, the axial momentum equation, and two lateral momentum equations that couple each element with its neighbors. The momentum equations used in THINC-IV are similar to the Euler equations, (80) except that frictional loss terms have been incorporated which represent the combined effects of frictional and form drag caused by the presence of grids and fuel assembly nozzles in the core. The crossflow resistance model used in the lateral momentum equations was developed from experimental data for flow normal to tube banks.(68, 81) The energy equation for each element also contains additional terms that represent the energy gain or loss because of the crossflow between elements.
The unique feature in THINC-IV is that lateral momentum equations, which include both inertial and crossflow resistance terms, have been incorporated into the calculational scheme. This differentiates THINC-IV from other thermal hydraulic programs in which only the lateral resistance term is modeled. Another important consideration in THINC-IV is that the entire velocity field is solved en masse, by a field equation, while in other codes such as THINC-I (82) and COBRA (83) the solutions are obtained by step-wise integration throughout the array.
The resulting formulation of the conservation equations are more rigorous for THINC-IV; therefore, the solution is more accurate. In addition, the solution method is complex and some simplifying techniques must be employed. Since the reactor flow is chiefly in the axial direction, the core flow field is primarily one-dimensional, and it is reasonable to assume that the lateral velocities and the parameter gradients are larger in the axial direction than the lateral direction. Therefore, a perturbation technique can be used to represent the axial and lateral parameters in the conservation equations. The lateral velocity components are regarded as perturbed quantities which are smaller than the unperturbed component equaling the core average value at a given elevation and the perturbed value as the difference between the local value and the unperturbed component. Since the magnitudes of the unperturbed and perturbed parameters are significantly different, they can be solved separately. The unperturbed equations are one-dimensional and can be solved with the resulting solutions becoming the coefficients of the perturbed equations. An iterative method is then used to solve the system of perturbed equations which couples all the elements in the array.
Experimental Verification An experimental verification (73) of the THINC-IV analysis for core-wide, assembly-to-assembly enthalpy rises, as well as enthalpy rise in a nonuniformly heated rod bundle, has been obtained.
In these experimental tests, the system pre ssure, inlet temperature, mass flowrate, and heat fluxes were typical of present PWR core designs.
During the operation of a reactor, various incore monitoring systems obtain measured data indicating the core performance. Assembly power distributions and assembly mixed mean FNP-FSAR-4 4.4-28 REV 27 4/17 temperature are measured and can be converted into the proper three-dimensional power input needed for the THINC programs. These data can then be used to verify the Westinghouse thermal hydraulic design codes.
One standard startup test is the natural circulation test in which the core is held at a very low power (~2 percent) and the pumps are turned off. The core will then be cooled by the natural circulation currents created by the power differences in the core. During natural circulation, a thermal siphoning effect occurs, resulting in the hotter assemblies gaining flow, thereby creating significant interassembly crossflow. As described in the preceding section, the most important feature of THINC-IV is the method by which crossflow is evaluated. Thus, tests with significant crossflow are of more value in the code verification. Interassembly crossflow is caused by radial variations in pressure. Radial pressure gradients are, in turn, caused by variations in the axial pressure drops in different assemblies. Under normal operating conditions (subcooled forced convection) the axial pressure drop is caused mainly by friction losses. Since all assemblies have the same geometry, all these assemblies have nearly the same axial pressure drops, and crossflow velocities are small. However, under natural circulation conditions (low flow) the axial pressure drop is caused primarily by the difference in elevation head (or coolant density) between assemblies (axial velocity is low and therefore axial friction losses are small). This phenomenon can result in relatively large radial pressure gradients and, therefore, higher crossflow velocities than at normal reactor operating conditions.
The incore instrumentation was used to obtain the assembly-by-assembly core power distribution during a natural circulation test. Assembly exit temperatures during the natural circulation test on a 157-assembly, three-loop plant were predicted using THINC-IV. The predicted data points were plotted as assembly temperature rise vs. assembly power and a least squares fitting program was used to generate an equation which best fit the data. The result is the straight line presented in figure 4.4-11. The measured assembly exit temperatures are reasonably uniform, as indicated in this figure, and are predicted closely by the THINC-IV code. This agreement verifies the lateral momentum equations and the crossflow resistance model used in THINC-IV. The large crossflow resistance used in THINC-I reduces flow redistribution so that THINC-IV gives better agreement with the experimental data.
Data have also been obtained for Westinghouse plants operating from 67 percent to 101 percent of full power. A representative cross-section of the data obtained from a two-loop and a three-loop reactor were analyzed to verify the THINC-IV calculational method. The THINC-IV predictions were compared with the experimental data as shown in figures 4.4-12 and 4.4-13.
The predicted assembly exit temperatures were compared with the measured exit temperatures for each data run. The standard deviations of the measured and predicted assembly exit temperatures were calculated and compared for both THINC-IV and THINC-I and are given in table 4.4-3. As the standard deviations indicate, THINC-IV generally fits the data somewhat more accurately than THINC-I. For the core inlet temperatures and power of the data examined, the coolant flow is essentially single phase. Thus, one would expect little interassembly crossflow and small differences between THINC-IV and THINC-I predictions as seen in the tables. Both codes are conservative and predict exit temperatures higher than measured values for the high-powered assemblies.
As experimental verification of the THINC-IV subchannel calculation method has been obtained from exit temperature measurements in a nonuniformly heated rod bundle.
(95) The inner nine FNP-FSAR-4 4.4-29 REV 27 4/17 heater rods were operated at approximately 20 percent more power than the outer rods to create a typical PWR intrassembly power distribution. The rod bundle was divided into 36 subchannels and the temperature rise was calculated by THINC-IV using the measured flow and power for each experimental test.
Figure 4.4-14 shows, for a typical run, a comparison of the measured and predicted temperature rises as a function of the power density in the channel. The measurements represent an average of two-to-four measurements taken in various quadrants of the bundle. It is seen that the THINC-IV results predict the temperature gradient across the bundle very well. In figure 4.4-15, the measured and predicted temperature rises are compared for a series of runs at different pressures, flows, and power levels.
Again, the measured points represent the average of the measurements taken in the various quadrants. It is seen that the THINC-IV predictions provide a good representation of the data.
Extensive additional experimental verification is presented in reference 73.
The THINC-IV analysis is based on a knowledge and understanding of the heat transfer and hydrodynamic behavior of the coolant flow and the mechanical characteristics of the fuel elements. The use of the THINC-IV analysis provides a realistic evaluation of the core performance and is used in the thermal analysis as described above.
Transient Analysis The THINC-IV thermal-hydraulic computer code does not have a transient capability. Since the third section of the THINC-I program(82) does have this capability, this code (THINC-III) continues to be used for transient DNB analysis.
The conservation equations needed for the transient analysis are included in THINC-III by adding the necessary accumulation terms to the conservation equations used in the steady-state (THINC-I) analysis. The input description must now include one or more of the following
time dependent arrays:
A. Inlet flow variation.
B. Heat flux distribution.
C. Inlet pressure history.
At the beginning of the transient, the calculation procedure is carried out as in the steady-state analysis. The THINC-III code is first run in the steady-state mode to ensure conservatism with respect to THINC-IV and in order to provide the steady-state initial conditions at the start of the transient. The time is incremented by an amount determined either by the user or by the program itself. At each new time step, the calculations are carried out with the addition of the accumulation terms which are evaluated using the information from the previous time step. This procedure is continued until a preset maximum time is reached.
FNP-FSAR-4 4.4-30 REV 27 4/17 At preselected intervals, a complete description of the coolant parameter distributions with the array, as well as DNBR, is printed out. In this manner the variation of any parameter with time can be readily determined.
At various times during the transient, steady-state THINC-IV is applied to show that the application of the transient version of THINC-I is conservative.
The THINC-III code does not have the capability for evaluating fuel rod thermal response. This is treated by the methods described in subsection 15.1.9.
As discussed in paragraph 4.4.2.2, the fuel rod behavior is evaluated utilizing a semiempirical thermal model which considers, in addition to the thermal aspects, such items as clad creep, fuel swelling, time-dependent densification, fission gas release, release of absorbed gases, cladding corrosion and elastic deflection, and helium solubility.
A detailed description of the thermal model can be found in references 5, and 100.
The analytical methods used to assess hydraulic instability are discussed in paragraph 4.4.3.5.
Boiling flow may be susceptible to thermohydrodynamic instabilities (reference 85). These instabilities are undesirable in reactors since they may cause a change in thermohydraulic conditions that may lead to a reduction in the DNB heat flux relative to that observed during a steady flow condition, or to undesired forced vibrations of core components. Therefore, a thermohydraulic design criterion was developed which states that modes of operation under Condition I and II events shall not lead to thermohydrodynamic instabilities.
Two specific types of flow instabilities are considered for Westinghouse PWR operation. These are the Ledinegg or flow excursion-type of static instability and the density wave-type of dynamic instability.
A Ledinegg instability involves a sudden change in flowrate from one steady state to another. This instability occurs (reference 85) when the slope of the reactor coolant system pressure drop-flowrate curve INTERNAL G p 1 becomes algebraically smaller than the loop supply (pump head) pressure drop-flowrate curve EXTERNAL G p. The criterion for stability is thus EXTERNAL G p INTERNAL G p. The Wpump head curve has a negative slope 0 EXTERNAL G p, whereas the reactor coolant system pressure drop-flow curve has a positive slope FNP-FSAR-4 4.4-31 REV 27 4/17 0 INTERNAL G p over the Condition I and Condition II operational ranges. Thus, the Ledinegg instability will not occur.
The mechanism of density wave oscillations in a heated channel has been described by Lahey and Moody (reference 86). Briefly, an inlet flow fluctuation produces an enthalpy perturbation.
This perturbs the length and the pressure drop of the single-phase region and causes quality or void perturbations in the two-phase regions which travel up the channel with the flow. The quality and length perturbations in the two-phase region create two-phase pressure drop perturbations. However, since the total pressure drop across the core is maintained by the characteristics of the fluid system external to the core, then the two-phase pressure drop perturbation feeds back to the single-phase region. These resulting perturbations can be either attenuated or self-sustained.
A simple method has been developed by Ishii (reference 87) for parallel, closed-channel systems to evaluate whether a given condition is stable with respect to the density wave-type of dynamic instability. This method had been used to assess the stability of typical Westinghouse reactor designs (references 88, 89, 90) under Condition I and II operation. The results indicate that a large margin to density wave instability exists, e.g., increases on the order of 200 percent of rated reactor power would be required for the predicted inception of this type of instability.
The application of the method of Ishii (reference 87) to Westinghouse reactor designs is conservative because of the parallel open-channel feature of Westinghouse PWR cores. For such cores, there is little resistance to lateral flow leaving the flow channels of high power density. There is also energy transfer from channels of high power density to lower-power density channels. This coupling with cooler channels has led to the opinion that an open-channel configuration is more stable than the above closed-channel analysis under the same boundary conditions. Flow stability tests (reference 91) have been conducted where the closed-channel systems were shown to be less stable than when the same channels were cross-connected at several locations. The cross-connections were such that the resistance to channel crossflow and enthalpy perturbations would be greater than that which would exist in a PWR core which has a relatively low resistance to crossflow.
Flow instabilities which have been observed have occurred almost exclusively in closed-channel systems operating at low pressure relative to the Westinghouse PWR operating pressures. Kao, Morgan, and Parker (reference 92) analyzed parallel closed-channel stability experiments simulating a reactor core flow. These experiments were conducted at pressures up to 2200 psia. The results showed that for flow and power levels typical of power reactor conditions, no flow oscillations could be induced above 1200 psia.
Additional evidence that flow instabilities do not adversely affect thermal margin is provided by the data from the rod bundle DNB tests. Many Westinghouse rod bundles have been tested over wide ranges of operating conditions with no evidence of premature DNB or of inconsistent data which might be indicative of flow instabilities in the rod bundle.
In summary, it is concluded that thermohydrodynamic instabilities will not occur under Condition I and II modes of operation for Westinghouse PWR reactor designs. A large power margin, greater than doubling rated power, exists to predicted inception of such instabilities. Analysis FNP-FSAR-4 4.4-32 REV 27 4/17 has been performed which shows that minor plant-to-plant differences in Westinghouse reactor designs such as fuel assembly arrays, core power flow ratios, fuel assembly length, etc., will not result in gross deterioration of the above power margins.
Waterlogging damage of a fuel rod could occur as a consequence of a power increase on a rod after water has entered the fuel rod through a cladding defect. Water entry will continue until the fuel rod internal pressure is equal to the reactor coolant pressure. A subsequent power increase raises the temperature and, hence, could raise the pressure of the water contained within the fuel rod. The increase in hydrostatic pressure within the fuel rod then drives a portion of the water from the fuel rod through the water entry defect. Cladding distortion and/or rupture can occur if the fuel rod internal pressure increase is excessive because of insufficient venting of water to the reactor coolant. This occurs when there is both a rapid increase in the temperature of the water within the fuel rod and small defect. Zircaloy-clad fuel rods which have failed because of waterlogging (93, 94) indicate that very rapid power transients are required for fuel failure. Normal operational transients are limited to about 40 cal/g-min (peak rod) while the Spert tests(93) indicate that 120- to 150-cal/g is required to rupture the cladding even with very short transients (5.5 ms period). Release of the internal fuel rod pressure is expected to have minimal effect on the reactor coolant system (93) and is not expected to result in failure of additional fuel rods.
(94) Ejecting of fuel pellet fragments into the coolant stream is not expected.(93, 94) A cladding breach because of waterloggi ng is thus expected to be similar to any fuel rod failure mechanism which exposes fuel pellets to the reactor coolant stream.
Waterlogging has not been identified as the mechanism for cladding distortion or perforation of any Westinghouse Zircaloy-4/ZIRLO/Optimized ZIRLO clad fuel rods.
The fuel rod experiences many operational transients (intentional maneuvers) during its residency in the core. A number of thermal effects must be considered when analyzing the fuel rod performance.
The clad can be in contact with the fuel pellet at some time in the fuel lifetime. Clad pellet interaction occurs if the fuel pellet temperature is increased after the clad is in contact with the pellet. Clad pellet interaction is discussed in paragraph 4.2.1.3.1.
The potential effects of operation with waterlogged fuel are discussed in paragraph 4.4.3.6, which concluded that waterlogging is not a concern during operational transients.
Clad flattening, as noted in paragraph 4.2.1.3.1, has been observed in some operating power reactors. Thermal expansion (axial) of the fuel rod stack against a flattened section of clad could cause failure of the clad. This is no longer a concern because clad flattening is precluded during the fuel residence in the core. (See paragraph 4.2.1.3.1.)
There can be a differential thermal expansion between the fuel rods and the guide thimbles during a transient. Excessive bowing of the fuel rods could occur if the grid assemblies did not FNP-FSAR-4 4.4-33 REV 27 4/17 allow axial movement of the fuel rods relative to the grids. Thermal expansion of the fuel rods is considered in the grid design so that axial loads imposed on the fuel rods during a thermal transient will not result in excessively-bowed fuel rods (see paragraph 4.2.1.3.2).
As discussed in paragraph 4.4.3.3, the core is protected from going through DNB over the full range of possible operating conditions. At full power nominal operating conditions, the minimum DNBR is 2.23 for the VANTAGE 5 fuel and 3.02 for the LOPAR fuel as compared to the DNBR limits of 1.23 and 1.24, respectively. This means that at nominal conditions, the probability of a rod going through DNB is negligible based on the statistics used with RTDP to determine the DNBR limit. In the extremely unlikely event that DNB should occur, the clad temperature will rise because of the steam blanketing at the rod surface and the consequent degradation in heat transfer. During this time there is a potential for a chemical reaction between the cladding and the coolant. However, because of the relatively good film-boiling heat transfer following DNB, the energy release resulting from this reaction is insignificant compared to the power produced by the fuel.
DNB With Physical Burnout - Westinghouse (95) has conducted DNB tests in a 25-rod bundle where physical burnout occurred with one rod. After this occurrence, the 25-rod test section was used for several days to obtain more DNB data from the other rods in the bundle. The burnout and deformation of the rod did not affect the performance of neighboring rods in the test section during the burnout or the validity of the subsequent DNB data points as predicted by the W-3 correlation. No occurrences of flow instability or other abnormal operation were observed.
DNB With Return to Nucleate Boiling - Additional DNB tests have been conducted by Westinghouse (96) in 19- and 21-rod bundles. In these tests, DNB without physical burnout was experienced more than once on a single rod in the bundles for short periods of time. Each time, a reduction in power of approximately 10 percent was sufficient to reestablish nucleate boiling on the surface of the rod. During these and subsequent tests, no adverse effects were observed on this rod or any other rod in the bundle as a consequence of operating in DNB.
A full discussion of waterlogging, including energy release, is contained in paragraph 4.4.3.6. It is noted that the resulting energy release is not expected to affect neighboring fuel rods.
Coolant flow blockages can occur within the coolant channels of a fuel assembly or external to the reactor core. The effects of fuel assembly blockage within the assembly on fuel rod behavior is more pronounced than external blockages of the same magnitude. In both cases, the flow blockages cause local reductions in coolant flow. The amount of local flow reduction, where it occurs in the reactor, and how far along the flow stream the flow reduction persists are considerations which will influence the fuel rod behavior. The effects of coolant flow blockages, FNP-FSAR-4 4.4-34 REV 27 4/17 in terms of maintaining rated core performance, are determined both by analytical and experimental methods. The experimental data are usually used to augment analytical tools. Inspection of the DNB correlations (paragraph 4.4.2.3 and references 40, 46, 47, 48, and 51) shows that the predicted DNBR is dependent upon the local values of quality and mass velocity.
The THINC-IV code is capable of predicting the effects of local flow blockages on DNBR within the fuel assembly on a subchannel basis, regardless of where the low blockage occurs. In reference 73, it is shown that for a fuel assembly similar to the Westinghouse design, THINC-IV accurately predicts the flow distribution within the fuel assembly when the inlet nozzle is completely blocked. Full recovery of the flow was found to occur about 30 in. downstream of the blockage. With the reactor operating at the nominal full power conditions specified in table 4.4-1, the effects of an increase in enthalpy and decrease in mass velocity in the lower portion of the fuel assembly would not result in a minimum DNBR below the DNBR limit.
From a review of the literature, it is concluded that flow blockage in "open-lattice cores" similar to the Westinghouse cores causes flow perturbations which are local to the blockage. For instance, A. Oktsubo, et al.
(97) show that the mean bundle velocity is approached asymptotically about 4 in. downstream from a flow blockage in a single flow cell. Similar results were also found for 2 and 3 cells completely blocked. Basmer (98), et al., tested an open-lattice fuel assembly in which 41 percent of the subchannels were completely blocked in the center of the test bundle between spacer grids. Their results showed that the stagnant zone behind the flow blockage essentially disappears after 1.65 L/De, or about 5 in. for their test bundle. They also found that leakage flow through the blockage tended to shorten the stagnant zone or, in essence, the complete recovery length. Thus, local flow blockages within a fuel assembly have little effect on subchannel enthalpy rise. The reduction in local mass velocity is then the main parameter which affects the DNBR. If the Farley reactor were operating at full power and nominal steady-state conditions, as specified in table 4.4-1, a significant reduction in local mass velocity (60 percent in the VANTAGE 5 fuel and 85 percent in the LOPAR fuel) would be necessary to reduce the DNBR to the DNBR limit based on the assumption of fully developed flow along the full channel length. In reality, a local flow blockage is expected to promote turbulence and, thus, would likely not effect DNBR at all.
Coolant flow blockages induce local crossflows as well as promote turbulence. Fuel rod behavior is changed under the influence of a sufficiently high crossflow component. Fuel rod vibration could occur, caused by this crossflow component, through vortex shedding or turbulent mechanism. If the crossflow velocity exceeds the limit established for fluid elastic stability, large-amplitude whirling results. The limits for a controlled vibration mechanism are established from studies of vortex shedding and turbulent pressure fluctuations. The crossflow velocity required to exceed fluid elastic stability limits is dependent on the axial location of the blockage and the characterization of the crossflow (jet flow or not). These limits are greater than those for vibratory fuel rod wear.
FNP-FSAR-4 4.4-35 REV 27 4/17
A reactor coolant flow test is performed following fuel loading, but prior to initial criticality. Coolant loop pressure drop data are obtained in this test. These data, in conjunction with coolant pump performance information, allow determination of the coolant flowrates at reactor operating conditions. This test verifies that proper coolant flowrates were used in the core thermal and hydraulic analysis.
Following initial criticality, periodic testing in accordance with the technical specification DNB surveillance for RCS flow will ensure that actual core flowrates are bounded by the assumptions found in the core thermal and hydraulic analysis.
Core power distribution measurements are made at several core power levels (see paragraph 4.3.2.2.7). These tests are used to ensure that conservative peaking factors are used in the core thermal and hydraulic analysis.
Additional demonstration of the overall conservatism of the THINC analysis was obtained by comparing THINC predictions to incore ther mocouple measurements. These measurements were performed on the Zion reactor.
(99) No further inpile testing is planned.
An additional test is provided which measures how the N35 and N36 detector currents are affected by Control Bank D insertions at a constant power level between 30 and 35 percent. The results of this rod shadowing test are used to optimize the calibration of the IR instruments.
Inspections performed on the manufactured fuel are delineated in paragraph 4.2.1.4. Fabrication measurements critical to thermal and hydraulic analysis are obtained to verify that the engineering hot channel factors employed in the design analyses (paragraph 4.4.2.3.4) are
met.
The movable neutron detector with the fixed ther mocouple system is used to provide information on the radial, axial, and azimuthal core characteristics for all core quadrants.
FNP-FSAR-4 4.4-36 REV 27 4/17 The incore instrumentation system is comprised of thermocouples positioned to measure fuel assembly coolant outlet temperatures at preselected positions and fission chamber detectors, positioned in guide thimbles, which run the length of selected fuel assemblies to measure the neutron flux distribution. Figures 4.4-16 and 4.4-17 show the number and location of instrumented assemblies in the core for Units 1 and 2, respectively.
The movable incore neutron detector system is the primary means for monitoring core power distribution. Routine collection of incore data is used to determine fission power density distribution, coolant enthalpy distribution, and fuel burnup distribution.
The core exit thermocouples provide an independent means for monitoring radial core power distribution. The core exit thermocouples are also utilized as post-accident instrumentation for monitoring of adequacy of core cooling.
The incore instrumentation can be used to obtain data from which fission power density distribution in the core, coolant enthalpy distribution in the core, and fuel burnup distribution may be determined.
The overtemperature T trip protects the core against low DNBR. The overpower T trip protects against excessive power (fuel rod rating protection).
As discussed in paragraph 7.2.1.1.2, factors included in establishing the overtemperature T and overpower T trip setpoints include the reactor coolant temperature in each loop and the axial distribution of core power through the use of the two-section, ex-core neutron detectors.
The output of the three ranges (source, intermediate, and power) of detectors, with the electronics of the nuclear instruments, are used to limit the maximum power output of the reactor within their respective ranges.
A total of eight neutron flux detectors are installed in six locations around the reactor in the primary shield. Two proportional counters for the source range are installed on opposite "flat" portions of the core containing the primary startup sources at an elevation approximately one-quarter of the core height. Two compensated ionization chambers for the intermediate range, located in the same instrument wells and detector assemblies as the source range detectors, are positioned at an elevation corresponding to one-half of the core height; four dual-section, uncompensated ionization chamber assemblies for the power range are installed vertically at the four corners of the core and located equidistant from the reactor vessel at all points and, to minimize neutron flux pattern distortions, within 1 ft of the reactor vessel. Each power range detector provides two signals corresponding to the neutron flux in the upper and in the lower sections of a core quadrant. The three ranges of detectors are used as inputs to monitor FNP-FSAR-4 4.4-37 REV 27 4/17 neutron flux from a completely shutdown condition to 120 percent of full power, with the capability of recording overpower excursions up to 200 percent of full power.
The difference in neutron flux between the upper and lower sections of the power range detectors is used to limit the overtemperature-T and overpower-T trip setpoints and to provide the operator with an indication of the core power axial offset. In addition, the outputs of the power range channels are used for:
A. The rod speed control function.
B. To alert the operator to an excessive power imbalance between the quadrants.
C. Protecting the core against the consequences of rod ejection accidents.
D. Protecting the core against the consequences of adverse power distributions resulting from dropped rods.
Details of the neutron detectors and nuclear instrumentation design and the control and trip logic are given in chapter 7. The limits on neutron flux operation and trip setpoints are given in subsection 16.2.3.
On Unit 1 there will be 13 thermocouples positioned at preselected positions to measure the coolant temperatures in the reactor vessel head plenum and two stanchions installed on the internals upper support plate as shown on figures 4.4-18 and 4.4-19. Up to four additional thermocouples will also be installed on the outside surface of the reactor vessel head to obtain additional information above fluid temperatures in this region. Data collected with this instrumentation will be provided to Westinghouse for use in a generic program. The conclusions from this program will be reported to the NRC by Westinghouse.
The heated junction thermocouple (HJTC) system is part of an inadequate core cooling monitoring system (ICCMS). This section addresses the HJTC reactor coolant inventory measurement capability. The remainder of the system is described in subsection 7.5.4.
The HJTC probe assembly in each ICCMS channel consists of eight HJTC sensors, a separator tube, a seal plug, and electrical connectors. The sensors are physically independent and located at key level points from the reactor vessel head to the fuel alignment plate.
As pictured in figure 4.4-20, an HJTC sensor consists of a Chromel-Alumel thermocouple near a heater (or heated junction) and another Chromel-Alumel thermocouple positioned away from the heater (or unheated junction or reference junction). In a fluid with relatively good heat transfer properties, the temperature difference between the adjacent thermocouples is very small. In a FNP-FSAR-4 4.4-38 REV 27 4/17 fluid with relatively poor heat transfer properties, the temperature difference between the thermocouples is large.
The heated and unheated thermocouples in the HJTC probes are connected as shown in figure 4.4-21. When water surrounds the thermocouples, their voltage outputs are approximately equal. Therefore V T is low.
In the absence of liquid, the heated thermocouple temperature increases in relation to the unheated thermocouple, causing V to rise. When V T passes a predetermined setpoint, the system considers the sensor uncovered, changing the display level.
Another determination of the absence of liquid is when the absolute temperature of the sensor (V TR) rises beyond the normal maximum coolant temperature. Then, the system will consider the sensor uncovered.
Two design features ensure proper operation under saturation conditions. First, each HJTC is shielded to avoid overcooling due to direct water contact during two-phase fluid conditions. The HJTC probe with the splash shield is referred to as the HJTC sensor. Second, a string of HJTC sensors is enclosed in a tube that separates the liquid and gas phases that surround it.
The separator tube creates a collapsed liquid level that the HJTC sensors measure. This collapsed liquid level is directly related to the average liquid fraction of the fluid in the reactor head volume above the fuel alignment plate. The mode of direct in-vessel sensing reduces spurious effects due to pressure, fluid properties, and nonhomogeneities of the fluid medium.
The probe assembly is housed in a stainless steel structure that protects the sensors from flow loads and serves as the guide path for the sensors.
The equipment required to install the HJTC consists of a probe holder shroud assembly and a
head port adapter.
There are two probe holder shroud assemblies in the upper internals assembly at core locations N-5 and C-11. The probe holder shroud assembly is similar in design to control rod drive mechanism guide tubes. The probe holder shrouds support, vent, and shroud the Combustion Engineering (CE) HJTC probe and probe holder.
The shroud consists of a lower and upper assembly that are bolted together. A probe holder, provided by CE, is inserted into the center of the probe holder shroud. The probe holder is held in place at four locations. A guide plate assembly is located at three elevations on the inside of the lower assembly of the probe holder shroud assembly. An interference fit exists between the probe holder shroud and each guide plate assembly. The probe holder is bolted to the upper flange of the upper assembly of the probe holder shroud. The shroud, with the CE supplied probe holder, is installed into the upper internals in a manner similar to the CRDM guide tubes.
The shroud is bolted to the upper support plate and has a support pin type arrangement at the bottom.
The head port adapter is compatible with the reactor vessel head penetration on one end and is provided with an integrally machined "Grayloc" hub feature on the other end. The head port FNP-FSAR-4 4.4-39 REV 27 4/17 adapter is machined from a single homogenous piece of metal. The head port adapter is part of the primary pressure boundary and extends approximately 164 in. above the reactor vessel mating surface. The head port adapter provides the HJTC probe access into the vessel.
The only ASME Section III item is the head port adapter. The applicable code for this item is ASME III, 1998 Edition through 2000 Addenda. The applicable material specification is SA-182, Type 316 Stainless Steel.
FNP-FSAR-4 4.4-40 REV 27 4/17
- 1. Cristensen, J. A., Allio, R. J., and Biancheria, A.,"Melting Point of Irradiated UO 2 ," WCAP-6065, February 1965.
- 2. Friedland, A. J. and Ray, S., "Revised Thermal Design Procedure," WCAP-11397-P-A, April 1989.
- 3. Andre', S. V., et.al., "RCS Flow Verification Using Elbow Taps at Westinghouse 3-Loop PWRs," WCAP-14750-P-A (Proprietary), Rev. 1, September 1999.
- 4. Moomau, W. H., and Andre', S. V. "Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Alabama Power Farley Nuclear Plant Units 1 and 2 (Uprating to 2785 MWt NSSS Power)," WCAP-12771, Rev. 1, (Proprietary), September, 1996.
- 5. Weiner, R. A., et al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-P-A, August 1988.
- 6. Kjaerheim, G. and Rolstad, E., "Inpile Determination of UO 2 Thermal Conductivity, Density Effects and Gap Conductance," HPR-80, December 1967.
- 7. Kjaerheim, G., Inpile Measurements of Centre Fuel Temperatures and Thermal Conductivity Determination of Oxide Fuels, paper IFA-175 presented at the European Atomic Energy Society Symposium on Performance Experience of Water Cooled Power Reactor Fuel, Stockholm, Sweden, October 21-22, 1969.
- 8. Cohen, I., Lustman, B., and Eichenberg, J. D., "Measurements of the Thermal Conductivity of Metal-Clad Uranium Oxide Rods During Irradiation," WAPD 228, 1960.
- 9. Clough, D. J. and Sayers, J. B., "The Measurement of the Thermal Conductivity of UO 2 under Irradiation in the Temperature Range 150-1600°C," AERE-R-4690, UKAEA Research Group, Harwell, December 1964.
- 10. Stora, J. P., DeBernardy DeSigoyer, B
., Delmas, R., Deschamps, P., Ringot, C., and Lavaud, B., "Thermal Conductivity of Sintered Uranium Oxide under Inpile Conditions,"
EURAEC-1095, 1964.
- 11. Devold, I., "A Study of the Temperature Distribution in UO 2 Reactor Fuel Elements," AE-318, Aktiebolaget Atomenergi, Stockholm, Sweden, 1968.
- 12. Balfour, M. G., Christensen, J. A., and Ferrari, H. M., "Inpile Measurement of UO 2 Thermal Conductivity," WCAP-2923, 1966.
- 13. Leech, W. J., et al., "Revised PAD Code Thermal Safety Model," WCAP-8720, Addendum 2, October 1982.
FNP-FSAR-4 4.4-41 REV 27 4/17 14. Duncan, R. N., "Rabbit Capsule Irradiation of UO 2," CVTR Project, CVNA-142, June 1962.
- 15. Nelson, R. C., Coplin, D. H., Lyons, M. F., and Weidenbaum, B., "Fission Gas Release from UO 2 Fuel Rods with Gross Central Melting," GEAP-4572, July 1964.
- 16. Hellman, J. M., ed., "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8219, October 1973.
- 17. Howard, V. C. and Gulvin, T. G., "Thermal Conductivity Determinations on Uranium Dioxide by a Radial Flow Method," UKAEA IG-Report 51, November 1960.
- 18. Lucks, C. F. and Deem, H. W., "Thermal Conductivity and Electrical Conductivity of UO 2," in Progress Reports Relating to Civilian Applications, BMI-1448 (Rev.) for June 1960; BMI-1489 (Rev.) for December 1960; and BMI-1518 (Rev.) for May 1961.
- 19. Daniel, J. L., Matolich, J., Jr., and Deem, H. W.,"Thermal Conductivity of UO 2 ," HW-69945, September 1962.
- 20. Feith, A. D., "Thermal Conductivity of UO 2 by a Radial Heat Flow Method," TID-21668, 1962. 21. Vogt, J., Grandell, L., and Runfors, U., "Determination of the Thermal Conductivity of Unirradiated Uranium Dioxide," AB Atomenergi Report RMB-527, quoted by IAEA Report on Thermal Conductivity of Uranium Dioxide, 1964.
- 22. Nishijima, T., Kawada, T., and Ishihata, A., "Thermal Conductivity of Sintered UO 2 and Al 2 O 3 at High Temperatures," J. American Ceramic Society, 48, pp 31-34, 1965.
- 23. Ainscough, J. B. and Wheeler, M. F., "The Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide," in Proceedings of the Seventh Conference on Thermal Conductivity, p. 467, National Bureau of Standards, Washington, 1968.
- 24. Godfrey, T. G., Fulkerson, W., Killie, T. G., Moore J. P., and McElroy, D. L., "Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique," ORNL-3556, June 1964.
- 25. Stora, J. P., et al., "Thermal Conductivity of Sintered Uranium Oxide Under Inpile Conditions," EURAEC-1095, August 1964.
- 26. Bush, A. J., "Apparatus for Measuring Thermal Conductivity to 2500°C," Westinghouse Research Laboratories Report 64-1P6-401-R3, (Westinghouse Proprietary), February 1965.
- 27. Asamoto, R. R., Anselin, F. L., and Conti, A. E., "The Effect of Density on the Thermal Conductivity of Uranium Dioxide," GEAP-5493, April 1968.
FNP-FSAR-4 4.4-42 REV 27 4/17 28. Kruger, O. L., Heat Transport Properties of Uranium and Plutonium Dioxide, paper presented at the fall meeting of Nuclear Division of the American Ceramic Society, Pittsburgh, PA, September 1968.
- 29. Gyllander, J. A., "Inpile Determination of the Thermal Conductivity of UO 2 in the Range 500-2500°C," AE-411, January 1971.
- 30. Lyons, M. F., et al., "UO 2 Powder and Pellet Thermal Conductivity During Irradiation," GEAP-5100-6, 1966.
- 31. Coplin, D. H., et al., "The Thermal Conductivity of UO 2 by Direct In-Reactor Measurements," GEAP-5100-1, March 1968.
- 32. Bain, A. S., "The Heat Rating Required to Produce Center Melting in Various UO 2 Fuels," ASTM Special Technical Publication, No. 306, p 30.
- 33. Stora, J. P., "In-Reactor Measurements of the Integrated Thermal Conductivity of UO 2 - Effect of Porosity," Trans. ANS, 13, p 137, June 1970.
- 34. International Atomic Energy Agency, "Thermal Conductivity of Uranium Dioxide," Report of the Panel held in Vienna, April, 1965, IAEA Technical Reports Series, No. 59, Vienna, The Agency, 1966.
- 35. Poncelet, C. G., "Burnup Physics of Heterogeneous Reactor Lattices," WCAP-6069, June 1965.
- 36. Nodvick, R. J., "Saxton Core II Fuel Performance Evaluation," WCAP-3386-56. Part II, Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel, July 1970.
- 37. Skaritka, J., ed., "Fuel Rod Bow Evaluation, WCAP-8691, Revision 1, July 1979.
- 38. Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1, letter NS-EPR-2515, E. P. Rahe, Jr., (Westinghouse) to J. R. Miller (NRC),
October 9, 1981 and Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1, letter NS-EPR-2572, E. P. Rahe, Jr., to R. J.
Miller, March 16, 1982.
- 39. Letter from C. Berlinger (NRC) to E. P. Rahe Jr. (W),
Subject:
"Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum Rod Bow Penalty, June 18, 1986. 40. Davidson, S. L. and Kramer, W. R., ed. "Reference Core Report VANTAGE 5 Fuel Assembly," WCAP-10444-P-A, September 1985.
- 41. Davidson, S. L. and Iorii, J. A., "Reference Core Report - 17 x 17 Optimized Fuel Assembly," WCAP-9500-A, May 1982.
FNP-FSAR-4 4.4-43 REV 27 4/17 42. Letter from E. P. Rahe (W) to Miller (NRC), NS-EPR-2573, WCAP-9500, and WCAPS-9401/9402 NRC SER Mixed Core Compatibility Items, March 19, 1982.
- 43. Letter from C. O. Thomas (NRC) to Rahe (W) - "Supplement Acceptance No. 2 for Referencing Topical Report WCAP-9500," January 1983.
- 44. Schueren, P. and McAtee, K. R., "Extension of Methodology for Calculating Transition Core DNBR Penalties," WCAP-11837-P-A, January 1990.
- 45. Letter from S. R. Tritch (W) to R. C. Jones (NRC) "VANTAGE 5 DNB Transition Core Effects," ET-NRC-91-3618, September 1991.
- 46. Motley, F. E., et al., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids," WCAP-8762-P, July 1984.
- 47. Tong, L. S., "Critical Heat Fluxes in Rod Bundles, Two Phase Flow and Heat Transfer in Rod Bundles," Annual Winter Meeting ASME, p 3146, November 1968.
- 48. Tong, L. S., "Boiling Crisis and Critical Heat Flux," NRC Critical Review Series, TID-25887, 1972.
- 49. Letter from A. C. Thadani (NRC) to W. J.
Johnson (Westinghouse),
Subject:
Acceptance for Referencing of Licensing Topical Report, WCAP-9226-P/9227-NP, "Reactor Core Response to Excessive Secondary Steam Releases," January 31, 1989.
- 50. Motley, F. E. and Cadek, F. F., "DNB Test Results for R-Grid Thimble Cold Wall Cells," WCAP-7695-L, Addendum 1, October 1972.
- 51. Tong, L. S., "Prediction of Departure from Nucleate Boiling for an Axially Nonuniform Heat Flux Distribution," J. Nucl. Energy, 21, pp 241-248, 1967.
- 52. Motley, F. E. and Cadek, F. F., "DNB Tests Results for New Mixing Vane Grids (R)," WCAP-7695-L, (Westinghouse Proprietary), July 1972 and WCAP-7958-A, January 1975.
- 53. Shefcheck, J., "Application of the THINC Program to PWR Design," WCAP-7359-L, August 1969 (Westinghouse Proprietary), and WCAP-7838, January 1972.
- 54. Cadek, F. F., Motley, F. E., and Dominicis, D. P., "Effect of Axial Spacing on Interchannel Thermal Mixing with the R Mixing Vane Grid," WCAP-7941-P-A, (Westinghouse Proprietary), June 1972 and WCAP-7959-A (Nonproprietary), October 1972. 55. Rowe, D. S. and Angle, C. W., "Crossflow Mixing Between Parallel Flow Channels During Boiling," Part II, "Measurement of Flow and Enthalpy in Two Parallel Channels," BNWL-371, Part 2, December 1967.
FNP-FSAR-4 4.4-44 REV 27 4/17 56. Rowe, D. S. and Angle, C. W., "Crossflow Mixing Between Parallel Flow Channels During Boiling", Part III, "Effect of Spacers on Mixing Between Two Channels," BNWL-371, Part 3, January 1969.
- 57. Gonzalez-Santalo, J. M., and Griffith, P., "Two-Phase Flow Mixing in Rod Bundle Subchannels," ASME Paper 72-WA/NE-19.
- 58. Motley, F. E., Wenzel, A. H., and Cadek, F. F., "The Effect of 17 x 17 Fuel Assembly Geometry on Interchannel Thermal Mixing," WCAP-8299, March 1974.
- 59. Hill, K. W., Motley, F. E., and Cadek, F. F., "Effect of Local Heat Flux Spikes on DNB in Nonuniform Heated Rod Bundles," WCAP-8174, (Westinghouse Proprietary), August 1973 and WCAP-8202 August 1973.
- 60. Cadek, F. F., "Interchannel Thermal Mixing with Mixing Vane Grids," WCAP-7667-L, (Westinghouse Proprietary), May 1971 and WCAP-7755, September 1971.
- 61. Hochreiter, L. E., "Application of the THINC-IV Program to PWR Design," WCAP-8054, (Westinghouse Proprietary), October 1973 and WCAP-8195, October 1973.
- 62. Nakazato, S. and DeMario, E. E., "Hydraulic Flow Test of the 17 x 17 Fuel Assembly," WCAP-8279, February 1974.
- 63. Dittus, F. W. and Boelter, L. M. K., "Heat Transfer in Automobile Radiators of the Tubular Type," Calif. Univ. Publication in Eng., 2, No. 13, pp 443-461, 1930.
- 64. Weisman, J., "Heat Transfer to Water Flowing Parallel to Tube Bundles," Nucl. Sci. Eng., 6, pp 78-79, 1959.
- 65. Thom, J. R. S., Walker, W. M., Fallon, T. A., and Reising, G. F. S., "Boiling in Subcooled Water During Flowup-Heated Tubes or Annuli," Proc. Instn. Mech. Engrs., 180, Pt. C, pp 226-246, 1965-66.
- 66. Hetsroni, G., "Hydraulics Tests of the San Onofre Reactor Model," WCAP-3269-8, June 1964. 67. Hetsroni, G., "Studies of the Connecticut-Yankee Hydraulic Model," NYO-3250-2, June 1965.
- 68. Idel'chik, I. E., Handbook of Hydraulic Resistance, NRC-TR-6630, 1960.
- 69. Moody, L. F., "Friction Factors for Pipe Flow," Transaction of the American Society of Mechanical Engineers, 66 pp 671-684, 1944.
- 70. Maurer, G. W., "A Method of Predicting Steady-State Boiling Vapor Fractions in Reactor Coolant Channels," WAPD-BT-19, pp 59-70, June 1960.
FNP-FSAR-4 4.4-45 REV 27 4/17 71. Griffith, P., Clark, J. A., and Rohsenow, W. M., "Void Volumes in Subcooled Boiling Systems," ASME Paper No. 58-HT-19.
- 72. Bowring, R. W., "Physical Model, Based on Bubble Detachment, and Calculation of Steam Voidage in the Subcooled Region of a Heated Channel," 4PR-10, December 1962.
- 73. Hochreiter, L. E., Chelemer, H., and Chu, P. T., "THINC-IV, An Improved Program for Thermal Hydraulic Analysis of Rod Bundle Cores," WCAP-7956, June 1973.
- 74. Carter, F. D., "Inlet Orificing of Open PWR Cores," WCAP-9004 (Westinghouse Proprietary), January 1969 and WCAP-7836, January 1972.
- 75. Novendstern, E. H. and Sandberg, R. O., "Single-Phase Local Boiling and Bulk Boiling Pressure Drop Correlations," WCAP-2850 (Westinghouse Proprietary), April 1966 and WCAP-7916, June 1972.
- 76. Owens, W. L., Jr., "Two-Phase Pressure Gradient,"International Developments in Heat Transfer, Part II, pp 363-368, ASME, New York, 1961.
- 77. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-L (Westinghouse Proprietary), March 1972 and WCAP-7912, March 1972.
- 78. Friedland, A. J. and Ray, S., "Improved THINC IV Modeling for PWR Core Design, WCAP-12330-P-P, August 1989.
- 79. Letter from Stolz, J. F. (NRC) to Eic heldinger, C., (Westinghouse) Regarding Staff Evaluation of WCAP-7956, WCAP-8054, WCAP-8567, and WCAP-8762, April 1978.
- 80. Vallentine, H. R., Applied Hydrodynamics, Buttersworth Publishers, London, 1959.
- 81. Kays, W. M., and London, A. L., Compact Heat Exchangers, National Press, Palo Alto, 1955.
- 82. Chelemer, H., Weisman, J., and Tong, L. S., "Subchannel Thermal Analysis of Rod Bundle Cores," WCAP-7015, Revision 1, January 1969.
- 83. Rowe, D. S., "COBRA-III, a Digital Computer Program for Steady-State and Transient Thermal Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," BNWL-B-82, 1971.
- 84. Deleted
- 85. Boure, J. A., Bergles, A. E., and Tong, L. S., "Review of Two-Phase Flow Instability," Nucl. Eng. Design 25, pp 165-192, 1973.
- 86. Lahey, R. T. and Moody, F. J., "The Thermal Hydraulics of a Boiling Water Reactor," American Nuclear Society, 1977.
FNP-FSAR-4 4.4-46 REV 27 4/17 87. Saha, P., Ishii, M., and Zuber, N., "An Experimental Investigation of the Thermally-Induced Flow Oscillations in Two-Phase Systems," J. of Heat Transfer, pp 616-622, November 1976.
- 88. Summer, V. C., FSAR, Docket No. 50-395.
- 89. Byron/Braidwood, FSAR, Docket No. 50-456.
- 91. Kakac, S., Veziroglu, T. N., Akyuzlu, K., Berkol, O., Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upflow System, Proc. of 5th International Heat Transfer Conference, Tokyo, September 3-7, 1974.
