ML20108D246

Nonproprietary Technical Justificaton for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for Byron & Braidwood Units 1 & 2 Nuclear Power Plants

ML20108D246
Person / Time
Site:	Byron, Braidwood
Issue date:	04/30/1996
From:	Bhowmick D, Kaihwa Hsu, Vora V WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19355D096	List: ML20108D229 ML20108D236 ML20108D246
References
WCAP-14560, WCAP-14560-R01, WCAP-14560-R1, NUDOCS 9605070354
Download: ML20108D246 (89)

Text

r WESTINGHOUSE NON-PROPRIETARY CLASS 3 l

i WCAP-14560 Revision 1 l

i TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE BYRON AND BRAIDWOOD UNITS 1 AND 2 NUCLEAR POWER PLANTS I APRIL 1996 l

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D. C. Bhowmick V. V. Vora K. R. Hsu VERIFIED. W '

D. E. Prager APPROVED: . - -

/4.< r"

5. 'A' Swamy, Manhgfr Structural Mechanics Technology Work Performed Under Shop Order BGCP-950 l

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! WESTINGHOUSE ELECTRIC ~ CORPORATION Systems and Major Projects Division P. O. Box 355

, Pittsburgh, Pennsylvania 15230-355 l

C 1996 Westinghouse Electric Corporation All Rights Reserved m:\2172w.wpf:lb/0422%

9605070354 960430 PDR ADOCK 05000454 p PDR

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l TABLE OF CONTENTS Section Title g l EXECUTIVE

SUMMARY

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1.0 INTRODUCTION

1-1 1.1 Purpose 1-1 j 1.2 Background Information 1-1 1.3 Scope and Objectives 1-2 1.4 References 1-3 2.0 OPERATION AND STABILITY OF THE REACTOR 2-1 COOLANT SYSTEM 2.1 Stress Corrosion Cracking 2-1 2.2 Water Hammer 2-2 2.3 Low Cycle and High Cycle Fatigue 2-3 2.4 References 2-3 3.0 PIPE GEOMETRY, LOADS AND STRESSES 3-1 3.1 Introduction 3-1 3.2 Calculation of Loads and Stresses 3-2 3.3 Loads for Leak Rate Evaluation 3-3 3.4 Load Combination for Crack Stability 3-3 Analyses 3.5 References 3-4 4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Primary Loop Pipe, Fittings and Weld Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-2 4.4 References 4-3 l

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TABLE OF CONTENTS Section Title .P_ age 5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA 5-1 5.1 Critical Locations 5-1 5.2 Fracture Criteria 5-1 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 References 8-2 9.0 ASSESSMENT OF MARGINS 9-1

10.0 CONCLUSION

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TABLE OF CONTENTS Section Title Page APPENDIX A - Limit Moment A-i APPENDIX B - Toughness Criteria for Byron and Braidwood Units B-1 1 and 2 Cast Primary Loop Components i

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l LIST OF TABLES l

Table Title Page 3-1 Normal Loads and Normal Stresses for Byron and Braidwood 3-5 Units 1 and 2 1

3-2 Faulted Loads and Stresses for Byron and Braidwood Units 1 and 2 3-6 4-1 Measured Tensile Properties for Byron Unit 1 Primary Loop Piping 4-5 Material SA376 Gr. 304N 4-2 Measured Tensile Properties for Byron Unit 2 Primary Loop Piping 4-7 Material SA376 Gr. 304N l

4-3 Measured Teesile Properties for Braidwood Unit 1 Primary Loop 4-9 Piping Material SA376 Gr. 304N l 4-4 Measured Tensile Properties for Braidwood Unit 1 Primary Loop 4-10 Piping Material SA376 Gr. 304N 4-5 Measured Room Temperature Tensile Properties for 4-11 Byron Unit 1 Primary Loop Elbow Fittings 4-6 Measured Room Temperature Tensile Properties for 4-12 l Byron Unit 2 Primary Loop Elbow Fittings 4-7 Measured Room Temperature Tensile Properties for 4-13 t

l Braidwood Unit 1 Primary Loop Elbow Fittings 4-8 Measured Room Temperature Tensile Properties for 4-14 Braidwood Unit 2 Primary Loop Elbow Fittings 4-9 Mechanical Properties for Byron and Braidwood Units 1 and 2 4-15

Materials at Operating Temperatures i

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LIST OF TABLES (cont)

Table Title P_ ate, 4-10 Fracture Toughness Properties for Byron and Braidwood Units 4-16 1 and 2 Primary Loops for Leak-Before-Break Evaluation at Critical Location 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the 6-4 Governing Locations 7-1 Stability Results for Byron and Braidwood Units 1 and 2 Based on 7-5 Elastic-Plastic J-Integral Evaluations 7-2 Stability Results for Byron and Braidwood Units 1 and 2 Based on 7-6 Limit Load i

8-1 Summary of Reactor Vessel Transients 8-3 i

8-2 Typical Fatigue Crack Growth at [ 8-4

]"# (40 Years) 9-1 Leakage Flaw Sizes, Critical Flaw Sizes and Margins 9-2 for Byron and Braidwood Units I and 2 B-1 Chemistry and Fracture Toughness Properties of the B-2 Material Heats of Byron Unit 1 B-2 Chemistry and Fracture Toughness Properties of the B-4 Material Heats of Byron Unit 2 B-3 Chemistry and Fracture Toughness Propenies of the B-6 Material Heats of Braidwood Unit 1 B-4 Chemistry and Fracture Toughness Properties of the B-8 Material Heats of Braidwood Unit 2 mA2172w.wpf:1b/022096 viii

LIST OF FIGURES Figure Title P_ age 3-1 Cold Leg Coolant Pipe 3-7 l

3-2 Schematic Diagram of Byron and Braidwood Units 1 and 2 Primary 3-8 Loop Showing Weld Locations I

4-1 Representative Lower Bound True Stress - True Strain 4-17 Curve for SA351 CF8A at 617 F l I

4-2 Pre-Service J vs. Aa for Cast Stainless Steel 4-18 at 600 F 4-3 J vs. An at Different Temperatures for Aged Material 4-19

[ ]"# (7500 Hours at 400 C) 6-1 Analytical Predictions of Critical Flow Rates of 6-5 St am Water Mixtures 6-2 [ ]"# Pressure Ratio as a Function 6-6 of IJD 6-3 Idealized Pressure Drop Profile Through a Postulated 6-7 Crack m:\2172w.wpf:Ib/021996 ix

d LIST OF FIGURES (cont)

Figure Title Page a

7-1 [ ]"" Stress Distribution 7-7 7-2 Critical Flaw Size Prediction - Hot Leg at Location 3 7-8 7-3 Critical Flaw Size Prediction - Cold Leg at Location 11 7-9 8-1 Typical Cross-Section of [ ]"# 8-5 8-2 Reference Fatigue Crack Growth Curves for [ 8-6 3us 8-3 Reference Fatigue Crack Growth Law for [ ]"# .

8-7 in a Water Environment at 600 F 1 A-1 Pipe with a Through-Wall Crack in Bending A-2 .

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b EXECUTIVE

SUMMARY

The original structural design basis of the reactor coolant system for the Commonwealth Edison Company Byron and Braidwood Units 1 and 2 Nuclear Power Plant required consideration of dynamic effects resulting from pipe break and that protective measures for l

such breaks be incorporated into the design. Subsequent to the original Byron and Braidwood I design, additional concem of asymmetric blowdown loads was raised as described in Unresolved Safety Issue A-2 (Asymmetric Blowdown Loads on the Reactor Coolant System) and Generic Letter 84-04 (Reference 1-1). However, research by the NRC and industry l coupled with operating experience determined that safety could be negatively impacted by placement of pipe whip restraints on certain systems. As a result, NRC and industry I initiatives resulted in demonstrating that Leak-Before-Break (LBB) criteria can be applied to reactor coolant system piping based on fracture mechanics technology and material toughness.

