ML20148A540

Nonproprietary Technical Justification for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for Trojan Plant

ML20148A540
Person / Time
Site:	Trojan File:Portland General Electric icon.png
Issue date:	02/29/1988
From:	Johnson E, Kim C, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19302D318	List: ML20148A525 ML20148A533 ML20148A540
References
TAC-12369, WCAP-11700, NUDOCS 8803210005
Download: ML20148A540 (88)

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WCAP-11700 4

TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE TROJAN PLANT FEBRUARY 1988 S. A. Swamy E. R. Johnson F. J. Witt C. C. Kim C. B. Bond V. M. Bhambri Y. S. Lee O

C. SM Verifiedby:d.'~C.Schmertz g

Approved by:

5. 5. Pflusamy, Manager Structural Materials Engineering Work Performed Under Shop Order PPLJ950 WESTINGHOUSE ELECTRIC CORPORATION Generation Technology Systems Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 9

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TABLE OF' CONTENTS Section Title Page 1-1 1.0 INTRDDUCTION 1-1 1.1 Purpose 1-1 1.2 Scope 1-1 1.3 Objectives 1-1 1.4 Background Information 1-4 1.5 References 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 2-1 2.1 Stress Corrosion Cracking 2-1 2-3 2.2 Water Hammer 2-4 2.3 Low Cycle and High Cycle Fatigue 2-4 2.4 References 3.0 PIPE GEOMETRY AND LOADING 3-1 3.1 Loads far Leak Rate Evaluation 3-2 3.2 Loads for Crack Stability Analysis 3-3 3.3 Alternate Load Combination for Crack Stability Analysis Based on Absolute Summation 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 4-4 4.4 References 4-6 5.0 LEAK RATE PREDICTIONS 5-1 5.1 Introduction 5-1 5.2 General Considerations 5-1 5.3 Calculation Method 5-1 5.4 Leak Rate Calculations 5-3 5.5 References 5-3 1

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TABLE OF CONTENTS (Cont'd.)

Section Title Page 6.0 FRACTURE NECHANICS EVALVATION 6-1 6.1 Local Failure Mechanism 6-1 6.2 Global failure Mechanism 6-2 6.3 Results of Crack Stability Evaluation 6-3 6.4 References 6-6 7.0 FATIGUE CRACK GROWTH ANALYSIS 7-1 7.1 References 7-3 l

8.0 ASSESSMENT

OF MARGINS 8-1

9.0 CONCLUSION

S 9-1 APPENDIX A LIMIT M0HENT A-1 l

APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE TROJAN CAST PRIMARY LOOP COMPONENTS B-1 2'res4445s 01204810 j

n

LIST OF TABLES Table No.

Title Page 3-1 Normal Condition (Dead Weight and Pressure and Thermal)

Loads for Trojan 3-5 3-2 T-ojan Primary Loop Data Including Faulted Loading Conditions 3-6 4-1 Mechanical Properties of the Primary, Loop Materials of 4-7 the Trojan Plant 4-2 Mechanical Properties of SA351 CF8M Mater,ial at 650*F (From a Typical PWR Plant) 4-9 4-3 Mechanical Properties of SA351 CF8M Material at Room Temperature (From a Typical PWR Plant) 4-10 4-4 Trojan Material Properties 650*F 4-11 4-5 Pressure and Temperature Conditions in the Primary Loop for Trojan 4-12 4-6 Fracture Toughness Criteria Used in the Leak-Before-Break Evaluation 4-13 5-1 Leak Rate Results 5-4 6-1 Results of Stability Analysis -- Margin on Flaw Size 6-7 6-2 Results of Stability Analysis -- Margin on Loads 6-8 7-1 Summary of Reactor Vessel Transients 7-4 3?Ms4dels C22eu 10 y

LIST OF TABLES (Cont'd.)

Table No.

Title Page 7-2 Fatigue Crack Growth at [

]" d 7-5 (40 years)

B-1 Chemistry and Calculated KCU Values for Each Primary Loop Piping of the Trojan Nuclear Plant B-4 B-2 fracture Toughness Criteria for the Cast Primary Piping B-8 Components of the Trojan Nuclear Plant j

nu.So. en.u io yj i

LIST OF FIGURES Figure Title Page 3-1 Reactor Coolant Pipe 3-7 3-2 Schematic Diagram of Trojan RCL Showing Weld Identification 3-8 4-1 True Stress Strain Curve for SA351 CF8A Stainless Steel at 617'F 4-14 4-2 True Stress Strain Curve for SA351 CF8A Stainless Steel at 553*F 4-15 4-3 True Stress Strain Curve for SA351 CF8M Stainless Steel at 617'F 4-16 4-4 True Stress Strain Curve for SA351 CF8M Stainless Steel at 552'F 4-17 4-5 J vs. aa for SA351 CF8H Stainless Cast Steel at 600'F 4-18 4-6 J vs. Aa at Different Temperatures for Aged Material

(

Ja c.e (7500 hours0.0868 days <br />2.083 hours <br />0.0124 weeks <br />0.00285 months <br /> at 400*C) 4-19 5-1 Analytical Predictions of Critical Flow Rates for Steam-Water Mixture 5-5 5-2

(

]a,c.e Pressure Ratio as a Function of L/D 5-6 5-3 Idealized Pressure Drop Profile Through a Postulated Crack 5-7 i

2?Ms/Duls 02244410 yjj

LIST OF FIGURES (Cont'd.)

