ML20133F838
| ML20133F838 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 08/31/1985 |
| From: | Congedo T, Kaiser W, Yanichko S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20133F744 | List: |
| References | |
| TAC-59902, WCAP-10910, NUDOCS 8510110246 | |
| Download: ML20133F838 (25) | |
Text
_-.
i l
WESilNGHOUSE CLASS 3 l l
CUSTOMER DESIGNATED DISTRIBUTION l WCAP-10910 ,
i HEAlVP AND C00LDOWN LIMIT CURVES FOR 1HE ALABAMA POWER COMPANY JOSEPH M. FARLEY UNIT 2 REACTOR VFSSEL W. T. Kaiser T. V. Congedo S. E. Yanichko
- 1. R. Mager August 1985 Approved: 1 J. NT Chirigos, Managhin Structural Materials Engineering Prepared by Westinghouse for the Alabama Power Company.
Work Performed Under Shop Order AIVJ-139 Although information contained in this report is nonproprietary, no distribution shall be made outside Westinghouse or its licensees without the customer's approval.
Westinghouse Electric Corporation Nuclear Energy Systems P.O. Box 355 Pittsburgh, Pennsylvania 15230 8510110246 BTWD930 PDR ADOCK 05000364 P pm 1281E:lD/082885
, 1. INTRODUCTION Heatup and cooldown limit curves are calculated using the most limiting value of RTNDT (refennce n%cMy temperature). He most H m W ng RTHDT of the material in the core region of the reactor vessel is determined by using the preservice reactor vessel material properties and estimating the radiation-induced ART is designated as the higher of either NDT* NDT the drop weight nil-ductility transition temperature (TNDT) r the
, temperature at which the material exhibits at least 50 ft Ib of impact energy and 35-mil lateral expansion (normal to the major working direction) minus 60*F.
RT increases as the material is exposed to fast-neutron radiation. Thus, NDT to find the most limiting RT at any time period in the reactor life, NDT i ART due to the radiation exposure associated with that time period must MDT be added to the original unirradiated RT ** " 5 "
NDT*
RT is enhanced by certain chemical elements (such as copper, nickel and NDT phosphorus) present in reactor vessel steels. Westinghouse, other NSSS vendors, the U.S. Nuclear Regulatory Comission and others have developed trend curves for predicting adjustment of RTNDT ** * ""C " "'"C' ""
copper, nickel and/or phosphorus content. The Nuclear Regulatory Comission (NRC) trend curve is published in Regulatory Guide 1.99 (Effects of Residual Elementr, on Predicting Radiation Damage to Reactor Vessel Materials)II} .
Regulatory Guide 1.99 was originally published in July 1975 with a Revision 1 being issued in April 1977. Currently, a Revision 2 to Regulatory Guide 1.99 is under consideration within the NRC. The chemistry factor, "CF", 'F, a function of copper and nickel content identified in Regulatory Guide 1.99, Revision 2 is given in Table I for welds and Table 11 for base metal (plates and forgings). Interpolation is permitted. The value, "f", given in Figure 1 is the calculated value of the neutron fluence at the location of interest (inner surface,1/4T, or 3/4T) in the vessel at the location of the postulated defect, n/cm (E > NeV) divided by 10 . The fluence factor is determined f rom Figure 1.
e a
1287E:lD/082385
, Given tha ccpper and nickel contents of the most limiting material, the radiation-induced ART NDT can be estimated from Tables I and II and Figure
~
- 1. Fast-neutron fluence (E > 1 MeV) at the inner surface,1/4T (wall i
thickness) and 3/4T (wall thickness) vessel locations are given as a function i of full-power service in Figure 2. The data for all other ferritic materials in the reactor coolant pressure boundary are examined to ensure tfiat no other component will be limiting with respect to RT NDT*
t
- 2. FRACTURE TOUGHNESS PROPERTIES 4
The preirradiation fracture-toughness properties of the Farley Unit 2 reactor vessel materials are presented in Table III. The fracture-toughness l properties of the ferritic material in the reactor coolant pressure boundary are determined in accordance with the NPC Regulatory Standard Review Plan (2) . The postirradiation fracture-toughness properties of the reactor vessel beltline material were obtained directly from the Farley Unit 2 Vessel Material Surveillance Program.