The Byron and Braidwood Units 1 and 2 primary loop piping analyses by Westinghouse for the application of LBB was documented in WCAP-10553 (Reference 1-2) and approved by the NRC letter dated October 28,1985 (Reference 1-3). By letter dated October 28,1985, l

the NRC stated that:

"In a letter dated September 28,1984, Commonwealth Edition Company (CECO) requested an exemption from a portion of the requirements of General Design Criterion (GDC) 4 of Appendix A to l 10 CFR Part 50. You provided the Westinghouse report " Technical Bases for Eliminating Large Primary Loop Pipe Rupture as a l Structural Design Basis for Byron Units 1 and 2 and Braidwood Units i

1 and 2, "WCAP-10554 (Westinghouse Non-Proprietary) and WCAP-10553 (Westinghouse Proprietary) as an enclosure to this letter which l serves as the technical basis in support of the request. The Westinghouse report addresses the " leak-before-break" concept as an I

alternative to providing protective devices against the dynamic effect of postulated rupture in the primary coolant loops. Your submittal dated June 28,1985 provided a value-impact analysis associated with l the exemption request and requested that a partial exemption to GDC-4 be granted for the first two cycles of operation for Byron Station, Units 1 and 2, and Braidwood Station, Units 1 and 2 (the facilities).

Your letter dated August 14,1985 withdrew, without prejudice, the request for the exemption for Byron Station, Unit 1.

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On the basis of the Staff's evaluation of these submittals the Commission has granted your exemption request for periods ending at the completion of the second refueling outage of each of the facilities, pending the outcome of the Commission's ongoing rulemaking on this subject. The staff has also received your September 25,1985 letter requesting amendments to the construction permits for Byron Station, Unit 2, and Braidwood Station, Units 1 and 2. The exemption granted will become effective upon the date ofissuance. The enclosed exemption is being forwarded to the office of the Federal Register for publication, accordingly."

This report demonstrates compliance with LBB technology for the Byron and Braidwood reactor coolant system piping based on the latest criteria. The report documents the plant specific geometry, loading, and material properties used in the fracture mechanics evaluation.

Mechanical properties were determined at operating temperatures. Since the piping systems include cast stainless steel fittings, fracture toughnesses considering thermal aging were determined for each heat of material, i

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 cause leak at a rate of ten times the leakage j detection system capability of the plant. Large margins for such flaw sizes were demonstrated against flaw instability. Finally, fatigue crack growth was shown not to be an j

issue for the primary loops.

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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 Byron and Braidwood Units 1 and 2 Nuclear Power Plants.

Revision 1 is to modify the second paragraph of section 2.3 and to add reference 2-3.

The revisions are identified by vertical lines in the column. '

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L i o SECTION

1.0 INTRODUCTION

- 1.1 Purpose  !

This report applies to the Byron and Braidwood Units 1 and 2 Reactor Coolant System (RCS) primary loop piping. It is intended to demonstrate that for the specific parameters of the l Byron and Braidwood Units 1 and 2 Nuclear Power Plants, RCS primary loop pipe breaks need not be considered in the structural design basis. The approach taken has been accepted l by the Nuclear Regulatory Commission (NRC) (Reference 1-1). l l

1.2 Background Information l

Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated from the stmetural design basis of all l Westinghouse plants. The concept of eliminating pipe breaks in the RCS primarf loop was

- first presented to the NRC in 1978 in WCAP-9283 (Reference 1-4). That topical report employed a deterministic fracture 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. l 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 1-5).

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 1-6 and 1-7). 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 per reactor year and the mean probability of an indirect LOCA to be 10-7 per reactor year. Thus, the results previously obtained by Westinghouse (Reference 1-4) were confirmed by an independent NRC research study.

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4 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 leak-Before-Break (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 Rupture" (Reference 1-8).

1.3 Scope and Objective The general purpose of this investigation is to demonstrate leak-before-break for the primary loops in Byron and Braidwood Units 1 and 2. The recommendations and criteria proposed in Reference 1-9 are used in this evaluation. These criteria and 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.

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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 size of the flaw should

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be large enough so that the leakage is assured of detection with margin using the installed leak detection equipment when the pipe is subjected to normal operating loads. A margin of 10 is demonstrated between the calculated leak rate and the leak l

detection capability.

5) Using faulted loads, demonstrate that there is a margin of 2 between the leakage size 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 I and high cycle fatigue. j m:\2172w.wpf:1b/021996 1-2

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7) For the materials actually used in the plant provide the properties including toughness '

! and tensile test data. Evaluate long term effects such as thermal aging where applicable.

8) Demonstrate margin on applied load.

l This report provides a fracture mechanics demonstration of primary loop integrity for the j Byron and Braidwood Units 1 and 2 Plants consistent with the NRC position for exemption i 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 NRC's Standard Review Plan 3.9.1 (Reference 1-10). The fracture mechanics calculations are independently verified (benchmarked).

1.4 References 1-1 USNRC Generic letter 84-04,

Subject:

" Safety Evaluation of Westinghouse Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary Main Loops," February 1,1984.

I l-2 WCAP-10553, " Technical Justification for Eliminating Large Primary Loop Pipe Rupture as the Stmetural Design basis for the Byron Units 1 and 2 and Braidwood Units 1 and 2," May 1984.

1-3 Nuclear Regulatory Commission Docket Nos. STN 50-455, STN 50-456, and STN 50-457, dated October 28,1985, Letter from B. J. Youngblood,' Chief Licensing Branch No.1 Division of Licensing, NRC, to Dennis L. Farrar, Director of Nuclear Licensing, Commonwealth Edison Company.

1-4 WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events," March,1978.

1-5 Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G. Eisenhut),

Westinghouse Proprietary Class 2, November 10,1981.

1-6 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated i April 25,1983.

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1.4 References (cont) 1-7 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25,1983.

I l-8 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 RegisterNol. 52, No. 207/ruesday, October 27,1987/ Rules and Regulations, pp. 41288-41295.

1-9 Standard Review Plan: Public Comments Solicited; 3.6.3 leak-Before-Break Evaluation Procedures; Federal RegisterNol. 52, No.167/ Friday August 28, 1987/ Notices, pp. 32626-32633.

1-10 Nuclear Regulatory Commission, Standard Review Plan Section 3.9.1, "Special Topics for Mechanical Component," NUREG-0800, Revision 2, July 1981.

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l-11 WCAP-7211, Revision 3, " Energy Systems Business Unit Policy and Procedures for Management, Classification, and Release of Information," March,1994.

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s SECTION 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Cracking The Westinghouse reactor coolant system primary loops 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 corrosion cracking (IGSCC)). This operating history totals over 800 reactor-years, including five plants each having over 18 years of operation and 15 other plants each with over 13 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 (PCSG) established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second 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 Corrosion Cracking in Piping of Light Water Reactor Plants." In that report the PCSG stated:

l "The PCSG has determined that the potential for stress-conosion cracking in PWR primary system piping is extremely low because the ingredients that produce IGSCC are not all present. The use of hydrazine 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 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 reponed in NUREG-0691 (Reference 2-2) further confirmed that no occurrences of IGSCC have been reponed 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 qualifies the PCSG's fm' dings.

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For stress corrosion cracking (SCC) to occur in piping, the following three conditions must l exist simultaneously: high tensile stresses, susceptible material, and a corrosive environment.

Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized 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 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 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.

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 normal power operation, oxygen concentration in the RCS is expected to be in the ppb range by controHing charging flow chemistry and maintaining hydrogen in the reactor coolant at specified concentrations. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and fluorides within the specified limits.

Thus during plant operation, the likelihood of stress corrosion cracking is minimized.