Figure Title Page 6-1

(

Ja,c.e Stress Distribution 6-9 6-2 "Critical" Flaw Size Prediction - Hot Leg at Critical Location 1

'6-10 7-1 Typical Cross-Section of (

Ja c.e 7-6 7-2 Reference Fatigue Crack Growth Curves for (

,)a,c.e 7,7 7-3 Reference Fatigue Crack Growth Law for (

la.c.e in a Water Environment at 600*F 7-8 A-1 Pipe with a Through wall Crack in Bending A-3 B-1 Typical Layout of the Primary Loops for a Westinghouse Four-Loops Without Isolation Valves B-9 B-2 Indentification of Heats with Location for Cold Leg B-10 B-3 Identification of Heats with Location for Hot Leg B-11 1

B-4 Identification of Heats with Location for Crossover Leg B-12 nee,ws, uwa yii1

PREFACE Portland General Electric Company submitted a leak-before-break analysis for the elimination of primary loop pipe ruptures for the Trojan Nuclear plant (Impell Report No. 01-0300-1395, Rev. 1).

After completing their review, the NRC transmitted to Portland General Electric Company a request for additional information.

Portland General '

Electric Company contracted with Westinghouse Electric Corporation to respond to the NRC request including the performance of analyses.

This report presents a detailed leak-before-break evaluation including the response to the NRC request. The NRC request is reproduced here:

REQUEST FOR ADDITIONAL INFORMATION ON ELIMINATION OF POSTULATED PRIHARY LOOP PIPE RUPTURES AS A DESIGN BASIS By letter dated October 31, 1986, Portland General Electric Company (the licensee) submitted the technical basis for the elimination of primary loop pipe ruptures using "leak-before-break" (LB3) methodology for the Trojan Nuclear Plant in Impell Report No. 01-0300-1395, Rev. 1, dated October 1986.

The staff has reviewed the licensee's submittal for compliance with the revised General Design Criterion 4 (GDC-4) of appendix A, 10 CFR 50.

The evaluation criteria are discussed in detail in NUREG-1061, Volume 3. The staff requires the following additional information from the licensee before our review can continue.

(1) The LBB criteria in NUREG-1061, Volume 3 contain margins on leakage rate, crack size, and applied loads to account for uncertainties inherent in the analyses.

The LBB margins are discussed in detail in NUREG-1061, Volume 3.

The margins on crack size and applied loads were not addressed by the licensee in the submitted report.

The licensee should discuss compliance with the margin of 2 on the crack size and the margin of "the square root of 2" on the apolied loads. The licensee should demonstrate the stability of a through-wall crack twice the size cf the leakage-size crack under 27Ms 9445, 0224M 10 jy

combined normal and safe shutdown earthquake (SSE) leads. Also, the licensee should demonstrate the stability of the leakage-size crack if the loads are increased to the square root of 2 times the combination of normal and SSE loads.

(2)

In Section 2.5.1 of the submitted report, the Licensee discussed the thermal aging effects of cast stainless steel on the material fracture toughness.

However, the specific extent of thermal aging depends on the chemistry and the ferrite content of the cast material. The 'iicensee should demonstrate that the data shown in figure 5 of the submitted report provide a lower-bound estimate of 1

the toughness properties for the specific Trojan material.

Alternatively, the licensee should determine the thermally-aged f

fracture toughness for each cast stainless steel piping component material in the primary loop of Trojan based on its specific chemistry and the ferrite content.

(3) The LBB analysis should be performed for the functional piping system from anchor point to anchor point.

In figure 1 of the submitted report, it is indicated that LBB analyses were performed j

at break locations.

In Section 2.4.1 of the submitted report, the licensee indicated that several intermediate locations were also considered. The LBB methodology should be separated from the methodology in Standard Review Plan 3.6.2 which discusses break locations.

In the application of LBB, the piping system from anchor j

to anchor, including all intermediate locations, should be considered. The limiting location for LBB analyses is the location k

with the highest stresses coincident with the poorest material properties for base materials, weldments, and safe ends. The effects of thermal aging as discussed in item 2 above must be considered.

(4) Linear elastic fracture mechanics (LEFM) was used for the fracture stability analysis.

However, from the calculated fracture mechanics

{

parameter "J-integral", it appears that the associated Irwin l

plane-stress piastic zone sizes are not small, compared with the l

I 37us9649e4224M to X

i i

l i

half-crack length. The licensee should use elastic plastic fracture i

mechanics (EPFM) instead of LEFM procedures. Although_the licensee attempted to compare the plastic zone with the uncracked 4

circumference in table 6 of the submitted report, the proper comparison should be with the crack length because the crack length is less than the uncracked circumference.

Furthermore, the licensee should benchmark the EPFM procedure against experimental pipe test data.

(5)

In Section 2.4.1 of the submitted report, it is indicated that SSE loads were obtained by multiplying the maximum operating basis earthquake (OBE) loads by a factor of 1.67.

Describe the validity of obtaining SSE loads from OBE loads.

(6)

In Section 2.5.2 of the submitted report, it is indicated that four typical weld procedures were reviewed to determine the welding process of the primary loop.

From this review, the licensu -

determined that the welding processes were shielded metal arc welding (SMAW) and gas-tungsten are welding (GTAW). The licensee should review all the welding procedures for the primary loop to determine if submerged are walds (SAW) were used. Also, indicate if solution annealing was performed.

Furthermore, the staff disagrees with the licensu's assertion in Section 2.5.3 of the submitted report that the fracture toughness of GTAW would bound that of SMAW.

1 (7)

In Section 2.4.2 of the submitted report, it is indicated that ASME Code minimum material properties were used.

