ML20066J566
| ML20066J566 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 01/31/1991 |
| From: | Bhowmick D, Swamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19310D162 | List: |
| References | |
| WCAP-12826, NUDOCS 9102040249 | |
| Download: ML20066J566 (106) | |
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BESTINGHOUSE CLASS 3 l l WCAP-12826 TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP P!PE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE JOSEPH W. FARLEY UNITS 1 AND 2 NUCLEAR POWER PLANTS JANUARY 1991 F. J. Witt D. C. Bhowmick S. A. Swamy C. C. Kim Y. S. Lee VERIFIED: N 08 <b4j GC.Schmertz f' . f: APPROV [ " --r-^ D. C. Adamonis, Acting Manager Structural Mechanics Technology WESTINGHOUSE ELECTRIC CORPORATION Power Systems Division P. O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 e 1991 Westinghouse Electric Corp. a300.* 011051 10
1 FOREWORD This document contains Westinghouse Electric Corporation preorietary information and data which has been icentified by brackets. Coding associated with the brackets sets forth the basis on which the information is considered proprietary. These codes are listed with their meanings in WCAP-7211. The preprietary information and data contained in this report were obtained at considerable Westinghouse exoense and its release could seriously affect our competitive position. This information is to be withheld from public , disclosure in accordance with the Rules of Practice 10 CFR 2.790 and the information presented herein be safeguarded in accordance with 10 CFR 2.903. Withholding of this information does not adversely affect the public interest. This information has been provided for your internal use only and should not be released to persons or organizations outside the Directorate of Regulation and the ACRS without the express written approval of Westinghouse Electric Corporation. Should it become necessary to release this information to suen persons as part of the review procedure, please contact Westinghouse Electric i Corporation, which will make the necessary arrangements required to protect the Corooration's proprietary interests. The proprietary information is deleted in the unclassified version of this report (WCAP-12826). l l l l an mmo 4$$
l TABLE OF CONTfNTS Section Title Pa;e EXECUTIVE
SUMMARY
xiii
1.0 INTRODUCTION
1-1 1.1 Purpose 1-1 1.2 Scope 1-1 1.3 Objectives 1-1 1.4 Background Information 1-2 1.5 References 1-3 l 2.0 OPERATION AND STABILITY OF 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 References 2-4 3.0 PIPE GEOMETRY AND LOADING 3-1 3.1 Introduction to Methodology 3-1 3.2 Calculation of Leads and Stresses 3-2 3.3 Loads for Leak Rate Evaluation 3-2 3.4 Lead Combination for Crack Stability Analysis 3-3
- 3.5 References 3-4 4.0 MATERIAL CHAP.ACTERIZATION 4-1 4.1 Primary Loop Pipe and Fittings Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-2 4.4 References 4-4 5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA 5-1 5.1 Critical Locations 5-1 5.2 Fracture Criteria 5-2
..... .i in y
1 ) TABLE OF CONTENTS (Cont'd.) 1 q Section Title Page i I 6.0 LEAK RATE PREDICTIONS 6-1 6.1 Introduction 6-1 ' 6.2 General Considerations 6-1
, 6.3 Calculation Method 6-1 6.4 Leak Rate Calculations 6-2 6.5' References 6-3 7.0 FRACTURE MECHANICS EVALUATION 7-1 [
7.1 Local Failure Mechanism 7-1 7.2 Global Failure Mechanism 7-2 7.3 Results of Crack Stability Evaluation 7-3 7.4 References 7-4 8.0 FATIGUE CRACK GROWTH ANALYSIS 8-1 8.1
Reference:
8-2 9.0 ASSESSMENT OF MARGINS 9-1
10.0 CONCLUSION
S 10-1 APPENDIX A
- Limit Moment A-1 APPENDIX B - Alternate Toughness Criteria for the Farley Units 1 B-1
. and 2 Cast Primary Loop Components
- I
- B.1 Introduction- B-1 B.2 Chemistry-and KCU Toughness B-1 B.3 Alternative. Toughness Criteria for the Farley B-1 Primary Loop Material on a Component by Component Basis B.4 References B-2
.onm to y3
LIST OF TABLES Table Title Pace 3-1 Dimensions, Normal Loads and Normal Stresses for farley Unit 1 35 3-2 Dimonsions, Normal loads and Normal Stresses for Farley Unit 2 3-6 3-3 Faulted Loads and Stresses for Farley Unit 1 3-7 3-4 Faulted Loads and Stresses for Farley Unit 2 3-8 4-1 Measured Tensile Properties for Farley Unit 1 4-5 Primary Loop Piping and Fittings 4-2 Measured Tensile Properties for the Primary Loop 4-6 Piping and Fittings (all SA351 CFBA) of Farley Unit 2 4-3 Mechanical Properties for the Farley Unit 1 Materials 4-7 at 544*F and 611'F 4-4 Mechanical Properties for the Farley Unit 2 Materials 4-8 (SA351 CF8A) at 544'F and 611*F 4-5 Enveloped Fracture Toughness Properties for Farley 4-9 Units 1 and 2 Primary Loops for Leak-Before-Break Evaluation 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the 6-4 Four Locations 4
..... .o mei in y$$
i LIST OF TABLES (Cont'd)
; Table Title Pace 7-1 Stability Results for Farley Units 1 and 2 Based on 7-6 Elastic-Plastic J Integral Evaluations 7-2 Stability Results Based on Limit Lead Analyses 7-7 8-1 Summary of Reactor Vessel Transients B-4 8-2 Typical Fatigue Crack Growth at ( 8-5 Ja.c.e (40 Years)
B-1 Chemistry and Fracture Toughness Properties of the B-3 Material Heats of Farley Unit 1 B-2 Chemistry and Fracture Toughness Properties of the B-12 Material Heats of Farley Unit 2 e l anos. ine o io y$g4
1 LIST OF FIGURES 4 Fioure Title Pace 3-1 Het Leg Coolant Pipe of Farley Unit 1 3-9 3-2 Schematie Diagram of the Farley Plants RCL Showing 3-10 Weld Locations 4-1 Average True Stress-True Strain Curve for the 4-10 SA351 CFBA Material of Farley Unit 1 at 544'F 4-2 Lower Bound True Stress-True Strain Curve for the 4-11 SA351 CFBA Material of Farley Unit 1 at 544'F 4-3 Average True Stress-True Strain Curve for the 4-12 SA351 CFBA Material of Farley Unit 1 at 611'F 4-4 Lower Bound True Stress-True Strain Curve for the 4-13 SA351 CF8A Material of Farley Unit 1 at 611*F 4-5 Average True Stress-True Strain Curve for the 4-14 SA351 CF8M Material of Farley Unit 1 at 544*F 4-6 Lower Bound True Stress-True Strain Curve for the 4-15 SA351 CFBM Material of Ferley Unit 1 at 544*F 4-7 Average True Stress-True Strain Curve for the 4 16 SA351 CF8M Material of Farley Unit 1 at 611'F 4-8 Lower Bound True Stress-True Strain Curve for the 4-17 SA351 CFSM Material of Farley Unit 1 at 611'F 4064s 122e4310 g
LIST OF FIGURES (Cont'd) Figure Title Pace 1 4-5 Average T se Stress-True Strain Curve for the 4-16 SA351 CFBA Material of Farley Unit 2 at 544'F 4 10 Lower Bound True Stress-True Strain Curve for the 4-19 SA351 CF8A Material of Farley Unit 2 at 544*F 4-11 Average True Stress-True Strain Curve for the 4-20 SA351 CFBA Material of Farley Unit 2 at 611*F 1 4-12 Lower Bound True Stress-True Strain Curve for the 4-21 SA351 CFBA Material of Farley Unit 2 at 611'F 1 4-13 J vs. aa for SA351 CF8N Cast Stainless Steel at 600'F 4-22 l 4-14 J vs. aa at Different Temperatures for Aged Material 4-23 [ Ja.c.e (7500 Hours at 400*C)
- 6-1 Analytical Predictions of Critical Flow Rates of 6-5 l Steam-Water Mixtures 6-2 ( Ja,c.e Pressure Ratio as a Function 6-6 of L/0 l l
l l 6-3 Ideali:ed Pres >ure Drop Profile Through a Postulated 6-7 r Crack l 7-1 ( lC Stress Distribution 7-8 l 7-2 Critical Flaw Size Prediction - Hot leg at Location 1 7-9 for Farley Unit 1 l m ..in m io ,
l LIST OF FIGURES (Cont'd) Ficure t Ti,tle Pace 7-3 Critical Flaw Size Prediction - Hot Leg at Location 2 7-10 for Farley Unit 1 7-4 Critical Flaw Size Prediction - Hot leg at location 3 7-11 for Farley Unit 1 ' 7-5 Critical Flaw Size Prediction - Hot leg at Location 1 7-12 for Farley Unit 2 8-1 Typical Cross-Section of ( Ja,c,e 8-6 8-2 Reference Fatigue Crack Growth Curves for ( 8-7 ja.c.e 8-3 Reference Fatigue Crack Growth Law for ( c- Ja,c.e g.g in s Water Envirenment at 600'F A-1 Pipe with a Through-Wall Crack in Bending A-2 aspes.12266010 j
EXECUTIVE
SUMMARY
The existing structural design basis for the reactor coolant systems of the Joseph M. Farley Units 1 and 2 nuclear reactor power plants re vires that the dynamic effects of pipe breaks ce evaluated and that protective measures for such breaks be incorporated into the design. However, within the last decade, such breaks have been shown to be highly unlikely and shculd not be included. in general, in the structural design basis of Westinghouse type pressuri:ed water reactors, for example. To eliminate primary loop pipe creaks from the design basis, it must be demonstrated to the satisfaction of tne U.S. Nuclear Regulatory Commission that a leak-before-break situation exists. This recort provides such a demonstration for the Joseph M. Farley Units 1 anc 2 nuclear power plants. In this report it is shown that the primary loops are highly resistant to stress corrosion cracking and high and low cycle fatigue. Water hammer is mitigated by system design and operating procedures, the primary loops were extensively examined. The as-built geometries for the
-pipe and elbows and loadings were obtained. The materials were evaluated using the Certified Materials Test Reports. Mechanical properties were determined at operating temperatures. Since the piping systems are faericated from cast stainless steel, fracture toughnesses considering thermal aging were determined for each heat of material.