2.2 Water Hammer 1

( l Overall, there is a low potential for water hammer in the RCS since it is designed and I operated to preclude the voiding condition in normally filled lines. The reactor coolant system, including piping and primary components, is 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 vahe opening are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system mA2172w.wpf:lb/021996 2-2 l

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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 rod 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 j 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.

2.3 Low Cycle and High Cycle Fatigue Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the rules of Section III of the l

ASME Code. A further evaluation of the low 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.

High 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 through periodic monitoring by plant personnel through instrumentation located in the Byron and Braidwood Auxiliary Electrical Equipment Room. Field measurements have been made on a number of plants during hot functional testing, including plants similar to Byron and Braidwood 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.

l These stresses are well below the fatigue endurance limit for the material and would also l result in an applied stress intensity factor below the threshold for fatigue crack growth.

l During preoperational testing at Byron and Braidwood, vibration 'evels of the Reactor Coolant l System were measured and determined by Commonwealth Edison to be significantly below l the Westinghouse acceptance criteria (Reference 2-3).

2.4 References l

l l 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.

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2.4 References (cont) 2-3 Letter # CAW-5738; " Pre-Operational loop Testing," from Westinghouse (W. E.

Kortier) to J. D. Deress of Commonwealth Edison Company dated May 19,1983.

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SECTION 3.0 PIPE GEOMETRY, LOADS AND STRESSES 3.1 Introduction The general approach is discussed first. As an example a segment of the primary coolant cold leg pipe is shown in Figure 3-1. The as-built outside diameter and minimum wall thickness of the pipe are 32.14 in, and 2.215 in., respectively, as shown in the figure. The normal stresses at the weld locations are from the load combination procedure discussed in Section 3.3 whereas the faulted loads are as described in Section 3.4. The components for normal loads are pressure, dead weight and thermal expansion. An additional component, Safe Shutdown Earthquake (SSE), is considered for faulted loads. As seen from Table 3-2, the highest stressed location in the entire loop is at Location 11 at the reactor coolant pump t outlet nozzle to pipe weld. This highest stressed location is a load critical location and is one of the locations at which, as an enveloping location, leak-before-break is to be established.

Essentially a circumferential flaw is postulated to exist at this location which is subjected to both the normal loads and faulted loads to assess leakage and stability, respectively. The loads (developed below) at this location are also given in Figure 3-1.

Since the elbows are made of cast stainless steel, thermal aging must be considered (Section 4.0). Thermal aging causes in lower fracture toughness of the cast materials; thus, location other than the load critical location must be examined taking into consideration both fracture toughness and stress. The enveloping location so determined is called the toughness critical location. One most critical location is identified after the full analysis is completed.

Once loads (this section) and fracture toughnesses (Section 4.0) are obtained, the load critical and toughness critical locations are determined (Section 5.0). At these locations, leak rate evaluations (Section 6.0) and fracture mechanics evaluations (Section 7.0) are performed per the guidance of Reference 3-1. Fatigue crack growth (Section 8.0) and stability margins are also evaluated (Section 9.0).

All the weld locations including load critical location and toughness critical location for evaluation are shown in Figure 3-2.

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3.2 Calculation of Loads and Stresses The stresses due to axial loads and bending moments are calculated by the following equation:

M o = _F +_ (3-1)

A z

where, o = stress F = axial load M = bending moment A = pipe cross-sectional area Z = section modulus The bending moments for the desired loading combinations are calculated by the following equation:

(3*2)

M = }Mr* + Mi *

where, M = bending moment for required loading My = Y component of bending moment M = Z component of bending moment 7

The axial load and bending moments for leak rate predictions and crack stability analyses are computed by the methods to be explained in Sections 3.3 and 3.4.

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1 3.3 Loads for Leak Rate Evaluation The normal operating loads for leak rate predictions are calculated by the following equations:

F =

FDW + FTH + Fp (3-3)

M y= (My)DW + (My)TH + (My)p (3-4)

M =

(M7)DW + (MZ)TH + (Mg)p (3-5)

Z The subscripts of the above equations represent the following loading cases:

DW = deadweight TH = normal thermal expansion P = load due to intemal pressure l

This method of combining loads is often referred as the algebraic sum method (Reference 3-1).

The loads based on this method of combination are provided in Table 3-1 at all the locations identified in Figure 3-2.

3.4 Load Combination for Crack Stability Analyses l

In accordance with Standard Review Plan 3.6.3 (Reference 3-1) the absolute sum of loading components can be applied which results in higher magnitude of combined loads. If crack stability is demonstrated using these loads, the LBB margin on loads can be reduced from 42 to 1.0. The absolute summation of loads are shown in the following equations:

F = l Fow l + l Fm l + l F, l + l PsstrusaTtA l + l FSSEAM l (3-6) l My = l (My)ow l + l (My)m l + l (My)p l + l (My)sseinsartA l + l (My)3:siu l (3-7)

Mz = l (Mz)ow l + l (Mz)m ! + l (Mz)p l + l (Mz)SSEINER11A + (Mz)33exu l Q-8) where subscripts SSE, INERTIA and AM mean safe shutdown earthquake, inertia and anchor motion, respectively.

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. . -. _ _ _. . _ _ _ _ _ _ _ _ _ . _ . _ . _ . _ - _ . _ _ _ . . . ~ _ _ _ _.

The loads so determined are used in the fracture mechanics evaluations (Section 7.0) to demonstrate the LBB margins at the locations established to be the governing locations.

These loads at all the locations of interest (see Figure 3-2) are given in Table 3-2.

3.5 References 3-1 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.

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Table 3-1 i

Normal Loads and Normal Stresses for Byron and Braidwood Units 1 and 2 i

Outside Minimum Axial Bending Total Diameter Thickness Load" Moment Stress 1

Location * (in) (in) (kips) l (in kips) (ksi)  !

1 33.90 2.345 1470 17909 16.76 1 2 33.90 2.345 1

1472 917 6.87 3 34.21 2.500 1472 8635 10.60 4 37.56 3.175 1616 14023 9.86 I 5 37.57 3.180 1638 5161 6.66 6 36.20 2.495 1737 4014 8.50 7 36.20 2.495 1624 3846 7.99 8 36.20 2.495 1673 663 6.64 9 36.20 2.495 1673 2898 7.72 10 37.57 3.180 1770 7052 7.74 11 32.14 2.215 1404 5481 10.50 1 12 32.14 2.215 l 1404 5191 10.30 l

13 32.14 2.215 1404 3"' 9.10 14 32.14 2.215 1389 4770 9.94 15 32.43 2.360 1387 4937 9.38

$ See Figure 3-2 Includes pressure m:\2172w.wpf:1b/0219% 3-5 l

i l

Table 3 2 Faulted Loads and Stresses for Byron and Braidwood Units 1 and 2 Axial Load 6 Bending Moment Total Stress Location' (kips) (in kips) (ksi) 1 2917 30318 30.21 2 2892 9965 18.24 3 2573 22713 22.67 4 2686 31649 19.46 5- 2095 30190 17.17 6 1977 22177 18.12 7 1984 12980 13.74 8 1988 10434 12.53 9 1953 13057 13.66 10 2017 39666 20.42 1 11 2275 31856 32.77 12 2275 28561 30.51 13 2306 19245 24.27 l 14 2293 13283 20.12 15 2226 15558 19.93  ;

• See Figure 3-2 b Includes pressure l

l m:\2172w.wpf:1b/021996 3-6

l l

1 l

1 d

+ Crack

] Thickness = t 4

. ^

t ......