- 92. Kao, H. S., Morgan, T. D., and Parker, W. B., "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS, Vol. 16, pp 212-213, 1973.
- 93. Stephan, L. A., "The Effects of Cladding Material and Heat Treatment on the Response of Water-logged UO 2 Fuel Rods to Power Bursts," IN-ITR-111, January 1970.
- 94. Western New York Nuclear Research Center Correspondence with the NRC on February 11 and August 27, 1971, Docket 50-57.
- 95. Weisman, J., Wenzel, A. H., Tong, L. S., Fitzsimmons, D., Thorne, W., and Batch, J., "Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressures," Chem. Eng. Prog. Symp. Ser. 64, No. 82, pp 114-125, 1968.
- 96. Tong, L. S., et al., Critical Heat Flux (DNB) in Square and Triangular Array Rod Bundles, presented at the Japan Society of Mechanical Engineers Semi-International Symposium held at Tokyo, Japan, pp 25-34, September 4-8, 1967.
- 97. Ohtsubo, A. and Uruwashi, S., "Stagnant Fluid Due to Local Flow Blockage," J. Nucl. Sci. Technol. 9, No. 7, pp 433-434, 1972.
- 98. Basmer, P., Kirsh, D., and Schultheiss, G. F., "Investigation of the Flow Pattern in the Recirculation Zone Downstream of Local Coolant Blockages in Pin Bundles,"
Atomwirtschaft, 17, No. 8, pp 416-417, 1972. (In German).
- 99. Burke, T. M., Meyer, C. E., Shefcheck, J., "Analysis of Data from the Zion (Unit 1) THINC Verification Test," WCAP-8453 (Westinghouse Proprietary) and WCAP-8454 (Non-proprietary), December 1974.
100. Foster, J. P., et al., "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0)," WCAP-15063-P-A, Revision 1, with Errata, July 2000.
FNP-FSAR-4 4.4-47 REV 27 4/17 101. Westinghouse letter ALA-15-97, dated December 8, 2015, "Westinghouse Resolution Plan and Technical Basis for NSAL-14-5, 'Lower than Expected Critical Heat Flux Results Obtained During DNB Testing.'"
FNP-FSAR-4 TABLE 4.4-1 (SHEET 1 OF 3)
THERMAL AND HYDRAULIC COMPARISON TABLE FOR FNP UNITS 1 AND 2
REV 25 4/14 Design Parameters Reactor core heat output (MWt) 2775 Reactor core heat output (10 6 Btu/h) 9469 Heat generated in fuel (%)
97.4 System pressure, nominal (psia) 2250 System pressure, minimum steady-state (psia) 2200 Coolant temperature Nominal inlet (°F) 530.6 - 541.1 Average rise in core (°F) 78.2 - 77.0 Average rise in vessel (°F) 73.2- 72.2 Average in core (°F) 571.7 - 581.8 Average in vessel (°F) 567.2 - 577.2 Nominal core outlet (°F) 608.8 - 618.1 Nominal vessel outlet (°F) 603.8 - 613.3 Coolant conditions (b) Vessel minimum measured flowrate (MMF)(c) 10 6 lbm/h 101.5 - 100.1 gal/min 263,400 (k) Vessel Thermal Design flowrate (TDF) 10 6 lbm/h 99.4 - 98.1 gal/min 258,000 Effective flowrate for heat transfer (based on TDF) 10 6 lbm/h 92.3 - 91.1 gal/min 239.680 LOPAR VANTAGE 5 Minimum DNBR at nominal conditions Typical flow channel 3.20 2.36 Thimble (cold wall) flow channel 3.02 2.23 Minimum DNBR for design transients Typical flow channel 1.25 1.24 Thimble (cold wall) flow channel 1.24 1.23 FNP-FSAR-4 TABLE 4.4-1 (SHEET 2 OF 3)
Design Parameters LOPAR VANTAGE 5 REV 25 4/14 DNB correlation (a) WRB-1 WRB-2 Effective flow area for heat transfer (ft 2)(d) 41.55 44.04 Average velocity along fuel rods (ft/s)(d) 13.5 12.8 Average mass velocity 10 6 lbm/h-ft 2 (based on TDF)(d) 2.22 - 2.19 2.10 - 2.07 Heat transfer Active heat transfer, surface area (ft 2)(d) 48,598 46,779 Average heat flux (Btu/h-ft 2)(d) 189,820 197,200 Maximum heat flux for normal operation (Btu/h-ft 2)(d,e) 440,380 493,000 Average linear power (kW/ft)(f) 5.45 5.45 Peak linear power for normal operation (kW/ft)(e,f) 12.63 13.61 Peak linear power resulting from overpower transients/operator errors, assuming a maximum overpower of 120% (kW/ft)(g) < 22.4 < 22.4 Peak linear power for prevention of centerline melt (kW/ft)(h) 22.4 22.4 Power density (kW/1 of core)(i) 104.5 104.5 Specific power (kW/kg uranium)(d,i) 37.3 40.7 Fuel Central Temperature Peak at peak linear power for prevention of centerline melt (°F) 4700 4700 Pressure drop Across Core (psi) (l) 23.7 +/- 2.4 (j) Across Vessel, Including Nozzle (psi) (l) 42.3 +/- 4.2 FNP-FSAR-4 TABLE 4.4-1 (SHEET 3 OF 3)
REV 25 4/14 _________________
- a. See paragraph 4.4.1.1 for the use of the W-3 correlation.
- b. Flowrates are based on 15-percent average and 20-percent peak steam generator tube plugging.
- c. Inlet temperature (°F) = 531.3 - 541.8.
- d. Assumes all LOPAR or VANTAGE-5 core.
- e. Based on 2.32 FQ peaking factor for LOPAR and 2.50 FQ peaking factor for VANTAGE-5.
- f. Based on densified active fuel length.
- g. See paragraph 4.3.2.2.6.
- h. See paragraph 4.4.2.2.6.
- i. Based on cold dimensions and 95 percent of theoretical density fuel.
- j. Maximum core pressure drop is based on 0% SGTP, thimble plugging devices installed and the Best Estimate Reactor Flow Rate of 96,200 gpm/loop for Unit 2 (bounds Unit 1).
Thimble plug removal results in a lower pressure drop even though the Best Estimate Flow increases slightly. This pressure drop remains a bounding value for both Unit 1 and Unit 2 containing fuel assemblies with the standardized debris filter bottom nozzle (SDFBN) and a maximum best estimate reactor flow rate of 98,600 gpm/loop.
- k. Value includes a 2.1-percent flow uncertainty (0.1-percent feedwater venturi fouling bias included). The minimum measured flow (MMF) rate is the flow used in the reactor core DNB analyses which were performed with the Revised Thermal Design Procedure. The DNB analyses also bound a MMF of 264,200 gpm which reflects a flow measurement uncertainty of 2.4-percent (0.1-percent feedwater venturi fouling bias included).
- l. The pressure drop for LOPAR fuel is bounded by the pressure drop for VANTAGE-5 fuel.
FNP-FSAR-4 TABLE 4.4-2 VOID FRACTIONS AT NOMINAL REACTOR CONDITIONS
REV 21 5/08 Average (percent) Maximum (percent)
Core (LOPAR) 0.12 -- (VANTAGE 5) 0.23 Hot subchannel (LOPAR) 0.4 0.9 (VANTAGE 5) 8.6 26.5 FNP-FSAR-4 TABLE 4.4-3 COMPARISON OF THINC-I AND THINC-IV PREDICTIONS WITH DATA FROM REPRESENTATIVE WESTINGHOUSE TWO- AND THREE-LOOP REACTORS
REV 21 5/08 Power (Mwt) % Full Power Measured Inlet Temp (°F) rms (°F) THINC-I (°F) THINC-IV Improvement (F
°) for THINC-IV over THINC-I Ginna Reactor 847 65.1 543.7 1.97 1.83 0.14 854 65.7 544.9 1.56 1.46 0.10 857 65.9 543.9 1.97 1.82 0.15 947 72.9 543.8 1.92 1.74 0.18 961 74.0 543.7 1.97 1.79 0.18 1091 83.0 542.5 1.73 1.54 0.19 1268 97.5 542.0 2.35 2.11 0.24 1284 98.8 240.2 2.69 2.47 0.22 1284 98.9 541.0 2.42 2.17 0.25 1287 99.0 544.4 2.26 1.97 0.29 1294 99.5 540.8 2.20 1.91 0.29 1295 99.6 542.0 2.10 1.83 0.27 Robinson Reactor 1427.0 65.1 548.0 1.85 1.88 0.03 1422.6 64.9 549.4 1.39 1.39 0.00 1529.0 88.0 550.0 2.35 2.34 0.01 2207.3 100.7 534.0 2.41 2.41 0.00 2213.9 101.0 533.8 2.52 2.44 0.08
REV 21 5/08 LOPAR PEAK FUEL AVERAGE AND SURFACE TEMPERATURES DURING FUEL ROD LIFETIME VS. LINEAR POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-1 (SHEET 1 OF 2)
REV 21 5/08 ZIRLO CLAD VANTAGE-5 PEAK FUEL AVERAGE AND SURFACE TEMPERATURES DURING FUEL ROD LIFETIME VS. LINEAR POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-1 (SHEET 2 OF 2)
REV 21 5/08 LOPAR PEAK FUEL CENTERLINE TEMPERATURE DURING FUEL ROD LIFETIME VS. LINEAR POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-2 (SHEET 1 OF 2)
REV 21 5/08 ZIRLO CLAD VANTAGE-5 PEAK FUEL CENTERLINE TEMPERATURE DURING FUEL ROD LIFETIME VS. LINEAR POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-2 (SHEET 2 OF 2)
REV 21 5/08 THERMAL CONDUCTIVITY OF UO 2 (DATA CORRECTED TO 95% THEORETICAL DENSITY)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-3
REV 21 5/08 TYPICAL AXIAL VARIATION OF AVERAGE CLAD TEMPERATURE FOR ROD OPERATING AT 5.43 kW/ft JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-4
REV 21 5/08 MEASURED VERSUS PREDICTED CRITICAL HEAT FLUX WRB-1 CORRELATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-5 (SHEET 1 OF 2)
REV 21 5/08 MEASURED VERSUS PREDICTED CRITICAL HEAT FLUX WRB-2 CORRELATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-5 (SHEET 2 OF 2)
REV 21 5/08 TDC VERSUS REYNOLDS NUMBER FOR 26-IN. GRID SPACING JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-6
REV 21 5/08 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT4 FT-ELEVATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-7
REV 21 5/08 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT 8-FT ELEVATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-8
REV 21 5/08 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT 12-FT ELEVATION - CORE EXIT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-9
REV 21 5/08 VOID FRACTION VERSUS THERMODYNAMIC QUALITY H-H SAT/H g-H SAT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-10
REV 21 5/08 PWR NATURAL CIRCULATION TEST JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-11
REV 21 5/08 COMPARISON OF A REPRESENTATIVE W TWO-LOOP REACTOR INCORE THERMOCOUPLE MEASUREMENTS WITH THINC-IV PREDICTIONS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-12
REV 21 5/08 COMPARISON OF A REPRESENTATIVE W THREE-LOOP INCORE THERMOCOUPLE MEASUREMENTS WITH THINC-IV PREDICTIONS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-13
REV 21 5/08 HANFORD SUBCHANNEL TEMPERATURE DATA COMPARISON WITH THINC-IV JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-14
REV 21 5/08 HANFORD SUBCIRTICAL TEMPERATURE DATA COMPARISON WITH THINC-IV JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-15
REV 21 5/08 UNIT 1 DISTRIBUTION OF INCORE INSTRUMENTATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-16
REV 21 5/08 UNIT 2 DISTRIBUTION OF INCORE INSTRUMENTATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-17
REV 21 5/08 UNIT 1 UPPER HEAD THERMOCOUPLE SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-18
REV 21 5/08 UNIT 1 UPPER HEAD THERMOCOUPLE SCHEMATIC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-19
REV 21 5/08 TYPICAL HJTC PROBE/SENSOR CONFIGURATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-20
REV 21 5/08 ELECTRIAL DIAGRAM OF HJTC JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 4.4-21
FNP-FSAR-5 5.0 REACTOR COOLANT SYSTEM AND CONNECTED SYSTEM TABLE OF CONTENTS
Page 5-i REV 21 5/08 5.1
SUMMARY
DESCRIPTION........................................................................................5.1-1 5.1.1 Schematic Flow Diagram............................................................................5.1-5
5.1.1.1 System Operation ......................................................................................5.1-5
5.1.2 Piping and Instrumentation Diagram .........................................................5.1-8
5.1.3 Elevation Drawing.......................................................................................5.1-9
5.2 INTEGRITY OF THE REACTOR COOLANT PRESSURE BOUNDARY...................5.2-1 5.2.1 Design of Reactor Coolant Pressure Boundary Components....................5.2-1
5.2.1.1 Performance Objectives.............................................................................5.2-1 5.2.1.2 Design Parameters.....................................................................................5.2-2 5.2.1.3 Compliance With 10 CFR 50.55a...............................................................5.2-2 5.2.1.4 Applicable Code Cases..............................................................................5.2-2 5.2.1.5 Design Transients.......................................................................................5.2-3 5.2.1.6 Identification of Active Pumps and Valves................................................5.2-10 5.2.1.7 Design of Active Valves ...........................................................................5.2-11 5.2.1.8 Inadvertent Operation of Valves...............................................................5.2-11 5.2.1.9 Stress and Pressure Limits.......................................................................5.2-11 5.2.1.10 Stress Analysis for Structural Adequacy...................................................5.2-11 5.2.1.11 Analysis Method for Faulted Condition.....................................................5.2-22 5.2.1.12 Protection Against Environmental Factors................................................5.2-25 5.2.1.13 Compliance With Code Requirements......................................................5.2-26 5.2.1.14 Stress Analysis for Emergency and Faulted Condition Loadings.............5.2-26 5.2.1.15 Stress Levels in Category I Systems........................................................5.2-26 5.2.1.16 Analytical Methods for Stresses in Pumps and Valves ............................5.2-27 5.2.1.17 Analytical Methods for Evaluation of Pump Speed and Bearing Integrity........................................................................................5.2-27 5.2.1.18 Operation of Active Valves Under Transient Loadings ............................5.2-27 5.2.1.19 Field Run Piping ......................................................................................5.2-28
FNP-FSAR-5 TABLE OF CONTENTS
Page 5-ii REV 21 5/08 5.2.2 Overpressurization Protection..................................................................5.2-28
5.2.2.1 Location of Pressure Relief Devices.........................................................5.2-28 5.2.2.2 Mounting of Pressure Relief Devices........................................................5.2-29 5.2.2.3 Report on Overpressure Protection..........................................................5.2-33 5.2.2.4 RCS Pressure Control During Low Temperature Operation.....................5.2-34
5.2.3 General Material Considerations..............................................................5.2-38
5.2.3.1 Material Specifications..............................................................................5.2-38 5.2.3.2 Compatibility With Reactor Coolant..........................................................5.2-39 5.2.3.3 Compatibility With External Insulation and Environmental Atmosphere...5.2-39 5.2.3.4 Chemistry of Reactor Coolant...................................................................5.2-39
5.2.4 Fracture Toughness..................................................................................5.2-41
5.2.4.1 Compliance With Code Requirements......................................................5.2-41 5.2.4.2 Acceptable Fracture Energy Levels..........................................................5.2-41 5.2.4.3 Operating Limitations During Startup and Shutdown................................5.2-43 5.2.4.4 Compliance With Reactor Vessel Materials Surveillance Program Requirements.............................................................................5.2-44 5.2.4.5 Reactor Vessel Annealing........................................................................5.2-44
5.2.5 Austenitic Stainless Steel.........................................................................5.2-44
5.2.5.1 Cleaning and Contamination Protection Procedures................................5.2-44 5.2.5.2 Solution Heat Treatment Requirements...................................................5.2-46 5.2.5.3 Material Inspection Program.....................................................................5.2-46 5.2.5.4 Unstabilized Austenitic Stainless Steels...................................................5.2-46 5.2.5.5 Avoidance of Sensitization........................................................................5.2-46 5.2.5.6 Retesting Unstabilized Austenitic Stainless Steels Exposed to Sensitizing Temperatures.........................................................................5.2-48 5.2.5.7 Control of Delta Ferrite.............................................................................5.2-48
5.2.6 Pump Flywheel.........................................................................................5.2-49
5.2.6.1 Compliance with NRC Regulatory Guide 1.14..........................................5.2-49
5.2.7 Reactor Coolant Pressure Boundary (RCPB) Leakage Detection Systems....................................................................................................5.2-50
5.2.7.1 Leakage Detection Methods.....................................................................5.2-50 FNP-FSAR-5 TABLE OF CONTENTS
Page 5-iii REV 21 5/08 5.2.7.2 Indication in Control Room ......................................................................5.2-54 5.2.7.3 Limits for Reactor Coolant Leakage.........................................................5.2-54 5.2.7.4 Unidentified Leakage ...............................................................................5.2-54 5.2.7.5 Maximum Allowable Total Leakage..........................................................5.2-56 5.2.7.6 Differentiation Between Identified and Unidentified Leaks ......................5.2-56 5.2.7.7 Sensitivity and Operability Tests...............................................................5.2-56
5.2.8 Inservice Inspection Program ..................................................................5.2-56
5.2.8.1 Provisions for Access to Reactor Coolant System Boundary...................5.2-56 5.2.8.2 Equipment for Inservice Inspections.........................................................5.2-58 5.2.8.3 Recording and Comparing Data ..............................................................5.2-58 5.2.8.4 Reactor Vessel Acceptance Standards...................................................5.2-58 5.2.8.5 Coordination of Inspection Equipment with Access Provisions................5.2-59 5.2.8.6 Preservice and Inservice Inspection and Inservice Testing Programs ....5.2-59 5.2.8.7 Ultrasonic Calibration Blocks....................................................................5.2-62
5.2.9 Loose Parts Monitoring Program (Metal Impact Monitor System)............5.2-62
5.3 THERMAL HYDRAULIC SYSTEM DESIGN..............................................................5.3-1 5.3.1 Analytical Methods and Data......................................................................5.3-1 5.3.2 Operating Restrictions on Pumps...............................................................5.3-1 5.3.3 Boiling Water Reactor (BWR).....................................................................5.3-1 5.3.4 Temperature-Power Operating Map...........................................................5.3-1 5.3.5 Load Following Characteristics ..................................................................5.3-1 5.3.6 Transient Effects.........................................................................................5.3-1 5.3.7 Thermal and Hydraulic Characteristics Summary Table............................5.3-2
5.4 REACTOR VESSEL AND APPURTENANCES.........................................................5.4-1 5.4.1 Design Bases..............................................................................................5.4-1
5.4.1.1 Codes and Specifications...........................................................................5.4-1 5.4.1.2 Design Transients.......................................................................................5.4-1 5.4.1.3 Protection Against Nonductile Failure.........................................................5.4-2 5.4.1.4 Inspection...................................................................................................5.4-2
FNP-FSAR-5 TABLE OF CONTENTS
Page 5-iv REV 21 5/08 5.4.2 Description..................................................................................................5.4-2
5.4.2.1 Fabrication Processes................................................................................5.4-3 5.4.2.2 Protection of Closure Studs........................................................................5.4-6
5.4.3 Evaluation...................................................................................................5.4-6
5.4.3.1 Steady-State Stresses................................................................................5.4-6 5.4.3.2 Fatigue Analysis Based on Transient Stresses..........................................5.4.6 5.4.3.3 Thermal Stresses Caused By Gamma Heating..........................................5.4-6 5.4.3.4 Thermal Stresses Caused By Loss-of-Coolant Accident............................5.4-6 5.4.3.5 Heatup and Cooldown................................................................................5.4-8 5.4.3.6 Irradiation Surveillance Program................................................................5.4-8 5.4.3.7 Capability for Annealing the Reactor Vessel............................................5.4-16
5.4.4 Tests and Inspection ................................................................................5.4-16
5.4.4.1 Ultrasonic Examinations...........................................................................5.4-16 5.4.4.2 Penetrant Examinations ...........................................................................5.4-16 5.4.4.3 Magnetic Particle Examination.................................................................5.4-17 5.4.4.4 Inservice Inspection .................................................................................5.4-17 5.4.4.5 Inspection of Rod Cluster Control Assemblies (RCCAs)..........................5.4-18
5.5 COMPONENT AND SUBSYSTEM DESIGN .............................................................5.5-1 5.5.1 Reactor Coolant Pumps..............................................................................5.5-1
5.5.1.1 Design Bases .............................................................................................5.5-1 5.5.1.2 Design Description .....................................................................................5.5-1 5.5.1.3 Design Evaluation.......................................................................................5.5-3 5.5.1.4 Tests and Inspections.................................................................................5.5-7
5.5.2 Steam Generator........................................................................................5.5-7
5.5.2.1 Design Bases..............................................................................................5.5-7 5.5.2.2 Design Description......................................................................................5.5-8 5.5.2.3 Design Evaluation.......................................................................................5.5-9 5.5.2.4 Tests and Inspections...............................................................................5.5-11
FNP-FSAR-5 TABLE OF CONTENTS
Page 5-v REV 21 5/08 5.5.3 Reactor Coolant Piping.............................................................................5.5-12
5.5.3.1 Design Bases............................................................................................5.5-12 5.5.3.2 Design Description....................................................................................5.5-13 5.5.3.3 Design Evaluation.....................................................................................5.5-16 5.5.3.4 Tests and Inspection.................................................................................5.5-17
5.5.4 Main Steam Line Flow Restrictions..........................................................5.5-17
5.5.4.1 Design Basis.............................................................................................5.5-17 5.5.4.2 Description................................................................................................5.5-18 5.5.4.3 Evaluation.................................................................................................5.5-18 5.5.4.4 Tests and Inspections...............................................................................5.5-18
5.5.5 Main Steam Line Isolation System...........................................................5.5-18
5.5.5.1 Design Bases............................................................................................5.5-19 5.5.5.2 System Description...................................................................................5.5-19 5.5.5.3 Design Evaluation.....................................................................................5.5-20 5.5.5.4 Tests and Inspections...............................................................................5.5-20
5.5.6 Reactor Core Isolation Cooling System....................................................5.5-21 5.5.7 Residual Heat Removal System...............................................................5.5-21
5.5.7.1 Design Bases............................................................................................5.5-21 5.5.7.2 System Description...................................................................................5.5-22 5.5.7.3 Design Evaluation.....................................................................................5.5-26 5.5.7.4 Tests and Inspections...............................................................................5.5-28
5.5.8 Reactor Coolant Cleanup System............................................................5.5-28 5.5.9 Main Steam Line and Feedwater Piping...................................................5.5-29 5.5.10 Pressurizer................................................................................................5.5-29
5.5.10.1 Design Bases............................................................................................5.5-29 5.5.10.2 Design Description....................................................................................5.5-30 5.5.10.3 Design Evaluation.....................................................................................5.5-31 5.5.10.4 Tests and Inspections...............................................................................5.5-34
5.5.11 Pressurizer Relief Tank............................................................................5.5-34
5.5.11.1 Design Bases ...........................................................................................5.5-34 5.5.11.2 Design Description ...................................................................................5.5-35 FNP-FSAR-5 TABLE OF CONTENTS
Page 5-vi REV 21 5/08 5.5.11.3 Design Evaluation.....................................................................................5.5-35
5.5.12 Valves ......................................................................................................5.5-36
5.5.12.1 Design Bases ...........................................................................................5.5-36 5.5.12.2 Design Description ...................................................................................5.5-36 5.5.12.3 Design Evaluation.....................................................................................5.5-37 5.5.12.4 Tests and Inspections...............................................................................5.5-37
5.5.13 Safety and Relief Valves ..........................................................................5.5-38
5.5.13.1 Design Bases ...........................................................................................5.5-38 5.5.13.2 Design Description ...................................................................................5.5-38 5.5.13.3 Design Evaluation.....................................................................................5.5-39 5.5.13.4 Tests and Inspections...............................................................................5.5-39
5.5.14 Component Supports ...............................................................................5.5-39
5.5.14.1 Description................................................................................................5.5-40 5.5.14.2 Evaluation ................................................................................................5.5-41 5.5.14.3 Tests and Inspections...............................................................................5.5-42
5.5.15 Reactor Vessel Head Vent System..........................................................5.5-42
5.5.15.1 Design Basis ............................................................................................5.5-42 5.5.15.2 System Description ..................................................................................5.5-42 5.5.15.3 Design Evaluation.....................................................................................5.5-43 5.5.15.4 Tests and Inspections...............................................................................5.5-44
5.6 INSTRUMENTATION APPLICATION........................................................................5.6-1
FNP-FSAR-5 LIST OF TABLES
5-vii REV 21 5/08 5.1-1 System Design and Operating Parameters
5.2-1 Hardship Exceptions to 10 CFR 50.55a
5.2-2 Summary of Reactor Coolant System Design Transients
5.2-2a Component Cyclic or Transient Limits
5.2-3 Load Combinations and Operating Conditions
5.2-4 Loading Conditions and Stress Limits: Class 1 Components
5.2-5 Loading Conditions and Stress Limits: Nuclear Power Piping 5.2-6 Faulted Condition Stress Limits for Class 1 Components
5.2-7 Allowable Stresses for Primary Equipment Supports
5.2-8 Active and Inactive Valves in the Reactor Coolant System Pressure Boundary
5.2-9 Stresses Caused by Maximum Steam Generator Tubesheet Pressure Differential (2485 psig)
5.2-10 Steam Generator Primary-Secondary Boundary Components
5.2-11 Steam Generator Primary-Secondary Boundary Components
through 5.2-13
5.2-14 51,500 Square Foot Steam Generator Usage Factors (Individual Transients) Primary and Secondary Boundary Components
5.2-15 51,500 Square Foot Steam Generator Usage Factors (Individual Transients) Center of Tubesheet
5.2-16 Tubesheet Stress Analysis Results for 51,500 Square Foot Steam Generators
5.2-17 Limit Analysis Calculation Results-Tables of Strains, Limit Pressures, and Fatigue Evaluations for 51,500 Square Foot Steam Generator
5.2-18 Relief Valve Discharge to the Pressurizer Relief Tank
5.2-19 Reactor Coolant System Design Pressure Settings (psig)
FNP-FSAR-5 LIST OF TABLES
5-viii REV 21 5/08 5.2-20 Reactor Coolant System Boundary Materials Class 1 Primary Components
5.2-21 Typical Reactor Coolant System Boundary Materials Auxiliary Components
5.2-22 Reactor Coolant Water Chemistry Specification
5.2-23 Materials for Reactor Vessel Internals for Emergency Core Cooling
5.2-24 Unit 1 Reactor Vessel Toughness Properties
5.2-25 Unit 2 Reactor Vessel Toughness Data
5.2-26 Faulted Condition Loads for the Reactor Coolant Pump Foot
5.2-27 Reactor Coolant Pump Outlet Nozzle Faulted Condition Loads
5.2-28 Steam Generator Lower Support Member Stresses
5.2-29 Steam Generator Upper Support Member Stresses
5.2-30 Reactor Coolant Pump Support Member Stresses
5.2-31 Pressurizer Upper Support Member Stresses
5.2-32 CRDM Heat Adaptor Bending Moments
5.2-33 Farley Nuclear Plant Unit 2 Preservice Inspection Program ASME Code Class 1 Components
5.2-34 Farley Nuclear Plant Unit 2 Preservice Inspection Program ASME Code Class 2 Components
5.2-35 Type B-4 Weld Wire and Linde 0091 Flux Tests
5.2-36 Farley Nuclear Plant Unit 2 Lower Shell Course Charpy V Notch Data
5.2-37 Farley Nuclear Plant Unit 2 Intermediate Shell Course Charpy V Notch Data
5.2-38 Farley Nuclear Plant Unit 2 Nozzle Shell Course Charpy V Notch Data
5.2-39 Steam Generator Pressurizer Fracture Toughness Properties
5.2-40 Load Combinations and Acceptance Criteria for Pressurizer and Relief Valve Piping - Upstream of Valves - Class 1 Piping FNP-FSAR-5 LIST OF TABLES
5-ix REV 21 5/08 5.2.41 Load Combinations and Acceptance Criteria for Pressurizer and Relief Valve Piping - Downstream of Valve - NNS Piping 5.2.42 Safety Line Pipe Stress and Strain Summary for Emergency Conditions
5.2.43 Farley Nuclear Plant TMI Action NUREG-0737.11.D.1 Units 1 and 2 PSARV Line Pipe Supports Anchor Bolt Data for Supports with Factor of Safety F.S. < 4
5.3-1 Natural Circulation Reactor Coolant Flow Versus Reactor Power
5.4-1 Reactor Vessel Design Parameters
5.4-2 Reactor Vessel Quality Assurance Program
5.4-3 Identification of Unit No. 1 Reactor Vessel Beltline Region Base Material
5.4-4 Predicted End of License (54 EFPY) Upper Shelf Energy Values - Farley Unit No. 1 Reactor Vessel Beltline Plates
5.4-5 Identification of Unit No. 1 Reactor Vessel Beltline Region Weld Metal
5.4-6 Predicted End of License (54 EFPY) Upper Shelf Energy Values - Farley Unit No. 1 Reactor Vessel Beltline Welds
5.4-7 Identification of Unit No. 2 Reactor Vessel Beltline Region Base Material
5.4-8 Predicted End of License (54 EFPY) Upper Shelf Energy Values - Farley Unit No. 2 Reactor Vessel Beltline Plates
5.4-9 Identification of Unit No. 2 Reactor Vessel Beltline Region Weld Metal
5.4-10 Predicted End of License (54 EFPY) Upper Shelf Energy Values - Farley Unit No. 2 Reactor Vessel Beltline Welds
5.4-11 Surveillance Material - Beltline Location and Fabrication History
5.4-12 Surveillance Material Chemical Composition
5.5-1 Reactor Coolant Pump Design Parameters
5.5-2 Reactor Coolant Pump Quality Assurance Program
5.5-3 Steam Generator Design Data
FNP-FSAR-5 LIST OF TABLES
5-x REV 21 5/08 5.5-4 Steam Generator Quality Assurance Program
5.5-5 Reactor Coolant Piping Design Parameters
5.5-6 Reactor Coolant Piping Quality Assurance Program
5.5-7 Design Bases for Residual Heat Removal System Operation
5.5-8 Residual Heat Removal System Component Data
5.5-9 Pressurizer Design Data
5.5-10 Pressurizer Quality Assurance Program
5.5-11 Pressurizer Relief Tank Design Data
5.5-12 Reactor Coolant System Boundary Valve Design Parameters
5.5-13 Reactor Coolant System Valv es Quality Assurance Program
5.5-14 Pressurizer Valves Design Parameters
5.5-15 Main Steam Valve Design Parameters - Main Steam Isolation Valves
5.5-16 Reactor Vessel Head Vent System Equipment Design Parameters
FNP-FSAR-5 LIST OF TABLES
5-xi REV 21 5/08 5.1-1 Pump Head - Flow Characteristics
5.2-1 Primary - Secondary Boundary Components Shell Locations of Stress Investigations
5.2-2 Primary and Secondary Hydrostatic Test Stress History for the Center Hole Location
5.2-3 Plant Heatup and Loading Operational Transients (With Steady-State Plateau)
Stress History for the Hot Side Center Hole Location
5.2-4 Large Step Load Decrease and Loss of Flow Stress History for the Hot Side Center Hole Location
5.2-5 Reactor Coolant Loop/Supports Syst em Dynamic - Structural Model
5.2-6 STHRUST RCL Model Showing Hydraulic Force Locations
5.2-7 (Deleted)
5.2-8 (Deleted)
5.2-9 (Deleted)
5.2-10 (Deleted)
5.2-11 K ID Lower Bound Fracture Toughness A533V (Reference WCAP 7623) Grade B Class 1
5.2-12 (Deleted)
5.2-13 Tool Details (Vessel Scanner) (Deleted)
5.2-14 Tool Details (Nozzle and Flange Scanner) (Deleted)
5.2-15 Sample Weld Data Sheet
5.2-16 Pressurizer Safety Line Structural Model
5.2-20 Reactor Coolant Pump Casing With Support Feet
5.2-21 Bolt Hold Radial Centerline
5.2-22 Nonlinear CRDM Center Row Model
5.4-1 Surveillance Capsule Elevation View FNP-FSAR-5 LIST OF TABLES
5-xii REV 21 5/08 5.4-2 Surveillance Capsule Plan View
5.4-3 Identification and Location of Farley Unit No. 1 Reactor Vessel Beltline Region Material
5.4-4 Identification and Location of Farley Unit No. 2 Reactor Vessel Beltline Region Material
5.5-1 Reactor Coolant Controlled Leakage Pump
5.5-2 Reactor Coolant Pump Performance Curve
5.5-3 Reactor Coolant Pump Spool Piece and Motor Support Stand
5.5-4 Steam Generator
5.5-5 Steam Generator Flow Limiting Device
5.5-6 Pressurizer
5.5-7 Reactor Vessel Supports
5.5-8 Dry Containment Steam Generator Supports
5.5-9 Reactor Coolant Pump Supports
5.5-10 Pressurizer Supports
5.5-11 (Deleted)
5.5-12 CRDM Seismic Support Platform Pipe Support Clamp
5.5-13 Sideview RVHVS and Supports FNP-FSAR-5
5.1-1 REV 25 4/14 5.0 REACTOR COOLANT SYSTEM AND CONNECTED SYSTEM
5.1
SUMMARY
DESCRIPTION The reactor coolant system (RCS) shown on drawings D-175037, sheet 1, D-205037, sheet 1, D-175037, sheet 2, D-205037, sheet 2, D-175037, sheet 3, and D-205037, sheet 3, consist of
similar heat transfer loops connected in parallel to the reactor pressure vessel. Each loop
contains a reactor coolant pump, steam generator, and associated piping and valves. In
addition, the system includes a pressurizer, a pressurizer relief tank, interconnecting piping, and
instrumentation necessary for operational control. All of the above components are located in
the containment building.
During operation, the reactor coolant system transfers the heat generated in the core to the
steam generators, where steam is produced to drive the turbine generator. Borated, demineralized water is circulated in the reactor coolant system at a flowrate and temperature
consistent with achieving the reactor core thermal hydraulic performance. The water also acts
as a neutron moderator and reflector, and as a solvent for the neutron absorber used in
chemical shim control.
The reactor coolant system pressure boundary provides a barrier against the release of
radioactivity generated within the reactor, and is designed to ensure a high degree of integrity
throughout the life of the plant.
Reactor coolant system pressure is controlled by the pressurizer, where water and steam are maintained in equilibrium by electrical heaters and water sprays. Steam can be formed (by the
heaters) or condensed (by the pressurizer spray) to minimize pressure variations caused by
contraction and expansion of the reactor coolant. Spring-loaded safety valves and
power-operated relief valves are mounted on the pressurizer and discharge to the pressurizer
relief tank, where the steam is condensed and cooled by mixing with water.
The extent of the reactor coolant system is defined as:
A. The reactor vessel, including control rod drive mechanism housings.
B. The reactor coolant side of the steam generators.
C. Reactor coolant pumps.
D. A pressurizer attached to one of the reactor coolant loops.
E. Safety and relief valves.
F. The interconnecting piping, valves, and fittings between the principal components listed above.
G. The piping, fittings, and valves leading to connecting auxiliary or support systems up-to-and-including the second isolation valve (from the high-pressure side) on
each line.
FNP-FSAR-5
5.1-2 REV 25 4/14 Reactor Coolant System Components A. Reactor Vessel
The reactor vessel is cylindrical, with a welded hemispherical bottom head and a removable, flanged, and gasketed hemispherical upper head. The vessel
contains the core, core supporting structures, control rods, and other parts
directly associated with the core.
The vessel has inlet and outlet nozzles located in a horizontal plane just below the reactor vessel flange, but above the top of the core. Coolant enters the
vessel through the inlet nozzles and flows down the core barrel vessel wall
annulus, turns at the bottom, and flows up through the core to the outlet nozzles.
The steam generators are vertical shell and U-tube evaporators with integral moisture separating equipment. The reactor coolant flows through the inverted
U-tubes, entering and leaving through the nozzles located in the hemispherical
bottom head of the steam generator. Steam is generated on the shell side and
flows upward through the moisture separators to the outlet nozzle at the top of
the vessel.
C. Reactor Coolant Pumps
The reactor coolant pumps are identical, single-speed, centrifugal units driven by air-cooled, three-phase induction motors. The shaft is vertical with the motor
mounted above the pumps. A flywheel on the shaft above the motor provides
additional inertia to extend pump coastdown. The inlet is at the bottom of the
pump; discharge is on the side.
D. Piping
The reactor coolant loop piping is specified in sizes consistent with system requirements.
The hot leg inside diameter is 29 in. and the cold leg return line to the reactor vessel is 27-1/2 in. The piping between the steam generator and the pump
suction is increased to 31 in. in diameter to reduce pressure drop and improve
flow conditions to the pump suction.
E. Pressurizer
The pressurizer is a vertical, cylindrical vessel with hemispherical top and bottom heads. Electrical heaters are installed through the bottom head of the vessel
while the spray nozzle, relief, and safety valve connections are located in the top
head of the vessel.
FNP-FSAR-5
5.1-3 REV 25 4/14 F. Safety and Relief Valves
The pressurizer safety valves are of the totally enclosed pop-type. The valves are spring-loaded and self-activated, with back-pressure compensation. The
power-operated relief valves limit system pressure for large power mismatch.
They are operated automatically or by remote manual control. Remotely operated valves are provided to isolate the inlet to the power-operated relief
valves if excessive leakage occurs.
Reactor Coolant System Performance Characteristics Tabulations of important design and performance char acteristics of the reactor coolant system are provided in table 5.1-1.
Reactor Coolant Flow The reactor coolant flow, a major parameter in the design of the system and its components, is established with a detailed design procedure supported by operating plant performance data, by
pump model tests and analyses, and by pressure-drop tests and analyses of the reactor vessel
and fuel assemblies. Data from all operating plants have indicated that the actual flow has been
well above the flow specified for the thermal design of the plant. By applying the design
procedure described below, it is possible to specify the expected operating flow with reasonable accuracy.
Three reactor coolant flowrates are identified for the various plant design considerations. The
definitions of these flows are presented in the following paragraphs, and the applications of the
definitions are illustrated by the system and pum p hydraulic characteristics on figure 5.1-1.
Best Estimate Flow The best estimate flow is the most likely value for the actual plant operating condition. This flow
is based on the best estimate of the reactor vessel, steam generator and piping flow resistance, and on the best estimate of the reactor coolant pump head, with no uncertainties assigned to either the system flow resistance or the pump head. System pressure losses based on best
estimate flow are presented in table 5.1-1. Although the best estimate flow is the most likely
value to be expected in operation, more conservative flowrates are applied in the thermal and
mechanical designs.
Thermal Design Flow Thermal design flow is the basis for the reactor core thermal performance, the steam generator
thermal performance, and the nominal plant parameters used throughout the design. To
provide the required margin, the thermal design flow accounts for the uncertainties in reactor
vessel, steam generator and piping flow resistances, reactor coolant pump head, and the
methods used to measure flowrate. The combination of these uncertainties is equivalent to
increasing the best estimate reactor coolant system flow resistance by approximately 15 percent.
[HISTORICAL][
The intersection of this conservative flo w resistance with the best estimate pump curve, as shown in figure 5.1-1, established the original/plant thermal design flow. This procedure provides a flow margin for thermal design of approximately 4 percent.
] For this plant, changes FNP-FSAR-5
5.1-4 REV 25 4/14 subsequent to the original specification of thermal design flow have resulted in additional
margin. The thermal design flow is confirmed when the plant performs precision RCS flow measurements at the beginning of each cycle. Tabulations of important design parameters
based on the thermal design flow are provided in table 5.1-1.
Mechanical Design Flow Mechanical design flow is the conservatively high flow used in the mechanical design of the reactor vessel internals, fuel assemblies, and other system components.
[HISTORICAL][
To ensure that a conservatively high flow is specified, th e original plant mechanical design flow was set at least 4% higher than the original best estimate flow.
] The mechanical design flow is 101,800 gpm/loop, which is 5.5% above the current best estimate flow of 97,800 gpm/loop with 0%
steam generator tube plugging and thimble plugs removed after best estimate flow is adjusted to account for measured RCS flow. This best estimate flow is based on Unit 2, since it yields the minimum margin to mechanical design flow.
Pump overspeed, because of a turbine generator overspeed of 20 percent, results in a peak
reactor coolant flow of 120 percent of the mechanical design flow. The overspeed condition is
applicable only to operating conditions when the reactor and turbine generator are at power.
Interrelated Performance and Safety Functions The interrelated performance and safety functions of the reactor coolant system and its major
components are listed below:
A. The reactor coolant system provides sufficient heat transfer capability to transfer the heat produced during power operation and when the reactor is subcritical, including the initial phase of plant cooldown, to the steam and power conversion system.
B. The system provides sufficient heat transfer capability to transfer the heat produced during the subsequent phase of plant cooldown and cold shutdown to
the residual heat removal (RHR) system.
C. The system heat removal capability under power operation and normal operational transients, including the transition from forced to natural circulation, will ensure no fuel damage within the operating bounds permitted by the reactor
control and protection systems.
D. The reactor coolant system provides the water used as the core neutron moderator and reflector and as a solvent for chemical shim control.
E. The system maintains the homogeneity of soluble neutron poison concentration and rate of change of coolant temperature so that uncontrolled reactivity changes
do not occur.
F. The reactor vessel is an integral part of the reactor coolant system pressure boundary and is capable of accommodating the temperatures and pressures FNP-FSAR-5
5.1-5 REV 25 4/14 associated with the operational transients. The reactor vessel functions to
support the reactor core and control rod drive mechanisms (CRDM).
G. The pressurizer maintains the system pressure during operation and limits pressure transients. During the reduction or increase of plant load, reactor
coolant volume changes are accommodated in the pressurizer via the surge line.
H. The reactor coolant pumps supply the coolant flow necessary to remove heat from the reactor core and transfer it to the steam generators.
I. The steam generators provide high-quality steam to the turbine. The tube and tube sheet boundary are designed to prevent the transfer of activity generated
within the core to the secondary system.
J. The reactor coolant system piping serves as a boundary for containing the coolant under operating temperature and pressure conditions and for limiting
leakage (and activity release) to the containment atmosphere. The reactor
coolant system piping contains demineralized, borated water, which is circulated
at the flowrate and temperature consistent with achieving the reactor core
thermal and hydraulic performance.
Interlocks on critical motor-operated valves are discussed in subsection 7.6.2 and paragraph
6.3.2.15.
5.1.1 SCHEMATIC FLOW DIAGRAM The reactor coolant system is shown on drawings D-175037, sheet 1, D-205037, sheet 1, D-175037, sheet 2, D-205037, sheet 2, D-175037, sheet 3, and D-205037, sheet 3, and
principal pressures, temperatures, flowrates, and coolant volume data under normal
steady-state, full-power operating conditions are provided in table 5.1-1.
5.1.1.1 System Operation Brief descriptions of normal, anticipated system operations are provided below. These
descriptions cover plant startup, power generation, hot shutdown, cold shutdown and refueling.
5.1.1.1.1 Plant Startup Plant startup encompasses the operations which bring the reactor plant from cold shutdown to
no-load power operating temperature and pressure. Before plant startup, the reactor coolant
loops and pressurizer are filled completely, by the use of the charging pumps, with water
containing the cold shutdown concentration of boron. The loops are vented using either the
Reactor Coolant Vacuum Refill System (RCV RS) or the dynamic venting process. The secondary side of the steam generator is filled to normal startup level with water which meets
the steam plant water chemistry requirements.
FNP-FSAR-5
5.1-6 REV 25 4/14 If the RCVRS is used, air is removed from t he RCS by a skid-mounted vacuum pump system.
The RCVRS is connected to the RCS via a special connection to the pressurizer relief tank (PRT) inlet line. The RCS evacuation path includes the pressurizer surge line (while at midloop
conditions), the reactor vessel head vent paths, and the pressurizer spray line (once the surge
line is submerged). Transportation of the air from the hot legs to the cold legs occurs through
the air gap between the internal and external hot leg reactor vessel nozzles and the core
bypass flow nozzles.
Initial conditions are as follows: the RCS level is at midloop and the PRT level is below the
sparging header. The vacuum pump skid suction hose is connected to the PRT inlet line
connection. The RHR flow is adjusted to prevent vortexing and to ensure adequate NPSH. The
air evacuation path is established by opening the reactor vessel head vent valves, the
pressurizer spray valves, the PORV block valves and the PORVs.
