ML20066K977

Nonproprietary Technical Justification for Eliminating Pressurizer Surge Line Rupture as Structural Design Basis for Sequoyah Units 1 & 2

ML20066K977
Person / Time
Site:	Sequoyah
Issue date:	12/31/1990
From:	Palusamy S, Schmertz J, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML080650640	List: ML20066K977 ML20066L006
References
WCAP-12776, NUDOCS 9102060188
Download: ML20066K977 (87)

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WESTINGHOUSE CLASS 3 WCAP-12776 TECHNICAL JUSTIFICATION FOR ELIMINATING PRESSURIZER SURGE LINE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR SEQUOYAH UNITS 1 AND 2 December 1990 D. C. Bhowmick S. A. Swamy Y. S. Lee D. E. Prager J. F. Petsche Verified: Id C' le/- Jrf. Schmertz ) SMuctural Mechanics Technology Approved: ~ yI Nx ~,[ e ,/'5. 5. P,aiusamy, Manager Diagne'stics and Monitoring Technology Work Performed Under Shop Order: SIXP-9508 WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 c 1990 Westinghouse Electric Corp. l 4857s/121890 10

i TABLE OF CONTENTS Section Title Pace

1.0 INTRODUCTION

1-1

1.1 Background

1-1 1.2 Scope and Objective 1-1 1.3 References 1-3 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 2-1 2.1 Stress Corrosion Cracking 2-1 2.2 Water Hammer 2-3 2.3. Low Cycle and High Cycle Fatigue 2-4 2.4 Summary Evaluation of Surge Line for Potential Degradation During Service 2-4 2.5 References 2-5 3.0 MATERIAL CHARACTERIZATION 3-1 3.1 Pipe and Weld Materials 3-1 3.2 Material Properties 3-1 3.3 References 3-2 4.0 LOADS FOR-FRACTURE MECHANICS ANALYSIS 4-1 4.1 Loads for Crack Stability Analysis 4-2 4.2 Loads for Leak Rate Evaluation 4-2 4.3 Loading Condition 4-2 4.4 Summary of Loads Geometry and Materials 4-4 4.5 Governing Locations 4-5 e 4451s/112090 to [

TABLE OF CONTENTS (cont.) Section Title Pace 5.0 FRACTURE MECHANICS EVALUATION 5-1 5.1 Global failure Mechanism 5-1 5.2 Leak Rate Predictions 5-2 5.3 Stability Evaluation 5-4 5.4 References 5-5

6.0 ASSESSMENT

OF FATIGUE CRACK GROWTH 6-1 6.1 Introduction 6-1 6.2 Initial Flaw Size 6-2 6.3 Results of FCG Analysis 6-2 6.4 References 6-3

7.0 ASSESSMENT

OF MARGINS 7-1

8.0 CONCLUSION

S 8-1 s APPENDIX A Limit Moment A-1 e e f 4 44S F t >112000 10 qjj

LIST OF TABLES '~ Table Title Pace 3-1 Room Temperature Mechanical Properties of the Pressurizer Surge Line Materials and Welds of the Sequoyah Unit 1 3-3 3-2 Room Tempersture Mechanical Properties of the Pressurizer Surge Line Materials and Welds of the Sequoyah Unit 2 3-4 3-3 Room Temperature ASME Code Minimum Properties 3-5 3-4 Representative Tensile Properties for Sequoyah Unit 1 3-6 3-5 Representative Tensile Properties for Sequoyah Unit 2 3-7 3-6 Modulus of Elacticity (E) 3-B 4-1 Types of Loadings 4-6 4-2 Normal and faulted Leading Cases for Leak-Before Break Evaluations 4-7 4-3 Associated Load Cases for Analyses 4-8 4-4 Summary of LEB Loads and Stresses by Case for Sequoyah Unit 1 4-9 4 Summary of LBB Loads and Stresses by Case for Sequoyah Unit 2 4-10 44$ Ft o 1200010 jy

LIST OF TABLES (cont.) Table Title Pace 5-1 Leak Rate Crack length for Sequoyah Unit 1 5-6 5-2 Leak R?.te Crack Length for Sequoyah Unit 2 5-7 5-3 Summary of Critical Flaw Size for Sequeyah Unit 1 5-8 5-4 Summary of Crit 4;al Flaw Size for Sequeyah Unit 2 5-9 6-1 Fatigue crack Growth Results for 10% of Wall Initial Flav Size 6-4 7-1 '.eakage Flaw Sizes, Critical Flaw Sizes and Margins for Sequoyah Unit 1 7-2 7-2 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Sequoyah Unit 2 7-3 7-3 LBB Conservatisms 7-4 o 4457s/112090 to

l LIST OF f!GURES 4 Figure Title Page 3-1 Sequoyah Unit 1 Surge Line Layout 3-9 3-2 Sequoyah Unit 2 Surg 0 Line Layout 3-10 4-1 Sequoyah Unit 1 Surge Line Showing the Governing Locations 4-11 4-2 Sequoyah Unit 2 Surge Line $howing the Governing Locations 4-12 5-1 Fully Plastic Stress Distribution 5-1( 5-2 Analytical Predictions of Critical flow Rates of Steam-Water Mixtures 5-11 5-3 [ l"'CPressureRatioasa Function of L/D 5-12 5-4 Idealized Pressure Drop Profile through a Postulated Crack 5-13 5-5 Loads Acting on the Model at the Governing Location 5-14 5-6 Critical Flaw Size Prediction for Sequoyah Unit 1 Node 1020 Case D 5-15 5-7 Critical Flaw Si:e Prediction for Sequoyah Unit 1 Node 1020 Case E 5-15 l 5-8 Critical Flaw Size Prediction for Sequoyah Unit 1 Node *020 Case F 5-17 .eu.mmon y3 l

LIST OF FIGURES (cont.) ~ Fioure Title Page 5-9 Critical Flaw Size Prediction for Secuoyah Unit 1 Node 1020 Case G 5-1B 5-10 Critical Flaw Size Prediction for Sequoyah Unit 1 Node 1080 Case D 5-19 5-11 Critical Flaw Size Prediction for Sequoyah Unit 1 Node 1080 Case E 5-20 5-12 Critical Flaw Size Prediction for Sequoyah Unit 1 Node 1080 Case F 5-21 5-13 Critical Flaw Size Prediction for Saquoyah Unit 1 Node 1080 Case G 5 22 5-14 Critical Flaw Size frediction for Sequoyah Unit 2 Node 1020 Case 0 5-23 5-15 Critical Flaw Size Prediction for Sequoyah Unit 2 Node 1020 Case E 5-24 5-16 Critical Flaw Size Prediction for Sequoyah Unit 2 Node 1020 Case F 5-25 5-17 Critical Flaw Size Prediction for Seauoyah Unit 2 Node 1020 Case G 5-26 5-18 Critical Flaw Size Prediction for Sequoyah Unit 2 Node 1080 Case D 5-27 5-19 Critical Flaw Size Prediction for Secuoyah Unit 2 Node 1080 Case E 5-28 <sw nnove p;

