ML20138N375
| ML20138N375 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 10/25/1985 |
| From: | ALABAMA POWER CO. |
| To: | |
| Shared Package | |
| ML20138N363 | List: |
| References | |
| TAC-60075, NUDOCS 8511050174 | |
| Download: ML20138N375 (72) | |
Text
{{#Wiki_filter:r. Attachment 1 Proposed Changed Pages Unit 1 Page 3/4 4-29 Page 3/4 4-30 Page B3/4 4-6 Page-B3/4 4-7
.Page B3/4 4-8 Page B3/4 4-8a Page B3/4 4-9 Page B3/4 4-10 Page B3/4 4-10a Page B3/4 4-14 8511050174 851025 PDR ADOCK 05000348 P PDR 4
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lKTERIAL PRDPERTY RASISr Controlling Material : Wald Metal Copper Content : 0.22 WTX Nickel Content : 0.2gWT5 Initial RT ET : -56 F RT ET After 24 EFPY : 1/47,147'g
- 3/4T, 102 F Curve applicable for heatup rates up to 60*F/hr for the service perfod up to 24 EFPY 3000.0 LEAK TEST LIMIT
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, CRITICALITY LIMIT BASED ON INSERVICE HYDRO 5TATIC TEST TE R RATURE (275'F)
FOR TE SERVICE PERIOD - UP TO 24 EFPY i _ i I 0.0 0.0 100.0 360.0 300.0 444.s tes.4 l 1801CAtte TEnttRATURE totG.5) Fic3ure 3.4-2 Farley Unit 1 Reactor Coolant System Heatup Limitatior.s Applicable for the First 24 EFPY
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l l FARLEY - UNIT 1 3/4 4-29 AMENDMENT NO. 1 I
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MATERIAL PROPERTY BASIE : Controlling Mat rial : Wald Metal Copper Content : 0.22 WT3 Nickel. Content : 0.2gWT5 4 Initial RT : -56 F ET WT After 24 EFPY RT : 1/4T,147g
- 3/4T, 102 F l
Curves applicable for cooldown rates up to 100*F/hr for the service period up to 24 EFPY 3000.0
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100 - 8.8 ' 9.4 169.9 399.0 398.0 498.0 188 8 It01CATED TEMPERATURE (DEG.F) Figure 3.A-3 Farley Unit 1 Reactor Coolant Syster Cocidowr. Limitatices Applicable for the First 24 EFPY FAPLEY - UNIT 1 3/4 4-30 AMENDMENT NO.
REACTOR COOLANT SYSTEM BASES _____ - - - - - - - - - - - - , _ _ _ _ _ , _ _ =- ___ ___ Reducing Tayo to less than 500*F prevents the release of activity should a steam generator tu6e rupture since the saturation pressure of the primary coolant is below the lift pressure of the atmospheric steam relief valves. The surveillance requirements provide adequate assurance that excessive specific activity levels in .the primary coolant will be detected in sufficient tine to take corrective action. Information obtained on iodine spiking will be used to assess the parameters associated with spiking phenomena. A reduction in frequency of isotopic analyses following power changes may be permissible if justified by the data obtained. 3/4.4.10 PRESSURE / TEMPERATURE LIMITS The temperature and pressure changes during heatup and cooldown are limited to be consistent with the requirements given in the ASME Boiler and Pressure Vessel Code, Section III, Appendix G as required per 10CFR Part 50 Appendix G.
- 1) -The reactor coolant temperature and pressure and system heatup and cooldown rates (with the exception of the pressurizer) .shall be limited in accordance with Figures 3.4-2 and 3.4-3.
a) Allowable combinations of pressure and temperature for specific temperature change rates are below and to the.right of the limit lines shown. Limit lines for cooldown rates between those presented may be obtained by interpolation. b) Figures 3.4-2 and 3.4-3 define limits to assure prevention of nonductile failure only. For normal operation, other inherent plant characteristics, e.g., pump heat addition and pressurizer heater capacity, mgy limit the heatup and cooldown rates that can be achieved over certain pressure-temperature ranges.
- 2) These limit lines shall be calculated periodically using methods provided bel ow.-
- 3) The secondary side of the steam generator must not be pressurized above 200 psig if the temperature of the steam generator is below 70*F.
e 5-i- 5 FARLEY-UNIT 1 B 3/4 4-6 AMENDMENT N0. s
' REACTOR COOLANT SYSTEM
. BASES
- 4) The pressurizer heatup' and cooldown rates shall not-exceed 100*F/hr and 200*F/hr respectively. The spray shall not be used if the temperature difference between the pressurizer and the spray fluid is greater than 320*F.
-5) System preservice hydrotests and in-service leak and hydrotests shall be performed at pressures in accordance with the requirements of ASME Boiler and Pressure Vessel Code, Section XI.
The fracture toughness properties of the ferritic materials in the reactor vessel .are determined in accordance with ASTM E185-82, and in accordance with additional reactor vessel requirements. These properties are then evaluated in accordance with Appendix G of the 1976 Summer Addenda to Section III of the ASME Boiler and Pressure Vessel Code and the calculation methods described in WCAP-7924-A, " Basis for Heatup and Cooldown Limit Curves, April 1975." Heatup and cooldown limit curves are calculated using the most limiting value of the nil-ductility reference temperature, RTndt, at the end of 24 effective full power years of service life. The 24 EFPY service life period is chosen such that the limiting RTndt at.the 1/4T location in the
. core region is greater than the RTndt of the limiting unirradiated material. The selection 'of such a limiting RTndt assures that all components in the Reactor Coolant System will be operated conservatively in accordance with applicable Code requirements.
The reactor vessel materials have been tested to determine their initial RTndt; the results of these tests are shown in Table B 3/4.4-1. Reactor operation and resultant fast neutron (E greater than 1 MEV) irradiation can cause.-an increase in the RTndt. Therefore, an adjusted reference temperature, based upon the fluence and copper and nickel contents of the material .in question, can be predicted using Figure B 3/4.4-1 and the recommendations of Regulatory Guide 1.99, Revision 2, " Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials". The l Charpy test results from Capsule U'(from the Alabama Power Company, Joseph M.. Farley Unit 1 Reactor Vessel Radiation Surveillance Program) were used to determined the ARTndt due to irradiation effects. These Charpy test specimens from Capsule U irradiated to 1.65 x 1019 n/cm2 indicate that the representative core region weld metal and limiting core shell plate B6919-1 exhibited maximum shifts in RTndt of 80*F and 90'F, respectively. The weld metal located between the intermediate and lower-shell is considered to be the limiting vessel material. The ARTndt prediction for the material at 1.65 X 1019 n/cm2 is computed'as follows:
. ARTndt = I.CF] [FF] = 125' F where CF = Chemistry Factor = 112 (from Revision 2 of R.G.1.99 for a weld having a copper content of 0.22 WT% and nickel content of 0.20 WT%)
FF = Fluence Factor = 1.12 (from Figure B 3/4.4-2 at a fluence of 1.65'x 1019 n/cm2)
.FARLEY-UNIT 1 B 3/4 4-7 AMENDMENT NO.
REACTOR COOLANT SYSTEM BASES
-, -- - - -_ -- ===_
Because the surveillance weld was not fabricated with the same weld wire and lot of flux as the limiting intermediate to lower shell girth weld, the larger chemistry factor resulting for the limiting weld was used to predict A RTndt'. s. These ARTndt's which were used to compute the heatup and cooldown curves are more conservative than those obtained from the surveillance capsule results. Although the initial RTndt for the welds was conservatively estimated (see Table B 3/4.4-1) as 0 degrees-F per NUREG-0800, "USNRC Standard Review Plan", Branch Technical Position MTEB 5-2, a generic value of -56 degrees-F was used for the initial RTndt along with a margin as permitted per Regulatory Guide 1.99 Revision 2. The margin used was calculated as follows: Margin = 2 + 2 g = 66*F where (I = 17 degrees-F
/A = 28 degrees-F.