Based on loading, pipe geometry and fracture toughness considerations, enveloping critical locations were determined at which leak-before-break crack stability evaluations were made. Through wall flaw sizes were found which would leak at a rate of ten times the leakage detection system capabilities of the plants. Large margins in such flaw sizes were shown aga. inst flaw instability. Finally, fatigue crack growth was shown not to be an issue for the primary loops, l it is concluded that dynamic effects of reactor coolant system primary loop pipe breaks need not be considered in the structural design basis of the Joseph M. Farley Units 1 and 2 nuclear power plants.
.sm one io 3444
SECTION 1.0
, INTRODUCTION i J
1.1 Purpose This report applies to the Joseph M. Farley Nuclear Power Plant Units 1 and 2 (Farley) Reactor Coolant System (RCS) primary loop piping. It is intended to 1 demonstrate that for the specific parameters of the Farley plants, RCS primary loop pipe breaks need not be considered in the structural design basis. The approach taken has been accepted by the Nuclear Regulatory Commission (NRC) (reference 1-1). 1.2 Scope
-The existing structural design basis for the RCS primary loop requires that i
dynamic effects of pipe breaks be evaluated. Specifically, the LOCA design basis for the Farley plants includes eleven breaks postulated in the RC5
- primary loop pioing
- the six terminal ends in the cold, hot, and crossover legs; a split in the steam generator inlet elbow, the loop closure weld in the crossover leg; and the nozzle welds for the three large branch lines
'(accumulator, residual heat removal, and surge lines). However, Westinghouse has demonstrated on a generic basis that RCS primary loop pipe breaks are highly unlikely and should not be included in the structural design basis of destinghouse plants (ses reference 1-2). In order to demonstrate this applicability of the generic evaluations to the farley plants, Westinghouse has performed a fracture meenanics evaluation, a determination of leak rates from a through wall crack --a fatigue crack growth evaluation, and an assessment of margins against crack instability consistent with the i leak-before-break (LBB) methodology. Through this successful application of I 'the LBB methodology, the eight break locations in the RCS primary loop piping t-
.(the-branch lines have not been included in this evaluation) are eliminated frem Farley's structural design basis.
1.3 Objectives la order to validate the elimination of RCS primary loop pipe breaks for the i-Farley piants, the following objectives must be acnieved: :
.m..c o mi io 11
I
- a. Demonstrate that margin exists between the critical crack size and a postulated crack which yields a detectable leak rate. .
- b._ Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the Farley plants.
l
- c. Demonstrate margin on applied load. l I
- d. Demonstrate that fatigue crack growth is negligible.
l 4 1.4 Background Information Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated frcm the structural design basis of all Westinghouse plants. The concept of eliminating pipe breaks in the.RCS primary loop was first presented to the NRC in 1978 in WCAP-9283 ; (reference 1-3). That topical report employed a deterministic fracture ' 3 mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. That approach was then used as a means of addressing Generic issue A-2 and Asymmetric LOCA Loads. d Westinghouse performed additional testing and analysis to justify the
- elimination of RCS primary loop pipe breaks. This material was provided to the NRC along with Letter Report NS-EPR-2519 (reference 1-4).
e The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to address this same issue using a probabilistic approach. As part of the
. LLNL research effort, Westinghouse performed extensive evaluations of specific
- plant loads, material properties, transients, and system geometries to demonstrate that the analysis and testing previously performed by Westinghouse and the research performed by LLNL applied to all Westinghouse plants (references 1-5 and 1-6). The results from the LLNL study were released at a March 28. 1983 ACRS Subcommittee meeting. These studies which are applicable to all Westinghouse plants east of the Rocky Mountains determined the mean probability of a direct LOCA (RCS primary loop pipe break) to be 4.4 x 10 -12
. n no 12
per reactor year and the mean probability of an indirect LOCA to be 10'7 per reactor year. Thus, the results previously cbtained by Westinghouse (reference 1-3) were confirmed by an independent NRC research study. Based on the studies by Westinghouse, LLNL, the ACRS, and the AIF, the NRC completed a safety review of the Westinghouse reports submitted to address asymmetric blowdown loads that result from a number of discrete break locations on the PWR primary systems. The NRC Staff evaluation (reference 1-1) concludes that an acceptable technical basis has been provided so that asymmetric blowdown loads need not be considered for those plants that can demonstrate the applicability of the modeling and conclusions contained in the Westinghouse response or can provide an equivalent fracture mechanics demonstration of the primary coolant loop integrity. In a more formal ' recognition of LBB methodology applicability for PWRs, the NRC appropriately modified 10 CFR 50, General Design Criterion 4 " Requirements for Protection Against Dynamic Effects for Postulated Pipe aupture" (reference 1-7). This report provides a fracture mechanics demonstration of primary loop integrity for the farley plants consistent with the NRC position for exemption from consideration of dynamic effects. Several computer codes are used in the evaluations. The main-frame computer programs are under Configuration Control which has requirements conforming to Standard Review Plan 3.9.1. The fracture mechanics calculations are independently verified (benchmarked). 1.5 References ! 1-1 USNRCGenericletter84-04,
Subject:
" Safety Evaluation of Westinghouse .
Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary Main Loops," February 1, 1984, i l 1-2 Letter from Westinghouse (E. P. Rahe) to NRC (R. H. Vollmer), i NS-EPR-2768, dated May 11, 1983. 1-3 WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse ; Nuclear Power Plants During Postulated Seismic Events," March, 1978. j
.an. onni in 33 J
3
1-4 Letter Report NS-EPR-2519 Westinghouse (E. P. Rahe) to NRC (D. G. 3 Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.
- 1-5 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983.
