ML20087N443

Reactor Vessel Outlet Nozzle-to-Shell Weld Flaw Indication Fracture Mechanics Evaluation for Point Beach Nuclear Plant, Unit 1

ML20087N443
Person / Time
Site:	Point Beach
Issue date:	03/27/1984
From:	Nair P SOUTHWEST RESEARCH INSTITUTE
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ML20087N436	List: ML20087N435 ML20087N438 ML20087N443
References
TAC-54446, NUDOCS 8404030464
Download: ML20087N443 (43)

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{{#Wiki_filter:__ l 1 REACTOR VESSEL OUTLET N0ZZLE-TO-SHELL WELD FLAW INDICATION FRACTURE MECHANICS EVALUATION FOR POINT BEACH NUCLEAR PLANT, UNIT I SwRI Project 17-7222-130 by P. K. Nair Prepared for: Wisconsin Electric Power Company 231 W. Michigan, PO Box 2046 . Milwaukee, Wisconsin 53201 March 27, 1984 8404030464 840330 PDR ADOCK 05000266. _P_ PDR f

TABLE OF CONTENTS Page

1.0 INTRODUCTION

1 2.0 IECHNICAL APPROACH 2 2.1 ASME BVP Section XI Appendix A 2 2.2 Flaw Characterization 2 3.0 FRACTURE MECHANICS INPUT DATA 4 3.1 Material Data. 4 3.2 Loading and Stress Conditions 10 3.2.1 Level I and II 10 3.2.2 Level III and IV 10 3.3 Stress Intensity Factors 19 4.0 FRACTURE INTEGRITY EVALUATION 30 4.1 Normal and Upset Conditions (I & II) 30 4.2 Emergency and Accident Conditions (III and IV) 33

5.0 CONCLUSION

S 36 REFERENCES 37 i I

LIST OF TABLES

3. age 3.1 Design Transients and Cycles 11 3.2 Stress Data for the Outlet Nozzle 12 3.3 Stress Ranges and Cycles Data 13 4.1 Fatigue Crack Growth - Flaw No. 2 32 4.2 Fatigue Crack Growth - Flaw No. 3 32 m

11 1.:

I 1 .k. LIST OF FIGURES Page '1_ FIGURE 3-1. LONGITUDINAL DISTANCE VERSUS MULTIPLYING FACTOR FOR PEAK FLUENCE 6 ' U-l i FIGURE 3-2. RADIAL FLUENCE DISTRIBTUION (MAXIMUM i FLUENCE, 32 EFPY, 40 CALENDAR YEARS) 7 ' l FIGURE 3-3. (FIG. A-4200-1) LOWER BOUND Kia AND Kie ,,3 TEST DATA FOR SA-533 GRADE B CLASS 1, _ F SA-508 CLASS 2, AND SA-508 CLASS 3 P STEELS 9 ! L ' "r FIGURE 3-4. TEMPERATURE DISTRIBUTION BROUGH THE VESSEL I WALL FOR A LARGE LOSS-OF-COOLANT ACCIDENT 15 FICURE 3-5. THERMAL HOOP STRESS THROUGH THE VESSEL WALL ~ P FOR A LARGE LOSS-OF-COOLANT ACCIDENT 16 - 3 FIGURE 3-6. LARGE STEAM BREAK WITH REACTOR COOLANT PUMPS RUNNING; REACTOR COOLANT SYSTEM -5 PRESSURE VERSUS TIME (SECONDS) 17 P FIGURE 3-7. LARGE STEAM BREAK WITH REACTOR COOLANT PUMPS 'h # RUNNING; COLD LEG TEMPERATURE VERSUS TIME (SECONDS) 18 ? f FIGURE 3-8. TEMPERATURE DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE STEAM LINE BREAK WITH 3. OFF-SITE POWER AVAILABLE 20 ~ 'E ,E FIGURE 3-9. PRESSURE AND IHERMAL HOOP STRESS U L DISTRIBTUION THROUGH THE VESSEL WALL FOR = i A LARGE STEAM LINE BREAK WITH OFF-SITE POWER 2 AVAILABLE 21 'i d FIGURE 3-10. LARGE STEAM BREAK WITH REACTOR COOLANT PUMPS - '= c TRIPPED; REACTOR COOLANT SYSTEM PRESSURE ~- VERSUS TIME (SECONDS) 22 '1 FIGURE 3-11. LARGE STEAM BREAK WITH REACTOR COOLANT PUMPS . ; j TRIPPED; COLD LEG TEMPERATRUE VERSUS TIME