2 O a 4, F 4 s-w-- -*- GP r e w

'r ,0 M

. = =

oD 4

OD* = 32.14 in t' = 2.215 in 2

1 Normal Loads' Faulted Loads b force': 1404 kips force': 2275 kips bending moment: 5481 in kips bending moment: 31856 in-kips 4

• See Table 3-1

• See Table 3 2

' Includes the force due to a pressure of 2305 psia Figure 31 Cold Leg Coolant Pipe m:\2 I72w.wpf:Ib/II2295 3-7

Reactor Pressure vessel 15 9

O HOT LEG -

COLD LEG 3

g

^

@ - 1 O S l

8.h it REACTOR COOLANT PUMP STEAM GENERATOR l

CROSSOVER LEG HOT LEG 4 x 94 Temperature 617'F, Pressure: 2250 psia CROSSOVER LEG n a Temperature 557'F, Pressure: 2215 psia COLD LEG Temperature 557'F, Pressure: 2305 psia Figure 3-2 Schematic Diagram of Byron and Braidwood Units 1 and 2 Primary Loop Showing Weld Locations m:\2172w.wpf:I b/I l 2295 3-8

1 SECTION 4.0 MATERIAL CHARACTERIZAT.f0N I

4.1 Primary Imop Pipe, Fittings, and Weld Materials The primary loop piping material for Byron and Braidwood Units 1 and 2 is SA376 Gr. i l

304N. The elbows are made of SA351 CF8A material. The weld processes used were TIG and SMAW combination.

l 4.2 Tensile Properties The Pipe Certified Materials Test Repons (CMTRs) for Byron and Braidwood Units 1 and 2 were used to establish the tensile properties for the leak-before-break analyses. The CMTRs l include tensile properties at room temperature for each of the heats of material. These properties are given in Table 4-1 through Table 4-4 for piping and in Table 4-5 through Table 4-8 for elbow fittings.

For the SA376 Gr. 304N material, the representative properties a: 557 F were established from the tensile propenies at room temperature given in Table 4-1 through 4-4 by utilizing Section III of the 1989 ASME Boiler and Pressure Vessel Code (Reference 4-6). Code tensile properties at 557 F were obtained by interpolating between the 500 F and 600 F tensile properties. Ratios of the code tensile properties at 557 F to the corresponding tensile propenies at room temperature were then applied to the room temperature tensile properties given in Table 4-1 through 4-4 to obtain the plant specific properties for SA376 Gr. 304N at 557*F.

The Elbow Fittings Certified Materials Test Reports (CMTRs) for Byron and Braidwood Units 1 and 2 were used to establish the tensile properties for the leak-before-brruk analyses.

The CMTRs for elbow fittings include tensile properties at room temperature foz each of the heats of material. These properties are given for Byron and Braidwood Units 1 and 2 in Table 4-5 through Table 4-8.

For the SA351 CF8A material, the representative propenies at 617*F were established from the tensile properties at room temperature given in Table 4-5 through 4-8 by utilizing Section III of the 1989 ASME Boiler and Pressure Vessel Code. Code tensile properties at 617 F were established by interpolating between the 600 F and the 650 F tensile propenies. Ratios of the code tensile properties at 617 F to the corresponding properties at room temperature m:\2172w.wpf:lb/021996 4-1

4

, \

! were then applied to the room temperature properties given in Table 4-5 through 4-8 to obtain the plant specific representative properties for SA351 CF8A at 617*F.

The average and lower bound yield strengths and ultimate strengths are given in Table 4-9.

The ASME Code modulus of elasticity are also given, and Poisson's ratio was taken as 0.3.

For leak-before-break fracture evaluations of the toughness critical location the true stress-tme strain curve for SA351 CF8A at 617'F must be available. This curve was obtained using the i Nuclear Systems Materials Handbook (Reference 4-1). The lower bound true stress-true

, strain curve is given in Figure 4-1.

1 4.3 Fracture Toughness Properties The pre-servne fracture toughnesses of both forged and cast stainless steels ofinterest here f have in terms of Ju been four.d to be very high at 600 F. Typical results for a cast material are given in Figure 4-2. Ju is observed to be over 2500 in-lbs/in 2. Forged materials are even higher. However, cast stainless steels are subject to thermal aging during service. This l thermal aging causes an elevation in the yield strength of the material and a degradation of the fracture toughness, the degree of degradation being somewhat proportional to the level of 4

ferrite in the material. >

To determine the effects of thermal aging on piping integrity, a detailed study was carried out t

in Reference 4-2. In that report, fracture toughness results were presented for a material

[

t

] The effects of the aging process on the end-of-service life fracture toughness are further discussed in Appendix B.

End-of-service life toughnesses for the heats are established using the alternate toughness criteria methodology of Reference 4-5 (Appendix B). By that methodology a heat of material is said to be as good as [ ] if it can be demonstrated that its end-of-service fracture 3

m:\2172w.wpf:1b/021996 4-2

_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ~

toughnesses equal or exceed those of [

]"#, The fracture toughness value for Byron and Braidwood Units 1 and 2 loops at the toughness critical location, as established in Appendix B, is given in Table 4-10.

Available data on aged stainless steel welds (References 4-2 and 4-3) indicate that Ju values for the worst case welds are of the same order as the aged material. However, the slope of the J-R curve is steeper, and higher J-values have been obtained from fracture tests (in 2

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 the temperature". Therefore, weld regions are less limiting than the cast material.

It is thus conservative to choose the end-of-service life toughness properties of [ ]"# as representative of those of the welds. Also, such pipes and fittings have an end-of-service life calculated room temperature Charpy U-notch energy, (KCU), greater than that of [ ]"#

are also conservatively assumed to have the properties of [ ]"#.

l In the fracture mechanics analyses that follow, the fracture toughness properties given in Table 4-10 will be used as the criteria against which the applied fracture toughness values will be compared.

Forged stainless steel piping such as SA376 Gr. 304N does not degrade due to thermal aging.

Thus fracture toughness values well in excess of that established for the cast material and welds exist for this material throughout service life and are not limiting.

4.4 References 4-1 Nuclear Systems Materials Handbook, Part I - Structural Materials, Group 1 - High Alloy Steels, Section 2, ERDA Report TID 26666, November,1975.

4-2 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.

In the report all the applied J value, were conservatively determined by using base metal strength properties.

m:\2172w.wpf:lb/021996 4-3

.. . .-. ~ . _ - . . - - . - - _ -. - -

I O

4.4 References (cont) 4-3 Slama, G., Petrequin, 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, Monterey, CA.

4-4 Appendix II of letter from Dominic C. Dilanni, NRC to D. M. Musolf, Northern States Power Company, Docket Nos. 50-282 and 50-306, December 22,1986.

4-5 Witt, F.J., Kim, C.C., " Toughness Criteria for Thermally Aged Cast Stainless Steel,"

WCAP-10931, Revision 1, Westinghouse Electric Corporation, July 1986, (Westinghouse Proprietary Class 2).

4-6 ASME Boiler and Pressure Vessel Code Section III, " Rules for Construction of Nuclear Power Plant Components; Division 1 - Appendices." 1989 Edidon, July 1, 1989.

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r k

s Table 4-1 l Measured Tensile Properties for Byron Unit 1 Primary Loop Piping i Material SA376 Gr. 304N I Component Heat Number Yield Room Temp. Ultimate (psi) Room Temp I (psi)

Hot leg L1360 Ser.14480 45900 89900 flot leg L1360 Ser.14480 48500 91400 Hot leg L1360 Ser.14482 43500 85700 Hot leg L1360 Ser.14482 46200 89200 Hot Leg K2980 Ser.12585Y 42200 84900 Hot leg K2980 Ser.12585Y 43000 87200 Hot Leg L1360 Ser.14481 44700 87900 Hot leg L1360 Ser.14481 48200 90700 Cold leg K2980 Ser.11044Y 42200 84900 Cold leg K2980 Ser.11044Y 43700 88100 l

Cold leg ,

L1336 Ser.14459 43400 86100 Cold leg L1336 Ser.14459 41700 81600 Cold Leg K3660 Ser.14458 45400 89900 l

j Cold leg K3660 Ser.14458 47200 91700 Cold leg K3723 Ser.14477X 45000 87400 Cold leg K3722 Ser.16100 41200 84900 Cold leg K3722 Ser.16100 41400 86600 Cold Leg K3723 Ser.14477Y 47700 92400 Cold leg L1334 Ser.14457 40800 84800 Cold leg L1334 Ser.14457 41400 87400 X-Over leg L1359 Ser.14467Y 46200 99600 X-Over leg L1359 Ser.14467Y $0900 93400 X-Over leg L1334 Ser.14465Y 44700 88900 X-Over Leg L1334 Ser.14465Y 44900 84400 X-Over leg L1336 Ser.14466Z 41400 83600 X-Over Leg L1336 Ser.14466Z 43200 87400 X-Over Leg Ll336 Ser.14466W 41400 83600 m:\2172w.wpf:1b/021996 4-5

Table 4-1 (Cont.)