- 3. FLUENCE CALCULATIONS i For the purpose of revising heatup and cooldown curves for Farley Unit 2,
- which has limiting embrittlement characteristics in the vessel base metal, it is necessary to know vessel fast fluence ($ (E > 1 MeV)) at the azimuthal
! peak location. This peak location is at O', and at this angle, fast fluences are required at vessel inner radius, vessel 1/4T, and vessel 3/4T. The calculations performed for this purpose consist of adjoint analyses, relating the fast flux ($ (E > 1 MeV)) at the vessel IR to the power distributions in the reactor core. The adjoint (importance) functions used, when combined with cycle specific core power distributions, yield the plant specific exposure data for each operating fuel cycle.
The adjoint function was generated using the DOT discrete ordinates code (3) and the SAILOR cross-section library ) The SAILOR library is a 47 group, .
ENDF-B/IV based data set produced specifically for light water reactor
, applications. In generating the adjoint function, anisotropic scattering was treated with a P3 expansi n f the cross-sections. The adjoint source i
1287E:10/082985
,- , , e.--,--,.,---,-,..----,,...---,,,,,---,v-- ,,---,-n ,,--n-..-.,----,-,,-.,.---__gr.n-- _ . - ,.,,,.--..,--,,.n~.--,,.-wn-,y--.
- . + , -- ,--
. location was chosen along the inner diameter of the pressure vessel. This calculation was run in R, 6 geometry to provide a power distribution importance function for the exposure parameter of interest ($ (E > 1 MeV)).
llaving the adjoint importance function and appropriate core power distributions, the response of interest is calculated as ,
R R,0 " I R IO I(R,0) F(R,0) R dR de where:
R
= Response of interest ($ (E > 1.0 MeV), dPa, etc.) at radius R,0 R and azimuthal angle 9.
I(R,0) = Adjoint importance function at radius R and azimuthal angle G.
F(R,0) = Full power fission density at radius R and azimuthal angle 0.
It should be noted that as written in the above equation, the importance function 1(R,0) represents an integral over the fission distribution so that the response of interest can be related directly to the spatial distribution of fission density within the reactor core.
Core power distributions for Farley Unit 2 were taken from the following Westinghouse fuel cycle design reports for each operating cycle to date:
Fuel Cycle Report 1 WCAP-9710 2 WCAP-10187 3 WCAP-10410 4 WCAP-10674 1287E:1D/082385
, Of thssa, Cycles 1 and 2 utilizzd cut-in fumi lesding patterns, and Cycles 3 and 4 implemented low leakage fuel loading patterns.
The power distributions employed represent cycle averaged relative assembly powers. Therefore, the adjoint results are in terms of fuel cycle averaged neutron flux, which when multiplied by the fuel cycle length yields the incremental fast neutron fluence. Fast fluences at 1/4T and 3/4T are obtained from those at vessel IR through fast flux ratios obtained from the DOT I
transport analysis performed in support of WCAP-10425, " Analysis of Capsule U from the Alabama Power Company, Joseph M. Farley Unit 2 Reactor Vessel Radiation Surveillance Program". As a result, the following neutron fluences for E > 1.0 MeV were calculated:
Cumulative Fluence (E > 1 MeV) at 0" 2
Lifetime (n/cm )
EFPY Vessel IR Vessel 1/4T Vessel 3/4T I 18 1.09 2.027 x 10 1.200 x 10 2.789 x 10" 18 18 1.86 3.625 x 10 2.147 x 10 4.989 x 10" 18 18 2.95 5.527 x 10 3.273 x 10 7.606 x 10 "
18 18 3.20 5.898 x 10 3.493 x 10 8.116 x 10" I9 18 32.0 4.999 x 10 2.960 x 10 6.878 x 10
- 4. CRITERIA FOR ALLOWA8LE PRESSURE-TEMPERATURE RELATIONSHIPS The ASME approach for calculating the allowable limit curves for various heatup and cooldown rates specifies that the total stress intensity factor, Kg , for the combined thermal and pressure stresses at any time during heatup and cooldown cannot be greater than the reference stress intensity factor, K e e me a empuatum at hat dme. K s oMaind kom W IR, IR reference fracture toughness curve, defined in Appendix G to the ASME Code (5) The K IR curve is given by the equation:
l K = 26.78 + 1.223 exp (0.0145 (T-RTNDT +
IR 1287E:10/082385
where K is the reference stress intensity factor as a function of the l
IR
. metal temperature T and the metal reference nil-ductility temperature '
Thus, the governing equation of the heatup-cooldown analysis is l defined in Appendix G to the ASME Code ( } as follows:
CKgg + kit I IR (2) where:
K gg is the stress intensity factor caused by membrane (pressure) stress K is the stress intensity factor caused by the thermal gradients It K s a func n emperature to the RT f the material IR NDT C = 2.0 for Level A and Level B service limits C = 1.5 for hydrostatic and leak test conditions during which the reactor core is not critical.