Prior to starting the air evacuation via the RCVRS, letdown flow and charging flow are adjusted
to maintain a constant VCT level with RCP seal injection in service. The RCVRS is then used to
pull the air from the RCS via the connection to the PRT inlet line. The RCS is filled via one
charging path while maintaining the vacuum in the RCS. Once the RCS is filled to a pressurizer
level approximately equal to the steam generator tube elevation, the RCS vacuum is broken.
Charging is continued until a level increase is detected in the PRT. Finally, the PORVs, pressurizer spray valves and reactor vessel head vent valves are closed. This completes the
RCS filling and venting operation.
If the RCVRS is not used, the RCS is pressurized, by use of the low pressure control valve and
one centrifugal charging pump, to obtain the required pressure drop across the number one
seal of the reactor coolant pumps. The pumps may then be operated intermittently to assist in
venting operations.
During operation of the reactor coolant pumps, one charging pump and the low pressure
letdown path from the residual heat removal loop to the chemical and volume control system (CVCS) are used to maintain the reactor coolant system pressure in an appropriate range.
Plant operating experience and instrument inaccuracy are used to establish a pressure range
which ensures that all RCP support conditions are met and that the LTOP relief valves are not
challenged during RCP start, the ensuing transient, and any subsequent operation. The
fracture prevention temperature limitations of the reactor vessel impose an upper limit of approximately 450 psig. The charging pump supplies seal-injection water for the reactor
coolant pump shaft seals. A nitrogen atmosphere and normal operating temperature, pressure, and water level are established in the pressurizer relief tank.
Upon completion of venting, the reactor coolant system is pressurized, the reactor coolant
pumps are started, and the pressurizer heaters are energized to begin heating the reactor coolant. When the cold leg temperature reaches between 175-180
°F and the pressurizer temperature increases to the saturation temperature corresponding to a saturation pressure of about 375 psig, a steam bubble is formed in the pressurizer while the reactor coolant pressure
is maintained in an appropriate range. Plant operating experience and instrument inaccuracy
are used to establish a pressure range which ensures that all RCP support conditions are met
and that the LTOP relief valves are not challenged during RCP start, the ensuing transient, and
any subsequent operation. The pressurizer liquid level is reduced until the no-load power level
volume is established. During the initial heatup phase, hydrazine is added to the reactor FNP-FSAR-5
5.1-7 REV 25 4/14 coolant to scavenge the oxygen in the system. The heatup is not taken beyond 250°F until the oxygen level has been reduced to the specified level.
An alternative to water-solid operation to establish RCS pressure for RCP operation is the use
of a pressurizer steam bubble. In this case, the RCVRS is used to remove most of the system
air. Hydrazine is then added to the pressurizer via auxiliary spray to remove dissolved oxygen
from the pressurizer liquid. The pressurizer heaters are actuated to establish a steam bubble to
pressurize the RCS and RHR flow is reduced or bypassed to allow the RCS to heat up to 150-160°F. The combination of RCS letdown flow diversion to the recycle holdup tanks and RHR flow adjustment is used to maintain a constant pressurizer level as the RCS expands.
When the pressurizer pressure reaches the appropriate range, the RCPs are started to remove
the small volume of air trapped in the top of the steam generator tubes. Plant operating
experience and instrument inaccuracy are used to establish a pressure range which ensures
that all RCP support conditions are met and that the LTOP relief valves are not challenged
during RCP start, the ensuing transient, and any subsequent operation.
The VCT is then burped as required to reduce the oxygen in the gas space. Additional
hydrazine is then added by the normal charging flow path to reduce the RCS dissolved oxygen
concentration within Technical Requirements Manual limits before the RCS is allowed to heat up above 250
°F. The reactor coolant pumps and pressurizer heaters are used to raise the reactor coolant
temperature and pressure to normal operating levels.
As the reactor coolant temperature increases, the pressurizer heaters are manually controlled to
maintain adequate suction pressure for the reactor coolant pumps. When the normal operating
pressure of 2235 psig is reached, pressurizer heat and spray controls are transferred from
manual to automatic control.
5.1.1.1.2 Power Generation and Hot Shutdown Power generation includes steady-state operation, ramp changes not exceeding the rate of 5
percent of full power per minute, step changes of 10 percent of full power (not exceeding full
power), and step load changes with steam dump not exceeding the design step load decrease.
During power generation, reactor coolant system pressure is maintained by the pressurizer
controller at-or-near 2235 psig, while the pressurizer liquid level is controlled by the charging
letdown flow control of the chemical and volume control system.
When the reactor power level is less than 15 percent, the reactor power is controlled manually.
At power above 15 percent, the reactor control system controls automatically maintain an
average coolant temperature, consistent with the power relationships, by control rod movement.
During the hot shutdown operations, when the reactor is subcritical, the reactor coolant system
temperature is maintained by steam dump to the main condenser. This is accomplished by a
controller in the steam line, operating in the pressure control mode, which is set to maintain the
steam generator steam pressure. Residual heat from the core or operation of a reactor coolant
pump provides heat to overcome reac tor coolant system heat losses.
FNP-FSAR-5
5.1-8 REV 25 4/14 5.1.1.1.3 Plant Shutdown Plant shutdown is the operation which brings the reactor plant from no-load power operating
temperature and pressure to cold shutdown. Concentrated boric acid solution from the chemical and volume control system is added, as necessary, to the reactor coolant system to
increase the reactor coolant boron concentration to ensure adequate shutdown margin is
maintained as required by plant Technical Specifications. If the reactor coolant system is to be
opened during the shutdown, the hydrogen and fission gas in the reactor coolant is reduced by
degassing the coolant in the volume control tank.
Plant shutdown is accomplished in two phases; the first is by the combined use of the reactor
coolant system and steam systems, and the sec ond is by the residual heat removal system.
During the first phase of shutdown, residual core and reactor coolant heat is transferred to the
steam system via the steam generator. Steam fr om the steam generator is dumped to the main condenser. At least one reactor coolant pump is kept running to assure uniform reactor coolant
system cooldown. The pressurizer heaters are de-energized and spray flow is manually controlled to cool the pressurizer while maintaining the required reactor coolant pump suction
pressure.
When the reactor coolant temperature is below approximately 350°F and the pressure is in the
range of 400 to 450 psig, the second phase of shutdown commences with the operation of the
residual heat removal system.
When the reactor coolant temperature is below 200
°F, the pressurizer steam bubble is collapsed. One reactor coolant pump (either of those in a loop containing a pressurizer spray line) remains in service as the coolant temperature approaches 160°F. One or more RCPs may
remain in service after the steam bubble is collapsed to facilitate mixing of the RCS.
Pressurizer cooldown is continued by initiati ng auxiliary spray flow from the chemical and volume control system. Plant shutdown continues until the reactor coolant temperature is 140°F
or less.
5.1.1.1.4 Refueling Before removing the reactor vessel head for refueling, the system temperature has been
reduced to 140°F or less and hydrogen and fission product levels are reduced. A clear plastic
tube is attached to one of the reactor coolant loops to indicate when the water has been drained
below the reactor vessel head vent. Draining continues until the water level is below the reactor
vessel flange. The vessel head is then raised. Upon completion of refueling, the system is
refilled for plant startup.
5.1.2 PIPING AND INSTRUMENTATION DIAGRAM A piping and instrumentation diagram of the r eactor coolant system is shown on drawings D-175037, sheet 1, D-205037, sheet 1, D-175037, sheet 2, D-205037, sheet 2, D-175037, sheet
3, and D-205037, sheet 3. The diagrams show the extent of the systems located within the containment, and the points of separation between the reactor coolant system and the FNP-FSAR-5
5.1-9 REV 25 4/14 secondary (heat utilization) system. The isolation provided between the reactor coolant
pressure boundary and connected systems is discussed in subsection 6.2.4.
5.1.3 ELEVATION DRAWING Figures 1.2-6 and 1.2-7 are plant general arrangements which show the elevations and relative
locations of the major components in the reactor coolant loop.
REV 21 5/08
[PUMP HEAD - FLOW CHARACTERISTICS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.1-1
]
FNP-FSAR-5
5.2-1 REV 27 4/17 The reactor coolant system (RCS) boundary for Westinghouse (W) pressurized-water reactors (PWR) is defined as stated in American Nuclear Society document N18.2 "Nuclear Safety
Criteria for the Design of Stationary Pressurized-Water Reactor Plants", January 1972, paragraph 5.4.3.2. This definition of the RCS boundary is consistent with the definition of the
reactor coolant pressure boundary (RCPB) as defined in 10 CFR 50.2, part V, as applied to
codes and standards required by 10 CFR 50.55a (with appropriate footnotes), and ASME Section XI requirements for inservice inspection.
The RCS boundary is designed to accommodate the system pressures and temperatures
attained under all expected modes of plant operation, including all anticipated transients, and to
maintain the stresses within applicable stress limits.
The system is protected from overpressure
by means of pressure-relieving devices as required by applicable codes. Materials of
construction are specified to minimize corrosion and erosion and to provide a structural system
boundary throughout the life of the plant. Fracture prevention measures are taken to prevent
brittle fracture. Inspections in accordance with applicable codes and provisions are made for
surveillance of critical areas to enable periodic assessment of the boundary integrity, as
described in subsection 5.2.8.
The performance objectives of the RCS for normal operation are described in section 5.1. The
performance objectives for upset and faulted conditions are given in subsection 5.2.1 above.
No transient is classified as an emergency condition.
Equipment code and classification lists for the components within the reactor coolant system
boundary are given in table 3.2-1.
The RCS, in conjunction with the reactor control and protection systems, is designed to
maintain the reactor coolant at conditions of temperature, pressure, and flow adequate to protect the core from damage. The design requirement for safety is to prevent conditions of
high power, high reactor-coolant temperature, or low reactor-coolant pressure or combinations
of these which could result in a departure from nucleate boiling ratio (DNBR) less than the
safety analysis limit.
The RCS is designed to provide controlled changes in the boric acid concentration and the
reactor coolant temperature. The reactor coolant is the core moderator, reflector, and solvent
for the chemical shim. As a result, changes in coolant temperature or boric acid concentration
affect the reactivity level in the core.
The following design bases have been selected to ensure that the uniform RCS boron
concentration and temperature will be maintained:
FNP-FSAR-5
5.2-2 REV 27 4/17 A. Coolant flow is provided by either a reactor coolant pump or a residual heat removal (RHR) pump to ensure uniform mixing whenever the boron
concentration is decreased.
B. The design arrangement of the RCS eliminates deadended sections and other areas of low coolant flow in which nonhomogeneities in coolant temperature or
boron concentration could develop.
C. The RCS is designed to operate within the operating parameters, particularly the coolant temperature change limitations.
The design pressure for the RCS is 2485 psig, except for the pressurizer relief line from the
safety valve to the pressurizer relief tank, which is 600 psig, and the pressurizer relief tank, which is 100 psig. For components with design pressures of 2485 psig, the normal operating
pressure is 2235 psig. The design temperature for the RCS is 650°F, except for the pressurizer
and its surge line, which are designed for 680°F, and the pressurizer relief line from the safety
valve to the pressurizer relief tank, which is designed for 600°F. The seismic loads for Farley
Nuclear Plant (FNP) are given in section 3.7.
Reactor coolant system and component test pressures are discussed in paragraph 5.2.1.5.
The components of the RCPB are designed and fabricated in accordance with the rules of
10 CFR 50, Section 50.55a, Codes and Standards, except as noted in table 5.2-1. This
table lists the components, the code to which the components were designed and fabricated, the code required by Section 50.55a based on the August 1972 construction permit date, and
the differences between the code requirements as designed and fabricated and as required by
Section 50.55a.
All of the exceptions listed result from the issuance of a construction permit being delayed
beyond June 30, 1972, because of the extensive per iod for environmental review of the Farley
project after the safety evaluation was essentially completed. Total time from filing the
construction permit application was 34 months, which was substantially beyond the period
anticipated at the times that the components were purchased. Efforts were made to upgrade the
components beyond the codes listed in the Prelimi nary Safety Analysis Report (PSAR) in order to comply with Section 50.55a; those areas in which the efforts were not completely successful
are listed in table 5.2-1.
The ASME Code case interpretations that may have been applied to the components of the
RCS boundary are tabulated in table 3.2-5.
FNP-FSAR-5
5.2-3 REV 27 4/17 The following five ASME operating conditions are considered in the design of the RCS.
A. Normal Conditions
Any condition in the course of startup, operation in the design power range, and hot standby and system shutdown other than upset, emergency, faulted, or
testing condition.
B. Upset Conditions
Any deviations from normal conditions anticipated to occur often enough that design should include a capability to withstand the conditions without operational
impairment. The upset conditions include those transients resulting from any
single operator error or control malfunction, transients caused by a fault in a
system component requiring its isolat ion from the system, and transients because of loss-of-load or power. Upset conditions include any abnormal
incidents not resulting in a forced outage and also forced outages for which the
corrective action does not include any repair of mechanical damage. The
estimated duration of an upset condition is included in the design specifications.
C. Emergency Conditions
Those deviations from normal conditions that require shutdown for correction of the conditions or repair of damage in the system. The conditions have a low
probability of occurrence, but are included to provide assurance that no gross
loss of structural integrity will result as a concomitant effect of any damage
developed in the system. The total number of postulated occurrences for such
events will not cause more than 25 stress cycles having an S A value greater than that for 10 6 cycles from the applicable fatigue design curves of the ASME Code Section III.
D. Faulted Conditions
Those combinations of conditions associated with extremely low-probability, postulated events whose consequences are such that the integrity and
operability of the nuclear energy system may be impaired to the extent that
considerations of public health and safety are involved. Such considerations
require compliance with safety criteria as may be specified by jurisdictional
authorities.
E. Testing Conditions
Testing conditions are those tests in addition to the hydrostatic or pneumatic tests permitted by the ASME Code Section III, including leak tests or subsequent
hydrostatic tests.
FNP-FSAR-5
5.2-4 REV 27 4/17 To provide the necessary high degree of integrity for the equipment in the RCS, the transient
conditions selected for equipment fatigue evaluation are based upon a conservative estimate of
the magnitude and frequency of the temperature and pressure transients resulting from various
operating conditions in the plant. To a large extent, the specific transient operating conditions to
be considered for equipment fatigue analyses are based upon engineering judgment and
experience. The transients selected are repres entative of operating conditions which prudently might be considered to occur during plant operation and are sufficiently severe or frequent to be
of possible significance to component cyclic behavior. The transients selected may be regarded
as a conservative representation of transients that, used as a basis for component fatigue
evaluation, provide confidence that the component is appropriate for its application over the design life of the plant. As required by Technical Specifications administrative controls, the
components identified in table 5.2-2a are designed and shall be maintained within the cyclic or
transient limits of table 5.2-2a. The Fatigue Monitoring Program, as described in chapter 18, subsection 18.3.2, will be used to monitor plant transients that are significant contributors to the
fatigue cumulative usage factor to ensure the design limit on fatigue usage is not exceeded
during the period of extended operation.
The following five transients are considered normal conditions:
A. Heatup and Cooldown
For design evaluation, the heatup and cooldown cases are represented by continuous heatup or cooldown at a rate of 100°F/h. The heatup occurs from
ambient to the no-load temperature and pressure condition and the cooldown
represents the reverse situation.
In actual practice, the rate of temperature change of 100°F/h will not usually be attained because of other limitations such as:
- 1. Criteria for prevention of nonductile failure, which establish maximum permissible temperature rates of change as a function of plant pressure
and temperature.
- 2. Slower initial heatup rates when using pumping energy only.
- 3. Interruptions in the heatup and cooldown cycles because of such factors as drawing a pressurizer steam bubble, rod withdrawal, sampling, water
chemistry, and gas adjustments.
B. Unit Loading and Unloading
The unit loading and unloading cases are conservatively represented by a continuous and uniform ramp power change of 5% min between 15% load and
full load. This load swing is the maximum possible consistent with operation with
automatic reactor control. The reactor coolant temperature varies with load as
prescribed by the temperature control system.
FNP-FSAR-5
5.2-5 REV 27 4/17 C. Step Increase and Decrease of Ten Percent
The 10-percent step change in load demand is a control transient assumed to be a change in turbine control valve opening that might be occasioned by disturbances in the electrical network into which the plant output is tied. The
reactor control system is designed to restore plant equilibrium without reactor trip, following a 10-percent step change in turbine load demand initiated from nuclear plant equilibrium conditions in the range between 15-percent and 100-percent full load, the power range for automatic reactor control. In effect, during
load change conditions, the reactor control system attempts to match turbine and
reactor outputs in such a manner that peak reactor coolant temperature is
minimized and reactor coolant temperature is restored to its programmed
setpoint, at a sufficiently slow rate to prevent excessive pressurizer pressure
decrease.
Following a step-load decrease in turbine load, the secondary-side steam pressure and temperature initially increase, since the decrease in nuclear power
lags behind the step decrease in turbine load. During the same increment of
time, the RCS average temperature and pressurizer pressure also initially
increase. Because of the power mismatch between the turbine and reactor and
the increase in reactor coolant temperat ure, the control system automatically inserts the control rods to reduce core power. With the load decrease, the
reactor coolant temperature is ultimately reduced from its peak value to a value
below its initial equilibrium value at the inception of the transient. The reactor
coolant average temperature setpoint change is made as a function of turbine
generator load as determined by first stage turbine pressure measurement. The
pressurizer pressure also decreases from its peak pressure value and follows the
reactor coolant decreasing temperature trend. At some point during the
decreasing pressure transient, the saturated water in the pressurizer begins to
flash, which reduces the rate of pressure decrease. Subsequently, the
pressurizer heaters come on to restore the plant pressure to its normal value.
Following a step-load increase in turbine load, the reverse situation occurs; i.e., the secondary-side steam pressure and temperature initially decrease and the
reactor coolant average temperature and pressure initially decrease. The control
system automatically withdraws the cont rol rods to increase core power. The
decreasing pressure transient is reversed by actuation of the pressurizer heaters
and eventually the system pressure is restored to its normal value. The reactor
coolant average temperature is raised to a value above its initial equilibrium
value at the beginning of the transient.
D. Large Step Decrease in Load
This transient applies to a step decrease in turbine load from full power of such magnitude that the resultant rapid increase in reactor coolant average
temperature and secondary-side steam pre ssure and temperature automatically initiates a secondary-side steam dump syst em that prevents a reactor shutdown or lifting of steam generator safety valves. Thus, when a plant is designed to
accept a step decrease of 95 percent from full power, it signifies that a steam FNP-FSAR-5
5.2-6 REV 27 4/17 dump system provides a heat sink to accept 85 percent of the turbine load. The
remaining 10 percent of the total step change is assumed by the rod control
system. If a steam dump system were not provided to cope with this transient, there would be such a large mismatch between what the turbine is demanding
and what the reactor is furnishing that a reactor trip and lifting of steam generator
safety valves would occur.
Although Farley has been designed for a 50-percent step change, the transient for the 95-percent step-load decrease is considered since it represents a more
severe condition than the lower percentages.
E. Steady-State Fluctuations
The reactor coolant average temperature, for purposes of design, is assumed to increase or decrease a maximum of 6°F in 1 min. The temperature changes are
assumed to be around the programmed value of Tavg , (Tavg + 3°F). The corresponding reactor coolant average pressure is assumed to vary accordingly.
The following six transients are considered upset conditions:
A. Loss of Load Without Immediate Turbine or Reactor Trip
This transient applies to a step decrease in turbine load from full power occasioned by the loss of turbine load without immediately initiating a reactor trip
and represents the most severe transient on the RCS. The reactor and turbine
eventually trip as a consequence of a high pressurizer level trip initiated by the
reactor trip system. Since redundant means of tripping the reactor are provided
as a part of the reactor protection system, transients of this nature are not
expected, but are included to ensure a conservative design.
B. Loss of Power
This transient applies to a blackout situation involving the loss of outside electrical power to the station with a reactor and turbine trip. Under these
circumstances, the reactor coolant pumps are deenergized and, following the
coastdown of the reactor coolant pumps, natural circulation builds up in the
system to some equilibrium value. This condition permits removal of core
residual heat through the steam generators which, at this time, are receiving
feedwater from the auxiliary feed system operating from diesel generator power.
Steam is removed for reactor cooldown through atmospheric relief valves
provided for this purpose.
C. Loss of Flow
This transient applies to a partial loss-of-flow accident from full power in which a reactor coolant pump is tripped out of service as a result of a loss of power to the
pump. The consequences of such an accident are a reactor and turbine trip, on
low reactor coolant flow, followed by automatic opening of the steam dump FNP-FSAR-5
5.2-7 REV 27 4/17 system and flow reversal in the affected loop. The flow reversal results in reactor
coolant at cold leg temperature being passed through the steam generator and
cooled still further. This cooler water then passes through the hot leg piping and
enters the reactor vessel outlet nozzles. The net result of the flow reversal is a
sizable reduction in the hot leg coolant temperature of the affected loop.
D. Reactor Trip From Full Power
A reactor trip from full power may occur for a variety of causes resulting in temperature and pressure transients in the RCS and in the secondary side of the
This is the result of continued heat transfer from the reactor coolant in the steam generator. The transient continues until the reactor coolant and steam generator
secondary side temperatures are in equilibrium at zero power conditions. A
continued supply of feedwater and controlled dumping of secondary steam
remove the core residual heat and prevent the steam generator safety valves
from lifting. The reactor coolant temperature and pressure undergo a rapid
decrease from full power values as the reactor trip system causes the control
rods to move into the core.
E. Inadvertent Pressurizer Auxiliary Spray Initiation
The inadvertent pressurizer auxiliary sp ray transient will occur if the auxiliary spray valve is opened inadvertently during normal operation of the plant. This
will introduce cold water into the pressurizer with a very sharp pressure decrease
as a result.
The temperature of the auxiliary spray water is dependent upon the performance of the regenerative heat exchanger. The most conservative case is that in which
the letdown stream is shut off and the charging fluid enters the pressurizer
unheated. Therefore, for design purposes, the temperature of the spray water is
assumed to be 100°F. The spray flowrate is assumed to be 200 gal/min.
The pressure decreases rapidly to the low pressure reactor trip point. At this pressure, the pressurizer low pressure reactor trip is assumed to be actuated.
This accentuates the pressure decrease until the pressure is finally limited to the
hot leg saturation pressure. At 5 min the spray is stopped and all the pressurizer
heaters return the pressure to 2250 psia. This transient is more severe on a
two-loop plant than on a three-loop plant; e.g., a bigger and more rapid pressure
decrease. Therefore, the transient for a two-loop plant is used as design basis
for the FNP.
For design purposes it is assumed that no temperature changes in the RCS will occur as a result of initiation of auxiliary spray except in the pressurizer.
FNP-FSAR-5
5.2-8 REV 27 4/17 F. Operating Basis Earthquake (OBE)
The earthquake loads are a part of the mechanical loading conditions specified in the equipment specifications. The origin of their determination is separate and
distinct from those transient loads resulting from fluid pressure and temperature.
Their magnitude however, is considered in the design analysis for comparison
with appropriate stress limits.
The following four transients are considered faulted conditions:
A. Reactor Coolant System Boundary Pipe Break
This accident involves the postulated rupture of a pipe belonging to the RCS boundary. It is conservatively assumed that the system pressure is reduced
rapidly and the emergency core cooling system (ECCS) is initiated to introduce
water into the RCS. The safety injection signal also will initiate a turbine and
The criteria for locating design basis pipe ruptures used in the design of the supports and restraints of the RCS in order to assure continued integrity of vital
components and engineered safety systems is given in section 3.6.
Analyses reported in reference 1 and service experiences show that the criteria given in section 3.6 offer a practical equivalent to ensure the same degree of
protection to public health and safety as postulating both longitudinal and
circumferential breaks at any location. Westinghouse nuclear steam supply
system (NSSS) piping and support components are designed to these criteria.
Protection criteria against dynamic effects associated with pipe breaks are covered in section 3.6.
B. Steam Line Break
For component evaluation, the following conservative conditions are considered:
- 1. The reactor is initially in hot, zero-power subcritical condition, assuming all rods in, except the most reactive rod, which is assumed to be stuck in
its fully withdrawn position.
- 2. A steam line break occurs inside the containment resulting in a reactor and turbine trip.
- 3. After the break the reactor coolant temperature cools down to 212°F.
- 4. The ECCS pumps restore the reactor coolant pressure.
The above conditions result in the most severe temperature and pressure variations which the component will encounter during a steam-break accident.
FNP-FSAR-5
5.2-9 REV 27 4/17 The dynamic reaction forces associated with circumferential steam line breaks will be considered in the design of supports and restraints in order to assure
continued integrity of vital components and engineered safety features.
Protection criteria against dynamic effects associated with pipe breaks are
covered in section 3.6.
C. Steam Generator Tube Rupture
This accident postulates the double-ended rupture of a steam generator tube resulting in a decrease in pressurizer level and RCS pressure. Reactor trip
occurs because of a safety injection signal on low pressurizer pressure. The
planned procedure for recovery from this accident calls for isolation of the steam
line leading from the affected steam generator (reference section 15.4).
Therefore, this accident results in a transient which is no more severe than that
associated with a reactor trip.
D. Safe Shutdown Earthquake (SSE)
The mechanical stress transient resulting from the safe shutdown earthquake (SSE) is considered on a component basis.
The above design conditions are given in the Equipment Specifications which are written in
accordance with the ASME Code.
The design transients and the number of cycles of each that are normally used for fatigue
evaluations are shown in table 5.2-2. In accordance with the ASME Boiler and Pressure Vessel
Code, faulted conditions are not included in fatigue evaluations.
Prior to plant startup the following tests are carried out:
A. Turbine Roll Test
This transient is imposed upon the plant during the hot functional test period for turbine cycle checkout. Reactor coolant pump power is used to heat the reactor
coolant to operating temperature and the steam generated is used to perform a
turbine roll test. However, the plant cooldown during this test exceeds the
100°F/h maximum rate.
B. Hydrostatic Test Conditions
The pressure tests are outlined below:
- 1. Primary-Side Hydrostatic Test Before Initial Startup
The pressure tests covered by this section include both shop and field hydrostatic tests which occur as a result of component or system testing.
This hydro test is performed, prior to initial fuel loading, at a water
temperature which is compatible with reactor vessel fracture prevention
criteria requirements and a maximum test pressure of 3107 psig, or 1.25 FNP-FSAR-5
5.2-10 REV 27 4/17 times the design pressure. In this test, the primary side of the steam
generator is pressurized to 3107 psig coincident with no pressurization of
the secondary side. To hydrostatically test the RCS, a separate
hydro-test pump is provided.
- 2. Secondary-Side Hydrostatic Test Before Initial Startup
The secondary side of the steam generator is pressurized to 1360 psia, or 1.25 times the design pressure of the secondary side coincident with the
primary side at 0 psig.
- 3. Primary-Side Leak Test
After each time the primary system has been opened, a leak test is performed. For design purposes, the primary system pressure is
assumed to be raised to 2500 psia during the test, with the system
temperature above design transition te mperature, while the system is checked for leaks. In actual practice, the primary system will be
pressurized to < 2500 psia to prevent the pressurizer safety valves from
lifting during the leak test.
During this leak test, the secondary side of the steam generator will be pressurized so that the pressure differential across the tubesheet does
not exceed 1600 psi. This is accomplished by closing off the steam lines.
Since the tests outlined under items 1 and 2 occur prior to plant startup, the number of cycles is independent of plant life.
The design loading combinations and the associated stress and deformation limits are provided
in tables 5.2-3 through 5.2-7.
ASME Code Class 1 active and inactive valves in the RCPB, as defined in 10 CFR 50.2, are
reflected in table 5.2-8. Active valves are those in the pressure boundary whose operability
through a mechanical motion is relied on to perform a safety function (as well as reactor
shutdown function) during the transients or events considered in each operating condition
category. Pressure boundary valves which have no required motion and must only retain their
structural integrity, are not classified as active valves.
There are no active pumps in the RCPB. The reactor coolant pumps, which are the only pumps
within the RCS boundary, are classified as inactive.
Every valve and pump is hydrostatically te sted by the manufacturer to ASME Boiler and Pressure Vessel Code requirements to ensure the integrity of the pressure boundary parts.
This test is followed by a seat leak test to MSS-SP-61 criteria to ensure that no gross
deformation is caused by the hydrostatic test.
FNP-FSAR-5
5.2-11 REV 27 4/17 The design methods and procedures used to show that active valves listed in table 5.2-8 will
operate during a faulted condition are described in section 3.9. The control and instrumentation
are discussed in chapter 7.0.
Valves required to open or close during or following any specified plant design transient
condition have been designed in accordance with various codes and procedures that have been
widely used by the nuclear industry. These codes and procedures are based on engineering
judgment, inservice performance, and fundamental principles of engineering mechanics rather
than the requirements of a detailed stress analysis. This basis has resulted in conservative
designs which, in conjunction with periodic inspections, ensure that these components will
function as required.
Those remotely-operated valves that are used in the isolation of the RCPB during normal plant
operation, and are not relied on to function after an accident, are redundant. The inadvertent
operation of one of these redundant valves, excluding the pressurizer power operated relief
valve, does not increase the severity of any transient. Should the pressurizer power-operated
relief valve inadvertently open, operator action is required to close the pressurizer
power-operated relief valve block valve to ensure that the severity of any transient is not
increased.
The loading combinations and associated stress or deformation limits for inactive components
are provided in tables 5.2-3 through 5.2-7.
Allowable stress limits for active Code Class 1 valves are provided in paragraph 3.9.4.1. There
are no active Code Class 1 pumps within the RCPB.
The design evaluation of the RCS, including the types of analyses that are performed to ensure
the performance and the structural adequacy of the RCS, is provided below in
paragraph 5.2.1.10.1.
The RCS provides for heat transfer from the reactor to the steam generators under conditions of
forced circulation flow and natural circulation flow. The heat transfer capabilities of the RCS are
analyzed in chapter 15 for various transients.
FNP-FSAR-5
5.2-12 REV 27 4/17 During the second phase of plant cooldown and during cold shutdown and refueling, the heat
exchangers of the RHR system are employed. Thei r capability is discussed in section 5.5.
The pumps of the RCS ensure heat transfer by forced circulation flow. Design flowrates are
discussed in conjunction with the reactor coolant pump description in section 5.5.
Initial RCS tests are performed to determine the total delivery capability of the reactor coolant
pumps. Thus, it is confirmed prior to plant operation that adequate circulation is provided by the
RCS.
To ensure a heat sink for the reactor under conditions of natural circulation flow, the steam
generators are at a higher elevation than the reactor. In the design of the steam generators, consideration is given to provide adequate tube area to ensure that the RHR rate is achieved
with natural circulation flow.
Whenever the boron concentration of the RCS is reduced, plant operation will be such that good
mixing is provided in order to ensure that the boron concentration is maintained uniformly
throughout the RCS.
Although mixing in the pressurizer will not be achieved to the same degree, the fraction of the
total RCS volume which is in the pressurizer is small. Thus, the pressurizer liquid volume is of
no concern with respect to its effect on boron concentration.
Also, the design of the RCS is such that the distribution of flow around the system is not subject
to the degree of variation which would be required to produce nonhomogeneities in coolant
temperature or boron concentration as a result of areas of low coolant flow rate. An exception
to this is the pressurizer, but for the same reasons as discussed above, it is of no concern.
Operation with one reactor coolant pump inoperable is possible under certain conditions and, in
this case, there would be backflow in the associated loop even though the pump itself is
prevented from rotating backwards by its antir otation device. The backflow through the loop would cause departure from the normal temperature distribution around the loop, but would
maintain the boron concentration in the loop the same as that in the remainder of the RCS.
The range of coolant temperature variation during normal operation is limited and the
associated reactivity change is well within the capability of the rod control group movement.
For design evaluation, the heatup and cooldown transients are analyzed by using a rate of
temperature change equal to 100°F/h. Over cert ain temperature ranges, fracture prevention criteria will impose a lower limit to heatup and cooldown rates.
Concentrated boric acid solution from the chemical and volume control system is added, as
necessary, to the reactor coolant system to increase the reactor coolant boron concentration to
ensure adequate shutdown margin is maintained as required by plant Technical Specifications.
Therefore, it is concluded that the temperature changes imposed on the RCS during its normal
modes of operation do not cause any abnormal or unacceptable reactivity changes.
FNP-FSAR-5
5.2-13 REV 27 4/17 The design cycles as discussed in the preceding section are conservatively estimated for
equipment design purposes and are not intended to be an accurate representation of actual
transients or, for all cases, to reflect operating experience.
Certain design transients, with an associated pressure and temperature curve, have been
chosen and assigned an estimated number of design cycles for the purpose of equipment
design. These curves represent an envelope of pressure and temperature transients on the
RCS boundary with margin in the number of design cycles chosen based on operating
experience.
To illustrate this approach, the reactor trip transient can be mentioned. Four hundred design
cycles are considered in this transient. One cycle of this transient would represent any operational occurrence which would result in a reactor trip. Thus, the reactor trip transient
represents an envelope design approach to various operational occurrences.
This approach provides a basis for fatigue evaluation to ensure the necessary high degree of
integrity for the RCS components.
System hydraulic and thermal design parameters are used as the basis for the analysis of
equipment, coolant piping, and equipment support structures for normal and upset loading
conditions. The analysis is performed using a static model to predict deformation and stresses
in the system. Results of the analysis give six generalized force components, three bending
moments and three forces. These moments and forces are resolved into stresses in the pipe in
accordance with the applicable codes. Stresses in the structural supports are determined by
the material and section properties assuming linear elastic small deformation theory.
In addition to the loads imposed on the system under normal and upset conditions, the design
of mechanical equipment and equipment supports requires that consideration also be given to
abnormal loading conditions, such as seismic and pipe rupture.
Analysis of the RCLs and support systems for seismic loads is based on a three-dimensional, multimass elastic dynamic model with nonlinear bumper and tie-rod supports. This model is
coupled to a simplified reactor containment building model (reference 32). The seismic model is
then subjected to time-history seismic OBE and SSE excitation (reference 33) with all the SG
snubbers removed. The piping, equipment nozzle, and equipment support loads from this
seismic analysis are obtained and evaluated.
The dynamic analysis employs the displacement method, lumped parameter, and stiffness matrix formulations and assumes that all components behave in a linearly elastic manner. The
reduced modal analysis method and modal superposition method are used in the time-history
seismic analyses. Seismic analyses are covered in detail in section 3.7.
Analysis of the RCLs and support systems for blowdown loads resulting from a loss-of-coolant
accident (LOCA) is based on the time-history response of simultaneously applied blowdown
forcing functions on a single broken and unbroken loop dynamic model. The forcing functions
are defined at points in the system loop where changes in cross-section or direction of flow
occur such that differential loads are generated during the blowdown transient. Stresses and
loads are checked and compared to the corresponding allowable values.
FNP-FSAR-5
5.2-14 REV 27 4/17 The stresses in components resulting from normal sustained loads and the blowdown analysis
are combined with the seismic analysis to determine the maximum stress for the combined
loading case. This is considered a very conservative method since it is highly improbable that
both maxima will occur at the same instant. These stresses are combined to determine that the
RCLs and support system will not lose its intended functions under this highly improbable
situation.
Protection criteria against dynamic effects associated with pipe breaks are covered in
section 3.6.
For fatigue evaluations, in accordance with the ASME Boiler and Pressure Vessel Code, maximum stress intensity ranges are derived from combining the normal and upset condition
transients given in paragraph 5.2.1.5. Note that there are no emergency conditions designated.
The stress ranges and number of occurrences are then used in conjunction with the fatigue
curves in the ASME Boiler and Pressure Vessel Code to get the associated cumulative usage
factors.
The criterion presented in the ASME Boiler and Pressure Vessel Code is used for the fatigue failure analysis. The cumulative usage factor is 1.0 and hence, the fatigue design is adequate. Metal fatigue, including the effect of environmentally assisted fatigue, was evaluated for license renewal as a TLAA in accordance with 10 CFR 54.21. The results for the period of
extended operation are summarized in chapter 18, subsection 18.4.2.
The reactor vessel vendor's stress report is reviewed and approved by Westinghouse Electric
Corporation. The stress report includes a summary of the stress analysis for regions of
discontinuity analyzed in the vessel, a discussion of the results (including a comparison with the
corresponding code limits), a statement of the assumptions used in the analyses, descriptions
of the methods of analysis and computer programs used, a presentation of the actual
calculations used, a listing of the input and output of the computer programs used, and a
tabulation of the references cited in the report. The contents of this stress report and other
Class I component stress reports are in accordance with the requirements of the ASME Boiler
and Pressure Vessel Code. These stress reports are available inhouse for review.
The Westinghouse analysis of the steam generator tube-tubesheet complex is included as part
of the stress report requirement for ASME Code Class 1 Nuclear Pressure Vessels. The
evaluation is based on the stress and fatigue limitations outlined in ASME Section III.
The stress analysis techniques utilized include all factors considered appropriate to
conservative determination of the stress levels used in evaluation of the tube-tubesheet
complex. The analysis of the tubesheet complex includes the effect of all appurtenances
attached to the perforated region of the tubesheet that are considered appropriate for
conservative analysis of the stresses for evaluation on the basis of the ASME Code Section III
stress limitations. The evaluation involves the heat conduction and stress analysis of the
tubesheet, channel head, secondary shell structure for particular steady design conditions for
which code stress limitations are to be satisfied, and for discrete points during transient
operation for which the temperature/pressure conditions must be known to evaluate maximum
and minimum stresses for fatigue life usage. In addition, limit analyses are performed to
determine tubesheet capability to sustain faulted conditions for which elastic analysis does not
suffice. The analytic techniques utilized are computerized and significant stress problems are
verified experimentally to justify the techniques when possible.
FNP-FSAR-5
5.2-15 REV 27 4/17 The major concern in fatigue evaluation of the tube weld is the fatigue strength reduction factor
to be assigned to the weld root notch. For this reason, Westinghouse has conducted low cycle
fatigue tests of tube material samples to determine the fatigue strength reduction factor, and
has applied them to the analytic interaction analysis results in accordance with the accepted
techniques in the Boiler and Pressure Vessel Code for experimental stress analysis.
The steam generator tube-tubesheet complex integrity is verified by analysis for most adverse
conditions resulting from a rupture of either primary or secondary piping.
It has been established that for such accident conditions, where a primary-to-secondary-side
differential pressure exists, the primary membrane stresses in the tubesheet ligaments, averaged across the ligament and through the tubesheet thickness, satisfy the conditions given
in table 5.2-3 for this faulted event. Also, for such accident conditions, the primary membrane
stress plus primary bending stress in the t ubesheet ligaments, averaged across the ligament width at the tubesheet surface location giving maximum stress, must not exceed the faulted
condition criteria. In the case of a primary pressure loss accident, the secondary primary
pressure differential is somewhat higher than the primary secondary design pressure
differential. However, rigorous analysis shows that no stresses in excess of those covered by
the ASME Boiler and Pressure Vessel Code for faulted conditions are experienced by the
tubesheet for this accident.
Table 5.2-9 summarizes the tubesheet stress results for a pressure differential of 2485 psig.
Tabulations of significant results for the tube-tubesheet complex are presented in tables 5.2-10
through 5.2-17 and figures 5.2-1 through 5.2-4.
The tubes have been designed to the requirements of the ASME Boiler and Pressure Vessel
Code assuming 2485 psig as the design pressure differential. Hence, neither a primary nor a
secondary pressure-loss accident impose stresses beyond those normally expected and
considered as normal operation by the Code. ASME Section VIII design curves for
iron-chromium-nickel steel cylinders under external pressure indicate a collapse pressure of
2310 psi for tubes having the minimum properties required by ASTM specifications. This
indicates a minimum factor of safety of 2.4 against collapse. Collapse tests of 7/8-in. diameter, 0.050-mil-wall straight tubes at room temperature indicate actual tube strengths are significantly higher than specification and a collapse pressure of 6000 psi was recorded for the straight tube.
The code charts indicate a collapse pressure of 2740 psi for this tube. The difference is
attributed to the fact that the yield strength of the tube tested was 44,000 psi and the code
charts are based on a yield strength of approximately 29,000 psi at room temperature.
Consideration has been given to the superimposed effects of secondary-side pressure loss and
the safe shutdown earthquake loading. For the case of the tubesheet, the safe shutdown
earthquake loading will contribute an equivalent static pressure loading over the tubesheet of <
10 psi (for vertical shock). Such an increase is small when compared to the pressure
differentials (up to 2485 psig) for which the tubesheet is designed and does not result in
stresses exceeding the allowable stresses. The fluid dynamic forces on the internals under
secondary steam-break accident conditions indicate, in the most severe case, that the tubes are
adequate to constrain the motion of the baffle plates with some plastic deformation, while
boundary integrity is maintained.
FNP-FSAR-5
5.2-16 REV 27 4/17 A complete tube-tubesheet complex analysis is also performed to verify structural integrity for a
primary pressure loss accident plus the safe shutdown earthquake.
Although the ASME Boiler and Pressure Vessel Code provides for rules and techniques in
analysis of perforated plates, it should be noted that the stress intensity levels for perforated
plate are given for triangular perforation arrays. Westinghouse tubesheets contain square hole
arrays. Hence, Westinghouse utilizes its own data and that obtained from Pressure Vessel
Research Committee research in square array perforation patterns for development of similar
charts for stress intensity factors and elastic constants. The resulting stress intensity levels and
fatigue stress ranges are evaluated according to the stress limitation of the code.
The vessels, piping, valves, pumps, and associated supports of the RCPB are designated ANS
Safety Class 1.
Loading combination and allowable stresses for ASME Section III, Class 1 components, piping
and supports are given in tables 5.2-3 through 5.2-7.
Valves in sample lines are not considered to be part of the RCS boundary, i.e., not ANS Safety
Class 1. This is because the nozzles where these lines connect to the RCS are orificed to a
3/8-in. hole. This hole restricts the flow such that loss through a severance of one of these lines
can be made up by normal charging.
The load combinations that are considered in the design of structural steel members of component supports are given in paragraph 5.2.1.5. The design is described in paragraph 5.5.14.2.
A. Deadweight
The deadweight loading imposed by the piping on the supports is defined to consist of the dry weight of the coolant piping and the weight of the water
contained in piping during normal operation. In addition, the total weight of the
primary equipment components, including water, forms a deadweight loading on
the individual component supports.
B. Thermal Expansion
The free vertical thermal growth of the reactor vessel nozzle centerlines is considered to be an external anchor movement transmitted to the RCL. The
weight of the water in the steam generator and reactor coolant pump is applied
as an external force in the thermal analysis to account for equipment nozzle
displacement as an external movement. The cold and hot moduli of elasticity, the coefficient of thermal expansion at the metal temperature, external movements transmitted to the piping as described
above, and the temperature rise above the ambient temperature define the
required input data to perform the flexibility analysis for thermal expansion.
FNP-FSAR-5
5.2-17 REV 27 4/17 C. Earthquake Loads
The intensity and character of the earthquake motion that produces forced vibration of the equipment mounted within the containment building are specified
in terms of the ground acceleration time history. The ground acceleration time
history for earthquake motions is given in reference 33.
D. Pressure
The steady-state hydraulic forces based on the system initial pressure are applied as external loads to the RCL model for determination of the RCL/support
system deflections and support forces.
E. Pipe Rupture Loads
Blowdown loads are developed in the broken and unbroken RCLs as a result of the transient flow, pressure fluctuations following a postulated LOCA in one of
the RCL accumulator or RHR branch nozzles. The postulated LOCA is assumed
to have 1-ms opening time to simulate the instantaneous occurrence.
F. Analytical Methods
The static and dynamic structural analyses assume linear elastic behavior and employ the displacement (stiffness) matrix method and the normal mode theory
for lumped parameter, multimass structural representation to formulate the
solution. The complexity of the physical system to be analyzed requires the use
of a computer for solution. Herein lies the need for accurate and adequate
representation of the physical system by means of an idealized (mathematical) model.
The loadings on the component supports are obtained from the analysis of an integrated RCL support system dynamic structural model as shown on
figure 5.2-5.
G. Reactor Coolant Loop Model
The RCL model is constructed for the WESTDYN, WECAN, and the PS +
CAEPIPE computer programs. These are special purpose programs designed
for the static and dynamic analysis of redundant piping systems with arbitrary
leads and boundary conditions. The RCL lumped mass model represents an
ordered set of data that numerically de scribes the physical system to the WESTDYN, WECAN, and the PS + CAEPIPE programs. The node point
coordinates and the incremental lengths of the elements are calculated. The
lumping of distributed mass of a segment or elbow is accomplished by locating
the total mass at the mass center of gravity.
A valid representation of the effect of the equipment motion on the RCL piping and its support system is ensured by modeling the mass and stiffness
characteristics of the equipment in the overall RCL model. Since the reactor FNP-FSAR-5
5.2-18 REV 27 4/17 pressure vessel is very massive and relatively rigid, for LOCA analysis, it is
represented by a fixed boundary condition for the RCL model. The requirement
in the time-history dynamic analysis, that the external hydraulic forcing functions
be applied at only mass points, influences the construction of the steam
generator and reactor coolant pump model described below. A simplified reactor
pressure vessel model is incorporated into the time-history seismic analysis.