Measured Tensile Properties for Byron Unit 1 Primary Loop Piping Material SA376 Gr. 304N Component Heat Number Yield Room Temp. Ultimatc (psi) Room Temp (psi)

X-Over Leg Ll334 Ser.14465X 44700 88900 X-Over leg L1336 Ser.14466X 43200 87400 X-Over Leg L1334 Ser.14465W 44900 84400 m:\2172w.wpf:lb/021996 4-6

i s

Table 4-2 Measured Tensile Properties for Byron Unit 2 Primary Loop Piping Material SA376 Gr. 304N Component Heat Number Yield Room Temp. Ultimate (psi) Room Temp (psi)

Hot leg K3660 Ser.14483 47000 89900 Hot leg 47500 90900 Hot leg K3660 Ser.14484 45700 89900 Hot leg 47000 88900 Hot Leg K3660 Ser.14485 45200 88400 l Hot Leg 50000 91400 l-l Hot leg K3660 Ser.14486 46200 88700 Hot Leg 48500 90200 Cold leg K2980 Ser.11044U 42200 84900 Cold leg K3810 Ser.16116 42200 85900 l Cold Leg 42700 86400 Cold leg K2980 Ser.11044W 43700 88100 Cold Leg Ll360 Ser.14478 43700 86400 Cold leg "

50400 92200 Cold Leg K3810 Ser.16115 42500 85900 Cold leg 42900 85700 Cold Leg K3722 Ser.16099 41200 84900 Cold leg "

43400 85900 X-Over Leg K3722 Ser.15691Y 43500 87200 X-Over leg L1359 Ser.14467Z 46200 89600 X-Over leg 50900 93400 X-Over Leg K3723 Ser.15692Z 45000 87400 X-Over leg 48000 91400 X-Over Leg L1334 Ser.14465Z 44700 88900 MW W%

i X-Over Leg K3722 Ser.15691X 43700 86900 I

m:\2172w.wpf:1b/021996 4-7

i Table 4-2 (Cont.)

Measured Tensile Properties for Byron Unit 2 Primary Loop Piping i Material SA376 Gr. 304N Component Heat Number Yield Room Temp. Ultimate (psi) Room Temp i (Psi)

X-Over Leg L1359 Ser.14467W 46200 89600 X-Over Leg L1359 Ser.14467X 50900 93400 l

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m:\2172w.wpf:1b/021996 4-8

i

)

l Table 4-3 l

Measured Tensile Properties for Braidwood Unit 1 Primary Loop Piping Material SA376 Gr. 304N Component Heat Number Yleid Ultimate Room Temp. Room Temp (psi) (psi)

Hot Leg L3384 Ser. 27223 42700 84100 l

Hot leg L3384 Ser. 27224 42700 84100 Hot leg L3384 Ser. 27225 42700 84100 Hot leg L3387 Ser. 27229Z 42900 83600 Cold leg J6199 Ser. 21741Z 39900 82700 l l- Cold leg J6199 Ser. 21741Y 39900 82700 Cold leg 16200 Ser. 27146Z 42400 80000 Cold leg J6200 Ser. 27146Y 42400 80000 Cold Leg J6200 Ser. 27145Z 42400 80000 Cold leg J6200 Ser. 27145Y 42400 80000 Cold leg 16200 Ser. 27144Z 42400 80000

! Cold leg 36200 Ser. 27146Y 42400 80000 X-Over leg L3385 41200 83900 X-Over Leg L3385 41200 83900 l

X-Over Leg L3386 42200 85900 X-Over leg L3386 42200 85900 t

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mA2172w.wpf:1b/021996 49

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Table 4-4 Measured Tensile Properties for Braidwood Unit 2 Primary Loop Piping Material SA376 Gr. 304N Component Heat Number Yield Ultimate Room Temp. Room Temp (psi) (psi)

Hot leg J6343 Ser. 27698Z 41200 83900 Hot leg J6343 Ser. 27698Y 41200 83900 Hot leg 16349 Ser. 27700Z 39900 82400 Hot leg J6349 Ser. 27700Y 39900 82400 Cold Leg 16344 Ser. 27672Y 41700 84600 42400 84400 Cold leg J6344 Ser.27673Z 41700 84600 42400 84400 Cold leg J6344 Ser. 27672X 41700 84600 42400 84400 Cold Leg 16344 Ser. 27672Z 41700 84600 42400 84400 Cold Leg 36347 Ser. 27674Y 41900 84100 Cold leg 36347 Ser. 27675Y 41900 84100 Cold leg J6347 Ser. 27675X 41900 84100 Cold leg J6347 Ser.27674Z 41900 84100 X-Over leg 16342 Ser. 27645Y 40400 83600 X-Over leg J6341 Ser. 27642Y 41800 84900 X-Over Leg L3386 42600 87000 X-Over 12g L3386 42600 87000 m:\2172w.wpf:1b/021996 4-10

1 1

l Table 4 5 l

Measured Room Temperature Tensile Properties for Byron Unit 1 Primary Loop Elbow F ttings Component Heat Number Yleid Ultimate Material Room Temp. Room Temp (Psi) (psi)

Hot leg 97192 2 35600 84950 SA351-CF8A Hot leg 02158 1 36400 81600 SA351-CF8A Hot leg 00763-5 37750 82300 SA351-CF8A Hot leg 99556-2 38200 84050 SA351-CF8A Cold Leg 02588-4 38500 79750 SA351-CF8A Cold leg 97577-7 35500 80800 SA351-CFBA Cold Leg 04252 2 43050 86050 SA351-CF8A Cold leg 04303 2 42250 93100 SA351-CF8A X-Over Leg 01064-2 38500 82700 SA351-CF8A X-Over Leg 99599-3 41350 86200 SA351-CF8A l X-Over leg 03109 1 39400 82350 SA351-CF8A X-Over Leg 01285-2 39100 85450 SA351-CF8A X-Over leg 02231 1 39250 85800 SA351-CF8A l X-Over Leg 04506-2 40600 85200 SA351-CF8A X-Over Leg 02772-1 40300 83100 SA351-CF8A X-Over leg 04171 1 40000 82800 SA351-CF8A X-Over Leg 99386-1 41950 88400 SA351-CFBA l

X-Over 12g 02930-1 40000 82300 SA351-CF8A X-Over leg 02727 1 39850 83950 SA351-CF8A

! X-Over leg 07128-1 44400 89350 SA351-CFBA l

l l

l m:\2172w.wpf:1b/021996 4-11 l

Table 4 6 Measured Room Temperature Tensile Pro prties for Byron Unit 2 Primary Loop Elbow F ttings j C- - , rat Heat Number Yield Ultimate Material i

Room Temp. Room Temp (psi) (psi)

Hot leg 05337-1 40600 89700 SA351-CF8A Hot leg 04342-3 40350 85400 SA351-CF8A Hot leg 03755-2 39250 83300 SA351-CF8A Hot Leg 03692-2 41450 86950 SA351-CF8A Cold leg 04567-1 38050 84200 SA351-CF8A Cold leg 04640-3 37900 82850 SA351-CF8A Cold leg 04640 4 37900 82850 SA351-CFBA Cold leg 04606-2 42250 86300 SA351-CF8A X-Over leg 04567-1 38050 84200 SA351-CFBA X-Over leg 04603-2 39900 84400 SA351-CF8A X-Over leg 04817 2 37900 79850 SA351-CF8A X-Over Leg 046401 37900 82850 SA351-CF8A  ;