The adjusted reference temperature (ART) for the limiting weld in the beltline at 1/4 T was determined by the following expression: ART = Initial RTndt + ARTndt + Margin or ART = -56 degrees-F + 137 degrees-F-(24 EFPY) + 66 degrees-F
= 147 degrees-F.
The resulting heatup and cooldown limit curves of Figures 3.4-2 and 3.4-3 include predicted adjustments for this shif t in RTndt at the end of 24 EFPY. Values of ARTndt determined in this manner may be used until the next results from the material surveillance program, evaluated according to ASTM E185-82, are available. Capsules will be removed in accordance with the requirements of ASTM E185-82 and 10 CFR 50, Appendix H. The surveillance specimen withdrawal schedule is shown in Table 4.4-5. The heatup and cooldown curves must be recalculated when the ARTndt determined from the next surveillance capsule exceeds the calculated ARTndt for the equivalent l capsule radiation exposure. Allowable pressure-temperature relationsh.ps for various heatup and cooldown rates are calculated using methods derived from Appendix G in Section III of the ASME Boiler and Pressure Vessel Code as required by Appendix G to 10 CFR Part 50 and these methods are discussed in detail in WCAP-7924-A. FARLEY-UNIT 1 B 3/4 4-8 AMENDMENT NO.
REACTOR COOLANT SYSTEM BASES The. general method for calculating heatup and cooldown limit curves is based upon the principles of the linear elastic fracture mechanics (LEFM) technology. . In the calculation procedures a semi-elliptical surface defect with a depth of one-quarter of the wall thickness, T, and a length of 3/2T j is assumed to exist at the inside of the vessel wall as well as at the outside of the vessel wall. The dimensions of this postulated crack,
; referred to in Appendix G of ASME Section III as the reference flaw, amply ; exceed the current capabilities of inservice inspection techniques.
Therefore, the reactor operation limit curves developed for this reference crack are conservative and provide sufficient safety margins for protection against non-ductile failure. To assure that the radiation embrittlement effects are accounted for in the calculation of the limit curves, the most limiting value of the nil-ductility reference temperature, RTndt, is used and this includes the radiation induced shift, 6RTndt, corresponding to the end of the period for which heatup and .cooldown curves are generated. 4 i f i ) 4 FARLEY-UNIT 1 B 3/4 4-8a AMENDMENT N0. Y d
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Table B 3/4.4-1 5 _ i FARLEY UNIT 1 REACTOR VESSEL TOUGHNESS PROPERTIES , ,
$ Material Cu P- Ni T NOT RT NDT Upper Shell Energy
' ~ [ Component Code No. Type (%) (%)- .(%) (*F) (*F) MWD (c) NMWO(d) Closure head dome B6901 A533,B,C1.1 0.16 .0.009 0.50 -30 -20[a] 140 - Closure head segment B6902-1 A533,B,C1.1 0.17 0.007 0.52 -20 -20[a] 138 - Closure head flange B6915-1 A508 C1.2 0.10 0.012 0.64 60[a] 60[a] 75[a] _ Vessel flange B6913-1 A508, C1.2 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 4 Inlet nozzle B6917-2 A508, C1.2 - 0.008 0.80 60[a] 60[a] - 80 Inlet nozzle B6917-3 A508, C1.2 - 0.008 0.87 60[a] 60[a] - 98 Outlet nozzle B6916-1 A508, C1.2 - 0.007 0.77 60[a] 60[a] - 96.5 i Outlet nozzle B6916-2 A508, C1.2 - 0.011
- 0.78 60[a] 60[a] -
97.5 , Outlet nozzle B6916-3 A508, C1.2 - 0.009 0.78 60[a] 60[a] _ 100 Nozzle shell B6914-1 A508, C1.2 - 0.010 0.68 30 30[a] 148 -
' 5 Inter. shell 86903-2 A533,B,C1.1 0.13 0.011 0.60 0 0 151.5 97 R Inter shell 86903-3 A533,B,C1.1 0.12 0.014 0.56 10 10 134.5 100
- Lower shell B6919-1 A533,B,C1.1 0.14 0.015 0.55 -20 15 133 90.5
? Lower shell B6919-2 A533,B,C1.1 0.14 0.015 0.56 -10 5 134 97
- Bottom head ring B6912-1 A508 C1.2 -
0.010 0.72 10 10[a] 163.5- - Bottom head segment B6906-1 A533,5,C1.1 0.15 0.011 0.52 -30 -30[a] 147 _ Bottom head dome B6907-1 A533,B,C1.1 0.17 0.014 0.60 -30 -30[a] 143.5 3 - l Inter. shell long. M1.33 ; , Sub Arc Weld 0.25 0.017 0.21 0[a] o[a] _ 4 weld seas s , , ,' Inter. to lower G1.T 8 Sub Arc Weld 0.22 0.011 <0.20[b] 0[a] o[a] _ shell weld seams , Lower shell long. G1.08 Sub Arc Weld 0.17 0.022 <0.20[b] 0[a] o[a] _ weld seams ,
- [a] Estimate per NUREG-0900 "USNRC Standard Review Plan" Branch Technical. Position MTEB 5-2.
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l FARLEY - UNIT 1 B 3/4 4-10a AMENDMENT N0. i 1
. REACTOR COOLANT SYSTEM
. : BASES
=- _____ _ _____ =,,,,-
The use of the composite curve is necessary to set conservative heatup , limitations because it is possible for conditions to exist such that 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. Finally, the 10 CFR Part 50, Appendix G Rule which addresses the metal temperature of the closure head flange and vessel flange must be considered.
'This Rule states that the minimum metal temperature of the closure flange regions be at least 120*F higher than the limiting RTndt for these regions when the pressure exceeds 20 percent of the preservice hydrostatic test pressure (621 psig.for Farley Unit 1). In addition, the new 10 CFR Part 50 Rule states that a , . plant specific fracture evaluation may be perfonned to justify 1.ess limiting requirements. As a result, such a fracture analysis was performed for Farley Unit 2. These Farley Unit 2 fracture analysis results are applicable to Farley-Unit 1 since the pertinent parameters are identical for both plants. Based.upon this fracture analysis, the 24 EFPY heatup and cooldown curves are impacted by the new 10 CFR Part 50 Rule 'as shown on Figures 3.4-2 and 3.4-3.
_ Although' the pressurizer operates in temperature ranges above those for which there is reason-for concern of non-ductile failure, operating limits are . provided to assure compatibility of operation with the fatigue analysis perfonned in accordance with the ASME Code requirements. The OPERABILITY of two RHR relief valves or an RCS vent opening of greater than or equal to 2.85 square inches ensures that the RCS will be protected from _ pressure transients which could exceed the limits of Appendix G to 10CFR Part 50 , when one or more of the RCS cold legs are less than or equal' to 310*F. Either RHR relief. valve has adequate relieving capability to protect the RCS from
, overpressurization when the transient is limited to either. (1) the start of an
- ' idle RCP with the secondary water temperature of the steam generator less than or equal to 50*F above the RCS cold leg temperatures or (2) the start of 3 charging pumps and their injection into a water solid RCS.
l 3/4.4.11 STRUCTURAL INTEGRITY
.The inservice inspection and testing programs for ASME Code Class 1, 2 and 3 components ensure that- the struct' u ral integrity and operational ~ readiness of these components will be maintained at an acceptable level throughout the life of the plant. These programs are in accordance with Section XI of the ASME
~
- . Boiler and Pressure Vessel Code and applicable Addenda-as required by 10CFR Part 50.55a(g) except where specific written relief has been granted by the 4 Commission pursuant to 10CFR Part 50.55a(g)(6)(i).
