ML20133F835
| ML20133F835 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 09/30/1985 |
| From: | ALABAMA POWER CO. |
| To: | |
| Shared Package | |
| ML20133F744 | List: |
| References | |
| TAC-59902, NUDOCS 8510110245 | |
| Download: ML20133F835 (44) | |
Text
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Attachment 1 Proposed Changed Pages Unit 2 l l ! Page 3/4 4-29 l- Page 3/4 4-30 l Page B3/4 4-6 Page B3/4 4-7 i Page B3/4 4-8, Page B3/4 4-9 Page B3/4 4-10 Page B3/4 4-10a Page B3/4 4-14 l l i l l l l t i l l l-B510110245 8 364 l PDR ADOCK 0 l P PM _.. .. .,-.____._.,_,.~. - _.._-, _.-_, ._____,___ -
MATE'RTAL PROPERTY BASIS
~
CO EROLLING MATERIAL : R. V. IhTERMIDIATE SHELL COPPER CONTENT : 0.20 WI4 NICKEL CONTEAT : 0.6RWI1 INITIAL RTNDT : -10T RT NDT UTER 9 EPY : 1/47,146% *
- 3/47, 83 7 .,
CURVES APPLICABLE FOR HEAWP RATES UP 10 60T/HR FOR THE SERVICE PERIOD UP TO 9 EPY 3000.0 LEAK TEST LIMIT. / / / ! N ' J ( / 1 ; ; ; UNACCEPTABLE OPERATION [ [ / i f r r i
;;; ee.... 2 2
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= HEAWP RATES UP l I E TO 607/HR g )
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y j E / / l , e g ( / ACCEPTABLE ! g' / OPERATION
- = 2 2
' ( 3
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/
/
/ - CRITICALITY LIMIT
/ BASED ON INSERVICE HYDROSTATIC TEST TEMPERAWRE (274*F)
FOR THE SERVICT TIRIOD UP TO 9 EFPY , l l e.e ' 0.9 900.0 800.9 999.9 400.9 988.0 3RDICATED TERPERATURE (MS.F) i < Figure 3.4-2 i FARLEY UNIT 2 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS APPLICABLE FOR THE FIRST 9 EFPY l UNIT 2 3/4 4-29 AMENDMEMT NO. \ . fan 1EY -- . . .
MATEPTAL PROPERTI BASIS CONTROLLING MATERIAL : R. V. IhTERMEDIATE SHELL
- 0.20 WT5 COPPER C0hTEhT NICKEL COhTEhT : 0.6gWr5 INITIAL RTET : -10 F
- 1/4T, 146 F RT AFTER 9 EFPY '
ET
- 3/4T, 83 F .
, CURVES APPLICABLE FOR COOLDOWN RATES UP TO 100 F/HR FOR THE SERVICE PERIOD UP 10 9 FPY '
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ACCEPTABLE OPERATION
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l e.e e.e tes.e ses.e see.e see.e see.e ISDICATED TERPERATURE (DEC.F) Figore-3.4 -3 FARLEY UNIT 2 REACTOR COOLANT SYSTEM COOLDOWN LIMITATIONS APPLICABLE l FOR THE FIRST 9 EFPY FARLEY - UNIT 2 3/4 4-30 AIENDMENT NO.
REACTOR COOLANT SYSTEM BASES Reducing Tavg to less than 500*F prevents the release of activity should a steam generator tube 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 time 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 for the first full-power service period.
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 liuits to assure prevention of nonductile failure only. For normal operation, other inherent plant characteristics, e.g., pump heat addition and pressurizer heater capacity, may 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 below.
- 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.
FARLEY-UNIT 2 B 3/4 4-6 AMENDMENT NO.
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 9 effective full power years of service life. The 9 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, 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 Charpy test results from Capsule U (from the Alabama Power Company, Joseph M. Farley Unit 2 Reactor Vessel Radiation Surveillance Program) were used to determined the ARTndt due to irradiation eff hese Charpy test specimens from Capsule U irradiated to 5.61 x 10gts. n/c indicate that the representative care region weld metal and limiting core shell plate B7212-1 exhibited maximum shif ts in RTndt of 10*F and 133*F, respectively. FARLEY-UNIT 2 B 3/4 4-7 AMENDMENT NO.