1 l LIST OF FIGURES (cont.) 1. ~ Figure Title Pace i 5-20 Critical Flow Size Prediction for Sequoyah Unit 2 Nede 1080 Case F 5-29 5-21 Critical Flaw Size Prediction for Sequoyah Unit 2 Node '080 Case G 5-30 6-1 Dett-nation of the Effects of Thermal Stretification on Fatigue Crack Growth 6-5 6-2 Fatigue Crack Growth Methodology 5-6 63 Fatigue Crack Growtn Rate Curve for Austenitic } Stainless Steel 6-7 6-4 Fatigue Crack Growth Rate Equation for Austenitic Stainless Steel 6-B 6-5 Fatigue Crack Growth Critical Locations 9 A-1 Pipe with a Through-Wall Crack in Bending A-3 4 G4 S ? t '11209010 yjjj

F SECTION 1.0 INTRODUCTION 1.1 Backcround The current structural design basis for the pressurizer surge line requires postulating non-mechanistic circumferential and longitudinal pipe breaks. This results in additional plant nardware (e.g. pipe whip restraints and jet shields) which would mitigate the dynamic consecuences of the pipe breaks. It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for the surge line. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that can be used for establishing that a circumferential type break will not occur within the pressurizer surge line. The evaluations considering circumferential1y oriented flaws cover longitudinal cases. The pressurizer surge line is known to be subjected to thermal stratification and the effects of thermal stratification for Sequoyah surge lines have been evaluated and documented in WCAP-12777. The results of the stratification evaluation as described in WCAP-12777 have been used in the leak before-break evaluation presented in this report. 1.2 Scoce and Objective The general purpose of this investigation is to demonstrate leak-before-break for the pressurizer surge line. The scepe of this work covers the entire pressurizer surge line from the primary loop nozzle junction to the pressurizer nozzle junction. A schematic drawing of the piping system is shown in Section 3.0. The recommendations anc criteria proposed in NUREG 1061 Volume 3 (1-1) are used in this evaluation. The criteria and the resulting steps of the evaluation procedure can be briefly summarized as follows:

1) Calculate the applied loads.

Identify the location at which the highest stress occurs. 2) Identify the materials and the associated material properties. <wsmtwo 11

. ~- _ 3) Postulate a surface flaw at the governing location. Determine fatigue crack growth. Show that a through wall crack will not result. 4) Postulate a through-wall flaw at the governing location. The si:e of the flaw should be large enough so that the le:kage is assured of detection with margin using the installed leak detection equipment when the pipe is subjected to nornal operating loads. A margin of 10 is demonstrated between the calculated leak rate and the leak detection capability. 5) Using maximum faulted leads, demonstrate that there is a margin of at least 2 between the leakage si:e flaw and the critical size flaw. 6) Review the operating history to ascertain that operating experience has indicated no particular susceptibility to failure from the effects of corrosion, water hammer or low and high cycle fatigue. 7) For the base and weld metals actually in the plant provide the material properties including toughness and tensile test data. Justify that the properties used in the evaluation are representative of the plant specific material. Evaluate long term effects such s thermal aging where applicable. l 8) Demonstrate margin on applied load. The flaw stability analyses is performed using the methodology described in SRP 3.6.3 (1-2). The leak rate is calculated for the normal operating condition. The leak rate prediction model used in this evaluation is an (' Ja.c,e The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through wall flaw to normal operating loads (1-3). Surface roughness is accounted for in determining the leak rate through the postulated flaw. <w mm vo 1.g

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!..s ' References 1-1 Report of the U.S. Nuclear Regulatory Commission Piping Review Committee

• Evaluation of Potential for Pipe Breaks, NUREG 1061, Volume 3 November

1984, 1-2 Standard Review Plan; public comments solicited; 3.6,3 Leak-Before-Break Evaluation Procedures; Federal Register /Vol. 52, No. 167/ Friday, August 28, 1987/ Notices, pp. 32626-32633.

1-3 NUREG/CR-3464, 1983, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks." 4 1 5 .m.mm io 13 ,1i ..-...s...vi m. ._,w. y ...,y---c.,- w. -..v.--iw,, -.--,,,..e,., ,,--e-4 w, , -,m-.

SECTION 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 2.1 gress Corrosion Cracking The Westinghouse reactor coolant system primary loop and connecting Class 1 lines have an operating history that demonstrates the inherent operating stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosioncracking). This operating history totals over 400 reactor years, including five plants each having over 15 years of operation and 15 other plants each with over 10 years of operation, in 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group. (The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pressurized Water Reactors (PWR's). The results of the study performed by the PCSG were presented in NUREG-0531 (Reference 2-1) entitled " Investigation and Evaluation of Stress (srrosion Cracking in Piping of Light Water Reactor Plants." In that report the PCSG stated: "The PCSG has determined that the potential for stress-corrosion cracking in PW't primary system piping is extremely low because the ingredients that produce IGSCC are not all present. The use of hydra:ine additives and a hydrogen overpressure limit the oxygen in the coolant to very low levels. Other impurities that might cause stress-corrosion cracking, such as halides or caustic, are also rigidly controlled. Only for brief periods during reactor shutdown when-the coolant is exposed to the air and during the subsequent startup are conditions even marginally capable of producing stress-corrosion cracking in the primary systems of PWRs. 4m.mme 2-1

Operating experience in PWRs supperts this determination. To date, no stress-corrosion cracking has been reported in the primary piping or safe ends of any PWR." During 1979, several instances of cracking in PWR feedwater piping led to the establishment of the third PCSG, The investigations of the PCSG reported in NUREG-0691 (Reference 2-2) further confirmed that no occurrences of !GSCC have been reported for PWR primary coolant systems. As stated above, for the Westinghouse plants there is no history of cracking failure in the reactor coolant system loop or connecting Class 1 piping. The discussion below further cualifies the PCSG's findings. For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously: high tensile stresses, susceptible material, and a corrosive environment. Since some residual stresses and seme degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimiced by properly selecting a material immune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other material in the system, applicable ASME Code rules, fracture toughness, welding, fabrication, and processing. The elements of a water environment kncan to increase the susceptibility of austen4 tic stainless steel to stress corrosion are: oxygen, fluorides, chlorides, hydroxides, hydrogen peronide, and reduced forms of sulfur (e.g., sulfides, sulfites, and thionates), Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment, Prior to being put into service, the piping is cleaned internally and externally, During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping, ~. I g

During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Centaminant concentratione are kept below the thresholds known to be conducive to stress corresion cracking with the major water chemistry control standards being included in the p'ent operating procedures as a condition for plant operation. For example, during r'ormal power operation, oxygen concentration in the RCS and connecting Class 1 lines is expected to be in the ppb range by controlling charging flow chem-istry and maintaining hydrogen in the reactor coolant at specified concentr6-tions. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and fluorides within the specified limits. This is assured by controlling charging flow chemistry. Thus during plant opera-tion, the likelihood of stress corrosion cracking is minimi:ed. 2.2 Water Hammer Overall, there is a low potential for water hammer in the RCS and connecting surge lines since they are designed and cperated to preclude the voiding ro~-tition in normally filled lines. The RCS and connecting surge line including piping and components, are designed for normal, upset, emergency, and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control red position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled -in the design process. Adcitionally, Westinghouse has instrumented typical

• eactor conlant systems to verify the flow and vibration characteristics of the system and connecting surge lines.