X-Over Leg 02576-1 38950 82650 SA351 CF8A X-Over leg 04146-1 43950 91450 SA351-CF8A X-Over leg 04050-1 38500 82700 SA351-CF8A X-Over leg 04377-1 41100 86750 SA351-CFBA X-Over leg 02833 1 39250 82550 SA351-CF8A X-Over leg 12426 1 39550 84600 SA351-CF8A X-Over Leg 04932 1 41500 87800 SA351-CF8A X-Over leg 12669-1 42100 83200 SA351-CF8A I

m:\2172w.wpf:lb/021996 4-12

Table 4-7 Measured Room Temperature Tensile Properties for Braidwood Unit 1 Primary Loop Elbow Fittings Component Heat Number Yield Ultimate Material Room Temp. Room Temp (Psi) (psi)

Hot Leg 04270 3 39700 87400 SA351-CF8A Hot leg 04171 2 40000 82800 SA351-CF8A Hot Leg 03652-2 36100 78700 S A351-CF8A Hot Leg 09464-2 37900 78850 SA351-CF8A Cold leg 04606-1 42250 86300 SA351-CF8A Cold leg 04641-3 37900 82850 SA351-CF8A Cold Leg 04464-1 40000 87050 SA351-CF8A Cold Leg 04641-2 37900 82850 SA351-CF8A X-Over Leg 02772-2 40300 83100 SA351-CF8A X-Over Leg 03002-1 36250 84800 SA351-CF8A X-Over Leg 04567-2 38050 84200 SA351-CF8A X-Over Leg 04568-2 43200 87650 ' SA351 CF8A X-Over leg 04431 1 37000 79350 SA351-CF8A X-Over leg 04252-1 43050 86050 SA351-CF8A X-Over leg 14936-1 37000 81550 SA351-CF8A X-Over Leg 15822 2 37150 78300 SA351-CF8A X-Over Leg 13670-1 36100 79450 SA351-CF8A X-Over Leg 16141 1 38350 79050 SA351-CF8A X-Over Leg 02801-1 ,41100 82200 SA351-CF8A X-Over leg 18524-1 40600 86000 SA351-CF8A m:\2172w.wpf:1b/021996 4-13

I Table 4-8  !

Measured Room Temperature Tensile Properties for Braidwood Unit 2 Primary Loop Elbow Fittings Component Heat Number Yleid Ultimate Material Room Temp. Room Temp (psi) (psi)

Hot leg - 04407-1 39400 86900 SA351-CF8A Hot lag 04253-1 39100 85700 SA351-CF8A Hot leg 04146-2 43950 91450 SA351-CF8A Hot leg 03804-2 36550 78900 SA351-CFBA Cold Leg 04567-4 38050 84200 SA351-CF8A Cold leg 04407-1 39400 86900 SA351-CFBA Cold leg 07006 1 37900 86400 SA351-CF8A Cold leg i1431-2 39100 85950 SA351-CFBA X-Over leg 16778-1 38350 81350 SA351-CFBA l

X-Over leg 17743-2 38450 82950 SA351-CF8A X-Over leg 15271 1 38950 81700 SA351-CP8A X-Over leg 16875-1 36550 79900 SA351-CF8A X-Over Leg 13674-1 40600 80200 SA351-CF8A X-Over leg 18683-1 39950 83850 SA351-CF8A X-Over leg 18264-1 39750 84050 SA351-CF8A X-Over Leg 18587 1 38800 84350 SA351 CF8A X-Over Leg 04640-2 37900 82850 SA351-CFBA X-Over Leg 05337-3 40600 89700 SA351-CFBA X-Over Leg 07006-3 37900 86400 SA351-CF8A X-Over leg 07006-2 37900 86400 SA351-CF8A m:\2172w.wpf:1b/021996 4-14

Table 4-9 Mechanical Properties for Byron and Braidwood Units 1 and 2 Materials at Operating Temperatures a,C.c m:\2172w.wpf:1b/0219% 4-15

._ - _ _ _ . . ._ . . _ _ _ _ _ _ _ _ _ . . ~ _ .. _ .__.

Table 410 ,

Fracture Toughness Properties for Byron and Braidwood Units 1 and 2 Prhnary Loops for  ;

Leak-Before-Break Evaluation at Critical Location

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4 m:\2172w.wpf:Ib/021996 4-16

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i Figure 4-1 Representative Lower Bound True Stress - True Strain Curve for SA351 CF8A at 617'F l

P i

m:\2172w.wpf:lb/021996 4-17

.4

+

a,c.e 1.

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1 Figure 4-2 Pre-Service J vs. As for Cast Stainless Steel at 600*F m:\2172w.wpf:Ib/021996 4.}g

4 a.C,e

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i' Fipm 43 res for Aged Material l I 5 H rs i

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m:\2172w.wpf:ltW21996 4-19

4 SECTION 5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA 5.1 Critical Locations The leak-before-break (LBB) evaluation margins are to be demonstrated for the limiting locations (governing locations). Candidate locations are designated load critical locations or  !

toughness critical location as discussed in Section 3.0. Such locations are established based l on the loads (Section 3.0) and the material properties established in Section 4.0. These locations are defined below for Byron and Braidwood Units 1 and 2. Table 3-2 as well as Figure 3-2 are used for this evaluation.

l Load Critical Location l The highest stressed location for the SA376 Gr. 304N straight pipes is Location 11 at the '

reactor coolant pump outlet nozzle to pipe weld. Furthermore, since it is on a straight pipe, it is a high toughness location.

Toughness Critical Location Low toughness locations are the elbows. All the elbows for the Byron and Braidwood Units 1 and 2 primary loop as indicated in page B-1 of Appendix B, exceed the toughness of [

].'" Hence , for all the elbows the toughness allowables are[ ]'" The highest stressed elbow is at location 3 where the temperature is 617'F. Location 3 is most limiting since it has the highest stress and the temperature is also higher. It is thus concluded that the enveloping location is 3. The allowable toughness for the critical location are shown in Table 4-10.

5.2 Fracture Criteria As will be discussed later, fracture mechanics analyses are made based on loads and postulated flaw sizes related to leakage. The stability criteria against which the calculated J and tearing modulus are compared are:

(1) If J,pp < Ja, then the crack is stable; mA2172w.wpf:1b/021996 5-1

6 (2) If J,pp 2 Ju, but, if Tapp < Tmat and J,pp < Jmax, then the crack is stable.

Where: J = Applied J app J

Ic

= J at Crack Initiation T,pp

= Applied Tearing Modulus T = Maedal Tearing Wulus max J

mu

= Maximum J value of the material These criteria apply to the toughness critical locations. For critical locations, the limit load method discussed in Section 7.0 is used.

1 i

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m:\2172w.wpf:1b/021996 5-2 .