REACTOR COOLANT SYSTEM BASES The weld metal ARTndt of 10*F is well within the ARTndt prediction method from Revision 2 of Regulatory Guide 1.99. However, the shell plate ARTndt of 133*F exceeds the ARTndt )rediction computed as follows: A RTndt = [CF] [FF] = 123*F where CF = Chemistry Factor = 149 (from Revision 2 of Regulatory Guide 1.99 for a copper content of 0.20 WT% and nickel content of 0.60 WT%) FF = Fluence Factor = 0.82 (from Figg1 B 3/4.4-2 at a fluence of 5.61 x 10 d n/cm ) Therefore, the fluence factor is adjusted to reflect the shift of the shell plate surveillance capsule so that the ARTndt's used to compute the heatup and cooldown curves include the surveillance capsule results. These resulting heatup and cooldown limit curves of Figures 3.4-2 and 3.4-3 include predicted adjustments for this shift in RTndt at the end of 9 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, 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 capsule radiation exposure. Allowable pressure-temperature relationships 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. 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 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, A RTndt, corresponding to the end of the period for which heatup and cooldown curves are generated. FARLEY-UNIT 2 B 3/4 4-8 AMENDMENT NO.
Table B 3/4.4-1 FARLEY UNIT 2 REACTOR VESSEL TOUGHNESS DATA Average Upper Shelf Enerov Normal to Principal Principal Werking Working Cu P Ni T RT Direction Comnenent Code No. HDT MDT Direction
$71!Lt M M $ {'f.), l'[1 (ft-1b1 (ft-1b1 CL. HD. Dome 87215-1 A533,B.CL.1 0.17 0.010 CL. HD. Flange 0.49 -30 16(a) g3(a) 128 B7207-1 A508,CL.2 0.14 0.011 0.65 VES. Flange 6C(a) 60(a) >56(a) >B6(C) 67206-1 A508,CL.2 0.10 0.012 0.67 inlet Noz. 60(a) 60(a) >7)(a) > tog 87218-2 A508,CL.-2 -
0.010 0.68 Inlet Noz. 50(8) 50(a) 103(a) 158 B7218-1 A508,CL.2 - 0.010 0.71 32(a) Inlet Noz. 87218-3 A508,CL.2 32(a) tig(a) 172 0.010 0.72 60(a) 60(a) 98ta) Outlet Noz. B7217-1 A508,CL.2 - 0.010 0.73 150
' Outlet Noz. 60(a) 60(a) loo (a) 154 87217-2 A508,CL.2 -
0.010 0.72 6(a) log (a) Outlet Noz. B1217-3 A508,CL.2 6(a) 167 Upper Shell 0.010 0.72 48(a) 4s(a) to3(a) 158 87216-1 A508,CL.2 - 0.010 0.73 30 g7(a) Inter Shell 87203-1 30(a) 149 A533.B.CL.1 0.14 0.010 0.60 -40 15 99 12ter Shell B7212-1 A533,B.CL.1 0.20 0.018 0.60 140 Lower Shell -30 -10 99 134 B7210-1 A533.B.CL.) 0.13 0.010 0.56 Lower Shell -40 18 103 128 B7210-2 A533 B.CL.1 0.14 0.015 0.57 1rans. Ring -30 0 99 145 B7208-1 A508,CL.2 - 0.010 0.73 40 gg(a) Bot. HD. Dome 87214-1 40(a) 137 A533.B.CL.1 0.11 0.007 0.48 -30 -2(a) 87(a) IQter. Shell A1.46 SMAW 0.02 0.009 0.96 134 Long Seams A1.40 0(a) o(a) 313) . SMAW 0.02 0.010 0.93 -60 -60 Inter Shell ~
>106 -
to Lower Shell- 61.50 SAW 0.13 0.016 <.20(b) ' -40 Lower Shell -40 >102 - Long Seams 61.39 SMAW 0.05 0.006 c.20(b) 70 -13 >126 - (a) Estimate per NURE6 0000 'USNRC Standard Review Plan' Branch Technical Position MTEB 5-2. (b) Estimated. (t) upper shelf not available, value represents minimum energy at the* highest test temperature. 1281E:10/082385 FARLEY-UNIT 2 B 3/4 4-9 AMENDMENT NO.