Precoerational testing and operating a experience have verified the Westinghouse aporoach. The operating transients i 48 S ?s q 1209010 2-3 l

of the RCS primary piping and connected surge '4nes are such that no significant water hammer can occur. 2.3 Low Cycle 6nd High Cycle Fatigue Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage f a' tor evaluation to show compliance with the rules of Section 111 of the ASME Code. A further evaluation of the low cycle fatigue leading is discussed in Section 6.0 as part of this study in the form of a fatigue crack growth analysis. Pump vibrations during operation would result in high cycle fatigue loads in the piping system. During operation, an alarm signals the exceedance of the RC pump shaft vibration limits. Field measurements have been made on the reactor coolant loop piping of a number cf plants during hot functional testing. Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest. Recent field measurements on typical PWR plants indicate vibration amplitudes less than 1 ksi. When ^ translated to the connecting surge line, these stresses would be even lower, well below the fatigue endurance limit for the surge line material and would result in an applied stress intensity factor below the threshold for fatigue crack growth. 2.4 Summary Evaluatien of Surge Line for Potential Degradation During Service Tnere has never been any service cracking or wall thinning identified in the pressurizer surge lines of Westinghouse PWR design. Sources of such degradation are mitigated by the design, construction, inspection, and cperation of the pressurizer surge piping. There is no mechanism for water hammer in the pressurizer / surge system. The pressurizer safety and relief piping system which is connected to the top of the pressurizer could have loading from water hammer events. However, these leads are effectively mitigated by the pressurizer and have a negligible ? effect on the surge line. .m.m m in 2-4

Wall thinning by erosion and erosion-corrosion effects will not occur in the surge line due to the low velocity, typically less than 1.0 ft/see and the material, austenitic stainless steel, which is highly resistant to these degradation mechanisms. Per NUREG-0691, a study of pipe cracking in PWR piping, only two incidents of wall thinning in stainless steel pipe were reported and these were not in the surge line. Although it is not clear from the report, the cause of the wall thinning was related to the high water velocity and is therefore clearly not a mechanism which would affect the surge -line. it is well known that the pressurizer surge lines are subjected to thermal strati /ication and the effects of stratification are particularly significant during certain modes of heatup and cooldown operation. The effects of stratification have been evaluated for the Sequoyah plant surge lines and the loads, accounting for the stratification effects, have been derived in WCAP-12777. These loads are used in the leak-before-break evaluation described in this report. The Se:/oyah Units 1 & 2 surge line piping and associated fittings are forged product forms (see Section 3) which are not susceptible to toughness degradation due to thermal aging. Finally, the maximum operating temperature of the pressurizer surge piping, which is about 650'F, is well below the temperature which would cause any creep damage in stainless steel piping. 2.5 References 2-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of Light Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979. 2-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980. <w.mm to g,5

4 d SECTION 3.0 MATERIAL CHARACTERIZATION 4 3.1 Pjpe and Weld Materials The pipe material of the pressurizer surge line for the Sequoyah Units 1 & 2 is A376/TP316. These are a wrought product form of the type used for the primary loop piping of several PWR plants. The surge line is connected tt) the primary loop nozzle at one end and the other end of the surge line is connected to the pressurizer nozzle. The surge line system dees not include any cast pipe or cast fitting. The welding processes used are shielded metal arc (SMAW) and submerged arc (SAW). Weld locations are identified in Figures 3-1 and 3 2. In the following section the tensile properties of the materials are presented for use in the leak-before-break analyses. 3.2 Material Precerties The room temperature mechanical properties of the Sequoyah Units 1 & 2 surge line materials were obtained from the Certified Materials Test Reports and are given in Table 3-1 and 3-2. The room temperature ASME Code minimum properties are given in Table 3-3. It is seen that the measured properties well exceed those of the Code. The representative minimum and average tensile properties - were established from the Certified Material Test Report. The material properties at temperatures (135'F, 205'F, 300*F, 330'F and 653*F) are required for the leak rate and stability analyses discussed later. The minimum and average tensile properties were calculated by using the ratio of the ASME Section 111 properties at the temperatures of interest stated above. Tables 3-4 and 3 5 show the tensile properties at various temperatures for the Sequoyah Units 1 & 2. The modulus of elasticity values were established at various temperatures from the ASME Section III (Table 3-6). In the leak-before-break evaluation, the representative minimum properties at .. ~,,. m 31

temperature are used for the flan stability evaluatiens and the representative average preperties are used for the leak rate preci:tiens, ine minimum ultimate stresses are used for stability analyses. These pregerties are summarized in Tables 3-4 and 3-5. 3,3 Peferences 3-1 ASME Boiler and Pressure Vessel Ce:e Se: tion !!!, Division 1, Accendices July 1, 1989. l l l 4 g. 4sS7s't i2090 ic 3y

TABLE 3-1 Room Temperature Mechanical Preperties of the Pressurizer Surge Line Materials and Welds of the Secuoyah Unit 1 ULTIMATE YlELO 10 HEAT NO./ SERIAL NO. MATERIAL STRENGTH STRENGTH ELONG. R/A psi psi (%) (%) 1 J2469/6559 A376/TP316 86,100 43,700 51.4 67.5 83,400 42,400 52.5 72.3 2 J2471/6551 A376/TP316 03,400 41,800 53.1 69.3 84,300 39,900 50.8 65.7 3 J2617/7044 A376/TP316 85,900 42,500 50.5 65.9 88,400 43,900 50.0 63.5 4 J2469/6538 A376/TP316 87,400 44,900 51.7 73.0 87,100 48,000 47.2 67.3 5 J2469/6538 A376/TP316 87,400 44,900 51.7 73.0 87,100 48,000 47.2 67.3 SW - Shop Weld All shop welds were fabricated by SAW FW - Field Weld All field welds were fabricated by GTAW and SMAW combination O nivnem io 33 1. 4-