}

(SECONDS) 23 FIGURE 3-12. TEMPERATURE DISTRIBUTION THROUGH THE VESSEL =I WALL FOR A LARGE STEAM LINE BREAK WITHOUT -F OFF-SITE POWER. 24 -:( p , 'r iii

i (f

Li a I

_p LIST OF FIGURES Page FIGURE 3-13. PRESSURE AND 'DiERMAL HOOP STRESS DISTRIBUTION THROUGH 111E VESSEL WALL FOR A LARGE STEAM LINE BREAK WITHOUT OFF-SITE POWER 25 FIGURE 3-14. LOCKED ROTOR PRESSURE TRANSIENT 26 iv

1.0 INTRODUCTION

During the spring of 1984's in-service examination of the Point Beach Udit reactor pressure vessel, several ultrasonic reflectors were detected in the outlet nozzle to shell welds. No of the indications were found to be unacceptable by size limits of Table IWB-3512 of the ASME Section XI Code [1]. The purpose of this document is to provide a detailed fracture mechanics analysis of the marginally unacceptable indications by IWB-3512. Using the alternate acceptance criteria in IWB-3600, the long-term integrity of the vessel will be established.

2 2.0 TECHNICAL APPROACH 2.1 ASME BVP Code Section XI Appendix A The detected flaw indications are analyzed using the linear elastic fracture mechanics method described in Appendix A of the ASME Section XI Code (1977). The analysis involves characterizing flaws in a regular elliptic shape and determining their stability using representative material properties t and applicable loadings. The Appendix A approach provides algorithms to compute stress intensity factors. The tcughness properties are only available for lower-shelf and transition regions. Hence, upper-shelf toughnesses have to be estimated from surveillance and other weld data with similar chemistries. Irradiation degradation estimation procedures are also presented. Flaw acceptance criteria is based on the alternate criteria in IWB-3600 and expressed as follows: Kla/K > /1 for Levels I and II loadings 1 and Kic/g > /2"for Levels III and IV loadings 1 2.2 Flaw Characterization Two major flaw indications are considered. These were individually focad unacceptable by inspection criteria in Table IWB-3512 of Section XI. l The location of these indications are, Flaw #2 at 28.5' vessel azimuth and Flaw #4 at 208.5' vessel azimuth. Associated with Flaw #2 is a satellite flaw indication with about 1% area of -the Flaw #2 area. They are spaced in excess of the larger flaw diameter distance apart. The effect of the small flaw on L Flaw #2 is considered negligible. Flaw #2 will be treated as an independent - flaw for the analysis in this report. The #2 and #4 indications can be assumed to be radial-axial flaws in the vessel of 9.125 inches thick. This l'

3 assumption permits the flaw to be oriented perpendicular to the maximum principle stress direction. The indications are completely circumscribed by ell'iptic area according to the IWA-3300 method. J l l-l l l l l l

l 4 - 3.0 FRACTURE MECHANICS INPUT DATA To erform a fracture mechanics based analysis, several input data are required. These data include, in general, the material properties under i irradiated conditions, loading / transient information and the stress intensity factor computation algorithms. 3.1 Material Data Point Beach. Unit I Raactor Vessel was manufactured by Babcock & Wilcox Co. in accordance with ASME Code Section III 1965 Edition and satisfying Code Case 1332-2. The outlet nozzles (A-508 Class 2) were welded to the shell with Mu-Mo-Ni wire HT #8T1554B and Linde 80 flux lot #8479. Weld Chemistry C Mn P-S Si Cr Ni Mo Cu 0.08 1.58 0.014 0.012 0.45 0.07 0.60 0.40 0.19 Charpy V-Notch at (10*F) 53 ft-lbs 42 ft-lbs 50 ft-lbs Weld Tensile Properties UTS Yield Point % Elongation jg{ 88000 psi 73500 psi 25.0% 64.3% 86500 psi 71500 psi 26.5% 66.2% The production welds were stress relieved at 1100'F-1150* F for approximately 11 hours. RTNDT and Upper-Shelf Charpy V-Notch Preirradiated test data for both RTNDT and Charpy V-notch (upper-shelf) is not available for the outlet nozzle weld. However, data from tha f-