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I _ _ . _ _ _ _ l t b p; #' _ ._ _ _ _ _ _- I .-g ,1 - ! 1 2 3 4 5 6 7891 2 3 4 5 6 7891 2 3 4 5 6 7891 ! P z 10 " 10 10 10 i H z Fluence, n/cm s(E > 1 MeV)
.O FLUENCE FACTOR FOR USE IN . THE EXPRESSION FOR ORTNOT i
I i Figure B 3/4 4-2 , I
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. 4 Finally, the 10 CFR Part 50, Appendix G Rule which addresses the inetal 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 2). In addition, the new 10 CFR Part 50 Rule states that a plant specific fracture evaluation may be performed to justify less limiting l requirements. Based upon such a fracture analysis for Farley Unit 2, the 9 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 performed 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. . 3/4.4.11 STRUCTURAL INTEGRITY The inservice inspection and testing programs for ASME Code Class 1, 2 and 3 components ensure that the structural 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 Commission pursuant to 10CFR Part 50.55a(g)(6)(1). 3/4.4.12 REACTOR VESSEL HEAD VENTS The OPERABILITY of the Reactor Head Vent System ensures that adequate core 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 2 B 3/4 4-14 AMENDMENT NO.
b 1 0 ATTACHMENT 2 Response to NRC Letter dated May 2, 1985
" Response to NRC Comments on
. Farley Unit 2", ALA-85-706, July 1985 F l i i .i i 1 e
WFSTINGHOUSE RFRPONSE TO NRC COMMENTS ON FARLEY UNIT 2 Reference 1 contained NRC comments on the Westinghouse analysis of Farley Unit 2. The Westinghouse anlaysis is contained in Attachment 2 of Reference 2. 'Ihese comments and associated Westinghouse answers are given in'the following sections: 1.0 E=nants on No. 1 The licensee's consultant has indicated that the moment arm of the bolt force about the center of gravity of the flange bearing pressure diagram 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 fla. ge junctions than that calculated by the finite element method used by the licensee's consultant. The licensee must provide an analysis that accounts for this larger moment arm, which results from a realistic distribution of the bearing pressure. 2.0 Answer to C=nant 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 i duplicates the bearing pressure load distribution. Froof 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 ti21 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 c =nant No. 2 The finite element stress analysis must account for the stress concentration effect of fillets at the flange junctions. The licensee
~
must cxplain how the finite element cnalysis dsterinined tha cffect of fillets on the localized stresses 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 section 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 from Table 2 are:
og = -14 32 ksi (1) og= 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. Specifically, footnote no. 5 of Table NB-3217-1 of the ASME Code, Section III [3] 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: (og)g = -12.58 ksi (2) (og)g = 17.02 ksi
The peak ctrccses cra d:;ttrained by subtracting th2 lineirized surfaca stresses in Equation (2) from the stresses in Equation (1). As a result, the peak stresses are: (og)p = o g - (og)g = -1.74 ksi (3) i (oo)p = o o - (oo)g = 5 72 ksi where og = -14.32 ksi from Equation (1) oo= 22.74 ksi from Equation (1) (og)g= -12.58 ksi from Equation (2) (oo)g= 17.02 ksi from Equation (2) 5.0 h nt No. ~4 Describe the dimensional analysis that was performed to detennine that the finite element stress analysis performed for the Comanche Peak ve::elt will be conservative for the Farley 2 vessel. 6.0 Answer to Comnent No. ~4 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 cross section 3 of Figure 2, the critical dimensions for the typical 4-loop plant are: a 3 = inner radius = 85.60 in. bj = outer radius = 96.35 in. (4) tg = thickness = 10.75 in.