The steam generator is represented by a multimass, lumped model. The lower mass position is located at approximately the intersection of the inlet and outlet
nozzles of the steam generator. The other masses are located at various
locations on the steam generator.
The reactor coolant pump is represented by a two-mass, lumped model. The lower mass position is located at the intersection of the pump suction and
discharge nozzles. The upper mass position is located at the center of gravity of
the pump motor.
H. Support Structure Models
The equipment support structure models are dual purpose since they are required to quantitatively represent (in terms of 6-x-6 stiffness matrix) the elastic
restraints which the supports impose upon the loop; and to evaluate the
individual support member stresses caused by the forces imposed upon the
supports by the loop.
Models for the STRUDL (2) computer program are constructed for the steam generator lower, steam generator upper lateral, and reactor coolant pump lower
support structures. The structure geometry and member properties are obtained
from the certified construction structural drawings.
I. Hydraulic Models
The hydraulic model is constructed to quantitatively represent the behavior of the coolant fluid within the RCLs in terms of the concentrated time-dependent loads it
imposes upon the loops.
In evaluating the hydraulic forcing functions during a LOCA, the pressure and the momentum flux terms are dominant. I nertia and the gravitational terms are neglected; however, they are taken into account to evaluate the local fluid
conditions.
Thrust forces resulting from a LOCA are calculated in two steps using two digital computer codes. The first code, MULTIFLEX, calculates transient pressure, flowrates, and other coolant properties as a function of time. The second code, THRUST, uses the results obtained from the first code and calculates time
history of forces at locations where there is a change in either direction or area of
flow within the RCL. These locations for the broken loop are shown in
figure 5.2-6.
FNP-FSAR-5
5.2-19 REV 27 4/17 In MULTIFLEX blowdown analysis, both the broken and the unbroken loops are represented. The NRC approved MULTIFLEX 1.0 computer code (reference 3)
is used to generate the transient coolant properties throughout the RCS. The MULTIFLEX code calculates the thermal-hydraulic transient within the RCS and
considers subcooled, transition, and two-phase (saturated) blowdown regimes.
The code employs the method of characteristics to solve the conservation laws, assuming one-dimensional flow and a homogeneous liquid and vapor mixture.
The RCS is divided into subregions in which each subregion is regarded as an
equivalent pipe. A complex network of these equivalent pipes is used to
represent the entire primary RCS.
A coupled fluid-structure interaction is incorporated into the MULTIFLEX code by accounting for the deflection of the constraining boundaries, which are
represented by separate spring-mass osc illator systems. For steam generator and other RCS component analyses, MULTIFLEX provides pressure and other
coolant property transients at select locations. For loop piping analyses, the time-history RCS properties as computed by MULTIFLEX are used as input to
the THRUST code to calculate the LOCA hydraulic forces at various locations
along the RCS piping for the broken and unbroken loops. In the THRUST
calculation of blowdown forces, the RCS is represented by the same model employed in the MULTIFLEX code. Twenty-six node points are selected along
the geometric model of the RCL where the vector forces and their coordinate
components are calculated.
The force components at each aperture are vectorially summed to obtain the total force components in global coordinate system at the nodes. These forces
are stored on electronic media and, after proper coordinate transformation, applied as external loadings on the RCL dynamic model.
J. Static Load Solutions
The static solutions for deadweight, thermal expansion, and pressure load conditions are obtained by using the WESTDYN computer program. The
computer input consists of the RCL mode, stiffness matrices representing various
supports for static behavior, and the appropriate load condition. Coordinate
transformations for rotation from the local or support coordinate system to the
global system are applied to the stiffness matrices prior to their input.
K. Time-History Dynamic Solution for Seismic Loading
The reduced modal analysis method and modal superposition method are used in the time-history seismic analyses. The reduced modal analysis is used to
determine the natural frequencies and mode shapes for a linear, undamped
structure. This analysis requires the specification of dynamic or active degrees
of freedom (DOF) for the model, which are a subset of the total number of DOF.
The selection of dynamic DOF must be such that the low frequency spectrum
can accurately be represented while a reduced eigenvalue problem is solved. In
other words, the selected dynamic (or active) DOF should be able to describe the
frequency modes of interest.
FNP-FSAR-5
5.2-20 REV 27 4/17
The modal superposition method gives a time-history solution for the response of an arbitrary structure subjected to known nodal forces or ground acceleration
time histories. The structure may include linear and nonlinear elements. The
uncoupled modal equations are integrated analytically.
The input to the time-history seismic analysis is in the form of time-history seismic motions applied individually for all three components at the base of the
soil springs in the north-south, east-west, and vertical directions. These time-
history seismic motions were provided in reference 33. The total response is
obtained by determining the maximum response from absolutely combining each
of the two horizontal responses with the vertical seismic response.
L. Time-History Dynamic Solution for LOCA Loading
The initial displacement configuration of the mass points is defined by applying the initial steady-state hydraulic forces to the unbroken RCL model. For this
calculation, the support stiffness matrices for the static behavior are incorporated
into the RCL model. For dynamic solution, the unbroken RCL model is modified
to simulate the physical severance of the pipe caused by the postulated LOCA
under consideration. This model includes definition of the support stiffness
matrices for dynamic behavior. The natural frequencies and normal modes for
the modified RCL dynamic model are determined. After proper coordinate
transformation to the RCL global coordinate system, the hydraulic forcing
functions to be applied at each lumped mass point are stored on magnetic tape for later use as input to the FIXFM program. FIXFM is a part of program
WESTDYN.
The initial displacement conditions, natural frequencies, normal modes, and the time-history hydraulic forcing functions from the input to the FIXFM program
which calculates the dynamic time-history displacement response for the
dynamic degrees of freedom in the RCL model. The displacement response is
plotted at all mass points. The displacement response at support points is
reviewed to validate the use of the chosen support stiffness matrices for dynamic
behavior. The time-history displacement response from the valid solution is
saved on electronic media for later use to compute the support loads and to
analyze the RCL piping stresses.
M. Evaluation of Support Structures
The support loads are computed by multiplying the support stiffness matrix, and the displacement vector, at the support point. The support loads are saved on
magnetic tape for use in support member evaluation.
The STRUDL computer program is used to obtain support stiffness matrices and member influence coefficients for the equipment supports. Unit forces along and
unit moments about each coordinate axis are applied to the models at the
equipment vertical centerline joints. Stiffness analysis is performed for each unit
load for each model. Printed output includes all six components of displacement FNP-FSAR-5
5.2-21 REV 27 4/17 at the joint at which loads are applied and six force components at each end of
each member in the support system.
Joint displacements for applied unit loads are formulated into flexibility matrices.
These are inverted to obtain support stiffness matrices which are included in the
RCL model.
Loads acting on the supports obtained from the RCL analysis (including time-history LOCA forces), support structure member properties, and influence
coefficients at each end of each member, are input into the THESSE program.
This program accomplishes the following for each support case used:
- 1. Combines the various types of support plane loads to obtain operating condition loads (normal, upset, or faulted).
- 2. Multiplies member influence coefficients by operating condition loads to obtain all member internal forces and moments. The 6-x-6 force arrays
are printed for each end of each member. Diagonal terms in the array are the maximum (or minimum) values of each internal member force
component and the other terms are the corresponding values of all other
components. In addition, all member force components are printed along
with the time of occurrence for the LOCA time-history loading producing
the highest stresses in each member. This output gives a complete
tabulation of all worst force and stress conditions in each member in the
supporting system. It also provides maximum loads on the supporting
concrete.
- 3. Solves appropriate stress or interaction equations for the specified operating condition. Maximum normal stress, shear stress, and
combined load interaction equation values are printed as a ratio of
maximum actual values divided by limiting values. The time of
occurrence of the maximum value of each equation is also printed for the
faulted condition, which includes time-history LOCA forces. Stress and
interaction equations are used with limits specified for the operating
conditions considered.
The stresses that were calculated are given in tables 5.2-28 and 5.2-31 as a percentage of the allowable. The member number refers to the
identification used in the computer code. The members are identified by
general classifications, such as "lower bumpers" for the steam generator
supports. Those members which have no stress entries such as steam
generator upper support members for the normal condition see no load in
that plant condition. The largest percentage of allowable for any member
of the steam generator, reactor coolant pump, and pressurizer supports
for the normal condition is 34 percent; for the upset conditions, 44
percent; and for the faulted conditions of the SEE, combined with LOCA, 92 percent.
FNP-FSAR-5
5.2-22 REV 27 4/17 The reactor vessel supports were analyzed using a detailed finite element model. The maximum horizontal and maximum vertical loads were simultaneously applied to the support model, and the corresponding
stresses were determined. For the reactor vessel support box (figure 3M-2), the percentage of allowable stress for the normal condition
is 41 percent; for the upset condition, 47 percent; and 38 percent for the
faulted condition.
The reactor vessel support shoe (figure 3M-1) is stressed to 37 percent of the allowable stress for the upset condition and 48 percent of the
allowable stress for the faulted condition.
N. LOCA Evaluation of the Control Rod Drive Mechanisms
The response of the control rod drive mechanisms (CRDM) to the postulated reactor vessel inlet nozzle and outlet nozzle limited displacement breaks has
been evaluated. The time-history analysis of the mechanism has been
performed for the vessel motion developed previously. A one-row model of the
CRDMs was formulated with gaps at the upper CRDM support modeled as
nonlinear elements. The CRDMs were represented by beam elements with
lumped masses. The translation and rotation of the vessel head were applied to
this model (see figure 5.2-22). The resulting loads and stresses were compared
to allowables to verify the adequacy of the system. The highest loads occur at
the head adapter, the location where the mechanisms penetrate the vessel head.
The bending moments at this location are presented in table 5.2-32 for the
longest and shortest CRDM. The combined effect, including seismic loads, is
shown to be less than the allowable bending moment at this location.
The heat transfer capability of the steam generators is sufficient to transfer to the steam and power conversion system the heat generated during normal
operation, and during the initial phase of plant cooldown under natural circulation
conditions.
When the components and systems for the Farley units were being designed, only general
design requirements existed for faulted conditions. There were no specific stress limits or
associated methods of analysis established for faulted conditions. To provide a conservative
basis for the analysis of Class 1 components, the collapse curves given in the PSAR were
developed. The criterion represented by the collaps e curves has evolved into the criteria of table 5.2-6 of the FSAR. The methods and criteria in table 5.2-6 should thus be reviewed with
respect to the criterion agreed to in the PSAR, rather than with the more recently derived
methods and limits established in the nonmandatory Appendix F of the ASME Code,Section III.
These methods of analysis, in conjunction with the faulted condition stress limits, ensure that
the general design requirements of the NRC for faulted conditions will be met and the plant can
thus be safely shut down under accident conditions.
FNP-FSAR-5
5.2-23 REV 27 4/17 For the RCL and components, the elastic system analysis option of table 5.2-6 was used.
Elastic component analyses were used on all components except those discussed below.
Inelastic component analysis was used for the reactor coolant pump support feet. The pump
casing with the pump support feet is shown on figure 5.2-20. The pump foot was analyzed for a
set of umbrella loads which are greater than the loads expected in any plant. The umbrella
loads are calculated for the faulted condition and each of the maxima of the six load
components, F x , F y , ..., M z , are assumed to occur simultaneously. For example, the maximum F is chosen by surveying many past plants, and th is is applied simultaneously with the maximum F x , F y , ..., M z , all determined similarly. The actual plant loads are calculated and compared to the umbrella loads. Conformance indicates adequacy of the component for the specific plant
application. If conformance is not demonstrated, an individual plant analysis would be
performed. Table 5.2-26 indicates the relationship between the Farley specific plant loads for
three different faulted conditions (from three different break locations) and the umbrella loads for
which the pump foot was designed. The actual plant loads are, in themselves, also
conservative since the maximum for each of the six load components is determined and
assumed to act concurrently with the others.
For the LOCA condition, the dynamic time-history
analyses show that the maximum values of the six load components do not act concurrently.
The seismic event, although evaluated by response spectra analysis, is also dynamic and the
load component maximums at the foot clearly will not coincide. Note from table 5.2-26 that the
umbrella loads are greater than these actual plant loads by a factor ranging from 1.0 to 20.4.
From the preceding discussion, the conservatisms in the actual plant loads and the adequacy of
the umbrella loads are therefore demonstrated.
The entire casing foot was analyzed by means of a 3-dimensional stress analysis. The foot
model utilized symmetry about the bolt hold radial centerline (figure 5.2-21). The completed
model contains 1584 node points and 1518 3-dimensional solid elements with 4088 active
degrees of freedom in the model. The 3-dimensional finite elements are a mixture of
rectangular prisms, triangular prisms, and tetrahedrons. The vertical side and horizontal plate
sections have a minimum of four elements through the thickness. The model therefore yields
bending stresses as well as direct stresses through the thickness. The higher stress regions
have a finer model mesh consisting of smaller tetrahedron and triangular prism elements.
The ANSYS computer code (11) plastic analysis options were employed. The plasticity program is based upon incremental strain equations with the Prandtl-Reuss flow rule (12). The virgin stress-strain option was used to incur the true stress-true strain material curve. To yield the
required accuracy, loading increments were computed to keep the size of the plastic strain
increments near the size of the material yield strain. The smaller load steps keep the solution
process from diverging from the input stress-strain curve.
The resulting faulted condition plastic analysis stress intensity was compared with the faulted
condition criteria of 0.7 S ut = 59, 950 psi for 304 SS at 600°F. This is the limit for the primary membrane plus tending stress intensities as given in table 5.2-6. Since the foot is similar to a
beam-type structure, the average stress across the section is very low. The primary tending stresses therefore control. The true ultimate stress, S ut is determined from the engineering ultimate stress (the engineering stress at the point of maximum load) by assuming constancy of
volume. Using this assumption, the true ultimate stress (S ut) is given by:
S ut = S u (1 + )
FNP-FSAR-5
5.2-24 REV 27 4/17 Where is the engineering strain corresponding to the point of maximum load.
The stresses in the pump foot-to-casing attachment zone and weld-filled region were not
controlling. The maximum stress in the foot occurred in the horizontal plate member near the
vertical to horizontal plate intersection and in line with the bolt. Since the faulted allowables are
based upon primary stresses and not peak stresses, the stress components in the high stress
region were linearized through the plate thickness. The resulting maximum stress intensity of
the section was found from these linearized maximum principal stresses. The stress intensity
was
( I)max = 59,614 psi which was less than the inelastic allowable.
The maximum localized outer-fiber strain corresponding to this stress was approximately
12-14 percent. The incremental strains, however, for each load step were kept to approximately
0.2 percent. The maximum deflection calculated by the statically-applied loads was
approximately 1 in. at the radial symmetry line passing through the hole. If geometry
modifications had been made for this deflection, the load induced in the high stress regions
would have been lowered since the moment arm for the beamline structure would decrease.
The present analysis is therefore considered conser vative from the analysis as well as the loads standpoint.
The stress and deflection analysis is based on a static application of loads which are physically
short duration, dynamically applied loads. For this reason, the actual deflections caused by the short duration peak loads could be expected to be much lower than those calculated by the
static analysis. The actual plant loads are also, in general, considerably lower than the design
loads so that this will further reduce the true magnitude of the deflections.
The reactor coolant pump outlet nozzle was analyzed for the faulted condition using the limit
analysis option of table 5.2-6. These limits are identical to the limits of Appendix F of the ASME
Boiler and Pressure Vessel Code, subparagraph F-1323.2(a)/NB-3213.22. A set of umbrella
loads was used in the analysis. These umbrella loads were developed using methods similar to
those described for the determination of the umbrella loads for the pump foot. These umbrella
loads, along with the actual plant loads for the faulted seismic condition, combined with the four
worst pipe-break cases, are given in table 5.2-27. (Note that the umbrella loads exceed the
actual plant loads by ratios of 2.15 to 9.77.) A three-dimensional finite element model was
developed and these worst-case umbrella loads were applied. The complete model contains
792 node points and 1512 elements with 4676 active degrees of freedom. The plastic options
of the ANSYS computer code were used with an elastic perfectly-plastic stress-strain curve. An
iterative loading technique utilizing 25 load steps took the model from the elastic condition to the
maximum load. The maximum load was increased by 10/9 to reflect the criteria in table 5.2-6.
This requires that the load be < 90 percent of the limit load. At the final load step, the load
deflection curve was increasing, indicating that the nozzle could take additional loads.
Therefore, the faulted limit analysis requirements had been satisfied.
The reactor vessel support pads are also qualified using the test option of table 5.2-6.
FNP-FSAR-5
5.2-25 REV 27 4/17 The reactor pressure vessel support pads and shoes are designed to restrain unidirectional
horizontal motion in addition to supporting the vessel. The design of the shoes, which are in
contact with pads attached to the nozzles of the vessel, allows radial growth of the vessel, but
restrains the vessel from horizontal displacements since each shoe prevents tangential
displacement of the vessel at the location of the support.
To duplicate the loads that act on the pads during faulted conditions, the tests, which utilized a
1/8 linear scale model of the support system (nozzle pad, shoes, shims, and hold down bolts), were performed by applying a unidirectional static load to the nozzle pad. The load on the
nozzle pad was reacted by the support shoe which was mounted to the test fixture with the
bolt-down bolts.
The above modeling and application of load thus duplicates the actual case and allows the
maximum load capacity of the support system to be accurately established. The test load, L t , was then determined by multiplying the maximum collapse load by 64 (ratio of prototype area to
model area) and including temperature effects in accordance with the rules of the ASME Code
Section III.
The loads on the shoes, as calculated in the analysis of the components for faulted conditions, are limited to the value of 0.80 L t in table 5.2-6.
The tests performed and the limits established for the test load method ensure that the
experimentally obtained value for L t is accurate and that the support system for the reactor pressure vessel will perform its intended function.
A discussion of the protection provided for the principal components of the RCS against
missiles is found in chapter 3.0.
External flooding protection for the containment and the RCS is provided as described in
appendix 3A. Internally-generated flooding of the containment could be caused only by
inadvertent generation of the safety injection system, including the containment spray system, or a LOCA condition. The maximum amount of water injected into the containment during a
spurious spray system operation is the volume of water contained in the refueling water storage
tank. All safety-related components inside the containment are designed to withstand the effect
of a water spray solution containing boric acid and sodium hydroxide. The maximum level of
water inside the containment that would result from the containment spray system would be below the level of the RCPB and any of the safety-related equipment. Therefore, the flooding of
this equipment is effectively precluded.
FNP-FSAR-5
5.2-26 REV 27 4/17 Fire protection for the RCS is provided by the following means: first, the minimum use of
combustible materials within the containment reduces the possibility of fire; second, environmental design specifications for electrical components and cables in the RCSs and all
safety-related equipment inside the containment are discussed in paragraph 3.11.2.1. These
requirements in design minimize the possibility of electrical shorts because of environmental
effects. If shorts do occur, the selective tripping feature described in subsection 8.3.1 instantly
removes power to the faulty equipment, minimizing damage. Also discussed in this
subsection is the single-failure criterion imposed on the safety-related equipment, which
ensures adequate protection for the RCS.
A brief description of the analyses and methods used to assure compliance with the applicable
codes is provided in paragraph 5.2.1.10.1.
The stress analyses used for faulted condition loadings are discussed in paragraph 5.2.1.10.1.
There are no emergency conditions specified.
The stress intensity evaluations for the normal, upset, and faulted conditions show that the
stress intensities in the piping are below the code-allowable values established in the design
specifications.
RCL piping minimum wall thickness, t m , was calculated in accordance with equation 1, subparagraph NB-3641.1, of the code. The as-built pipe minimum wall thickness meets the
code requirement.
The maximum combined primary stress intensity caused by DBE pressure, and weight in the
RCL is 19,010 psi, which is less than the code allowable stress intensity value of (1.5 S m) 26,700 psi, using equation 9 of NB-3652.
The primary-plus-secondary stress intensity range calculations outlined in the code were
performed. They show compliance with the code stress and fatigue requirements.
The cumulative usage factors calculated in accordance with the rules described in the code are
less than the allowable value of unity for all piping components. All normal, upset, and test
conditions having contributions to the usage factors were included in this evaluation. The code
limit on the fatigue damages, measured by cumulative usage factors, is satisfied at all locations FNP-FSAR-5
5.2-27 REV 27 4/17 on the RCL piping. The maximum cumulative factor obtained from the analysis is 0.632 at the
reactor pressure vessel outlet nozzle.
The RCL piping stress intensity ranges and fatigue damages are in conformance with the
requirements of the code for the fatigue damage evaluation performed under all normal, upset, and test conditions.
The primary stress intensity contribution during the faulted condition can be an increase in the
operating pressure of the RCL. The maximum pressure variation above the normal operating
pressure for all faulted condition transients is 780 psi, caused by a control rod ejection transient.
This pressure increase indicates that the permissible pressure of 2.0 P, where P is the design
pressure as defined in the design specification, is not exceeded for the faulted condition.
The calculated maximum values of stress intensity for high stress points in the unbroken legs of
the broken loop and the unbroken loop piping meet the code allowable stress intensity value for
equation 9 for all LOCA cases and main steam line rupture. The maximum primary stress
intensity for the primary stress intensity for the faulted condition loading combinations listed in
table 5.2-3 is 47,700 psi, which is less than the code allowable primary stress intensity value of
(3 S m) 53,400 psi.
Therefore, the reactor coolant piping as designed is adequate and will maintain its structural
integrity and meet the safety-related design requirements under all specified operating
conditions.
Pumps and valves within the RCS boundary are designed to meet the stress limits given in
table 5.2-4. Analytical methods are in accordance with the applicable codes described in
table 3.2-1.
Reactor coolant pump overspeed evaluations are covered in paragraph 5.5.1.3.
Valves required to open or close during or following any specified plant design transient
condition have been designed in accordance with various codes and procedures that have been
widely used by the nuclear industry. These codes and procedures are based on engineering
judgment, inservice performance, and fundamental principles of engineering mechanics, rather
than the requirements of a detailed stress analysis. This basis has resulted in conservative
designs which ensure that these components will function as required.
FNP-FSAR-5
5.2-28 REV 27 4/17 Normally, pipe 2-in. and under will be field run with the following exception:
- Piping classified under ASME Section III, Class 1.
These pipes require certain physical routing considerations for protection from such events as
pipe break and missiles and provisions for other design considerations such as separation and
redundancy. It is necessary, therefore, that all 2-in. and under piping in the above category not be permanently installed by the field until the field isometric sketch is reviewed and analyzed by
the responsible design engineer on the project.
Piping classified under ASME Section III, Classes 2 and 3, and ANSI B31.1 that require seismic
stress analysis, were routed on the piping design drawings and dimensioned in the field. Detail
isometrics were prepared for those pipes that were dimensioned in the field and forwarded to
the project for review and analyses by the responsible engineer for seismic stress, thermal
stress, shielding, and thermal insulation requirements as needed. The approved isometrics
were then released for permanent installation. Only piping in the ANSI B31.1 class that does
not require seismic analysis is run and dimensioned in the field without design engineering
approval being required for permanent installation.
The RCS is protected against overpressuriza tion by two independent relief systems whose operability is governed by the mode of plant operation. During startup and shutdown
operations, when the RCS is in the solid condition, low temperature overpressurization
protection is provided by an overpressurizati on mitigating system which utilizes the two RHR
system relief valves. Detailed information c oncerning the design of the overpressurization mitigating system is discussed in paragraph 5.2.2.4.
The RCS is protected by pressure relief devices comprising the three pressurizer safety valves
and the two power-operated relief valves (PORVs) in other modes of plant operation when the
overpressurization mitigating system is not in use. The following provides a detailed description of the pressurizer safety valves and the PORVs.
Pressure relief devices for the RCS include the three pressurizer safety valves and two PORVs shown on drawings D-175037, sheet 2 and D-205037, sheet 2; these discharge to the
pressurizer relief tank by common header. Other relief valves that discharge to the pressurizer
relief tank are itemized in table 5.2-18.
FNP-FSAR-5
5.2-29 REV 27 4/17 The pressure relief devices, as specified in paragraph 5.2.2.1, are mounted and installed as
follows:
A. The pressurizer safety valve inlet piping forms a loop to ensure a water seal on the valve seat. The water volume in the loop seal is minimized to keep the
reaction forces on the downstream piping as low as possible.
B. The loop seal piping is insulated to maximize loop seal water temperature. This maximizes the water volume expected to flash to steam upon lifting of the safety
valves and thus, reduces downstream forces on discharge piping.
C. A support is provided on the discharge piping as close as possible to each safety and relief valve discharge nozzle so that forces and moments (including pipe
whip and reactions following an assumed discharge pipe rupture) will not
jeopardize the integrity of the valves, the inlet lines to the valves, or the nozzles
on the pressurizer.
D. The support on the valve discharge is connected to the pressurizer instead of adjacent structures in order to minimize differential thermal expansion and seismic interactions.
E. Each straight leg of discharge piping is supported to take the force along that leg.
During original plant licensing, static and dynamic analyses were performed to verify the
adequacy of the pressurizer safety and relief valves for FNP.
Under NUREG 0737 (18),Section II.D.1, "Performance Testing of BWR and PWR Relief and Safety Valves," all operating plant licensees and applicants were required to conduct testing to
qualify the RCS relief and safety valves under expected operating conditions for design-basis
transients and accidents. In addition to the qualification of valves, the functionability and
structural integrity of the as-built discharge piping and supports was also required to be
demonstrated on a plant-specific basis.
In response to these requirements, a program for the performance testing of PWR safety and
relief valves was formulated by EPRI (19). The primary objective of the test program was to provide full scale test data confirming that functionability of the RCS PORVs and safety valves
are capable of performing their design function for expected operating and accident conditions.
The second objective of the program was to obtain sufficient piping thermal hydraulic load data
to validate models utilized for plant-unique analysis of PSARV discharge piping systems. Based
on the results of the aforementioned EPRI Safety and Relief Valve Test Program, additional
thermal hydraulic analyses were required to adequately define the loads on the piping system
due to valve actuation.
FNP-FSAR-5
5.2-30 REV 27 4/17 The results of the analysis for FNP were provided to the NRC in reference 20. NRC acceptance
of the FNP analysis is documented in reference 21. A summary of the FNP evaluation follows.
The safety valve discharge loads were calculated for the fluid transient condition that will produce the most severe loading on the piping system.
This occurs during a high pressure transient where steam from the pressurizer forces the water
in the water seal through the safety valve down the piping system to the relief tank. Forcing
functions are normally generated for hot or cold loop seals depending on the temperature in the
loop seal. The hot and cold loop seal conditions for Farley plants are consistent with the hot
and cold loop seal conditions defined in 1982 EPRI tests. Thermal hydraulic analysis for the
Farley pressurizer safety valve system was originally analyzed in 1982 for both the hot and cold
loop seal conditions. The hydraulic forces generated when the safety valves open are much
higher for the cold loop seal condition compared to those forces from the hot loop seal
condition. To reduce the loads from cold loop seal condition, modification to piping insulation
was necessary to ensure sufficient heat was conducted to the loop seal water. However, the
resulting loop seal piping temperatures were not high enough for classification as a hot loop
seal. The measured temperature profiles at the three loop seal systems fall between the bounds of hot and cold. The thermal hydraulic forces resulting from this intermediate
temperature loop seal are significantly less than predicted for the cold loop seal condition.
Based on the WCAP-10105 (22) report "Review of Pressurizer Safety Valve Performance as Ob served in the EPRI Safety and Relief Valve Test Program," (June 1982), the valve opening characteristics are not linear. The valve stem actually lifts partially, allowing the water seal to pass through the valve. Once the steam behind the
water slug reaches the valve stem, the valve stem will lift up fully in about .04 s. These valve opening characteristics are consistent with Figure 4-12 of the WCAP-10105 report and the loop
seal purge delay curve (Figure 8) for a Cr osby 6M6 forged safety valve. The opening
characteristics of the Crosby 6M16 safety va lves in Farley plants behave similarly with the Crosby 6M6 safety valves. Furthermore, a revi ew of EPRI data confirmed that the pressure increase ramp rate from 2 to 375 psi/s envelops the ramp rate for Farley.
Nonlinear opening area time-history valve characteristics are considered in the latest thermal hydraulic analysis. In addition, an average loop seal temperature of about 200 F, which is below the average Farley loop seal temperature, is used for the loop seal water slug properties.
This method used along with programs ITCH and FORFUN was benchmarked against the
previous EPRI test results and good correlations were documented. For the Farley plant
specific application, the thermal hydraulic forces were generated using the nonlinear valve
opening area time-history method. The application of this method results in a reduction in the
hydraulic thrust forces due to the water slug being more slowly passed through the valve (with 5
to 10% opening area) before the valve is fully open. The water hammer effect is thus reduced.
The thermal hydraulic forces generated by considering time-history variable valve opening were
determined for 5% and 10% initial valve opening areas. The forces with the 10% initial valve
opening area are more conservative than those with the 5% initial valve opening area and are
used to perform the time-history structural analysis of the pressurizer safety valve piping
system. For the Farley plant-specific safety va lves, the actual initial valve opening area is 5%
as determined by documented valve characteristics calculations.
FNP-FSAR-5
5.2-31 REV 27 4/17 The computer program used for the thermal hydraulic analysis was ITCH on Sun Workstation (23). This program was upgraded several times from original program ITCHVALVE (24,25) since 1982 and was renamed to ITCHVENT once on the mainframe computer. The program ITCHVENT was converted to Sun
workstation in 1992. Program ITCHVALVE was benchmarked against the EPRI test data.
ITCHVALVE is a 1-D thermal hydraulic code that calculates the time-history fluid properties within the pressurizer safety and relief valve system for the condition when the safety or relief
valves open. The thermal hydraulic forces are calculated by another program called
FORFUN (26) considering the momentum changes for the fluid in each element of the piping segment.
The structural modeling and analysis of the pressurizer safety valve piping system were
performed using the WECAN computer code (27). The piping system was modeled by pipe, elbow, support stiffness elements with both elastic and elastic/plastic capabilities. Consistent
mass effect was considered in the analysis. For the analysis of the piping system with
combination of deadweight and safety valve thrust discharge loadings, WECAN dynamic
transient time-history analysis option was chosen. The input time-history was determined by
ITCH and FORFUN computer programs and was applied to the piping system structural model.
Figure 5.2-16 shows the structural model of the Unit 2 safety line system, which contains three
6-in. safety valves on three lines before meeting a 12-in. common header. The 12-in. common header leading to the pressurizer relief tank is also in the model. Part of the relief line piping was modeled in the structural system to account for the structural system interactions.
Structural analyses were performed for both Units 1 and 2.
The time-history solution for the dynamic thrust analysis of safety valve discharge with loop seal
water slug was obtained from WECAN computer programs using direct integration methods.
Since the purpose of this analysis is to determine the elastic behavior of the piping system
under the extreme loading of valve thrust, the linear-elastic option of the WECAN program was
used. The resulting stress at 8 equally spaced circumferential points of a given
cross-section was calculated for a 1.0-s time history following the simultaneous discharge at the
three safety valves.
The pressurizer safety and relief valve piping system was originally qualified to its design basis
allowables prior to NUREG-0737 requirements. The design basis was the requirements of
ASME B&PV Code Section III, 1971 edition, including summer 1971 addenda for Class I piping
and the ANS B31.1-1967 Code with 1971 addenda for nonnuclear safety (NNS) piping. In
1982, Westinghouse performed additional evaluations to address TMI-related issues by
considering the cold loop seal loads for these piping systems (28). Criteria used in that analysis was based on the recommendation from piping subcommittee of the PWR Pressurizer Safety
and Relief Valve (PSARV) test program and was documented in a WCAP-10105 (22). Those criteria were reviewed and accepted by the NRC in a 1986 SER(29).
FNP-FSAR-5
5.2-32 REV 27 4/17 In the FNP evaluation, the loading combination and piping evaluation criteria of WCAP-10105
were applied with the exception of an allowable stress of 2.4 S h for the emergency condition for the NNS portion of the piping system. This exception was approved by the NRC as
documented in reference 21.
Using elastic analysis techniques, the Class I piping (which connects the pressurizer safety line
nozzle to the 6-in. safety valve), was qualified to the allowables listed in table 5.2-40 with the
effect of valve thrust under both emergency and faulted conditions. The NNS portions of the
piping system area also qualified to meet the allowables listed in table 5.2-41. The most limiting
stresses for the emergency conditions are shown in table 5.2-42.
One additional means to ensure that the safety valve remains operable after the loop seal water
is discharged is to assess the valve nozzle loads with respect to the valve operability limit
provided in the equipment specification. For emergency condition, the calculated valve nozzle
loads from the combination of deadweight, pressure, and valve thrust effects are within the
equipment specification allowable. This allowable requires the maximum total valve nozzle
stress to be 75% of the yield stress of the nozzle material at temperature. In addition, it further
requires that the maximum bending stress be 50% and the maximum torsion stress also be 50%
of the yield stress of the nozzle at temperature.
The piping system loading conditions considered for the pipe support evaluation consisted of the valve thrust loadings discussed above in combination with the existing design basis deadweight, normal thermal expansion, transient thermal expansion, and the OBE & SSE seismic loadings.
Since the pipe supports had previously been qualified for the Normal, Upset, Emergency, and
Faulted conditions, the supports were only evaluated for the worst case load combination
including the valve thrust loads from the piping system analysis. The loading combination used
for support evaluation is:
2 Thrst 2 SSE max/minThmDWP The purpose of the support evaluation was to demonstrate that the supports retained their integrity for the controlling combined loads. This was accomplished by generally limiting the actual support
member stresses to the allowable stress limits established by the ASME Boiler and Pressure
Vessel Code,Section III, Subsection NF and Appendix F, 1974 Edition. The code of record, AISC 7th Ed., does not address the faulted loading combination. ASME Subsection NF was
used for this evaluation since it is essentially the same as AISC for the normal and upset
conditions, and it provides criteria for the extreme faulted loading combination. In addition, the FNP-FSAR-5
5.2-33 REV 27 4/17 Subsection NF criteria are consistent with the pipe support criteria utilized by most other nuclear
plants.
In accordance with NRC IE Bulletin 79-02, concrete expansion anchors (CEA) on Class I pipe
support base plates were limited to manufacturer's allowables, including a Factor of Safety of
4.0. However, for four CEAs on NNS Class Pipe Support Base Plates, the manufacturer's
allowable including a factor of safety of 3.0, was applied. These bolts are identified in
table 5.2-43. The use of this safety factor for the 4 bolts was approved by the NRC for this
application as documented in reference 21.
Class I supports - the results of the pipe support evaluations based on the as-built support data provided to Westinghouse show that all the
Unit 1 and Unit 2 pipe support standard Grinnell components, structural members, and base
plate element stress levels are within the allowable stress limits of ASME Subsection NF and
Appendix F and will maintain their structural integrity and stability for the faulted loading
combination provided above. All concrete ex pansion anchor for class I supports have a minimum safety factor of 4.0.
NNS supports - all Unit 1 and Unit 2 NNS pipe supports satisfied the ASME Subsection NF and
Appendix F faulted stress criteria. Therefore, all the NNS pipe supports will maintain their
structural integrity for the specified loading combination. Most expansion anchors have safety
factor > 4.0. Table 5.2-43 provides a summary of only those NNS class pipe supports which
have concrete expansion anchors with safety factor < 4.0 but > 3.0 in their qualification.
The pressurizer is designed to accommodate pressure increases (as well as decreases) caused
by load transients. The spray system condens es steam to prevent the pressurizer pressure from reaching the setpoint of the PORVs during a step reduction in power level of 10 percent of
load.
The spray nozzles are located on the top of the pressurizer. Spray is initiated when the
pressure controlled spray demand signal is above a given setpoint. The spray rate increases
proportionally with increasing pressure rate and pressure error until it reaches a maximum
value.
The pressurizer is equipped with PORVs which lim it system pressure for a large power mismatch and thus prevent actuation of the fixed high pressure reactor trip. The relief valves
are operated automatically or by remote manual c ontrol. The operation of these valves also
limits the undesirable opening of the spring-loaded safety valves.
Remotely-operated block valves are provided to is olate the PORVs if excessive leakage occurs.
The relief valves are designed to limit the pressurizer pressure to a value below the high
pressure trip setpoint for all design transients up to and including the design percentage step
load decrease with steam dump, but without reactor trip.
FNP-FSAR-5
5.2-34 REV 27 4/17 Output signals from the pressurizer pressure control channels are used for pressure control.
These are used to control pressurizer spray and heaters and PORVs. Pressurizer pressure is
sensed by fast response pressure transmitters with a time response of better than 0.2 s.
In the event of a complete loss of heat sink, i.e., no steam flow to the turbine, protection of the
RCS against overpressure is afforded by pre ssurizer and steam generator safety valves along with any of the following reactor trip functions:
A. Reactor trip on turbine trip (if the turbine is tripped).
B. High pressurizer pressure reactor trip.
C. Overtemperature-T reactor trip.
D. Low-low steam generator water level reactor trip.
Continued integrity of the RCS during the maximum transient pressure is assured by design
within the applicable codes as discussed in reference 4. The code safety limit is 110 percent of
the 2485 psig design limit.
A detailed functional description of the process equipment associated with the high pressure trip
is provided in reference 5.
The upper limit of overpressure protection is based upon the peak surge into the pressurizer of
the reactor coolant produced as a result of turbine trip under full load, assuming no reactor trip.
The self-actuated safety valves are sized on the basis of steam flow from the pressurizer to
accommodate this surge at a setpoint of 2500 psia and a total accumulation of 3 percent. Note
that no credit is taken for the relief capability provided by the PORVs during this surge.
The RCS design and operating pressure, together with the safety, power relief and pressurizer
spray valve setpoints, and the protection system setpoint pressures, are listed in table 5.2-19.
System components whose design pressure and temperature are less than the RCS design
limits are provided with overpressure pr otection devices and redundant isolation means.
System discharge from overpressure protection dev ices is collected in the pressurizer relief tank in the RCS. Isolation valves are provided at all connections to the RCS.
Administrative procedures have been developed to aid the operator in controlling RCS pressure
during low temperature operation. However, to minimize the frequency of RCS overpressurization, an overpressure mitigating system is provided to mitigate pressure
excursions initiated by inadvertent mass and/or heat additions when the RCS temperature is less than or equal to the low temperature overpressu re protection (LTOP) System applicability
temperature specified in the PTLR.
FNP-FSAR-5
5.2-35 REV 27 4/17 The overpressure mitigation system employ s the RHR system relief valves (RHRSRV) to mitigate RCS overpressure transients. One relief valve is installed in each RHR suction line.
The RHRSRVs are spring-loaded, bellows-type valves which have a setpoint of 450 psig. The current methodology requirements use a setpoint of 436 +13 psig. At 495 psig the valves deliver full design flow. There are two isolation valves between each of the RHRSRVs and the
RCS. The RHR suction lines are automatically isolated when the RCS pressure exceeds
700 psig on Unit 2. The autoclosure interlock of the suction/isolation valves was removed from Unit 1. Two main control board annunciator windows are installed to alert the operators when the RHR suction/isolation valve(s) is not fully closed and the RCS pressure exceeds the alarm setpoint on Unit 1.The RHR suction valves inside containment are open when the RCS temperature is less than or equal to the LTOP System applicability temperature specified in the
PTLR, thereby aligning the RHR relief valves for RCS overpressurization protection. As
additional protection against RCS overpressurization, power is removed from the Unit 2 RHR isolation valves when the RCS temperature is below 180°F to prevent an inadvertent isolation of the RHRSRVs. Power is reinstated to the isolation valves prior to exceeding an RCS
temperature of 180°F via strict administrative controls, which assure the operability of the RHR
isolation valves and associated interlocks. On Unit 1, power is removed from the RHR isolation valves in Modes 1, 2, and 3.
The RHR relief valves have no electrical components. The autoclosure (Unit 2) and open-permissive circuits of the RHR motor-operated isolation valves meet the requirements of
IEEE-279-1971. Power supplies for the RHR isolation valves, the pressurizer pressure sensors, and the RCS temperature sensors are designed so that no single failure of the electrical system
or the loss of offsite power would isolate both of the RHR relief valves.
In addition, several control room alarms have been provided. A Seismic Category I alarm
designed to the requirements of IEEE-279-1971 alerts the operator if the RHR isolation valves
are not fully open when the RCS temperature is 300°F. Another alarm provides indication to the operator of any overpressure transient occurring when the RCS pressure > 450 psig.
The RHRSRVs and the associated discharge piping up to the pressurizer relief tank are
designed in compliance with Regulatory Guide 1.29, Rev. 1. The RHRSRVs were
manufactured by the Crosby Valve and Gage Company, which has certified that the
performance of these valves will not be degraded by an OBE event with a horizontal acceleration of 1.725 g and 1.455 g and a vertical acceleration of 1.221 g. This certification is
based on valves of similar construction and charac teristics as the subject relief valves. The piping downstream of the RHR system isolation valves up to the RHRSRVs, including the
RHRSRVs, meets the ANSI Nuclear Safety Criteria for the design of stationary
pressurized-water reactor plants, August 1970 draft. The piping upstream of the RHR system
isolation valves including these valves is Quality Group A per 10 CFR 50.55(a). The RHRSRV
discharge piping up to the pressurizer relief tank and the pressurizer relief tank are Nonnuclear
Class (B.31.1 piping); however, they are seismically supported.
Thus, the overpressure mitigating system is capable of functioning following a seismic event.
FNP-FSAR-5
5.2-36 REV 27 4/17
ASME,Section III, Appendix G, establishes guidelines for RCS pressure during low temperature operation ( 350°F). The relief system discussed in paragraph 5.2.2.4.1 serves to mitigate overpressure excursions to within these allowable limits. The worst-case mass input event was assumed to be the inadvertent operation of three high-head safety injection pumps with a maximum total flowrate of 1000 gal/min at 0 psig backpressure at RCS temperatures 180 F. Due to Technical Specification restrictions that allow only one operable charging pump at RCS
temperatures < 180 F, the worst-case mass injection is limited to the start of a single charging pump at RCS temperatures < 180 F. The worst heat input event was assumed to be the starting of a single reactor coolant pump with a temperature differential of 50°F existing between the RCS and the steam generator. The maximum calculated RCS pressures for these
postulated worst mass and heat input events remained below the pressures allowed by the
Appendix G curves for transients initiated below 325°F. For transients above 325°F, the
pressurizer code safety valves would relieve pressure to prevent violation of Appendix G limits.
Although the system described in paragraph 5.2.2.4.1 mitigates pressure excursions to address
the allowable pressure limits, administrative procedures are employed to minimize the potential
for the development of any transient that would challenge the system.
Of primary importance is the basic mode of operation of the plant. Normal operating procedures
maximize the use of a pressurizer cushion (steam bubble) during periods of low temperature
operation. A steam bubble is formed in the pressurizer at a cold leg temperature in the range of
approximately 130 to 180°F when the plant is being started up. It is collapsed at a cold leg
temperature of < 200°F when the plant is being cooled down.
This cushion dampens the plant response to potential transient generating inputs, thereby
providing easier pressure control with slower response rates.
This cushion substantially reduces the severity of some potential transients such as
RCP-induced heat input and slows the rate of pressure rise for others. This provides
reasonable assurance that most potential transients can be terminated by operator action
before an overpressure condition exists.
Administrative controls employed to minimize the potential for overpressure developing include
the following:
A. Only one charging pump may be operational when the RCS temperature is <
180°F. Power is removed from the two nonoperating charging pumps when the RCS temperature is 180°F, except during pump sump operations, by removing the motor circuit breakers from their electrical power supply circuits.
FNP-FSAR-5
5.2-37 REV 27 4/17 B. The letdown heat exchanger control valve is placed in the manual control position prior to starting or stopping an RHR pump when the RCS is in a water
solid condition. C. The RHR suction isolation valves are open and the RHR relief valves are available to mitigate an overpressure event or the RCS is vented whenever the
RCS temperature is 325°F or less.
D. A reactor coolant pump shall not be started with one or more of the RCS cold leg temperatures is < 325 F unless 1) the pressurizer water volume is < 770 ft 3 (24% of wide range, cold, pressurizer level indication) or 2) the secondary water
temperature of each steam generators is < 50°F above each of the RCS+ cold
leg temperatures.
E. The accumulators are isolated and power is locked out from the accumulator isolation valve operators at RCS pressure below 1000 psig. These actions are
completed prior to reducing RCS pressure to < 900 psig.
F. The low pressurizer pressure and low steam line pressure safety injection signals are blocked during heatup and cooldown to preclude an inadvertent ECCS
actuation.
G. During cooldown all steam generators should be connected to the steam header to assure a uniform cooldown of the RCS loops.