Because various materials were used in the fabrication of the primary loop, describe j

the Code minimum of which material was used for the calculation and justify this selection. Also, provide the elastic modulus, yield j

strength, ultimate strength, and stress-strain curve at the limiting i

location and at the operating temperature. The licensee would have to select these material properties in order to perform a EPFM l

i evaluation as discussed in item 4 above.

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i.

xi

(8)

In Section 2.4.2 of the submitted report, the licensee indicated that a fully plasti: displacement-controlled tearing modulus (J-T) analysis was performed.

Describe whether the J-T result shown in figure 8 of the submitted report was calculated for a 7-inch crack in the hot leg vessel outlet nozzle.

If the calculation was based on a 25-inch crack as discussed in Section 2.4.2 of the submitted report, it is not conservative to use the J-integral value for a 7-inch crack on the J-T curve for a 25-inch crack. Also, the licensee should use elastic plastic load-controlled J-T analysis because a fully plastic displacement-controlled J-T analysis may not be conservative. Furthermore, if crack growth is predicted before instability, crack growth should be considered in the J-T analysis.

(9)

In appendix A of the submitted report, the licensee indicated that the leakage prediction computer code has been benchinarked against test data.

The licensee should submit the benchmarking results for s

staff review.

(10) In table 1 of the submitted report, the licensee listed the diameters and wall thicknesses of the hot leg, crossover leg, and the cold leg.

Describe whether the nominal dimensions or minimum dimensions were tabulated.

If nominal dimensions were used, the licensee should review the as-built configuration of the primary loop weldments to determine whether the actual thickness is less than the nominal thickness. Due to weld fit-up, the actual thickness may be less than the nominal pipe thickness.

(11) In appendix C of the submitted report, a fatigue crack growth analysis was discussed. Describe the design transients conc Mered in the analysis. Also, show the crack growth material data used for the analysis, l

2794s/0445s*022444 10 gjj

(12) Net section plastic analysis was discussed in Section 2.6 of the submitted report.

However, net section analysis does not account

~

for material toughness limitations.

In particular, low toughness thermally-aged cast stainless steel is involved in the present LBB evaluation. The licensee should use a fracture stability analysis which accounts for material toughness.

SUMMARY

RESPONSE

This report includes detailed response to the above requests.

A summary of responses is provided below by addressing each request individually.

Reauest 1 The response to this request is given in Section 6.3.

The margins on crack size and loads are discussed in detail in Section 6.3 and Section 8.0, Recuest 2 The thermally-aged fracture toughness (end of life) for each cast stainless steel piping component material in the primary loop of the Trojan plant was established based on its chemistry and ferrite content.

Details are provided in Section 4.3 and Appendix B.

Reauest 3 The entire primary loop (from anchor to anchor) was considered in the evaluation. Five critical locations were identified based on detailed review of loads, material chemistry and material properties for the entire system (Sections 3 and 4).

Effects of thermal aging have been addressed in detail.

(Sections 4 and Appendix B).

Request 4 was used for determining J,pp and Elastic plastic fracture mechanics (EPFMi T,pp in the stabil,ity analyses of Section 6.0.

Thus, the analyses of Section 6.0 comply with this request.

27Hs/044 ts-0224M 10 jjj i

a y

e.

Request 5 The response to this request is provided in Section 3.2.

Request 6 The response to this request is provided in Section 4.0.

Request 7 Representative minimum and average properties were established based on Trojan plant specific material certification records.

Lower-bound (rainimum) properties were used for crack stability analyses and average properties were used for leak rate predictions. All the information re, quested is provided in Section 4.0, Recuest 8 Elastic plastic load controlled J-T analyses were performed. Detailed discussion is provided in Section 6.0.

Recuest 9 The Westinghouse computer code for calculating leak rates has been benchmarked and placed under Configuration Control. Results of benchmark calculations have been reviewed and accepted by the NRC in connection with other LBB applications.

Reauest 10 j

The minimum wall-thicknesses at weld undercuts were used in the calcul~ations presented in this report (Section 3.0).

27us *>His c224u to gjy i

Request 11 Typical design transients are listed in Table 7-1 of this report. The crack growth rate data used for the fatigue crack growth analysis is provided in

~

Section 7.0.

Request 12 Limit load analyses do not produce limiting flaw sizes and such results have no impact on the governing stability and margin evaluations presented in this report. Fracture stability analyses which account for material toughness are presented in this report.

Thus, the analyses presented comply with this request.

i 27Me4445e-022tu 10 g

i i

SECTION

1.0 INTRODUCTION

1.1 Purpose This report applies to the Trojan Nuclear Power Plant Reactor Coolant System (RCS) primary loop piping.

It is intended to demonstrate that for the specific parameters of the Trojan plant, RCS primary loop pipe breaks need not be considered in the structural design basis.

The approach taken has been accepted by the Nuclear Regulatory Commission (NRC) (reference 1-1).

1.2 Scope The existing structural design basis for the RCS primary loop requires that dynamic effects of pipe breaks be evaluated.

In addition, protective measures for the dynamic effects associated with RCS primary loop pipe breaks have been incorporated in the Trojan plant design.

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 Trojan plant, 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.

1.3 Objectives In order to validate the elimination of RCS primary loop pipe breaks for the Trojan plant, 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.

am,+o. emu in 11

1 b.

Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the Trojanplant, c.

Demonstrate margin on applied load.

d.

Demonstrate that fatigue crack growth is negligible.