~
, R;firenca 4 yicids the fc11owing equaticn fcr longitudin:1 strzss:
,2
=P I o# 2 _,2
= 3.746 P (5) 1 b i
At the same location (through locations 3 and 4 in Reference 5) in the Farley Unit 2 vessel the critical dimensions are: a2 = inner radius = 77.938 in, bp = outer radius = 87.063 in. (6) tg = thickness = 9.125 in. The resulting longitudinal stress is: .-' a
=P 2
= 4.034 P .
(7) b 2,32 o4 2 The constant C that the 4-loop axial stress should be multiplied by to make it applicable to the 3-loop Farley Unit 2 plant is: C = 4.034P3.746P = 1.077 (g) The 4-loop 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 to 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 due to the boltup load contribution which are applicable to Farley Unit 2. First, a dimensional analysis is done to determine 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' = - 2 (9) t
1 ,
'diere S' = meridional bending stress M,= bending moment i t = thickness Also, Reference 4 yields M
OR* 2 (10) 2DA 1 M i and 0 e= AD (11) , l where aR = radial displacement b = change in slope 3
- {' 3 p_ Et 2
12 (1-v ) A* s. ( ., R t v = Poisson's Ratio R = mean radius of curvature of wall normal to meridian t = thickness f E = Young's Modulus Combining Equation (10) and (11) yields: A = (12) R = 2DA 2A For the 4-loop plant: e 1 aR j
* = 12.17 03 (13) 2q
.\
s.
, whera eg = changa in cicpe 2
3(1-v ) = 0.04110/in. A1= R21 t21 v =.3 R) = 1/2 (a) + b g) = 90.975 in. a 3 = 85.60 in. b) = 96.35 in, tg = 10.75 in. For Farley Unit 2: . e
^R 2 2
= 10.67 0 2
IIN) where 2 3(1 v ) = 0.04685/in. A p= R2 t 2 v = .3 Rp = 1/2 (a p + bp) = 82.500 in.
- a2= 77.938 in.
bp= 87.063 in. t p= 9.125 in. The following relationship exists between A and AR 2 R) R A Rg * ^R 2 E R
] = 1.103 23 (15)
I l where R3 = 90.975 in. Rp = 82.500 in.
. Therefore, from Equations (13), (14), and (15):
A Rj = 12.17 ej = 1.103 tg = (1.103) (10.67 e2}
= 11.77 0 2 (16)
Solve for og as a ibnction of e 3, such that: 12.17e j e 2* 11.77 1.034 e 3 (17)
, Using Equations (9) and (11), the relationship between the bending stressSjforthe4-loopplantgeometryandSjforFarleyUnit2can be written as follows:
[ 6Mgj [e jA) Dj [e j A) Dj ) 2 [t I /= ( t 2 I
/ t 2
I j
= 1.000
[S' 2
= = (18)
( 6Mo2) [e2 2 A Dh 2 i/1.034 ej2A D) 2 ( tf / tf / \ tl / 2 where Aj= 3(1-v 2 2
) = 0.04110/in.
R t 3 Et I 6 D j= 2
= 3.413 x 10 in-kip 12(1-v )
Rj = 90.975 in, t) = 10.75 in. v = .3 3 E = 30 x 10 ksi = 3 x 104 ksi A2* (1 " ) = 0.04685 /in. Rft 2 3 Et 6
= 2.087 x 10 in-kip D2= 2 12(1- v )
R2 = 82.500 in. t2 = 9.125 in.