E 3/4.4.12 REACTOR VESSEL HEAD VENTS The OPERABILITY of the Reactor Head Vent System ensures that adequate core L cooling can be maintained in the event of the accumulation of non-condensable gases in the reactor vessel. This system is in.accordance with 10CFR50.44(c)(3)(iii). FARLEY-UNIT 1 B 3/4 4-14 AMENDMENT NO. l l L _
1
.. o..
9 ATTACHMENT 2 Response to NRC Letter 4 dated May 2,1985
" Response to NRC Comnents on Farley Unit 2", ALA-85-706, July 1985 i -
4 P e r l
's I
i WEFTINGHOUSE RES90NSF TO NRC COMMEhTS ON FAP1EY UNIT 2 : Reference 1 contained NRC ccaments on the Westinghouse analysis of Farley Unit 2. The Westinghouse anlaysis is contained in Attachment 2 of Reference 2. These concents and associated Westinghouse answers are given in the following sections: 1.0 r - nts on No. 1 The licensee's consultant has indicated that the moment arm of the bolt force about the center of gravity of the flarge bearing pressure diagrar was measured to the outer edge 'for reasons having to do with the finite element modeling. In a real reactor vessel the bearing pressure will be distributed over the mating surface (core barrel to flange and head to flange), which will result in a greater moment arm at the flange junctions than that calculated by the finite element method used by the licensee's consultant. The licensee rust provide an analysis that accounts for this larger moment arm, which results troc a realistic distribution of the bearing pressure. 2.0 Aaswer to Cnment No. 1 The finite element model was changed to account for the fact that the bearing pressure will be distributed over the mating surface (core bearing to flange and head to flange). This was done by coupling nodes 902/233, 903/234, and 904/235 shown by Figure 1. This duplicates the bearing pressure load distributien. Proof that enough nodes are coupled is shown by Table 1. Table 1 shows that the uncoupled nodes 900/231 and 901/232 move apart when the primary load (boltup plus 621 psig pressure) is applied. This is proof that nodes 900/231 and 901/232 should not be coupled because there is nothing to prevent such separation in the real reactor vessel. As a result, Westinghouse has now provided the more realistic finite element model which was requested. 30 Cnment No. 2 b j The finite element stress analysis nost account for the stress l concentration effect of fillets at the flange junctions. The licensee i l
sust cxplcin how the finite element analysis d2termined the sffect of fillets on the localized stresser, concentration. Indicate the peak stress values adjacent to the closure flange fillets. 4.0 -Answer to C - nt No. 2 The location in the flange region with the largest stresses is cross sect'i'on 3 shown by Figure 2. There were no fillets put into the finite element model in this region. Therefore, the stresses generated by this finite. element model are conservative with respect to stresses which would have"resulted if the actual fillet had been included.
, The fracture mechanics analysis which is presented in Section 8 of this report is based on the stress profile of,the boltup plus 621 psig pressure condition shown by Figure 3 and tabulated in Table 2. The
- inside and outside surface stresses frem Table 2 are
c 3 = -14.32 ksi (1) co= 22.74 ksi A complete explanation of how this stress profile was developed is contained in Section 6. The peak stress at cross section 3 is shown on Figure 4. t Specifically, footnote no. 5 of Table NB-3217-1 of the ASME code, E33 Section III states that the equivalent linear stress profile produces the same bending moment on the cross section as the actual stress profile imposes. This principle was applied to produce the equivalent linear stress profile shown by Figure 4, and the linearized inside and outside surface stresses are: (c1)g = -12.58 ksi (2) (co)g = 17.02 ksi l
Th2 peak str:sses ar2 d2termined by cubtracting the linerrized gurfcca stresses in Equation (2) from the stresses in Equation (1). As a result, the peak stresses are: (eg)p = e g - (e g)g = -1.74 ksi ! (3) i (co)p = c o - (o g)g = 5.72 ksi where c1 = -14.32 ksi from Equation (1) e ,= 22.74 ksi from Equation (1) (cg)g: -12.58 ksi from Equation (2) (c,)L= 17.02 ksi from Equation (2) 5.0 Cement No. 9 Describe the dimensional analysis that was performed to determine that the finite element stress analysis performed for the Coranche Peak versels will be conservative for the Farley 2 vessel. 6.0 Anwer to Cmrent No. 9 The dimensional analysis is described in detail in this section. The purpose of this analysis is to show that the typical 4-loop plant finite element model stress results from Attachment 2 to Reference 2 can be used to generate stresses which are applicable to the Farley ! Unit 2 plant (a 3-loop plant). The results from Attachment 2 of Reference 2 show that the largest stresses occur at cross section 3 in the axial direction. Therefore, this dimensional analysis only considers stresses at cross section 3 in the axial direction. At l cross section 3 of Figure 2, the critical dimensions for the typical 4-Ioop plant are: i a l 3 = inner radius = 85.60 in. b3 = outer radius = 96.35 in. (4) tg = thickness = 10.75 in.
l R2farance 4 yisids the following equation fer' longitudinal ctress: 2 a c =P I f = 3.746 P (5) I b 2 ,,2
]
At the same location (through locations 3 and 4 in Reference 5) in the Farley Unit 2 vessel the critical dimensions are: ap = inner ~ radius = 77.938 in. by = outer radius = 87.063 in. (6) t2 = thickness = 9.125 in. The resulting longitudinal stress is: t
, 2
=P
,2
= 4.034 P .
(7) e'2 b 2 ,,2 The constant C that the 4 1oop axial stress should be multiplied by to make it applicable to the 3-loop Farley Unit 2 plant is: C= = 1.077 6 The bloop pressure stresses for an internal pressure of 621 psig (20 percent of the preservice hydrostatic test pressure) have been multiplied by C = 1.077 to obtain the pressure stresses applicable te Farley Unit 2. Figure 5 shows the resultant pressure stress profile for Farley Unit 2. Now, compute the stresses at cross section 3 in the axial direction l due to the boltup load contribution which are applicable to Farley Unit 2. First, a dimensional analysis is done to detencine the bending stress due to the boltup load contributions that is applicable to Farley Unit 2. The bending stress S' in a cylinder can be expressed as follows from Reference 4 6M s' = - (9) t
whera S' s meridional bending stress M, = bending moment t = thickness Also3 Reference 4 yields M
'R
- 2 (10) 2.DA M
and 0 e= AD (11) where tR = radial displacement e = change in slope 3 D= Et 2 12 (1-v ) 4 p (1 v2) 2 R t v = Poisson's Ratio R = mean radius of curvature of wall norral to meridian t = thickness E = Young's Modulus Combining Equation (10) and (11) yields: 0 3 = e>D = (12) 2 201 2A For the 4-loop plant: s t 3
= = 12.17 6 3 (13) 2q i
L
whera 63 = change in clope 4 2 3(1-v ) = 0.04110/in. AI= R2g 2 v =.3 R) = 1/2 (a) + b 3) = 90.975 in. aj = 85.60 in. b) = 96.35 in. t3 = 10.75 in. For Farley Unit 2: .