5 surveillance weld for Unit I.is applicable. The vessel welds have comparable chemistry and heat treatment. Weld Chemistry C-Mn P S Si Cr Ni Mo Cu 0.09 1.47 0.019 0.024 .0.49 0.13 0.57 0.39 0.18-0.24 Stress Relief at 1125'F, 11-1/4 hours, furnace cooled. I WCAP-8739 (3) presents the preirradiated and irradiated data for Unit I capsules. For the purposes of the fracture analysis, the surveillance weld data will be used. These include: RTNDT = 25' F preirradiated 'Cy(upper-shelf) = 65 ft lbs preirradiated Fluence Levels at Flaw Locations. 4 Figure 3-1 shows the decay of fluence levels as a function of distance from the active core. The flaw indications are an average distance of 35 inches from the active case. From Figure 3-1, the multiplying factor is obtained as 1.5 x 10-3, Figure 3-2 presents the end of life radial fluence distribution for (E- > 1.0 MeV) in the 6.5 inch thick beltline region. The nozzle region is 9.125 inches thick. Therefore, using Figure 3.2 to represent the outlet nozzle region would be conservative. Since one of the indication crack tips is located at about 0.55t, the end-of-life fluence level is taken to be 1.1 x 1019 (at the beltline) x 1.5 x 10 3 (multiplying-factor for nozzle). Therefore, the maximum fluence level at the flaw locations will be 1.65 x 1016 n/cm,. 2 On the basis of 1.65 x'1016 2 n/cm (E 1.0 MeV) fluence and Cu = 0.20%, s (upper-shelf) P = 0.0.019 there is no irradiation effect on both RTNDT and Cy I ~ ,_m., __-.,_-_-.,f.. ...___r...,,,_,

10 I I l 1 l 1 6 8 6 4 ( 2 wy 10-i 5 8_ g 6 t E 4 i j g N y m-m 2 7 g g d t 2 5 10-2

>

c 'a' 8 C0ag 6 4 4 d r 2 ) io-a I I I l 0 5 10 15 20 25 30 35 l DISTANCE FROM FUEL CORE ASSEMBLIES d (INCHES) l l FIGURE 3-1. LONGITUDINAL DISTANCE VERSUS MULTIPLYING FACTOR FOR PEAK FLUENCE

1 4.0 ^m 3.0 o u "a .$2 u5 3 w $ 2.0 x R A E E a: E E 1.0 O 0 0.1

0. 2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 FRACTIONAL DISTANCE THROUGH WALL (a/t) l l

l FIGURE 3-2. RADIAL FLUENCE DISTRIBUTION (MAXIMUM FLUENCE, 32 EFPY, 40 CALENDAR YEARS)

8 [ Reference Reg. Guide 1.99, April 1977 Rev 1 (2)]. Therefore, the end of life values will be: RTNDT = 25' F C (upper-shelf) = 65 ft lbs y Fracture Toughness Properties For the lower-shelf and transition temperatures, Figure 3-3, i.e., Figure A 4200-1 of ASME Section XI Appendix A, will be used. This provides both the Kic(static) and Kla(arrest) toughness values. At upper-shelf temperatures, both Kla and Kle are assumed to be equal. A very conserative estimate of toughness is obtained using the Barsom and Rolfe correlation from ASTM STP 466 (4). i.e., K 5 CVN ojr Ie = ey oy 20 ey is the yield stress (ksi) where CVN is the Charpy energy (f t lbs) using oy = 72.500 kai (average value) and CVN = 65 ft lbs KIc = KIa = 149 kai /In~ Fatigue Crack Growth Data presented in ASME Section XI Appendix A Figure A-4300-1 (in air environment) will be used. ci g = 2.67 x 10-11 ag3.726 j dN where da is the crack growth per cycle dN and AK is the stress intensity factor range. i

220 goo iso iso k Ky ido .2 } Ky u 120 E 5 100 h .a g so E.t so d' to - nrnor iL 0 a a i -soo -se o +s0 +sco + iso

• 200 tr-avno rl. *F (Fig. A-4200-1) LOWER BOUND K, TEST DATA FOR SA-533 GRADE B FIGURE 3-3.