5 0 o Since S$ ' = Sf (as shown in Equation (18)) the same boltup stresses apply to Farley Unit 2 as did apply to the 4-loop plant geometry, 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 flaw size is assumed in the fracture analysis in this report as was asstaned 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 from the ASME Code Section XI, Appendix A I63,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 representation of the boltup plus pressure stresses for the region of the assumed flaw. Figure 7 shows that the inside surface stress og and outside surface stress ao of the linearized pressure stress profile are: i = -59.930 ksi (19) o = 22 740 ksi According to their ASME Code Section XI I63 definitions, the Primary membrane go ard bending b stresses are computed as follows: o, = f (og + c )o = -18.595 ksi (20) 1 b (inside) = 7 (og - c o) = -41 335 ksi
- b(outside)=-f(og - c ) o= 41.335 ksi Using the formula from Reference 6, the Primary Ky (outside) is:
K = (o ,M,+ ob y = .68 ks 6 (20
I l where o ,= -18.595 kai b = 41.335 ksi M ,= 1.1 for a/t .068 and a/L = .167 using Figure A-3300-3 of Reference 6 4 = .98 for a/t = .068 and a/t = .167 using Figure A-3300-5 of Reference 6 t = 9 125 in, a = .625 in. o Q = 1.24046 .212 ( m + ob)2 = 1.197(Formula from Ref.10) y t ey = 50 ksi . Now, compute the thermal stresses at cross section 3 in the axial direction for Farley Unit 2. First, compute the thermal stresses by hand calculations for both 3-loop and 4-loop plant gecinetries. Then, the ratio of the stresses computed by hand calculations are multiplied by the atresses computed by the 4-loop finite element analysis to yielo stresses which are applicable to Farley Unit 2. The heatup transient considered in the hand-calculated therv.a1 stresses has a 100 F/ hour heatup rate from 70 F to 2790F. The parameters used in the analysis are listed as follows: 2 h = 660.6 Btu /ft -hr- F = 0.07646 Btu / min-in2,op k = 30.46 stu/hr-ft- F = 7.05 x 10-N Btu /sec-in- F 2 c = I b p
= 0.02128 in /sec. (22)
Ea = 191.4 psi / F = 0.1914 ksi/ F v =03
, e o time required to he: tup from 70 F to 279 F = 7524 sec. E : 2.938 x 10I psi = 2.938 x 10" ksi a = 6.516 x 10 in/in/ F p = 0.2836 lb/in3 Cp= 0.1168 Btu /lb- F . Using the formulas from Reference 7, the heatup thermal stresess for the 4-loop plant are: Ea(ATf ) (ej ) =- j,y (N) j = -13.029 ksi Ea(aTf ) (23)
- (og) =
(Ng ) = 7.029 ksi j,y where ATf = 279"F - 70"F = 209 F (Ng )$ = 0.228 for 1f2k = 0.05152 and A = 'l = 1.385 t) from Figure A.3-5 of Reference 7 (N,))=0.123forf2k t
= 0.05152 and A =
2 = 1 385 1 t; from Figure A.3-6 of Reference 7 tg = 10 75 in. The heatup thennal stresses for Farley Unit 2 are: (og)2
=- Ea(AT j,y f
9
)(N= )-9.715 ksi 2
Ea (aTf ) (24) (" } I o)2 l'" 2 = 5.372 ksi
where (N3 )2 = 0.170 for = 0.06069 and A = ce 2 = 1.923 from Figure A.3-5 of Reference 7. t 2 k *0 (No )g = 0.094 for = 0.06069 and A = = 1 923 2 fromFigureA.3-6ofkererence7. t 2 t 2 = 9.125 in. The ratios of the heatup stresses computed by hand calculations are computed as follows: (, ) 2 Kg= = 0.7456 1 (o) o . (E) 2 , K = = 0 7643 (o) o1 where Kg = ratio for inside surface stress K,= ratio for outside surface stress At cross section 3 in the 4-loop plant the heatup thermal stresses from Table 1 of Attachment 2 of Reference 2 are: og = -8.96 ksi (26) oo = 3.70 ksi For Farley Unit 2 the heatup thermal stresses are: o g = -8.96 Kg = -6.68 ksi (27) oo = 3.70 Ko= 2.83 ksi
, According to their ASME Code Section XIE03 definitions, the Secondary membrane and bending stresses during heatup are: o ' = f(o g + on ) = -1 92 ksi e' (inside)={(og-o,)=-4.76ksi (28) o{(outside)=-f(og-o)=4.76ksin Using the formula from Reference 6, the Secondary K; (outside) for the heatup transient is: kit *(m"m+"bM) b = 5.87 ksi G (29) i where og = 0 ksia o' = 4.76 ksi M ,= 1.1 for a/t = 0.068 and a/t = 0.167 using Figure A-3300-5 of Reference 6 Mb = 0 98 for a/t = 0.068 and a/t = 0.167 using Figure A-3300-3 of Reference 6 t = 9.125 in, a = 0.625 in. Q
= 1.24046 - 0.212 ( m + 'b)2= 1.239(Formula from 3er.10) y o
y = 50 ksi 8 o, = 0 ksi is used to avoid using a negative Secondary o' in the calculation.
