ML16341G487

Nonproprietary Technical Justification for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for Diablo Canyon Units 1 & 2 Nuclear Power Plants.

ML16341G487
Person / Time
Site:	Diablo Canyon
Issue date:	11/30/1991
From:	Adamonis D, Bhowmick D, Schmertz J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
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ML16341G488	List: ML16341G487 ML17083C404 ML17083C405
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WCAP-13038, NUDOCS 9203230343
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Westinghouse Class 3 WCAP-13038 TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE DIABLO CANYON UNITS 1 AND 2 NUCLEAR POWER PLANTS NOVEHBER 1991 J. C. Schmertz S. A. Swamy F. J. Witt Y. S. Lee VERIFIED:

D. Bho ic APPROVED

. C. Adamonis, Acting Manager Structural Mechanics Technology Work Performed Under Shop Order LgXP 950 WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P. 0. Box 2728 Pittsburgh, Pennsylvania 15230-272&

1991 Westinghouse Electric Corp.

All Rights Reserved 9203230343 920316 PDR ADOCK 05000275 P

PDR llP F0924 J/112791: 10

5 ~l

'1 0

FOREWORD This document contains Westinghouse Electric Corporation proprietary information and data which has been identified 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 proprietary information and data contained in this report were obtained at considerable Westinghouse expense and its release could seriously affect our competitive position. This information's 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 such persons as part of the review procedure, please contact Westinghouse Electric Corporation, which will make the necessary arrangements required to protect the Corporation's proprietary interests.

The proprietary information is deleted in the unclassified version of this report (WCAP-13038).

lP F0921 J/112691: 10

0 TABLE OF CONTENTS Section Title ~Pa e EXECUTIVE

SUMMARY

X1

1.0 INTRODUCTION

1-1

1. 1 Purpose 1.2 Scope 1.3 Objectives 1-2 1 ' Background Information 1-2 1.5 References 1-4 2.0 OPERATION AND STABILITY OF THE REACTOR 2-1 COOLANT SYSTEM

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 2-5 3.0 PIPE GEOMETRY ANO LOAOING 3-1

3. 1 Introduction to Methodology 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 Analysis 3.5 References 3-,4 4.0 MATERIAL CHARACTERIZATION 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

'MPF0921J/112691:10 iv

TABLE OF CONTENTS Section Title Pacae 5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA 5-1

5. 1 Critical Locations 5-1 5.2 Fracture Criteria 5-2 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 8-3 9.0 ASSESSMENT OF MARGINS 9-1

10.0 CONCLUSION

10-1 MP F 0921 J/112691: 10

TABLE OF CONTENTS Section Title Pacae APPENDIX A- Limit Moment A-1 APPENDIX B- Alternate Toughness Criteria for the B-1 Diablo Canyon Units 1 and 2 Cast Primary Loop Components B. 1 Introduction B-1 B.2 Chemistry and KCU Toughness B-1 B.3 Alternative Toughness Criteria for B-1 the Diablo Canyon Primary Loop Material on a Component by Component Basis B.4 References B-2 MP F 0921 J/112691: 10 V1

LIST OF TABLES Table Title Pacae 3-1 Dimensions, Normal Loads and,Normal Stresses for 3-5 Diablo Canyon 3-2 Faulted Loads and Stresses for Diablo Canyon 3-6 4-1 Measured Room Temperature Tensile Properties for 4-6 Diablo Canyon Unit 1 Primary Loop Piping and Fittings 4-2 Measured Room Temperature Tensile Properties for 4-10 Diablo Canyon Unit 2 Primary Loop Piping and Fittings Typical Tensile Properties of SA376 TP316, SA351 4-14 CFSA and Welds of Such Material for the Primary Loop 4-4 Mechanical Properties of SA351 CFSM Material at 4-15 Room Temperature (From a Typical PWR Plant) 4-5 Mechanical Properties of SA351 CFSM Material at 4-17 650'F (From a Typical PWR Plant) 4-6 Mechanical Properties for Diablo Canyon Units 1 5. 2 4-19 Materials at 544'F and 618'F 4-7 Enveloped Fracture Toughness Properties for Diablo 4-20 Canyon 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 MP F0921 J/112691:10 vii

LIST OF TABLES (Cont'd)

Table Ti tl e ~Pa e 7-1 Stability Results for Diablo Canyon Units 1 7-6 and 2 Based on Elastic-Pl'astic J- Integral Evaluations 8-1 Summary of Reactor Vessel Transients 8-4 8-2 Typical Fatigue Crack Growth at [

]

' (40 Years) 9-1 Summary Table 9-2 8-1 Chemistry and Fracture Toughness Properties of the 8-3 Material Heats of Diablo Canyon Unit 1 B-2 Chemistry and Fracture Toughness Properties of the Material Heats of Diablo Canyon Unit 2 llPF0921 J/112691:10 ,

V111

LIST OF FIGURES

~Fi ere Ti tl e Pa(ac 3-1 Hot Leg Coolant Pipe 3-7 3-2 Schematic Diagram of Diablo Canyon Primary Loop 3-8 Showing Weld Locations Lower Bound True Stress-True Strain Curve for the 4-21 SA351 CFBH Haterial of Diablo Canyon Unit 1 at 544'F 4-2 J vs. ha for SA351 CF8H Cast Stainless Steel at 600'F 4-22 4-3 J vs. ha at Different Temperatures for Aged Haterial 4-23

[ ]

' (7500 Hours at 400'C)

-i Analytical Predictions of Critical Flow Rates of Steam-Water Hixtures 6-5 6-2 ] ' Pressure Ratio as a Function 6-6 of L/0 6-3 Idealized Pressure Drop Profile Through a Postulated 6-7 Crack MP F 0921 J/112691: 10 1X

LIST OF FIGURES (Cont'd)

~Fi uee Title Pacae 7-1 ]

' Stress Distribution 7-7 7-2 Critical Flaw Size Prediction - Hot Leg at Location 1 7-8 for Diablo Canyon 8-1 F

Typical Cross-Section of [ >a,c,e 8-6 8-2 Reference Fatigue Crack Growth Curves for [

)a,cee 8-3 Reference Fatigue Crack Growth Law for [ )a,c,e 8-8 in a Water Environment at 600'F A-1 Pipe with a Through-Wall Crack in Bending A+

WP F0921 J/112691: 10 X

EXECUTIVE SUHHARY The existing structural design basis for the reactor coolant systems of the Diablo Canyon Units 1 and 2 nuclear reactor power plants requires that the dynamic effects of pipe breaks be 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 should not be included, in general, in the structural design basis of Westinghouse type pressurized water reactors, for example. To eliminate primary loop pipe breaks from the design basis, it must be demonstrated to the satisfaction of the U.S. Nuclear Regulatory Commission that a leak-before-break situation exists. This report provides such a demonstration for the Diablo Canyon Units 1 and 2 nuclear power plants.