H. NRC acceptance criteria for GL 90-06 is as follows: When an LTOP channel is inoperable and the RCS is not water-solid (water-solid is defined as a pressurizer
level of 30% [cold calibrated], a trained, dedicated operator will be assigned to
monitor and control RCS pressure. The operator will have two independent
alarms available to identify the occurrence of an overpressure event, and will be
specifically trained to respond to these alarms.
I. It is recommended that if all reactor coolant pumps have been stopped for more than 5 min during plant heatup, and the reactor coolant temperature is greater
than the charging and seal injection water temperature, there should be no
attempt to restart a pump unless a steam bubble is formed in the pressurizer.
This precaution will minimize the pressure transient when the pump is started
and the cold water previously injected by the charging pumps is circulated
through the warmer reactor coolant components. The steam bubble will
accommodate the resultant expansion as the cold water is rapidly warmed.
If all reactor coolant pumps are stopped and the RCS is being cooled down by the residual heat
exchangers, a nonuniform temperature distribution may occur in the RCLs. No attempt should
be made to restart a reactor coolant pump unless a steam bubble is formed in the pressurizer.
These special precautions back up the normal operational mode of maximizing periods of steam
bubble operation so that cold overpressure transient prevention is continued during periods of
transitional operations.
FNP-FSAR-5
5.2-38 REV 27 4/17 Recommended procedures for ECCS testing include the following to preclude the development
of cold overpressurization transients:
A. The normal procedure for periodic ECCS pump performance testing is to test the pumps during normal operation or at hot shutdown conditions. Performance
testing of the ECCS pumps with the RCS in a water-solid condition is prohibited.
B. The SI/LOSP test is performed during Mode 6 operation or with the reactor defueled.
C. The ECCS branch line flow verification and charging pump low discharge head flow tests are performed in Mode 6 with the reactor vessel head removed or with
the reactor defueled.
The above procedural recommendations covering normal operations with a steam bubble, transitional operations where potentially water solid, followed by specific testing operations, provide in-depth cold overpressure prevention or mitigation, augmenting the installed
overpressure relief system.
The material specifications used for the principal pressure retaining applications in each
component comprising the Reactor Coolant System boundary are listed in table 5.2-20 for
Class 1 Primary Components and table 5.2-21 for Class I and II Auxiliary Components. These materials are procured in accordance with the specification requirements and include
supplemental requirements of the applicable ASME Code rules.
The welding materials used for joining the ferritic base materials of the reactor coolant boundary
conform to, or are equivalent to, ASME Material Specifications SFA 5.1, 5.2, 5.5, 5.17, 5.18, and
5.20. They are tested and qualified to the requirements of ASME Section III rules.
The welding materials used for joining the austenitic stainless steel base materials of the reactor
coolant boundary conform to ASME Material Specifications SFA 5.4 and 5.9. They are tested
and qualified according to the requirements stipulated in subsection 5.2.5.
The welding materials used for joining nickel-chromium-iron alloy in similar base material
combination and in dissimilar ferritic or austenitic base material combinations of the reactor
coolant boundary conform to ASME Material Specifications SFA 5.11 and 5.14. They are tested
and qualified to the requirements of ASME Section III rules and are used only in procedures that
have been qualified to these same rules.
FNP-FSAR-5
5.2-39 REV 27 4/17 Materials used in components within the RCPB are listed in tables 5.2-20, 5.2-21, and 5.2-23.
All of the ferritic low-alloy and carbon steels used in principal pressure-retaining applications are
provided with a 0.125-in. minimum thickness of corrosion-resistant cladding on all surfaces that
are exposed to reactor coolant. This cladding materia l has a chemical analysis which is at least equivalent to the corrosion resistance of types 304 and 316 austenitic stainless steel alloys or
nickel-chromium-iron alloy. The other base materials which are used in principal
pressure-retaining applications that are exposed to the reactor coolant are austenitic stainless
steel, nickel-chromium-iron alloy, and martensitic stainless steel. Ferritic low-alloy and carbon
steel nozzles are safe-ended with stainless steel weld metal analysis A-7 or
nickel-chromium-iron alloy weld metal F-Number 43 using weld buttering techniques followed by
a post-weld heat treatment. The latter buttering material requires further safe-ending with
austenitic stainless steel base material after completion of the post-weld heat treatment when the nozzle is larger than 4 in. nominal I.D. and/or the wall thickness is 0.531 in.
The cladding on ferritic-type base materials receives a post-weld heat treatment.
All of the austenitic stainless steel and nickel-chromium-iron alloy base materials are used in the
solution-anneal-heat-treat-condition. The heat treatments are as required by the material
specifications. During subsequent fabrication, these pressure-retaining materials are not
heated above 800°F other than instantaneously and locally by welding operations. The
solution-annealed surge line material is subsequently formed by hot bending followed by a
resolution-annealing heat treatment. Corrosion tests are performed in accordance with ASTM A
393.
In general, all of the materials listed in tables 5.2-20 and 5.2-21, which are used in principal
pressure retaining applications and are subject to elevated temperature during system
operation, are in contact with thermal insulation that covers their outer surfaces.
The thermal insulation used on the RCS, including the pressure vessel, is of the stainless steel
reflective-type.
In the event of coolant leakage, the ferritic materials will show increased general corrosion
rates. Where minor leakage is anticipated from service experience, such as valve packing, pump seals, etc., materials that are compatible with the coolant are used. These are shown in
tables 5.2-20 and 5.2-21. Ferritic materials exposed to coolant leakage can be observed as
part of the inservice visual and/or nondestructive inspection program to ensure the integrity of
the component for subsequent service.
The RCS chemistry specifications are given in table 5.2-22.
FNP-FSAR-5
5.2-40 REV 27 4/17
- a. The Water Chemistry Control Program is credited as a license renewal aging management
program (see chapter 18, subsection 18.2.2).
The RCS water chemistry is selected to minimi ze corrosion. A periodic analysis of the coolant chemical composition is performed to verify that the reactor coolant quality meets the
specifications.(a)
The chemical and volume control system (CVC S) provides a means for adding chemicals to the RCS to control the pH of the coolant during initial startup and subsequent operation, to
scavenge oxygen from the coolant during startup, and to control the oxygen level of the coolant
caused by radiolysis during all power operations subsequent to startup. The oxygen content
and pH limits for power operations are shown in table 5.2-22.
The pH control chemical employed is lithium hydroxide. This chemical is chosen for its compatibility with the materials and water chemis try of borated water, stainless steel, zirconium, and Inconel systems. In addition, lithium is produced in solution from the neutron irradiation of
the dissolved boron in the coolant. The lithium hydroxide is introduced into the RCS via the
charging flow. The solution is prepared in the laboratory and poured into the chemical mixing
tank. Reactor makeup water is then used to flush the solution to the suction manifold of the
charging pumps. The concentration of lithium hy droxide in the RCS is maintained as a function of boron concentration in the range specified for pH control. If the concentration exceeds this
range, either the cation-bed demineralizer or the mixed-bed demineralizer is employed in the letdown line to reduce the lithium concentration to within range.
During reactor startup from the cold condition, hydrazine is employed as an oxygen scavenging agent. The hydrazine solution is introduced into the RCS in the same manner as described
above for the pH control agent.
Dissolved hydrogen is employed during power operation to control and scavenge oxygen
produced because of radiolysis of water in the core region. Sufficient partial pressure of
hydrogen is maintained in the volume control tank so that the specified equilibrium
concentration of hydrogen is maintained in the reactor coolant. A self-contained pressure
control valve maintains a minimum pressure in the vapor space of the volume control tank. This can be adjusted to provide the correct equilibrium hydrogen concentration.
Components with stainless steel sensitized in the manner expected during component
fabrication and installation will operate satisfactorily under normal plant chemistry conditions in
pressurized-water reactor systems because chlorides, fluorides, and, particularly, oxygen, are
controlled to very low levels.
Assurance of adequate fracture toughness of ferritic materials in the reactor coolant system boundary is provided by compliance with Section III of the 1968 ASME Boiler and Pressure
FNP-FSAR-5
5.2-41 REV 27 4/17
- a. Reactor vessel neutron embrittlement was evaluated as a TLAA for license renewal in
accordance with 10 CFR 54.21 (see chapter 18, subsection 18.4.1).
Vessel Code, plus applicable Addenda and Code Cases. Test results for reactor pressure
vessel materials are given in tables 5.2-24 and 5.2-25.
The initial NDTT of plate materials in the reactor vessel beltline will not be greater than the
criteria for fracture energy levels as given in paragraph 5.2.4.3.
Although two test specimens for weld metal used in weld seam 10-923 of Unit 2 exhibited impact energies of 75 ft-lb at a test temperature of 10°F, it is expected that the upper shelf impact energy requirement of 75 ft-lb identified in paragraph IV.A.1.a of 10 CFR 50 Appendix G would easily be exceeded if tests had been perfor med at test temperatures representative of the upper shelf. A review of many weld test certificates provided by the vessel fabricator indicates
that the upper shelf energy of welds of chemical composition and fabrication history similar to
weld seam 10-923 and fabricated with the same type of wire and flux (type B-4 weld wire and
Linde 0091 Flux) used in seam 10-923 exceeds 75 ft-lb by a considerable margin. Four
examples of the vessel fabricator test results for weld material similar to that of seam 10-923 are shown in table 5.2-35. Like weld seam 10-923, two of these four examples did not exhibit 75 ft-lb for all test specimens at 10°F; however , at higher temperatures, 75 ft-lb was exceeded.
Individual data points obtained from Charpy V-notch impact tests for each of the base metal
heats in the Farley Unit 2 reactor vessel beltline are presented in tables 5.2-36, 5.2-37, and
5.2-38.
The Farley Unit 2 pressurizer was designed and fabricated in accordance with the requirements
of the 1971 Edition of the ASME Code Section III through the Winter 1970 Addendum. The
current 10 CFR 50 Appendix G requirements, which became effective on August 16, 1973, are
more stringent than the applicable code requirements for Farley Unit 2.
The Farley Units 1 and 2 replacement steam generators were designed and fabricated in
accordance with the requirements of the 1989 edition of the ASME Code Section III which
includes provisions consistent with 10 CFR 50 Appendix G.
The actual fracture toughness data for RCPB pressure-retaining applications in the steam
generators and pressurizer are tabulated in table 5.2-39. In all cases, the applicable ASME
Code requirements, as well as the intent of 10 CFR 50 Appendix G, are satisfied.
SA 508 Class 2a material and SA 533 Class 2 material was used in the Farley Unit 2
pressurizer. Neither of these materials was used in primary-side (RCPB) pressure retaining
applications of the Farley Unit 2 steam generators. The fracture toughness data for these
materials are included in table 5.2-39. The adequacy of the fracture toughness properties of
these materials has been documented in reference 10.
The following discussion demonstrates that the intent of the Appendix G, Paragraph III.B.3
requirements is satisfied.
FNP-FSAR-5
5.2-42 REV 27 4/17 Reactor Vessel - Combustion Engineering (CE) calibrated Charpy V notch test machines in accordance with Watertown Arsenal Standards every 6 months. Temperature instruments, calibrated in accordance with ASTM-E-23, were purchased every 3 months.
These calibrations were performed in accordance with the requirements of the ASME Code 1968 Edition through Summer 1970 Addenda (Appendix IX-221 and 260), which is the
applicable Code for the Farley Unit 2 reactor vessel. The Charpy V notch test machine
calibrations were recorded. The temperature instrument calibrations were not recorded;
however, thermometers qualified to ASTM standards were purchased, used for the certified time
period, and replaced with new qualified thermometers.
CE required that all of its vendors who furnished materials or parts (for Farley Unit 2) to be on
an approved vendors list. Each vendor was requi red to have a quality control system in accordance with #N-335 of the 1968 ASME Code through Summer 1970 Addenda. Periodic
audits of these vendors were performed by CE QA personnel.
It should be noted that the Farley Unit 2 reactor vessel was partially furnished by B & W.
Material furnished by B & W was accepted on the basis of material certifications; therefore, no
QA audits were performed for those by CE.
Pressurizers - Charpy V-notch test machine ca libration at W Tampa plant was performed yearly using samples obtained from Watertown Arsenal. Temperature instrument calibration was
performed with standards traceable to the National Institute of Standards and Technology.
All material suppliers have been either survey ed by ASME auditors or W Tampa Plant Product Assurance to obtain supplier certifications. A sampling of one of the major material suppliers
indicated that Charpy V-notch test machine calibrations were recorded and that calibrated
temperature instruments were purchased (a s replacements) on a yearly basis.
The following discussion demonstrates that the intent of the Appendix G, Paragraph III.B.4
requirements is satisfied.
Reactor Vessel - The personnel performing the Charpy testing at Combustion Engineering were qualified by schooling, training, and many years of experience. Their qualifications to perform this work have been certified by qualified supervisory personnel. This meets the requirements of the applicable ASME Code 1968 Edition through Summer 1970 Addenda (Appendix IX 221d).
Pressurizer - Charpy impact tests were per formed at W Tampa Plant by Level III and Level II personnel who had a minimum of 5 years directly-related testing experience.
Steam Generators - The replacement steam generators are constructed to an edition of the ASME code,Section III that has incorporated provisions consistent with 10 CFR 50 Appendix G.
Compliance with the applicable portions of the ASME Code,Section III for the design and
testing of pressure boundary materials, welding, and weld filler metal provides a vessel in
compliance with the fracture toughness requirements of 10 CFR 50, Appendix G.
FNP-FSAR-5
5.2-43 REV 27 4/17 The heatup and cooldown curves for Units 1 and 2 are based on the fracture toughness
properties of each vessel, as given in tables 5.2-24 and 5.2-25 and the calculation methods
described in WCAP-14040-NP-A, Revision 2 (30) as amended by the methodology approved by NRC letter dated March 31, 1998. Tables 5.2-24 and 5.2-25 indicate that the original maximum
reference nil-ductility temperatures (RT NDT) of the Unit 1 and 2 reactor vessels are not higher than +60°F. Allowable pressures as a function of the rate of temperature change and the actual
temperature relative to the vessel RT NDT are established according to the methods given in Appendix G, "Fracture Toughness for Protection Against Failure," of Section XI of the ASME
Pressure Vessel and Boiler Code. As required by the Technical Specifications, curves showing
RCS heatup and cooldown limitations are provided in the Pressure Temperature Limits Report (PTLR).
These curves are based on temperature scale relative to the limiting RT NDT of the vessels, including appropriate estimates of RT NDT caused by radiation.(a) Predicted RT NDT values are derived by using the recommendations of Regulatory Guide 1.99, Revision 2 (16), and the maximum fluence at 1/4 and 3/4 of vessel wall thickness corresponding to the beltline material
in question and the selected service period. Heatup and cooldown limits are then calculated
using the most limiting RT NDT for the selected service period in accordance with the methods described in WCAP-14040-NP-A, Revision 2 (30) as amended by the methodology approved by NRC letter dated March 31, 1998. The selection of such a limiting RT NDT ensures that all components in the RCS are operated conservatively in accordance with ASME code
requirements. The heatup and cooldown curves are in compliance with the NRC acceptance
criteria contained in Appendices G and H of 10 CFR Part 50 and Regulatory Guide 1.99, Revision 2.
The results of the radiation surveillance programs are used to verify that the predicted RT NDT is appropriate, or to make necessary changes if the RT NDT determined from the surveillance capsules is different from the predicted RT NDT.
The use of an RT NDT that includes a RT NDT to account for radiation effects on the core region material automatically provides additional conser vatism for the nonirradiated regions. However, 10 CFR 50, Appendix G requires licensees to address the metal temperature of the closure head flange and vessel flange in the determination of heatup and cooldown rate limitations.
This rule states that the minimum metal temperature of the closure flange regions must be at
least 120°F higher than the limiting RT NDT for these regions when the pressure exceeds 20% of the preservice hydrostatic test pressure. The rule also states that a plant-specific fracture
evaluation may be performed to justify less limiting requirements. As a result, a fracture
analysis was performed for Unit 2.
(17) The fracture analysis results are also applicable to Unit 1 since the pertinent parameters are identical for both units. The impact of the 10 CFR 50,
- a. Reactor vessel neutron embrittlement was evaluated as a TLAA for license renewal (see
chapter 18, subsection 18.4.1).
FNP-FSAR-5
5.2-44 REV 27 4/17 Appendix G rule and the results of the fracture analysis are reflected in the heatup and
cooldown curves shown in the PTLR.
Changes in fracture toughness of the core region plates, weldments, and associated
heat-affected zones because of radiation damage will be monitored by a surveillance program which conforms with ASTM E-185-82, "Standard Practice for Conducting Surveillance Tests for
Light-Water Cooled Nuclear Power Reactor Vessels." The evaluation of the radiation damage
in this surveillance program is based on pre-irradiation and post-irradiation testing of Charpy V-notch and tensile specimens carried out during the lifetime of the reactor vessel. Specimens
are irradiated in capsules located near the core mid-height and removed from the vessel at
specified intervals. For additional details of the irradiation surveillance program, refer to
paragraph 5.4.3.6.
See paragraph 5.4.3.7 for a discussion of reactor vessel annealing.
The unstabilized austenitic stainless steel material specifications used for the RCS boundary, systems required for reactor shutdown, and systems required for emergency core cooling, are
listed in tables 5.2-20 and 5.2-21.
The unstabilized austenitic stainless steel material specifications used for the reactor vessel
internals that are required for emergency core cooling for any mode of normal operation, or
under postulated accident conditions, and for core structural load bearing members, are listed in
table 5.2-23.
All of the above tabulated materials are procured in accordance with the specification
requirements and include supplemental requirements of the applicable ASME Code rules.
It is required that all austenitic stainless steel materials used in the fabrication, installation, and
testing of nuclear steam supply components and systems be handled, protected, stored, and
cleaned according to recognized and accepted methods and techniques.
The rules covering these controls are stipulated in the following Westinghouse Electric
Corporation process specifications. These process specifications supplement the equipment specification and purchase order requirements of every individual austenitic stainless steel
component or system which Westinghouse procur es for a nuclear steam supply system, FNP-FSAR-5
5.2-45 REV 27 4/17 regardless of the ASME Code Classification. They are also given to the architect-erector and to
the owner of the power plant for use within their scope of supply and activity.
To ensure that manufacturers and installers adhere to the rules in these specifications, surveillance of operations by Westinghouse personnel is conducted either in-residence, at the
manufacturer's plant and the installer's construction site, or during periodic engineering and
quality assurance visitations and audits at these locations.
The process specifications which establish these rules and which are in compliance with the
more current American National Standards Institute N-45 Committee specifications are as
follows:
Process Specification Number
82560HM Requirements for Pressure Sensitive Tapes for Use on Austenitic Stainless Steels.
83336K Requirements for Thermal Insulation Used on Austenitic Stainless Steel Piping and Equipment.
83860LA Requirements for Marking of Reactor Plant Components and Piping.
84350HA Site Receiving Inspection and S torage Requirements for Systems, Material and Equipment.
84351NL Determination of Surface Chloride and Fluoride on Austenitic Stainless Steel Materials.
85310QA Packaging and Preparing Nuclear Components for Shipment and Storage.
292722 Cleaning and Packaging Requirements of Equipment for Use in the NSSS.
597756 Pressurized-Water Reactor Auxiliary Tanks Cleaning Procedures.
597760 Cleanliness Requirements During Storage, Construction, Erection and Startup Activities of Nuclear Power Systems.
The cleaning and contamination protection procedures for Bechtel-supplied equipment made of
austenitic stainless steel materials are detailed in the individual equipment specifications.
These procedures assure that austenitic stainless steel material is cleaned and protected
against contaminants capable of causing stress corrosion cracking. The cleaning procedures
consist of the removal of all mill scale, rust, grease, and other contaminants and cleaning with
both solvent and demineralized water.
During storage of austenitic stainless steel components, special precautions are taken to ensure
suitable environmental conditions. Strictly controlled working procedures are followed in order
to maintain the necessary cleanliness of all austenitic stainless steel components.
FNP-FSAR-5
5.2-46 REV 27 4/17 All of the austenitic stainless steels listed in tables 5.2-20, 5.2-21, and 5.2-23 are procured from
raw material producers in the final heat-treated condition required by the respective ASME Code
Section II material specification for the particular type or grade of alloy.
All of the wrought austenitic stainless steel alloy raw materials which require corrosion testing
after the final mill heat treatment are tested in accordance with ASTM A 393, using material test
specimens obtained from specimens selected for mechanical testing. The materials are
obtained in the solution-annealed condition.
The unstabilized austenitic stainless steels used in the RCPB and components are listed in
tables 5.2-20 and 5.2-21.
These materials are used in the as-welded condition, as discussed in paragraph 5.2.5.2. The
control of the water chemistry is stipulated in paragraph 5.2.3.4. These chemistry controls, coupled with the satisfactory experience wi th components and internals using unstabilized austenitic stainless steel materials which have been post-weld heat treated above 800°F, show
acceptability of these heat-treatments for stainless steel in the PWR chemistry environment (7). Actual observations of post-weld, heat-treated, austenitic stainless steel after actual operation
indicate no effects of such treatments. Internals heat-treated above 800°F from H. B. Robinson, Unit 2, Zorita, Connecticut Yankee, San Onofre, Beznau 1, R. E. Ginna, Yankee Rowe, Selni, and SENA have been examined after service and show acceptable material condition.
The unstabilized austenitic stainless steels used for core structural load bearing members and
component parts of the RCPB are processed and fabricated using the most practicable and
conservative methods and techniques to avoid partial or local severe sensitization.
After the material has been heat-treated as described in paragraph 5.2.5.2, the material is not
heated above 800°F during subsequent fabrication, except as described in paragraph 5.2.3.2
and in the paragraphs below.
Methods and material techniques that are used to avoid partial or local severe sensitization are
as follows:
FNP-FSAR-5
5.2-47 REV 27 4/17 A. Nozzle Safe Ends
Weld deposit with Inconel (Ni-Cr-Fe weld metal F No. 43), then attach safe-end after final post-weld heat-treatment, which was used for the reactor vessel, pressurizer, accumulators, and replacement steam generators.
B. For internals, the austenitic stainless steels have been given a stress-relieving treatment above 800°F; i.e., a high temperature stabilizing procedure is used.
This is performed in the temperature range of 1600-1900°F, with holding times
sufficient to achieve chromium diffusion to the grain boundary regions to limit the
effects of sensitization on Cr-carbide precipitation in the grain boundary. The
stainless nozzles on the pressurizer were given a post-weld treatment associated
with the fabrication of the head. No intergranular tests are planned because of
satisfactory service experience, as noted in paragraph 5.2.5.4.
C. All welding is conducted using those procedures that have been approved by the ASME Code rules of Section III and IX.
D. All welding procedures have been qualified by nondestructive and destructive testing according to the ASME Code rules of Section III and IX.
When these welding procedure tests are being performed on test welds that are made from base metal and weld metal mater ials that are from the same lot(s) of materials used in the fabrication of components, additional testing is frequently
required to determine the metallurgical, chemical, physical, corrosive, etc.,
characteristics of the weldment. The additional tests that are conducted on a
technical case basis are as follows: light and electron microscopy, elevated
temperature mechanical properties, chemical check analysis, fatigue tests, intergranular corrosion tests, and static and dynamic corrosion tests within
reactor water chemistry limitations.
E. The following welding methods have been tested individually and in multiprocess combinations as outlined in (D) above, using these prudent energy input ranges
for the respective method, as calculated by the following formula:
H = E x I x 60 S
where E = volts
I = amperes
S = travel speed in in./min
H = joules/in.
FNP-FSAR-5
5.2-48 REV 27 4/17 ENERGY INPUT RANGE WELDING PROCESS METHOD (Kilojoules/in.)
Manual shielded tungsten arc 20 to 50 Manual shielded metallic arc 15 to 120
Semi-automatic gas shielded metallic arc 40 to 60 Automatic gas shielded tungsten arc-10 to 50 hot wire Automatic submerged arc 60 to 140
Automatic electron beam - soft vacuum 10 to 50 F. The interpass temperature of all welding methods is limited to 350°F maximum.
G. All full-penetration welds require inspections in accordance with Article 6 of the ASME Section III Code rules.
In general, it is not feasible to remove samples from fabricated production components to
prepare specimens for retest to determine the susceptibility to intergranular attack. These tests
are performed only on test welds when meaningful results would predicate production material
performance and are as described in paragraph 5.2.5.5. No intergranular tests are planned
because of satisfactory service experience (see paragraph 5.2.5.5).
The austenitic stainless steel welding material used for joining Class 1 pipe, pump, fittings, and
applications is described in paragraph 5.2.3.1. The welding material conforms to ASME Weld
Metal Analysis A-8 for all applications. Bare weld filler metal materials, including consumable
inserts used in inert gas welding processes, conform to ASME SFA-5.9 and are procured to
contain not less than 5-percent delta ferrite. All weld filler metal materials used in flux-shielded
welding processes conform to ASME SFA-5.4 or SFA-5.9 and are procured in a wire flux
combination to be capable of providing not less than 5-percent delta ferrite in the deposit.
All welding materials are tested by the fabricator using the specific process(es) and the
maximum welding energy inputs to be employed in production welding. These tests are in
accordance with the requirements of ASME Section II, Material Specification, and, in addition, include delta ferrite determinations. The delta ferrite determinations are made by calculation
using the "Schaeffler or Modified Schaeffler Constitution Diagram for Stainless Steel Weld
Metal."
FNP-FSAR-5
5.2-49 REV 27 4/17 When subsequent in-process delta ferrite determinations are required, and since the welding
material conformance is proved by the init ial material testing described above, any of the recognized methods for measurement of delta ferrite is acceptable by mutual agreement. In
these instances, sound welds (as determined by visual, penetrant and volumetric examinations)
that display more than 1-percent-average delta ferrite content are considered to be
unquestionably acceptable. All other sound welds are considered acceptable also, providing
there is no evidence of deviation from qualified procedure parameters or use of malpractices. If
evidence of the latter prevails, sampling for chemical and metallurgical analysis is required to
determine the integrity and acceptability of the weld(s). The sample size is required to be
10 percent of the welds, but not less than 1 weld, in the particular component or system. If any
of these weld samples are defective, that is, fail to pass bend tests as prescribed by ASME Section IX, or if the chemical analysis deviates from the material specification, then all
remaining welds are sampled and all defective welds are removed and replaced.
All other applications use type 308 or type 316 which normally contain 3 to 15% delta ferrite and
1 to 5% delta ferrite in the deposit analyses, respectively. The successful experience with
austenitic stainless steel welds for these applications, supplemented by nondestructive
examination, provides assurance for avoiding microfissuring in welds.
The qualification of welding procedures is discussed in paragraph 5.2.5.5.
The integrity of the reactor coolant pump flywheel is assumed on the basis of the following design and quality assurance procedures.(a)
The calculated stresses at operating speed are based on stresses caused by centrifugal forces.
The stress resulting from the interference fit of the flywheel on the shaft is 2000 psi at 0 speed, but this stress becomes 0 at approximately 600 rpm because of radial expansion of the hub. The primary coolant pumps run at approximately 1190 rpm and may operate briefly at
overspeeds up to 109% (1295 rpm) during loss of outside load. For conservatism, however, 125% of operating speed was selected as the design speed for the primary coolant pumps. The
flywheels are given a preoperational test of 125% of the maximum synchronous speed of the
motor.
The flywheel consists of two plates, approximately 5-in. and 8-in. thick, bolted together. Each
plate is fabricated from electro-slag refined A-533 Grade B Class I steel . Supplier certification
reports are available for all plates and demonstrate the acceptability of the flywheel material on
the basis of the following requirements of NRC Regulatory Guide 1.14.
- a. Reactor coolant pump flywheel fatigue is evaluated as a TLLA for license renewal (see
chapter 18, paragraph 18.4.2.3).
FNP-FSAR-5
5.2-50 REV 27 4/17 A. The nil-ductility transition (NDT) temperature of the flywheel material should be no higher than +10°F.
B. The Charpy V notch (C v) energy level in both the parallel and normal orientation with respect to rolling direction of the material should be at least 50 ft-lb at the
normal operation temperature of the flywheel.
A lower bound K ID reference curve (see figure 5.2-11) has been constructed from dynamic fracture-toughness data generated in A533 Grade B Class I steel (8). All data points are plotted on the temperature scale relative to the NDT temperature. The construction of the lower bound
below which no single test point falls, combined with the use of dynamic data when flywheel
loading is essentially static, together represent a large degree of conservatism. Reference of
this curve to the guaranteed NDT temperature of +10°F indicates that, at the predicted flywheel
operating temperature of 110°F, the minimum fracture toughness is in excess of 100 KSI-in 1/2. This conforms to NRC Regulatory Guide 1.14 requirement (6.1) that the dynamic stress
intensity factor must be at least 100 KSI-in 1/2.
Flywheel blanks are flame-cut from the plate, with allowance for exclusion of heat affected
material. The finished flywheels are subjected to 100-percent volumetric ultrasonic inspection.
The finished machined bores are also subjected to magnetic particle or liquid penetrant
examinations.
Precautionary measures taken to preclude miss ile formation from primary coolant pump components ensure that the pumps will not produce missiles under any anticipated accident
condition. Each component of the primary pump motors has been analyzed for missile
generation. Any fragments of the motor rotor would be contained by the heavy stator. The
same conclusion applies to the pump impeller because the small fragments that might be
ejected would be contained by the heavy casing.
Thus it is concluded that flywheel plate materials are suitable for use and can meet NRC
Regulatory Guide 1.14 acceptance criteria on the basis of suppliers certification data.
The reactor coolant leakage detection system provides the capability of detecting the presence of significant radioactive or nonradioactive leakage from the RCLs to the containment
atmosphere during normal operation. Variations in the particulate activity, gaseous activity, and
specific humidity of the containment atmosphere above a preset level give positive indications in the control room to the reactor operators. A leakage estimate is then made from either the
functional variation during the transients or the new steady state. These leakage detection
provisions are sufficiently sensitive so that small increases in leakage rates can be detected
while the total leakage rate is still below a value consistent with safe operation of the plant.
FNP-FSAR-5
5.2-51 REV 27 4/17 Instrumentation is also provided to monitor pressu re and flow conditions in auxiliary system lines penetrating the RCPB. Protection is also provided against possible overpressurization
resulting from excessive check valve leakage, either by relief valves or by circuits permitting periodic tests. Provisions are also made to isolate the primary grade water within the
containment should excessive intersystem leakages occur.
The particulate and gaseous activity are monitored by the containment air particulate and
radiogas monitors. The specific humidity is monitored by the condensate measuring system and the dewpoint temperature system.
The reactor coolant leakage detection system consists of the air particulate monitor, the
radiogas monitor, condensate measuring devices, and humidity detectors.
A. Containment Air Particulate Monitor
This monitor takes continuous-flowing air samples from the containment atmosphere and measures the air particulate beta radioactivity. The samples are
drawn outside the containment in a closed, sealed system and are monitored by a beta scintillation detector assembly. The fixed filter paper collects 99 percent of the particulate matter > 1.0 in size, which is viewed by a hermetically sealed combination photomultiplier tube. This monitor is series connected to the containment radioactive gas monitor and uses the pumping system common to both. This monitor has a measuring range of 10
-12 to 10-6 Ci/cc.
The detector assembly is in a completely closed housing. The signal will be processed by the skid mounted microprocessor and will be transmitted to the
radiation monitoring system cabinet in the control room. Lead shielding is
provided to reduce the background radiation to a level where it does not interfere
with the detector's sensitivity.
The activity is indicated on meters and monitored by the plant process computer.
High activity alarm indications are di splayed on the radiation monitoring cabinets.
Local alarms provide operational status of supporting equipment such as pumps, motors, and flow and pressure controllers. The activity is indicated by a control
and display module instead of a meter.
The sensitivity of the air particulate monitor to an increase in reactor coolant leak-rate is dependent upon the magnitude of the normal baseline leakage into
the containment.
For cases where the baseline reactor coolant leakage falls with the detectable limits of the air particulate monitor, the instrument can be adjusted to
alarm on leakage increase from 2-to-5 times the baseline value.
FNP-FSAR-5
5.2-52 REV 27 4/17 B. Containment Radioactive Gas Monitor
This monitor measures the gaseous beta radioactivity in the containment by taking the continuous air sample from the containment atmosphere. The sample
first passes through the air particulate monitor where particulate matter is
removed, and then through a closed, sealed system to a gas monitor assembly.
After passing through the gas monitor, the gas sample is returned to the
containment atmosphere.
Each sample is constantly mixed in fixed, shielded volumes, where it is viewed by photomultiplier tubes. This monitor has a measuring range of 10
-12 to 10-3 Ci/cc.
The detector is in a completely enclosed housing containing a beta-sensitive photomultiplier tube mounted in a constant gas volume container. Lead shielding
is provided to reduce the background radiation level to a point where it does not
interfere with the detector's sensitivity.
The detector outputs are transmitted to the radiation monitoring system cabinets in the control room. The activity is indicated by a control and display module and
monitored by the Analog Data Managem ent System computer or the plant process computer. High activity alarm indications are displayed on the control
board annunciator in addition to the radiation monitoring system cabinets. Local
alarms annunciate the supporting equipment's operational status.
The air particulate and radiogas monitors have a pump unit common to both monitors.
The pump unit consists of:
- 1. A pump to obtain the air sample.
- 2. A digital mass flow indicator/controller to adjust and indicate the flow rate.
- 3. A flow-control valve to provide steady flow.
- 4. A flow-alarm assembly to provide low- and high-flow alarm signals.
The air particulate and radiogas monitors will be qualified to function following a safe shutdown earthquake as described in paragraph 11.4.2.2.3.
C. Specific Humidity Monitoring Devices
The containment specific humidity monitoring devices offer another means of detection of leakage into the containment. The devices, namely the condensate
measuring system and the dewpoint monitor s, are not as sensitive as the air particulate and the radiogas monitors, but have the advantage of being sensitive
to vapor originating from all sources:
the reactor coolant system, the steam FNP-FSAR-5
5.2-53 REV 27 4/17 system, and the feedwater system. T hus, these devices are able to detect leakage from nonradioactive or radioactive sources during the initial period of
plant operation when the coolant activity may be low.
D. Condensate Measuring System
The condensate measuring system permi ts measurements of liquid runoff from the drain pans under each containment fan cooler unit. It consists of a vertical
standpipe, valves, and standpipe level instrumentation installed in the drain
piping of the reactor containment fan cooler unit.
The condensation from the containment coolers flows to the vertical standpipes.
A differential pressure transmitter provides standpipe level signals. The system
provides measurement capability of condensate runoff by monitoring standpipe level increase versus time.
Depending on the number of reactor containment fan cooler units in operation, the sum of the drainage flowrate from each operating cooler unit represents the
total normal condensation. With the initiation of an additional or abnormal leak, the containment atmosphere humidity and condensation runoff rate will begin to
increase, the water level will rise in the vertical pipe, and the high-condensate
level alarm will be actuated.
The containment specific humidity will increase proportionally to time and leakage until the dewpoint is reached at the fan cooler units cooling coils. With
the increasing specific humidity, the heat removal capacity needed to cool the air
steam mixture to its dewpoint temperature decreases. Increases in specific
humidity and available heat removal capacity from the cooling coils will result in
added condensate flow. The condensate flowrate then is a function of specific
humidity. Through accurate measurements of condensate level and dewpoint
variations or RCS inventory (i.e., water inventory balance calculations), a reliable
indication of the reactor coolant leakage rate can be made.
Detection of hot water leakage can be obtained from the condensate flow and dewpoint increase during the transient. A better estimate of leakage can be
determined from the steady-state condensate flow when equilibrium has been
reached. The device will alarm on a level equivalent to or below a condensate
flowrate corresponding to a postulated 1.0 gal/min RCS leakage considering a
flashing factor of approximately 40%.
E. Dewpoint Temperature Monitoring
The dewpoint measuring system consists of ten dew-cell elements. One element is located at the inlet and outlet of each of the containment fan cooler units and
one each in the upper and lower compartments of the containment. The basis of
operation of these elements is the behavior of a hygroscopic salt in the presence
of water vapor.
FNP-FSAR-5
5.2-54 REV 27 4/17 When dry lithium chloride is exposed to the atmosphere under average room conditions, it will absorb moisture and dissolve, forming a salt solution. If this
solution is heated, the water tends to escape back to the atmosphere. A state of
equilibrium is reached at a temperature where the tendency of water to escape is
equal to the tendency of the salt to absorb moisture. At this equilibrium point, the
temperature of the salt and the saturated solution (temperature of the dew-cell
element) is a measure of the partial pressure of the water vapor surrounding
atmosphere, i.e., dewpoint temperature. The range of the dewpoint temperature
measuring system is 50° to 130°F. Its accuracy is +1°F.
Because of the slow response of containment atmosphere specific humidity for an abnormal
increase device, dewpoint temperature recordings may prove useful in establishing the location
and the history of the leak.
Positive indications in the control room of leakage of coolant from the RCS to the containment
are provided by equipment which permits conti nuous monitoring of containment air activity and humidity.
The limits for reactor coolant leakage are delineated in the FNP Technical Specifications.
The total, normally expected leakage from the RCS is expected to be about 40 lb/day. The
sensitivities and response times of subsystems are as follows:
The following system sensitivities are based upon discrete values of input parameters and
assumptions as documented in Westinghouse WCAP-8009 (31). Actual performance may vary based on plant conditions.
A. Containment Air Particulate Monitor
The containment air particulate monitor is the most sensitive instrument available for detection of reactor coolant leakage into the containment. This instrument is
capable of detecting particulate activity in concentration as low as 10
-9 to 10-6 c/cc of containment air sampled.
Leakage rates of about 0.01 gal/min to leaks > 10 gal/min can be observed, assuming specific values of corrosion-product activity and no fuel cladding damage. Assuming a corrosion-product activity (Fe, Mn, Co, Cr) of 0.4 c/cc, a FNP-FSAR-5
5.2-55 REV 27 4/17 low, but detectable, background of containment air particulate activity, and
complete dispersion of leaking radioactive solids into the air, leak rates of about
0.01 gal/min are detectable within 50 min after they occur. A 1.0 gal/min leak
would be detectable within 0.5 min.
B. Containment Radioactive Gas Monitor
The containment radioactive gas monitor is inherently less sensitive (threshold at 10-6 c/cc) than the containment air particulate monitor and would function in the event that significant reactor coolant gaseous activity exists because of fuel cladding defects. Assuming a reactor coolant gaseous activity of 22 c/cc (corresponding to about 0.1 percent fuel defects), the occurrence of a leak of
1.0 gal/min would double a zero leakage background in approximately 40 min.
C. Condensate Measuring and Dewpoint Monitoring System
These systems provide indications that allow determination of leakage losses from water and steam systems within the containment. The condensate measuring system collects and measures the moisture condensed from the
containment atmosphere onto the cooling coils of the containment cooling units.
The dewpoint and condensate measuring system provide a dependable and
accurate means of measuring integrated total leakage, including leaks from the
cooling coils themselves. Condensate flows 0.1 gal/min can be identified by the condensate measuring system. Dewpoi nts can be observed to within 1°F. Leaks smaller than 1 gal/min can be measured by periodic observation of the
level changes in the condensate collection system. If leakage is to another
closed system, it will be detected by the plant radiation monitors and/or inventory control. For a condensate high-level alarm setpoint corresponding to 0.1 gal/min
per cooler flowrate, the occurrence of a 1.0-gal/min leak would be detected within
1 h, assuming approximately 40 percent of the leakage enters the containment
atmosphere as vapor.
The 1-gal/min maximum permissible leakage rate from unidentified sources within the RCPB is
well below the leakage rates calculated for critical through-wall cracks in pipes of 3-in. diameter
and larger. The lengths of through-wall cracks that are calculated to leak 0.5- gal/min in 2-in.
lines, 1 gal/min in 3-in. lines, and 2 gal/min in lines of 4-in. diameter and larger are given in
reference 1. Included in this report are the ratios of critical through-wall cracks to computed
lengths for these leakage values, as a function of pipe diameter and wall thickness based on the
application of the principles of fracture mechanics, as well as the mathematical model and data
used in the analyses.
Although the 1-gal/min maximum permissible unidentified leakage rate is larger than the 0.5-
gal/min leakage rate analyzed for cracks in 2-in. lines, core cooling analyses have shown that
for "small breaks," that is, for breaks up to the equivalent of the cross-sectional area of a 4-in.-
diameter line, acceptable peak clad temperature results are obtained.
FNP-FSAR-5
5.2-56 REV 27 4/17 The maximum allowable total leakage from the RCPB from other than controlled sources is
10 gpm. This leakage rate is approximately 10 per cent of the makeup control system while in the automatic mode of operation. Normal background leakage (40 lb/day) does not influence
this value significantly. Gross leakage or condensate overflow accumulates in the containment
sump, which has a removal rate of 50 gal/min, a more than adequate capacity.
The methods described in paragraph 5.2.7.1.1 will allow detection of RCPB leakages occurring
within the containment. The location of specific leaks will in general have to be determined
visually, although the systems that are indicating that a leak exists should aid in determining its
source.
The air particulate and radiogas monitors are provided with their own test circuitry which tests
electronics and the photomultiplier tube. The electronics test provides a precalibrated pulse
signal that can be recorded. A remotely-operated long half-life radiation check source is
provided with energy emission ranges similar to the radiation energy spectra being monitored.
The source-strength is sufficient to cause approximately 30 percent of full-scale indication.
These units can be tested at any time at the discretion of the operator. For the condensate
measuring system, the level indicators will be calibrated prior to plant operation.
The provision of adequate access was verified by a review of all the drawings applicable to the layout and arrangement of the Reactor Coolant and Associated Auxiliary Systems within the
boundaries established in accordance with the requirements of IS-120.
The general design features of the nuclear plant reactor vessel, system layout, and other major
primary coolant components to ensure compliance with the requirements of IS-141 and IS-142
are as follows (Specific provision to be made for inspection access in the design of the reactor
vessel, system layout and other major prim ary coolant components also is listed):
FNP-FSAR-5
5.2-57 REV 27 4/17 A. All reactor internals are completely removable. The tools and storage space required to permit reactor internals removal for these inspections have been
provided.
B. The reactor vessel shell in the core area is designed with a clean, uncluttered cylindrical inside surface to permit future positioning of test equipment without
obstruction.
C. The reactor vessel cladding was improved in finish by grinding to the extent necessary to permit meaningful examination of the vessel welds and adjacent
base metal in accordance with the code.
D. The cladding-to-base-metal interface was ultrasonically examined to ensure satisfactory bonding to allow the volumetric inspection of the vessel welds and
base metal from the vessel inside surface.
E. The reactor closure head is stored in a dry condition on the operating deck during refueling, allowing direct access for inspection.
F. The insulation on the vessel closure and lower heads is removable, allowing access for the visual examination of head penetrations.
G. All reactor vessel studs, nuts, and washers are removed to dry storage during refueling, allowing inspection in parallel with refueling operations.
H. Access holes are provided in the core barrel flange, allowing access for the remote visual examination of the clad surface of the vessel without removal of
the lower internals assembly.
I. Removable plugs are provided in the primary shield, providing access for the surface and visual examination of the primary nozzle safe-end welds.
J. Manways are provided in the s team generator channel head to provide access for internal inspection.
K. A manway is provided in the pressurizer top head to allow access for internal inspection.
L. The insulation covering all component and piping welds and adjacent base metal is designed for ease of removal and replacement in areas where external
inspection will be planned.
M. Removable plugs are provided in the primary shield concrete above the main coolant pumps to permit removal of the pump motor to provide internal inspection
access to the pumps.
N. The primary loop compartments are designed to allow personnel entry during refueling operations, to permit direct inspection access to the external portion of
piping and components.
FNP-FSAR-5
5.2-58 REV 27 4/17 The use of conventional, nondestructive, volumetric test techniques can be applied to the
inspection of all primary loop components except for the reactor vessel. The reactor vessel
presents special problems because of the radiation levels and the remote underwater
accessibility to this component.
As indicated above, the only sophisticated remote inspection equipment currently required is for
inspection of the reactor vessel. The baseline inspection was performed by Westinghouse, utilizing a remote reactor vessel ultrasonic inspection tool to perform the code-required
inspection of the circumferential and longitudinal shell welds, the flange-to-vessel weld, the
ligaments between the flange holes, the nozzle-to-vessel welds, and the nozzle-to-safe-end-
to-pipe welds. Because of access restrictions imposed by the location of the lower radial core
support blocks, only 50 percent of the total length of the lower head-to-shell weld was examined
from inside the vessel. The remainder of the weld was examined manually from the outside of
the vessel.