At any time during the heatup or cooldown transient, K IR is determined by the metal temperature at the tip of the postulated flaw, the appropriate value of RTNDT, and the reference fracture toughness curve. The thermal stresses resulting from temperature gradients through the vessel wall are calculated and then the corresponding (thermal) stress intensity factors, K It, or the reference flaw are computed. From Equation (2), the pressure stress intensity factors are obtained and, from these, the allowable pressures are calculated.
i For the calculation of the allowable pressure-versus-coolant temperature during cooldown, the Code reference flaw is assumed to exist at the inside of the vessel wall. During cooldown, the controlling location of the flaw is always at the inside of the wall because the thernal gradients produce tensile stresses at the inside, which increase with increasing cooldown rates.
Allowable pressure-temperature relations are generated for both steady-state and finite cooldown rate situations. From these relations, composite limit curves are constructed for each cooldown rate of interest.
i 1287E:lD/082385
The use of the composite curve in the cooldown analysis is necessary because
- control of the cooldown procedure is based on measurement of reactor coolant temperature, whereas the limiting pressure is actually dependent on the material temperature at the tip of the assumed flaw. During cooldown, the 1/4T vessel location is at a higher temperature than the fluid adjacent to the vessel ID. This condition, of course, is not true for the steady-state situation. It follows that, at any given reactor coolant temperature, the AT developed during cooldown results in a higher value of K IR at the 1/4T location for finite cooldown rates than for steady-state operation.
i FurthermorelifconditionsexistsuchthattheincreaseinK IR exceeds K , the calculated allowable pressure during cooldown will be greater than the steady-state value.
The above procedures are needed because there is no direct control on temperature at the 1/4T location and, therefore, allowable pressures may unknowingly be violated if the rate of cooling is decreased at various intervals along a cooldown ramp. The use of the composite curve eliminates this problem and ensures conservative operation of the system for the entire cooldown period.
Three separate calculations are required to determine the limit curves for finite heatup rates. As is done in the cooldown analysis, allowable pressure-temperature relationships are developed for steady-state conditions as well as finite heatup rate conditions assuming the presence of a 1/4T defect at the inside of the vessel wall. The thermal gradients during heatup produce compressive stresses at the inside of the wall that alleviate the ;
tensile stresses produced by internal pressure ~ The metal temperature at the
)
i crack tip lags the coolant temperature; therefore, the K IR f r the 1/4T 1 crack during heatup is lower than the K IR f r %e W crack dudng steady-state conditions at the same coolant temperature. During heatup, especially at the end of the transient, conditions may exist such that the effects of compressive thermal stresses and lower K 3
,'s do not offset each other, and the pressure-temperature curve based on steady-state conditions no l longer represents a lower bound of all similar curves for finite heatup rates when the 1/4T flaw is considered. Therefore, both cases have to be analyzed l l
l 1 l 1287E:lD/082385 ;
- l
in order to ensure that at any coolant temperature the lower value of the allowable pressure calculated for steady-state and finite heatup rates is obtained.