I n 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 Haterials Test Reports. Hechanical properties were determined at operating temperatures. Since the piping systems are fabricated 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 against flaw instability. Finally, fatigue crack growth was shown not to be an issue for the primary loops.

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 Diablo Canyon Units 1 and 2 nuclear power plants.

MP F 0921 J/112691: 10 xi

SECTION

1.0 INTRODUCTION

1.1 ~Pur oee This report applies to the Diablo Canyon Nuclear Power Plant Units 1 and 2 (Diablo Canyon) Reactor Coolant System (RCS) primary loop piping. It is intended to demonstrate that for the specific parameters of the Diablo Canyon 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 ~Sco e The existing structural design basis for the RCS primary loop requires that dynamic effects of pipe breaks be evaluated. Specifically, as part of the LOCA design basis for the Diablo Canyon plants the following breaks are postulated in the RCS primary loop piping: the six terminal ends in the cold, hot, and crossover legs; a split in the steam generator inlet elbow, and the loop closure weld in the crossover leg. 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 Westinghouse plants (see reference 1-2). In order to demonstrate this applicability of the generic evaluations to the Diablo Canyon plants, Westinghouse has performed a fracture mechanics 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 leak-before-break (LBB) methodology. Through this successful application of the LBB methodology, the above eight break locations in the RCS primary loop piping are eliminated from Diablo Canyons plants'tructural design basis.

MP F0921 J/112691: 10

1.3 ~0b'ectives In order to validate the elimination of RCS primary loop pipe breaks for the Diablo Canyon plants, the following objectives must be achieved:

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 Diablo Canyon plants..

c. Demonstrate margin on applied load.

d. Demonstrate that fatigue crack growth is ne'gligible.

1.4 Back round Information Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated from 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 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.

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).

The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to address this same issue using a probabilistic ap'proach. 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 NPF0921J/112691:10 1-2

(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

-12 probability of a direct LOCA (RCS primary loop pipe break) to be 4.4 x 10 per reactor year and the mean probability of an indirect LOCA to be 10 per reactor year. Thus, the results previously obtained 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 conclusioris 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 Rupture" (reference 1-7).

This report provides a fracture mechanics demonstration of primary loop integrity for the Diablo Canyon 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 USNRC Generic letter 84-04,

Subject:

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

MPF0921J/112691:10 1-3

1-2 Letter from Westinghouse (E. P. Rahe) to NRC (R. H. Vollmer),

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.

1-4 Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G.

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.

1-7 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/Vol. 52, No.

207/Tuesday, October 27, 1987/Rules and Regulations, pp. 41288-41295.

NP F09210/112691:10 1-4

SECTION 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Crackin 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). This operating history totals over 450 reactor-years, including five plants each having over 17 years of operation and 15 other plants each with over 12 years of operation.

In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group. (The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of'the objectives of the- second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pressurized Water Reactors (PWR's). The results of the study performed by the PCSG were presented in NUREG-0531 (reference 2-1) entitled " Investigation and Evaluation of Stress 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 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."

MP F 0921 J/112691: 10 2-1

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 qualifies the PCSG's findings.

For stress corrosion cracking (SCC) to occur in piping, the following th'ree conditions must 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 ASHE 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 MP F0921 J/112691: 10 2-2

the ppb range by controlling 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 Overall, there is a low potential for water hammer in the RCS since it is designed and 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 valve opening are considered in .the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control 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 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 C cle and Hi h C cle Fati ue Low cycle fatigue considerations are generally. accounted for in the design of the piping system. 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 llP F 0921 J/112691: 10 2-3

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 exeeedence of,the vibration limits. Field measurements have been made on a number of plants during hot functional testing, including plants similar to the Diablo Canyon 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.

E 2-4

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 Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.

VP F0921 J/112691: 10 2-5

e SECTION 3.0 PIPE GEOMETRY AND LOADING

3. I Introduction to Methodolo The general approach is discussed first. As an example a segment of the primary coolant hot leg pipe is shown in figure 3-1. The as-built outside diameter and minimum wall thickness of the pipe are 33.99 in. and 2.395 in.,

respectively, as seen in the figure. Normal stresses at the weld locations result from the load combination procedure discussed in section 3.3 while faulted loads are developed as outlined 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 later the highest stressed location in the entire loop is at the reactor vessel outlet nozzle to pipe weld. This location is called the 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 thus the normal loads and faulted loads must be available to assess leakage and stability, respectively. The loads (developed below) at this location are also given in Figure 3-1.

Since the elbows 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 tou hness critical locations.

The single most critical location is apparent only after the full analysis is completed. Once loads (this section) and fracture toughnesses (section 4.0) are available, the load critical and toughness critical locations 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 stabilit mar g ins are also evaluated ( see section 9.0).

The locations for evaluation are those shown in figure 3-2.

MPF0921J/'l12691:10 3-1

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

Q = F+ M A Z

where, g m stress f m axial load H 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= My+Mz 2

where, H bending moment for required loading Y component of bending moment MZ

= Z component of bending moment 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.