TABLE 3-2 Room Temperature Nechanical Preperties of the Pressurizer Surge Line Materials and Welds of the Sequoyah Unit 2 ULTIMATE YlELD 10 HEAT NO / SERIAL NO. MATERIAL STRENGTH STRENGTH ELONG. R/A psi psi (%) (%) 1 J2471/6553 A376/TP316 83,600 41,800 50.0 68.2 83,600 41,800 51.4 6B.2 2 J2469/6562 A376/TP316 83,600 41,800 52.2 71.4 83,500 40,800 52.5 66.9 3 J2469/6562 A376/TP316 83,600 41,800 52.2 71.4 83,500 40,800 52.5 66.8 4 J2470/6541 A376/TP316 83,000 41,600 50,9 69,9 83,900 40,700 53.5 68.9 SW - Shop Weld All shop welds were fabricated by SAW FW - Field Weld All field welds were fabricated tj GTAW and SMAW combination l l .m.m rao io 34

TABLE 3 3 I Room Temperature ASME Code Minimum Preperties Material Yield Stress Ultimate Stress (psi) (esi) i A376/TP316 30,000 75,000 9 6 s e 9 91 U e .m.., uan o 3-5 re - -n-a ,,-m---, ...n,,-. .-.--w.--r..,-.w.-..,vn.,n=>.. .,,,...r.--,,---w ,.~--+..,w...w...-e--..,..-e. ,.n.nn. van.n,. r-e n,-n,w, v.n, ,n,-e,- ,,,,,. wwws

TABLE 3 4 Representative Tensile Properties for Sequoyah Unit 1 i Minimum Temperature Minimum Average Vitimate 1-Material (*F) Yield (psi) Yield (psi) (psi) A376/TP316 100 39,900 44,000 83,400 135-37,940 41,840 83,400 205 34.150 37,660 83,300 300 30,990 34,170 81,620 330 30,230 33,330 81,080 653 24,570 27,100 79,840-- 9 4 e e 4 i atl7,/11209010 3.g ,m._ ~. * -.,m. -..-..,-,c,.,%,. ~.,,. -. s,. .y., ,,.,,.[ , ~,. - _, -, _ -y

r d i TABLE 3-5 j Representative Tensile Properties for Sequoyah Unit 2 Minimum Temperature Minimum Average Ultimate Material ('f) Yield hsi) Yield (psi) (psi) A376/TP316 100 40,700 41,390 83,000 135 38,710 39,360 83,000 205 34,840 35,430 82,910 300 31,600 32,140 81,220 i 330 30,830 31,360 80,690 653 25,070 25,490 79,460 9 I k 4 e 9 .m.mme io 3-7 .~.-..

. _=_ -.. O TABLE 3-6 Modulus of Elasticity (E) Temocrature E (ksi) (*F) 100 28,138 135 27,950 205 37,600 300 27,050 330 26,885 653 25,035 4 4 5 L .mcu rm io 3.g

I m C 3) SW SW 1, ,- r FW PRESSURIZER FW , HOT LEG FW 7,'. D 4 .b3 SW FW - Field Weld SW - Shop Weld l l Figure 3-1 Secuoyah Unit 1 Surge Line layout .mems.o io 39 l t

i + SW _m FW PRESEURIZER k/ ~ HOT LEG FW o FW u; .a SW l FW - Field Weld SW - Shop Weld 1 Figure 3-2 Sequoyah Unit 2.iurge Line Layout .m. nim io 3-10

l SECTION 4.0 1 LOADS FOR FRACTURE MECHANICS ANALYS!$ 1 Figures 3-1 and 3-2 show schematic layouts of the surge lines for Sequoyah Units 1 & 2 and identify the weld locations. The stresses due to axial leads and bending moments were calculated by the following equat un: o=k+ (4-1)

where, o

= stress axial load F a bending moment M = metal cross-sectional area A = Z section modulus = The-bending moments for the desired leading combinations were calculated by the following equation: M = (M 2 y 2) 0.5 g y (4-2)

where, 9

M a g bending moment for required loading M = y Y component of bending moment 7 Z component of bending moment M = j l The axial load and bending moments for crack stability analysis and leak rate predictions are computed by the methods to be explained in Sections 4.1 and -4.2 which follow. .m n mu to 41 + +. ~. + - w-.-- m ...m..-.-

1 4.1 Leads for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the following equations:

4 !F I * 'F I ^ 'F ! + iF (4'3) F = 0W TH 'p SSE (4'4) l(M )DWI

• IM I*

Y SSE M = y Y TH y l(M )DWI + IH2 TH' * ' Z SSE' (4-5) M = 7 7 Deadweight DW = Applicable thermai load (normal er stratified) TH = lead due to internal pressure P = SSE loading including seismic anchor motion SSE = 4,2 Leads for leak Rate Evaluation The normal operating loads for leak rate predictions were calculated by the following general equations: F0W + FTH + F (4-6) F = p (M )DW * (N )TH (4-7) M = y Y y (M )DW * ("Z)TH (4'0) M = Z 7 The parameters and subscripts are the same as those explained in Section 4.1. 4.3 Leading. Conditions Because thermal stratification can cause large stresses at heatup and coolcown temperatures in the range of 455'F, a review of stresses was used to identify the worst situations for LBB applications. The loading states so identified are given in Table 4-1. I i 44 $7311209010 /

Seven loading cases were identified for LBB evaluation as given in Table 4 2. Cases A, B, C are cases for leak rate calculations with the remaining cases being the corresponding faulted situations for stability evaluations. The cases postulated for leak-before-b eak are summari:ed in Table 4-3. The cases of primary interest are the postulation of a detectable leak at normal poner conditions ( 3a,c.e The combint. tion ( ja.c.e .m.mem to 43

_- - -. -. -. - -. -. -_. ~. - -. - - The more realistic cases ( i a.c.e j [ 1 l l Jac.e The logic for this ai ( Ja e,e is based on'the following: Actual-practice, based on experience of other plants with this tyce of situation, indicates that the plant operators complete the cooldown as cuickly as possible once a leak in the primary system is detected. Technical $pecifications may require cold shutdown within 36 hours but actual practice is that the plant depressurizes f.se system as soon as possible once a primary system leak is detected. Therefore, the hot leg is generally on the warmer . side of the limits (-200'F) when the pressurizer bubble is quenched. Once the bubble is-quenched, the pressurizer is cooled down fairly quickly reducirg the AT in the system.- l: 4.4 Summary of 1. cads and Geometry The load combinations were evaluated at the various weld Ircations. Normal loads were. determined using-the algebraic sum method whereas faulted loads were combined using the absolute sum method. 1 4 4417,/I12090 to 44 m .~*---e .em.-,..- --...-...e-* c --w-w .c-,---%-.-,-.,%.--- .. = r-e-v-,-c--m,---s

f v 4.5 Governine Locations All the welds at Secuoyah Units 1 and 2 surgelines are fabricated using the SMAW and SAW procedure, iht,fo11 ewing governing locations were estabiished for each type of the weld. SMAW Weld w Nede 1020 (het leg nozzle junction) f er Secuoyah Units 1 and 2 sw SAW Weld Node 1080 fer Sequoyah Units 1 and 2 The loads and stresses at these critical locations for all the loading combinations are shown in Tables 4-4 and 4-5. Figure 4-1 and 4-2 show the governing locations. 9 Y ..u.a n..e 4-5

TABLE 4-1 Types of Loadings Pressure (P) Dead Weight (DW) Normal Operating Thermal Expansion (TH) Safe Shutdown Earthquake and Seismic Anchor Motion (SSE)a a,c.e pfk ~ 'SSE is used to refer to the absolute sum of these loadings. i f l e savsim uc,o 45 .L, . - _.. _ _ _ _.,. ~ ~.