CLASS 1,SA-508 CLASS 2,ANDSAandK.508ClhSS3 STEELS

10 3.2 Loading and Stress Conditions The various loading and stress conditons experienced under Levels I and II and postulated under Levels III and IV are discussed in this section. 3.2.1 Level I and II Table 3.1 presents the Design Transients and Cycles for the vessel. (5) Stress data for the outlet nozzle is presented in Table 3.2. Table 3.3 develops membrane and bending stress ranges with cycle information for the outlet nozzle. 3.2.2 Level III and IV Three transients are considered in the evaluation. In selecting the transients, the relative magnitudes of the pressure and thermal stress components of the transients were studied. Thermal transients that decrease the coolant temperatures tend to create tensile stresses at the vessel,ID. However, at the flaw locations of interest, compressive stresses are generated. These compressive stresses reduce the total tensile stress magnitude due to pressure at the flaw locations. In selecting the transients for evauation, the higher pressure transients have been evaluated. For the purposes of the present analysis, the transient and thermal stresses temperatures data developed in WCAP-8742 (6) for the 6.5 inch thick vessel wall region will be conservatively assumed to be valid at 'the nozzle section. The temperature and stress distributions are developed for fractional distances of the vessel wall thickness. The transients analyzed below provide a bound on fracture conditions for all other postulated accidents.

11 TABLE 3.1 Design Transients and Cycles DESIGN CTCLIS HOT LEG TDe (*F) FOR PRESSURE CURRENT TO TRANSIENTS TDfE SUTLET N0ZZLE (PSI) AT START UP END OF LIFE 7 Heatup 4.5 hr 547 2250 i Cooldown 0 hr $47 2250 j 100 150 . Plant Loading 20 min 612 2250 Unioading 20 min 547 2250 14500 10a75 small Step Load 220 sec 610 2275 Increase 2000 1500 small Step Load 35 sec 622 Decrease 2320 Large Sten Load 2 min 588 2375 200 150 Decrease Loss of Load 10 sec 658 2775 80 60 Loss of Power .75 hr 634 2550 40* 30 Loss of Flow 0 sec 614 2250 80 60 Reactor r On sec 614 2250 400 300 p Turbine Fall 0 min 614 2250 10 9 8 +fluce. 620 2350 1.0 x 10 2.5 x 10 6 5 Cold Hydro 5 hr 100 1125 5 3 Hot Hydro 5 hr 400 2500 40 30

• ESTIMATED 9

12 TABLE 3.2 Stress Data for the Outlet Nozzle Total Stress Primarv Stresses (ksi) Secondarv Stresses (ksil (ksi) Load 1,oad Condition Stage e e e, b i o m b m b o c Heatup 4.5 ER 33.53 23.43 28.48 5.05 -12.66 7.03 -2.R2 -9.85 25,66 -4.8 Cooldown 0 HR 33.53 23.43 28.48 5.05 0.0 0.0 0.0 0.0 18.48 5.05 Plant loading 20 MIN 33.53 23.43 18.48 5.05 -11.46 4.30 -3.58 -7.88 24.90 -2.83 Unioading 20 MIN 33.53 23.43 28.48 5.05 11.46 -4.30 3.58 7.88 12.06 12.93 Small load 220 SEC 33.88 23.78 28.33 5.05 0.87 -0.20 0.34 0.53 29.17 5.58 step increase Small load 35 SEC 34.50 24.40 24.45 5.05 -1.96 n.13 -0.87 -0.99 28.58 4.06 step dec. 1.arge load 2 MIN 35.27 25.17 30.22 5.05 -2.72 0.45 -1.14 -1.58 29.08 3.47 step dec. 'I 10 SEC 18.80 28.70 31.75 5.05 -10.39 0.42 -4.99 -5.40 28.76 .o,33 .75 HR 37.00 26.90 31.95 5.05 -2.58 1.26 -0.66 -1.92 31.29 3.13 er " f 0 SEC 33.53 23.43 28.48 5.05 0.0 0.0 0.0 0.0 28.48 5.05 Reactor trip from full 0 SEC 33.53 23.43 28.48 5.05 0.0 0.0 0.0 0.0 28.48 5.05 power 1 0 MIN 33.53 23.42 28.48 5.05 0.0 0.0 0.0 0.0 28.48 5.05 6 7 + F11*CT. 34.92 24.82 29.87 5.05 -1.36 0.13 -0.62 -0.74 29.25 4.31 hydro 5 HR 45.65 35.55 40.60 5.05 0.0 0.0 0.0 0.0 40.60 5.05 ",cydra , HR 37.00 26.90 31.95 5.05 _e.55 4.72 -1.42 -6.13 30.33 -1.08

13 TABLE 3.3 Stress Ranges and Cycles Cata CYCLES STRESS RANCES (KSI) TRANSIENT TO NO. LOAD CONDITION END OF LIFE as a: a b 1 Heacup .150 27.07 0.125 Cooldown 2 Plant loading 10.875 1.41 9.975 Unioadina 3. Small load 1,500 1.82 4.82 Increase / Decrease 4 Large step 150 2.01 3.345 Load Decrease 5 Loss of load 60 1.69 -0.475 6 Loss of power 30 4.22 3.005 7 Loss of flow 60 1.41 4.925 8 Reactor trip 300 1.41 4.925 from full power 9 Turbine roll 8 1.41 4.925 10 Steady-state 2.5 x 10' 2.18 4.180 11 Cold hydro 3 13.53 4.925 12 Hot hydro 30 3.44 -1.205 6 _____.___.--_-m-