QP F0921 J/112691: 10 3-2

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

DW TH P (3-3)

Y

( Y)DW ( Y)TH ( Y)P (3-4)

Z ( Z)DW ( Z)TH ( Z)P (3-5)

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

OW = deadweight TH = normal thermal expansion P - load due to internal pressure This method of combining loads is often referred as the al ebraic sum method.

The loads based on this method of combination are provided in table 3-1 at all the locations identified in Figure 3-2. The as-built dimensions are also given.

3.4 Load Combination for Crack Stabilit Anal sis In accordance with Standard Review Plan 3.6.3 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 <2 to 1.0. The absolute summation of loads results in the following equations:

I DMI I TH I I P I I SSEIHERTIA I I (3-6)

SSEAHi'Y

=I ("Y)DMI I ("Y)THI I ("Y)PI I ("Y)SSEIHERTIAI (g-7)

I ( Y) SSEAH I HP F0921 J/112691: 10 3-3

2 I ( Z)0'Ml+I (~Z)THI I ( Z)PI I ( Z)SSEIHERTIAI (3-8)

I ( Z)SSEAHI where subscripts SSE, INERTIA and AH mean safe shutdown earthquake, inertia and anchor motion, respectively.

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. The loads at all the locations of interest (see Figure 3-2) are summarized in Tables 3-1 and 3-2.

3.5 References 3-1 USNRC Standard Review Plan 3.6.3, Leak-Before-Break Evaluation Procedures, NUREG-0800.

HP F 0921 J/112691: 10 3-4

TABLE 3-1 DIMENSIONS, NORMAL LOADS AND NORMAL STRESSES FOR DIABLO CANYON Outside Axiaak Bending Diameter Thickness Load Moment Stress Location (in.) (in.) (kips) (in-kips) (ksi) 33.99 2.395 1495 25057 20.57 33.99 2.395 1478 1202 6.98 34.81 2.80 1582 7865 9.41 37.19 2.99 1491 19818 12.44

37. 19 2.99 1601 4689 6.84 37.19 2.99 1596 4349 6.69 37.19 2.99 1591 4318 6.64 37.19 2.99 1692 5848 7.51 37.19 2.99 1692 8471 5.86 10 37.19 2.99 1833 5982 8.07 32.26 2.275 1348 3456 8.59 12 32.26 2.275 1348 3422 8.57 13 33.06 2.675 1347 3491 7.23 14 33.06 2.675 1348 2728 6.80 See figure 3-2 Includes pressure MP F0921 J/112691:10 3-5

TABLE 3-2 FAULTED LOADS AND STRESSES FOR DIABLO CANYON Location

'xial Load (kips)

Bending Moment (in-kips)

Stress (ksi) 2161 46674 36.39 1933 6583 11.88 1943 12833 13.05 2360 41261 23.56 2006 28517 17.44 1955 23992 15.52 7 r 1919 16968 12.67 1845 12891 10.8 1839 12853 10.77 10 2053 42718 23.17 1826 31245 29.32 12 1668 12728 16.25 13 1667 13018 13.78 14 1701 14482 14.72 See Figure 3-2 See table 3-1 for dimensions Includes pressure MPF0921J/'l12691:10 3-6

Crack I

00 OD = 33.99 in t = 2.395 in Normal Loads Faulted Loads forcea: 1495 kips force: 2161 kips bending moment: 25057 in-kips bending moment: 46674 in-kips Includes the force due to a pressure of 2250 psi Figure 3-1 Hot Leg Coolant Pipe f

0515. wp /11279 1: 10 3-7

RSLCTON NfSSQRg VHSQ, 14 13 10 HOT LEG Temperature 6184F, Pressure: 2250 psi Temperature 544 F, Pressure: 2250 psi COLI~Lg Temperature 5444F, Pressure: 2250 psi Figure 3-2 Schematic Diagram of Diablo Canyon Primary Loop Showing Weld Locations 0515.wpf/080791:10 3-8

SECTION 4.0 MATERIAL CHARACTERIZATION 4.1 Primar Loo Pi e and Fittin s Materials The primary loop pipe materials are SA376 TP316 for both Diablo Canyon units, and the elbow fittings are SA351 CFSH for both units.

4.2 Tensile Pro erties The Certified Materials Test Reports (CMTRs) for Diablo Canyon Units 1 and 2 were used to establish the tensile properties for the leak-before-break analyses. The CHTRs include tensile properties at room temperature for each of the heats of material. These properties are given for Diablo Canyon Units 1 and 2 in Tables 4-1 and 4-2 respectively. The average properties are given and the lower bound properties are identified. The 1989 Code minimum properties are also given in these tables.

The properties at 544'F and 618'F were established from the tensile properties at room temperature given in Tables 4-1 and 4-2 by utilization of Tables 4-3, 4-4, and 4-5. These last three tables provide typical tensile properties at room temperature and at 650'F for both of the materials of concern. The tensile properties for typical materials at 544'F and 618'F were obtained by interpolating between the room temperature and the 650'F tensile properties given in Tables 4-3, 4-4, and 4-5. Ratios of the strengths at 544'F and 618'F to the strengths at room temperature for 'the typical materials were then applied to the room temperature values given in Tables 4-1 and 4-2 to obtain the Diablo Canyon properties at 544'F and 618'F.

In Table 4-3, the SA376 TP316 material properties of 'plant A'ere closer to the Diablo Canyon material properties than Plant C, and therefore the 'plant A'alues were used. In Tables 4-4 and 4-5 the tensile properties for all the listed components were averaged, to obtain the typical tensile properties for SA351 CFSM.

'NP F0921 J/112691: 10 4-1

The average and lower bound yield strengths and ultimate strengths, to encompass both Diablo Canyon Units I and 2, are given in Table 4-6. The ASME Code Modulus of Elasticity is also given, and Poisson's ratio was taken as 0.3.

For leak-before-break fracture evaluations of the toughness critical locations the true stress-true strain curve must be available. This curve was obtained using the Nuclear Systems Materials Handbook (reference 4-1). The lower bound true stress-true strain curve is given in Figure 4-1.