The second portion of the heatup analysis concerns the calculation of pressure-temperature limitations for the case in which a 1/4T deep outside surface flaw is assumed. Unlike the situation at the vessel inside surface, the thermal gradients established at the outside surface during heatup produce stresses which are tensile in nature and thus tend to reinforce any pressure stresses present. These thermal stresses are dependent on both the rate of heatup and the time (or coolant temperature) along the heatup ramp. Since the thermal stresses at the outside are tensile and increase with increasing heatup rates, each heatup rate must be analyzed on an individuct basis.
Following the generation of pressure-temperature curves for both the steady-state and finite heatup rate situations, the final limit curves are produced as follows: A composite curve is constructed based on a point-by-point comparison of the steady-state and finite heatup rate data. At any given temperature, the allowable pressure is taken to be the lesser of the three values taken from the curves under consideration. The use of the composite curve is necessary to set conservative heatup limitations because it is possible for conditions to exist wherein, over the course of the heatup ramp, the controlling condition switches from the inside to the outside and the pressure limit must at all times be based on analysis of the most critical criterion. Then, composite curves for the heatup rate data and the cooldown rate data are adjusted for possible errors in the pressure and temperature sensing instruments by the values indicated on the respective curves in Figures 3 through 6. In addition, heatup and cooldown curves without instrument errors are presented in Figures 7 through 10.
Based on the Farley Unit 2 fracture analysis results from Reference 6, the 9 EFPY heatup and cooldown curves without instrument errors are impacted by the new 10CFR50 rule as shown by Figures 7 and 8. The 9 EFPY heatup curve with instrument errors is impacted by the 10CFR50 rule as shown by Figure 3.
However, Figure 4 indicates that the 9 EFPY cooldown curves are- not impacted 1304E:10/091385
by the 10CFR50 rule. The 32 EFPY heatup and cooldown curves with and without instrument errors in Figures 5, 6, 9, and 10 are not impacted by the 10CFR50 rule. Since there are many conservatisms (safety factor of 2 on pressure, K
gg toughness and 1/4T flaw) built into the ASME Appendix G analysis
! d in t alys re o p ant ra n can e se n heatup and cooldown curves without instrument errors.
i 1
An evaluation has been performed to determine the acceptability of the Overpressure Mitigation System (OMS) presently in Farley Unit 2 (Technical Specification 3/4.4.10.3) with respect to the 9 EFPY heatup and cooldown curves shown in Figures 7 and 8 respectively. For the purpose of the
]
evaluation it was assumed that the RHR relief valve lifts at 495 psig which includes 10% accumulation. The heatup curve in Figure 7 does not fall below 495 psig at any temperature. A comparison to cooldown curves in Figure 8 shows that in the low temperature range (<l40*F) cooldown rates of 20*F/Hr and l
lower fall well above 495 psig. Although the cooldown curves for rates of 40*F/Hr and above do fall below 495 psig, it is not expected that the Appendix 6 curves will be violated during an actuation of the OMS since cooldown rates greater than or equal to 40*F/Hr are highly unlikely at low temperature conditions. Therefore, the Appendix G curves as illustrated in Figures 7 and 8 will not be violated as the result of an actuation of the OMS.
- 5. HEATUP AND COOLDOWN LIMIT CURVES Limit curves for normal heatup and cooldown of the primary Reactor Coolant System have been calculated using the methods discussed previously. The derivation of the limit curves is presented in the NRC Regulatory Standard Review Plan I} .
Transition temperature shifts occurring in the pressure vessel materials due to radiation exposure have been obtained directly from the reactor pressure
+ vessel surveillance program.
Allowable combinations of temperature and pressure for specific. temperature change rates are below and to the right of the limit lines shown on the heatup and cooldown curves. The reactor must not be made critical until 1304E:10/091385
pressure-temperature combinations are to the right of the criticality limit line, shown in Figures 3, 5, 7 and 9. This is in addition to other criteria which must be met before the reactor is made critical. ,
l l
The leak test limit curve shown in Figures 3, 5, 7 and 9 represent minimum temperature requirements at the leak test pressure specified by applicable i codes (2,5) ,
i
- 6. AVAILABLE SURVEILLANCE CAPSULE DATA 18 2 Charpy test specimens from Capsule U irradiated to 5.61 x 10 n/cm indicate that the representative core region weld metal and limiting core shell plate 87212-1 exhibited maximum shifts in RT NDT f a 1334 ,
respective 1yIII.