14 3.2.2.1 Loss of Coolant Accident During LOCA, the reactor vessel rapidly depressurizes and safety injection commences. Initially, water at 90'F enters the vessel via the downconer and 20 seconds later the injection pumps pump fluid from the BAST at a temperature of 155'F. Af ter 53 seconds into the accident pump, suction transfers to the refueling water storage tank and the temperature is 33'F. The injection flow from the safety injection pumps enters the RV downcomer through the inlet nozzle from the cold leg of the piping system. The outlet nozzles do not experience the initial thermal shock as much as the region near the inlet nozzles. Overall, the pressure drops from 2250 psia to approximately 30 psia. In reference (6) WCAP-8747, the resulting LOCA temperature and stress profiles in the vessel wall were developed. These are reproduced here in Figures 3-4 and 3-5. 3.2.2.2 Steam Line Break Accident During a large steam line break, the coolant temperature and pressure decreases rapidly. Safety injection is triggered when the coolant pressure falls below 1560 psia. The injection flow rates depend on the assumptions for reactor coolant system back pressure and for one or two loop cold leg flows. The details of the transients are described in WCAP-8742 (6). For the purposes of the current evaluation, two cases of the large steam line break are considered. In the first case, off site power is assumed available throughout and the coolant flow is maintained. In the second case, a loss of off site power and the coolant system flow coastdown is assumed. Figure 3-6 shows the pressure variation and Figure 3-7

600 300 f" SEC 500 /- / 180 0 / 1000 g p" SEC t. 300 E / 2000 200 [" / / 4000 SEC y r r 0 0 0.154 0.308 0.462 0.615 0.77 0.923 1.007 DISTANCE TilROUGH THE VESSEL WALL (x/t) FIGURE 3-4. TEMPERATURE DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE LOSS-0F-COOLANT ACCIDENT

16 100 NN E NN h \\ h \\\\ , 6000 SEC 0 y-- 4000 SEC N 3000 SEC a. { 2000 SEC - 1000 SEC g N g 2 300 SEC 5 N s j -100 0 0.154 0.308 0.462

• 0.615 0.77 0.923 1.077 DISTANCE THROUGH THE VESSEL WA'LL (x/t)

FIGURE 3-5. THERMAL H0OP STRESS THROUGH THE VESSEL WALL FOR A LARGE LOSS-OF-COOLANT ACCIDENT

l 1 l1 1l l O 0002 R O TCA E R 0 G 0 N 5 I N 1 NUR ) SS PD MN UO PC E TS N( ALE OM 0I CT ) 0 S RS 0 D OU 0 M TS O CR 1 C AE E EV S R ( E HR E TU M I S WS I T E KR AP ERM BE T MS AY ES 0 T 0 ST 5 N EA GL RO AO LC 6 3 E R UG (- I F 0 ~- 0 0 0 0 0 0 0 0 0 0 0 0 0 4 0 6 2 8 4 2 2 1 1 2 g t u o"a E 5 h m. ! *s. $ 5

• E is 1

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1li l lll 1jl 1ll lllj11' i 1l l 00 0 2 G E L D N D O C L E O S C / P OO G L N / 0 I 3 0 N 1 F l 5 NU 1 R 3 8 S 9 P 1 M = U P W O T L N F A L L A O N O) I CS M O D N RN ) OO 0 S TC 0 D CE I 0 N AS O E( 1 C R E E S HM ( TI IT E W H S l KU T AS ER RE BV ME AR EU TT SA 0 R I 0 EE 5 GP RM AE LT 7 3 E R UG I F 0' i 0 0 0 0 0 0 o 0 0 0 0 0 0 6 5 4 3 2 1 C~= 3E 5.- ' O OO l ll