4.3 Fracture Tou hness Pro erties The pre-service fracture toughnesses of cast materials in terms of J have been

,found to be very high at 600'F. Typical results for a cast material are given in figure 4-2 taken from reference 4-2. JIIc is observed to be over 5000 in-lbs/in . However, cast stainless steels are subject to thermal aging during -service. This 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 ferrite in the material.

To determine the effects of thermal aging on piping integrity, a detailed study was carried out in reference 4-3. In that report, fracture toughness results were presented for a material [

]

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

vpF0921 J/112691:10 =4-2

End-of-service life toughnesses for the heats are established using the alternate 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 it can be demonstrated that its end-of-service fracture toughnesses equal or exceed those of f

~a,c,e The worst case fracture toughness values for all the loops of each plant at each location (see figure 3-2), as taken from Appendix B, are given in table 4-7.

Available data on aged stainless steel welds (references 4-3 and 4-4) indicate that JI 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 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 temperature . Therefore, weld regions are less =limiting than the cast material.

In the report all J l. d values were conservatively determined by using base metal strength properties.

MPF0921 J/112691:10 4-3

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 having 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 [ ',c,e

]

In the fracture mechanics analyses that follow, the fracture toughness properties given in table 4-7 will be used as the criteria against which the applied fracture toughness values will be compared.

4.4 References 4-1 Nuclear Systems Materials Handbook, Part I - Structural Materials, Group 1 - High Alloy Steels, Section 2, EROA Report TID 26666, November, 1975, 4-2 "Mechanistic Fracture Evaluation of Reactor Coolant 0

WCAP-9558 Rev. 2, 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., 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.

4-5 Appendix II of Letter from Dominic C. Di Ianni, NRC to D. M. Musolf, Northern States Power Company, Docket Nos. 50-282 and 50-306, December 22, 1986.

IIP F0921 J/112691: 10 4 4

4-6 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).

le f0921 J/112691:10 4-5

TABLE 4-1 MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 1 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Cold Leg V0630/3259 SA376 TP316 43 ' 87.6 42.4 84.1 Cold Leg V0629/3334 SA376 TP316 42.4 84. 1

41. 9 81.2 Cold Leg K2011/3862 SA376 TP316 34.7 77.0 36.0 78.4 Cold Leg V0630/3261 SA376 TP316 38.5 78.5 42.9 86.3 Cold Leg '1478/3258 SA376 TP315 37.0 77 33.0 81 Cold Leg K2010/3684 SA376 TP316 38.6 81. 8 35.2 77.2 Cold Leg E1478/3256 SA376 TP316 45.5 84. 1 43.4 87.9 Cold Leg K2011/3680 SA376 TP316 34.1 75. 3 43 ' 82.9 Hot Leg E1490/3358 SA376 TP316 43.5 85'. 4 44.9 84.9 Hot Leg E1485/33524 SA376 TP316 42.3 85.4 45.4 87. 4 Hot Leg V0688/3539 SA376 TP316 39.9 81.7 40.2 81.7 Hot Leg E1484/3349X SA376 TP316 39.0 79. 1 42.9 85.9 0515. wpf /080791: 10

TABLE 4-1 (continued)

MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 1 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Hot Leg E1484/33494 SA376 TP316 Hot Leg E1483/3350X SA376 TP316 44 F 1 86.2 47.3 83.4 Hot Leg E1483/3350Y SA376 TP316 44. 1 86.2 47.3 83.4 Hot Leg V0684/3537X SA376 TP316 38.9 79. 8.

40+6 80.7 Crossover Leg 1 F-0222/2902X SA376 TP316 43.5 85'. 1 44.3 85. 6 C

Crossover Leg 2 F0222/29024 SA376 TP316 43 ' 85. 1 44.3 85.6 Crossover Leg 3 F0216/2863X SA376 TP316 43.7 86.7 39.5 82 Crossover Leg 4 E1485/3361X SA376 TP316 38.3 83. 9 45.9 91.4 Crossover Leg 1 E1493/3381X SA376 TP316 41 0~

43 ' 86.9 Crossover Leg 2 E1493/33814 SA376 TP316 41 '

43.5 86.9 Crossover Leg 3 K2029/44384 SA376 TP316 Q.3 40.0 82.4 Crossover Leg 4 E1468/3365 SA376 TP316 42.0 ~ ~

44.0 Average 0515. wpi/080791: 10 4-7

TABLE 4-1 (continued)

MEASURED ROOM TVG'ERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 1 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Cold Leg 30845-1 IT NG '. BOWS SA351 CF8M 36.5 76.9 Cold Leg 33047-4 SA351 CF8M 42.3 85.3 Cold Leg 3242514 SA351 CF8M 42.5 85.8 Cold Leg 32469-2 SA351 CF8M 47.25 89.75 Hot Leg 10966-1 SA351 CF8M 45.0 87.5 Hot 'Leg 11937-2 SA351 CF8M 43.5 88.0 Hot Leg 12198-2 SA351 CF8M 42.0 85.0 Hot Leg 10563-2 SA351 CF8M 45.0 Crossover Leg 1 13174-2 SA351 CF8M 48.0 85.0 Crossover Leg 2 13704-1 SA351 CF8M 48.0 86.5 Crossover Leg 3 16654-2 SA351 CF8M 36.0 74.0 Crossover Leg 4 16690-2 SA351 CF8M 39 '

Crossover Leg 1 13421-1 SA351 CF8M 46.5 80 3 14576-1 SA351 CF8M 40.5 Crossover Leg 2 14804-1 SA351 CF8M 51.0 93 15560-1 SA35 1 CF8M 43.5 Crossover Leg 3 15987-2 SA351 CF8M 45. 0 15823-1 SA351 CF8M 42.0 83.