. The weld metal ARTNDT of 10*F is well within the ART NDT prediction method from Reference 8. However, the shell plate ART NDT 133 4 exceeds the ART NDT prediction computed as fo11ows(8)
= [CF] [FF) = 123*F where CF = Chemistry Factor = 149 (from Table II for a copper content of 0.20 WT% and nickel content of 0.60 WT%)
FF = Fluence Factor = 0.82 (from Figure 1 at a fluence of 18 2 5.61 x 10 n/cm )
Therefore, the fluence factor is adjusted to reflect the shift of the shell plate surveillance capsule so that the ART NDT 's used to compute the heatup and cooldown curves include the surveillance capsule results.
- 7. SURVEllLANCE CAPSULE REMOVAL SCHEDULE The surveillance capsule withdrawal schedule for Unit 2 (Table IV) should remain the same as identified in the Technical Specifications and WCAP-10425I I. The dosimetry analysis of the second capsule to be removed af ter 4 EFPY should be used to re-evaluate the withdrawal schedule for the remaining capsules.
( 1304E:10/091385
REFERENCES 4
(1) Regulatory Guide 1.99, Revision 1, "Ef fects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials," U.S. Nuclear Regulatory Commission, April 1977.
(2) " Fracture Toughness Requirements," Branch Technical Position - MTE8 No.
5-2, Chapter 5.3.2 in Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants, LWR Edition, NUREG-0800,1981.
(3) Soltesz, R. G., Disney, R. K., Jedruch, J. and Ziegler, S. L., " Nuclear Rocket Shielding Methods, Modification, Updating and Input Data Preparation Vol. 5 - Two Dimensional, Discrete Ordinates Transport Technique," WANL-PR(LL)034, Vol. 5, August 1970.
(4) " Sailor RSIC Data Library Collection DLC-76," Coupled, Self-Shielded, 47 Neutron, 20 Gamma-Ray, P , Cross Section Library for Light Water 3
Reactors.
(5) 6SME Boiler and Pressure Vessel Code,Section III, Division 1 -
Appendices, " Rules for Construction of Nuclear Vessels," Appendix G,
" Protection Against Nonductile Failure," pp. 559-564, 1983 Edition, American Society of Mechanical Engineers, New York, 1983.
(6) " Response to NRC Comments on Farley Unit 2," ALA-85-706, July 1985.
(7) Kunko, M. K., Yanichko, S. E., Cheney, C. A. and Kaiser, W. T. " Analysis of Capsule U f rom the Alabama Power Company Joseph M. Farley Unit 2 Reactor Vessel Radiation Surveillance Program," WCAP-10425, October 1983.
(8) Regulatory Guide 1.99, Revision 2, "Ef fects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials" (Proposed Draf t),
- U.S. Nuclear Regulatory Commission, June 1984.