1.4 Backoround Information Westinghouse has performed considerable testing and analysis to demonstrate 1

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 approach. As part of the LLNL research effort, Westinghouse performed extensive evaluations of specific plant loads, material properties, transients, and system geometries to demonstrate that the analysis and testing previously performed by Westinghouse and the research performed by LLNL applied to all Westinghouse plants (references 1-5 and 1-6). The results from the LLNL study were released at a March 28, 1983 ACRS Subcommittee meeting.

These studies which are applicable to all Westinghouse plants east of the Rocky Mountains determined the mean

-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, nu,*o, emu so 1-2

~

To determine if the probability of double ended guillotine breaks is small enough for the plants located on the west coast also, the NRC contracted with the LLNL to conduct a probabilistic assessment of the primary coolant loop piping of the west coast plants. The study performed by LLNL included a plant specific evaluation of the Trojan plant (reference 1-7).

That evaluation determined the mean probabilii.y of a direct LOCA (RCS primary loop pipe break)

~13 to be 2.2 x 10 per reactor year.

It should be noted that the plant

-13 specific probability of a direct LOCA for the Trojan plant, 2.2 x 10 is even lower than the mean probability of all the Westinghouse plants east of the Rocky Mountains.

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 tnat result from 'a number of discrete break locations on the PWR primary systems.

The NRC Staff evaluation (reference 1-1) concludes that an acceptable technical basis has been provided so that asymmetric blowdown loads need not be considered for those plants that can demonstrate the applicability of the modeling and conclusions contained in the Westinghouse response or can provide an equivalent fracture mechanics demonstration of the primary coolant loop integrity.

In a more formal recognition of LBB methodology applicability for PWRs, the NRC appropriately modified 10CFR50, General Design Criterion 4, "Requirements for Protection Against Dynamic Effects for Postulated Pipe Rupture" (51FR 12502).

This report provides a fracture mechanics demonstration of primary loop integrity for the Trojan plant consistent with the NRC position for exemption from consideration of dynamic effects.

i 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 j

independently verified (benchmarked).

vu. o..e:r.u io 1-3

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 Main I. oops," February 1, 1984.

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

NS-EPR-2768, dated May 11, 1903.

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 t'estinghouse (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 NUREG/CR-3660. Vol. 4, UCID-19988, Vol. 4, "Probability cf Pipe Failure in the Reactor Coolant Loops of Westinghouse PWR Plants Volume 4:

Pipe Failure Induced by Crack Growth in West Coast P1:nts," July 1985.

2764s/0445s 021444 10

{.4

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 opera. ting 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 400 reactor years, including five plants each having over 15 years of operation and 15 other plants each with over 10 years of operation.

In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group.

(The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pressurized Water Reactors (PWR's). The results of the study performed by the PCSG were presented in NUREG-0531 (reference 2-1) entitled ' Investigation and Evaluation of Stress Corrosion Cracking in Piping of Light Water Reactor Plents."

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 rigidiy controlled.

Only for brief periods during reactor shutdown when the cooisnt is exposed to the air and during the subsequent startup are conditions m1 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."

2764s/0445s-0224M 10 2-1

During 1979, several instances of cracking in PWR feedwater piping led to the establisnment 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 syst.em loop.

The discussion below further qualifies the PCSG's findings.

For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously:

high tensile stresses, susceptible material, and a corrosive environment. Since some residual stresses and 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 procesring.

The elements of a water er N 7t known tu increase the susceptibility of austenitic stainless steei s..i.ess 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 operc'. ion 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 27H e S41s

• 0234H 10 2-2

normal power operation, oxygen concentration in the RCS is expected to be in 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 Ha'mmer 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 etability, 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 characteris-tics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics, are controlled in the design process.

Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system. Preoperational testing and operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer can occur.

1 l

vu.

,...wu w 23

2.3 Low Cycle and High Cycle Fatigue i

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 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 7.

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 operation.

During operation, an alarm signals the exceedance of the vibration limits.

Field measurements have been j

made on a number of plants during hot functional testing, including plants similar to Trojan.

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.

f 2.4 References 2-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of j

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 R6gulatory Commission, September 1980.

l nu,mn, on,u sa g.4

SECTION 3.0 MPE GFuMETRY AND LOAClNG The general analjtical approach is <*scussed first. A segment of the primary coolant hot le.g pipe shown belew to be hmicing in terms of stresses is sketched in figure 3-1.

This segment is postulated to contain a circumferential thecugh-wall flaw.

The inside diameter and minimum wall thickr.oss er the pipo at the weld undercut are 29.2 and 2.33 inches, respectively.

The pipe is subjected to a normal operating pressure of 2235 psig.

Figure 3-2 identifies the loop weld locations.

The material properties and the loads at these locations resulting from deadweight, thermal expansion, and Safe Shutdown Earthquake are evaluated.

The junction of the hot leg and the reactor vessel outlet nozzle is the worst location for crack stability analysis based on the highest stress due to combined pressure, dead weight, thermal expansion, and SSE (Safe Shutdown Earthquake) loadings. At this location, the axial load (F ) and the bending moment (H ) cre 1918 kips x

b (including axial force due to pressure) and 31,795 in-kips, respectively. This location is referred to as the load critical location.

However, as seen later, significant degradation of end-of-service life fracture toughnesses due to thermal aging occurs in some of the pipe heats and fittings.

The highest stresses and lowest toughness locations for which pipes fittings suffer such degradstien are referred to as toughness critical locations.

The associated heats of material or welds with low toughness are called the toughness critical materials.