The cooldown tranzient consid: red in the hand-calculcted thermal stresses has a 100 F/ hour cooldown rate from 557 F to 120 F. The parameters used in the analysis are listed as follows: 2 h = 660.6 Btu /ft -hr- F = 0.07646 Btu / min-in2,og 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 =03 (30) e = 15,732 sec. .
E = 2.856 x 107 psi = 2.856 x 104 ksi a = 7.214 x 10~0 in/in/ 0F o = 0.2836 lb/in3 cp= 0.1229 Btu /lb- F Using the formulas from Reference 7, the cooldown thermal stesses for the 4-loop plant are: Ea(ATf )
=-
(og )1 j, (Ng ) = 15.561 ksi 1 (31) Ea(ATf ) (og ) = 3,y (Ng ) = -9.002 ksi
, whera Tfa 557 F - 120*F o 437 F (Ng )j = 0.121 for = 0.04875 and A = = 2.608 froan Figure A.3-5 of Reference 7.
(No )3 = 0. M0 for ht = 0.04875 and A= = 2.608 from Figure A.3-6 of Reference 7. tj = 10.75 in. The cooldown thermal stresses for Farley Unit 2 are: Ea(AT) f (0 12*- 3 1.y (N3 )2 = 11.%0 ksi Ea(AT) (32) f (co)2
- 1.y (No )2 = -6.430 ksi where (N)g=0.170forff=0.05743andA=
3 2
$=3.620 t
2 from Figure A.3-5 of Reference 7. (No )2 = 0.094 for = 0.05743 and A= = 3.620 2 from Figure A.3-6 of Reference 7. t2 = 9.125 in. The ratios of the cooldown stresses computed by hand calculations are: (ej)2 l K j = (, y = 0.7686 (oo ) (33)
= 2 K
g (, ) = 0.7143 l where Kg = ratio for inside surface stress Ko a ratio for outside surface stress
- d
At cross section 3 in the 4-loop plant the cooldown therinal stresses from Table 3 of Attachment 2 of Reference 2 are:
g a 20.26 kai (34)
,3 = -7.79 kai For Farley Unit 2 the cooldown thermal stresses are:
og : 20.26 Kga 15.57 kai (35) o o = -7.79 K,= -5.56 kai
, According to their ASME Code Section XIE03 definitions, the Secondary membrane and bending stresses during cooldown Jre:
c'=f(og+o)=5.00ksi g of(inside)=f(ej-o)=10.56ksi g (36) of (outside) = - f (og - o g) = -10.56 ksi Using the formula from Reference G, the Secondary Kg (outside) for the cooldown transient is: kit
- I m"m+ch) = 6.93 ksi G (37) where o' = 5.00 ksi o[=0kai*
t 8o'=0kaiisusedtoavoidusinganegativeSecondaryc(inthe calculation.
- M ,a 1.1 for a/t c 0.068 and a/t o 0.167 using Figuro A-3300-3 of Reference 6 43 0.98 for a/t = 0.068 and a/t = 0.167 using Figure A-3300-5 of Reference 6 t a 9.125 in, a = 0.625 in.