Crossover Leg 4 14535-2 SA351 CF8M 41. 5 16037-3 SA351 CF8M 37. 5 0515. @pi'/080791: 10 4-8

TABLE 4-1 (continued)

MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 1 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Streng Component No. Heat No. Material (ksi) (ksi)

Crossover Leg 1 16690-3 SA351 CF8M 40.5 81.75 17388-2 SA351 CF8M 37.5 75.0 Crossover Leg 2 13930-5 SA351 CF8M 48.0 89.0 14251-1 SA35 1 CF8M 39 ' 82.5 Crossover Leg 3 11743-1 SA351 CF8M 45. 0 85.0 11556-1 SA351 CF8M 43.5 83.5 Crossover Leg 4 12281-1 SA351 CF8M 42.0 84.75 11974-1 SA351 CF8M 43.0 83.0 Average 42.89 83.58 a - indicates not available b - underline indicates the lower bound values 1989 Code Minimum Properties:

SA376 TP316 30 '75 SA35 1 CF8M 30 70 0515. wpf /080791: 10 4-9

TABLE 4-2 MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 2 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Cold Leg F0654/4153 SA376 TP316 38 ' 81.6 42.4 86. 1 Cold Leg 5167b/4555 SA376 TP316 38.2 81.6 41.2 83 '

Cold Leg F0654/4152 SA376 TP316 38 ~ 7 83. 1 43 ' 86.0 cold Leg K2029/4557 SA376 TP316 38.7 NL0,6 36.4 88. 1 Cold Leg F0653/4151X SA376 TP316 40 ' 83.1 41 ' 83.6 Cold Leg 51676/4556X SA376 TP316 37. 5 41.9 Cold Leg J1677/4554X SA376 TP316 33 ' 81.6 40 ' 82.6 Cold Leg J1681/4677 SA376 TP316 43.7 87.4 41 2 ~ 83.6 Hot Leg F0655/4132 SA376 TP316 42.4 86.9 48.3 91.6 Hot Leg F0656/4131 SA376 TP316 38.7 83.0 39.1 81. 5 Hot Leg F0654/4136 SA376 TP316 38 F 9 82. 1 47.4 86. 5 Hot Leg J1682/4627 SA376 TP316 44 ' 87.5 51.0 93.1 0515.vpf/080791:10 4-10

TABLE 4-2 (continued)

MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 2 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Hot Leg K2027/4480Y SA376 TP316 37.4 79.9 36 ' 80.9 Hot Leg F0656/4131X SA376 TP316 49 ' 88.6 50 ' 92.1 Hot Leg K2027/4499 SA376 TP316 37 2~ 80.6 44.2 87.1 Hot Leg J1676/4631X SA376 TP316 38 ' 81.9 47.8 87.3 Crossover Leg 1 J-1681/4560X SA376 TP316 45 ' 84.9 40.7 81. 1 Crossover Leg 2 J-1681/4561X SA376 TP316 37.7 81.9

41. 2 84.8 Crossover Leg 3 J-1682/4669X SA376 TP316 43.7 85.9 45 ' 87.5 Crossover Leg 4 J1681/4562Y SA376 TP316 43 ' 83.9 40 ' 83.6 Crossover Leg 1 J-1681/4562X SA376 TP316 43.3 83.9 40.9 83.6 Crossover Leg 2 J-1682/4563Y SA376 TP316 46 ' 84.3 Crossover Leg 3 J-1681/4561Y SA376 TP316 37 ' 81.9 41 ' 84.9 Crossover Leg 4 J 1682/4563X SA376 TP316 39 ' 82.0 46.7 84.3 Average g4.'4 0515. vpf /080791: 10 4-11

TABLE 4-2 (continued)

MEASURED ROOM TEHPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 2 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

ITTINGS i .e. OWS Cold Leg 40874-4 SA351 CFBH 43.8 86.1 Cold Leg 40874-3 SA351 CFBH 43.8 86.1 Cold Leg 39716-4 SA351 CFBH 43.8 87.1 Cold Leg 39716-1 SA351 CF8M 43.8 87.1 Hot Leg 33756-2 SA351 CFBH 44.25 84.0 Hot Leg 41192-2 SA351 CFBH 34.3 73.3 Hot Leg 37168-1 SA351 CF8M 38.65 76.3 Hot Leg 39792-3 SA351 CFBM 42.3 84.3 Crossover Leg 1 38408-3 SA351 CFBH 43.6 83,1 Crossover Leg 2 39231-3 SA351 CFBH 46.8 89.6 Crossover Leg 3 39445-2 SA351 CFBH 42.3 83.5 Crossover Leg 4 34900-2 SA351 CFBH 33.75 70.5 Crossover Leg 1 55485-1 SA351 CF8H 43.3 86.35 55485-2 SA351 CF8M Crossover Leg 2 519223-3 SA351 CFBH 42.45 86.9 52665-1 SA351 CFBH 43.05 86.3 Crossover Leg 3 55282-2 SA351 CFBM 41.7 85.65 52369-2 SA351 CF8M 41.85 85.75 Crossover Leg 4 51712-2 SA351 CF8H 39.4 80.35 52665-2 SA351 CFBH 42.45 88.7

&F0921 J/112691: 10 4-12

TABLE 4-2 (continued)

MEASURED ROOM TEMPERATURE TENSILE PROPERTIES FOR DIABLO CANYON UNIT 2 PRIMARY LOOP PIPING AND FITTINGS Yield Ultimate Loop Strength Strength Component No. Heat No. Material (ksi) (ksi)

Crossover Leg 1 49686-1 SA351 CF8M 46.95 89.95 50362-2 SA351 CF8M 41 7~ 86. 15 Crossover Leg 2 41423-1 SA351 CF8M 40 ' 79. 25 52449-1 SA351 CF8M 42.75 87. 5 Crossover Leg 3 51712-1 SA351 CF8M 38.25 79.50 52369-1 SA351 CF8M, 42.60 88.00 Crossover Leg 4 34900-1 SA351 CF8M 70.0 50115-1 SA351 CF8M 42.45 81.9 Average 41.57 83.38.

a - indicates not available b - underline indicates the lover bound values I

1989 Code Minimum Properties:

SA376 .TP316 30 75 SA351 CF8M 30 70 0515 ~ vpf /080791: 10 4-13

TABLE 4-3 TYPICAL TENSILE PROPERTIES QF SA376 TP3 16 g SA351 CF8A AND WELDS OF SUCH MATERIAL FOR THE PRIMARY LOOP Test Temperature Ave a e Tens' o erties Plant Material ('F) Yield (psi) Ultimate (psi) a,Ce4 0515. vpf /080791: 10 4-14