. - _ = - - _ _ -. TABLE 4-2 Normal and Faulted Leading Cases for Leak-Before-Break Evaluations CASE A: This is the normal eperating case at 653*F consisting of the algebraic sum of the leading components due to P. DW and TH. -. a,:,e CASE B: CASE C: CASE D: This is tha faulted operating case at 653*F consisting of the absolute sum (every component lead is taken as positive) of P, DW, TH and SSE. CASE E: a,c.e CASE F: . CASE G: e ensamma i .m. m see io 4,7 n.-. -w---

TABLE 4-3 Associated Lead Cases for Analyses A/D This is here-to-fore standard leak-before-break evaluation. a,c.e A/F l B/E B/F 8 B/G i i C/Ga b 8 These are judged to be low probability events. t amim inc io 43

a J TABLE 4-4 Summary of LBB Loads and Stresses by Case for Sequoyah Unit 1 Node Case F (lbs) S (PSI) H (in-lb) S IPSI) S (PSI) X X B B T 1020 A 251742 5025 1642805 11191 16216 2 1020 l a,c.e 1020 d 1020 0 258370 5157 2122075 14456 19613 1020 a,c.e 1020 1020 1080 A 248830 4967 628052 4278 9245 1080 a,c.e 1080 1080 0 253777 5065 1686229 11487 16552 1080 a,c,o 1080 1080 ~_ s. asstsettaoso 10

TA8LE 4-5 Summary of L8B Loads and Stresses by Case for Sequoyah Unit 2 Node Case F (ibs) S (PSI) M (in-lb) S (psi) S lPSI) X X B B T 1020 A 251742 5025 1642805 11191 16216 ~l a.:,e 1020 1020 1020 0 258370 5157 2122075 14456 19613 i a,:,e 1020 1020 l 1020 o 1080 A 248830 4967 628052 4278 9245 __ 1080 a,,e 1080 1080 0 253777 5065 1686229 11487 16552 __ 1080 a,:,e 1080 1080 e .sv.. u ma ia 4-10

O Pipe 14" Schedule 160 O Minimum Wall Thickness is 1,251" Highest Stressed Weld Location (SWJ) -- Highest Stressed Weld Location (SMAW) 1080 m PRESGURIZER 1020 HOT LEG O Figure 4-1 Secucyan Unit 1 Surge Line Showing Governing Locations ent.,n is,e io 41 4

o Pipe 14" Sct.cGule 160 o Minimum Wall Thickness is 1.?51" Highest Stressed Weld Location (SAW) Highest Stressed hWeldLocation (SMAW) 080 e s PRESSURIZER 1020 ( N _ HOT LEG Figure 4-2 Sequoyah Unit 2 Surge Line Showing Governing Locations l 4457:<t1I690 to 4 12

SECTION 5.0 FRACTURE MECHANICS EVALUATION

5. '1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel should ce done with plastic fracture rethodology because of the large amount of deformation accomoanying fracture.

One method for precieting the failure of ductile material is the ( Ja,c.e method, based on traditional clastic limit lea:: concepts, but accounting for [ Ja,c.e and taking into account tne cresence of a flaw. The flawed component is predicted to fail wnen tne remaining net section reaches a stress level a which a plastic hinge is formed. The stress level at which this occurs is ter.ned as the flow stress. [ )C This methodology has been shown to be applicable to ductile piping through a large number of experiments and is used here to_ predict the critical flaw size in-the pressuricer surge line, ire failure criterion has been obtained by reauiring equilibrium of the section containing the flaw-(Figure 5-1) when loads are applied. The detailed cevelopment is provided in Appendix A for a through wall qircumferential fisw in a pipe secti:n with internal pressure, axial force, and imposed bending moments. The limit mcment for such a pipe is gisan by: a,c.e [ ] (5-1) where: ( )a,c.e 1!- L i .m nime io 51

( ja.c.e (5-2) The analytical model described above accuratoly accounts for the internal pressure as well as imposed axial force as they affect the limit niement. Good agreement was found between the analytical predictions and the experimental results (reference 5-1). Flaw stability evaluations, using this analytical t model, are presented in section 5.3. 1 5.2 Leak Rate Predictions Fracture mechanics analysis shows in general that postulated through wall cracks in the surge line would remain stable and do not cause a gross failure of this component. However, if such a through wall crack did exist, it woulc be desirable to detect the leakage such that the plant could ce brought to a safe shutdown condition. The purpose of this section is to discuss the metn0d which will be used to predict the flow through such a postulated crack and present the leak rate calculation results for through wall circumferential cracks. 5,2,1 General Considerations The flow of-hot pressuri:ed water through an opening to a lower back oressure (causing choking) is taken into account. For long channels where the ratio of the channel length, L, to hydraulic diameter, O, W D ) is greater than g H L -( Ja,c,e, both ( la,c.e must be considered. 1 In this situation the flow can be described as being single phase through tne channel until the local pressure equals tne saturation pressure of the fluid. me,m mo io 5-2

6 At this point, the flow begins to flash and choking occurs, Pressure losses due to momentum changes will dominate for ( Jac.e However, for large L/D values, the friction pressure drop will become important and must g be considered along witn the momentum losses due to flashing. 5.2.2 Calculational Method in using the ( ja.c.e, The flow rate through a crack was calculated in the following manner. Figure 5-2 from reference 5-2 was used to estimate the critical pressure, Pc, for tne primary loop enthalpy condition and an assumed flow. Once Pc was founa for a given mass flow, the ( Ja,c,e was found from figure 5-3 taken from reference 5-2. For all cases consicerec, since ( 3a,c.e Therefore, this method will yield the two phase pressure drop due to momentum effects as illustrated in figure 5-4. Now using the assumed flow rate, G, the frictional pressure drop can te calculated using aPf=[ la.c e (5,3) where the friction factor f is determined using the (. )8'C The crack relative roughness, e, was obtained fecm fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was ( Ja,c,e py3, ass 7en treso to g.3

The-frictional pressure drop using Equation 5-3 is then calculated for the assumed flow and added to the [ Fauske modella,c,e to obtain the total pressure drop from the system under consideration to the atmosphere.