1-6 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July , 25, 1983. ' 17 Nuclear Regulatory Commission, 10 CFR 50, Modification of General Design Criteria 4 Requirements for Protection Against Dynamic Effects of Postulated Pipe Ruptures, Final Rule Federal Register / Vel. 52, No. 207/ Tuesday, Oct:bar 27, 1987/ Rules anc Regulatiens, pp. 41288-41295. .) l 1 l l I 4464e = Il296010 g.g
I i l j SC. TION 2.0 j OPERATION AND STAB!LITf Or THE REACTOR COOLANT SYSTEM t j 2.1 Stress Corresien Crackino The Westinghouse reactor coolant system primary loeps have an operating 3 history that demonstrates the inherent operating stability characteristics of j the design. This includes a low susceptibility to cracking failure from the l effects of corrosion (e.g., intergranular stress corrosion cracking). This 4 operating history totals over 450 reactor years, including five plants each l having over 17 years of eperation and 15 other plants each with over 12 years l of operation. I j 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 St.;dy Group (PCSG) was to include a review l 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 alnvestigation and Evaluation of Stress t
Corrosion 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 PWR primary system piping is extremely low because the ingredients that produce IGSCC are not all present. The use of hydrazins 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. Operating experience in PWRs supports this determination. To dt.te, no stress- corrosion cracking has been reported in the primary piping or safe ends of any PWR."
. ..oom ie 2-1
_ , _ . _ . . _ . . . . ..__.2-. ___
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 IGSCC 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. 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 envirenment. Since some residual stresser and some
- degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimi:ed 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.
L
- The elements of a water environment known to increase the susceptibility of austenitic stainless steel to stress corrosion are
- oxygen,' fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., -
sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to i 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 i 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 pioing. During plant operation, the reactor coolant water chemistry is monitored and ! _ maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the major water chemistry control standards being included in the plant operating procedures as a condition for plant operation. For example, during e646 sal 3240010 g ,, g i x,,,.,,,,~ + , w- -,,,n--- v-, n--,-a-,,-~~---w-----+n
I normal power operation, oxygan concentratien in the RCS is expected to be in the ppb range by controlling charging flow chemistry and maintaining hydrogen in the reacter coolant at specified concentratiens. Halogen concentrations are also stringently controlled by maintaining concentrations of chierides and fluorides within the specified limits. Thus during plant operation, 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 since it is designed and operated to preclude the voiding condition in ncrmally filled lines. The reactor coolant system, including piping and primary components, is designed for normal, upset, emergency, and faulted condition transients. The design recuirements are conservative relative to both the numoer 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 leads. To ensure dynamic system stability, reactor coolant parameters are stringently centrolled. Temperature during normal Operation is maintained within a narrow range cy control roc position; pressure is controlled by pressuri:er heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteris-tics 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. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system. Preoperational testing and operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer Can occur.
*see, 't soo is 2-3
-- - ---- - - -- .. - . - ... - - - - - . - . - - .~ - - . .~.-.-
a 7 4 2.3 Low Cycle and High Cycle Faticut 1 Low cycle fatigue considerations are acccunted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the
- rules of Section 111 of the ASNE Code. A further evaluation of the icw cycle fatigue loadings was carried out as part of this study in the form of a fatigue crack growth analysis, as discussed in section 8.0 H @ cycle fatigue loads in the system would result primarily from pump '
vibrations. These are minimized by restrictions placed on shaft vibrations during hot functional testing and operation. During operation, an alarm signals the exceedence of the vibration limits. Field measurements have been made on a number of plants during hot functional testing, including plants-similar to the Farley Units 1 and 2. Stresses in the elbow below the reactor coolant' pump resulting from system vibration have been found to be very small, between 2 and~3 ksi at the highest. These stresses are well below the fatigue
- endurance limit for the material and would also result in an applied. stress intensity factor below the threshold for fatigue crack growth.
2,4 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 Reacto'rs, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980. i 2., l d
~or-v ,w m--r e s w,~we, --,-+.-a- ---v, -,.,-.mm-y,n-,,,,,..ewm.w ,,w-,--me, e,,-,,-,- -e .-e, , ~ , , ,, g,,m-n,,- ,. ,w, w, w m : , 4 -,-ww,,-e-- I
l SECTION 3.0 PIFE GEOMETRY AND LCADING 3.1 Introduction to Methodoloey ' The general approach is discussed first. As an example a segment of the primary coolant hot leg pipe of Farley Unit 1 is shown in figure 3 1. The as-built outside ciameter and minimum aall thickness of the pipe are 33.78 in, and 2.28 in., respectively, as seen in the figure. Nermal stresses at the weld locations result from the load combination procedure discussed in section 3.3 while faulted lead are developed as outlined in section 3.4. The i components for normal loacs are pressure, dead weight and thermal exN.' An additional component, Safe Shutdown Earthouake (SSE), is consid faulted loads. As seen later the highest stressed location in th* sep l is at the reactor vessel outlet noz:le to pipe weld. This location ind the load critical location and is one of the locations at which, as an enveloping location, leak-before-break is to be established. Essentiall33 circumferential flaw is postulated to exist at this location thus the r . loads and faulted loads must be available to assess leakage and stability, respectively. The loads (developed below) at this location fer farley Unit 1 are also given in Figure 3-1. j Since the pipe anc fittings are cast stainless steel, thermal aging must be considered (see section 4.0). Thermal aging results in lower fracture toughness criteria; thus, other locations than the highest stressed location must be examined taking into consideration both fracture toughness and stress. The enveloping locations so determined are called touchness critical locations. The single most critical location is apparent only after the full analyses is completed. Once loads (this section) and fracture toughnesses (section 4.0) are avai.lable, the load critical and toughness critical locatiens are determined (see section 5.0). At these locations, leak rate evaluations (see section 6.0) and fracture mechanics evaluations (see section 7.0) are performed per the guidance of reference 3 1. Fatigue crack growth (see section 8.0) and stability margins are also evaluated (see section 9.0). l cm.. mw o 3,3 1
j j s The locations for evaluation are those shown in figure 3 2. l' l 3.2 Calculatien of Loads and Stresses o l' The stresses due to axial leads and bending moments are calculated b,, the folloaino equation: l I ! o=(+f (3-1) i where, i i o = stress F = axial load M = bending moment
; A- = pipe cross-sectional area F
2 = section modulus i. The bending moments for the desired loading combinations are calet. lated by the following equation: M=/M y2+M 2 (3-2) where, 1 M = bending moment for required leading _ My = Y component of bending moment M = 2 component of bending moment 7 , The axial load and bending moments for leak rate predictions and crack stability analysis are computed by the methods to be explained in sections 3.3 and 3.4. , 3.3 Loads for Leak Rate Evaluation The normal operating loads for leak rate predictions are calculated by the following equations: p.
.. i i .
32
F = F0W + FTH + Fp (3-3) My = (My )DW * (NY )TH + (MY)P (3'4) M Z (NZ )0W + (MZ )TH + INZ)P (30) The subscripts of the above equations represent the following loading cases: DW = deadweight TH = normal thermal expansion P = load due to internal pressure This method of combining loads is often referred as the algebraic sum method. The loads based on this method of combination are provided in tables 3-1 and 3-2 for Farley Units 1 and 2, respectively, at all the locations identified in figure 3-2. The as-built dimensions are also given. 3.4 Load Combination for Crack Stability Analysis in accordance with standard review plan 3.6.3 the absolute sum of leading components can be applied which results in higher magnitude of combined loads. If crack stability is demonstrated using these 1%ds, the LBB margin on loads can be reduced from / 2 to 1.0. The absolute summation of loads results in the following equations: F = lF 0W I + IF TH I+IFl+IF p SSEINERTIA I + IF SSEAM I (3'0) NY ' IIN Y)DWI ^ l("Y)THI
- I(N Y)P !
- l(M )SSEINERTIA Y I
- IIN )SSEAM Y I (3*7)
Mg = l(Mg )DWI
- I(NZ )TH I
- IINZ )P I* ("Z)SSE!NERTIA I
- IIN )SSEAMI Z
(3'3) where subscripts SSE, INERTIA and AM mean safe shutdown earthquake, inertia and anchor motion, respectively. The loads so e termined are used in the fracture mechanics evaluations (section 7.0) to demonstrate the LBB margins at the locations established to be the governing locations. The loads at all the locations of interest (see
)
figure 3-2) are summarized in tables 3-3 and 3-4 for Farley Units 1 and 2, resoectively. l l
... .o m i i.
