ML18312A072

Revision 28 to Updated Final Safety Analysis Report, Chapter 4, Reactor and Chapter 5, Reactor Coolant System and Connected System

ML18312A072
Person / Time
Site:	Farley
Issue date:	10/30/2018
From:	Southern Nuclear Operating Co
To:	Office of Nuclear Reactor Regulation
Shared Package
ML18312A093	List: ML18312A062 ML18312A068 ML18312A069 ML18312A070 ML18312A071 ML18312A072 ML18312A073 ML18312A074 ML18312A077 ML18331A123 NL-18-1299, Technical Requirements Manual
References
NL-18-1299
Download: ML18312A072 (659)

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 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 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.

90. South Texas, FSAR, Docket No. 50-498.

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

5.1-1 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.

B. Steam Generators

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

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.

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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.

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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)

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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

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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

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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)

ANSI B36.19

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

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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.

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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

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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.

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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)

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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.

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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.

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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

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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

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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

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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.

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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.

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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

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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

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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%

1. 5 - 80%

1. 12 - 70%

1. 16 - 80%

1. 20 - 80%

1. 24 - 70%

1. 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%

1. 32 - 50%

1. 14 - 90%

1. 11 - 30%

1. 20 - 30%

1. 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%

1. 4 80% #2 80%

1. 1-5 - 80% #2 80%

1. 1 80% #2 80%

1. 1 80% #2 80%

1. 1 80% #3-5 - 80%

1. 1 80% #3 80%

1. 1 80% #3 80%

1. 1 80% #3 80%

1. 1 80% #3 80%

1. 2-5 - 80% #3 80%

1. 2 80% #3 80%

1. 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 5.2-40 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

1. 4 3.57 3.57 #3 and #4 3/4" HILTI KWIK 2 1 2RC-131X 5 3 2 #2

1. 5 3.77 3.20 #2 AND #5 1/2" HILTI KWIK

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

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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).

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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

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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.)

149-9/16 OD 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

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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

]

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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

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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

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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

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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

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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

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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

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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 - - -

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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).

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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.

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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

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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

]

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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.

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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)

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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)

a. RT - Radiographic UT - Ultrasonic PT - Dye penetrant MT - Magnetic particle

b. Flat surfaces only.

c. Weld deposit areas only.

]

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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)

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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

a. RT - Radiographic UT - Ultrasonic PT - Dye penetrant]

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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)

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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

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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

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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

a. RT - Radiographic UT - Ultrasonic PT - Dye Penetrant MT - Magnetic Particle

b. Or a UT and ET.

]

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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

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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)

a. RT - Radiographic UT - Ultrasonic PT - Dye penetrant

]

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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

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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.