TABLE 4-4 MECHANICAL PROPERTIES OF SA351 CF8M MATERIAL AT ROOM TEMPERATURE (FROM A TYPICAL PHR PLANT)

TEST 0 '4 OFFSET ULTIMATE PRODUCT FORM PIECE YIELD STRESS STRENGTH \ REDUCTION HEAT NR NR MATERIAL (PSI) (PSI) ELONGATION IN AREA a,c,e 0515.wpf/080791:10 4-15

TABLE 4-4 (continued)

MECHANICAL PROPERTIES OF SA351 CF8M MATERIAL AT ROOM TEMPERATURE (FROM A TYPICAL PNR PLANT)

TEST 0. 24 OFFSET ULTIMATE PIECE YIELD STRESS STRENGTH REDUCTION PRODUCT FORM HEAT NR NR MATERIAL (PSI) (PSI) ELONGATION IN AREA a,c,e 051 /080791:10

TABLE 4-5 MECHANICAL PROPERTIES OF SA351 CF8M MATERIAL AT 650 F (FROM A TYPICAL PWR PLANT)

TEST 0.2\ OFFSET ULTIMATE PIECE YIELD STRESS STRENGTH \ \ REDUCTION PRODUCT FORM HEAT NR NR MATERIAL (PSI) (PSI) ELONGATION IN AREA B.c,e 4-17

TABLE 4-5 (continued)

MECHANICAL PROPERTIES OF SA351 CFSM MATERIAL AT 650~F (FROM h TYPICAL PWR PLANT)

TEST 0. 24 OFFSET ULTIMATE PIECE YIELD STRESS STRENGTH 4 REDUCTION HEAT NR NR (PSI) (PSI) ELONGATION IN ARE1 a,c.e 0 c 05] fj080791:10

TABLE 4-6 HECHANICAL PROPERTIES POR DIABLO CANYON UNITS 1 AND 2 HATERIALS AT 5iiiP AND 618iP Temperature Yield Stress Ultiaate Strength Yield Stress Ultiaate Strength Material ( r) (psi) . (Vsi) (psi) (psi) a,c,e Hodulus of Elasticity for Both Haterials:

SA376 TP316 E ~ 25.21x106 psi at 618'F SA351 CF8H - E ~ 25.58x106 psi at 5iiiF Poisson's Ratio: 0.3 4-19 0515. vpf /080791: 10

TABLE i-7 ENVELOPED FRACTURE TOUGHNESS PROPERTIES FOR DIABLO CANYON UNITS 1 AND 2 PRIMARY LOOPS FOR LEAK-BEFORE-BREAK EVALUATION J IC aat aax L tiki.b HT NO. (daJ/ca ) (in-lb/in2 ) (non-dia) (in-lb/in2 )

a,c,e 60 2200 111 location ?i.3 750 The locations are shorn in figure 3-2.

The lover of the values for all the loops are given here.

reference i-5.

A value of 3000 in-lb/in is acceptable per 051 f /080791: 10

a.c.e Figure 4-1 Representative Lower Bound True Stress True Strain Curve for the SA351 CFSM at 5444F 0515.wpf/080791:10 4-21

a,c,e Figure 4-2 J vs. ha for SA351 CF8M Cast Stainless Steel at 600'F WP0515.wpf/080291:10 .4-22

Figure 4-3 J Vs. 4a at Different Temperatures for Aged Material

]

' (7500 Hours at 400'C)

WP0515.wpf/080291:10 4-23

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 location (governing location). Candidate locations are designated load critical locations or toughness critical locations as discussed in Section 3.0. Such locations are established considering the loads (section 3.0) and the material properties established in section 4.0. These locations are defined below for Diablo Canyon Units 1 and 2. Table 3-2 as well as Table 4-7 are used for this evaluation, along with Figure 3-2.

Diablo Can on Unit 1 Location 1 is the highest stressed location and is the load critical location by definition. Furthermore, since it is on a straight pipe, it is a high toughness location. Low toughness locations are at the ends of every elbow.

On an elbow-by-elbow basis, these are location pairs 3 and 4, 5 and 6, 7 and 8, 9 and 10, and lastly 13 and 14. The higher stressed of each location pair are weld locations 4, 6, 7, 10, and 13. Weld Location 10 has a faulted stress comparable to weld location 4, but it is less tough. Because of the lower toughness, weld location 10 is more limiting. Weld location 10 has a toughness comparable to weld location 13, but it has a much higher faulted stress. Because of the higher stress, location 10 is more limiting. It is thus concluded that the enveloping locations in Diablo Canyon Unit 1 for which LBB methodology is to be applied are locations 1, 6, 7 and 10.

Diablo Can on Unit 2 Location 1 is the highest stressed location and is thus the load critical location. Since this location is at the higher temperature (i.e. has the worst tensile properties) and none of the other Unit 2 locations has toughnesses as low as the low toughness locations of Unit 1, the Unit 1 toughness critical locations envelope the Unit 2 toughness critical locations.

MPF0921J/112691:10 5-1

5.2 Fracture Criteria As 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 (i.e. J ) and tearing modulus (T ) are compared app are:

(I) If J app < JI, Ic' then the crack is stable; Ic'hen, if Ta mat and Ja < Jma, the crack is stable.

The toughness criteria at each location have previously been determined and are given in table 4-7.

MPF0921 J/112691:10 5-2

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 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, DH, (L/DH) is greater than [ ] ' , both [

]a,c,e 6.3 Calculation Method The basic method used in the leak rate calculations is the method developed by

[Fauske (Reference 6-1) for the two-phase choked flow, and then adding to it.

the additional frictional pressure loss upstream of the choked exit plane.]

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.