3-3 i w* -
3.5 References 3-1 USNRC Standard Review Plan 3.6.3, Leak-Before-Break Evaluation Procedurer, NUREG-0800.
.. ..i um in 34
TABLE 3-1 O! MENS 10NS, NORMAL LOADS AND NORMAL STRESSES FOR FARLEY UNIT 1 Outside Bending Diameter Thickness Axiaf Load Moment Stress Location' (in.) (in.) (kips) (in-kips) (ksi) 1 33.78 2.28 1458 25150 21.6 2 33.78 2.28 1456 12519 14.0 3 36.96 2.88 1519 20635 13.4 4 36.76 2.88 1644 6412 8.0 5 36.05 2.42 1630 5168 8.9 6 36.05 2.42 1625 4929 8.8 7 36.05 2.42 1711 7406 7.1 8 36.05 2.42 1711 3091 8.2 9 37.16 2.98 . 1819 8586 9.1 10 32.03 2.16 1369 4427 9.9 11 32.03 2.16 1370 4389 9.9 12 32.03 2.56 1365 4314 8.1 8 See figure 3-2 Includes pressure l l -
..... i zi..o i .
35 , j
i i TABLE 3-2 DIMENSIONS, NORMAL LOADS AND NORMAL STRESSES FOR FARLEY UNIT 2 Outside Bending Diameter Thicknest, Axief Lead Moment Stress location e (in.) (in.) (kips) (in-kips) (ksi) , 3 1 33.81 2.30 1456 25150 21.4 2 33.81 2.30 1456 12519 13.8 3 36.20 2.50 1519 20635 15.6
~
4 35.20 2.50 1644 6412 9.3 5 36.11 2.45 1630 5168 S.8 S 36.11 2.45 1625 4929 8.7 7 36.11 2.45 1711 7406 7.0 8 36.11 2.45 1711 3051 8.1 9 37.52 3.16 1819 8586 8.5 10 32.07 2.18 1369 4427 9.8 11 32.07. 2.18 1369 4389 9,8 12 32.14 2.22 1365 4314 9.5 a See figure 3-2 Includes pressure i l
..... 122 o e 3-6
TABLE 3I 3 FAULTED LOADS AND STRESSE! 'OR FA'll ' UNIT 1 AxialLeaE Bending Moment Stress e Location .b (kips) (in-kips) (ksi) i 1768 30130 25.9 h 2 1765 16101 17.5 3 2072 26015 17.4 ' 4 1937 19498 14.3 5 1895 13837 14.3 6 1890 6889 10.8 7 1840 7274 10.8 x L 8 1834 10206 12.2 9 1896 15486 12.4 10 1576 10665 15.3 11 1550 8615 13.9 12 1539 8659 11.4 See Figure 3-2 b See table 3-1 for dimensions C Includes pressure j m.. . m.m o 37
i TABLE 3-4 FAULTED LOADS AND STRESSES FOR FARLEY UNIT 2 Axial Load C Bending Moment Stress Location ,b a (kips) (in-kips) (ksi) 2 1767 30130 25.7 7 1765 16101 17.3 3 2072 26015 20.3 4 1937 19498 16.7 6 1895 130t? 14.1 6 1890 6889 10.7 7 1840 7274 10.7 8 1834 10206 12.1 9 1896 16486 11.7 10 1576 10665 15.1 12 1550 8615 13,7 12 }539 8689 13.3 a See Figure 3-2 See table 3-2 f:r d mencions C Includes pressure
..... .o i i o. i i s 3-8 l .-- .
- s Crack I n.
t L _. ._ _ - - - i 1 _ _ _ . -_____ l l h % j+ W
)
1 P + I
% - _ - _ - - - . - - - - - - - - - - f) m L 00 33.7s in
- 00
- t = 2.28 in i
a Normal Loads Faulted Leads fcrcea : 1458 kips force": 1768 kips bending mcment: 25150 in-kips bending moment: 30130 in-kips
" Includes the force due to a pressure of 2235 psig Figure 3-1. Hot leg Coolant Pipe of Farley Unit 1 ..... . m e.o i c 39
LJ fy, REACTOR PRESSURE VESSEL
#2 H
COLD WT gg LIG . 3 . 2
. . ,L.
REACTOR COOLANT PUMP STEAM GENERATOR CROSSOYER LES
- -D Hot Leg Temperature: 611'F j +7 @ -
Cold Leg Temperature: 544*F ;, u O G Figure 3-2 Schematic Diagram of the Farley Plants RCL Showing Weld Locations
..a. . ms.o in 3-10
m ,:o 6' ,
~
SECTION 4.0
- w. MATERIAL CHARACTERIZATION 4.1 Primary loco Pipe and Fittinos Materials The primary loop piping material for both Fariey Unit 1 and Farley Unit 2 are SA351 CF8A. The elbow fittings for Farley Unit I are SA?51 *.F8M. while for Farley Unit 2, they are SA351 CFBA. The field welds e,e SMAW following GTAW root passes. The shop welds are SAW.
4.2 Tensile Properties The Certified Materials Tes' Reports (CMTRs) 'or Farley Units 1 and 2 were used to establish the tensile properties for the leak-before-break analyses. The CMTRs include tensile properties at room temperature for each of the heats of material and some tensile properties at 650*F. The properties for the heats of Farley Unit 1 and Farley Unit 2 are given in Tables 4-1 and 4-2, respectively. The average properties are given and the lower bound properties are identified. The 1989 ASME Code minimum on properties are also given. The properties at $44*F and 611*F were established by scaling the plant specific values by the ratio of the ASME Code minimum properties. Preference was given to values at 650'F when available; otherwise, a scaling based on room temperature was made. The average and lower bound yield stresses and ultimate strengthr., so scaled, are given in tables 4-3 and 4-4 for Farley Units 1 and 2, respectively. The ASME Code Modulus of Elasticity is also given at each temperature. Poisson's Ratio was taken as 0.3. The study of critical locations (Section 5.0) showed that the prcperties at only 611'F are < required for the LBB evaluation. The 544'F properties are included for completeness, u... ome, so 41
_ ._- ._ . _ . _ . . _ . _ _ . _ ... _ ..___. .. _ _ _ . .m. .___ -._ . . _ _ . . . _ .
?