'2 g,L =
2x
= 10.67 s p (14) 2 where 4
3 2) = 0.04685/ir..
>p =
y = .3 Rp = 1/2 (a p + bp) = 82.500 in. ap= 77.938 in. bp= 87.063 in.
' t 2= 9.125 in.
The following relationship exists between op and 4p:
' R L *#
R [ g ] = 1 103 g (15) R) 2
whera R$ a 90.975 in. R2 = 82.500 in. Therefore, from Equations (13), (14), and (15): og = 12.17 e3 = 1.103 tg = (1.103) (10.67 e2 }
= 11.77 s p (16)
Solve for e2 as a function of 6 3, auch that: 12.17e) ep= 11.77 ' 1.034 e 3 (17) Using Equations (9) and (11), the relationship between the bending
' stress Sj for the 4-loop plant geometry and Sj for Farley Unit 2 can be written as follows:
6Mg ) [e)A)D) e) A) D) h 2 S' i t I / 1 2
\ t 2
I j gr = = = = 1.000 (18) 2 [ 6Mo2 62 A2 D2 /1.034 ejp AD 2 t tf tf k tf 4 , 2 where A) = ~"2 ) = 0.04110/in. E l t) 3 Et I 6 D) = 2
= 3.413 x 10 in-kip 12(1-v )
R) = 90.975 in. t) = 10.75 in. v = .3 3 E = 30 x 10 ksi = 3 x 104 ksi A2= I " ) = 0.04685 /in. Rft 2 3 Et 6 D2= 2
= 2.087 x 10 in-kip 12(1- v )
R2 = 82.500 in. t2 = 9.125 in.
.-~ _-. .__
Sinca S3 ' = Sj (as shown in Equation (18)) the same boltup stresses apply to Farley Unit 2 as did apply to the 4-loop plant gecmetry, and the resultant stress profile is shown in Figure 6. The stress profile of the boltup plus 621 psig pressure condition is shown in Figure 7. The same new size is asstaied in the fracture analysis in this report as was anstaned in Attachment 2 of Reference 2. This' flaw is a 0.625 inch deep surface flaw with an aspect ratio of 1:6. Figure 8 is froic the ASME Code Section XI, Appendix A E03, and it shows the procedure.used to make a linearized representation of stresses. Figure 7 shows that the second stress value from the outer edge occurs at an a/t of .9221 or .711 inch from the outer surface. As a result, a straight line can be drawn through the two stress
, values in Figure 7 which encompass the 0.625 inch deep surface flaw.
This straight line is shown as a dashed line in Figure 7, and it is the appropriate linearized representrtion of the boltup plus pressure stresses for the region of the asmmied flaw. Figure 7 shows that the inside surface stress c1 and outside surface stress e of the linearized pressure stress profile are: c 3 = -59.930 k:1 (19) e ,= 22.740 ksi According to their ASME Code Section XI Ib3 definitions, the Primary l membrane me and bending b stresses are computed as follows: 1 c, = 7 (c; + c ,) = -18.595 ksi (20) b(inside)=f(eg - c ) o= -41.335 ksi 1 b ( utside) = 7 (eg - e g) = 41 335 ksi Using the formula from Reference 6. the Primary Ky (outside) is: l K = (c M +cy) = 25.68 ksi 6 (21) l l l
~'.. '.. wher2 c ,a -18.595 ksi b= 1.335 ksi M, = 1.1 for a/t .068 and a/L = .167 using Figure A-3300.-3 of Reference 6 4=.98 fora /t.=.068anda/L=.167usingFigureA-3300-5 of Reference 6 t = 9.125 in. . a = .625 in. e +e Q = 1.24046 .212 ( m b)2 = 1.197(Formula from Ref.10)
- t
,y e
y = 50 ksi . Now, compute the thermal stresses at cross section 3 in the axial direction for Farley Unit 2. First, compute the therical stresses by hand calculations for both 3-loop and 4-loop plant geometries. Then, the ratio of the stresses cceputed by hand calculations are multiplied by the stresses computed by the 4-loop finite element analysis to yle.Lc stresses which are applicable to Farley Unit 2. The heatup transient considered in the hand-calculated thermal I stresses has a 100 F/ hour heatup rate from 70 F to 279 F. The parameters used in the analysis are listed as follows: 2 h = 660.6 Btu /ft -hr- F = 0.07646 Btu / min-in2,cp k = 30.46 Btu /hr-ft- F = 7.05 x 10~4 Btu /sec-in- F 2 c = I b p
= 0.02128 in /Sec- (22)
Ea = 191.4 psi / F = 0.1914 ksi/ F v = 0.3
. e a time required to hestup from 70 F to 279"F = 7524 sec. E = 2.938 x 107 psi = 2 938 x ION ksi a = 6.516 x 104 in/in/ F e = 0.2836 lb/in3 Cp= 0.1168 Btu /lb- F . Using the formulas from Reference 7, the heatup thermal stresess for the bloop plant are: ,
=-
Ea(tTf ) (eg ) ),y (N$ ) = -13.029 ksi . Ea(ATf ) (23)
- . (co) =
),y (Ng ) = 7.029 ksi 0 D where LTp = 279 F - 70 F = 209e r (N1 ); = 0.228 for
~
1 = 0.05152 and A =t h = 1.385 j from FiEure A.3-5 of Reference 7
=
(N)3=0.123forf2 g 0.05152 and A = = 1 385 1 2 t j from Figure A.3-6 of Beterence 7 tg = 10.75 in. The heatup therral stresses for Farley Unit 2 are: Ea(aT )
=-
(eg)2 ),y f (Ng ) =
-9.715 ksi 2
Ea (tTf) (24) (Co)2 I-" ("O}2 = 5.372 ksi
.O
i where (N3 )p = 0.170 for = 0.06069 and A = '8 = 1.923 2 froc Figure A.3-5 of Reference 7. tg (N,)g=0.094foryk = 0.06069 and A = = 1 923 3 from Figure A.3-6 of teference 7. tg 2 = 9.125 in.
, t The ratios o? the heatup stresses computed by hand calculations are cocputed as follows:
(, ) K 2 g={ )
= 0.7456
- 1 (co ) .
(25) 2 K = 0.7643 U = (co ) 1 ubere K$ = ratio for inside surface stress
>:c = ratio for outside surface stress At cross section 3 in the 4-loop plant the heatup thermal stresses free Table 1 of Attachment 2 of Reference 2 are:
o g = -8.96 ksi (26) c o = 3.70 ksi For Farley Unit 2 the heatup therral stresses are: o g = -8.96 Kg = -6.68 ksi (27: c o = 3.70 K,= 2.83 ksi
~
According to their ASME Cod 2 Section XII definitions, the Secondary membrane and bending stresses during heatup are: c ' = f(eg + co) = -1 92 ksi e' (inside)=f(og o ) = 4.76 ksi
- (28) o{ (outside) = - f(eg on) = 4.76 ksi Using the formula from Reference 6, the Secondary Kg (outside) for the heatup transient is:
K ' It
- I b"m+ "ht) = 5.87 ksi/Tii (29) where c ' = 0 ksia -
c' = 4.76 ksi M g = 1.1 for a/t = 0.068 and a/t = 0.167 using Figure A-3300-5 of Reference 6 P b = 0.98 for a/t = 0.068 and a/l = 0.167 using Figure A-3300-3 of Reference 6 t = 9.125 in, a = 0.625 in. Q
= 1.24046 - 0.212 C
( m + 'b)2= 1.239(Formula from Ref.10) y c y = 50 ksi
c, = 0 ksi is used to avoid using a negative Secondary c' in the calculation.