Thus, Absolute Pressure - 14.7 = {

Ja,c.e (5-4) for a given assumed flow G. If the right-hand side of equation 5-4 does not agree with the pressure difference between the piping under consideration and the atmosphere -then the procedure is repeated until equation 5-4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. 5.2.3 Leak Rate Calculations Leak rate calculations were performed as a function of postulated through wall crack length for the critical locations previously identified. The crack opening area was estimated ut,ing the methed of reference 5-3 and the leak rates were calculated using the calculational methods described above. The leak rates were calculated using the normal ooerating loads at the governing nodes identified in section 4.0. The crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for critical location at the Sequoyah Unit 1 & 2 pressuri:er surge lines are shown in Tables 5-1 anc 5-2. 5.3 Stability Evaluation A typical segment of the no::le under maximum loads of axial force F and bending moment M is senematically illustrated as shown in figure 5-5. In order to calculate the critical flaw size, plots of the limit moment versus crack length are generated as shown in figures 5-6 to 5-21. The critical flaw l-sizo correspo'd:, to the intersection of this curve and the maximum load line, The critical flaw site h calculated using the lower bound base metal tensile l properties established in section 3.0. 4457 a d 1209010 g f, l l

l l The weld at the locations of interest (i.e. the governing location) are SMAW welds. Therefore, "Z" factor corrections for SMAW and SAW welds were applied (references 5-4 and 5-5) as follows: Z = 1.15 (1 + 0.013 (0.0. - 4)] (for SMAW) (5-5) Z

• 1.30 (1 + 0.010 (0.0. - 4)] (for SAW)

(5-6) where 00 is the outer diameter in inches. Substituting 00 = 14.00 inches, the Z factor was calculated to be 1.2995 for SMAW and 1.43 for SAW. The applied loads were increased by the Z factors and the plots of limit load versus crack length were generated as shown in figure 5-6 to 5-21. Tables 5-3 and 5-4 show the summary of critical flaw sizes for Sequoyah Units 1 & 2. 5.4 References 5-1 Kanninen, M. F. et al., " Mechanical Fracture Predictions for Sensiti:ed Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976. 5-2 (Fauske, H. K., " Critical Two-Phase, Steam Water Flows," Proceedings of the Heat Transfer and Fluid Mechanics Institute, Stanford, California, Standford University Press, 1961.)a,c,e 5-3 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, Statember 19S3, 5-4 NRC letter from M. A. Miller to Georgia Power Company, J. P. O'Reilly, dated September 9, 1987. 5-5 ASME Code Section XI, Winter 1985 Addendum, Article IWS-3640. 5-6 Standard Review Plan; Public Comment Solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal Register /Vol. 52, No. 167/ Friday, August 28, 1987/ Notices, pp. 32626-32633. 4857s/111093 to g.g

TABLE 5-1 Leak Rate Crack Length for Sequoyah Unit 1 Node Point Lead Case Temeerature Crack Leneth (in.) (*F) (for 10 gpm leakage) a,:,e 1020 a,:,e 1080 e i i 1 i l 44S?vti2000 to

.g

TABLE 5-2 9 Leak Rate Crack length for Sequoyah Unit 2 Node Point Lead Cas;! Temeerature Crack length (in.) ('F) (for 10 gpm leakage) a,C,0 1020 a,c e 1080 I l i 4417 rttl59010 5-7 .w -c.-

TABLE 5-3 ~. Summary of Critical Flaw Si:e for Sequoyah Unit 1 Critical Node Point Lead Case Temeerature Flaw Size (in) (*f) a,:,e 1020 t A C,e 1080 G ~ .m.,m s.o i. 53

TABLE 5-4 Summary of Critical Flaw Si:o for Sequoyah Unit 2 Critical Node Point Load Case Temoerature flaw Size (in) ('F) a.c.e i 1020 a,c e ~1080 mammum 9 l-l l-44 $ 74 /III S9010 .}

, a,c,e i Figure 5-1 Fully Plastic Stress Distribution 5-10 f

- a,:,e I ( -C E U i >c 8 a W> M<3 .J 2 STAGNATION ENTHALPY (10 Stu/lb) Figure 5-2 Analytical Predicti:ns Of Critical Flow Rates of Steam-Water viitures .m.n u m io

=.... _.. _ -4 d T a :,e f SE 9-a a4 - W = .- a u o LENGTH / DIAMETER RATIC (L/D) o 1 Figure'5-3 (- ]8'C Pressure Ratio as a Function of L/D ..immimo 5-12 l -w^ n- -,, ,x- ~v--, --,,-~,c r v-wwr -w w w-e--- ,-,.rs-- e-1- ~ w

1 _ a :,e I a,c.e / / J 3 i g _- - \\- Figure 5-4 Idealized Pressure Deco Profile Through a Postulated Crack i i 40$I,/f1'S80 to 5-13 i

.e r .. h 4_t _ U'% ) J ~ ~ z 1 I l 1 l l l I I I l I l 1 l ^ \\ -m g I I l l 1 l I lt I l e JL \\/ l 1 \\ l Figure 5-5. Loads Acting on t e " :e1 a ae Governing Lecation

.4

.ast niiseo ie m

O .1.e.e j d an PIPE OD=14.99 T=1.250 SICY=24.6 SIGU=79.8 Fa=258. -M:.212E+04 1 i Figure 5-6. Critical Flaw Si:e Precietien for Sequoyah Unit 1 Node 1020 Case 0 ) 4m.n n s.o ia

5. g i

j

,a.c.e Ib PIPE OD=14.99 T=1.250 SICV=24.6 SICU=79.8 Fa=258. Ms.214E+94 Figure 5-7. Critical Flaw Size Prediction for Sequoyah Unit 1 Node 1020 Case E .m.nns o i. 5 16

Q C .1 e C t E ,J PIPE OD=14.OG T=1.250 SIGY=34.2 SIGU=83.3 Fa=56.1 M=.399E+94 4 Figure 5-8 Critical Flaw Si:e Prediction for Secuoyah Unit 1 Node 1020 Case F 4 ir.nusea io 5, g

.. ~ _. a.c.e f 'i leN0DE;.020(SKG)00ADCASEG PIPE OD=14.GG T=1.259 SICY=37.9 SICU=83.4 Fa=61.8 -M=.43GE+94 Critical Flaw Si:e Prediction for Sequoyah Unit 1 Figure 5-9 Node 1020 Case G .m.n um io 5 1g

a.c.e 4 in 7 P!PE OD=14.GG T=1.259 SIGY=24.6 SIGU=79.8 Fa=254. pt=.169E+94 .4 Figure 5-10 Critical Flaw Si:e "rediction for Sequoyah Unit 1 Nede 1C80 Case O stSPs/t t ilM 10