For' leak-before-break ~ fracture evaluations the true stress-true strain curves must be available.- Such curves were obtained using.the Nuclear Systems Materials Handbook (reference 4-1). The average and' lower bound true stress-true strain. curves are given for Farley Unit 1 in Figures 4-1 through 4-8. Similar curves for'Farley Unit 2 are given in figures 4-9 through 9-12. Curves at both 544*F and 611*F are presented. 4.3 Fracture Toughness Procerties The pre-service fracture toughness of cast materials in terms of J have been ' found to be vary high at 600'F. Typical results for a cast material are given in figure 4-13 taken from reference 4-2. J;c is observed to be over-5000 in-lbs/in2. However cast stainless steels are subject to thermal aging during service. This thermal aging causes an elevation in the yield strength 4 of the material and a degradation of the fracture toughness, the degree of degradation being somewhat proportional to the level of ferrite in the material. To determine the effects of thermal aging on piping integrity, a detailed study was carrisd out in reference 4-3. In that-report, fracture toughness results were presented for a material ( Ja,c.e The effects of the aging process on the-end-of-service life fracture toughness are further discussed in Appendix B. I ? l l 1 j <m..mwo 4.g l l I
End-of-service life toughness for the heats are established using the alter-nate toughness criteria methodology of reference 4-6 (appendix B). By that methodology a heat of material is said to be as good as [ ]a.c.e if it can be demonstrated that its end-of service fracture toughnesses equal or i exceed thost, of ( . 1 3a,c.e , l The worst case fracture toughness values for all the loops of each plant a+ 1 each location (see figure 3-2), as taken from Appenoix B, are given in table 4-5. All locations for Fcrley Unit 2 qualify for the highest assignable toughness values as discussed above. Only four locations so qualify for Farley Unit 1. The lowest fracturo toughness values for Farley Unit 1 occur at locations 2 and 3. Available data on aged stainless steel welds (references 4-3 and 4-4) indicate that J gg values for the worst case welds are of the same order as the aged material. However, the slope of the J-R curve is stec7er, and higher J-values 2 have been obtained from fracture tests (in excess of 3000 in-lb/in ). The applied value of the J-integral for a flaw in the weld regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at temperature8 . Therefore, weld regions are less limiting than the cast material. It is thus conservative to choose the end-of-service life toughness properties of ( Ja,c.e as representative of those of the welds. Also, such pipes and fittings having an end-of-service life calculated room temoerature charpy U notch energy, (KCU), greater than that of ( ]a,c.e are also conserva-tively assumed to have the properties of ( ]a,c.e , In the fracture mechanics analyses that follow, the fracture toughness properties given in table 4-5 will be used as the criteria against which the applied fracture toughness values will be compared, a in the report all J applied values were conservatively determined by using base metal strength procerties. l j case 01048110 1 43
4.4
References:
4-1 Nuclear-Systems Materials Handbook, Part'I - Structural Materials, Group-1 - High Alloy Steels, Section 2, ERDA Report TIO 26666, November, 1975. 4-2 WCAP-9558 Rev. 2, " Mechanistic Fracture Evaluation of_ Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack," Westinghouse Proprietary Class 2 June 1981. . 4-3 WCAP-10456, "The Effects'of Thermal Aging on the Structural Integrity of . Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2 November 1983. . 4-4 Slama, G., Petroquin, P., Masson, S.H., and Mager, T.R., "Effect of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and , Welds", presented at SMiRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel-Reactor Pressure Boundary Components - August 29/30. 1983, Monter~ey, CA.
.4-5 Appendix 11 of. Letter from Dominic C. Oilanni, NRC to 0. M. Musolf, Northern States Power Company, Occket Nos. 50-282 and 50-306, December 22, 1986.
4-6 ' Witt, F.J. , Kim, C.C., "Toughr.ess Criteria for Thermally Aged Cast Stainless Steel," WCAP-10931, Revision 1, Westinghouse Electric Corporation, July 1986,L(Westinghouse Proprietary Class 2). 4 i l l-
.m.e.i io 44
TABLE 4-1 heEA$t; RED TENSILE PROPf 211ES FOR FARLEY (JNIT 1 PRIEEARY LOOP PIPING AtX) FITTINGS LOOD $ $$ P 8I out m- ma trengtfl pst) tMa t No- MJterial R 3 p [ PIPE m.c.e b e us 4891i/011001:10 l l
..~. ,
[ 1Q81E O-2 ME ASURE D If M5ll E PROPERilES f 0H IllE PR!aAARY t_OUP PIPING ' AND f f111NGS ( Alt. 57351 Cf 8A) Of f ARLEY UNIT 2 toop . Yield Stress-{pst) Ultimate Strength (psil Component- No, lleat No. Room Temp. 650*F at Rotw t emp . PIPE A.c.e i s u I
..y i
s e cn 1
-}
t 4 :
}
}
.y l
t t l t P 4096/t22690:10 ,
TABLE 4-3 NECHANICAL PkOPERTIES FOR fARLEY UNIT l-MATERIALS'AT 544*f AND 611*f
~
Average- Lower Bound lemperature Yield Stress Ultimate Strength Yield Stress Ultimate Strergth Material (*f) (psi) -(psi) (psi)- (psi) SA351 Cf8A d.C.e
- SA3SI Cf8N~ ;
~ , s Nodulus of Elasticity for Poth' Materials: 6 at 544 f. E = 25.6 x 10 p3g 6 ,g at 611*f, E = 25.2 x 10 Poisson's Ratio: 0.3
?
som. a sksses so
--. 4
, , +
+
i l l TABLE 4-4 . MECHANICAL-PROPERTIES FOR FARLEY UNIT.2 MATERIAL (SA351 CF8A) AT-544'F AND 611*F Average' Lower Bound-Temperature. Yield Stress Ultimate Strength Yield' Stress Ultimate Strength ('F) (psi) (psi) (psi) (psi) . I 544~ d'C' 611 Modulus of Elasticity: at 544*F, E = 25.6 x 106p3g at 611'F,.E.= 25,2 x 106 psi Poisson's Ratio:-.0.3
. t-l
. caos.ssasm a so 4.g L,..,_ . ._.
TABLE 4-5 ENVELOPED FRACTURE TOUGHNESS PROPERTIES FOR FARLEY UNITS 1 AND 2 PRIMARY LOOPS FOR LEAK-BEFORE-BREAK EVALVATION U I KCU lc mat max - Locationa ,b 2 2 HT. NO. (daJ/cm ) (in-lb/in ) (non-dim) (in-lb/in') Farley Unit 1
~ ~
y a,0,0
. 2 3
4 5 6 7 8 9 10 11 12 4E-4
. u.,i us.o ie 4.g
- a,c.e l
l 1 Figure 4-1 Average True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 1 at 544'F 84964'122640 10 4-10
a,c.e Figure 4-2: Lower Bound True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit I at 544*F ca..n um in 4.yy
I
~
a,c,e i Figure 4-3: Average True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 1 at 611*F me.n um se a.7g
I l 1 j a,c.e Figure 4-4: Lower Bound True Stress-True Strain Curve for the SA351 CFSA Material of Farley Unit 1 at 611*F l 4 13
a,c.e Figure 4-5: Average True Stress-True Strain Curve for the SA351 CF8M Material of Farley Unit 1 at 544*F
.m.n zaeo io 4 14
~ ,c.e a
1 l 1 I 1 l 1 1 l l l Figure 4-6: Lower Bound True Stress-True Strain Curve for the SA351 CF8H Haterial of Farley Unit 1 at 544*F u..., m.9e ic 4-15
~
a,c.e Figure 4-7: Average True Stress-True Strain Curve for the SA351 CF8M Material of Farley Unit 1 at 611*F .u.. o n o io 4-16
l l 1 i
""~
a,c.e Figure 4-8: Lower Bound True Stress-True Strain Curve for the SA351 CF8H Material of Farley Unit 1 at 611*F
....., u s o i o 4-17
a c.e l l l Figure 4-9: Average True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 2 at 544*F
.n..naruo io 4.ig
i 1 1
~' a,c.e Figure 4-10: Lower Bound True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 2 at 544*F ceu.,iasmo io 4 19
1 l a ,c .e Figure 4-11: Average True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 2 at 611*F u...,i us o io 4-20
'~
a,c.e l l l i I
~
1 l Figure 4-12: Lower Bound True Stress-True Strain Curve for the SA351 CF8A Material of Farley Unit 2 at Sil'F ao.,me.o ia 4 21
1
... a,c.e Figure 4-13: J vs. aa for SA351 CF8M Cast Stainless Steel at 600*F
<ses.,,useo ,e 4_
_ . .c. . Figure 4-14: J Vs. aa at Different Temperatures for Aged Material ( )"'C (7500 Hours at 400*C) ca..e u rm ' i 4-23 l
- l. SECTION 5.0 l
l CRITICAL LOCATIONS AND EVALUATION CRITERIA 5.1 Critical Locations The leak-before-break (LBB) evaluation margins are to b6 demonstrated for the limiting location (governing location). Candidate locations are cesignatea load critical locations or toughness critical locatians as discussed in g Section 3.0. Sud locations are established consicering the 'oads (section 3.0) and the mes. ciel properties established ie section 4.0. These locations are defined below for Farley Units 1 and 2. Tables 3-1 through 3-4 and 4-5 are used for this evaluation along with figu're 3-2. Farley Unit 1 Location 1 is the highest stressed location and is the load critical location by definition. The lowest toughness values are at locations 2 cnd 3 with the loads being about the same. These two locations are thus tcughness critical locations. These locations are now compared with the remaining locations in the crossover leg and cold leg. It is observed that (1) the temperature at [ locations 2 and 3 is higher thus the tonsile properties are worse (2) the stresses at locations 2 and 3 are higher and (3) the fracture toughness.