I" - The cooldown transient con;id: red in the hand-calculcted thermal atresses has a 100 F/ hour cooldown rate from 557*F to 120The F. parameters used in the analysis are listed as follows: 2 h = 660.6 Btu /ft -hr- F = 0.07646 Btu / min-in2,op k = 28.85 Btu /hr-ft- F = 6.678 x 10-N Btu /sec-in- F c = 2
= 0.01916 in /sec Ea = 206.0 psi / F = 0.2060 ksi/ F v = 0.3 (30) e = 15,732 see. .
E = 2.856 x 107 psi = 2.856 x 10 # ksi o = 7.214 x 10-6 in/in/ F
- = 0.2836 lb/ir.3 4 C:
p 0.1229 Btu /lb UF Using the forr::ulas from Reference 7, the cooldown thermal stesses for the 4-loop plant are: Ea(:Tf ) (es ) =- 3, (Nj ) = 15.561 ksi (31) l Ea(4Tf ) (co) - j,, (Ng ) = -9.002 ksi d
whera Tfa 557 F - 120 F a 437 F (Ng )) = 0.121 for - 1 = 0.04875 and A = $ = 2.608
' t) froxc Figure A.3-5 of Reference 7.
( v 1 = 0.070 for g * . 0.o4875 - A. g = 2. 08 1 t) freet Figure A.3-6 of Reference 7. t3 = 10.75 in. The cooldown thermal stresses for Farley Unit 2 are: Ea( Tf ) (eg)p = (Ng )2 = 11.960 ksi Ea( TI ) (eg)g = (No )g = -6.430 ksi where (Ng )p = 0.170 for = 0.05743 and 1 = g = 3.620 2 froc Figure A.3-5 of Reference 7. 12k ca (N,)2 = 0.094 for ht 2 = 0.05743 and A = t-P = 3.620 2 from Figure A.3-6 of Reference 7. t2 = 9.125 in. The ratios of the cooldown stresses cosnputed by hand calculations are: (ej)2 kg= )
= 0.7686 1
(co) (33) k, = g , ) 2 = 0.7143
)
where Kg = ratio for inside surface stress X, = ratio for outside surface stress
At cro:o sectirn 3 in the 4-loop plant the cooldown tiermal stresses frw. Table 3 of Attachment 2 of Reference 2 are: g = 20.26 ksi (34) g = -7.79 kai For Farley Unit 2 the cooldowr. thermal stresses are: o g = 20.26 Kg = 15.57 kai (35) co = -7.79 K,= -5.56 ksi According to their ASME Code Section XII0 definitions, the Secondary me=brane and bending stresses during cooldown are: c'=f(eg + e )g = 5.00 ksi of(inside)=f(eg - o ) g= 10.56 ksi (36) of(outside)=-f(eg - e ) g= -10.56 ksi Using the fomula froc Reference 5, the Secondary K3 (outside) for the cooldown transient is: i K yg = (e,M, + e/g) [ = 6.93 ksie'in- (37) where e' = 5.00 ksi - c' = 0 ksi * , e e' = 0 ksi is used to avoid using a negative Secondary c[ in the calculation.
M,a 1.1 fcr a/t a 0.068 and c4 a 0.167 using Figure A-3300-3 of Reference 6 Pg = 0.98 for a/t = 0.068 and a4 = 0.167 using Figure A-3300-5 of Reference 6 t = 9.125 in. a = 0.625 in. Q = 1.24046 - 0.2'12( n. + b) =2 1.238 (Fonnula from Ref.10)'
,y cy= 50 ksi 7.0 c - nt No. 4 -
In order to demonstrate that the beltline region is more lir.iting than the flange region, indicate the minimum metal temperature in the flange region which results from the fracture analysis. During a heatup and cooldown, what is the required minimum water temperature to ensure that the temperature at the limiting flange location will be equal to or greater than the required minimum inetal temperature? 8.0 Anwer to Cent Mc. 14 This section demonstrates that the beltline region is more limiting than cross section 3 shown by Figure 2. To do this, the required r.inistra metal temperature T at cross section J during heatup of the 4-loop plant is first deterv.ined by using the follwing equation from Section III of the ASME Code, Appendix ,GE03:
) [Kgp - 26.78 }-
T"(0.0145) \ 1.223 / -160 + RTNDT = 121.7'F (38) L
,t . '.
whera K IR e2K yp + kit a 57.23 k21 5i Kyp = 25.68 ksi/In from Equation (21) t
, kit = 5.87 ksiUIi from Equation (29) a c
RTET= 60 p Next, determine the temperature lag in Farley Unit 2 during normal heatup. A two dimensional finite element model for a typical bloop reactor vessel closure head flange and vessel flange geometry was used in the 4 analysis. The WECANE93 finite element program ws used to develop the model. This finite element model was used to obtain temperature gradients caused by the heatup transient. Two-dimensional axisynnetric elements were u ed to model the closure flange regions of the reactor vessel. Four node isoparametric 4 elements were used for all the four node elements. All exterior surfaces of the model were asstaned to be perfectly insulated, and therefore, adiabatic. Figure 9 shows the therral boundary conditions. When the inside surface of the vessel is subjected to thennal transients, the primary mechanism of heat transfer is forced { convection. The thennal properties of the metal are computed as linear 5 functions of temperature. A unironn film coefficient was assumed for the entire inside surface of the vessel. Since the thermal resistance across the flange mating surfaces will not be significant, all the nodes on the flange sating surfaces were thenna11y coupled on the finite element model. The fin'ite element model results show that the temperature lag through cross section 3 in the bloop plant is 51.2'F during the 60*F/ hour heatup transient. Since the Farley Unit 2 has a thinner wall t.hickness of 9.125 in. versus 10 75 in, for a 4-loop plant, the Farley Unit P plant will have less temperature lag. The decrease in temperature lag is considered to be directly proportional to the thickness, so that for Farley Unit 2 tne temperature lag t.T p is: t l
. _ _ , _ . - - ~ . . -
aTp=(t -) AT) = 43.5'F (39) where LTy = bloop plant temperature lag = 51.20 F ty = 41oop plant wall thickness = 10 75 in. tg = Farley Unit 2 wall thickness = 9.125 in. Now, compute the required minimum water temperature during heatup to ensure that the outside aur. face temperature at the critical . location will be greater than the required minimum metal temperature. The required mininnz water temperature yT is: s Ty = T + AT2 = 165.2 F (40) where T = 121.7 F froc Equation (38) LTp : 43.5 F from Equation (30) Figure 10 depicts the Appendix G heatup curve for Farley Unit 2. It can be seen that the 621 psig and 165.2 F data point is less limiting than the heatup curve in FIEure 10. Now, compute the required minimum water temperature during cooldowr to ensure that the temperature at the critical location will be greater than the required minimum metal temperature. Durire cooldown, the metal temperature lags the water teoperature (i.e., the metal temperature is higher than the water teeperature). No credit will be taken for the temperature lag in the cooldown analycir, and therefore the resdits of the calculation for cooldown will be conservative. The required minimum water temperatureyT is: Ty = T = 121.7 F (41) where T = 121.7 F froc Equation (38).
Figura 11 d:picts the Appendix G cooldown curv2 for Farley Unit 2. It can be seen that the C21 psig and 121.7% data point is less limiting than the cooldown curve in Figure II.
' Based on all the above, this fracture analysis has shown that the beltline region is scre limiting than the most critical flange region.