• .{g

1 a,c.e ) b PIPE OD=14.OG T=1.250 SICY=24.6 SICU=79.8 Ta=253. M= 163E+04 Figure 5-11 Critical Flaw Si:e P ediction for Sequoyah Unit 1 Node 1080 Case E . m in o m io 5-20

.I a,c.e e l t d l 2 IB P1PE OD=14.99 T=1.259 SICY=39.2 SICU=81.1 Fa=58.4 M=.227E+94 ~ Figure 5-12 Critical Flaw Si:e Prediction for Sequoyah Unit 1 y Nede 1080 Case F .m.m me io 5 31 _________________.______.m_

a.c.e I 1 l b I l PIPE OD=14.99 T=1.250 SICY=31.9 SICU=81.6 Fa=63.3 Mm.357E+94 Figure 5-13 Critical Flaw Si:e Prediction for Sequoyah Unit 1 Node 1080 Case G .munis,o to

4 e a a.c.e i l l l ( e <J si PIPE OD=14.99 T=1.259 SICY=25.1 SICU=79.5 Fa=258. M=.212E+94 1 Figure 5-14 Critical Flaw Size Prediction for Sequoyah Unit 2 Node 1020 Case 0 . si.n ni.o io 5-23 .-n

a.c.< r e .l( J i P1PE OD=14.99 T=t.259 SICV=25.1 S1C0=79.5 Fa=258. M=.214E+94 Figure 5-15. Critical Flaw Si:e Preci: tion for Sequoyah Unit 2 Nede 1020 Case E .w.iw w - 5.za ri.......

4 i .i.c.e 1 3 l-I ib d FIPE OD=14.99 Tui.259 SICYz34.8 SIGU=82.9 /a=56.1 Ms.399E+94 Figure 5-16. Critical Flaw Si:e Prediction for Sequoyal Unit 2 Node 1020 Case F .m.n ni.o io 5 25

i i I i .2. 0. e i e ll PIPE OD=14.99 T=1.259 SIGY=38.7 SICU=83.9 Fa=61.9 l M:.439E+94 Figure 5-17 Critical Flaw Si:e Precietion for Sequoyah Unit 2 Nede 1020 C ts G .m.m isiv o 5 26

i 2 1 s e a.c.e I l t e l d P!PE OD=14.Se T=1.258 5 ICY =25.1 SICU=79.5 Fa=254. M=.169E+94 I Figure 5-1B Critical Flaw Si:e reci: tion for Sequoyah Unit 2 Note 1:50 Case D estti"'isso ie g,g7

7. A A 4 a.c.e l l I 1 J i l PIPE OD=14.99 T=1.258 SICY=25.1 SICU=79.5 Fa=253. l M=.163E+94 [ L l 1' Figure 5-19 Critical Flaw Si:e Preci:: ion for Sequoyah Unit 2 Node 1050 Case E mv.n owo 5-28

i 9 i g d.;.e w e I i J i PIPE CD=14.99 T=1.259 SICV=39.8 STCU=SO.7 Ta=58.4 Pt=. 2 27 E+ 94 t figure 5-20 Critical Flaw Si:e Frediction for Sequoyah Unit 2 Nede 1080 Case F 446?6'111640 10 3 29 I ' ~ " ,-.k-_.,

e 1 a.c.e i 't i l o PIPE OD=14.99 T=1.250 SICV=31.6 SICU:01.2 Fa=63.3 M=.357E+94 Figure 5-21 Critical Flaw Si:e :recietion for Sequoyah Unit 2 Ncce 1050 Case 3 .iit.,o win o 3,3) 4

SECTION

6.0 ASSESSMENT

Of FAT!GUE CRACK GROWTH 5.1 Introduction To determine the sensitivity of the pressuri:er surge line to the presence of small cracks when subjected to the transients discussed in WCAP-12777, fatigue crack growth analyses were performed. This section summari:es the analyses and results, figure 6-1 presents a general fitw diagram of the overall process. The methodology consists of seven basic steps as shown in figure 6-2. Steps 1 through 4 are discussed in WCAP-12777. Stecs 5 through 7 are specifi: to fatigue crack growth and are discussed in this section. There is presently no fatigue crack growth rate curve in the ASME Code for austenitic stainless steels in a water environment. However, a great deal of work has been done recently which supports the development of such a curve, e An extensive study was performed by the Materials Prnnerty Council Working Group on Reference fatigue Crack Growth concerning the crack growth behavior ~ of these steels in air environments, published in reference 6-1. A reference curve for stainless steels in air environments, based on this work, is in the 1989 Edition of Section XI of the ASME Code. This curve is shown in figure 6-3. A compila6 ion of data for austenitic stainless steels in a PWR water environment was mace by Bamford (reference 6-2), and it was found that the effect of the environment on the crack growth rate was very small. For this reason it was estimated that the environmental factor should be set at 1.0 in the crack growth rate equation from reference 6-1. Based on these works (references 6-1 and 6-2) the fatigue crack growth law used in the analyses is as shown in figure 6-4 e 4 h h' ,-n----

6.2 Initial Flaw Size o Various initial surface flaas aere assumeo te exist. The flaws aere assumed to be temi-elliptical with a six-to-one aspect ratio. The largest initial flaw assumed to exist was ene with a depth ecual to 10% of the nominal wall thickness, the maximum flaw size that could te found acceptable by Section XI of the ASME Code. 6.3 Results of FCG analysis Fatigue crack growth analyses were performed at the reactor coolant loop nozzle junction at 10catien 1 (which ecreeseends to the highest usage facter in the surge line) and at location 2 as shown in Figure 6-5. Location 2 corresponds to the location of highest ASME Section 111 equation 12 stress. Results of the fatigue crack growth analysis are presented in table 6-1 for an initial flaw of 10% nominal wall thickness. Conservatisms existing in the fatigue crack gr:wth analysis are listed beloa. l. Plant operational transient cata nas shown that the conventional cesign transients contain significant conservatisms (2. 3. ja,c e 4 Fatigue crack growth calculations are based conservatively on elastic stresses 5. FCG neglects fatigue life Orier to initiation l .m n.no g.2

6.4 References s 6-1. James, L. A. and Jones, D. P., " Fatigue Crack Growth Correlations for Austenitic Stainless Steel in Air," in Predictive Capabilities in Environmentally Assisted Crackino, ASME publication PVP-99, December 1985. 6-2. Bamford, W. H., " Fatigue Crack Greath cf Stainless Steel Lesctor Coolant Piping in a Pressurized Water Reactor Environment," ASME Trans. Journal of Pressure vessel Technology, Feb.1979, e 4 l ..v.,i, nu 1e 33

TABLE 6-1 FATIGUE CRACK GROWTH RESULTS FOR 10% of WALL INIT!AL FLAW SIZE Initial Initial Final (40 yr) Final Flew Location Position Size (in) (% Wall) Size (in) (f.Wal) a.c.e 9 4 an-- 9 .sp.a sisee ie 6-4

4 OITERMINATION CF Twi EF8tCT5 0F THERM AL STR ATIFICATION a c.e e i i e f I i i i e i I l I i f a i I i i i I 4 4 Figure 5-1 Determination of tre ~f'e::s Of Thermal Stratifi:ati:n :n Fatigue Crack Gr:wtn .sem:nec 4e g,.