Q = 1.24046 - 0.212( *,y+): 1.238 (Formula from Ref.10) cy: 50 ksi 7.0 c-nt No. 4 - In order to demonstrate that the beltline region is more limiting than the flange region, Micate the minimum metal temperature in the flange region v .ts from the fracture analysis. During a heatup and coolde , 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 metal temperature? 8.0 Anmer to r e nt No. 4 1his section demonstrates that the beltline region is more limiting than cross section 3 shown by Figure 2. To do this, the required minisun metal temperature T at cross section J during heatup of the 4-loop plant is first deterinined by using the following equation from Section III of the ASME Code, Appendix OE03:
)
K;p-26.78p T=(0.0145) '" [, \ 1.223 / -160 + RTNDT = 121.7'F (38) t
, wher] K IR 2K yp oK It 57.23 ksi/In K yp = 25.68 kaidii from Equation (21) kit =5.87ksi/fitfromEquation(29)
RTET= 60 F Next, determine the temperature lag in Farley Unit 2 during nonnal heatup. A two dimensional finite element model for a typical 4-loop reactor vessel closure head flange and vessel flange geometry was used in the analysis. The WECAN finite element program ws used to develop the model. This finite element model was used to obtain temperature gradients caused by the heatup transient. hro-dimensional axisymmetric elements were used to model the closure famnge regions of the reactor vessel. Four node iseparametric elements w re used for all the four node elements. All exterior surfaces of the model were assumed to be perfectly insulated, and therefore, adiabatic. Figure 9 shows the therral boundary conditions. When the inside surface of the vessel is subjected to thernal transients, the primary mechanism of heat transfer is forced convection. The thermal properties of the metal are computed as linear functions of temperature. A uniform 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 mating surfaces were therinally coupled on the finite element model. The finite element model results show that the temperature lag through cross section 3 in the ti-loop 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 2 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 the temperature lag AT p is: O
e , l t l ,. ATp(t ) AT, a 43.5SF (39) where ATy = Aloop plant temperature lag = 51.20"F t ya bloop 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 surface temperature at the critical location will be greater than the required minimum metal temperature. The required minimum water temperatureyT is: t Ty : T + AT2 x 165.2 F (40) where T 121.7 F from Equation (38) ATp 43.5"F from Equation (39) 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 Figure 10. Now, compute the required minimum water temperature during cooldcom to ensure that the temperature at the critical location will be greater than the required minimum metal temperature. During cooldown, the metal temperature legs the water temperature (i.e., the metal temperature is higher than the water temperature). No credit will be taken for the temperature lag in the cooldown analysis, and therefore the results of the calculation for cooldown will be conservative. The required minimum water temperatureyT is: Ty:T 121 7 F (41) where T = 1217"F from Equation (38).
'. o Figure 11 depista the Appendix G cooldown curve for Farley Unit 2.
It can be seen that the 621 psig and 121.7"F data point is less limiting than the cooldown curve in Figure 11.
~ Based on all the above, this fracture analysis has shown that the beltline region is more limiting than the most critical flange region.
9.0 At+~'==nt Rareramen
- 1. Verga, S. A., " Pressure-Temperature Limit Calculations for Joseph M. Farley Nuclear Plant Unit No. 2", Docket No. 50-364, United States Nuclear Regulatory Caesnission, Washington, D.C., May 2, 1985.
, 2. Mcdonald, 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 Boiler and Pressure Vessel Code, Section III, Division 1 -
Subsection NB, " Rules for Construction of Nuclear Power Plant
, Components", p. 56, 1983 Edition.
- 4. Roark, R. J., "Fonnulas for Stress and Strain", 4th Edition, i 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",
4 Combustion Engineering, p. A-37, May 1974
- 6. ASME Boiler and Pressure Vessel Code, Section XI, Division 1 -
Appendix A, " Analysis of Flaw Indications", pp. 446, 449, 451, 1983 Edition. l
- 7. aTentative Structural Design Basis for Reactor Pressure Vessels and i Directly Associated Components (Pressurized, Water Cooled Systems), U.S. Department of Commerce, December 1,1958 and l February 27, pp. 58, 59, 60, Addendum No.1, February 27. 1959. l l
i
o ,
- 8. ASME Coda, Section III, Appendix G, CProtection Against Non-ductile Failure",1983 Edition.
- 9. WECAN Westinghouse Electric Computer Analysis User's Manual, Westinghouse R&D Center, Pittsburgh, Pennsylvania, September 17, 1979.