-J,,,, at locations 2 and 3 is at least a fsetor of three less, the factor '
being over 10 for T It is thus concluded that the enveloping locations mat. in Farley Unit I for which the LSB mnthodology is to be applied ars locations 1, 2' and 3. Farley Unit 2 Location 1 is the highest stressed location and is thus the load critical location. Since this location is at the higner temperature (i.e., has the worst tensile properties) and all locations are assigned the same toughnestes, this location envelopes the other locations and is the only one at wnich LEE evaluations are required.
<sss,n nsso <o 53 l l
5.2 Fracture Criteris As discussed later, fracture mechanics analyses are made based on loacs and postulated flaw sizes related to leakage. The stability criteria against which the calculated J (i.e. J,p3) and tearing modulus (T,pp) are ecmpared are: (1) If J,pp < J7c, ther the crack is stable; t (2) If Jaco 1 UI :, tnen, if T,p 4 T,,; D and Ja :p # Umax, the crack is stable. % The toughness criteria at each 1ccation have previously been determined and are given in table 4-5, F 4...... i s i i o 5-2
SECTIGH 6.0 ii.T! KAii PREDICTIONS
! . .i. introduc 't in The purpose of this snction is to discuss the method which is used to predict the flow through postulated through-wall cracks end present the leak rate Cdlculation results for through-wall circumfsrential cracks.
6.2 Gsneral Consideratiens The flow of hot pressurized water through an opening to a lGwer bar.k presser 0 causes flashing which can result in choking. For long channols share the ratio of the channel length, L, to hydre.ulic diameter, gO , U/0;) is greater than [ la,c.e, both [ t ja.c e , 6.3 Calculajion Method
'lb0 aasic method used in the leak rate calculations is the method develcped by
(
)
The flow rate through i. crack aas calculated in the following manner. Figure
'S-1 frem referetice 6-1 was used t's es'simate the critical pressure, Pc, for the primary loop enthalpy conditt an ar.c ar assumed flow. Once Pc was found for a given mass fiow, trei Ja,c.e was found free. figure 6-2 trkon from reference 6-1. For all cases considered, sines f, JC Therefore, this method will yield m...wni io 6-1
the two-phase pressure drop due to momentum effects as illustrated in figure 6 3. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using APf=(-
)"'" (6-1) where the friction factor f is determined using the ( }"'C
The crack relative roughness, e, was obtained from fatigue crick data on stainless steel samples. The relative roughness value used in these calculations us ( la.c.e The frictional pressure drop using equation 6-1 is then calculated for the l assumed flow and added to the ( 4
}"'C to obtain the total pressure drop from the ;,rimary system to the atmosphere. That is, for the primary loop l
y Absolute Pressure - 14.7 = [ la.c.e (6-2) for a given assumed flow G. If the right-hand side of equation 6-2 does not agree with the pressure difference between the primary loop and the atmosphere, then thre procedure is repeated until equation 6-2 is satisfied to within an acceptable tolerance and this results in the flow value through the crack.
~---
6.4 Leak Pete Calculations Leak rate calculations were made as a function of crack length at the four locations previously identified in section 5.2. TLe normal operating loads of
- tables 3 1 and 3 2 were applied as appropriate, in these calculations. The crack opening areas were estimated using the method-of reference 6-2 and the leak rates were calculated using the two phase flow formulation described above. The average material properties of section 4.0 were used for these calculations.
l
, .. . .. n o. o .
32
l l The flaw sizes to yield a leak rate cf 10 gum were calculated at the four locations and are given in Table 6-1. The flaw sizes so determined are called leakaoe flaws. The farley plant RCS pressure boundary leak detection system, as documented in FSAR Section 5.2.7 and the NRC Safety Evaluation Report Section 5.6, meets the intent of Regulatory Guide 1.45. Thus, to satisfy the margin of 10 on the leak rate, the flaw sizes (leakage flaws) are determined which yi61d a leak rate of 10 gpm. 6.5 References 6-1 [
)t.c.e, 6-2 Tada, H., "The Effects of Shell Corrections en Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section !!-1. NUREG/CR-3464, September 1983, t ... ,e u n i i.
6-3
TABLE 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the Four Locations Unit Location Flaw Site (in) 1 1 3.25 1 2 5.00 1 3 5.50 2 1 3.10 i
.an.>mino io 6-4
- a,c.e 5s
=
5 1 t. 8 w I N 3 1 Figure 6-1. Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures cses.:'um 'o s.5
8 c,e b 9 2 s w 5 w f i w 2 . 5 LENGTH / DIAMETER RATIO (L/D) Figure 6-2. ( Ja,c.e Pressure Ratio as a Function of L/D
..... i nno io 6-6
~
, , a,c.e a.c.e l f w __'
Figure 6-3. Idealized Pressure Orop Profile Through a Postulated Crack
. . ... , i s a. "
6-7
l l i SECTION 7.0 FRACTURE MECHANICS EVALVAT!0N 7.1 Local Failure Mechanism The local mechanism of failure is primarily dominated by the crack tip behavior in terms of crack tip blunting, initiation, extensien and finally crack instability. The local stability will be assumed if the crack dces not initiate at all. It has been accepted that the initiation tc,ghness measured in terms of J;g from a J-integral resistance curve is a material parameter defining the crack initiation. If, for a gi'>en load, +,he calculated J-integral value is shown to be less than the J;g cf the material, then the crack will not initiate. If the initiation criterion is not met, one can calculate the tearing modulun as defined by the following relation: d E T,pp = g) q f whers: T,pp = applied tearing modulus E = modulus of elasticity of = 0.5 (c y u) III
- 5t55) a a crack length ej , u = yield and ultimate strength of the material, respectively Stability is said to exist when ductile tearing occurs if T,pp is less than Tmat, the experimentally deteemined tearing modulus. Since a constant I must be mat is assumed a further restriction is placed in J, p. J, less than J g, where J,,,; is the maximum value of J for which the experimental T is greater than or eaval to the T used.
mat
.nss., m sev o y.1
As discussed in Section 5.2 the local crack stability will be established by the two-step criteria: (1) If K,pp < J;g, then the crack is stable. (2) If J,pp >,,,Jgg, then, if T,pp < T,,g and J,pp < J,,,, the crack is stable. 7.2 Global Failure Meehan_ ism Determination of the conditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the plastic instability method, based on traditional , plastic limit load concepts, but accounting for strain hardening and taking , into account the presence of a flaw. The flawed pipe is predicted to fail when the remaining not section reaches a stress level at which a plastic hinge is formed. The stress level at which this occurs is termed as the flow strass. The flow stress is generally taken as the average of the yield and ultimate tensile strength of the material at the temperature of interest. ! This methodology has been shown to be applicable to ductile piping through a large number of experiments and will be used here to predict the critical flaw . size in the primary coolant piping. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (figure 7-1) when loads are applied. The detailed development is provided in appendix A for a through wall circumferential flaw in a pipe with internal pressure, axial feree, and imposed bending moments. The limit moment for such a pipe _is given by:
'l a,c.e
[ ) l 1 where: ( I 3a,c.e
.. . i u . . .
7-2 l
. . _ . _ , . . . . . _ _ . _ _ - _ __ . - . _ . . .. ~ . - --
(
)a,ce The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental results (reference 7-1).
For ar. plication of the limit lead methodology, the material, including censideration of the configuration, must have a sufficient ductility and ductile tearing resistance to sustain the limit load. 7.3 Results of Crack Stability Evaluation Stability analyses were perforced at the critical locations established in section 5.1. The elastic plastic fracture mechanics (EPFM) J-integral analyses for through wall circumferential cracks in a cylinder were performed using the procedure in the EPRI fracture mechanics handbook (reference 7-2). The lower-bound material properties of section 4.0 were applied (see tables 4-3 and 4-4). The fracture toughness properties established in section 4.3 (see table 4-5) at.d the normal plus SSE icads given in tables 3-3 and 3-4 were used for the EPFM calculations. The flaw sizes were twice those giving a leak rato of 10 gpm as established in section 6.0 ( we table 6-1). Evaluations wore performed at the four critical location; identified in section 5.1. The resJits of the elastic plastic fracture mechanics J-integral evaluations are om.aum so y.3 i l
._,,,.-.-..m. .m._.m_._~_-.~.. -..m given in table 7 1. It is seen that the fracture criteria are met at all the locations. Specifically a margin of 2 on flaw size is demonstrated. Since the faulted load combination method used in this calculation is based on the absolute sum method, the required margin on load of 1.0 is also accomplished as discussed in SRP 3.6.3.