The toughness critical locations are 3, 7, 9 and 11 (see figure 3-2).

The stresses due to axial loads and bending moments are calculated by the fellowing equa'. ion:

o=f+f (3.1) 2788s/044Ss 022484 10 3-1

,here, stress o

=

axial load F

=

bending moment M

=

metal cross-sectional area A

=

section modulus Z

=

The bending moments for the desired leading combinations are calculated by the following equation:

M=/M 2,g (3.2) y 2

where, bending moment for required loeding M

=

Y component of bending moment M

=

y Z component of bending moment M

=

Z The axial load and bending moments for leak rate predictions and crack stability analysis are computed by the' methods to be explained in Sections 3.1, 3.2 and 3.3.

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

(3.3)

FDW+FTH + Fp F

=

(M )DW + (N )TH + (N )P (3'4)

M

=

y Y

Y y

(M )DW + (N )TH + (N )P (3.5)

M

=

Z Z

Z Z

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

deadweight OW

=

normal thermal expansion TH

=

load due to internal pressure P

=

4 vu.

.. :.u to 32

The loads based on this method of combination are provided in table 3-1 for all the critical locations (load critical as well as toughness critical.)

3.2 Loads for Crack Stability Analysis The faulted leads for the crack stability analysis are calculated by the following equations:

l (3.6)

TH + F l + lFSSE lFDW + F F

=

p l(M )DW + (M )TH + (N )P I + IIN )SSEI (3'7)

M y

y Y

Y y

l(M )DW + (N )TH + (N )P I + IIN )SSE l (3.8)

M Z

Z Z

2 z

Where the subscript SSE represents SSE loading including seismic anchor motion. For the above load combination, the SSE (inertia) and SSE (anchor motion) are combined by the square-root-sum-of-the-squares method.

The loads based on this method of combination are provided in table 3-2 for all the critical locations (load critical as well as toughness critical).

The seismic response spectra for the Trojan plant were generated for the OBE event by Bechtel Corporation and provided to Westinghouse for seismic analysis of the primary coolant loop. The response spectra for the SSE (DBE) event have been calculated on a conservative basis by multiplying the OBE spectra by the ratio of DBE to OBE peak ground acceleration.

The calculated ratio is 0.25 g's divided by 0.15 g's or 1.67.

Lower magnitude SSE spectra, based on higher structural damping for SSE than for OBE, were not generated or used in the loop seismic analysis.

The seismic analysis of the primary coolant loop consists of two portions:

inertia response of the primary coolant loop (steam generator, reactor coolant pump, hot leg, crossover leg and cold leg piping); and static anchor motion response for the reactor pressure vessel displacement. The building static anchor motions for SSE are negligibly small (only 0.010 inches of relative horizontal displacement).

The damping values used in the SSE analysis are j

0.5% for the inertia response of the primary coolant loop and 2.0% for the inertia response of the reactor pressure vessel.

These correspond to FSAR

)

' values for vital piping and welded structures, respectively.

v u......ons.u a 33 i

3.3 Alternate Load Combination for Crack Stability Analysis Based on Absolute Summation i

In accordance with draft standard review plan 3.6.3 an alternate combination 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 /i to 1.0.

The absolute summation of loads results in the following equations:

I + lF I

(3*9)

F = lF I + lFTHl + IF l + IFSSEINERTIA SSEAM DW p

l + l(M )SSEAM (3.10) l MY ' I(N )DWI + I(N )TH! + IIN )P I + I(N )SSEINERTIA Y

Y Y

Y y

I + IIN )SSEAM (3.11) l Mz = l(M )DWI + I(N )THI + IIN )P I + IIN )SSEINERTIA z

Z 2

2 Z

Based on this method of combination, the loads at the highest stressed location (i.e. location 1 - reactor vessol cutlet nozzle junction) are:

F, = 2235 kips, Mb = 34,716 in-kips.

These loads are used in the fracture mechanics evaluations (section 6.0) to demonstrate margin on loads at location 1.

nu,mn,-owu to 34

TABLE 3-1 NORMAL CONDITION (DEAD WEIGHT AND PRESSURE AND THERMAL) LOADS FOR TRDJAN Weld AxialLoag Bending Moment Location Fx (kips)

Mb (in-kips) bl 1438 24,754 3

1572 16,689 7

1695 4,372 9

1785 12,346 11 1363 5,945 a) Includes internal pressure b) Load critical location.

Remaining locations are toughness critical locations.

1 n u. w i. 4 n i.

3-5

TABLE 3-2 TROJAN PRIMARY LOOP DATA INCLUDING FAULTED LOADING CONDITIONS a

Faulted Loads Bending Direct Stress Yield Ultimate Flow Stress M **"I Stress Stress Axial Load (ksi)

Inside Wall Weld Radius Thickness "y

"u b

], c,,

(Kips)

(in Kips)

F, M

b Locations (in)

(in)

(ksi)

(ksi)

(ksi) x b

"a _ K-Z-

b I

14.6 2.33 20.4 60.0 40.2 1918 31795 26.98 C

3 15.6 2.48 23.8 52.9 38.4 2053 26432 20.59 c

7 15.6 2.48 24.7 52.9 38.8 1798 8901 11.15 c

l[

9 15.6 2.48 24.7 52.9 38.8 1901 15634 14.80 C

11 13.85 2.21 21.1 60.0 40.6 2146 11832 18.47 a Includes internal pressure b

Load critical location for the entire system C

Toughness critical location n

2780s 022488 10

i f

f Crack n

2.33 l

l

---

_____4 F

fT

)

I i+P

+i k

d M

M k29.2 M i

P = 2,235 psig F = 1918 kips j

M = 31,795 in-kips Figure 3-1 Reactor Coolant Pipe noe.-onses to l

i

REACTOR PRESSURE VESSEL

1. 2 Q

COLD HOT LEG LEG 3

2

'-l-REACTOR COOLANT PUMP STEAM GENERATOR CROSSOVER LEG a

O m

G Figure 3-2 Schematic Diagram of Trojan RCL Showing Weld Identifications nu..onna io 3-8

SECTION 4.0 MATERIAL CHARACTERIZATION 4.1. Primary Loop Pipe and Fittings Materials The primary loop piping for the Trojan plant is SA351 CF8A.