1281L:lD/082885
TABLE I CHEMISTRY FACTOR FOR WELDS, 'F Copper, Nickel, Wt. %
Wt. 5 0 0.20 0.40 0.60 0.80 1.00 1.20 0 20 20 20 20 20 20 20 0.01 70 70 20 20 20 20 20 0.02 21 26 27 27 27 27 27 0.03 22 35 41 41 41 41 41 0.04 24 43 54 54 54 54 54 0.05 ,
26 49 67 68 68 68 68 0.06 29 52 77 82 82 82 82 0.07 32 55 85 95 95 95 95 0.08 36 58 90 106 108 108 108 0.03 40 61 94 115 122 122 122 0.10 44 65 97 122 133 135 135 0.11 49 68 101 130 144 148 148 0.12 52 72 103 135 153 161 161 0.13 58 76 106 139 162 172 176 0.14 61 79 109 142 168 182 188 0.15 66 84 112 146 175 191 200 0.16 70 88 115 149 178 199 211 0.17 75 92 119 1 51 184 207 221 0.18 79 95 122 154 187 214 230 0.19 83 100 126 157 191 220 238 0.20 88 104 129 160 194 223 245 0.21 92 108 133 164 197 229 252 0.22 97 112 137 167 200 232 257 0.23 101 117 140 169 203 236 263 0.24 105 121 144 173 206 239 268 0.25 110 126 148 176 209 243 272 0.26 113 130 151 180 212 246 276 0.27 119 134 155 184 216 249 280 0.28 122 138 160 187 218 251 284 0.29 128 142 164 191 222 254 287 0.30 131 146 167 194 225 257 290 0.31 136 151 172 198 228 260 293 0.32 140 155 175 202 231 263 296 0.33 144 160 180 205 234 266 299 0.34 149 164 184 209 238 269 302 0.35 153 168 187 212 241 272 305 0.36 158 172 191 216 245 275 308 0.37 162 177 196 220 248 ~ 278 311 0.38 166 182 200 223 250 281 314 0.39 171 185 203 227 254 285 317 0.40 175 189 207 231 257 288 320 1287E:10/082285
TABLE II CHEMISTRY FACTOR FOR BASE METAL, 'F Copper, Nickel, Wt. %
Wt. % 0 0.20 0.40 0.60 0.80 1.00 1.20 l
0 20 20 20 20 20 20 20 l 0.01 20 20 20 20 20 20 20 0.02 20 20 20 20 20 20 20 0.03 20 20 20 20 20 20 20 0.04 22 26 26 26 26 26 26 0.05 25 31 31 31 31 31 31 0.06 28 37 37 37 37 37 37 0.07 31 43 44 44 44 44 44 0.08 34 48 51 51 51 51 51 0.09 37 53 58 58 58 58 58 0.10 41 58 65 65 67 67 67 0.11 45 62 72 74 77 77 77 0.12 49 67 79 83 86 86 86 0.13 53 71 85 91 96 96 96 0.14 57 75 91 100 105 106 106 0.15 61 80 99 110 115 117 117 0.16 65 84 104 118 123 125 125 0.17 69 88 110 127 132 135 135 0.18 73 92 115 134 141 144 144 0.19 78 97 120 142 150 154 154 0.20 82 102 125 149 159 164 165 0.21 86 107 129 155 167 172 174 0.22 91 112 134 161 176 181 184 0.23 95 117 138 167 184 190 194 0.24 100 121 143 172 191 199 204 0.25 104 126 148 176 199 208 214 0.26 109 130 151 180 205 216 221 0.27 114 134 155 184 211 225 230 0.28 119 138 160 187 216 233 239 0.29 124 142 164 191 221 241 248 0.30 129 146 167 194 225 249 257 0.31 134 151 172 198 228 255 266 0.32 139 155 175 202 231 260 274 0.33 144 160 180 205 234 264 282 0.34 149 164 184 209 238 268 290 0.35 153 168 187 212 241 272 298 0.36 158 173 191 216 245 275 303 0.37 162 177 196 220 248 278 308 0.38 166 182 200 223 250 281 313 0.39 171 185 203 227 254 285 317 0.40 175 189 207 231 257 288 320 1287E:10/082285
- = - . _. _. .