At the four critical locations identified in section 5.1 stability analyses based on limit load were also performed as described in section 7.2. The weld at these locations are either SMAW with GTAW root passes or SAW. Therefore, "Z" factor corrections for the SMAW welds were applied (reference 7-3) as follows: 2 = 1.15 (1.0 + 0.013 (00-4)) where 00 is the outer diameter of the pipe in inches. The Z-factor for SAW welds is as follows (reference 7-3): 2 = 1.30 (1 + 0.010 (00-4)) 7he Z-factors _were calculated for the four critical locations using the
- dimensions given in tables 3-1 and 3-2. These factors are given in table L 7-2. The applied loads were increased by the Z factor and a plot of limit load versus crack length was generated as shown in figures 7-2 through 7-5.
The critical flaw sizes at the four critical locations are given in table 7-2 along with the leakage flaw sizes. A margin well in excess of 2.is demonstrated at each location. The lower bound base metal properties established in section 4.0 were used for this purpose. 7.4 References 7-1. Kanninen, M. F. , et. al. . " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks," EPRI NP-192 September 1976.
. .on.i i. 7,4
-- - - .-,.._.- - ., - - . - . - . - . , - - . - . . . . - - = _ - , . - . - - . . - . . , . .,- , . _ . .-- .-..,
7-2. Kumar, V., German, M. D. and Shih, C. P., "An Engineering Approacn for Elastic-Plasti: Fracture Analysis," EPRI Report NP-1931 Project 1237-1, Electric Power Research Institute, July 1981. 7-3. ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640. l l
.ses.n nw o 75
l'1 TABLE 7 1
- STABILITY RESULTS FOR FARLEY UNITS 1 AND 2
, BASED ON ELASTIC-PLASTIC i J-INTEGRAL EVALUATIONS a Fracture Criteria Calculated Values i Flaw Size" J;g T mat U 2 max 2 app 2 "DD i Location (in) (in-lb/in ) (in-lb/in ) (in-lb/in ) i Farley Unit 1
- 6 c.e ~
1- 6.5
- 2 10.0 3 11.0
~ -
Farley Unit 2
~
L
~
a,c.e 1 6.2 4 . d Twice the 10 gpm leakage size flaw established in section 6.0 l N.A. - Not applicable; J,pp <Jyg m.m . 73
TABLE 7 2 - STABIL!'TY RESULTS BASED ON LIMIT LOAD ANALYSES Leakage 71aw Critical Flaw Location Weld Z Facter Si:e (in) ti:e (in) Margin l Farley Unit 1
~
gg, a.c.e 1 2 SAW l 3 SMAW Farley Unit 2
~
~
a.c.e e l ass.o useo ,o y7
i e
- 4,C,e
//////////// f, i
20 1, = Neutf 84 A set NN .)l
% i I
l ll 8 'C Figure 7-1. ( 1 Stress Distribution 44tw t 2200010 7,g
~
a,c.e Figure 7-2, Critical Flaw Si:e Prediction - Hot Leg at Location 1 for Farley Unit I c, ... . m s.o i s l 79 ;
.J
a c.e - t l Figure 7-3 Critical Flaw Size Prediction - Hot Leg at l Location 2 for Farley Unit 1
.....n n o io 7-10
I
'===
~~
a,c.e Figure 7-4 Critical Flaw Size Prediction - Hot leg at Location 3 for farley Unit 1 soes,?,aneo ,o
M, a ,c ,< Figure 7-5 Critical Flaw Size Prediction - Hot Leg at Location 1 for farley Unit 2 sees.nnw o y.yg
l l SECTION 8.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the presence of small cracks, a fatigue crack grcwth analysis was carried out for the ( Ja c.e region of a typical system (see Location ( Ja.c.e of figure 3-2). This regi0n was selected because Crack gr0wth calculated here will be typical of that in the entire primary loop. Crack growths calculated at other locatiens can be expectec to show less than IC% variation. A( Ja.c.e of a plant typical in geometry and operational characteris-tics te any Westinghouse PWR System. [ Ja c.e All normal, upset, and test conditions were considered. A summary.of the applied transients is provided in table B 1. Circumferentially oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in figure B-1. Scecifically, these were: Cross Section A: ( Ja.c.e Cross Section B: [ ]a,c.e Cross Section C: ( )8'C fatigue crack growth rate laws were used ( 8 'C 1 The law for stainless steel was derived from reference S-1, with a very conservative correction for the R ratic,, which is the ratio of minimum to maximum stress during a transient, ag66s !239010 g,{
For stainless steel, the fatigue crack growth formula is: ; i
'40 inches / cycle h*=(5.4x10-12) g aff where K,ff = K ,,, (1-R)0.5 j
R=Kmin/E max (
?
k e p ja.c.e .. a,c.e ( ) > where: ( Ja.c.e The calculated fatigue crack growth for semi-elliptic surface flaws of-circumferential orientation and various depths is summarized in table B-2, and shows that the_ crack growth is very small, (
.ja,c.e ,
l 8.1 References 8-1 Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Piping'in a Pressurized Water. Reactor Environment," Trans. ASME Journal of Pressure Vessel Technology, Vol. 101, Feb. 1979. Y l
.. ,in.o in 8-2
82(
)a.c.e s3(
)a.c.e D
4490s '!!i&90 to g,
TABLE 8-1
SUMMARY
OF REACTOR VESSEL TRANS!ENTS NUMBER TYPICAL TRANSIENT !DENTIFICATION NUMBER OF CYCLES Normal Conditions 1 Heatup and Cooldown at 100'F/hr 200 (pressurizar cooldown 200'F/hr) 2 Lead Follow Cycles 18300 (Unit leading and unloading at 5% of full power / min) 3 Steo lead increase and decrease 2000 4 Large step load decrease, with steem dump 200 5 Steady state fluctuations 10 6 Upset Conditions 6 Loss of lead, without immediate turbine 80 or reactor trip 7 Loss of power (blackout with natural circulation 40 in the Reactor Coolant System) l 8 Loss of Flow (partial loss of flow, one pump only) 80 9 Reactor trip from full power 400 Test Conditions 10 Turbine roll test 10 j 11 Hydrostatic test conditions Primary side 5 Primary side leak test 50 12 Cold Hydrostatic test 10 W
. . .. . . . o .. o .
8-4
TABLE 8-2 TYPICAL FATIGUE CRACK GROWTH AT [ Ja.c.e (40 YEARS) FINAL FLAW (in) [ INITIAL FLAW (IN) Ja,c.e [ 3a c.e [ 3a,c.e 0.292 0.31097 0.30107 0.30698 0.300 0.31949 0.30953 0.31626 0.375 0.39940 0.38948 0.40763 0.425 0.45271 0.4435 0.47421
....., m .. "
8-5
- e ,c.e
~ Dimensions in inches a Farley Unit 1 b Farley Unit 2 Figure 8-1. Typical Cross-Section of ( Ja c.e ases nuseo ,o g.g
F" - a.c. e w a O U E) w z O E 6 m M E E s 4 d r U z N O z o w U ( b Figure B 2. Reference Fatigue Crack Grewth Curves for (
)a,c,e
. . ... , , u..a , o g.,
e,c. .
. i.
y. C - l s c i .
., . i r
s
\
Figure 8-3, Reference Fatigue Crack Growth Law for [ Ja c,e in a Water Environment at 600*F 4496s/132640 10 g, L.