I

SECTION 6.0 LEAK RATE PREDICTIONS 6.1 Introduction The purpose of this section is to discuss the method which is used to predict the flow through postulated through-wall cracks and present the leak rate calculation results for through-wall circumferential cracks. )

6.2 General Considerations )

1 l

The flow of hot pressurized water through an opening to a lower back pressure causes flashing which can result in choking. For long channels where the ratio of the channel length, L, to hydraulic diameter, D ' H) is greater than [ ]'##, both [

H 1

l I

]'##. j l

J 6.3 Calculation Method l The basic method used in the leak rate calculations is the method developed by [

]'##

The flow rate through a crack was calculated in the following manner. Figure 6-1 from Reference 6-1 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass flow, the [

]'## was found from Figure 6-2 (taken from Reference 6-1). For all cases considered, since [ ]'##

Therefore, this method will yield the two-phase pressure drop due to momentum effects as illustrated in Figure 6-3, Po is the operating pressure. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using m:\2172w.wpf:1b/021996 6-1

4 AP, = [ -]"" (6-1) where the friction factor f is determined using the [ ]"' The crack relative roughness, e, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ ]"#

The frictional pressure drop using equation 6-1 is then calculated for the assumed flow rate and added to the [ momentum pressure drop calculated using the Fauske model]"' to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary loop Absolute Pressure - 14.7 = [ ]"# (6-2) for a given assumed flow rate 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 the procedure is repeated until equation 6-2 is satisfied to within an acceptable tolerance which in turn leads .o correct flow rate value for a given crack size.

6.4 Leak Rate Calculations l

l Leak rate calculations were made as a function of crack length at the governing locations previously identified in Section 5.1. The normal operating loads of Table 3-1 were applied, 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 l

The flaw sizes to yield a leak rate of 10 gpm were calculated at the governing locations and are given in Table 6-1. The flaw sizes so determined are called leakane flaws.

The Byron and Braidwood Units 1 and 2 RCS pressure boundary leak detection system 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 yield a leak rate of 10 gpm.

m:\2172w.wpf:Ib/021996 6-2

6.5 References 1

6-1 [

pu, I

6-2 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 II-1, NUREG/CR-3464, September 1983.

6-3 WCAP-9558 Rev. 2, " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack," Westinghouse Proprietary Class 2, June 1981.

e

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m:\2172w.wpf:1b/0221% 6-3

Table 6-1 Flaw Sizes Yleiding a Leak Rate of 10 gym at the Governing Locations l

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4 Figure 6-1 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures

.i

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't

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m -- -. >w-w- - ,-

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I 5m 62 [ }* Pressure Ratio as a Function of UD m:\2172w.wpf:ltW21996 6_6

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a,c.e a,c,e

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~j V~~

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---

r-r X l

Figure 6-3 Idealized Pressure Drop Profile: Through a Postulated Crack m:\2172w.wpf:lb/0219% 6-7

l l SECTION 7.0 FRACTURE MECHANICS EVALUATION 7.1 Local Failure Mechanism

I The local mechanism of failure is primarily dominated by the crack tip behavior in terms of crack-tip blunting, initiation, extension and finally crack instability. The local stability will be 4

assumed if the crack does not initiate at all. It has been accepted that the initiation toughness measured in terms of Ja from a J-integral resistance curve is a material parameter defining the crack initiation. If, for a given load, the calculated J-integral value is shown to be less than the Ju of the material, then the crack will not initiate. If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation: l i E '

T* = da 0 a2 J

4 where: 1 1

T,, = applied tearing modulus.

E = modulus of elasticity

, o, = 0.5 (o, + c,) (flow stress) a = crack length

o,,c, = yield and ultimate strength of the material, respectively

Stability is said to exist when ductile tearing occurs if T,, is less than T., the experimentally determined tearing modulus. Since a constant T is assumed a further restriction is placed in J.,. J , must be less than J where J. is the maximum value of J for which the j, experimental T is greater than or equal to the T used.

As discussed in Section 5.2 the local crack stability will be established by the two-step I criteria:

(1) If J app < I k, then the crack will not initiate.

1 (2) If J,,2 Ju, but, if T , < T.

4 i and J., < J., then the crack is stable.

s m:\2172w.wpf:lb/021996 7-1

l e

7.2 Global Failure Mechanism 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 l remaining net section reaches a stress level at which a plastic hinge is formed. The stress level at which this occurs is termed as the flow stress. The flow stress is generally taken as l the average of the yield and ultimate tensile strength of the material at the temperature of i 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 l provided in appendix A for a through-wall circumferential flaw in a pipe with internal  ;

l pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:  ;

l l

[ Jue i

where:

[

3u o, = 0.5 (a, + c,) (flow stress), psi

[ ju.e

[

1 1

i l 3"#

s 1

m:\2172w.wpf:lb/021996 7-2

L 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).

l For application of the limit load methodology, the material, including consideration of the configuration, must have a sufficient ductility and ductile tearing resistance to sustain the limit load.

! 7.3 Results of Crack Stability Evaluation 1 i

l Stability analyses were performed 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),

l h

The lower-bound material properties of Section 4.0 were applied (see Table 4-9). The  !

fracture toughness properties established in Section 4.3 and the normal plus SSE loads given l in Table 3-2 were used for the EPFM calculations. Evaluations were performed at the toughness critical locations identified in Section 5.1. The results of the elastic-plastic fracture mechanics J-integral evaluations are given in Table 7-1.

The critical locations were also identified in Section 5.1. A stability analysis based on limit load was performed for these locations as described in Section 7.2. The welds at these locations are TIG and SMAW combination. The "Z" factor correction for SMAW was l applied (Reference 7-3) as follows:

4 Z=1.15[1.0+0.013(OD-4)]

where OD is the outer diameter of the pipe in inches.

The Z-factors were calculated for the toughness critical and load criticallocations. The Z factors were 1.60 and 1.57 for locations 3 and 1I respectively. The applied loads were increased by the Z factors and plots of limit load versus crack length were generated as

! shown in Figures 7-2 and 7-3. Table 7-2 summarizes the results of the stability analyses

{ based on limit load. The leakage size flaws are also presented on the same table.

+

m:\2172w.wpf:Ib/0221% 7-3

A 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.

7-2. Kumar, V., German, M. D. and Shih, C. P., "An Engineering Approach for Elastic-Plastic Fracture Analysis," EPRI Report NP-1931, Project 1237-1, Electric Power Research Institute, July 1981.

7-3. Standard Review Plan; Public Comment Solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal RegisterNol. 52, No.167/ Friday, August 28,1987/ Notices, pp. 32626-32633.

I I

l i

m:\2172w.wpf:1b/021996 7-4

l 1

1 Table 7-1 Stability Results for Byron and Braidwood Units 1 and 2 Based on Elastic-Plastic J Integral Evaluations i

Fracture Criteria Calculated Values Flaw Size J,, Tmat J max J ap Tapp l IAcation (in) 2 (in-lb/in ) (in lb/in') 2 (in-ib[n ) j a,c.e l

1 i

l m:\2172w.wpf:1b/021996 7-5

=. - . - . .._ . . . . - _ - . _ - . . . . .

9 Table 7-2 Stability Results for Byron and Braidwood Units 1 and 2 Based on Limit Load Leakage Location Flaw Size (in.) Flaw Size (in) a,c,e m:\2172w.wpf:1b/021996 7-6

\

a,c,e I

20

) Neutral Axis

\N ))

V l

Figure 7-1 [ ]*' Stress Distribution m:\2172w.wpf:1b/021N 7~7

a.c.e 1 1

l I

i l

OD = 34.21 in a, = 21.57 ksi F, = 2573 kips t = 2.50 in o, = 66.30 ksi M. = 22713 in-kips SA351 CF8A Material With SMAW Weld 1

i Figure 7-2 Critical Flaw Size Prediction - Hot Leg at Location 3 m:\2172w.wpf:1b/021995 7-8

, , . _ $ , __A.. _. . . . . - _ . . _ _ _ _ _ , . _ _ _ . _ . s.__a - . . . .a . _ . .

l i

a,c e

(

l l

l l

l OD = 32.14 in o y= 23.11 ksi F, = 2275 kips t = 2.215 in o, =70.35 ksi M. = 31856 in-kips SA376 Gr. 304N Material With SMAW Weld l

l l

Figure 7-3 Critical Flaw Size Prediction - Cold Leg at Location 11 m:\2172w.wpf:lb/022096 79

~

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 growth analysis was carried out for the [ ]'" region of a typical system (see Location [ ]'" of Figure 3-2). This region was selected because crack growth calculated here will be typical of that in the entire primary loop. Crack growths calculated at other locations can be expected to show less than 10% variation.