Now using the assumed flow rate, G, the frictional pressure drop can be calculated using i% F0921 J/112691: 10 6-1

a,c.e where the friction factor f is determined using the [

]

' The crack relative roughness, e, was obtained from fatigue these calculations was [ )

crack data on stainless steel samples. The relative roughness value used in a,c,e The frictional pressure drop using, equation 6-1 is then calculated for the assumed flow and added to the [

] ' to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary loop Absolute Pressure - 14.7 = [

)a,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 the 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 Rate Calculations Leak rate calculations were made as a function of crack length at the four locations previously identified in section 5. 1 The normal operating loads of Table 3-1 were applied, in'hese 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.

T he flaw sizes to yield a leak rate of 10 gpm were calculated at the four locations and are given in Table 6-1 ~ The flaw sizes so determined are called k f'1 MP F0921 J/112691: 10 6-2

The Diablo Canyon plant 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.

6.5 References 6-1 [

~a,c,e 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-I, NUREG/CR-3464, September 1983.

MP F0921 J/112691: 10 ,'-3

TABLE 6-1 FLAM SZZES YZELDZNG A LEAK RATE OF 10 GPM AT THE FOUR.LOCATZONS 0515. vpf/080791: 10 6-4

Figure 6-1. Analytical Predictions of Critical Flow Rates of Steam-Hater Mixtures 6-5

a,c,e Figure 6-2. [ j ' Pressure Ratio as a Function of L/0 4895e>11RNO'.10 6-6

Figure 6-3. Idealized Pressure Orop Profile Through a Postulated Crack

~ 495szl2i490'll 6-7

~-

SECTION 7.0 FRACTURE MECHANICS EVALUATION

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, extension and finally crack instability. The local stability will be assumed if the crack does not initiate at all. It has been accepted that the initiation toughness measured in terms of J Ic 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 JI 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:

dJ' T 2 Qg where:

applied tearing modulus app E modulus of elasticity Gf 0.5 (o + cr) (flow stress) a crack length yield and ultimate strength of the material, respectively Stability is said to exist when ductile tearing occurs if T ,is less than T t, the experimentally determined tearing modulus. Since a constant T t is a further restriction is placed in J mat'ssumed

. J must be less than J where J is the maximum value of J for which the experimental T is greater than or equal to the T t used.

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

(1) If K < JI, then the crack is stable.

llP F 0921 J/112691: 10 7-1

(2) If Ja ) JI, then, if T <

Tmat and J < J , the crack is stable.

app max'.2 Global Failure Mechanism Oetermination 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 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 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 force, and imposed bending moments. The limit moment for such a pipe is given by:

]a,c,e where:

p internal pressure

]a,c,e MPF092'IJ/112691:10 7-2

] a,c,e 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* 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 Stabilit Evaluation 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).

The lower-bound material properties of section 4.0 were applied (see Table 4-6). The fracture toughness properties established in section 4.3 (see Table 4-7) and the normal plus SSE loads given 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 leakage size flaws are presented on the same table.

The load critical location was identified as location 1 in section 5. 1. A stability analysis based on limit load was performed for this location as described in section 7.2. Since the weld at this location is a SMAW weld, the "Z" factor correction for SMAW welds was applied (Reference 7-3) as follows:

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

NP F0921 J/112691:10 7-3

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

1-factor was calculated for critical location 1, using the dimensions 0

given in table 3-1. This factor is 1.6. The applied loads were increased by the Z factor and a plot of limit load versus crack length was generated as shown in Figure 7-2. The critical flaw size is [ ]""'nd the leakage flaw sizes is [ ]""'ower bound base metal properties taken from ASHE BPVC Section III were used for this purpose. They are o = 18.7 ksi and 0 71.8 ksi.

7.4 References 7-1. Kanninen, M. F., et. al., "Hechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks," EPRI NP-192, September 1976.

0 MP F0921 J/112691: 10 7-4

r 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. ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.

QPF0921 J/112691:10 7-5

TABLE 7-1 STABILITY RESULTS FOR DIABLO CANYON UNITS 1 AND 2 BASED ON ELASTIC-PLASTIC J- INTEGRAL EVALUATIONS Fracture Criteria Calculated Values Flaw Size Ic mat max app app Location (in) (in-lb/in2 ) (in-lb/in2 ) (in-lb/in2 )

~-

MP F0921 J/112691:10 7-6

Figure 7-1

]

' Stress Distribution WP0515.wpf/072991:10 7-7

a,c,e OD = 33.99 in cr Y

= 18.7 ksi F~ = 2161 kips t = 2 395 in e= 71. 8 ksi Mb = 46674 in-kips (SMAW Weld)

Figure 7-2 Critical Flaw Size Prediction Hot Leg at Location 1 0515.wpf/112791:10 7-8

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 [ ] 'f Figure 3-2). This region was selected because crack growth calculated her e 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 the 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: ]a,c,e Cross Section B: ]a,c,e Cross Section C: ]a,c,e Fatigue crack growth rate laws were used [

]

' 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:

i (5.4 x

= 10 '2) KeEE E~4 4'inches/cycle MPF0921J/112691:10 8-1

where Keff Kmax (1-R)0.5 min max

~a,c,e

]a,c,e where: ]a,c,e

'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, [

~a,c,e le F0921 J/112691: 10

'-2

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.