9.0 At$chment Rererences
- 1. Varga, S. A., " Pressure-Temperature Limit Calculations for Joseph M. Farley Nuclear Plant , Unit No. 2", Docket No. 50-364, United States Nuclear Regulatory Commission, Washington, D.C., May 2, 1985.
, 2. hcDonald, R. P., " Joseph M. Farley Nuclear Plant - Unit 2 Response to NRC Questions - Reactor Vessel Surveillance Capsule Report and Associated Technical Specification Change Requests", Alabama Power Company, Birmingham, Alabama, June 18, 1984
- 3. AsME soiler and Pressure vessel Code, Section III, Division 1 -
Subsection NB, nRules for Construction of Nuclear Power Plant components", p. 56, 1983 Edition.
- 4. Roark, R. J., "Fory:ulas for Stress and Strain", 4th Edition, McGraw-Hill Book Company, New York, N.Y., pp. 302, 308, 1965.
- 5. Watson, T. C. , et. al., " Analytical Report for Alabama Power and Light Company J. M. Farley Station Unit No. 2 Reactor Vessel",
Combustien Er.gineering, p. A-37, May 1974
- 6. ASME Boiler and Pressure Vessel Code, Section XI, Division 1 -
Appendix A, " Analysis of FInw Indications", pp. 446, 449, 451, 1983 Edition.
- 7. " Tentative Structural Design Basis for Reactc- Pressure Vessels and Directly Associated Components (Pressurized, Water Cooled Systees), U.S. Department of Cecraerce, Decerber 1,1958 and February 27, pp. 58, 59, 60 Addendum No.1, February 27. 1959.
, i
- 8. ASME Code, Section III, Appendix G, CProtection Against Non-ductile Failure",1983 Edition.
- 9. WECAN Westinghouse Dectric Computer Analysis User's Manual, Westinghouse R&D Center, Pittsburgh, Pennsylvania, September 17, 1979.
10.' *PVRC Reconnendations on Toughness Requirements for Ferritic Materials", WRC Bulletin No.175, Welding Research Council, New York, N.Y., p. 19, August 1972. s i f _ _ , . _ . _ . _ , _ . . - . . _ - _ - - , _ - . _ , _ _ _ - _ .,___,_.-__.. -- _ , _ _ - , - , . . ~. -
1 TABLE 1 SEPARATION OF NODES AT MATING SURFACES FOR BOLTUP PLUS 621 PSIG PRESSURE J fiODES AXIAL SEPARATION BETWEEN NODES (INCH) 231 & 900 0.0041 232 & 901 0.0007 4 233 & 902 0.0000 i nr G
- e TABLE 2 STRESS DATA AT CROSS SECTION 3 i
DISTANCE DISTANCE BOLTUP 621 PSIG BOLTUP & PRESSURE X/T X (INCH) STRESSES PRESSURE STRESS (KSI) (KSI) STRESS (KSI) 0.0 0.00 -16.35 2.03 -14.32 0.0121 0.1298 -15.66 2.01
-13.65 0.0342 0.3671 -14.39 1.97 -12.42 0.0627 0.6745 -13.03 1.90 -11.13 0.0966 1.038 -11.09 1.84 -9.25
. 0.1350 1.451 -9.19 1.78 -7.41 0.1775 1.908 -8.01 1.82 -6.19 0.2236 2.404 -6.74 1.87 -4.87 0.2732 2.937 -5.54 1,91 -3.63 0.3260 3.505 -4.27 1.99 -2.28 0.3818 4.105 -3.15 2.06 -1.09 0.4405 4.736 -1.85 2.15 0.30 0.5019 5.396 -0.63 2.27 1,64 0.566 6.084 0.90 2.38 3.28 0.6325 6.799 2.17 2.47 4.64 0.7015 7.541 4.03 2.64 6.67 0.7728 8.307 5.86 2.83 8.69 0.8463 9.098 8.14 3.00 11.14 0.9221 9.913 12.54 3.76 16.30 1.000 10.75 18.02 4.72 22.74 J
FIGURE 1 FINITE ELEMENT MODEL AT FLANGE MATING SURFACES
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FIG. 8 LINEARIZED REPRESENTATION OF STRESSES O e
FIGURE 9
. THERMAL BOUNDARY CONDITIONS
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+ 0" :l :m '
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. p:If .
l l l I l:. f!O' !,,j ji'iiii:iji ,[Iuli l
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FIGURE 11 - FARLEY UNIT 2 (APR) REACTOR COOLANT SYSTEM COOLDOWN LIMITATIONS APPLICABLE FOR THE FIRST 4.3EFPY
9 W WESTINGHOUSE CLASS 3 CUSTOMER DESIGNATED DISTRIBUTION WC AP-10934 HEATUP AND C00LDOWN LIMIT CURVES FOR THE ALABAMA POWER COMPANY JOSEPH M. FARLEY UNIT 1 REACTOR VESSEL W. T. Kaiser T. V. Congedo S. E. Yanichko T. R. Mager
-September 1985 Approved: -
f t 4N J.M.Chirigos,Managef 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 9119Q:10/091085
- 1. INTRODUCTION Heatup and cooldown limit curves are calculated using the most limiting value of RTNDT (reference nil-ductility temperature). The most limiting RT NDT 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 8"" #5
- I#
NDT' NDT the drop weight nil-ductility transition temperature (TNDT) r he
- temperature at which the material exhibits at least 50 ft lb of impact energy and 35-mil lateral expansion (normal to the major working direction) minus 60*F.
RT NDT increases as the material is exposed to fast-neutron radiation. Thus, to find the most limiting RT a any pe N in h e nac h W e, NDT ART due to the radiation exposure associated with that time period must NDT be added to the original unirradiated RT ** " ** " NDT* RT is enhanced by certain chemical elements (such as copper, nickel and NDT phosphorus) present in reacto'r vessel steels. Westinghouse, other NSSS vendors, the U.S. Nuclear Regulatory Comission and others have developed I trend curves for predicting adjustment of RTNDT ** * ""' " "'"C' "" copper, nickel and/or phosphorus content. The Nuclear Regulatory Commission (NRC) trend curve is published in Regulatory Guide 1.99 (Effects of Residual Elements on Predicting Radiation Damage to Reactor Vessel Materials)II) . Regtilatory 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.II 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 > MeV) divided by 10 . The fluence factor is determined from Figure 1.
i 91190:10/091695
J
- l Given the copper 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-heutron fluence (E > 1 MeV) at the inner surface,1/4T (wall thickness) and 3/4T (wall thickness) vessel locations are given as a function of full-power service in Figure 2. The data for all other ferritic materials in the reactor coolant pressure boundary are examined to ensure that no other component will be limiting with respect to RT NDT"
- 2. FRACTURE TOUGHNESS PROPERTIES The preirradiation fracture-toughness propertics of the Farley Unit 1 reactor vessel materials are presented in Table III. The fracture-toughness properties of the ferritic material in the reactor coolant pressure boundary are detennined in accordance with the NRC Regulatory Standard Review Plani I. The postirradiation fracture-toughness properties of the reactor vessel beltline material were obtained directly from the Farley Unit 1 Vessel Material Surveillance Program.
- 3. FLUENCE CALCULATIONS For the purpose of revising heatup and cooldown curves for Farley Unit 1, whir.'s has limiting embrittlement characteristics in the circumferential girth weld, 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 U'). The SAILOR library is a 47 group, ENDF-8/IY based data set produced specifically for light water reactor applications. In generating the adjoint function, anisotropic scattering was treated with a P expans n 3 e cms-secdons. h e a @ nt source 9119Q:10/091685
location was chosen along the inner diameter of the pressure vessel. This calculation was run in R, 8 geometry to provide a power distribution importance function for the exposure parameter of interest ($ (E > 1 MeV)). Having the adjoint importance function and appropriate core power distributions, the response of interest is calculated as R, 0 " AR IO I(R,0) F(R,0) R dR de where: R R,0
= Response of interest ($ (E > 1.0 MeV), dPa, etc.) at radius R and azimuthal angle 9.