8 a.:,e l i l l I k l 1 4 l e 1 Fatigue Crack Prewth MethW ology U e 6-2 e l l . i,,,n am ie 56

9 30s,04 i/ //H j l/ jl/

• ~ ~ ~ ' ~ * *

,e i ii! / "l,' 1 7 .' ; i lll I/ fi I i i i/ /} / / ~ i !/ //// c ! !/ /// / / // g6-a ap ies ll ll',! } f fi // 2 I/ ~!/ r/ o 2 1/ / // / l / / //// // I lf } j -- n. s o ,rs l l- 'l, l / / /,' l 2 / ' // / / f j r/ l l / /) / l l /// l {/ { t r J / / / // / ,rr z... IN suuJm o Figure 6-3 Fatigue Crack Growth Rate Curve for Austenitic Stainless Steel 4:n.ncsin io 5-7 t' .~

e 3 h=CFSEaK.30 where h = Crack Growth Rate in inches / cycle -20 C: = 2.42 x 10 t F = Frequency factor (F = 1.0 for temperature below 800*F) S = R ratio correction (S = 1.0 for R = 0; 5 = 1 + 1.8R for 0 < R <.8; and S = -43.35 + 57.97R for R > 0.8) E -= Environmental Factor (E = 1.0 for PWR) e 4K = Range of stress intensity factor, in psi /in R = The ratio of the minimum K (K! min) to the maximum K g (K,,,). g g i l l Figure 6-4 Fatigue Crack Growth Ecuation for Austenitic Stainless Steel l l <sn.nonso no 6-8 .=..:,..,-.---...

a. 2 _h. __._a__, m.A, _u.u_~ _i,__.me. ~ ,.m.a n i i LOC.2 Loc.1 _.e PDES$URIZER mr LEc, o-il N l 4 figure 6-5. fatigue Crack Greath Critical Locations . p emi." 6-9

SECTION 7.0 l ASSESSMENT OF MARGINS 3 In the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessment aere performed. Margins at the I critical locations are summarized below: In Secten 5.3 using the IWB 3640 approach (i.e. "2" factor approach), the " critical" flaw sizes at the governing locations are calculated. In Section 5.2 the crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for the critical locations are calculated. The leakage size flaws, the instability flaws, and margins are given in Tables 7-1 and 7-2. The margins are the ratio of instability flaw to leakage flaw. The margins for analysis combination cases A/0, ( Ja,c.e well exceed the factor of 2. The margin for the extremely low probability eventdefinedby( Jac.e has also exceeded the factor of 2, As stated in Section 4.3, the probability of simultaneous occurrence of SSE and maximum stratification due to shutdown because of leakage is estimated to be very low. l In this evaluation, the leak before-break methodology is applied conser_vatively. The conservatisms used in the evaluation are summarized in Table 7-3. k i l .m.mmvo 71 ~+,wr~ .c u - - ~ +, -p, n,-,--.+r-r- ~g, -w v~--v-~ ~,,rer~ ~. wv v

l 4 TABLE 7-1 n Leakage Flaw Sizes. Critical Flaw Sizes and Margins for Sequoyah Unit 1 -l Lead Critical Flaw Leakage Flaw Node Case Size (in) Size (in) Marcin 1020 A/D 14.60 3.80 3.84 a,c.e A/F B/E B/F a C/G B/Ga ' 1080 'A/D 15.08 S.55 2.72 a,c.e A/F B/E B/F C/Ga B/Ga a These are judgcd to be low probability events mr,muo io 7-2

TABLE 7-2 4 3 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Sequoyah Unit 2 Lead Critical Flaw Leakage Flaw Nede Case Si:e (in) Si:e (in) Marcin 1020 A/D 14.61 3.70 3.94 a,c.e ~ 1080 A/D 15.09 5.50 2.74 -' a, c. e a These are judged to be low probability events 4 k 4437e et 1209010 73 .r -m. _,_,.v.... e--.. 1-,-r,.-.,, ,y v-- 3

---

m-'---y..-

TABLE 7-3 n LBB Conservatisms o Factor of 10 on Leak Rate o factor of 2 on Leakage Flaw for all :sses o Algebraic Sum of loads for Leakage o Absolute Sum of Loads for Stability o Average Material Properties for Leakage o Minimum Material Properties for Stability 9 a . m inis u to 74

l SECTION 8.0 CONCLUSION 5 Thisreportjustifiestheeliminationofpressuritersurgelinepipebreaksas the structural design basis for Sequoyah Units 1 and 2 as follows: a. Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls en reactor coolant chemistry, temperature, pressure, and flow during normal operation, b. Water hammer should not occur in the RCS piping (primary loop and the attached class 1 auxiliary lines) because of system design, testing, and operational considerations, n c. The effects of low and high cycle fatigue on the integrity of the surge line were evaluated and shown acceptable. The effects of thermal stratification were evaluated and shown acceptable. d. Ample margin exists between the leak rate of small stable flaws and the criterion of Reg. Guide 1.45, e. Ample margin exists between the small staole flaw sizes of item d and the critical flaw size. f.' With respect to stabili.ty of the reference flaw, c.mple margin exists between the maximum postulated loads and the plant specific maximum f aulted loads. The postulated reference flaw will be stable because of the ample margins in d, e and f and will leak at a detectable rate which will assure a safe plant shutdown. Based on the above, it is concluded that pressurizer surgo line breaks should not be considered in the structural design basis of Sequoyah Units 1 & 2. l .m.m a.o a g.1 L

r APPENDIX A t.1MIT MOMENT es b 6 ( * .sp nnSoo to b'I e e g n--

APPENDIX A 1 LIMIT MOMENT [ <4 Ja,c.e J ..o.,m i.n o A-2

uni i gi u v e e V e I I O = m C W '0 L U ei.e

• E, m5 Ow 5

.t Aw . au L @a Sud I h W g f =. O = A-3 l l -.. _ _ _ _ _}}