The vessel inspection tool has two major components, the superstructure which holds the
examination assembly and the examination assemb ly which delivers the various ultrasonic transducers to the desired work point in the vessel. Design of the tool permits precise
positioning for accurate scanning of the examination volume. Reconfiguration of the
examination assembly and transducers permits examination of the desired welds and
components in the reactor vessel, e.g., circumferential shell welds, flange-to-shell welds, lower
head welds, and nozzle examinations. Appropriate ultrasonic transducers are installed in the
examination assembly to detect and size indications.
For reactor vessel automated ultrasonic examinations, the data recording and positioning
system is fully integrated with the vessel inspection tool to provide precise locations for
state-of-the art sizing and characterizing of indications.
For manual ultrasonic examinations, such as the examination of a circumferential pipe weld, procedures are used to ensure that inspection results are recorded in such a manner which will
avoid any ambiguity in interpretation. Procedures specify the location of weld reference points and the way in which indications must be recorded with respect to these reference points.
The data from various examinations is collected into a comprehensive report tabulating all of the
results in sufficient detail to ensure repeatability for each examination.
FNP-FSAR-5
5.2-59 REV 27 4/17
The only areas where it is expected that high radiation levels will prohibit the access of
personnel for direct examination of component areas or systems is the reactor vessel. The special design provisions and tooling required to perform the code-required examinations in
these areas have been discussed above.
FNP-FSAR-5
5.2-60 REV 27 4/17
The Units 1 and 2 Inservice Inspection Programs have been established in accordance with 10 CFR 50.55a(g). The inservice inspections are performed in accordance with Section XI of
the ASME Code with certain exceptions whenever specific written relief or alternative to ASME
Code requirements are granted by the NRC. The First Ten-Year Inservice Inspection (ISI)
Program for each unit was established to meet, to the extent practical, the requirements of the 1974 Edition through the Summer 1975 Addenda of the ASME Code Section XI. Following
completion of the first 10-year interval for Unit 1, the Second Ten-Year ISI Program was
established. Following completion of the second 10-year interval, the Third Ten-Year ISI
Program is effective from December 1, 1997 through November 30, 2007. The Code of record for the third 10-year interval is the ASME Code,Section XI, 1989 Edition.
For Unit 2, by letter dated August 31, 1988, the NRC granted approval of an exemption from certain requirements of 10 CFR 50.55a, regarding the update of the ISI Program. Rather than
requiring update of the Unit 2 ISI Program to the Code of record in effect on July 30, 1991, the
NRC approved updating the program 3 years early, as provided by 10 CFR 50.55a(g)(4)(iv), in
- a. The ISI Program is credited as a licens e renewal aging management program (see chapter 18, subsection 18.2.1).
FNP-FSAR-5
5.2-61 REV 27 4/17 conjunction with the Unit 1 ISI Program which was previously updated to the ASME Code,Section XI, 1983 Edition through Summer 1983 Addenda. Exemption to the requirements of
10 CFR 50.55a(g)(4)(ii) extended the date of record by which the Unit 2 program is required to
be updated from July 30, 1991, which marks the completion of the first 10 years of commercial
operation of Unit 2, through November 30, 1997, the date which marks completion of the
second 10 years of commercial operation for Unit 1. In this way, the Code of record in effect through November 30, 1997, for Unit 1-the ASME Code,Section XI, 1983 Edition through
Summer 1983 Addenda-is also applicable to Unit 2.
The updated Unit 2 program went into effect in March 1989 during the Unit 2 sixth refueling
outage and will continue through the third 40-month period of the first 10-year interval and
remain in effect through the first and second 40-month periods of the second 10-year interval
until December 1, 1997, the completion date for the second 10-year interval for Unit 1. At this
time, a new updated Unit 2 program is in effec t from December 1, 1997 through November 30, 2007. In this way, the Code of record in effect for the third 10-year interval for Unit 1-the
ASME Code 1989 Edition-is also applicable to Unit 2.
Beginning at the fourth ISI interval, inservice inspection of the metallic liner and the pressure
retaining concrete structure of the containments of both units meet the requirements of Subsections IWE and IWL of the appropriate edition of ASME Section XI as described in the
Containment Inspection Plan.
Reactor vessel examinations in accordance with the First Ten-Year ISI Program for each unit
included the Mandatory Appendix I requirements entitl ed "Ultrasonic Examination." Reactor vessel examinations in accordance with the Unit 1 Second Ten-Year ISI Program and the Unit 2
updated ISI Program are bound to Article 4 of Section V entitled "Ultrasonic Examination When
Dimensioning of Indications is Required."
Reactor vessel examinations performed under the Unit 1 Third Ten-Year ISI program and the new Unit 2 Updated Program (effective from
December 1, 1997 through November 30, 2007) will be accomplished per the requirements of Appendix I, Article I-2100 of the 1989 Edition of ASME Section XI.
While maintaining these requirements as the technical basis of the examination programs, Units 1 and 2 comply with the Augmented Reactor Vessel Examination Program developed in
response to NRC Generic Letter 83-15 and Regulatory Guide 1.150, Revision 1. This program
for Units 1 and 2 was submitted by letter from F. L. Clayton, Jr. (APC) to S. A. Varga (NRC) of
October 26, 1983.
The Units 1 and 2 Inservice Testing (IST) Programs have been established in accordance with
10 CFR 50.55a(g). The First Ten-Year IST Program for each unit was established to meet, to
the extent practical, the requirements of the 1974 Edition through the Summer 1975 Addenda of the ASME Code Section XI. Following completion of the first 10-year interval for Unit 1, the FNP-FSAR-5
5.2-62 REV 27 4/17 Second Ten-Year IST Program was established. Following completion of the second 10-year
interval, the Third Ten-Year IST Program is effective from December 1, 1997, through
November 30, 2007. FNP received approval to use the ASME OM Code - 1990 Edition as the
Code of record for the third 10-year interval.
For Unit 2, by letter dated August 31, 1988, the NRC granted approval of and exemption from certain requirements of 10 CFR 50.55a regarding the requirements for updating the IST
Program. Rather than requiring update of the Unit 2 IST Program to the Code of record in effect
on July 30, 1991, the NRC approved updating the program 3 years early, as provided by 10 CFR 50.55a(g)(4)(iv), in conjunction with the Unit 1 IST Program which was previously updated to the ASME Code,Section XI, 1983 Edition through Summer 1983 Addenda.
Exemption to the requirement of 10 CFR 50.55a(g)(4)(ii) extended the date of record by which
the Unit 2 program is required to be updated from July 30, 1991, which marks the completion of
the first 10 years of commercial operation of Unit 2 through November 30, 1997, the date which
marks completion of the second 10 years of commercial operation for Unit 1. In this way, the Code of record in effect through November 30, 1997, for Unit 1-the ASME Code,Section XI, 1983 Edition through the Summer 1983 Addenda-is also applicable to Unit 2.
The updated Unit 2 program went into effect in March 1989 during the Unit 2 sixth refueling
outage and will continue through the third 40-month period of the first 10-year interval and
remain in effect through the first and second 40-month periods of the second 10-year interval
until December 1, 1997, the completion date for the second 10-year interval for Unit 1. At this
time, a new updated Unit 2 program is in effec t from December 1, 1997 through November 30, 2007. In this way, the Code of record in effect for the third 10-year interval for Unit 1-
ASME OM Code-1990 Edition is also applicable to Unit 2.
The ASME Boiler and Pressure Vessel Code, Sections V and XI, were used for the design
ultrasonic calibration blocks as described in the Units 1 and 2 ISI Programs.
The metal impact monitor system in the Farley Nuclear Plant is designed to detect loose parts in the RCS. The developmental prototype of the Westinghous e metal impact monitor is installed in
the R. E. Ginna Plant to evaluate the long-term performance of the system in an operating plant.
The system consists of a detector, preamplifier, signal processor (with audio and record
outputs), and display alarm. The system is a general maintenance aid and is not necessary for
safe operation of the Farley Nuclear Plant.
Detector The detectors are high temperature accelerometers mounted on each steam generator and on
the reactor vessel.
FNP-FSAR-5
5.2-63 REV 27 4/17 Preamplifier
Preamplification of the detector signal is performed with a signal conditioning amplifier. This
consists of a remote charge preamp and a signal conditioner. The remote charge preamp is
located in close proximity to the accelero meter, on the outside of the primary system component. The signal conditioner is located in the MIMs cabinet outside of containment.
These amplifiers are used to convert the low level accelerometer charge signal to a voltage
signal for transmission to the signal processing equipment outside the containment.
Signal Processor and Display
The metal impact monitor was designed so that rate, as well as energy, of metal debris impact can be monitored continuously. Rate and amplitude latching-type alarms are displayed on the
front panel of the monitor. Common alarm outputs are provided for connection to the main
control room annunciator panel. An audio system produces the sound equivalent in parallel to
the impact signal.
FNP-FSAR-5
5.2-64 REV 27 4/17 1. Szyslowski, J. J., and Salvatori, R., "Determination of Design Pipe Breaks, for the Westinghouse Reactor Coolant System," WCAP-7503, Revision 1, February 1972.
- 2. Logcher, R. D., and Flachsbart, B. B., "ICES STRUDL-II, The Structural Design Language Frame Analysis," MTT-ICES-R68-1, Vol. 1, November 1968.
- 3. Takeuchi, K., et. al., "MULTIFLEX, A FORTRAN-IV Computer Program for Analyzing Thermal-Hydraulic-Structure System Dy namics," WCAP-8708-PA-V1 (Proprietary), WCAP-8709-A (Non-Proprietary) September, 1977.
- 4. Cooper, K., Starek, R. M., and Miselis, V., "Overpressure Protection for Westinghouse Pressurized Water Reactors," WCAP-7769, Revision 1, June 1972.
- 5. Nay, J. A., "Process Instrumentation for We stinghouse Nuclear Steam Supply Systems," WCAP-7671, April 1971.
- 6. "Fracture Toughness Requirements," Branch Technical Position - MTEB No. 5-2, Chapter 5.3.2 in Standard Review Plan for the Review of Safety Analysis Reports for
Nuclear Power Plants, LWR Edition, NUREG-0800, 1981.
- 7. Hazelton, W. S., "Sensitized Stainless Steel in Westinghouse PWR Nuclear Steam Supply Systems," WCAP-7735, August 1971.
- 8. Shabbits, W. O., "Dynamic Fracture Toughness Properties of Heavy Section A 533 Grade B Class 1 Steel Plate," WCAP-7623, December 1970.
- 9. "Final Seismic Response Spectra - Joseph M. Farley Nuclear Units 1 and 2 Alabama Power Company Containment and Internal Structure, Bechtel Report No. 7597 03/20, December 1972.
- 10. Logsdon, W. A., Begley, J. A., and Gottshall, C. L., "Dynamic Fracture Toughness of ASME SA508 Class 2a ASME SA533 Grade A Class 2 Base and Heat Affected Zone
Material and Applicable Weld Metals," WCAP-9292, March 1978.
- 11. DeSalvo, G. J. and Swanson, J. A., ANSYS User's Manual, Engineering Analysis Systems Report, October 1, 1972.
- 12. Mendelson, A., Plasticity: Theory and Application, MacMillian, New York, 1968.
- 13. Terek, E., "Farley Units 1 and 2 Heatup and Cooldown Limit Curves for Normal Operation and PTLR Support Documentation," WCAP-14689, Revision 4, April 1998.
- 14. Peter, P. A., et. al., "Analysis of Capsul e W from the Alabama Power Company Farley Unit 1 Reactor Vessel Radiation Surveillance Program," WCAP-14196, February 1995.
FNP-FSAR-5
5.2-65 REV 27 4/17 15. Terek, E., Lloyd, T. M., Albertin, L., Analysis of Capsule X from the Alabama Power Company Joseph M. Farley Unit 2 Reac tor Vessel Radiation Surveillance Program, "WCAP-12471, Revision 0, December 1989.
- 16. Regulatory Guide 1.99, Revision 2, " Radiation Embrittlement of Reactor Vessel Materials," U.S. Nuclear Regulatory Commission, May 1988.
- 17. Miller, J. C., "Response to NRC Comments on Farley Unit 2," ALA-85-706, July 31, 1985.
- 18. NUREG-0737, "Clarification of TMI Action Plan Requirements," NRC, November 1980.
- 19. "Application of RECARS/MODI for Calculation of Safety and Relief Valve Discharge Piping Hydrodynamic Loads," EPRI NP-2479, Final Report, December 1982.
- 20. SNC Letter from Dave Morey to the NRC, "Response to NUREG-0737, Item II.D.1,"
dated July 26, 1994.
- 21. NRC letter from Byron L. Siegel to D. N. Morey, "Safety Evaluation Regarding NUREG-0737, Item II.D.1., Sub-Item 8," dated July 18, 1995.
- 22. Westinghouse Report WCAP-10105, "Review of Pr essurizer Safety Valve Performance as Observed in the EPRI Safety Valve and Relief Valve Test Program," June 1982.
- 23. Westinghouse Program ITCH Sun Workstation Version 1.0 Configured 9/7/92.
- 24. W Report WCAP-9924, "ITCHVALVE Code Description and Verification," M. A. Berger &
K. S. Howe, July 1982.
- 25. Westinghouse Internal Letter SE&PT-CSE-465 1/31/90, "Release of Thermal Hydraulic and Related Computer Codes (ITCHVALVE, ITCHVENT, FORFUN, KJTRPLT2) for Production Use on the Cray X-MP."
- 26. Westinghouse Program FORFUN Sun Workstation Version 1.0, Configured 10/30/92.
- 27. W Computer Program WECAN, WECAN/PLUS User's Manual, Dec. 1, 1990, First Edition, Westinghouse Electric Corp., Pittsburgh, PA.
- 28. Letter, F. L. Clayton, Jr., to S. A. Varga, NRC, "Joseph M. Farley Nuclear Plant, Units 1 and 2 NUREG-0737, Item II.D.1," November 4, 1982.
- 29. Letter, E. A., Reeves, NRC to R. P. McDonald, Alabama Power Co., "Completion of Review of Item II.D.1 NUREG-0737 Safety and Relief Valve Testing for Joseph M.
Farley Nuclear Plant Unit Nos. 1 and 2," December 16, 1986.
- 30. Andracnek, J. D., et. al., "Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cool down Limit Curves," WCAP-14040-NP-A, Revision 2, January 1996.
FNP-FSAR-5
5.2-66 REV 27 4/17 31. R. M. Norris and C. W. Thomas, "Joseph M. Farley, Reactor Coolant Pressure Boundary Leak Detection System," WCAP-8009, April 2, 1973.
- 32. Letter by Southern Company Services, Inc., File: REA 97-1463, Log: FP 97-0530, October 24, 1997, "Steam Generator Snubber Elimination", Chris A. Byrd.
- 33. Letter by Southern Company Services, Inc., File: REA 97-1463, Log: FP 97-0536, October 28, 1997, "Steam Generator Snubber Elimination-Synthesized Time History,"
Chris A. Byrd.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-1 HARDSHIP EXCEPTIONS TO 10 CFR 50.55a As-constructed NRC-required Component Code Code Differences Reactor coolant pumps 1968 Pump and Valve ASME B & PV Code, (1) Major defect mapping- (Unit 1)(a) Code, March 1970 Addenda Section III 1971 1968 P & V; 1/5 of the Edition casting thickness. 1971 B & PV; lesser of 10% of casting thickness or 3/8 in.
(2) Hydrostatic test pressure.
Pumps will be tested to 4100 psi instead of 4900 psi.(c) Class I 1968 Pump & Valve ASME B & PV Code, Major differences in formal Valves Codes plus Addenda Section III 1971 documentation required.
Edition plus Summer 1971 Addenda Thermocouple (b) ASME B & PV Code, Formal documentation Lead Section III 1968 requirements Appurtenances Edition plus all Addenda thru Summer 1970
Notes a. The reactor coolant pumps for FNP Unit Number 2 will conform with ASME B & PV Code,Section III, 1971 Edition plus Summer 1972 Addenda.
- b. Prior to the Summer 1970 Addenda of the 1968 Edition of the ASME B & PV Code Section III, no specific code requirements existed for the internals vessel appurtenances. In lieu of any formal code requirements, the internals vessel appurtenances were designed to meet the intent of the 1968 Edition of the ASME B & PV Code Section III.
- c. Summer 1972 Addenda hydrostatic test pressure requirement is 3750 psi.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-2 (SHEET 1 OF 2)
SUMMARY
OF REACTOR COOLANT SYSTEM DESIGN TRANSIENTS Normal Conditions Occurrences Heatup and cooldown at 100°F/h 200 (each) (pressurizer cooldown 200°F/h)
Unit loading and unloading at 18,300 (each) 5 percent of full power/min
Step load increase and decrease 2,000 (each) of 10 percent full power
Large step load decrease, with 200 steam dump
Steady-state fluctuations Infinite Upset Conditions Loss of load, without immediate 80 turbine or reactor trip Loss of power (blackout with 40 natural circulation in the reactor coolant system)
Loss of flow (partial loss 80 of flow one pump only)
Reactor trip from full power 400 Inadvertent auxiliary spray 10 One-half safe shutdown earthquake 5 Faulted Conditions(a) Main reactor coolant pipe break 1
Steam pipe break 1
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-2 (SHEET 2 OF 2)
Test Conditions Occurrences
Steam generator tube rupture (included above in reactor trip from full power)
Safe shutdown earthquake 1 Turbine roll test 10 Hydrostatic test conditions Primary Side 5 Secondary side 10 Primary side leak test 50
- a. In accordance with the ASME Nuclear Power Plant Components Code, faulted conditions are not included in fatigue evaluations.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-3 LOAD COMBINATIONS AND OPERATING CONDITIONS Load Combination Operating Condition Normal condition transients, Normal condition deadweight
Upset condition transients, Upset condition deadweight, 1/2 SSE
Faulted condition transients, Faulted condition deadweight, SSE, or SSE and pipe rupture loads
FNP-FSAR-5 REV 21 5/08 TABLE 5.2-4 (SHEET 1 OF 2)
LOADING CONDITIONS AND STRESS LIMITS: CLASS 1 COMPONENTS Loading Conditions(a) Stress Intensity Limits Note Normal (a) P m S m (b) P L 1.5 s m (c) P m (or P L) + P B 1.5 S m 1 (d) P m (or P L) + P B + Q 3.0 S m 2 Upset condition (a) P m S m (b) P L 1.5 S m (c) P m (or P L) + P B 1.5 S m 1 (d) P m (or P L) + P B + Q 3.0 S m 2 Faulted condition Faulted condition limits in table 5.2-6 P m = primary general membrane stress intensity. P L = primary local membrane stress intensity.
P B = primary bending stress intensity.
Q = secondary stress intensity.
S m = stress intensity value from ASME B&PV Code,Section III, Nuclear Vessels. S y = minimum specified material yield (ASME B&PV Code,Section III, Table N-421 or equivalent).
- a. Emergency condition is not included since none have been specified.
FNP-FSAR-5 REV 21 5/08 TABLE 5.2-4 (SHEET 2 OF 2)
NOTES FOR TABLE 5.2-4 Note 1: The limits on local membrane stress intensity (P 1.5S m) and primary membrane plus primary bending stress intensity (P m (or P L) + P 1.5S m) need not be satisfied at a specific location if it can be shown by means of limit analysis or by tests that the specified loadings do not exceed 2/3 of the lower bound collapse load as per paragraph N-417.6(b) of the ASME B&PV Code,Section III, Nuclear Vessels.
Note 2: In lieu of satisfying the specific requirements for the local membrane (P L 1.5S m) or the primary plus secondary stress intensity (P m(or P L) + P + Q 3S m) at a specific location, the structural action may be calculated on a plastic basis and the design will be considered to be acceptable if shakedown occurs, as opposed to continuing deformation, and if the deformations which occur prior to shakedown do not exceed specified limits, as per paragraph N-417.6(a) (2) of the ASME B&PV Code,Section III, Nuclear Vessels.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-5 LOADING CONDITIONS AND STRESS LIMITS: NUCLEAR POWER PIPING Loading Conditions (b) Stress Intensity Limits (a)
Normal condition (a) P mS m (b) P L1.5 S m (c) P m (or P L) + P1.5 S m (d) P m (or P L) + P + P e + Q3.0 S m (e) P e3.0 S m Upset condition (a) P mS m (b) P L1.5 S m (c) P m (or P L) + P 1.5 S m (d) P m (or P L) + P + P e + Q3.0 S m (e) P e3.0 S m Faulted condition Faulted condition limits are shown in table 5.2-6. P m = primary general membrane stress intensity.
P L = primary local membrane stress intensity.
P B = primary bending stress intensity.
P e = secondary expansion stress intensity.
Q = secondary membrane plus bending stress intensity.
S m = allowable stress intensity from ASME Boiler & Pressure Vessel Code,Section III, Nuclear Power Plant Components, 1971.
- a. Alternatively, the rules and simplified analysis of sub sub articles NB-3640 and NB-3650 of ASME B&PV Code,Section III, Nuclear Power Plant Components, 1971, may be used in lieu of the stated
equations.
- b. Emergency condition is not included since none have been specified.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-6 FAULTED CONDITION STRESS LIMITS FOR CLASS 1 COMPONENTS System (or Subsystem)
Components Stress Limits for Analysis Analysis Components Test P m P m + P b Elastic Smaller of Smaller of 2.4 S m and 0.70 S u 3.6 S m and 1.05 S u Note (b)
ELASTIC Plastic Larger of Larger of 0.70 S u or 0.70 S ut or S y 1/3(S u - S y) S y + 1/3 (S ut - S y) Note (c)
Note (c) 0.8 L T Limit Analysis 0.9 L 1 Notes (a and c)
Plastic Larger of 0.70 S U Larger of 0.70 S ut Notes or or (c and d)
PLASTIC Elastic S + 1/3 (S u - S y) S + 1/3 (S ut - S y)
Notes:
- a. L 1 = Lower bound limit load with an assumed yield point equal to 2.3 S m .
- b. These limits are based on a bending shape factor of 1.5 for simple bending cases with different shape factors, the limits will be changed proportionally.
- c. When elastic system analysis is performed, the effect of component deformation on the dynamic system response should be checked.
- d. L T = The limits established for the analysis need not be satisfied if it can be shown from the test of a prototype or model that the specified loads (dynamic or static equival ent) do not exceed 80 percent of L T, where L T is the ultimate load or load combi nation used in the test. In using this method, account should be taken of the size effect and dim ensional tolerances similitude relationships) which may exist between the actual component and the tested models to assure that the loads obtained fr om the test are a conservative representation of the load carrying capability of the actual component under postulated loading for faulted conditions.
S y = Yield stress at temperature.
S u = Ultimate stress from engineering stre ss-strain curve at temperature.
S u = Ultimate stress from true stress-strain curve at temperature.
S m = Stress intensity from ASME Section III at temperature.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-7 ALLOWABLE STRESSES FOR PRIMARY EQUIPMENT SUPPORTS Loading Conditions Stress Limits Normal AISC, Seventh Edition (a), Part 1, Allowable Stresses Upset AISC, Seventh Edition Part 1, Allowable Stresses Faulted Stresses yield strength of material.
Local yielding is permitted but limited so that the structural integrity of the system is maintained.
As an alternative to the above, 80 per-cent of L T (see table 5.2-6) may be used.
- a. Specifications for the design, fabrication and erection of structural steel for buildings.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-8 (SHEET 1 OF 2)
ACTIVE AND INACTIVE (c) VALVES IN THE REACTOR COOLANT SYSTEM PRESSURE BOUNDARY Classification Actuation A-Active Environmental System Location Line Type Size Type I-Inactive Design Criteria(b)
RCS 8010 A,B, Pressurizer Safety 6-in. System pressure A (Internal fluid C safety (to PRT)
(over set point) characteristics specified) RCS 0460 Letdown Globe 3-in. Air-operated A 1,2 RCS 0459 Letdown Globe 3-in. Air-operated A 1,2 CVCS 8378 Charging Check 3-in. p A 2,3 CVCS 8347 Charging Check 3-in. p A 2,3 CVCS 8153, Excess Globe 1-in. Air-operated A (a) 1,2 8154 letdown CVCS 8377 Aux. spray Check 2-in. p A (a) 1,2 CVCS 8145 Aux. Spray Globe 2-in. Air-operated A (a) 1,2 SIS 8998 A,B, SIS injection Check 6-in. p A 2,3 C SIS 8973 A,B, RHR supply Check 6-in. p A 2,3 C SIS 8948 A, B, Accumulator Check 12-in. p A 2, 3 C disch. to C.L.
SIS 8956 A, B, Accumulator Check 12-in. p A 2, 3 C disch. to C.L.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-8 (SHEET 2 OF 2)
Classification Actuation A-Active Environmental System Location Line Type Size Type I-Inactive Design Criteria(b) SIS 8998 A, B, Cold leg Check 6-in. p A 2, 3 C LHSI SIS 8997 A, B, Cold leg Check 2-in. p A 2, 3 C HHSI SIS 8993 A, B, C Hot leg conn.
Check 6-in. p A 2, 3 SIS 8988 A, B Hot leg conn.
Check 6-in. p A 2, 3 CVCS 8346 Alternate Check 3-in. p A 2, 3 charging CVCS 8348 A, B, C RCP Seal Check 2-in. p A 2, 3 8367 A, B, C injection CVCS 8379 Alternate Check 3-in. p A 2, 3 charging SIS 8990 A, B, C HHSI Hot leg Check 2-in. p A 2, 3 8992 A, B, C HHSI Hot leg 8995 A, B, C HHSI Cold leg.
WDS 8057 A, B, C RCDT Drain Isolation 2-in. Manual I 2, 3 8058 A, B, C
- a. There is a possibility that these valves may be open when an accident occurs.
- b. Environmental Design Criteria
- 1. Ambient Temperature: 50°-150°F 2. Ambient Atmosphere: 8-15 psia, 100 percent Relative Humidity, 50 R/hr - Gamma Radiation 3. Ambient Temperature: 120°-150°F
- c. All other valves in this Reactor Coolant Pressure Boundary are considered inactive and are shown on FSAR project drawings D-175037 Sh. 1, D-175037 Sh. 2, D-175037 Sh. 3, D-205037 Sh. 1, D-205037 Sh. 2, and D-205037 Sh. 3.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-9 STRESSES CAUSED BY MAXIMUM STEAM GENERATOR TUBESHEET PRESSURE DIFFERENTIAL (2485 PSIG)
(600°F) Stress Computer Value Allowable Value Primary membrane stress 24,356 psi 37,000 psi
(.9 Sy) Primary membrane plus Primary bending stress 54,946 psi 55,600 psi (1.35 Sy)
Notes:
- 1. In addition to the foregoing evaluation, elasto-plastic limit analysis of the tubesheet-head-shell combination indicates a limit pressure of 3050 psi at operating temperature.
- 2. These values are for Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-10 STEAM GENERATOR PRIMARY-SECONDARY BOUNDARY COMPONENTS (b) Condition: 100% Load Operation - 2485/885 psig (a) 650/600°F Normal Operation Stress Limits Inside Limit Stress Limit Inside Surface Stress Center Limit Center Limit Center Surface Stress Location Description Outer Limit Stress Limit Outer Surface Stress 7 Jct of short 3 S m 80,100 -10,063 psi cyl with S m 26,700 + 8,597 psi tubesheet 3 S m 80,100 +27,247 psi 8 1/2 through 3 S m 80,100 + 9,514 psi short cyl S m 26,700 + 8,597 psi discontinuity 3 S m 80,100 + 7,670 psi 9 Jct of short 3 S m 80,100 +10,740 psi cyl with S m 26,700 + 8,597 psi shell 3 S m 80,100 6,443 psi 10 On shell 3 S m 80,100 +10,269 psi S m 26,700 + 8,597 psi 3 S m 80,100 + 6,912 psi 11 On shell 3 S m 80,100 + 9,746 psi S m 26,700 + 8,597 psi 3 S m 80,100 + 7,435 psi 12 Jct of pri 3 S m 80,100 +58,701 psi short cyl with S m 26,700 +14,528 psi tube plate 3 S m 80,100 -29,646 psi 13 1/2 through 3 S m 80,100 +50,836 psi prim short S m 26,700 +14,528 psi cyl discon.
3 S m 80,100 -27,781 psi 14 Jct of pri 3 S m 52,200 +42,286 psi short cyl S m 19,400 +14,528 psi with head 3 S m 52,200 -13,231 psi
- a. Based on 1600 psig design pressure differential.
- b. These values are for Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-11 STEAM GENERATOR PRIMARY-SECONDARY BOUNDARY COMPONENTS Condition: Primary Hydrotest - 3107/0 psig Primary Axial Primary Membrane Membrane Stress Description Code Limit Stress Limit Intensity Jct of short cyl with 0.9 Sy 45,000 0 psi tubesheet
1/2 through short cyl 0.9 Sy 45,000 0 psi discontinuity
Jct of short cyl with 0.9 Sy 45,000 0 psi shell On shell 0.9 Sy 45,000 0 psi On shell 0.9 Sy 45,000 0 psi Jct of pri short cyl 0.9 Sy 45,000 18,158 psi with tube plate 1/2 through prim short 0.9 Sy 45,000 18,158 psi cyl discon.
Jct of pri short cyl 0.9 Sy 36,000 18,158 psi with head
Note:
These values are for Model 51 Steam Generators. The values will be updated for the
replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-12 STEAM GENERATOR PRIMARY-SECONDARY BOUNDARY COMPONENTS Condition: Secondary Chamber Hydrotest - 0/1356 psig Primary Axial Primary Membrane Membrane Stress Location Description Code Limit Stress Limit Intensity 7 Jct of short cyl with tubesheet
.9 S y 45,000 13,169 psi 8 1/2 through short cyl discontinuity
.9 S y 45,000 13,169 psi 9 Jct of short cyl with shell
.9 S y 45,000 13,169 psi 10 On shell
.9 S y 45,000 13,169 psi 11 On shell
.9 S y 45,000 13,169 psi 12 Jct of pri short cyl with tube plate
.9 S y 45,000 0 psi 13 1/2 through pri short cyl discon
.9 S y 45,000 0 psi 14 Jct of pri short cyl with head
.9 S y 36,000 0 psi
Note:
These values are for Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) afte r the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-13 STEAM GENERATOR PRIMARY-SECONDARY BOUNDARY COMPONENTS (a) (b) Condition: Loss of Secondary Pressure (Steam Line Break)
Faulted Condition 2485/0 psig, 660°F Primary Membrane Stress Emergency Condition Limits Primary Membrane Location Description Code Limit Stress Stress 7 Jct of short cyl with S y 41,112 0 psi tubesheet 8 1/2 through short cyl S y 41,112 0 psi discontinuity 9 Jct of short cyl with S y 41,112 0 psi shell 10 On shell S y 41,112 0 psi 11 On shell S y 41,112 0 psi 12 Jct of pri short cyl S y 41,112 14,528 psi with tube plate 13 1/2 through pri short S y 41,112 14,528 psi cyl discon 14 Jct of pri short cyl S y 29,000 14,528 psi with head
- a. Complete Tubesheet Structure Complex also evaluated on Limit Analysis Basis.
- b. These values are for Model 51 Steam Generators. The values will be updated for the
replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-14 51,500 SQ FT STEAM GENERATOR USAGE FACTORS (INDIVIDUAL TRANSIENTS)
PRIMARY AND SECONDARY BOUNDARY COMPONENTS (b) JUNCTION OF HEAD AND TUBESHEET AND DIVIDER PLATE IN TUBESHEET ON TUBESHEET FACE NO. OF INLET Outlet INLET OUTLET NO. TRANSIENT CYCLES SASR(a) SASH SRSH SASR SASH SRSH SASR SASH SRSH SASR SASH SRSH 1. Reactor Cooldown 200 .008 .01 0 .008 .01 0 0 0 0 0 0 0 2. Loading Unloading 18,300 0 0 0 0 0 0 0 0 0 0 0 0 3. Small Step Increase 2,000 0 0 0 0 0 0 0 0 0 0 0 0 4. Small Step Decrease 2,000 0 0 0 0 0 0 0 0 0 0 0 0 5. Large Step Decrease 200 0 0 0 0 0 0 0 0 0 0 0 0 6. Loss of Load 80 0 0 0 0 0 0 0 0 0 0 0 0 7. Loss of Power 40 0 0 0 0 0 0 0 0 0 0 0 0 8. Loss of Flow 80 .008 .009 0 .008 .016 0 0 0 0 0 0 0 9. Reactor Trip 400 0 0 0 0 .001 0 0 0 0 0 0 0 10. Reactor Cooling Pipe Break 1 0 0 0 0 0 0 0 0 0 0 0 0 11. Steam-Line Break 1 0 0 0 0 0 0 0 0 0 0 0 0 12. Primary Hydrotest 5 .004 .007 0 .004 .007 0 0 0 0 0 0 0 13. Secondary Hydrotest 5 0 0 0 0 0 0 0 0 0 0 0 0 14. Turbine Roll Test 10 0 0 0 0 0 0 0 0 0 0 0 0 UNPERFORATED JUNCTION OF IN HEAD Outlet OUTER RING SHELL TO INLET PRIM. SEC. TUBESHEET NO. TRANSIENT SASR SASH SRSH SASR SASH SRSH INLET OUTLET INLET OUTLET MAHH HAHR HNHR 1. Reactor Cooldown .009 .003 0 .009 .003 0 0 0 0 0 0 0 0 2. Loading Unloading 0 0 0 0 0 0 0 0 0 .019 .082 .056 0 3. Small Step Increase 0 0 0 0 0 0 0 0 0 0 0 0 0 4. Small Step Decrease 0 0 0 0 0 0 0 0 0 0 0 0 5. Large Step Decrease 0 0 0 0 0 0 0 0 0 .002 .001 0 6. Loss of Load 0 0 0 .001 .001 0 0 0 .001 .002 .006 .005 0 7. Loss of Power 0 0 0 0 0 0 0 0 0 0 0 0 0 8. Loss of Flow .011 .003 .001 .012 .006 0 0 0 .002 .004 .002 .002 0 9. Reactor Trip 0 0 0 0 0 0 0 0 .001 .005 .012 .008 0 10. Reactor Cooling Pipe Break 0 0 0 0 0 0 0 0 0 0 0 0 0 11. Steam-Line Break 0 0 0 0 0 0 0 0 0 0 0 .001 0 12. Primary Hydrotest .005 .002 0 .005 .002 0 0 0 .001 .001 .001 .001 0 13. Secondary Hydrotest 0 0 0 0 0 0 0 0 0 0 .001 .002 0 14. Turbine Roll Test 0 0 0 0 0 0 0 0 0 0 0 0 0
- a. Principal Steam Differential Codes. b. These values are for the Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model
- 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-15 51,500 SQ FT STEAM GENERATOR USAGE FACTORS (INDIVIDUAL TRANSIENTS)
CENTER OF TUBESHEET (a) No. Of Primary Inlet Primary Outlet Secondary Inlet Secondary Outlet No. Transient Cycles Angle Angle Angle Angle 1 Heatup-cooldown 200 0° 15° 30° 45° 60° 75° 90° 0° 15° 30° 45° 60° 75° 90° 0° 15° 30° 45° 60° 75° 90° 0° 15° 30° 45° 60° 75° 90° 2 Loading-unloading 18,300 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 3 Small step increase 2,000 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 4 Small step decrease 2,000 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 5 Large step decrease 200 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 6 Loss of load 80 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 7 Loss of power 40 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 8 Loss of flow 80 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 9 Reactor trip 400 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 10 Reactor cooling pipe break 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 11 Steam line break 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 12 Primary hydrotest 5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 13 Secondary hydro 5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 14 Turbine roll test 10 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
- a. These values are for the Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
° Angular location around perforation
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REV 21 5/08 TABLE 5.2-16 TUBESHEET STRESS ANALYSIS RESULTS FOR 51,500 SQ FT STEAM GENERATORS (a) Maximum Primary Membrane Maximum Effective Plus Primary Bending Ligament Membrane Conditions Average Ligament Stress (psi)
Stress (psi) 100% normal operation 2485/885 psi 33,979 (40,050)(1) 15,853 (26,700)(2) 650/600°F Primary hydrotest 3107/0 psi 67,300 (67,500)(3) 30,365 (45,000)(4) 100°F Secondary hydrotest 0/1356 psi 29,811 (67,500)(3) 13,159 (45,000)(4) 100°F Steam line break 2485/0 psi 56,785 (Limit)(5) 24,356 (Limit)(5) (fault condition) 660°F
- a. These values are for Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
Parentheses Indicate Code Allowable Stress 1. 1.5 S m
- 2. 1.0 S m
- 3. 1.35 S y
- 4. 0.9 S y
- 5. Limit Analysis Results Apply
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-17 LIMIT ANALYSIS CALCULATION RESULTS TABLES OF STRAINS, LIMIT PRESSURES, AND FATIGUE EVALUATIONS FOR 51,500 SQ FT STEAM GENERATORS Circum- Peak Allowable Usage Limit Meridional ferential Stress Number of Number of Factor Pressure Case Location Strain, in/in Strain, in/in Intensity, psi Cycles, N 1 Cycles, N 2 N 2/N 1 psi Hot Channel/Primary Shell .0188 -.000559 508,000 46 10 .22 2500/0 psi Tubesheet/Secondary Shell -.00193 .00602 83,700 5,000 10 .0020 3,158 650 F Tubesheet Center .00159 .00159 77,400 6,600 10 .0015 Cold Hydro Tubesheet/Primary Shell .0145 -.000537 434,000 80 5 .053 3105/0 psi Tubesheet/Secondary Shell -.00220 .000684 106,000 3,500 5 .0014 3,887 70 F Tubesheet Center .00177 .00177 95,400 5,000 5 .0010 Cold Hydro with Tubesheet/Primary Shell .00730 -.000348 218,000 500 5 .010 Secondary Pressure Tubesheet/Secondary Shell -.000962 .000560 50,700 40,000 5 .0001 4,401 3105/700 Tubesheet Center .00147 .00147 79,000 8,000 5 .0005 psi 70 F Hot Hydro Tubesheet/Primary Shell .00777 -.000407 222,000 400 50 .13 2485/0 psi Tubesheet/Secondary Shell -.00176 .000551 80,900 7,000 50 .0071 3,354 400 F Tubesheet Center .00148 .00148 76,300 8,500 50 .0059 Note:
These values are for the Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-18 RELIEF VALVE DISCHARGE TO THE PRESSURIZER RELIEF TANK Reactor Coolant System 3 Pressurizer safety valves D-175037 Sh.2 (Unit 1) 2 Pressurizer power-operated D-205037 Sh.2 (Unit 2) relief valves Safety Injection System 1 SIS discharge to hot leg D-175038 Sh.2 (Unit 1) 2 SIS discharge to cold legs D-205038 Sh.2 (Unit 2)
Residual Heat Removal System 2 RHR pump suction line from D-175041 Sh.1 (Unit 1)
RCS hot legs D-205041 Sh.1 (Unit 2)
Chemical and Volume Control System 2 Charging pump suction D-175039 Sh.6 (Unit 1)
D-205039 Sh.2 (Unit 2) 1 Seal-water return line D-175039 Sh.1 (Unit 1)
D-205039 Sh.1 (Unit 2) 1 Letdown line D-175039 Sh.1 (Unit 1)
D-205039 Sh.1 (Unit 2)
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-19 REACTOR COOLANT SYSTEM DESIGN PRESSURE SETTINGS (PSIG)
Hydrostatic test pressure (cold) 3107 Design pressure 2485 Safety valves open 2485 High pressure reactor trip 2385 Power relief valves open 2335 High controller output alarm 100 psig + controller setpoint (nominal 2335) High pressure alarm 2310 Proportional spray full on 2310 Pressurizer spray valve begin to open 2260 Proportional spray off 2260 Proportional heaters off 2250 Design nominal operating 2235 Proportional heaters full on 2220 Backup heaters on 2210 Low pressure alarm 2185 P11 interlock 2000 Low pressure reactor trip 1865 Pressurizer level and pressure coincidence 1850
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-20 (SHEET 1 OF 3)
REACTOR COOLANT SYSTEM BOUNDARY MATERIALS CLASS 1 PRIMARY COMPONENTS Reactor Vessel Component
Shell (other than core region)
SA-533 B, Class 1 (vacuum treated)
Shell plates (core region)
SA-533 B, Class 1 (vacuum treated)
Head forging SA-508 Grade 3, Class 1
Shell, flange, and nozzle forgings nozzle safe ends SA-508 Class 2
SA-182 Type F316
CRDM, Instrumentation port and RVLIS head
adapters and vent pipe (lower part)
SB-167 UNS No. 6690
RVLIS and instrumentation port housings SA-182, F316
Vent pipe (upper part)
SA-312, Type 316
Instrumentation tube appurtenances - lower head SB-166 or -167 and SA-182 Type F304, F304L, or F316
Closure studs SA-540 Class 3 Gr B23 or B24
Closure nuts SA-540 Class 3 Gr B23 or B24
Closure washers SA-540 Class 3 Gr B23 or B24
Core support pads SB-166 with carbon less than 0.10%
Vessel supports, seal ledge SA-516 Gr 70 quenched and tempered or SA-533 Gr A, B, or C. (Vessel supports may
be of weld metal buildup of equivalent
strength.)
Head lifting lugs SA-533, Type B, Class 1
Steam Generator Components
Pressure forgings SA-508 Class 3 or 3a Nozzle safe ends SA-336 Class F Type 316LN FNP-FSAR-5
REV 21 5/08 TABLE 5.2-20 (SHEET 2 OF 3)
Tubes SB163 Ni-Cr-Fe, annealed
Closure bolting and studs SA193 Gr B-7
Closure nuts SA194 Gr 7
Pressurizer Components
Pressure plates SA533 Gr A, Class 2
Pressure forgings SA508 Class 2 or 3
Nozzle safe ends SA182 or 376 Type 316 or 316L and Ni-Cr-Fe Weld Metal F-Number 43
Closure bolting SA193 Gr B-7
Pressurizer safety valve forgings SA182 Type F316
Reactor Coolant Pump
Pressure forgings SA182 Type F304, F316 or F348
Pressure castings SA351 Gr CF8, CF8A or CF8M
Tube and Pipe SA213, SA376 or SA312 -
Seamless Type 304 or 316
Pressure plates SA240 Type 304 or 316
Bar material SA479 Type 304 or 316
Closure bolting SA193 Gr B7 or B8 or, SA540 Gr B23 or B24 or SA453 Gr 660
Reactor Coolant Piping
Reactor coolant pipe Code Case 1423-1 Gr F304N or 316N, or SA351 Gr CF8A or CF8M centrifugal
castings Reactor coolant fittings SA351 Gr CF8A or CF8M
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-20 (SHEET 3 OF 3)
Branch nozzles SA182 Gr F304 or 316 or Code Case 1423-1 Gr F304N or 316N
Surge line and loop bypass SA-376 Type 304 or 316 or Code Case 1423-1 Gr F304N or 316N
Auxiliary piping 1/2 in. through12 in. and wall
schedules 40S through 80S (ahead of second
isolation valve)
All other auxiliary piping(ahead of second isolation ANSI B36.10
valve)
Socket weld fittings ANSI B16.11
Piping flanges ANSI B16.5
Welding materials SFA 5.4 and 5.9 Type 308 or 308L
Control Rod Drive Mechanism
Pressure housing SA-182 Gr F316
Pressure forgings SA-182 Gr F316
Bar material SA-479 Type 304
Welding materials SFA 5.9 Type 316L
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-22 REACTOR COOLANT WATER CHEMISTRY SPECIFICATION Electrical conductivity Determined by the concentration of boric acid and alkali present.
Solution pH Determined by the concentration of boric acid and alkali present. Expected values range between 4.2 (high boric acid concentration) to10.5 (low boric acid concentration) at 25°C.
Oxygen, ppm maximum Oxygen concentration of the reactor coolant is maintained below 0.1 ppm for plant operation above 250°F. Hydrazine may be used to chemically scavenge oxygen during heatup.