The piping is centrifugally cast while the fittings are statically cast.

The elbow fittings are made of SA351 CF8M material. The field welds feature a gas tungsten arc weld (GTAW or TIG) root pass followed by shielded metal arc welding (SMAW) to completion. The shop welds are either SMAW or submerged arc (SAW) with a GTAW root pass. Weld repairs on shop welds >;ould be either SMAW or GTAW.

The welds have TP 308 stainless steel chemistry. No solution annealing was performed.

4.2 Tensile Properties Plant specific material certifications were used to establish the tensile properties for the leak-before-break analyses.

Table 4-1 shows the tensile properties of the SA351 CF8A material at 70'F and 650*F, as well as the properties of SA351 CF8M at 70*F, as taken from the material certifications ~

The properties of SA351 CF8A at 650*F from table 4-1 were used to obtain the representative minimum and average tensile properties of these materials at 650'F.

Properties at operating temperature were not available for the CF8M material in the Trojan plant.

['

Ja,c.e In table 4-4, the representative minimum and average properties of materials in the Trojan plant primary loop are summarized.

In that table the ASME Code minimum properties are also included for a comparison.

vu, esam in 41 4

y, -,,.--

Table 4-5 shows the temperature and pressurs conditions in the primary loop.

At location 1 of figure 3-2, for example, the normal operating temperature is 617'F.

(

Ja.c.e In brief, the following material properties were the ones used in the leak-before-break analyses set forth in this report.

Minimum SA351 CF8A Properties for Flaw Stability Analysis for Location 1 (617'F) a,c,e Averace SA351 CF8A Properties for Leak Rate Calculations for Location 1 (617'F) a,c.e Minimum SA351 CF8M Properties for Flaw Stability Analysis for Location 3 (617'F) a,c.e i

vu, onen so 42

~.. - -

Average SA351 CF8M Properties for Leak Rate Calculations for Location 3.

(617'F) a,c.e Minimum SA351 CF8M Properties for Flaw Stability Analysis for Locations 7 and 9 (552'F)

- a,c.e Average SA351 CF8M Properties for Leak Rate Calculations for Locations 7 and 9 (552*F)

- a,c.e Minimum SA351 CF8A Properties for Flaw Stability Analysis for Location 11 (553*F) a,c.e

_I Average SA351 CF8A Procerties for Leak Rate Calculations for Location 11 (553'F)

- a,c,e n u,-es u n so g.3

i 4.3 Fracture Toughness Properties The pre-service fracture toughness 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 4-5 taken from reference 4-2.

J is observed to be over 5000 Ic 2

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 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 (-

)

i

\\

ja,c.e The effects of the aging process on the end-of-service life fracture toughness are further discussed in appendix B.

End-of-service life toughness for the heats are established using the alter-nate toughness criteria methodology (appendix B). By that methodology a heat of material is said to be as good as [

Ja,c.e if it can be demonstrated that its end-of-service fracture toughneises equal or exceed those of (

Ja c.e Of the forty-seven heats examined in appendix B, five are seen to be not as good as [

Ja,c.e The fracture toughness criteria to be used in the 'racture mechanics evaluation, based on the alternate toughness methodology of appendix B, are given in table 4-5.

These toughness values are the lowest of all heats occurring at the critical locations 1, 3, 7, 9 and 11, shown in figure 3-2.

vu, an,u so 44

Available data on aged stainless steel welds (references 4-3 and 4-4) indicate that'J, values for the worst case welds are of the same order as the aged y

material. However, the slope of the J-R curve is steeper, and higher J-values 2

have been obtained from fracture tests (in excess of 3000 in-lb/in ).

The applied value of the J-integral for a flaw in the weld regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at temperature. Therefore, weld regions are less limiting than a

the cast material.

It is thus conservative to choose the end-of-service life toughness properties of (

Ja,c.e as representative of those of the welds. Also, such pipes and fittings having an end-of-service life calculated room temperature charpy U-notch energy, (KCU), greater than that of [

]a,c,e are also conserva-tively assumed to have the properties of (

Ja,c,e,

in the fracture mechanics analyses that follow, the fracture toughness properties given in table 4-6 will be used as the criteria against which the applied fracture toughness values will be compared, 4

values were conservatively determined by a

In the report all Japplied using base metal strength properties.

3168a C33 ASS 10 f.5

l 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-9558 Rev. 2, "Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack,"

Westinghouse Proprietary Class 2, June 1981.

4-3 WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2, November 1983.

4-4 Slama, G., 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 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),

s m, us m,o 4.g

TABLE 4-1 i'

MECHANICAL PROPERTIES OF THE PRIMARY LOOP MATERIALS OF THE TROJAN PLANT 0.2% OFFSET YIELD ULTIMATE STRENGTH

% REDUCTION i

STRESS (PSI)

(PSI)

ELONGATION IN AREA PRODUCT FORM HEAT NR MATERIAL 70*F 650*F 70'F 650*F.