- 10. "PVRC Recommendations on Toughness Requirements for Ferritic Materials", WRC Bulletin No.175, Welding Research Council, New York, N.Y., p. 19, August 1972.
s e I
9 - f' , e e TABLE 1 SEPARATION OF NODES AT MATING SURFACES FOR BOLTUP PLUS 621 PSIG PRESSURE AXIAL SEPARATION BETWEEN N0 DES NODES (INCH) 231 & 900 0.0041 232 & 901 0.0007 233 & 902 0.0000
1 [ ' i TABLE 2 l ,' I STRESS DATA AT CROSS SECTION 3 l 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 l 0.1775 1.908 -8.01 1.82 -6.19 0.2236 2.404 -6.74 1.87 -4.87 l 0.2732 2.937 -5.54 1.91 -3.63
! 0.3260 3.505 -4.27 1.99 -2.28 l 0.3818 4.105 -3.15 2.06 -1.09 i 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 j 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 l i
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l .. ,t FI,GURE4 7 S.TRES.S PROFILE.i FOR..BOL.TU.P P. LUS 621 PSIG PRESSURE 4 $ I ! t , I l - - i
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e e a e Actual nonlineer Actuel nonlineer
~f stress dis-tribution
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N etrees dis-tribution en i r , \ U e hI b N g% T
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I s Equivalent l [ 9 l l 8 I
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I l Eau,,eieniiinea,
, ,, n ,1,q I "'"' l l of stress I distribution ' l distribution 4- t/2 e4 -- t/2 , % t/2 ; e t/2 e 1 8
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, FIGURE 9 THERMAL BOUNDARY COND1TIONS
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ADIABATIC / / x f,
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' Leak Test Livr.it Curve Applicable For !!eatup Rates Up To j li '. l
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60*F/HR, For the Service Period lip To 4.3 . j ,1 l ,li!:Fji-d!i :E/ 7
!lb
- ! EFPY, and Contains Margins 0f 10*F And ll 1! ! . l IIIII l . '..!. l 1
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tlj Controlling Materials - RV Inter Shell i
- , ; 3!,, ::::
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21lw i s j Copper Content, o.20s
, ., Rrug7 Orwnal ,-iov i j i f
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e p, Test Temperature (277*F) ror the ; q!!i;;
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service Period up to 4.3EFPY. l r 7: w
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; i FIGURE 10 FARLEY UNIT 2 (APR) REACTOR COOLANT SYSTEM HEATUP LIMITATIONS FUR TiiE FIRST 4.3EFPY
----------_-_11_ m1
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jill
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- i l Curve Applfcable for the Service N I,j' j , lilll lllj @, !!!! ii bi!!!
l .h l Period up to 4.3EFPY, and Contains '2 2m
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li
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Materials Basis: J:.: Ig nou . . .; !. : : }: m:: , l, , -
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3 i li: lj.l g.h '!ii !r' il i. liig ! ti jjl I I i f I ? ll,I ;l - I j,i'
;W i y, Copper content, 0.20'l; f-g '***
Ti j'j! fi! 6 :'i: ( ji! !l ill O ;I hll '
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! 'iN i I, !NifIi 'O !! Il j! h l ' ' l i ll lil RT aftei4.3EFeY, 1/4T, ia7 r i-m.e!m,i kp ,j[ ,'j j nor !=ds? I si iU; .- i, .l9 fili pn;d i
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3/4T, 57'F
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- lbjl:li di >" I i
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' h l!!!l l !I;i I! in l l l m
l , o! i o 2. : m q l'i u m iqj g y g j q i iii g$j,g!isW 3,, ,
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l I l i . Mim :,, .ii' Mi lE :! 4 qW!,i!iii!!g nI:ditb'WM,k<m.hnihin l i . y e, ll il d
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Ii $ ihip g q:i ,
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ii FIGURE 11 - FARLEY UNIT 2 (APR) REACTOR COOLANT SYSTEM COOLDOWN LIMITATIONS APPLICABLE FOR THE FIRST 4.3EFPY
- - - - ~ - [~ ] _}}