1 SECTION 9.0 ASSESSMENT OF MARGINi ; The results of the leak rates of section ',.4 and the corresponding fracture toughness evaluations of section 7,3 are used in performing the assessment of margins. Leakage flaws were established at the four critical locations band on a leak rate of 10 gpm. A margin of at least two en the leakage flew size w3J established using both elastic-plastic fracture mechanics and limit load analysis (with the 2-factor correction). The faulted iceds were determined using the absolute sum method thus a margin on lead of 1.0 for the leakage flaw is adequate per SRP 3.6.3.. It follows, of course, that the leakage flaw is stable since a flaw twice as large n6s shown to be stable. In summary, at all the critical locations relative to:
- 1. Flaw Size - Using faulted loads obtainec by the absolute sum method, a margin of at least 2 exists between the critical flaw and the flaw having a leak rate of 10 gpm (the leakage flaw).
L
- 2. Leak Rate - A margin of 10 exists between the calculated leak rate fecm the leakage flaw and the leak detectier capability of 1 gpm.
- 3. Loads,- At the critical locations the leakage flaw was shown to be stable using the faulted Icads obtained by the absolute sum method.
u ss,n n sso ,o 93
SECT (ON
10.0 CONCLUSION
S This report justifies the elimination cf RiS : S imary loop eipe breaks fer the Farley Units 1 and 2 nuclear plants as follows:
- a. Stress corrosion cracking is precluded by use of fracture r asir: ant materialt in the piping system and controls on reacter coolant chemistry, temperature, pressure, and flow curing normal cperatien.
- b. Water hammer saould not occur in the RCS piping because of systam design, testing, and operational considerations,
- c. The effects of low and high cycle fatigue en the integrity of the primary piping are negligible.
- d. Acequate margin exists between the leak rate of small stable flaws and the capability of the Farley Units 1 and 2 reactor coolant system pressure boundary Leakage Detection System.
- e. Ample margin exists between the small stable flaw si:es of item d and larger stable flaws.
- f. Ample margin exists in the material properties used to demonstrate end-of service life (relative to aging) stability of the critical flaws.
For the critical locations flaws are identified that will be stable because of the ample margins in d, e, and f above. Based on the above, it is concluded that dynamic effects of RCS primary icop pipe breaks need not be considered in the structural design basis of the Farley Units 1 and 2 plants.
.ess.n nsso ,o 39.z
I APPENDIX A LIM!i MOMENT [ 8,C,0 3 m s,n nsso ,o g3
l ta c.e l. Figure A-1 Pipe with a Through-Wall Crack in Bending
. "uno in A-2
i 1 APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE FARLEY UNITS 1 AND 2 CAST PRIMARY LOOP 40MPONENTS B.1 Introduction " Not all of the individual cast piping components of the Farley primary loop piping satisfy the original { ]a,c.e criteria (reference 4-3). In this appendix, the alternate toughness criteria for thermally aged cast stainless steel developed in reference 4-6 will be used to categorize the various
'ndividual cast piping components thus establishing criteria based upon which the leak-before-break evaluations may be performed. Reference 4-6 has been reviewed by the NRC wherein the NRC concluded that reference 4-6 may be utilized for establishing the fracture criteria for thermally aged cast stainless piping applicable for the leak-before-break analyses (reference B-1).
B.2 Chemistry and KCU Toughness Per the procedure of reference 4-6 the correlat'.ons of reference 4-4 which are based on the chemistry of the cast stainless steel piping was used to calculate the associated KCU values. The chemistry and end of-service life KCU toughness values are given in table B-1 for Unit 1 and in table B-2 for Unit 2. B.3 Alternate Toughness Criteria for the Farley Primary Loop Material en a Comoonent by Comoonent Basis The alternate toughness criteria for the Farley Unit 1 and 2 cast primary loop material may be obtained by applying the methodology of reference 4-6 to the KCU values of tables B-1 and B-2. First, it is observed that 44 of the 50 heats fall into category 1, i.e., they are at least as tough as [
]a c,e The remaining heats fall into category 2.
Typical toughness calculations using the methodology of reference 4-6 are given below for a category 2 heat. sm.nnuo n 93
t h b For example c( )"'C (the reducing elbow at the steam generator inlet nozzle of Farley Unit 1) has the calculated end-of service ' l.ife'KCU at room temperature of (- )"'C'"daJ/cm2 which falls below that of ( Ja,c.e. The 6-ferrite content is ( )"'C. By reference 4-6, the (
).a.c.e Since the end-of service life KCU exceeds the fully aged KCU, t% heat falls into category 2. Thus:
Jte = [ j a,c.e Tmat
- I- 3
and max ' I
],,c,,.
The fracture toughness. values for each heat of material was calculated as 4 formulated in. references 4-6. These values are also given in tables B-1 and
- B-2.
B.4 References B-1 Letter: Dominic C., Dilanni, NRC to D. H.- Husolf, Northern States Power Company, dated December 22, 1986, Docket Nos 50-282 and 50-306. i l-l-
.inua io B-2 t- pyyg+y 7- ~ , , - - --rg--r-3-+ q ypy y - _a -
,-g--cm-+,- ~e--p-+---m eem-- m1vp + n-r- a m'- -------
TABLE B-1 CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEAIS OF FARLEY UNIT 1 1,c,e
.u.., m m "
B-3 l
TABLE B-1 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 1 8.0.0 i 1 I
=
f 4494vI2299010 i B-4
l TABLE B-1 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 1
~
a,c,e
..... . i :2..a i o B-5
. . - ~ - - _ . . - . . .. . . _ - .. . - . . - .
1 1 TABLE ~B-1-(Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF'THE MATERIAL HEATS OF FARLEY UNIT 1
~
8 ,C,0-
- J D
?
4 4 a 6 1
-o m. i .
B-6 s L _ _ . . , . _ _ . _ . - _ - . . - .,
TABLE B-1 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 1 a,c,e
.....o nuo in B-7
TABLE B-1 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 1 a,c.e I l 1
) .m.a n a in B-8
TABLE B-1 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 1 a,c.e 4496s/132900 10 g,g l
- ' - " '^
=_apa._ .e m -a.4r,A &a-..4--- + --'hM' E
- 4 J E M e+ 4- ri lW MM---S6+44d0.=-
-M-Ne-*d'N R J.bameMe n psM M d A. a Zeha 4A -se.4pJ.4'.2-' & Md.M .AAK 4 A
-TABLE B-1 (Co'nt'd) '
CHEMISTRY AND FRACTURE = TOUGHNESS PROPERTIES
- 0F THE MATERIAL HEATS OF FARLEY-UNIT-1
- ~
e 3,0,e-
'I fi l
m l 1 ; I' t'
-L I
l. l
= 'M
' 44Most 2200010 B-10
====
O iCte 4096s r 11250010 B-11
TABLE B CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY-' UNIT 2: . S,C,e i
\
l 1 f
- i. ;
m m 4444e113200010 B-12
TABLE B-2 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 2 a ,c.e u...,im o in B-13
1 1 TABLE B-2 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE NATERIAL HEATS OF FARLEY UNIT 2 , l ~
~~a,c,e i
J
.....n ueu in B-14
' TABLE B-2 (Cont'd)
CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 2 s,:,e
.....,i :2 =o i .
B-15
_ _ TABLE B-2 (Cont'd). CHEMISTRY AND-FRACTURE TOUGHNESS PROPERTIES OF-THE MATERIAL HEATS OF FARLEY UNIT 2 4 ,C,e. l-1. e
'- . g, B-16
I TABLE B-2 (Cont'd) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 2 a,c,e
... n um io B-17
~ . .. _, . . _ _ . ~ . _-
- . TABLE B-2-(Cont'd)
~ CHEMISTRY'AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 2 N
8,C,e: i i I 4 i j i n e MD
'I 44Ms/132eso 10 -
- B-18
TABLE B-2 (Cont'c) CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF FARLEY UNIT 2
~~~
'a,c,e 4 8 l
l
- c. ... , i r * '
B-19
8,C,9
, essa e
9
.. .n 2=o io B 20
- _ _ _ _ _ _ _ _ _ _ _}}