19 presents the cold leg temperature variation for the first case. The resulting vessel wall temperatures and the hoop stresses are presented in Figures 3-8 and 3-9, respectively. For the second case, Figures 3-10 and 3-11 describe the pressure and temperature transients respectively. The vessel wall temperatures and loop stresses are shown in Figures 3-12 and 3-13 respectively. Reactor coolant system flow is rapidly reduced and the reactor trips at the low-flow signal. 3.2.2.3 Locked Rotor Accident During this postulated accident, the instantaneous seizure of a reactor coolant pump rotor is assumed. The flow through the reactor coolant system is rapidly reduced and the reactor trips at the low-flow signal. Following the trip, heat stored in the fuel rods continues to heat up and expand the coolant. The transient is described in the Unit I FSAR and PSAR. The pressure transient is presented in Figure 3-14. The coolant temperature during the peak pressure of 2778 psia is assumed unchanged at the outlet nozzle and is at operating temperatures. 3.3 Stress Intensity Factors The flaw characterized earlier will be analyzed using the ASME Section XI Appendix A approach. Stress intensity factors for any given flaw is calculated as (Article A-3000). K1 = (om Mm + o b M) Ia 3.1 b Q where K1 = stress intensity factor omso b = membrane and bending stresses MasMb = membrane and bending correction factors 2a = through wall crack penetration Q = flaw shape parameter

20 603 p - 300 SEc 503 f 400 ic00 sEc C / 303 2000 CEC K f / e-= 3000 SEC p g = = = = = = = = = = = = ;ggg ;;; 2. = 100 0 0 0.154 0.308 0.462 0.615 0.77 0.923 1.077 DISTANCE THROUGH THE VESSEL WALL (x/t) FIGURE 3-8. TEMPERATURE DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE STEAM LINE BP,EAK WITH OFF-SITE POWER AVAILABLE

21 100 C .E. E \\ 5 x\\ sk 6000 SEC a. h N 4000 SEC N% ~ 0 SE s w o N 1000 CEC h 300 SEC W I E l c E S5 =5 o -100 0 0.154 0.308 0.462 0.615 0.77 0.923 1.077 DISTANCE THROUGH THE VESSEL WALL (x/t) FIGURE 3-9. PRESSURE AND THEFNAL HOOP STRESS DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE STEAM LINE ~ s BREAK WITH OFF-SITE POWER AVAILABLE

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1l1 l O 0002 D L O C D E P 0 P 0 I l 5 R T 1 S PM U P ~ TN) AS LD ON OO CCE RS ) O( 8 T 0 0 CE AM 0 0 l 0 0 EI 4 C RT 1 ES HS ( TU IS E WR M E KV I T AEE RR BU T i A lAR EE TP SME 0 ET 0 G l 5 RG AE LL 1 1 3 E R U = G I F 0 I 0 0 0 0 0 0 0 0 0 0 0 0 6 5 4 3 2 1 - O[m==g =s 38. ll11

24 600 300 SEC [ 400 1000 SEC / / c / / 2000 SEC 300 3000 SEC E 4000 SEC 200 6000 SEC 100 0 0 0.154 0.308 0.462 0.615 0.77 0.923 1.077 DISTANCE THROUGH THE VESSEL WALL (INCHES) FIGURE 3-12. TEMPERATURE DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE STEAM LINE BREAK WITHOUT OFF-SITE POWER

25 !00 0 b \\ \\ 'S \\\\ 6000 SEC

i N

N 001 E! ~N - 2000 sEC E O N 1000 sEC E 300 stC u R S E S 5 m E o -100 0 0.154 0.308 0.462 0.615 0.77 0.923 1.077 DISTANCE THROUGH THE VESSEL WALL (INCHES) FIGURE 3-13. PRESSURE AND THERMAL H00P STRESS DISTRIBUTION THROUGH THE VESSEL WALL FOR A LARGE STEAM LINE BREAK WITHOUT OFF-SITE POWER

i l l 26 ?cchet Rotor 2300 l 3 l

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d. :li '7!U L ' 11 *- 32':b..Y 28' i

'_1_. II i 2200 0 2 k 6 8 m Time, Seconds FIGURE 3-14. LOCKED ROTOR PRESSURE TRANSIENT

27 3.3.1 Flaw #2 The following are the flaw parameters: flaw location = 3/4t from ID t = 9.125 inches 2a = 1.24 inches 1 = flaw length = 1.935 inches a/t = 0.32 ~ Q = 1.55 for on + ob = 0.6 oys e = flaw eccentricity = 0.25t 2 = 0.14 From Figures A-33003-2 and A-3300-4, the correction factors for locations pti, and pt2 as shown in the sketch below are: M,= 1.03 for pti 1.02 for pt2 = Mb = 0.6 for pti = 0.45 for pt2-3/4 t e = o.25t IS51 pte y1 = 1.935,, 'gy I r-e- 2a = 1.24" t = 9.125 t W FIGURE 1. FLAW #2 PARAMETERS.