Chloride, ppm, maximum 0.15 Fluoride, ppm, maximum 0.15 Hydrogen, cc(STP)/kg H 2 O 25-50 (power operation)(a) Total suspended solids, 1.0 ppm, maximum
pH control agent (Li 7 0H) 0.20 - 4.36 (power operation)
(ppm Li )
Boric acid, ppm B Variable from 0 to approximately 2500
- a. Hydrogen concentration during transients (including preparation for shutdown, plant restart, etc.)
is controlled per plant procedures based on OEM (Westinghouse) recommendations.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-23 MATERIALS FOR REACTOR VESSEL INTERNALS FOR EMERGENCY CORE COOLING Forgings SA182 Type F304
Plates SA240 Type 304
Pipes SA312 type 304 seamless or SA376 Type 304
Tubes SA213 Type 304
Bars SA479 type 304 & 410
Castings SA351 Gr CF8 or CF8A
Bolting SA(Pending)Westinghouse
PE Spec. 70041EA
Nuts SA193 Gr B-8
Locking devices SA479 type 304
Weld buttering Stainless steel weld metal analysis A-7
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-24 UNIT 1 REACTOR VESSEL TOUGHNESS PROPERTIES Material Cu P Ni T NDT RT NDT Upper Shelf Energy Component Code No. Type (%) (%) (%) (°F) (°F) MWD (c) NMWD(d) Closure head dome B6901 A533,B,CL.1 0.16 0.009 0.50 -30 -20 140 - Closure head segment B6902-1 A533,B,CL.1 0.17 0.007 0.52 -20 -20(a) 138 - Closure head flange B6915-1 A508, CL.2 0.10 0.012 0.64 60(a) 60(a) 75(a) - Vessel flange B6913-1 A508, CL.2 0.17 0.011 0.69 60(a) 60(a) 106(a) - Inlet nozzle B6917-1 A508, CL.2 - 0.010 0.83 60(a) 60(a) - 110 Inlet nozzle B6917-2 A508, CL.2 - 0.008 0.80 60(a) 60(a) - 80 Inlet nozzle B6917-3 A508, CL.2 - 0.008 0.87 60(a) 60(a) - 98 Outlet nozzle B6916-1 A508, CL.2 - 0.007 0.77 60(a) 60(a) - 96.5 Outlet nozzle B6916-2 A508, CL.2 - 0.011 0.78 60(a) 60(a) - 97.5 Outlet nozzle B6916-3 A508, CL.2 - 0.009 0.78 60(a) 60(a) - 100 Nozzle shell B6914-1 A508, CL.2 - 0.010 0.68 30 30(a) 148 - Inter. shell B6903-2 A533,B,CL.1 0.13 0.011 0.60 0 0 151.5 97 Inter. shell B6903-3 A533,B,CL.1 0.12 0.014 0.56 10 10 134.5 100 Lower shell B6919-1 A533,B,CL.1 0.14 0.015 0.55 -20 15 133 90.5 Lower shell B6919-2 A533,B,CL.1 0.14 0.015 0.56 -10 5 134 97 Bottom head ring B6912-1 A508, CL.2 - 0.010 0.72 10 10(a) 163.5 - Bottom head segment B6906-1 A533,B,CL.1 0.15 0.011 0.52 -30 -30(a) 147 - Bottom head dome B6907-1 A533,B,CL.1 0.17 0.014 0.60 -30 -30(a) 143.5 -
Inter. shell long. M1.33 Sub Arc Weld 0.258 0.017 0.165 0(a) -56(e) - - weld seam (19-894A&B)
Inter. to lower G1.18 Sub Arc Weld 0.205 0.011 0.105 0(a) -56(e) - - weld seams (11-894)
Lower shell long. G1.08 Sub Arc Weld 0.197 0.022 0.060 0(a) -56(e) - - weld seams (20-894A&B)
(a) Estimate per NUREG-0800 "USNRC Standard Review Plan" Branch Technical Position MTEB 5-2. (b) Estimated (low nickel weld wire used in fabricating vessel weld seams). (c) Major working direction.
(d) Normal to major working direction.
(e) Estimate per 10 CFR 50.61.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-25 (SHEET 1 OF 2)
UNIT 2 REACTOR VESSEL TOUGHNESS DATA Average Upper Shelf Energy Normal to Principal Principal Working Cu P Ni T NDT RT NDT Working Direction Direction Component Code No. Grade (%) (%) (%) (°F) (°F) (ft-lb) (ft-lb) CL. HD. Dome B7215-1 A533,B,CL.1 0.17 0.010 0.49 -30 16 (a) 83 (a) 128 CL. HD. Flange B7207-1 A508,CL.2 0.14 0.011 0.65 60 (a) 60 (a) >56 (a) >86(c) VES. Flange B7206-1 A508,CL.2 0.10 0.012 0.67 60 (a) 60 (a) >71 (a) >109 Inlet Noz. B7218-2 A508,CL.2
- 0.010 0.68 50 (a) 50 (a) 103 (a) 158 Inlet Noz. B7218-1 A508,CL.2
- 0.010 0.71 32 (a) 32 (a) 112 (a) 172 Inlet Noz. B7218-3 A508,CL.2
- 0.010 0.72 60 (a) 60 (a) 98 (a) 150 Outlet Noz. B7217-1 A508,CL.2
- 0.010 0.73 60 (a) 60 (a) 100 (a) 154 Outlet Noz. B7217-2 A508,CL.2
- 0.010 0.72 6 (a) 6 (a) 108 (a) 167 Outlet Noz. B7217-3 A508,CL.2
- 0.010 0.72 48 (a) 48 (a) 103 (a) 158 Upper Shell B7216-1 A508,CL.2
- 0.010 0.73 30 30 (a) 97 (a) 149 Inter Shell B7203-1 A533,B,CL.1 0.14 0.010 0.60 -40 15 99 140 Inter Shell B7212-1 A533,B,CL.1 0.20 0.018 0.60 10 99 134 Lower Shell B7210-1 A533,B,CL.1 0.13 0.010 0.56 -40 18 103 128 Lower Shell B7210-2 A533,B,CL.1 0.14 0.015 0.57 -30 10 (d) 99 145 Trans. Ring B7208-1 A508,CL.2
- 0.010 0.73 40 40 (a) 89 (a) 137 Bot. HD. Dome B7214-1 A533,B,CL.1 0.11 0.007 0.48 2 (a) 87 (a) 134 Inter. Shell A1.46 SMAW 0.027 0.009 0.947 0(a) -56 (d) >131 - Long Seam (19-923A)
Inter Shell A1.40 SMAW 0.027 0.010 0.913 60
>106 - Long Seam (19-923A&B)
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-25 (SHEET 2 OF 2)
Average Upper Shelf Energy Normal to Principal Principal Working Cu P Ni T NDT RT NDT Working Direction Direction Component Code No. Grade (%) (%) (%) (°F) (°F) (ft-lb) (ft-lb) Inter Shell to Lower Shell(11-923) G1.50 SAW 0.153 0.016 0.077 40
>102 - Lower Shell Long Seams(20-923A&B) G1.39 SAW 0.05 0.006 0.096 70
>126 -
(a) Estimate per NUREG 0800 "USNRC Standard Review Plan" Branch Technical Position MTEB 5-2. (b) Estimated. (c) Upper Shelf not available, value represents minimum energy at the highest test temperature.
(d) Estimate per 10 CFR 50.61.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-26 FAULTED CONDITION LOADS FOR THE REACTOR COOLANT PUMP FOOT F F F M M M (kips) (kips) (kips) (kips) (kips) (kips)
Umbrella Loads +/-2605 +/-3305 +/-3340 +/-7050 +/-7050 +/-4010
Faulted 1 (a) 834 162 1334 2001 6023 337 Faulted 2 (a) 601 711 752 3682 2657 560 Faulted 3 (a) 876 170 1804 2859 7021 442 Ratio between umbrella loads and actual loads for the faulted condition
Case 1 3.12 20.40 2.50 3.52 1.17 11.90 Case 2 4.33 4.65 4.44 1.91 2.65 7.16 Case 3 2.97 19.44 1.85 2.47 1.00 8.97
- a. These faulted loads on the pump support feet are derived from both the pump tie rod and the
support column loads. At a particular foot, the maximums from the tie rods and the columns are
combined absolutely, although the time history LOCA analysis demonstrates clearly that the
maxima from the columns and tie rods do not occur at the same time-point. A time history
combination of the column and tie rod loads on a particular foot would significantly reduce these
loads.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-27 REACTOR COOLANT PUMP OUTLET NOZZLE FAULTED CONDITION LOADS F x F y F z M x M y M z Umbrella 3005 915 930 28,070 72,770 97,850 Case 1 575 213 239 9,667 17,519 10,001 Case 2 428 116 274 13,532 24,535 11,672 Case 3 467 148 113 4,648 8,735 11,585 Case 4 926 184 154 3,568 12,592 13,273 Ratio Between Umbrella And Actual Loads For The Faulted Condition Case 1 6.97 4.30 3.89 3.01 4.15 9.77 Case 2 9.36 7.89 3.39 2.15 2.97 8.38 Case 3 8.58 6.18 8.23 6.25 8.33 8.45 Case 4 4.33 4.97 6.04 8.15 5.78 7.37
Coordinate System z x y FNP-FSAR-5
REV 21 5/08 TABLE 5.2-28 STEAM GENERATOR LOWER SUPPORT MEMBER STRESSES Member Stresses, Percent of Allowable, for Loading Condition:
Member Normal Upset Faulted 7 to 12 Bumpers
-- 39 23 13, 14, 15 Beam
-- 31 23 20 to 23 Columns 34 44 92
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-29 STEAM GENERATOR UPPER SUPPORT MEMBER STRESSES Member Stresses, Percent of Allowable, for Loading Condition:
Member Normal Upset Faulted 25 to 29 Snubbers -- -- --
34 to 69 Bumpers & Girder
-- 18 18
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-30 REACTOR COOLANT PUMP SUPPORT MEMBER STRESSES Member Stresses, Percent of Allowable, for Loading Condition: Member Normal Upset Faulted 4 to 6 Tie Rod -- 26 36 7 to 9 Columns 30 31 42
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-31 PRESSURIZER UPPER SUPPORT MEMBER STRESSES Member Stresses, Percent of Allowable, for Loading Condition: Member Normal Upset Faulted 12 11 10 9 Upper Struts
--
--
--
-- 13.
- 10.
- 11.
- 16. 25.
- 30.
- 36.
- 29.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-32 CRDM HEAD ADAPTOR BENDING MOMENTS Combination of LOCA (a) SSE and LOCA
% of (in-kip) (in-kip)
Allowable Longest CRDM 48.0 68.2 28. Shortest CRDM 30.5 50.0 20.
- a. Maximum moments are from reactor vessel inlet break.
FNP-FSAR-5
REV 22 8/09
[HISTORICAL][TABLE 5.2-33 (SHEET 1 OF 8)
FARLEY NUCLEAR PLANT UNIT 2 PRESERVICE INSPECTION PROGRAM ASME CODE CLASS 1 COMPONENTS Table Table IWB-2500 Section XI IWB-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested B1.1 B-A Reactor Vessel Upper-to-middle-shell course Volumetric No circumferential weld B1.1 B-A Middle-to-lower-shell course Volumetric No circumferential weld B1.1 B-A Middle shell course longitudinal Volumetric No welds (2)
B1.1 B-A Lower shell course longitudinal Volumetric No welds (2)
B1.2 B-B Lower head-to-shell Volumetric No circumferential weld B1.2 B-B Lower head ring-to-disc Volumetric No circumferential weld B1.3 B-C Flange-to-vessel weld Volumetric No] B1.4 B-D Outlet nozzle-to-shell welds(3) and Volumetric No Nozzle inside-radiused sections (3)
B1.4 B-D Inlet nozzle-to-shell welds (3)
Volumetric No and nozzle inside-radiused sections (3)
B1.5 B-B CRDM, Vent and incore Visual No instrumentation penetrations and CRDM seal welds B1.6 B-F Primary nozzle-to-safe-end welds Volumetric & surface No B1.7 B-G-1 Closure studs (in place)
Not applicable No-note b
FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 2 OF 8)
Table Table IWB-2500 Section XI IWB-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested B1.8 B-G-1 Reactor Vessel (Cont'd)
Closure studs and nuts Volumetric & Surface No B1.9 B-G-1 Vessel flange ligaments Volumetric No B1.10 B-G-1 Closure washers Visual No B1.12 B-H Integrally-welded supports Not applicable No - note c B1.13 B-I-1 Closure head cladding Visual & No B1.14 B-I-1 Vessel cladding Visual No B1.15 B-N-1 Vessel interior surfaces and Visual No internals B1.16 B-N-2 Interior attachments and core Not applicable No - note d support structures B1.17 B-N-3 Core support structures Visual No B1.18 B-0 Control rod drive housings Volumetric No B1.19 B-P Exempted components Visual No B2.1 B-B Pressurizer Circumferential shell welds (5)
Volumetric Yes - note a note m B2.1 B-B Longitudinal shell welds (3)
Volumetric Yes - note a note m B-2.2 B-D Nozzle-to-vessel welds (6)
Volumetric Yes - note e and nozzle-to-vessel radiused note a sections (6) note m B2.3 B-E Heater penetrations Visual No B2.4 B-F Nozzle-to-safe-end welds (6) Surface &
No volumetric
FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 3 OF 8)
Table Table IWB-2500 Section XI IWB-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested B2.5 B-G-1 Pressurizer (Cont'd) Pressure-retaining bolting (in Not applicable No - note g place) B2.6 B-G-1 Pressure-retaining bolting Not applicable No - note g B2.7 B-G-1 Pressure-retaining bolting Not applicable No - note g B2.8 B-H Integrally welded support Volumetric No B2.9 B-I-2 Vessel cladding Visual No B2.10 B-P Exempted components Visual No B2.11 B-G-2 Manway Bolting Visual No B3.1 B-B Steam Generators (3) Channel head-to-tubesheet weld Volumetric No (primary side) (3) B3.2 B-D Nozzle-to-vessel welds and Not applicable No - note h nozzle inside-radiused sections B3.3 B-F Nozzle-to-safe-end welds (6) Volumetric & Yes - note f surface B3.4 B-G-1 Pressure-retaining bolting (in place)
Not applicable No - note g B3.5 B-G-1 Pressure-retaining bolting Not applicable No - note g B3.6 B-G-1 Pressure-retaining bolting Not applicable No - note g B3.7 B-H Integrally welded supports Not applicable No - note g B3.8 B-I-2 Vessel cladding Visual No B3.9 B-P Exempted components Visual No FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 4 OF 8)
Table Table IWB-2500 Section XI IWB-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested B3.10 B-G-2 Steam Generators (Cont'd)
Manway bolting Visual No B4.1 B-F Piping Pressure Boundary Safe-end-to-pipe welds Volumetric & Yes -note i surface B4.2 B-G-1 Pressure-retaining bolting (in place)
Not applicable No - note g B4.3 B-G-1 Pressure-retaining bolting Not applicable No - note g B4.4 B-G-1 Pressure-retaining bolting Not applicable No - note g B4.5 B-J Circumferential and Volumetric Yes - notes longitudinal pipe welds i & j B4.6 B-J Branch pipe connection welds Volumetric Yes- note l exceeding 6-inch diameter B4.7 B-J Branch pipe connection welds Surface No 6-inch diameter and smaller B4.8 B-J Socket welds Surface No B4.9 B-K-1 Integrally-welded supports Volumetric Yes- note k B4.10 B-K-2 Support components Visual No B4.11 B-P Exempted components Visual No B4.12 B-G-2 Pressure-retaining bolting Visual No B5.1 B-G-1 Reactor Coolant Pump Pressure-retaining bolts (in place)
Volumetric No B5.2 B-G-1 Pressure-retaining bolting Volumetric &
No surface B5.3 B-G-1 Pressure-retaining bolting Visual No FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 5 OF 8)
Table Table IWB-2500 Section XI IWB-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested B5.4 B-K-1 Reactor Coolant Pump (Cont'd) Integrally-welded supports Not applicable No - note g B5.5 B-K-2 Support components Visual No B5.6 B-L-1 Pump casing welds Not applicable No - note g B5.7 B-L-2 Pump casing Visual No B5.8 B-P Exempted components Visual No B5.9 B-G-2 Pressure-retaining bolting Not applicable No - note g B6.1 B-G-1 Valve Pressure Boundary Pressure-retaining bolting Not applicable No - note g (in place)
B6.2 B-G-1 Pressure-retaining bolting Not applicable No - note g B6.3 B-G-1 Pressure-retaining bolting Not applicable No - note B6.4 B-K-1 Integrally-welded supports Not applicable No - note g B6.5 B-K-2 Support Components Visual No B6.6 B-M-1 Valve-body welds Not applicable No-note g B6.7 B-M-2 Valve bodies Visual No B6.8 B-P Exempted components Visual No B6.9 B-G-2 Pressure-retaining bolting Visual No FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 6 OF 8)
Notes a. For the pressurizer, the requirements of I-3121 of S ection XI are impossible to meet. At the time the components were built, no excess material was saved fo r fabrication of calibration blocks. As an alternative, calibration blocks required for the ultrasonic examination of welds in these vessels will be fabricated from material of the same specifica tion, product form, and heat treatment as one of the materials being joined as allowed by Article T-434.1.1 in Section V of the ASME Boiler and Pressure Vessel Code.
- b. The reactor vessel closure studs are rem oved during the preservice inspection.
- c. The reactor vessel supports are integral with the primary nozzles and the examination requirements of IWB-2600 is covered by item B1.4.
- d. The requirements of IWB-2600 are applicable only to boiling water-type reactors and are thus not applicable to Farley Nuclear Plant.
- e. The geometric configuration of the weld surface prevents ultrasonic examinations being performed to the extent required by IWB-2600. Angle beam examinations will be performed from the vessel head and on top of the weld. All of the weld, the heat affected zone, and the required amount of base metal on the shell side of the weld will be examined.
Base metal on the nozzle side of the weld will be examined to the extent practical, which is approximately 25 percent. In addition, the welds will receive surface examination on those areas not scanned by UT.
- f. Examination of the steam generator primary no zzle safe-end-to-pipe welds is limited by the nozzle geometry and surface condition, and by the limited surface preparation on the pipe side of the weld.
The surface on the pipe side of the weld, which is a cast elbow, is machined for a distance of approximately 5-1/4 inches from the edge of the weld. Ultrasonic examination is limited to this distance from the edge of the weld. Examinations can be performed on the surface of the weld but are severely limited from the nozzle side by the con figuration of weld build up and weld overlay.
Ultrasonic examinations will be performed from bot h the pipe and weld surfaces as allowed by T-532 of Section V. All of the weld metal, including th e weld root, will be inspected. Since no UT can be performed on the nozzle side of the weld, the ext ent of examination is limited to approximately 90 percent of the code-required area. Surface examinations will be performed on essentially 100 percent of the required area.
- g. There are no items in this category that require examination under the requirements of IWB-2600.
- h. The steam generator nozzles are integrally forged with the channe l head and thus do not contain any welds.
FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 7 OF 8)
- i. The arrangements and details of the piping systems and components are such that some examinations as required by IWB-2600 are limited because of geom etric configuration or accessibility. The welds will be ultrasonically examined by angle beams to the extent allowed by geometric configuration. In all cases, 100 percent of the weld material will be examined. Also, surface examinations will be performed to supplement limited volumetric examinations.
Welds requiring supplemental surface examination, along with the estimated extent of volumetric examination, are as follows:
Loop 1 RTD return, weld #16 - 40% Loop 1 Cold Leg SIS, weld #8 - 60% Pressurizer Spray, Welds #42 - 70% and #43 - 70% Loop 3 RTD Return, weld #8 - 60% Pressurizer Relief, weld #14 - 50% Pressurizer Safety, welds #2 - 70%
- 5 - 80%
- 12 - 70%
- 16 - 80%
- 20 - 80%
- 24 - 70%
- 27 - 80%
Pressurizer safety welds 29, 31, 32, and loop 3, 2-in. safety injection (hot leg) weld 9 are inaccessible.
However, field data in the form of radiography and dye penetrant will be utilized for preservice inspection as allowed by IWC-2100(b).
- j. In instances where the locations of pipe suppor ts or hangers restrict the access available for the examination of pipe welds as required by IWB-2600, examinations will be performed to the extent practical unless removal of the support is permissible without unduly stressing the system.
- k. The piping system integrally welded supports are attached to the pipe by fillet welds. The configurations of such welds are such that exami nations cannot be performed to the extent required by IWB-2600 and only the base material of the pipe wall can be examined by ultrasonic techniques.
Surface examination will be performed on the integr ally welded attachments to supplement the limited volumetric examination.
- l. The geometric configuration of the weld surface prevents ultrasonic examinations from being performed to the extent required by IWB-2600. Examinations will be performed to the extent practical from the pipe and nozzle surfaces adjacen t to the weld. Surface examination of the weld will be performed to supplement the volumetric examinations.
Welds requiring supplemental surface examination along with the estimated extent of volumetric examination, are as follows:
FNP-FSAR-5
REV 22 8/09 TABLE 5.2-33 (SHEET 8 OF 8)
Reactor Coolant Loop #1, weld #16BC - 80% Reactor Coolant Loop #1, weld #21BC - 80% Reactor Coolant Loop #2, weld #16BC - 80% Reactor Coolant Loop #2, weld #21BC - 80% Reactor Coolant Loop #3, weld #16BC - 80% Reactor Coolant Loop #3, weld #21BC - 80%
- m. For the pressurizer, the requirements of I-3122 of Section XI cannot be met because of lack of cladding on the calibration blocks. However, only the top (O.D.) portions of the blocks are used for calibration. Specifically, the blocks contain side-dr illed holes at depths of 1/4 T, 1/2 T, and 3/4 T. The blocks also contain a 2% T I.D. notch, but it is used only as a reference. Since the lack of cladding does not affect the ultrasonic calibration, the existing unclad calibration blocks will be utilized].
FNP-FSAR-5 REV 22 8/09
[HISTORICAL][TABLE 5.2-34 (SHEET 1 OF 8)
FARLEY NUCLEAR PLANT UNIT 2 PRESERVICE INSPECTION PROGRAM ASME CODE CLASS 2 COMPONENTS Table Table IWC-2520 Section XI IWC-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested C1.1 C-A Letdown Heat Exchanger Head-to-shell weld Volumetric No (tube side)
C1.1 C-A Shell-to-flange weld Volumetric No C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a C1.3 C-C Integrally-welded supports Not applicable No - note b C1.4 C-D Tubesheet flange bolting Not applicable No - note b C1.1 C-A Excess Letdown Heat Head-to-flange weld Volumetric Yes - note k Exchanger (tube side)
C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a C1.3 C-C Integrally-welded supports Not applicable No - note b C1.4 C-D Tubesheet flange bolting Visual & No volumetric C1.1 C-A Regenerative Heat Exchanger Head-to-shell welds (6) Volumetric Yes - note g note k C1.1 C-A Shell-to-tubesheet welds (6) Volumetric Yes - note g note k C1.2 C-B Nozzle-to-vessel welds (12)
Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C1.1 C-A Residual Heat Exchangers (2)
Head-to-shell welds Volumetric No (tube side)
FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 2 OF 8)
Table Table IWC-2520 Section XI IWC-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested C1.1 C-A Shell-to-tubesheet welds Volumetric No C1.2 C-B Nozzle-to-vessel welds Not accessible Yes - note c C1.3 C-C Integrally-welded supports Not applicable No - note b C1.4 C-D Tubesheet flange bolting Visual and No volumetric C1.1 C-A Seal-Water Return Filter Cover weldment-to-shell weld Visual and Yes - note d surface C1.1 C-A Head-to-shell weld Visual and Yes - note d surface C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C1.1 C-A Volume Control Tank Upper head-to-shell weld Volumetric No C1.1 C-A Lower head-to-shell weld Volumetric No C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Manway bolting Visual and No volumetric C1.1 C-A Letdown Reheat Heat Head-to-shell weld Visual and Yes - note d Exchanger (tube side) surface note k C1.1 C-A Shell-to-flange weld Visual and Yes - note d surface note k C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a
FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 3 OF 8)
Table Table IWC-2520 Section XI IWC-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C1.1 C-A Seal-Water-Heat Exchanger Head-to-shell weld Visual and Yes - note d (tube side) surface C1.1 C-A Shell-to-flange weld Visual and Yes - note d surface C1.2 C-B Nozzle-to-vessel welds Not applicable No - note a C1.3 C-C Integrally-welded supports Not applicable No - note b C1.4 C-D Tubesheet flange bolting Not applicable No - note b C1.1 C-A Steam Generators (3)
Upper head-to-shell weld Volumetric No (shell side)
C1.1 C-A Barrel-to-tubesheet weld Volumetric No C1.2 C-B Feedwater inlet nozzle-to-shell Volumetric No weld C1.3 C-C Integrally-welded supports Not applicable No - note b C1.4 C-D Pressure retaining bolting > 2 In.
Not No - note b Applicable C1.1 C-A Reactor Coolant Filter Cover weldment-to-shell weld Visual and Yes - note d surface FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 4 OF 8)
Table Table IWC-2520 Section XI IWC-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested C1.1 C-A Head-to-shell weld Visual and Yes - note d surface C1.2 C-B Nozzle-to-vessel weld Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C1.1 C-A Letdown Delay Tanks (2)
Head-to-shell welds Volumetric No C1.2 C-B Nozzle-to-vessel welds Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C1.1 C-A Excess Letdown Delay Head-to-shell welds Volumetric No Tanks (2)
C1.2 C-B Nozzle-to-vessel welds Not applicable No - note a C1.3 C-C Integrally-welded supports Surface No C1.4 C-D Pressure-retaining bolting Not applicable No - note b C2.1 C-F; C-G Piping Systems - note i Circumferential butt welds Volumetric Yes - notes e & f C2.2 C-F; C-G Longitudinal weld joints in Volumetric No fittings C2.3 C-F; C-G Branch pipe-to-pipe welds Volumetric Yes - note e C2.4 C-D Pressure-retaining bolting Visual and No volumetric C2.5 C-E-1 Integrally-welded supports Surface No C2.6 C-E-2 Support components Visual No FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 5 OF 8)
Table Table IWC-2520 Section XI IWC-2600 Examination System or Method of Code Relief Item No. Category Component Area To Be Examined Examination Requested C3.1 C-F Residual Heat Removal Pumps Pump casing welds Not applicable No - note b (2) C3.2 C-D Pressure-retaining bolting Visual and No volumetric C3.3 C-E-1 Integrally-welded supports Not applicable No - note b C3.4 C-E-2 Support components Visual No C3.1 C-F Centrifugal Charging Pumps Pump casing welds Volumetric No (3) C3.2 C-D Pressure-retaining bolting Visual and No volumetric C3.3 C-E-1 Integrally-welded supports Surface Yes - note j C3.4 C-E-2 Support components Visual No C4.1 C-F; C-G Valves Valve-body welds Not applicable No - note b C4.2 C-D Pressure-retaining bolting Visual and No Volumetric C4.3 C-E-1 Integrally-welded supports Not applicable No - note b C4.4 C-E-2 Support components Visual No FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 6 OF 8)
Notes a. This item is excluded from the examination requirements of IWC-2600 by application of the criteria given in IWC-1220.
- b. There are no items in this category that require examination under the requirements of IWC-2600.
- c. The nozzle to vessel welds of the residual heat exchangers are covered by a reinforcement ring and are not accessible for examination as required by IWC-2600. The geometric configuration is such that alternative NDE methods cannot be substituted. The nozzles will
be subject to visual inspection for leakage.
- d. The thickness of the materials utilized for the construction of this component (0.165 to 0.185 in.) is such that meaningful results could not be expected with ultrasonic examination as required by IWC-2600. Surface and visual examination of these welds will be performed as an alternative method.
- e. The arrangement and details of the Class 2 piping system and components were designed and fabricated before the examination requirements of Section XI of the Code were formalized and some examinations as required by IWC-2600 are limited or not practical because of geometric configuration or accessibility. Generally these limitations exist at all
fitting to fitting welds such as elbow to tee, elbow to valve, reducer to valve, etc. where geometry and sometimes surface conditions preclude ultrasonic coupling or access for the required scan length. The limitations exist to a lesser degree at pipe to fitting welds, where examination can only be fully performed from the pipe side, the fitting geometry limiting or even precluding examination from the opposite side. The welds will be ultrasonically examined by angle beam to the extent allowed by geometric configuration; however, 100 percent of the weld material will be examined. Also, surface examinations will be performed to supplement the limited volumetric examinations. Welds requiring supplemental surface examination, along with the estimated extent of examination, are as follows:
RHR, welds #31 - 50%
- 32 - 50%
- 14 - 90%
- 11 - 30%
- 20 - 30%
- 18 - 50%
FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 7 OF 8)
In instances of branch pipe to pipe welds, ultrasonic examinations cannot be performed on the surface of the weld. Surface examination will be performed on 100 percent of the weld and adjacent base material. Welds requiring supplemental surface examination, along with the estimated extent of volumetric examination, are as follows:
Main Steam, welds #4 80% #2 80%
- 4 80% #2 80%
- 1-5 - 80% #2 80%
- 1 80% #2 80%
- 1 80% #2 80%
- 1 80% #3-5 - 80%
- 1 80% #3 80%
- 1 80% #3 80%
- 1 80% #3 80%
- 1 80% #3 80%
- 2-5 - 80% #3 80%
- 2 80% #3 80%
- 2 80% #3 80%
- f. In instances where the locations of pipe supports or hangers restrict the access available for the examination of pipe welds as required by IWC-2600, examinations will be performed to the extent practical unless removal of the support is permissible without unduly stressing the system.
- g. The regenerative heat exchanger shell is fabricated from centrifugally cast austenitic steel material which limits ultrasonic examination as required by IWC-2600 to the half node technique. The geometric configuration of the weld surface and the location of adjacent nozzles and supports provide limitations to the extent of examination coverage. Surface examinations will be performed to supplement the volumetric examination.
- h. The following components are exempt from the examination requirements of IWC-2520 by application of the criteria given in IWC-1220. These components will be examined in accordance with the requirements of IWC-2510.
- 1. CVCS seal water injection filters (2)
- 2. Safety injection accumulators (3)
- 3. Boron injection tank
- 4. Containment spray pumps (2)
- 5. Refueling water storage tank (RWST) and
- a. Suction piping from the RWST to the High Head Safety Injection Pumps.
FNP-FSAR-5 REV 22 8/09 TABLE 5.2-34 (SHEET 8 OF 8)
- b. Suction piping from the RWST to the Low Head Safety Injection/Residual Heat Removal Pumps.
- c. Suction piping from the RWST to the Containment Spray Pumps.
- i. All Class 2 piping with a nominal diameter of 4 in. or less is excluded from the examination requirements of IWC-2520 by the application of the criteria given in IWC-1220.
- j. Because of component and support designs, approximately 20 percent of each integrally-welded support is inaccessible for examination. The accessible portion of each support will receive visual and surface examinations.
- k. Table IWC-2520, Category C-A and IWC-2600, Item C1.1 require volumetric examinations "uniformly distributed among three areas around the vessel circumference." The location of adjacent nozzles provides limitations to the extent of examination coverage. Consequently, the requirement for three uniformly distributed areas cannot be met. One or two areas will be inspected, as accessibility permits, instead of the required three areas. The required 20 percent of each circumferential weld will be volumetrically inspected except where material thickness precludes ultrasonic testing as stated in note 4.]
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-35 (SHEET 1 OF 2)
TYPE B-4 WELD WIRE AND LINDE 0091 FLUX TESTS Example 1 - Type B-4 Weld Wire and Linde 0091 Flux (Test #1302)
Impact and or Fracture Tests Type Temp. °F Values Temp. °F Values NDT CVN Ft/Lbs %Shear Mils Lat Exp Drop Weights -80 3 0 1 -50 1 F -40°F -80 3 0 2 -40 1 F -80 9 0 4 -30 2 NF -40 26 10 19 -40 37 15 25 -40 38 15 24 +10 69 35 46 +100 117 90 83 +10 50 25 38 +100 114 90 82 +10 66 30 44 +100 120 90 83 +20 66 35 46 +160 124 100 83 +20 81 50 57 +160 136 100 89 +20 90 60 63 +160 135 100 88 Example 2 - Type B-4 Weld Wire and Linde 0091 Flux (Test #1388)
Impact and or Fracture Tests Type Temp. °F Values Temp. °F Values NDT CVN Ft/Lbs
%Shear Mils Lat Exp Drop Weights -80 11 0 3 -60 1 F -60°F -80 11 0 3 -50 2 NF -80 13 0 4 -40 1 NF -40 30 15 17 -40 27 15 15 -40 25 10 11 0 77 50 45 +100 143 100 84 0 72 50 40 +100 133 100 82 0 70 50 41 +100 145 100 86 +10 76 50 41 +180 143 100 82
+10 74 50 46 +180 149 100 86
+10 82 60 45 +180 148 100 85
+60 116 70 76 +60 118 70 74 +60 121 70 71 FNP-FSAR-5
REV 21 5/08 TABLE 5.2-35 (SHEET 2 OF 2)
Example 3 - Type B-4 Weld Wire and Linde 0091 Flux (Test #1389)
Impact and or Fracture Tests Type Temp. °F Values Temp. °F Values NDT CVN Ft/Lbs
%Shear Mils Lat Exp Drop Weights -60 16 0 9 -60 1 F -60°F -60 15 0 7 -50 2 NF -60 19 0 11 -40 1 NF -40 20 5 11 -40 28 10 16 -40 32 15 22 -20 85 50 53 +60 132 80 77
-20 88 50 56 +60 149 100 84
-20 76 40 47 +60 123 80 74 0 77 40 47 +100 142 100 82 0 75 40 45 +100 148 100 84 0 99 60 52 +100 140 100 82
+20 117 70 74 +20 105 60 65 +20 114 70 74 Example 4 - Type B-4 Weld Wire and Linde 0091 Flux (Test #1386)
Impact and or Fracture Tests Type Temp. °F Values Temp. °F Values NDT CVN Ft/Lbs
%Shear Mils Lat Exp Drop Weights -80 16 0 7 -60 1 F -60°F -80 18 0 8 -50 2 NF -80 18 0 7 -40 1 NF -40 38 20 26 -40 32 15 17 -40 34 15 19 0 79 40 52 +100 137 100 82 0 61 70 39 +100 132 100 82 0 95 70 60 +100 141 100 83
+10 96 70 62 +180 142 100 82
+10 101 70 60 +180 145 100 85
+10 84 60 58 +180 143 100 83
+60 118 80 78 +60 130 90 80 +60 117 80 75 FNP-FSAR-5
REV 21 5/08 TABLE 5.2-36 FARLEY NUCLEAR PLANT UNIT 2 LOWER SHELL COURSE CHARPY V NOTCH DATA (a) Plate Code No. B7210-1 Plate Code No. B7210-2 Test Energy Lat. Exp Shear Energy Lat. Exp Shear Temp. (°F) (Ft-Lb) (Mils) (%) Temp. (°F) (Ft-Lb) (Mils) (%) -50 10 6 9 -50 15 11 9 -50 14.5 8 15 -50 12.5 8 9 -50 11 7 9 -50 11 8 9 20 33 25 29 0 26 24 30 20 47 35 34 0 27.5 27 34 20 46 33 34 0 45 35 32 75 48.5 38 59 30 51 39 30 75 50 40 59 30 40 34 34 75 62 47 64 30 47 39 30 110 86 67 80 100 67 52 79 110 75 57 75 100 80.5 58 75 110 69.5 54 67 100 85 60 75 150 100 69 100 150 100 76 100 150 95 71 100 150 101 74 100 150 93 67 100 150 97 75 100 210 96 70 100 210 98 69 100 210 105.5 74 100 210 102 76 100 210 107 75 100 210 95.5 72 100
- a. Normal to major rolling direction of the plate.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-37 FARLEY NUCLEAR PLANT UNIT 2 INTERMEDIATE SHELL COURSE CHARPY V NOTCH DATA (a) Plate Code No. B7210-1 Plate Code No. B7210-2 Test Energy Lat. Exp Shear Test Energy Lat. Exp Shear Temp. (°F) (Ft-Lb) (Mils) (%) Temp. (°F) (Ft-Lb) (Mils) (%) -50 13.5 9 15 -50 18.5 11 12 -50 19 11 15 -50 15.5 11 12 -50 14 8 15 -50 19 11 12 0 28 25 30 0 35 27 27 0 34 26 28 0 34.5 27 25 0 44 36 34 0 30 27 25 20 55 41 40 30 43 35 32 20 51 38 45 30 48 36 35 20 43 32 34 30 52 39 43 75 50.5 50 56 100 76.5 55 73 75 61.5 40 52 100 74 56 73 75 65 46 61 100 70 54 69 150 91 68 100 150 95 67 100 150 97 76 100 150 98 68 100 150 92 70 100 150 106 76 100 210 105.5 69 100 210 89 68 100 210 97.5 74 100 210 94 70 100 210 95.5 72 100 210 88 69 100
- a. Normal to major rolling direction of the plate.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-38 (SHEET 1 OF 2)
FARLEY NUCLEAR PLANT UNIT 2 NOZZLE SHELL COURSE CHARPY V NOTCH DATA (a) FORGING CODE No. B7261-1 Test Energy Lat. Exp Shear Temp. (°F) (Ft-Lb) (Mils) (%) -80 2 0 0 -80 4 0 0 -80 8 4 0 -20 68 53 29 -20 37 25 16 -20 66 52 29 10 99 76 64 10 103 76 64 10 110 81 70 10 95 77 52 10 55 41 29 10 78 63 40 30 72 57 23 30 93 70 55 30 87 65 46 100 147 91 100 100 123 77 75
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-38 (SHEET 2 OF 2)
Test Energy Lat. Exp Shear Temp. (°F) (Ft-Lb) (Mils) (%) 100 110 80 70 180 146 90 100 180 151 88 100 180 149 88 100
- a. Major working direction of forging.
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-39 STEAM GENERATOR PRESSURIZER FRACTURE TOUGHNESS PROPERTIES Charpy Lateral Test Component Test Material V Notch Expansion Temperature T ndt RT ndt Component Part Number Specification (ft-lb)
(in)
(°F) (°F) (°F) Steam generator(b) Channel head T03848 SA 216 Gr. WCC 61.4, 59.0, 78.7 .070, .066, .089 70 10 10 (1551) 86.2, 84.2, 86.6 .127, .124, .135 70 Tubesheet T03254 SA 508 C1.2 88, 88, 74 .057, .057, .051 10 (a) 84, 58, 81 .057, .041, .055 10 Manway cover T03729 SA 533 Gr. A C1.1 53, 51, 52 .051, .047, .049 100 10 40 52, 53, 54 .049, .052, .052 100 Steam generator(b) Channel head T03843 SA 216 Gr. WCC 56, 68, 57 .046, .057, .043 70 0 10 (1552) 80, 80, 80 .061, .062, .068 70 Tubesheet T03302 SA 508 C1.2 80.5, 74.0, 85.0 .054, .047, .063 10 (a) 86.0, 108.0, 83.0 .059, .067, .052 10 Manway cover T03729 See data above for steam generator (1551)
Steam generator(b) Channel head T03909 SA 216 Gr. WCC 88.9, 70.1, 73.7 .075, .064, .062 70 (a) (1553) 62.9, 67.9, 67.9 .058, .063, .062 70 Tubesheet T03307 SA 508 C1.2 67, 55, 65 .045, .036, .046 10 (a) 75, 51, 54 .051, .035, .037 10 Manway cover T03729 See data above for steam generator (1551)
Pressurizer Lower head T03626 SA 533 Gr. A C1.2 75, 83, 76 .060, .064, .062 70 10 10 (1561) Surge nozzle forging T03386 SA 508 C1.2 88, 100, 96 .067, .082, .078 70 10 10 Upper head T03748 SA 533 Gr. A C1.2 63, 75, 72 .060, .056, .062 70 10 10 Manway nozzle forging T03336 SA 508 C1.2 113, 129, 120 .085, .086, .077 70 10 10 Safety nozzle forging T03381-3 SA 508 C1.2 82, 82, 78 .065, .066, .063 70 10 10 Safety nozzle forging T03284-1 SA 508 C1.2 64, 64, 55 .042, .042, .036 10 (a) Safety nozzle forging T04281-10 SA 508 C1.2a 139, 136, 141 .089, .079, .087 120 60 60 Relief nozzle forging T03380-3 SA 508 C1.2 84, 83, 92 .067, .070, .076 70 10 10 Spray nozzle forging T03722 SA 508 C1.2 74, 86, 71 .064, .072, .059 70 10 10 Manway cover T04405 SA 533 Gr. A C1.1 75, 79, 81 .069, .068, .064 120 60 60 Shell barrel T03630 SA 533 Gr. A C1.2 58, 58, 62 .054, .054, .054 70 10 10 Shell barrel T03741 SA 533 Gr. A C1.2 51, 54, 54 .049, .044, .050 80 10 20 Shell barrel T03355 SA 533 Gr. A C1.2 60, 64, 66 .046, .058, .058 70 10 10
- a. Drop weight test results not available. b. These values are for Model 51 Steam Generators. The values will be updated for the replacement steam generator (Model 54) after the issuance of the Replacement Steam Generator Design Report FNP-FSAR-5
REV 21 5/08 TABLE 5.2-40 LOAD COMBINATIONS AND ACCEPTANCE CRITERIA FOR PRESSURIZER AND RELIEF VALVE PIPING - UPSTREAM OF VALVES CLASS I PIPING Plant/System Operating Condition
Load Combination Piping Allowable Stress Intensity Normal N 1.5 S m Upset N + OBE 1.5 S m Upset N + SOT U 1.5 S m Upset N + OBE + SOT U 1.8 S m/1.5 S y (2) Emergency N + SSE + SOT E 2.25 S m/1.8 S y (2) Faulted N + SSE + SOT F 3.0 S m NOTES: 1. Use SRSS for combining dynamic load responses.
- 2. The smaller of the given allowable is to be used.
N = Sustained loads during normal plant operation SOT = System operating transient
SOT U = Relief valve discharge transient SOT E = Safety valve discharge transient SOT F = Max (SOT U; SOT E); or transition flow OBE = Operating basis earthquake SSE = Safe shutdown earthquake
S h = Basic material allowable stress at maximum (hot) temperature S m = Allowable design stress intensity S y = Yield strength value
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-41 LOAD COMBINATIONS AND ACCEPTANCE CRITERIA FOR PRESSURIZER AND RELIEF VALVE PIPING - DOWNSTREAM OF VALVES NNS PIPING Plant/System Operating Condition Load Combination Piping Allowable Stress Normal N 1.0 S h Upset N + OBE 1.2 S h Upset N + SOT U 1.2 S h Upset N + OBE + SOT U 1.8 S h Emergency N + SOT E 2.4 S h* Faulted N + SSE + SOT F 2.4 S h NOTE: Use SRSS for combining dynamic load responses.
- See reference (21)
N = Sustained loads during normal plant operation SOT = System operating transient
SOT U = Relief valve discharge transient SOT E = Safety valve discharge transient SOT F = Max (SOT U; SOT E); or transition flow OBE = Operating basis earthquake SSE = Safe shutdown earthquake
S h = Basic material allowable stress at maximum (hot) temperature S m = Allowable design stress intensity S y = Yield strength value
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-42 FARLEY UNITS 1 AND 2 SAFETY LINE PIPE STRESS AND STRAIN
SUMMARY
FOR EMERGENCY CONDITION Node Point Piping Components Code Maximum Stress (ksi)
Allowable Stress (ksi) 1290* Butt weld at valve end nozzle 15.1 18.8 1460* Long radius elbow 34.2 36.45 100** Branch connection 32.9 44.67 690** reducer 25.1+ 44.67 1490** Welded attachment at support R120***
54.97*** 55.42
- ASME Class 1 piping, upstream of safety valves
- ASME NNS piping, downstream of safety valves
- Based on ASME Code Case N-318 allowable
+ Stress Index based on ANSI B31.1-1967, including 1971 Addenda
FNP-FSAR-5
REV 21 5/08 TABLE 5.2-43 FARLEY NUCLEAR PLANT - TMI ACTION NUREG-0737.II.D.1 UNITS 1 AND 2 PSARV LINE PIPE SUPPORTS ANCHOR BOLT DATA FOR SUPPORTS WITH FACTOR OF SAFETY F.S. <4 Unit Serial Support Total No. of No. of Actual F.S.
Types of Bolts No. Mark No. No. of Bolts Bolts w/ F.S. 4 Bolts w/ F.S. <4 Bolt # F.S. with F.S. <4 1 1 RC-R61 4 2 2 #3
- 4 3.57 3.57 #3 and #4 3/4" HILTI KWIK 2 1 2RC-131X 5 3 2 #2
- 5 3.77 3.20 #2 AND #5 1/2" HILTI KWIK
REV 21 5/08 PRIMARY-SECONDARY BOUNDARY COMPONENTS SHELL LOCATIONS OF STRESS INVESTIGATIONS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-1
REV 21 5/08 PRIMARY AND SECONDARY HYDROSTATIC TEST STRESS HISTORY FOR THE CENTER HOLE LOCATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-2
REV 21 5/08 PLANT HEATUP AND LOADING OPERATIONAL TRANSIENTS (WITH STEADY-STATE PLATEAU) STRESS HISTORY FOR THE HOT SIDE CENTER HOLE LOCATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-3
REV 21 5/08 LARGE STEP LOAD DECREASE AND LOSS OF FLOW STRESS HISTORY FOR THE HOT SIDE CENTER HOLE LOCATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-4
REV 21 5/08 REACTOR COOLANT LOOP/SUPPORTS SYSTEM DYNAMIC - STRUCTURAL MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-5
REV 21 5/08 STHRUST RCL MODEL SHOWING HYDRAULIC FORCE LOCATIONS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-6
REV 21 5/08 UNIT 1 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS APPLICABLE FOR FIRST TIME 16 EFPY OF OPERATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-7 (SHEET 1 OF 2)
REV 21 5/08 FARLEY UNIT 2 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS APPLICABLE FOR THE FIRST 14 EFPY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-7 (SHEET 2 OF 2)
REV 21 5/08 UNIT 1 ALA REACTOR COOLANT SYSTEM COOLDOWN LIMITATIONS APPLICABLE FOR THE FIRST 16 EFFECTIVE FULL POWER YEARS OF OPERATION JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-8 (SHEET 1 OF 2)
REV 21 5/08 FARLEY UNIT 2 REACTOR COOLING SYSTEM COOLDOWN LIMITATIONS APPLICABLE FOR THE FIRST 14 EFPY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-8 (SHEET 2 OF 2)
THIS FIGURE HAS BEEN DELETED.