A[ ]'" of a plant typical in geometry and operational characteristics to any Westinghouse PWR System.

[

]'" All normal, upset, and test conditions were considered. A summary of generic applied transients is provided in Table 8-1. Circumferentially oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in Figure 8-1. Specifically, these were:

Cross Section A: [ ]'"

Cross Section B: [ l'"

Cross Section C: [ ]'"

Fatigue crack growth rate laws were used [

l'" The law for stainless steel was derived from Reference 8-1, with a very conservative correction for the R ratio, which is the ratio of minimum to maximum stress during a transient. For stainless steel, the fatigue crack growth formula is:

b = (5.4 x 10a2) K,y"' inches / cycle dn l

l where K,y = K (1-R)o.5

! R=K,iA m:\2172w.wpf:Ib/021996 8-1

[

]""

[ ]

where: [ ]""

where AK is the stress intensity factor range.

The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 8-2, and shows that the crack growth is very small, [ ]"#

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.

1 8-2 [

ju.e 8-3 [

.]"'

m:\2172w.wpf:lb/021996 8-2

Table 8-1 Summary of Reactor Vessel Transients Number Typical Transient Identification Number of Cycles Normal Conditions 1

1 Heatup and Cooldown at 100'F/hr 200 l (pressurizer cooldown 200 F/hr) l 2 Load Follow Cycles 18300 (Unit loading and unloading at 5%

of full power / min) 3 Step load increase and decrease 2000 1

4 Large step load decrease, with steam dump 200 5 Steady state fluctuations 6 10 Upset Conditions 6 Loss of load, without immediate turbine 80 or reactor trip 7 Loss of power (blackout with natural 40 I circulation in the Reactor Coolant System) l 8 Loss of Flow (partial loss of flow, one 80 pump only) 9 Reactor trip from full power 400 Test Conditions 10 Turbine roll test 10 11 Hydrostatic test conditions '

Primary side 5 Primary side leak test 50 12 Cold Hydrostatic test 10 m:\2172w.wpf:1b/021996 8-3

Table 8 2 Typical Fatigue Crack Growth at

[ ]C (40 years)

FINAL FLAW (in.)

Initial Flaw (in.) [ ]* [ ]* [ ]*

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:\2172w.wpf:Ib/021996 8-4 i

e s

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i Figure 8-1 Typical Cross-Section of [ jw l

l f

t m:\2172w.wpf:1b/0219% 85

u o

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I Figure 8-2 Reference Fatigue Crack Growth Curves for [

]*

m:\2172w.wpf:lb/021996 8-6

. . - . . ~

1 a.c.e 1 l

l l

1 l

l l

l i-l 1

! Figure 8-3 Reference Fatigue Crack Growth Law for [ ]* in a Water Environment at 600*F i

i i

m:\2172w.wpf:lb/021996 8-7 i

I

I  !

SECTION 9.0 l ASSESSMENT OF MARGINS 1

The results of the leak rates of Section 6.4 and the corresponding stability and fracture l toughness evaluations of Sections 7.1,7.2 and 7.3 are used in performing the assessment of margins. Margins are shown in Table 9-1.

i In summary, at all the critical locations relative to:

1. Flaw Size At location 3 using faulted loads obtained by the absolute sum method, a margin  !

of more than 2 exists between the critical flaw and the flaw having a leak rate of )

10 gpm (the leakage flaw).

At location 11 using faulted loads obtained by the absolute sum method, a margin of 1.91 exists between the critical flaw and the flaw having a leak rate of 10 gpm I (the leakage flaw).

2. Leak Rate - A margin of 10 exists between the calculated leak rate from the l leakage flaw and the leak detection capability of 1 gpm. I

3. Loads - At the critical locations the leakage flaw was shown to be stable using l

the faulted loads obtained by the absolute sum method.

l l

l l

m:\2172w.wpf:lb/021996 9-1 i

1 Table 91 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Byron and Braidwood Units 1 and 2

_ Location Leakane Flaw Size Critical Flaw Size Margin __

a,c.e l

based on limit load 6

based on J integral evaluation 1

m:\2172w.wpf:1b/021996 9-2

l

! SECTION

10.0 CONCLUSION

S This report justifies the elimination of RCS primary loop pipe breaks from the structural  ;

l design basis for the Byron and Braidwood Units 1 and 2 as follows:

l

a. Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, l

pressure, aid flow during normal operation.

b. Water hammer should not occur in the RCS piping because of system design, I l

l testmg, and operational considerations.

c. The effects of low and high cycle fatigue on the integrity of the primary piping l

are negligible.

l

) d. Ample margin exists between the leak rate of small stable flaws and the l capability of the Byron and Braidwood Units 1 and 2 reactor coolant system l

pressure boundary Leakage Detection System.

l

e. Ample margin exists between the small stable flaw sizes 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 described in d, e, and f above.

Based on the above, it is concluded that dynamic effects of RCS primary loop pipe breaks need not be considered in the structural design basis of the Byron and Braidwood Units 1 and 2 Nuclear Power Plant.

I 1

m:\2172w.wpf:lb/021996 10-1 l . _ _ _ _ _ _ - _ _

l l

APPENDIX A LIMIT MOMENT l

l

[

1 i

l l

l l

l l

l l

l Ja.c.e m:\2172w.wpf:1b/021996 A-1

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i d

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l Figure A-1 Pipe with a Through-Wall Crack in Bending m:\2172w.wpf:lb/021996 A-2

APPENDIX B TOUGHNESS CRITERIA FOR BYRON AND BRAIDWOOD UNITS 1 AND 2 CAST PRIMARY LOOP COMPONENTS All of the individual cast piping components of the Byron and Braidwood Units 1 and 2 primary loop piping satisfy the original [ ]'" criteria (Reference 4-5). [

Jae.e i

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l I

1 l

m:\2172w.wpf:lb/021996 B-1

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s 4

Table B-1

Chemistry and Fracture Toughness Properties of the Material Heats of Byron Unit 1 3

__ _ a,c,e l

I l

l l

i m:\2172w.wpf:Ib/021996 B-2 l

._ . __ _ _ _ _ . . . ~ _ . . _. . . .,_~._.. ___. _ _ _ _ _

Table B-1 (Cont.)

- _ a,c.e l

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l l

I 1

1 m:\2172w.wpf:1b/021996 B-3

Table B 2 Chemistry and Fracture Toughness Pronerties of the Material Heats of Byron Unit 2 ,

I

- _ 8,C,e m:\2172w.wpf:lb/021996 B-4

Table B-2 (Cont.)

_ a.c.e 1

1 i

~ _

I m:\2172w.wpf:lb/021996 B-5

. .. . . ~ . . . . - . . . . . . - . . . - - - . . - . .. .. .

Table B-3 Chemistry and Fracture Toughness Pr_operties of the Material Heats of Braidwood Unit 1

_ _ a.c.e m:\2172w.wpf:lb/021996 B-6

a

, i

. 1 Table B 3 (Cont.)

~

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l e

1 N

TA:\2172w.wPf :IWO21996 B-7

Table B-4 1

Chemist and Fracture Toughness Pros,erties of the aterial Heats of Braidwood Unit 2 l

_ _ a,c.e i

J l

l

)

l l

m:\2172w.wpf:1b/021996 B-8

Table B-4 (Cont.)

_ a,c.e e

1 m:\2172w.wpf:lb/021996 B-9

._ _. q O

o a,C,0 t

b B-10

l i

l l

ATTACHMENT A WCAP-14559 Revision 1, " Technical Justification for Eliminating large Primary Loop Pipe Rupture as the Structural Design Basis for the Byron and Braidwood Units 1 and 2 Nuclear Power Plants"- Proprietary Version l

4

~i i

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