28 Equation 1 for pti and pt2 can be written as: for pti Ki = 1.424 (1.03 cm 1 0.60 b) & 3.2 for pt2 K1 = 1.424 (1.02 cm 10.45c b) E 3.3 3.3.2 Flaw #4 The following are flaw parameters: flaw location = 0.56t from ID and at 311' around nozzle t = 9.125 inches 2a = 1.14 inches 1.29 inches i = a = 0.44 L Q = 2.1 for on + ob = 0.6 cys e = 0.06t 2a = 0.125 t n_n6t W e-Oli "E I*29" pt2"V + + ' 1.14" =- 2' FIGURE 2. FLAW #4 PARAMETERS.

29 Correction factors for pti and pt2 in Figure 2 are: Ma = 1.02 for pti = 1.02 for pt2 Mb = 0.2 for pel = 0.05 for pt2 Stress intensity factors for ptt and pt2 are: for pti K1 = 1.223 (1.02a,.1 0.2c b) '/*~ 3*4 for pt2 K1 = 1.223 (1.02c,10.05 eb),/T 3.5

30 4.0 FRACTURE INTEGRITY EVALUATION 4.1 Normal and Upset Conditions (Levels I and II) Under Levels I and II loading cases, the detected indications are analyzed for fatigue crack growth to end of life conditions. The flaw stability for the final crack size is evaluated. The acceptance criteria to be used is: ~ Kla = /TO K1 4.1.1 Fatigue Crack Crowth Analysis Fatigue crack growth is represented by (ASME Section XI App. A Figure A-4300-1). dat = 2.67 x 10-lla K3.726 4.1 dN where aK = (M,aa, + M oo b)$ 4.2 b Q Substituting equation 4.2 in equation 4.1 and integrating the following expression is obtained 1 af = [at-0.863 -0.863 2.67 x 10-11 (M,aa (IJ.863 aN] 0.863 4.3 1 m + M aob)

• b Q

where af = final crack depth at = initial crack depth AN = number of cycles Fatigue Crack Growth for Flaw #2 Substituting the various flaw parameters for flaw #2, Equation

4.3 becomes

for pti ag = [ai-0.863 - 8.593 x 10-ll(1.030m 1 0.6ab)3.726aNJ-1.16 4.4

31 for pt2 af = [ai-0.863 - 8.593 x 10-11(1.02a,1 0.450b)3.726a N]-1.16 4.5 Table 4-1 presents the end of life crack growth resulting from the expected transients. It is seen that the present crack of depth 2 x ai = 2 x 0.62 = 1.24 inches will grow to a crack 2 x af = 2 x 0.625 = 1.25 inches by end of life. Fatigue Crack Growth for Flaw #4 The following equations represent crack growth for Flaw #4. at pti ag = [ai-0.863 - 4.88 x 10-11(1.02o,10.20b)3.726aN]-1.16 4.6 at pt2 af = [at-0.863 - 4.88 x 10-11(1.02om 1 0.050b)3.726a N]-1.16 4.7 Table 4-2 presents the cumulative crack growth for crack #3 to end of life. The total growth is seen to be insignificant. 4.1.2 Flaw Integrity Assessment For all operating and upset load conditions, the minimum temperatures at the flaw locations are considerably in excess of the material dTmbt > KIa = 149 = 3.92 > M K1 38.02 Therefore, the flaw is acceptable.

i r 32

2.I :.: ritt;.e :en:< 3r: :n

n..c.

ORACK CRC'.'M CRACK CROL"TH AT At Pt TRANSIENT NO. OF g ?t, hO. CYCLES a inches a inenes g f a inches g a, inches 1 150 0.62 0.6211876651 0.62 0.62113 2 10,875 0.62119 0.621657 0.62113 0.62115 3 1,500 0.621657 0.621657 0.62115 0.62115 4 150 5 5 60 No growth No arowth 6 30 7, 8, 9 368 .I 10 2.5 x 10 0.621657 0.624778 0.62115 0.622813 11 3 1 ) i No growth 12 30 ]i No growth i ) TABLE 4-2. Fatigue Crack Growth - Flaw No. 3 CRACK CROWTH CRACK CROWTH AT AT 'I TRANSIENT No. OF 2 a inches a inches g g a inches a inches g g 1 150 0.57 0.570354 0.57 0.570353 2 10.875 0.570354 0.570354 3 1,500 4 150 5 60 7 6 30 No growth No growth calculated 7,8,9 368 calculated 10 2.5 x 10' 11 3 12 30 h