1(R,0) = Adjoint importance function at radius R and azimuthal angle 6. F(R,0) = Full power fission density at radius R and azimuthal angle
~
9. It should be noted that as written in the above equation, the importance function I(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 1 were taken from the following Westinghouse fuel cycle design reports for each operating cycle to date: Fuel Cycle Report 1 WCAP-8515 and Ref. 5 2 Ref. 5 and WCAP-9761 3 WCAP-9761 and WCAP-10036 4A WCAP-10036 and WCAP-10308 5 WCAP-10308 and WCAP-10525 6 WCAP-10525 and WCAP-10795 7 WCAP-10795 and Ref. 6 l 9119Q:10/091685 1
Of'these, Cycles 1 through 4A utilized cut-in fuel loading patterns, and Cycles 5, 6 and 7 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 tho'es at vessel IR through fast flux ratios obtained from the 00T transport analysis performed in support of WCAP-10474, " Analysis of Capsule U f rom the Alabama Power Company, Joseph M. Farley Unit 1 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 O' Lifetime (n/cm ) EFPY Vessel IR Vessel 1/4T Vessel 3/4T 18 1.05 1.943xlb 1.151 x 10 2.674 x 10" 18 1.61 3.526 x 10 2.088 x 10 4.852 x 10" 2.19 4.231 x 10 2.505 x 10 5.822 x 10"
?.98 I
5.738 x 10 3.398 x 10 7.896 x 10" I 3.81 7.041 x 10 4.169 x 10 9.688 x 10" ' 18 4.72 8.296 x 10 4.913 x 10 1.142 x 10 18 18 18 4.79 8.406 x 10 4.978 x 10 1.157 x 10 I 18 32.00 5.037 x 10 2.983 x 10 6.932 x 10 Projection of fluence at 32.0 EFPY was made assuming a power distribution unchanged from that used in Cycle 7 (i.e., that which generated the fluence estimate for 4.79 EFPY). l 9119Q:10/101685 l I
- 4. CRITERIA FOR ALLOWABLE 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 therwal and pressure stresses at any time during heatup and cooldown cannot be greater than the reference stress intensity factor, KIR, f r the metal temperature at that time. K IR is obtained from the reference fracture toughness curve, defined in Appendix G to the ASME CodeI I. The K IR curve is given by the equation:
KIR = 26.78 + 1.223 exp (0.0145 (T-RTNDT + 160)) (1) where K is the reference stress intensity factor as a function of the IR metal temperature T and the metal reference nil ductility temperature RT NDT. Thus, the governing equation of the heatup-cooldown analysis is defined in Appendix G to the ASME CodeI ) as follows: CKgg + kit IK IR (2) where: K gg is the stress intensity factor caused by membrane (pressure) stress K It is the stress intensity factor caused by the thermal gradients.- K IR is a function of temperature relative to the RT NDT f the material 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. l 1 i 91190:10/093085
At any time during the heatup or cooldown transient, K is determined by IR 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, kit, f r the 3 reference flaw are computed. From Equation (2), the pressure stress intensity ! factors'are obtained and, from these, the allowable pressures are calculated. For the calculation of the allowable pressure-versus-coolant temperature i during coo'ldown, the Code reference flaw is assumed to exist at the inside of the vessel wall. During cooldown, the controlling location of the flaw is j always at the inside of the wall because the thermal gradients produce tensile j 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 curve.s are constructed for each cooldown rate of interest. 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.10. This condition, of course, is not true for the steady-state situation. It follows that, at any given reactor coolant temperatu're, the AT developed during cooldown results in a higher value of K g at the 1/4T location for finite cooldown rates than for steady-state operation. Furthermore, if conditions exist such that the increase in K g exceeds Kg , 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. 9119Q:10/100785
Three separate calculations are required to determine the limit curves for 2 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 crack tip lags the coolant temperature; therefore, the K f r the 1/4T IR crack during heatup is lower than the K f r the 1/4T crack during IR 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 's do not offset each IR other, and the pressure-temperature curve based on steady-state conditions no 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
.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. j The second portion of the heatup analysis concerns the calculation of pressure-temperature limitations for the case in which a 1/4T deep outside i 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 I heatup and the time (or coolant temperature) along the heatup ramp. Since the { thermal stresses at the outside are tensile and increase with increasing l heatup rates, each heatup rate must be analyzed on an individual basis. ' i 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 l l three values taken from the curves under consideration. The use of the 4 composite curve is necessary to set conservative heatup limitations because it 1 1 9119Q:lD/100785
-~~ ~ . . -n.. . --- - . - .
. is possible for ccnditiens to exist whsrein, cver the eturse cf 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 4 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. The Farley Unit 2 fracture analysis results from Reference 8 are applicable to Farley Untt 1 since the pertinent parameters are identical for both plants. ! As a result, the 24 EFPY and 32 EFPY heatup curves with and without instrument errors are impacted by the new 10CFR50 rule as shown by Figures 3, 5, 7 and 9, and the 24 EFPY and 32 EFPY cooldown curves without instrument errors in
- Figures 8 and 10 are impacted by the 10CFR50 rule. However, the 24 EFPY and 32 EFPY cooldown curves with instrument errors shown by Figures 4 and 6 are not impacted by the 10CFR50 rule. Since there are many conservatisms (safety factor of 2 on pressure, K toughness and 1/4T flaw) built into the ASME Appendix G analysis method U) ,IRAppendix G does not require that instrument error margins be included in the analysis. Therefore, plant operation can be based on heatup and cooldown curves without instrument errors.
I i An evaluation has been performed to determine the acceptability of the
! Overpressure Mitigation System (OMS) presently in Farley Unit 1 (Technical Specification 3/4.4.10.3) with respect to the 24 EFPY heatup and co,oldown
, 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 of cooldown curves in Figure 8 shows that in the low temperature range (<130'F) cooldown rates of 20*F/hr and lower fall well above 495 psig. Although the cooldown curves for rates of l l 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 l 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. 91190:10/100185 l ---. -
. 5. HEATUP AND C00LDOWN 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 ( }.
Transitkontemperatureshiftsoccurringinthepressure_vesselmaterialsdue 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 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. The leak test limit curve sho in Figures 3, 5, 7 and 9 represent minimum temperature requirements at the leak test pressure specified by applicable codes ( ' .