70*F 650*F 70*F 650*F i

l Straight Pipe C1043 A351 CF8A 42,460 25,000 86,300 60,000 50.0 36.0 65.2 62.1 l

Straight Pipe C1217 A351 CF8A 37,960 28,750 87,400 69,250 51.0 39.0 65.9 61.3 j

Straight Pipe B2700-B A351 CF8A 41,293 21,300 79,303 60,250 59.0 41.0 67.6 60.3 Straight Pipe B2931 A351 CF8A 37,%0 23,700 83.100 63,000 57.0 36.5 66.8 59.1 i

i Straight Pipe C1798-A A351 CF8A 40,460 26,800 82,420 65,500 53.0 41.0 69.5 62.3 Straight Pipe C1798-B A351 CF8A 40,460 26,800 82,420 65,500 53.0 41.0 69.5 62.3 Straight Pipe B2942 A351 CF8A 38,100 23,150 83,200 63,000 50.0 37.5 66.5 61.8 i

Straight Pipe B2711 A351 CF8A 41,459 20,100 81,418 60,250 58.0 44.0 76.1 62.3 Straight Pipe B2700-D A351 CF8A 41,290 21,300 79,300 60,250 59.0 41.0 67.6 60.3 i Straight Pipe C1048 A351 CF8A 45,700 23,100 83,750 61,750 53.5 32.5' 67.0 56.8

)

Straight Pipe C2161-A A351 CF8A 39,960 24,400 81,420 65,500 47.0 42.0 69.3 53.3

)

Straight Pipe C2161-B A351 CF8A 39,960 24,400 81,420 65,500 47.0 42.0 69.3 53.3 Straight Pipe B2919 A351 CF8A 40,600 24,400 80,200 64,000 48.0 36.5 66.7 54.7 i

Straight Pipe B2828 A351 CF8A 41,958 20,900 82,300 61,000 59.0 42.3 75.7

-54.7 1

Straight Pipe B2700-A A351 CF8A 41,293 21,300 79,303 60,250 59.0 41.0 67.6 60.3 Straight Pipe B2849 A351 CF8A 50,949 24,900 83,116 66,500~

45.0 33.0 73.0.

56.0 Straight Pipe C1989-A A351 CF8A 38, % 0 21,300 80,420 62,250 48.0 41.5 70.1 53.3 Straight Pipe C1958-B A351 CF8A 38, % 0 21,300 80,420 62,250 48.0 41.5 70.1 53.3 Straight Pipe B2836 A351 CF8A 39,960 22,400 82,310 64,000 50.0 33.0 71.6 53.8 t

Straight Pipe 82%2 A351 CF8A 47,260 27,800 80,800 68,000 53.0 39.0 69.4 64.7 i

l Straight Pipe C1934-A A351 CF8A 42,957 22,200 83,516 66,250 48.0 39.5 62.8 52.8 Straight Pipe C1934-B A351 CF8A 42,957 22,200 83,516 66,250-48.0 39.5 62.8 52.8 Straight Pipe B2877 A351 CF8A 35,%0 22,800 77,600 60,000 51.0 42.0 75.3 67.0 Straight Pipe C1211 A351 CF8A 46,000 23,100 82,750 65,500 46.5 40.5 65.9 48.4 27-1/2" ID LR 22* ELL 41544-3 A351 CF8N 40,350 85,400 64.0 71.0 2

{

22' ELL 41544-2 A351 CF8M 41,800 86,500 67.0 75.0 22' ELL 41423-3 A351 CF8M 40,050 80,100 55.0 75.0 i

27-1/2" ID LR 22* ELL 41423-2 A351 CF8M 37,100 79,900 65.0 72.0 50' ELL 35222-1 A351 CF8M 48,320 86,820 49.0 70.0 3

50' ELL 34749-1 A351 CF8N 34,900 71,100 61.0 73.0 4

1

}

nee, eneen se i

i

i TABLE 4-1 (cont.)

l MECHANICAL PROPERTIES OF THE PRIMARY LOOP MATERIALS OF THE TROJAN PLANT l

0.2% OFFSET YIELD ULTIMATE STRENGTH

% REDUCTION STRESS (PSI)

(PSI)

ELONGATION IN AREA PRODUCT FORM HEAT NR MATERIAL 70*F 650*F 70*F 650*F 7D*F 650*F 70*F 650*F 50' ELL 61969-1 A351 CF8N 42,000 88,256 58.0 73.0 50' ELL 61969-2 A351 CF8M 42,000 88,250 58.0 73,0 9C* x 31" Elbow 65282-1 A351 CF8M 51,350 95,350 48.0 70.0 31" ID LR 90* ELL Suc. 66035-1 A351 CF8M 41,850 82,200 49.0 69.0 l

31" 10 LR 90* Elbow 64849-1 A351 CF8M 50,250 85,000 54.0 73.0 31" ID 90* Plenum Elbw 63317 A351 CF8M 40,850 81,150 52.0 72.0 31" IDx90* Elbow 65118-1 A351 CF8M 44,900 88,300 48.0 71.0 31" 10x90*Suct Elbw 67348-1 A351 CF8M 48,500 90,700 46.0 67.0 31" ID LR 90* ELL 63190-1 A351 CF8M 42,000 84,000 44.0 71.0 t 31" ID 90* Plenum Elbw 68774-1 A351 CF8M 42,100 81,950 48.0 67.0