33 4.2 Emergency and Accident Conditions (Levels III and IV) The acceptance criteria for flaw to withstand crack initiation under Level III or IV loading is given by KIe>[2~ K1 Loss of Coolant Accident From Figure 3-4 it can be seen that the temperatures during the 6000 see transient duration are in excess of the end of life RTNDT = 25'F for the flaw locations. Therefore, upper-shelf Kie = 149 kai /TEF. will be assumed. From Figure 3-5 the loop stresses are at all times dominated by bending stresses. Since the system is depressurized, there is no contribution due to pressure. The bending stresses at the flaw locations are compressise in nature. Hence, the stress intensity factor will be considered to be K1=0 Kt = 0 kai /TEI. Hence, crack initiation condition of Kic/g > /2~is satisfiad. Large Steam Line Break Case 1: With offsite power The coolant temperature during the transient remains at about 200*F (see Figure 3-7 and 3-8). Therefore, here again upper-shelf material properties are assumed at the flaw locations and will be considered to be at upper-shelf conditions. In Section 2, the minimum end of life toughness was shown to be Kte = Kla = 149 kai /InI. Flaw #2 End of life flaw dimension was calculated to be ag = 0.625 inches. The flaw parameters are

34 a/t = 0.32 2af = 1.25 inches The maximum stresses are obtained for the cold hydro transient (Refer Table 3-2). e, = 40.6 kai and a b = 5.05 kai The maximum stress intensity factor is obtained as Ki = 43.5 kai S. The acceptance criteria 149 = 3.45 > 8 K /g = TJ~5 I 1 is satisifed and hence, flaw is acceptable. Flaw #4 End of life flaw dimensions af = 0.5704 inches Flaw parameters a/t = 0.44 2af = 1.141 inches For the cold hydro transient, maximum '~ K1 = 38.02 kai,/in. Figure 3-9 presents the combined pressure and thermal hoop stresses. The large bending component of the stress provides compressive stresses and hence, K1 = 0 can be assumed. Case 2: Without offsite power Here also the temperatures are in excess of the material RTNDT (Refer Figures 3-31 and 3-12). The stresses shown in Figures 3-13 indicates the same behavior as Case 1 and Kt = 0.

)

E _l

35 Lockad Rotor Pressure Transient (Loss of Load) Conditions: Metal temperatures in excess of 550*F Max. Pressure = 2778 psia (Ref Figure 3-14) Stress intensity computation: = 33.75 kai o b = 5.05 kai om K1 = (for Flaw #2) = 36.2 kai,/Ti. Acceptance criteria: Klc/K " 4*1 > /I~ l 36.2

1 36

5.0 CONCLUSION

S The detected flaw indication #2 and #4 were analyzed using a linear elastic fracture mechanics approach. The approach conforms with the ASME Section XI requirements. The analysis demonstrates that the flaws are stable under all poetulated loading conditions. The margins of safety specified in ASME Section II are adequately met. It is recommended that no repair is necessary for the detected flaw indications. The indications locations should be reinspected and monitored at the next scheduled nozzle weld inspection. M)

37 REFERENCES 1. ASME Boiler and Pressure Vessel Code, Section XI, 1977 Edition. 2. USNRC Reg. Guide 1-99 " Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials", Rev. 1, April 1977. 3. S. E. Yanichko and S. L. Anderson " Analysis of Capsule S from the Wisconsin Electric Power Company and Wisconsin Michigan Power Company Point Beach Nuclear Plant Unit No.1 Reactor Vessel Radiation Surveillance Program," WCAP-8739, November 1976. 4. J. M. Barsom and S. T. Rolfe " Correlations Between KIe and Charpy V-Notch Test Results in the Transition-Temperature Range," Impact Testing of Metals, ASTM STP 466, June 1969, pp 281-302. 5. P. J. Fields, "ASME III, Appendix G Analysis of the Wisconsin Electric Power Company and Wisconsin Michigan Power Company Point Beach Nuclear Plant Unit No.1 Reactor Vessel", WCAP-8741, September 1976. 6. T. A. Meyer, J. M. Kramps, S. Palusany, J. ii. Phillips, P. J. Morris and R. W. Fleming " Fracture Mechenics Evaluation of the Wisconsin Electric Power Company and the Wisconsin Michigan Power Company Point Beach Nuclear Plant Unit No. 1 Reactor Vessel," WCAP 8742, February 1977. _ _ - _ _ _.}}