- 6. AVAILA3LE SURVEILLANCE CAPSULE DATA i
Charpy test specimens f rom Capsule U irradiated to 1.65 x 10 ' n/cm (representative of 10.5 EFPY) indicate that the representative core region 4 weld metal and limiting core region shell plate B6919-1 exhibited maximum shifts in RT NDT of 80*F and 90'F, respectively('}. The weld metal located between the intermediate and lower shell is considered to be the limiting vessel material. The ART prediction for the NDT I material at 1.65 x 10 I' n/cm2 is computed as follows(10) } ART "I NDT " where CF = Chemistry Factor = 112 (from Table I for a weld having a copper content of 0.22 WTY.and nickel content of ( 0.20 WT%) 9119Q:10/101685
FF = Fluence Factor = 1.12 (from Figure 1 at a fluence of , I 1.65 x 10 ' n/cm ) Because the surveillance weld was not fabricated with the same weld wire and lot of flux as the limiting intermediate to lower shell girth weld, the larger chemistry factor resulting for the limiting weld was used to predict ART s. These ART 's which were used to compute the heatup and NDT NDT cooldown curves are more conservative than these obtained from the surveillance capsule results. Although the initial RT NDT f r the welds was co'iservatively estimated (see Table III) as 0*F per Reference 2, a generic value of -56*F was used for the initial RT al ng with a margin as permitted per Reference 10. The margin OT used was as follows: Margin = 2 a2 , ,2 = 66*F where ay = 17'F a = 28'F 3 At the 1/4T location, the adjusted reference temperature ( ART) for the limiting weld in the beltline was detennined by the following expression: ART = Initial RTNDT + ARTNDT + Margin . or ART = -56*F + 137'F (24 EFPY) + 66*F
= 147'F
- 7. SURVEILLANCE CADSULE REMOVAL SCHEDULE The surveillance capsule withdrawal schedule for Unit 1 (Table IV) should repoin the same as identified in the Technical Specifications and WCAP-10474 I9) . The dosimetry analysis of the third capsule to be removed af ter 6 EFPY should be used to re-evaluate the withdrawal schedule for the remaining capsules.
9119Q:lD/101785
r REFERENCES ! (1) Regulatory Guide 1.99, Revision 1 " Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials," U.S. Nuclear l Regulatory Comission, April 1977. (2) " Fracture Tough iess Requirements," Branch Technical Position - MTEB No. 5-2, Chapter 5.3.2 in Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants, LWR Edition, NUREG-0800,1981. (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, Radiation Shield Information Center, Oak Ridge National Laboratory. 4 (5) Radcliffe, R. and Holmes, R., " Revised Nuclear Design Data for Cycle 2 of Farley Unit 1," Westinghouse Nuclear Fuels Division, NDS-79-096, April 25, 1979.
- 3. (6) Erwin, R., "New Power Distributions for Cycle 7," J. M. Farley Unit 1 Project File, Westinghouse Nuclear Fuels Division, August 19, 1985.
(7) ASNE Boiler and Pressure Vessel Code, Section III, Division 1 - Apper. dices, " 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.
(8) Miller, J. C., " Response to NRC Comments on Farley Unit 2," ALA-85-706, ! July 31, 1985. (9) Boggs, R. S., Yanichko, S. E., Cheney, C. A. and Kaiser, W. T. " Analysis of-Capsule U from the Alabama Power Company Joseph M. Farley Unit 1 Reactor Vessel Radiation Surveillance Program," WCAP-10474, February 1984. 91190:1D/100785
I- (10) Regulatory Guide 1.99, Revision 2, " Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials" (Proposed Draft), i U.S. Nuclear Regulatory Comission, June 1984. J j 4 9119Q:10/093085
- - - . - - - - - _-- ...__._y,_ .. _ - . - - - - --_v v
TABLE I
- CHEMISTRY FACTOR FOR WELOS, *F Copper, Nickel, Wt. %
Wt. % 0 0.20 0.40 0.60 0.80 1.00 1.20 0 ' 20 20 20 20 20 20 20 0.01 20 20 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 0.09 40 61 94 115 122 122 122 0.10 44 65 97 122 133 135 135 0.11 49 68 1 01 130 144 148 148 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 92 119 151 184 207 221 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 1 51 180 212 246 276 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 1 51 172 198 228 260 293 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 0.40 175 189 207 231 257 999 320 9119Q:10/093085
j
~
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 0 20 20 20 20 20 20 20 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 19Q 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
- 10/100785
~
TABLE III FARLEY UNIT 1 REACTOR VESSEL TOUGHNESS PROPERTIES Material Cu P N1 T RT Upper Shell Energy NDT NOT ' Component Code No. Type (%) (5). (%) (*F) , (*F)- MWO(C) NMWO(d) Closure head dome B6901 A533,B,C1.1 0.16 0.009 0.50 -30 -20[a] 140 _ Closure head segment 86902-1 A533,B,C1.1 0.17 0.007 0.52 -20 -20[a] 138 - Closure head flange B6915-1 A508, C1.2 0.10 0.012 0.64 60[a] 60[a] 75[a] _ Vessel flange B6913-1 'A508, C1.2 0.17 0.011 0.69 60[a] 60[a] 106[a] _ Inlet nozzle B6917-1 A508, C1.2 - 0.010 0.83 60[a] 60[a] - 110 Inlet nozzle B6917-2 A508, C1.2 - 0.008 0.P3 60[a] 60[a] - 80 Inlet nozzle B6917-3 A508, C1.2 - 0.008 0.87 60[a] 60[a] - 98 Outlet nozzle B6916-1 A508, C1.2 - 0.007 0.77 60[3] 60[a] - 96.5 Outlet nozzle B6916-2 A508, C1.2 - 0.011 'O.78 60[a] 60[a] - 97.5 Outlet nozzle B6916-3 A508, C1.2 - 0.009 0.78 60[a] . 60[a] - 100 Nozzle shell 86914-1 A508 C1.2 - 0.010 0.08 30 30[a] 148 - Inter. shell B6903-2 A533,B,C1.1 0.13 0.011 0.60 0 0 151.5 97 Inter, shell B6903-3 A533,B,C1.1 0.12 0.014 0.56 10 10 134.5 100 Lower shell 86919-1 A533,B,C1.1 0.14 0.015 0.55 -20 15 133 90.5 Lower shell 86919-2 A533,B,C1.1 0.14 0.015 0.56 -10 5 134 97 Bottom head ring B6912-1 A508 C1.2 - 0.010 0.72 10 10[a] 163.5 - Bottom head segment B6906-1 A533,B,C1.1 0.15 0.011 0.52 -30 -30[a] 147 - Bottom head dome B6907-1 A533,B,C1.1 0.17 0.014 0.60- -30 -30[a] 143.5 - Inter. shell long. M1.33 Sub Arc Weld 0.25 0.017 0.21 0[a] 0[a] _ _ weld seam Inter. to lower G1.18 Sub Arc Weld 0.22 0.011 <0.20[b] 0[a] 0[a] _ _ shell weld seams Lower shell long. G1.08 Sub Arc Weld 0.17 0.022 <0.20[b] 0[a] 0[a] _ _ weld seams [a] Estimate per NUREG-0800 "USNRC Standard Review Plan" Branch Technical. Position MTEB 5-2. [b] Estimated. [c] Major working direction. [d] Normal to major working direction 9119Q:10/101785
r -
]
TABLE IV 6 SURVEILLANCE CAPSULE REMOVAL SCHEDULE 1 Lead Estimated Fluence Capsule Factor I9 Removal Time "3 n/cm x 10
~
Y 3.12 Removed (1.13) .583 (Actual) U 3.12 Removed (3.02) 1.65 (Actual) X 3.12 6 3.05 W 2.70 12 5.28 EC3 V 2.70 21 9.28 Z 2.70 Standby - [a] Effective full power years from plant startup / [b] Approximates vessel end of life 1/4 thickness wall location fluence [c] Approximates vessel end of life inner wall location fluence 91190:10/091685
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- 0.22 WT5 l
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W TO 24 EFPY l l l l l 0.4 S.4 904.8 M4.0 500.0 440.e lee.s 180!CATED TERPERATURE (DEG.F) Figure 3 Farley Unit 1 Reactor Coolant ~ System Heatup Limitations Applicable for the First 24 EFPY
MATERIAL PRODERTY BASIS o
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Initial RT ET : -55 F RT ET After 24 EFPY : 1/4T,147'g
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