8-2 [

~a,c,e 8-3 [

)a,c,e MP F0921 J/112691: 10 8-3

TABLE 8-1

SUMMARY

OF REACTOR VESSEL TRANSIENTS NUMBER TYPICAL TRANSIENT IDENTIFICATION NUMBER OF CYCLES Normal Conditions Heatup and Cooldown at 100'F/hr 200 (pressurizer cooldown 200'F/hr)

Load Follow Cycles 18300 (Unit loading and unloading at 5%

'of full power/min)

Step load increase and decrease 2000 Large step load decrease, with steam dump 200 Steady state fluctuations 10 U set Co ditions Loss of load, without immediate turbine 80 or reactor trip Loss of power (blackout with natural 40 circulation in the Reactor Coolant System)

Loss of Flow (partial loss of flow, one 80 pump only)

Reactor trip from full power 400 Test Conditions 10 Turbine roll test 10 Hydrostatic test conditions Primary side 5 Primary side leak test 50 12 Cold Hydrostatic test 10 WP F0921 J/112691:10 8-4

TABLE 8-2 TYPICAL FATIGUE CRACK GROWTH AT

' (40 YEARS) t: ]

FINAL FLAW in

~a,c,e )a,c,e ~a,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 MP F0921 J/112691: 10 ,8-5

deCoO Figure 8-1 Typical Cross-Section of [

>a,c,e WP0515.wpf/080291:10 8-6

a.c. e Figure 8-2. Reference Fatigue Crack Growth Curves for

{'d,c,e tl94el112t00 10 8-7

e.c. e Figure 8-3. Reference Fatigue Crack Growth Law for [

in a Mater Environment at 600'F M04e/112500: I 0 8-8

SECTION 9.0 ASSESSMENT OF MARGINS The results of the leak rates of section 6.4 and the corresponding stability and fracture toughness evaluations of sections 7.2 and 7.3 are used in performing the assessment of margins. Margins are shown in Table 9-1.

In summary, at all the critical locations relative to:

1. Flaw Size - Using faulted loads obtained by the absolute sum method, a margin of about 2 exists between the critical flaw and the flaw having a leak rate of 10 gpm (the leakage flaw).

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

3. Loads - At the critical locations the leakage flaw was shown to be stable using the faulted loads obtained by the absolute sum method.

WP F 0921 J/112691: 10 9-1

TABLE 9-1

SUMMARY

TABLE Location Leakage Flaw Size Critical Flaw Size Margin a,c,e 0515. Mpt/080791: 10 9-2

SECTION

10.0 CONCLUSION

S This report justifies the elimination of RCS primary loop pipe breaks for the Diablo Canyon Units 1 and 2 nuclear plants as follows:

'a ~ Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.

b. Water hammer should not occur in the RCS piping because of system design, testing, and operational considerations.

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

d. Adequate margin exists between the leak rate of small stable flaws and the capability of the Diablo Canyon Units 1 and 2 reactor coolant system pressure boundary Leakage Detection System.

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 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 Diablo Canyon Units 1 and 2 plants.

WP F 0921 J/112691: 10 10-1

I APPENDIX A LIMIT MOMENT

~a,c,e NP F0921 J/112691: 10 A-I

a,c,e-Figure A-1 Pipe with a Through-Wall Crack in Bending WP0515.wpf/080291:10 A-2

For example, [ ]"'the 90'lbow in the crossover leg on the steam generator side of Diablo Canyon Unit 1) has the calculated end-of-service life KCU at room temperature of [ ]'"'aj/cm which falls below that of [ ]""'. The 8-ferrite content is [ ]""'. By Reference 4-6, the [

].""'ince the end-of-service life KCU exceeds the fully aged KCU, the heat falls into category 2. Thus:

J m

[

]a,c,e T, [

]a,c,e and J [

l a,c,e The fracture toughness values for each heat of material was calculated as formulated in References 4-6. These values are also given in Tables 8-1 and B-2.

B.4 References B-1 Letter: Dominic C., Di Ianni, NRC to D. M. Musolf, Northern States Power Company, dated December 22, 1986, Docket Nos. 50-282 and 50-306.

MP F0921 J/112691: 10 B-2

APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE DIABLO CANYON UNITS 1 AND 2 CAST PRIMARY LOOP COMPONENTS B. 1 Introduction Not all of the individual cast piping components of the Diablo Canyon primary loop piping satisfy the original [ ]""'riteria (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 individual 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 Chemistr and KCU Tou hness Per the procedure of Reference 4-6 the correlations 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-l for Unit 1 and in Table B-2 for Unit 2.

B.3 Alternate Toughness Criteria for the Diablo Canyon Primary Loo Material on a Com onent b Com onent Basis The alternate toughness criteria for the Diablo Canyon 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 47 of the 53 heats fall into category 1, i.e., they are at least as tough as [

]"'. 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.

MP F0921 J/112691: 10 . B-1

TABLE B- I CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF DIABLO CANYON UNIT I WP F0921 J/112691:10 8-3

a,c,e B-4

a,c,e B-5

B-6 a,c,e B-7

Leceend:

KCU - Charpy U-notch energy in daj/cm JIC40 - J at 40 years service in in-lb/in~

TMAT40 - T , at 40 years service JHAX40 - J at 40 years service in in-lb/in JICFA - The fully aged J in in-lb/in~

THATFA - The fully aged T , (always 0)

JHAXFA - The fully aged J in in-lb/in (always equals JICFA)

Ni - nickel, C - carbon, Cr - chromium, Si - silicon Ho - molybdenum, Cb - columbium, Cr(E) - chromium equivalent, Ni(E) - nickel equivalent, Chemistry is in percent weight DELTA is ferrite in percent volume (see reference 4-6 for discussion of category, fully aged toughness, etc.)

MP F0921 J/112691: 10 B-8

TABLE B-2 CHEMISTRY AND FRACTURE TOUGHNESS PROPERTIES OF THE MATERIAL HEATS OF DIABLO CANYON UNIT 2 NPF0921 J/112691:10 B-9

a,c B-10

a,c,e 1

4 I

t t

B-11

r yI/

B-12

a,c,e I'

Ill B-13

~Le end:

KCU - Charpy U-notch energy in daj/cm JIC40 - J at 40 years service in in-lb/ina THAT40 - T , at 40 years service JHAX40 - J at 40 years service in in-lb/in~

JICFA - The fully aged J in in-lb/in THATFA - The fully aged T , (always 0)

JHAXFA - The fully aged J in in-lb/in (always equals JICFA)

Ni - nickel, C - carbon, Cr - chromium, Si - silicon Ho - molybdenum, Cb - columbium, Cr(E) - chromium equivalent, Ni(E) - nickel equivalent, Chemistry is in percent weight DELTA is ferrite in percent 'volume (see reference 4-6 for discussion of category, fully aged toughness, etc.)

MP F0921 J/112691: 10 B-14