e 9
TABLE 111 FARLEY UNIT 2 REACTOR VESSEL TOUGHNESS DATA Aserage Upper i Shelf Er.erav Normal to Principal Principa' Working Working Cu P N1 T RT Direction Directior NOT NOT Component Code No. Grade LH LQ Lu {'fQ l'fQ fft-lb) (ft-lb1 CL. HD. Dome 87215-1 A533,8,CL.1 0.17 0.010 0.49 16(a) 83(a) 128 CL. HD. Flange 87207-1 A508,CL.2 0.14 0.011 0.65 -30(a) 60 60(a) >$6(a) >86(c)
VES. Flange 87206-1 A508,CL.2 0.10 0.012 0.67 60(a) 60(a) >7)(a) > jog Inlet Noz. 87218-2 A508,CL.-2 -
0.010 0.68 50(a) So(a) 103(a) 158 Inlet Noz. 87218-1 A508,CL.2 -
0.010 0.71 32(a) 32(a) 112(a) 172 Inlet Noz. 87218-3 A508,CL.2 - 0.010 0.72 60(a) 60(a) gg(a) 150 Outlet Noz. 87217-1 A508,CL.2 -
0.010 0.73 60(a) 60(a) 100(a) 154 Outlet Noz. 87217-2 A508,CL.2 - 0.010 0.72 6(a) 6(a) 108(a) 167 Outlet Noz. 87217-3 A508,CL.2 -
0.010 0.72 48(a) 4s(a) 10)(a) 158 Upper Shell 87216-1 A508,CL.2 - 0.010 0.73 30 30(a) g7ta) 14g Inter Shell 87203-1 A533,8,CL.1 0.14 0.010 0.60 -40 15 99 140 Inter Shell 87212-1 A533,8,CL.1 0.20 0.018 0.60 -30 -10 99 134 Lower Shell 87210-1 A533,8,CL.1 0.13 0.010 0.56 -40 18 103 128 Lower Shell 87210-2 A533,8,CL.1 0.14 0.015 0.57 -30 0 99 145 Trans. Ring 87208-1 A508,CL.2 -
0.010 0.73 40 40(a) 89(a) 137 Bot. HD. Dome 87214-1 A533,8,CL.1 0.11 0.007 0.48 -2(a) 87(a) 134 Inter. Shell A1.46 SMAW 0.02 0.009 0.96 -30 Ota ) o(a) >131 .
Long seems A1.40 SRAW 0.02 0.010 0.93 -60 -60
~
>106 -
Inter Shell to Lower Shell 61.50 SAW 0.13 0.016 <.20(b)
- 40
-40 >102 -
Lower Shell Long Seams G1.39 $ MAW 0.05 0.006 <.20(b) -70 -70 >126 -
(a) Estimate per NUREG 0000 'USNRC Standard Review Plan
- Branch Technical Position MTE8 5-2.
(b) Estimated.
(c) Upper shelf not available, value represents minimum energy at the highest test temperature.
8D 1287E:10/002385
o TABLE IV SURVEILLANCE CAPSULE REMOVAL SCHEDULE The following removal schedule is recommended for future capsules t,o be removed from the Farley Unit 2 reactor vessel:
Lead Estimated Fluence l9 Capsule Factor Removal Time "3 n/cm x 10 U
3.12 Removed (1.1) .56 (Actual)
W' 2.70 4 2.18 X 3.12 6 3.78 Z 2.70 12 6.54 ECl V 3.12 18 11.34 Y 2.70 Standby -
[a] Effective ful,1 power years from plant startup
[b] Approxinates vessel end of life 1/4 thickness wall location fluence
[c] Approximates vessel end of life inner wall location fluence 1287E:10/082985
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MATERIAL PROPERTY RARTM
- CO EROLLING MATERIAL : R. V. INTERMEDIATE SHELL COPPER CONTEE : 0.20 WT5 NICKEI. CONTENT : 0.6gWI1 INITIAL RT : -10 F ET RT ET AFTER 9 E PY : 1/4T, 146 0F i
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FIGURE 3 FARLEY UNIT 2 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS APPLICABLE FOR THE FIRST 9 EFPY
MATERIAL PROPERTY BAMS CONTROLLING MATERIAL : R. V. INTERMEDIATE SHELL
, COPPER CONTENT : 0.20 WT5 NICKEL CONTENT
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FIGURE 4 FARLEY UNIT 2 REACTOR COOLANT SYSTEM C00LDOWN LIMITATIONS APPLICABLE FOR THE FIRST 9 EFPY
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FIGURE 5 FARLEY UNIT 2 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS APPLICABLE UP TO 32 EFPY
, MATERIAL PROPERTY BASIS
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, COPPER CONTENT : 0.20 WT%
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, MATERIAL PROPERTY BASIS 9
COEROLLING MWERIAL : R. V. INTERMEDIATE SHD.L COPPER C0hTEM : 0.20 WIT, NICKEL CO R EhT : 0.6p WU INITIAL RTET : -107 RT ET AFTER 9 EFPY : 1/47, 146 *
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