REV 21 5/08 EFFECT OF FLUENCE AND COPPER CONTENT ON SHIFT OF RT NDT FOR REACTOR VESSEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-9
REV 21 5/08 UNIT 1 FAST NEUTRON FLUENCE (E > 1 MEV) AS A FUNCTION OF FULL POWER SERVICE LIFE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-10 (SHEET 1 OF 2)
REV 21 5/08 UNIT 1 FAST NEUTRON FLUENCE (E > 1 MEV) AS A FUNCTION OF FULL POWER SERVICE LIFE (45
° LOCATION)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-10 (SHEET 2 OF 2)
REV 21 5/08 K 1d LOWER BOUND FRACTURE TOUGHNESS A533V (REFERENCE WCAP 7623) GRADE B CLASS 1 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-11
THIS FIGURE HAS BEEN DELETED.
REV 21 5/08 CONDENSATE MEASURING SYSTEM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-12
REV 21 5/08 TOOL DETAILS (VESSEL SCANNER)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-13
REV 21 5/08 TOOL DETAILS (NOZZLE AND FLANGE SCANNER)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-14
REV 21 5/08 SAMPLE WELD DATA SHEET JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-15
REV 21 5/08 PRESSURIZER SAFETY LINE STRUCTURAL MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-16
REV 21 5/08 REACTOR COOLANT PUMP CASING WITH SUPPORT FEET JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-20
REV 21 5/08 BOLT HOLD RADIAL CENTERLINE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-21
REV 21 5/08 NONLINEAR CRDM CENTER ROW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.2-22
FNP-FSAR-5
5.3-1 REV 21 5/08 5.3 THERMAL HYDRAULIC SYSTEM DESIGN
5.3.1 ANALYTICAL METHODS AND DATA The thermal and hydraulic design bases of the reactor coolant system (RCS) are described in
sections 4.3 and 4.4 in terms of core heat generation rates, departure from nucleate boiling ratio (DNBR), analytical models, peaking factors, and other relevant aspects of the reactor.
5.3.2 OPERATING RESTRICTIONS ON PUMPS Plant operating experience and instrument inaccuracy are used to establish a pressure range
which ensures that all RCP support conditions are met and that the LTOP relief valves are not
challenged during RCP start, the ensuing transient, and any subsequent operation.
5.3.3 BOILING WATER REACTOR (BWR) 5.3.4 TEMPERATURE-POWER OPERATING MAP The effects of reduced core flow because of inoperative pumps is discussed in subsections
5.5.1, 15.2.5, and 15.3.4.
Natural circulation capability of the system is shown in table 5.3-1.
The issue of steam formation in the RCS was made part of TMI Action Plan Requirement
II.K.2.17. The potential for voids being generated in the RCS during anticipated transients is
accounted for in present analysis models. The transient analyses performed using these
models demonstrate that steam voids will not result in unacceptable consequences during
anticipated transients.
5.3.5 LOAD FOLLOWING CHARACTERISTICS The RCS is designed on the basis of steady-state operation at full-power heat load. The reactor
coolant pumps utilize constant speed drives as described in section 5.5, and the reactor power
is controlled to maintain average coolant temperature at a value which is a linear function of
load, as described in section 7.7.
5.3.6 TRANSIENT EFFECTS Transient effects are evaluated as follows: complete loss of forced reactor coolant flow (15.3.4);
partial loss of forced reactor coolant flow (15.2.5); startup of an inactive loop (15.2.6); loss of
load (15.2.7); loss of normal feedwater (15.2.8); loss of offsite power (15.2.9); and accidental
depressurization of the reactor coolant system (15.2.12).
FNP-FSAR-5
5.3-2 REV 21 5/08 5.3.7 THERMAL AND HYDRAULIC CHARACTERISTICS
SUMMARY
TABLE The thermal and hydraulic characteristics are given in tables 4.3-1, 4.4-1, and 4.4-2.
FNP-FSAR-5 REV 21 5/08 TABLE 5.3-1 NATURAL CIRCULATION REACTOR COOLANT FLOW VERSUS REACTOR POWER Reactor Power Reactor Coolant Flow
(% Full Power)
(% Nominal Flow) 3.5 4.8 3.0 4.6 2.5 4.4 2.0 4.1 1.5 3.7
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-1 REACTOR VESSEL DESIGN PARAMETERS Design/operating pressure (psig) 2485/2235 Design temperature (°F) 650 Overall height of vessel and closure 42 3/16 head (ft-in.) (bottom head OD to top of control rod mechanism adapter)
Thickness of insulation (min, in.)
3 Number of reactor closure head studs 58 Diameter of reactor closure head 6 studs (in.)
ID of flange (in.)
184 ID at shell (in.)
157 Inlet nozzle ID (in.)
27-1/2 Outlet nozzle ID (in.)
29 Clad thickness (min, in.)
5/32 Lower head thickness (min, in.)
5 Vessel beltline thickness (min, in.)
7-7/8 Closure head thickness (in.)
6-3/16
FNP-FSAR-5
REV 21 5/08
[HISTORICAL] [TABLE 5.4-2 REACTOR VESSEL QUALITY ASSURANCE PROGRAM RT (a) UT (a) PT (a) MT (a) Forgings 1. Flanges yes yes 2. Studs yes yes 3. Head adapters yes yes 4. Head adapter tube yes yes 5. Instrumentation tube yes yes 6. Main nozzles yes yes 7. Nozzle safe ends yes yes Plates yes yes Weldments
- 1. Main seam yes yes yes 2. CRD head adapter yes connection
- 3. Instrumentation tube yes connection
- 4. Main nozzles yes yes yes 5. Cladding yes yes 6 Nozzle safe ends yes yes yes (forging) 7 Head adapter forging to yes yes head adapter tube 8. All ferritic welds yes yes accessible after hydrotest
- 9. All nonferritic welds yes yes accessible after hydrotest
- 10. Seal ledge yes 11. Head lift lugs yes 12. Core pad welds yes yes yes
- a. RT - Radiographic
UT - Ultrasonic
PT - Dye penetrant MT - Magnetic particle
]
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-3 IDENTIFICATION OF UNIT NO. 1 REACTOR VESSEL BELTLINE REGION BASE MATERIAL Material Composition (Wt. %)
Component Code No. Heat No. Spec. No. C Mn P S Si Ni Mo Cu Cr AL
Inter. shell B6903-2 C6294 A533B, CL.1 0.20 1.32 0.011 0.013 0.21 0.60 0.55 0.13 - 0.017 Inter. shell B6903-3 C6308 A533B, CL.1 0.21 1.29 0.014 0.015 0.16 0.56 0.56 0.12 - 0.019 Lower shell B6919-1 C6940 A533B, CL.1 0.20 1.39 0.015 0.015 0.18 0.55 0.56 0.14 - 0.025 Lower shell B6919-2 C6897 A533B, CL.1 0.20 1.39 0.015 0.018 0.19 0.56 0.53 0.14 - 0.018
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-4 PREDICTED END OF LICENSE (54 EFPY) UPPER SHELF ENERGY VALUES FARLEY UNIT NO. 1 REACTOR VESSEL BELTLINE PLATES (Ref. 7) 1/4T Fluence Unirradiated Decrease in Projected EOL Beltline Material Wt. % Cu (10 19 n/cm 2) USE (ft-lb)
USE (%) USE (ft-lb) Intermediate Shell Plate B6903-2 0.13 4.0 99 30 69 Intermediate Shell Plate B6903-3 0.12 4.0 87 29 62 Lower Shell Plate B6919-1 0.14 4.0 86 31 59 Lower Shell Plate B6919-2 0.14 4.0 86 31 59
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REV 21 5/08 TABLE 5.4-5 IDENTIFICATION OF UNIT NO. 1 REACTOR VESSEL BELTLINE REGION WELD METAL Weld Wire Flux Composition (Wt. %)
Weld Weld Location Process Type Heat No. Type Lot No. C Mn P S Si Mo Cu Ni Inter. shell Sub-arc B4 33A277 Linde 1092 3889 0.11 1.27 0.015 0.010 0.14 0.49 0.258 0.165 long seams19-894 A&B
Inter. shell to Sub-arc B4 6329637 Linde 0091 3999 0.14 1.15 0.011 0.014 0.19 0.53 0.205 0.105 lower shell Circle Seam 11-894
Lower shell Sub-arc B4 90099 Linde 0091 3977 0.15 1.12 0.022 0.012 0.23 0.49 0.197 0.060 long seams20-894 A&B
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REV 21 5/08 TABLE 5.4-6 PREDICTED END OF LICENSE (54 EFPY) UPPER SHELF ENERGY VALUES FARLEY UNIT NO. 1 REACTOR VESSEL BELTLINE WELDS (Ref. 7)
1/4T Fluence Unirradiated Decrease in Projected EOL Beltline Material Wt. % Cu (10 19 n/cm 2) USE (ft-lb) USE (%) USE (ft-lb)
Intermediate Shell Longitudinal Welds 0.258 1.25 149 26 110 19-894 A & B using Surveillance Capsule Data Circumferential Weld 11-894 0.205 4.0 104 46 56 Lower Shell Longitudinal Welds 0.197 1.25 82.5 36 52.8 20-894 A & B
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-7 IDENTIFICATION OF UNIT NO. 2 REACTOR VESSEL BELTLINE REGION BASE MATERIAL (wt%)
Material Component Code No. Heat No. Spec. No.
C Mn P S Si Ni Mo Cu Cr Al Inter. shell B7203-1 C6319 A533B, CL.1 0.20 1.30 0.010 0.013 0.19 0.60 0.55 0.14 - 0.020
Inter. shell B7212-1 C7466 A533B, CL.1 0.21 1.30 0.018 0.016 0.24 0.60 0.49 0.20 0.15 0.040 Lower shell B7210-1 C6888 A533B, CL.1 0.24 1.28 0.010 0.014 0.20 0.56 0.56 0.13 - 0.020 Lower shell B7210-2 C6293 A533B, CL.1 0.19 1.30 0.015 0.015 0.18 0.57 0.59 0.14 - 0.026
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-8 PREDICTED END OF LICENSE (54 EFPY) UPPER SHELF ENERGY VALUES FARLEY UNIT NO. 2 REACTOR VESSEL BELTLINE PLATES (Ref. 7) 1/4T Fluence Unirradiated Decrease in Projected EOL Beltline Material Wt. % Cu (10 19 n/cm 2) USE (ft-lb) USE (%) USE (ft-lb)
Intermediate Shell Plate B7203-1 0.14 3.92 100 32 68 Intermediate Shell Plate B7212-1 0.20 3.92 100 42 58 using Surveillance Capsule Data Lower Shell Plate B7210-1 0.13 3.92 103 30 72 Lower Shell Plate B7210-2 0.14 3.92 99 32 67
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-9 IDENTIFICATION OF UNIT NO. 2 REACTOR VESSEL BELTLINE REGION WELD METAL Weld Wire Flux Composition (wt%) Weld Welding Location Process Type Heat Type Lot No. C Mn P S Si Mo Cu V Ni No.
Inter. shell SMAW E8018C3 HODA - - 0.09 1.00 0.009 0.010 0.38 0.25 0.027 0.010 0.947 long. seam 19-923A SMAW E8018C3 BOLA - - 0.09 0.95 0.004 0.014 0.34 0.23 0.027 0.006 0.913
Inter. shell SMAW E8018C3 BOLA - - 0.09 0.95 0.004 0.014 0.34 0.23 0.027 0.006 0.913 long. seam 19-923B Inter. shell to Sub-arc B4 5P5622 Linde 0091 1122 0.17 1.29 0.016 0.008 0.19 0.57 0.153 0.009 0.077 lower shell circle seam 11-923 Lower shell Sub-arc B4 83640 Linde 0091 3490 0.16 1.22 0.006 0.011 0.19 0.57 0.051 0.006 0.096 long. seams20-923 A&B
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-10 PREDICTED END OF LICENSE (54 EFPY) UPPER SHELF ENERGY VALUES FARLEY UNIT NO. 2 REACTOR VESSEL BELTLINE WELDS (Ref. 7) 1/4T Fluence Unirradiated Decrease in Projected EOL Beltline Material Wt. % Cu (10 19 n/cm 2) USE (ft-lb) USE (%) USE (ft-lb)
Intermediate Shell Longitudinal Welds 0.03 1.27 131 20 105 19-923A Intermediate Shell Longitudinal Welds 0.027 1.27 148 9.5 134 19-923B using Surveillance Capsule
data Circumferential Weld 11-923 0.153 3.92 102 40 61 Lower Shell Longitudinal Welds 0.051 1.27 126 20 101 20-923 A & B
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-11 (SHEET 1 OF 2)
SURVEILLANCE MATERIAL BELTLINE LOCATION AND FABRICATION HISTORY - FARLEY UNIT NO. 1 Surveillance Beltline Location of Material Surveillance Material Heat-Treatment Base metal Inter. shell plate B6919-1 1550 - 1650°F 4 hr-WQ 1200 - 1250°F 4 hr-AC 1125 - 1175°F 40 hr-FC to 600
°F Weld metal Inter. shell longitudinal 1125 - 1175°F 16 hr-FC Weld seams19-894 A & B SURVEILLANCE TEST SPECIMENS - TYPE, ORIENTATION, AND QUANTITY PER TEST CAPSULE - FARLEY UNIT NO. 1
Surveillance Specimen Material Orientation Charpy-V Tensile 1/2T-CT Bend Bar Base metal (plate B6919-1) Transverse 15 3 4 1 Base metal (plate B6919-1)
Longitudinal 15 3 4 - Weld metal Transverse 15 3 4 - HAZ metal (plate B6919-1)
Longitudinal 15 - - -
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-11 (SHEET 2 OF 2)
Surveillance Beltline Location of Material Surveillance Material Heat-Treatment Base metal Inter. shell plate B7212-1 1550 - 1650°F - 4 h-WQ, 1200 - 1250°F - 4 h-AC, 1125 - 1175°F - 18 h-FC Weld metal(a) Inter. shell long. weld seam 1125 - 1175°F - 13 h-FC HAZ metal Inter. shell plate B7212-1 1125 - 1175°F - 13 h-FC
SURVEILLANCE TEST SPECIMENS - TYPE, ORIENTATION, AND QUANTITY PER TEST CAPSULE - FARLEY UNIT NO. 2
Surveillance Specimen Material Orientation Charpy-V Tensile 1/2T-CT Base metal (plate B7212-1) Transverse 15 3 4 Base metal (plate B7212-1)
Longitudinal 15 3 4 Weld metal Transverse 15 3 4 HAZ metal (plate B7212-1)
Longitudinal 15 - -
- a. Surveillance weldment fabr icated using plate B7212-1 and B7203-1. Surveillance weldment was fabricated using the same type of wire (E8018C3) and the same heat of wire (heat No. BOLA) as was used to fabricate the intermediate shell longitudinal weld seam (19-923B) in the vessel. The same welding procedures (MA-511-D and A-244-110-8) were used by the vessel supplier to fabric ate the surveillance weldment and the intermediate shell longitudinal weld seam (19-923B).
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-12 (SHEET 1 OF 2)
SURVEILLANCE MATERIAL CHEMICAL COMPOSITION (wt%) -
FARLEY UNIT NO. 1 Element Plate B6919-1 Weld Metal Combustion Engineering Westinghouse Westinghouse
Analysis Analysis Analysis C 0.20 -- 0.13 S 0.015 0.013 0.009
N 2 -- 0.003 0.005 Co 0.008 0.16 0.018 Cu 0.14 0.10 0.014 Si 0.18 0.28 0.27 Mo 0.56 0.51 0.50 Ni 0.55 0.56 0.19 Mn 1.39 1.40 1.06 Cr -- 0.13 0.063 V -- <0.001 0.003 P 0.015 0.015 0.016 Sn -- 0.008 0.005 A1 0.025 -- 0.009 The surveillance weld was fabricated from sections of plate B6919-1 and adjoining
intermediate shell plate B6903-2, using weld wire representative of that used in the
original fabrication.
FNP-FSAR-5
REV 21 5/08 TABLE 5.4-12 (SHEET 2 OF 2)
SURVEILLANCE MATERIAL CHEMICAL COMPOSITION (wt%) -
FARLEY UNIT NO. 2 Element Plate B7212-1 Weld Metal C 0.21 <0.086 Mn 1.30 0.95 P 0.018 0.004 S 0.016 0.014 Si 0.24 0.34 Ni 0.60 0.89 Cr 0.15 <0.01 Mo 0.49 0.23 Cu 0.20 0.028 V 0.003 0.006 Co 0.027 0.010 Sn 0.011 0.002 A1 0.040 0.003 N 2 0.006 0.007
The surveillance weldment was fabricated with the same type of wire and the same heat of wire (wire type E8018C3 and wire heat No. BOLA) as was used to fabricate the longitudinal weld
seam (19-923 B) in the intermediate shell course of the vessel. The same welding procedures
were used to fabricate the surveillance weldment and the vessel weld seam (19-923 B).
REV 21 5/08 SURVEILLANCE CAPSULE ELEVATION VIEW JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.4-1
REV 21 5/08 UNIT 1 SURVEILLANCE CAPSULE PLAN VIEW JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.4-2 (SHEET 1 OF 2)
REV 21 5/08 UNIT 2 SURVEILLANCE CAPSULE PLAN VIEW JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.4-2 (SHEET 2 OF 2)
REV 21 5/08 IDENTIFICATION AND LOCATION OF FARLEY UNIT NO. 1 REACTOR VESSEL BELTLINE REGION MATERIAL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.4-3
REV 21 5/08 IDENTIFICATION AND LOCATION OF FARLEY UNIT NO. 2 REACTOR VESSEL BELTLINE REGION MATERIAL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.4-4
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REV 21 5/08 TABLE 5.5-1 REACTOR COOLANT PUMP DESIGN PARAMETERS
Design pressure (psig) 2485 Design temperature (°F) 650 Capacity per pump (gpm) 88,500 Developed head (ft) 264 NPSH required (ft) 170 Suction temperature (°F) 543.3 RPM nameplate rating 1200 Discharge nozzle, ID (in.)
27-1/2 Suction nozzle, ID (in.)
31 Overall unit height (ft-in.)
26-10 Water volume (ft
- 3) 57 Moment of inertia (ft-lb) 82,000 Weight, dry (lb) 197,000 Motor Type AC induction, single speed, air cooled Power (H.P.)
6000 Voltage, volts 4000 Insulation class Hot loop operation Class B Cold loop operation Class F Phase 3 Frequency (Hz) 60 Starting current 5120 @ 4000V Input, hot reactor coolant (kW) 4870 Input, cold reactor coolant (kW) 6165 Seal water injection (gpm) 8 Seal water return (gpm) 3
FNP-FSAR-5
REV 21 5/08
[HISTORICAL] [TABLE 5.5-2
REACTOR COOLANT PUMP QUALITY ASSURANCE PROGRAM
RT (a) UT (a) PT (a) MT (a) Castings yes yes Forgings
- 1. Main shaft yes yes 2. Main flange bolts yes yes 3. Flywheel (rolled plate) yes yes for bore Weldments, Pressure Boundary
- 1. Circumferential yes yes 2. Instrument connections yes
- a. RT - Radiographic
UT - Ultrasonic PT - Dye Penetrant MT - Magnetic Particle
]
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-3 STEAM GENERATOR DESIGN DATA (a)
Number of steam generators per Unit (No. ) 3 Design pressure, (psig) 2,485/1,085 RCS / Steam RCS hydrostatic test pressure (psig) 3,107 (tube side - cold)
Design temperature, (°F) 650/600 reactor coolant / steam Reactor coolant flow (lb/h) 32.7 x 10 6 Total head transfer surface area (ft 2) 54,500 Heat transferred (Btu/h) 3,168 x 10 6 Steam Conditions at full load, outlet nozzle:
Steam flow (lb/h) 4.08 x 10 6 Steam temperature
(°F) 515.5 Steam pressure (psig) 781 Maximum moisture carryover (wt %) 0.10
Feedwater (°F) 443.4 Overall height (ft-in. )
67-9 Shell OD, upper/lower (in. ) 177/136 Total number of U-tubes (No. ) 3,592 (plugged and unplugged)
U-tube outer diameter (in. ) 0.875 Tube wall thickness, (minimum) (in. ) 0.050 Number of manways/ID (No. ) 4 (16 inch) Number of inspection ports ID (No. ) 2 (4 inch) Number of inspection handholes ID (No. ) 6 (6 inch)
Rated Load No Load Reactor coolant water volume (ft
- 3) 1,168 1,168 Primary-side fluid heat (Btu) 30.6 x 10 6 29.9 x 10 6 content Secondary-side water volume (ft
- 3) 2,167 3,618 Secondary-side steam volume (ft
- 3) 3,645 2,193 Secondary-side fluid heat (Btu) 6.05 x 10 7 9.71 x 10 7 content
- a. Quantities are for each steam generator.
FNP-FSAR-5
REV 21 5/08 [HISTORICAL] [TABLE 5.5-4 (SHEET 1 OF 2)
STEAM GENERATOR QUALITY ASSURANCE PROGRAM
RT (a) UT (a) PT (a) MT (a) Tubesheet Forging yes yes Cladding yes (b) yes (c) Channel Head Forging yes yes Cladding yes Secondary Shell and Head Plates yes Tubes yes yes Nozzles (forgings) yes yes Weldments Shell, circumferential yes yes Cladding (channel head- yes tube sheet joint cladding restoration) Steam and feedwater yes yes nozzle-to-shell Support brackets yes Tube-to-tubesheet yes Instrument connections yes (primary and secondary) Temporary attachments yes after removal After hydrostatic test yes (all welds and complete channel head - where accessible)
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-4 (SHEET 2 OF 2)
RT (a) UT (a) PT (a) MT (a) Nozzle safe ends yes yes (if forgings)
- b. Flat surfaces only.
- c. Weld deposit areas only.
]
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-5 REACTOR COOLANT PIPING DESIGN PARAMETERS
Unit 1 Unit 2 Reactor inlet piping, ID (in.)
27.5 27.5 Reactor inlet piping, nominal wall 2.2975 2.3225 thickness (in.)
Reactor outlet piping, ID (in.)
29 29 Reactor outlet piping, nominal wall 2.420 2.445 thickness (in.)
Coolant pump suction piping, ID 31 31 (in.) Coolant pump suction piping, nominal 2.575 2.600 wall thickness (in.)
Pressurizer surge line piping, ID 11.188 11.188 (in.) Pressurizer surge line piping, 1.406 1.406 nominal wall thickness (in.)
Water volume, all loops and surge 1030 1030 line (ft 3) Design/operating pressure (psig) 2485/2235 2485/2235 Design temperature (°F) 650 650 Design temperature, pressurizer 680 680 surge line (°F)
Design pressure, pressurizer relief line From pressurizer to safety 2485 2485 valve (psig)
From safety valve to relief tank 600 600 tank (psig)
Design temperature, pressurizer relief line From pressurizer to safety 650 650 valve (°F) From safety valve to relief 600 600 tank (°F)
FNP-FSAR-5
REV 21 5/08
[HISTORICAL] [TABLE 5.5-6 REACTOR COOLANT PIPING QUALITY ASSURANCE PROGRAM
RT (a) UT (a) PT (a) Fittings and pipe (castings) yes yes Fittings and pipe (forgings) yes yes Weldments Circumferential yes yes Nozzle to runpipe (except no yes yes RT for nozzles less than 4 in. ) Instrument connections yes
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-7 DESIGN BASES FOR RESIDUAL HEAT REMOVAL SYSTEM OPERATION
Residual Heat Removal System Startup ~4 hours after reactor shutdown
Reactor Coolant System initial pressure
~425 (psig)
Reactor Coolant System initial temperature
~350 (°F)
Component cooling water design temperature 105 (°F) Cooldown time, hours after initiation of
~34 RHRS operation
Reactor Coolant System temperature, at 140 end of cooldown (°F)
Decay heat generation at 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> after 60.8 x 10 6 reactor shutdown (Btu/h)
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-8 RESIDUAL HEAT REMOVAL SYSTEM COMPONENT DATA
Residual Heat Removal Pump Number 2 Design pressure, (psig) 600 Design temperature (°F) 400 Design flow (gpm) 3750 Design head (ft) 280 Residual Heat Exchanger Number 2 Design heat removal capacity 29.5 x 10 6 Btu/h Tube-side Shell-side Design pressure (psig) 600 150 Design temperature (°F) 400 200 Design flow (lb/h) 1.87 x 10 6 2.8 x 10 6 Inlet temperature (°F) 140 105 Outlet temperature (°F) 124.3 115.6 Material Austenitic Carbon steel stainless steel Fluid Reactor Component coolant cooling water
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-9 PRESSURIZER DESIGN DATA
Item Value Design pressure (psig) 2485 Design temperature (°F) 680 Surge line nozzle diameter (in.)
14 Heatup rate of pressurizer using heaters only (°F/h) 55 Internal volume (ft
- 3) 1400
FNP-FSAR-5
REV 21 5/08
[HISTORICAL] [TABLE 5.5-10 PRESSURIZER QUALITY ASSURANCE PROGRAM
Item RT (a) UT (a) PT (a) MT (a) Heads Plates yes yes Cladding yes Shell Plates yes yes Cladding yes Heaters Tubing (b) yes yes Centering of element yes Nozzle yes yes Weldments Shell, longitudinal yes yes Shell, circumferential yes yes Cladding yes Nozzle safe-end (forging) yes yes Instrument connections yes Support skirt yes Temporary attachments after yes removal All welds, heads, and shell after yes hydrostatic test Final assembly All accessible exterior surfaces yes after hydrostatic test
]
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-11 PRESSURIZER RELIEF TANK DESIGN DATA
Item Value Design pressure (psig): Internal 100 External 15 Rupture disc release pressure (psig) 100 +/- 5% Design temperature (°F) 340 Total rupture disc relief capacity (lb/h at 100 psig) 1.14 x 10 6
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-12 REACTOR COOLANT SYSTEM BOUNDARY VALVE DESIGN PARAMETERS
Item Value Normal operating pressure (psig) 2235 Design pressure (psig) 2485 Preoperational plant hydrotest (psig) 3107 Design temperature (°F) 650 FNP-FSAR-5
REV 21 5/08
[HISTORICAL] [TABLE 5.5-13 REACTOR COOLANT SYSTEM VALVES QUALITY ASSURANCE PROGRAM
Boundary Valves, Pressurizer Relief and Safety Valves RT (a) UT (a) PT (a) Castings yes yes Forgings (no UT for valves yes yes 2 in. and smaller)
]
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-14 PRESSURIZER VALVES DESIGN PARAMETERS
Pressurizer Spray Control Valves Number 2 Design pressure (psig) 2485 Design temperature (°F) 650 Design flow for valves full open, each (gpm) 300 Pressurizer Safety Valves Number 3 Minimum relieving capacity, ASME rated flow 345,000 (lb/h)(per valve)
Set pressure (psig) 2485 Fluid Saturated steam Backpressure:
Normal (psig) 3 to 5 Expected during discharge (psig) 350 Pressurizer Power Relief Valves Number 2 Design pressure (psig) 2485 Design temperature (°F) 650 Relieving capacity at 2350 psig (lb/h) 210,000 (per valve)
Fluid (2335 psig)
Saturated steam
FNP-FSAR-5
REV 21 5/08 TABLE 5.5-15 MAIN STEAM VALVE DESIGN PARAMETERS MAIN STEAM ISOLATION VALVES
Number 6 Design Pressure (psig) 1085 Design temperature (°F) 600 Normal Operating Flow (lb/h) 3.875 x 10 6 Main Steam Bypass Valves Number 6 Design pressure (psig) 1085 Design Temperature (°F) 600 Actuator Type Piston FNP-FSAR-5
REV 21 5/08 TABLE 5.5-16 REACTOR VESSEL HEAD VENT SYSTEM EQUIPMENT DESIGN PARAMETERS
Reactor Vessel Head Vent Subsystem Valves Number (includes one manual valve) 5 Design pressure (psig) 2485 Design temperatures (°F) 650 Piping Vent line, nominal diameter (in.)
1 Design pressure (psig) 2485 Design temperature (°F) 650 Maximum normal operating temperature (°F) 620
REV 21 5/08 REACTOR COOLANT CONTROLLED LEAKAGE PUMP JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-1
REV 21 5/08 REACTOR COOLANT PUMP PERFORMANCE CURVE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-2
REV 21 5/08 REACTOR COOLANT PUMP SPOOL PIECE AND MOTOR SUPPORT STAND JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-3
REV 21 5/08 STEAM GENERATOR JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-4
REV 21 5/08 STEAM GENERATOR FLOW LIMITING DEVICE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-5
REV 21 5/08 PRESSURIZER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-6
REV 21 5/08 REACTOR VESSEL SUPPORTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-7
REV 21 5/08 DRY CONTAINMENT STEAM GENERATOR SUPPORTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-8
REV 21 5/08 REACTOR COOLANT PUMP SUPPORTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-9
REV 21 5/08 PRESSURIZER SUPPORTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-10
REV 21 5/08 CROSSOVER LEG RESTRAINTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-11
REV 21 5/08 CRDM SEISMIC SUPPORT PLATFORM PIPE SUPPORT CLAMP JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-12
REV 21 5/08 SIDEVIEW RVHVS AND SUPPORTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 5.5-13
FNP-FSAR-5
5.6-1 REV 21 5/08 5.6 INSTRUMENTATION APPLICATION
Process control instrumentation is provided for the purpose of acquiring data on the pressurizer
and, on a per-loop-basis, for the key process parameters of the reactor coolant system (RCS)
(including the reactor coolant pump motors), as well as for the residual heat removal system.
The pick-off points for the reactor coolant system are shown in drawings D-175037, sheet 1, D-205037, sheet 1, D-175037, sheet 2, D-205037, sheet 2, D-175037, sheet 3, and D-205037, sheet 3, and for the residual heat removal (RHR) system, in drawings D-175041 and D-205041.
In addition to providing input signals for the prot ection system and the plant control systems, the instrumentation sensors furnish input signals for monitoring and/or alarming purposes for the
following parameters:
A. Temperatures.
B. Flows.
C. Pressures.
D. Water levels.
E. Vibration.
In general, these input signals are used for the following purposes:
A. Provide input to the reactor trip system for reactor trips as follows:
- 1. Overtemperature-T.
- 2. Overpower-T.
- 3. Low-pressurizer pressure.
- 4. High-pressurizer pressure.
- 5. High-pressurizer water level.
- 6. Low primary coolant flow.
The following fluid parameter generates an input to the reactor trip system.
While not part of the reactor coolant system, it is included here for information.
(This is not a complete listing of reactor trip system inputs.)
- 7. Low-low steam generator level.
B. Provide input to the engineered safety features (ESF) actuation system as follows:
- 1. High differential pressure between any steam line and the other steam lines.
FNP-FSAR-5
5.6-2 REV 21 5/08 2. Low steam line pressure.
Although it is not part of the RCS, the following parameter, which also is sensed to generate an input to the reactor trip system, is included here for
purposes of completeness.
- 3. High steam flow coincident with low-low Tavg.
C. Furnish input signals to the nonsafety-related systems, such as the plant control systems and surveillance circuits so that:
- 1. Reactor coolant average temperature (Tavg) will be maintained within prescribed limits. The resistance temperature detector instrumentation is
identified on drawings D-175037, sheet 3, and D-205037, sheet 3.
- 2. Pressurizer level control, using Tavg to program the setpoint, will maintain the coolant level within prescribed limits.
- 3. Pressurizer pressure will be controlled within specified limits.
- 4. Steam dump control, using Tavg control, will accommodate sudden loss of generator load.
- 5. Information is furnished to the control room operator and at local stations for monitoring.
The following is a functional description of the system instrumentation. Unless otherwise stated, all indicators, recorders, and alarm annunciators are located in the plant control room.
A. Temperature Measuring Instrumentation
- 1. Mechanical
The individual loop temperature signals required for input to the reactor control and protection system are obtained using resistance temperature
detectors (RTDs) installed in each reactor coolant loop.
- a. Hot Leg
The hot leg temperature measurement on each loop is accomplished with three fast response, narrow range, dual
element RTDs mounted in thermowells. One element of the RTD
is considered active, and the other element is held in reserve as a
spare. To accomplish the sampling function of the RTD bypass
manifold system and to minimize the need for additional hot leg
piping penetrations, the thermowells are located within the three
existing RTD bypass manifold scoops wherever possible. A hole
is machined through the end of each scoop so that water flows in
through the existing holes in the leading edge of the scoop, past FNP-FSAR-5
5.6-3 REV 21 5/08 the RTD, and out through the new hole. Due to physical
limitations, several hot leg RTDs are located in independent
thermowells near the original scoop locations. These three RTDs
measure the hot leg temperature which is used to calculate the reactor coolant loop differential temperature (T) and average temperature (Tavg). One wide range RTD element is utilized in each hot leg. These elements, installed in dry thermowells, penetrate the reactor coolant piping and extend into the flow
stream. The wide range RTDs provide temperature indication on
temperature recorders.
- b. Cold Leg
One fast response, narrow range, dual element RTD is located in each cold leg at the discharge of the reactor coolant pump (RCP)
(as replacements for the cold leg RTDs located in the bypass
manifold). Temperature streaming in the cold leg is minimized by
the mixing action of the RCP. The cold leg RTD measures the
cold leg temperature which is used to calculate reactor coolant loop T and Tavg. The existing cold leg RTD bypass penetration nozzle was modified to accept the RTD thermowell. One element of the RTD is considered active, and the other element is held in
reserve as a spare. One wide range RTD element is utilized in
each cold leg. These elements, installed in dry thermowells, penetrate the reactor coolant piping and extend into the flow
stream. The wide range RTDs provide temperature indication on
temperature recorders.
- c. Crossover Leg
The RTD bypass manifold return line has been capped at the nozzle on the crossover leg.
- 2. Electrical
- a. Control and Protection System
The hot leg RTD measurements (three per loop) are electronically averaged in the process protection system. The averaged T hot signal is then used with the T cold signal to calculate reactor coolant loop T and Tavg which are used in the reactor control and protection systems. This is accomplished by additions to the
existing process protection system equipment. The T hot and T cold spare RTD elements are wired to the control rooms and
terminated at the 7300 rack input terminals. This arrangement
allows online accessibility to the spare elements for RTD cross
calibrations and facilitates connection of the spare RTD element in
the event of an RTD element failure.
FNP-FSAR-5
5.6-4 REV 21 5/08 The previous RCS loop tem perature measurement system used dedicated direct immersion RTDs for the control systems. This
was done largely to satisfy the IEEE Standard 279-1971 which
applied single failure criteria to control and protection system
interaction. The new thermowell mounted RTDs are used for both
control and protection. In order to continue to satisfy the requirements of IEEE Standard 279-1971, the Tavg and T signals generated in the protection system are electrically isolated and transmitted to the control system into median signal selectors for Tavg and T, which select the signal which is in between the highest and lowest values of the three loop inputs. This precludes an unwarranted control system response that could be caused by
a single signal failure.
- 3. Pressurizer Temperature
There are two temperature detectors in the pressurizer, one located in the vapor or steam space and one located in the water or liquid space. Both
detectors supply signals to temperature indicators and high-temperature
alarms. The steam space detector, located near the top of the
pressurizer, may be used during startup to determine water temperature
when the pressurizer is completely filled with water. The steam space
temperature is also used as part of an open permissive interlock to
prevent the residual heat removal system isolation valves from being
opened when the pressurizer steam space temperature is greater than 475°F. The liquid space temperature is used to determine the pressurizer spray differential temperature during heat up and cool down.
- 4. Surge Line Temperature
This detector supplies a signal for a temperature indicator and a low-temperature alarm. Low temperature is an indication that the
continuous spray rate is too small.
- 5. Safety and Relief Valve Discharge Temperatures
Temperatures in the pressurizer safety and relief valve discharge lines are measured and indicated. An increase in a discharge line temperature
is an indication of leakage through the associated valve or the valve being
open.
- 6. Spray Line Temperatures
Temperatures in the spray lines from two loops are measured and indicated. Alarms from these signals are actuated by low spray-water
temperature. Alarm conditions indicate insufficient flow in the spray lines.
FNP-FSAR-5
5.6-5 REV 21 5/08 7. Pressurizer Relief Tank Water Temperature
The temperature of the water in the pressurizer relief tank is indicated, and an alarm actuated by high temperature informs the operator that
cooling of the tank contents is required.
- 8. Reactor Vessel Flange Leakoff Temperature
The temperature in the leakoff line from the reactor vessel flange O-ring seal leakage monitor connections is indicated. An increase in
temperature above ambient is an indication of O-ring seal leakage. High
temperature actuates an alarm.
- 9. Reactor Coolant Pump Motor Temperature Instrumentation
- a. Thrust Bearing Upper and Lower Shoes Temperature
Resistance temperature detectors are provided, with one located in the shoe of the upper and one in shoe of the lower thrust
bearing. These elements provide a signal for a high-temperature
alarm and indication.
- b. Stator Winding Temperature
The stator windings contain six resistance-type detectors, two per phase, imbedded in the windings. A signal from one of these
detectors is monitored by the plant computer, which actuates a
high temperature alarm.
- c. Upper and Lower Radial Bearing Temperature
Resistance temperature detectors are located one in the upper and one in the lower radial bearings. Signals from these detectors
actuate a high-temperature alarm and indication.
B. Flow Indication
- 1. Reactor Coolant Loop Flow
Flow in each reactor coolant loop is monitored by three differential pressure measurements at a piping elbow tap in each reactor coolant
loop. These measurements on a two-out-of-three coincidence circuit
provide a low-flow signal to actuate a reactor trip.
FNP-FSAR-5
5.6-6 REV 21 5/08 C. Pressure Indication
- 1. Pressurizer Pressure
Pressurizer pressure transmitters provide signals for individual indicators in the control room for actuation of both a low-pressure trip and a
high-pressure trip.
One of the signals may be selected by the operator for indication on a pressure recorder.
Three transmitters provide low-pressure signals for safety injection initiation and for safety injection signal unblock during plant startup.
In addition, one transmitter is used, along with a reference pressure signal, to develop a demand signal for a three-mode controller. The lower
portion of the controller's output range operates the pressurizer heaters.
For normal operation, a small group of heaters is controlled by variable
power to maintain the pressurizer operating pressure. If the
pressure-error signal falls toward the bottom of the variable heater control
range, all pressurizer heaters are turned on. The upper portion of the
controller's output range operates the pressurizer spray valves and one
power relief valve. The spray valves are proportionally controlled in a
range above normal operating pressure with spray flow increasing as
pressure rises. If the pressure rises significantly above the proportional
range of the spray valves, a power relief valve (interlocked with P-11 to
prevent spurious operation) is opened. A further increase in pressure will
actuate a high-pressure reactor trip. A separate transmitter (interlocked
also with P-11 to prevent spurious operation) provides power relief valve
operation for a second valve upon high-pressurizer pressure.
- 2. Reactor Coolant Reference Pressure (Deadweight Test)
A differential pressure transmitter provides a signal for indication of the difference between the pressurizer pressure and a pressure generated by
a deadweight tester located outside the reactor containment. The
indication is used for online calibration checks of the pressurizer pressure
signals.
- 3. Reactor Coolant Loop Pressures
Two transmitting channels are provided. Each transmitting channel provides an indication of reactor coolant pressure on one of the hot legs.
This is a wide-range transmitter which provides pressure indication over
the full operating range. The wide range channel indicators serve as
guides to the operator for manual pressurizer heater and spray control
and letdown to the chemical and volume control system (CVCS) during
plant startup and shutdown. Amplified signals from the lower portion of
the range provide improved readability at the lower pressures.
FNP-FSAR-5
5.6-7 REV 21 5/08 The two wide-range channels provide the permissive signals for the residual heat removal loop suction line isolation valve interlock circuit. In
addition, the two channels each provide an input to both trains of the core
subcooling monitors.
There are also two local pressure gauges for operator reference during the shutdown condition located in two of the hot loops. These gauges are
equipped with auxiliary pointers which remain at the maximum pressure
measured until reset locally.
- 4. Pressurizer Relief Tank Pressure
The pressurizer relief tank pressure transmitter provides a signal to a pressure indicator and an annunciator on the main control board.
- 5. Reactor Coolant Pump Motor Pressure
- a. Oil Lift Switch
A dual-purpose switch is provided on the high-pressure oil lift system. Upon low oil pressure, the switch will actuate an alarm
on the main control board. In addition, the switch is part of an
interlock system that will prevent starting of the pump until the oil lift pump is started manually prior to starting the reactor coolant
pump motor. A local pressure gauge is also provided.
- b. Lower Oil Reservoir Liquid Level
A level switch is provided in the motor lower radial bearing oil reservoir. The switch will actuate a high- and low-level alarm on
the main control board.
- c. Upper Oil Reservoir Liquid Level
A level switch is provided in the motor upper radial bearing and thrust bearing oil reservoir. The switch will actuate a high- or
low-level alarm on the main control board.
D. Liquid Level Indication
- 1. Pressurizer Level
Three pressurizer liquid transmitters provide signals for use in the reactor control and protection system, the emergency core cooling system (ECCS), and the chemical and volume control system. Each transmitter
provides an independent high-water-level signal that is used to actuate an
alarm and a reactor trip. The transmitters also provide independent
low-water-level signals that will activate an alarm. Each transmitter also FNP-FSAR-5
5.6-8 REV 21 5/08 provides a signal for a level indicator that is located on the main control
board.
In addition to the above, signals may be selected for specific functions as follows:
- a. Any one of the three level transmitters may be selected by the operator for display on a level recorder located on the main control
board. This same recorder is used to display a pressurizer
reference liquid level.
- b. Two of the three transmitters perform the following functions. (A selector switch allows the third transmitter to replace either of
these two.)
(1) One transmitter provides a signal which will actuate an alarm when the liquid level falls to a fixed level setpoint.
The same signal will trip the pressurizer heaters "off" and
close the letdown line isolation valves.
(2) One transmitter supplies a signal to the liquid level controller for charging flow control and also initiation of a
low-flow (high-demand) alarm. This signal is also
compared to the reference level and actuates a high-level
alarm and turns on all pressurizer backup heaters if the
actual level exceeds the reference level. If the actual level
is lower than the reference level, a low alarm is actuated.
A fourth independent pressurizer level transmitter that is calibrated for low-temperature conditions provides water
level indication during startup, shutdown, and refueling
operations.
- 2. Pressurizer Relief Tank Level
The pressurizer relief tank level transmitter supplies a signal for an indicator and high- and low-level alarms.
E. Vibration Indication
Each of the reactor coolant pump assemblies is equipped for continuous monitoring of reactor coolant pump shaft and frame vibration levels. Shaft
vibration is measured by two relative shaft probes mounted on top of the pump seal housing; the probes are located 90
° apart in the same horizontal plane and mounted near the pump shaft. Frame vibration is measured by two velocity
seismoprobes located 90
° apart and mounted at the top of the motor support stand. Proximeters and converters provide output of the probe signals, which are displayed on meters in the electrical penetration room and annunciated in the
control room. These meters automatically indicate the highest output from the FNP-FSAR-5
5.6-9 REV 21 5/08 relative shaft probes and the frame seismoprobes. Manual selection allows
monitoring of individual probes. Indicator lights display caution and danger limits
of vibration, and are adjustable over the full range of the meter scale.
Process control instrumentation for the residual heat removal system is provided for the
following purposes:
A. Furnish input signals for monitoring and/or alarming purposes for:
- 1. Temperature indications.
- 2. Pressure indications.
- 3. Flow indications.
B. Furnish input signals for control purposes of such processes as follows:
- 1. Control valve in the residual heat removal pump bypass line so that it opens at flows below a preset limit and closes at flows above a preset
limit.
- 2. Residual heat removal inlet valves control circuitry. See section 7.6 for the description of the interlocks and requirements for automatic closure.
- 3. Control valve in the residual heat removal heat exchanger bypass line to control temperature of reactor coolant returning to reactor coolant loops
during plant cooldown.
- 4. Residual heat removal pump circuitry for starting residual heat removal pumps on "S" signal.