" 40* ELL 62648 A351 CF8M 40,850 87,000 65.0 69.0 40* ELL 62554 A351 CF8M 43,050 89,100 56.0 71.0 4

'4 2F90s 022484 le e

-~ ~

TABLE 4-2 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 90 DEG HALF ELBOW 4969 3/6 A351 CF8M 22225.9 53040.6 35.7 67.9 3/6A A351 CF8M 23008.0 53182.8 39.7 63.2 90 DEG HALF ELBOW 4663 3/1 A351 CF8M 23178.6 57733.2 33.3 73.L' 3/1A A351 CF8M 23463.0 57448.8 34.6 61.3 90 DEG HALF ELBOW 4747 3/2 A351 CF8M 23747.4 57448.8 34.2 58.3 3/3 A351 CF8M 23178.6 57022.2 28.9 47.5 90 DEG HALF ELBOW 4898 3/4 A351 CF8M 24174.0 59439.6 39.3 57.3 3/5 A351 CF8M 24245.1 59724.0 34.4 59.3 i' 90 DEG HALF ELBOW 5089 4/1 A351 CF8M 20988.7 55372.7 39.5 64.2 4/2 A351 CF8M 21017.2 56254.3 35.8 58.3 90 DEG HALF ELBOW 5839 4/5 A351 CF8M 21244.7 53026.4 31.1 62.3 4/6 A351 CF8M 21031.4 54107.1 40.0 59.3 90 DEG HALF ELBOW 5247 4/3 A351 CF8M 21400.1 55045.6 33.2 48.7 4/4 A351 CF8M 21371.6 54931.9 33.9 65.1 40 DEG ELBOW 4541 2/1 A351 CF8M 23463.0 57022.2 22.2 60.3 2/1A A351 CF8N 23889.6 57448.8 26.6 57.3 30 DEG ELBOW 4543 5/2 A351 CF8M 22041.0 57448.8 30.0 65.1 5/3 A351 CF8M 22609.8 57022.2 31.1 55.2 40 DEG ELBOW 5655 2/2 A351 CF8M 21244.7 52756.2 37.8 62.3 2/2A A351 CF8M 21258.9 52528.7 37.3 48.7 40 DEG ELBOW 5854 2/3 A351 CF8M 20960.3 54562.1 39.5 67.9 2/3A A351 CF8M 21287.3 54107.1 38.6 67.9 50 DEG ELBOW 4594 1/2 A351 CF8M 22894.2 57591.0 34.9 56.2 1/3A A351 CF8M 22680.9 57164.4 38.4 64.2 50 DEG ELBOW 4665 1/3 A351 CF8N 22894.2 56453.4 40.9 64.2 1/3A A351 CF8M 22680.9 56880.0 43.9 69.6 50 DEG ELBOW 4205 1/1 A351 CF8M 24245.1 57448.8 24.4 56.3 1

1/1A A351 CF8M 24031.8 57591.0 30.3 55.2 33 DEG ELBOW 64422 5/1 A351 CF8M 25169.4 57875.4 25.0 38.0

)

5/1A A351 CF8N 25027.2-58728.6 24.4 60.3

)

n u... n.w i.

TABIF 4-3 l

MECHANICAL PROPERTIES OF SA351 CF8N MATERIAL AT ROOM IEMPERATURE (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 90 DEG HALF ELBOW 1969 3/6 A351 CF8M 33843.6 71811.0 53.3 70.5 3/6A A351 CF8N 34270.2 70673.4 43.9 67.8 90 DEG HALF ELBOW 4663 3/1 A351 CF8N 35123.4 71953.2 41.1 63.2 3/1A A351 CF8M 34833.0 73659.6 45.1 68.5 90 DEG HALF ELB0W 4747 3/2 A351 CF8M 33701.4 71811.0 47.8 69.5 3/3 A351 CF8N 34839.0 71100.0 37.8 63.2 90 DEG HALF ELBOW 4898 3/4 A351 CF8M 35123.4 75792.6 36.0 52.2 3/5 A351 CF8M 36118.8 77072.4 51.8 67.0 i 90 DEG HALF ELBOW 5089 4/1 A351 CF8N 32137.2 71384.4 44.5 50.2 U

4/2 A351 CF8M 31284.0 70673.4 48.1 N/A 90 DEG HALF ELBOW 5839 4/5 A351 CF8N 33132.6 71100.0 47.9 66.5 4/6 A351 CF8N 32279.4 71526.6 40.5 68.4 90 DEG HALF ELBOW 5247 4/3 A351 CF8N 32420.0 70385.7 41.5 72.0 4/4 A351 CF8M 32846.6 71096.6 48.6 71.2 40 DEG ELBOW 4541 2/1 A351 CF8M 34981.2 72095.4 37.7 65.0 2/1A A351 CF8N 34981.2 71242.2 41.9 66.0 j

30 DEG ELBOW 4543 5/2 A351 CF8M 33132.6 72522.0 50.7 71.2 5/3 A351 CF8M 33417.0 72237.6 50.3 67.8 40 DEG ELBOW 5655 2/2 A351 CF8M 35123.4 72237.6 50.7 70.5 2/2A A351 CF8M 34128.0 71100.0 47.3 76.2 40 DEG ELBOW 5854 2/3 A351 CF8N 30999.6 72237.6 53.0 71.2 2/3A A351 CF8M 31852.8 71100.0 50.4 76.2 50 DEG ELBOW 4594 1/2 A351 CF8N 34128.0 72522.0 49.6 71.2 1/2A A351 CF8M 33843.6 72948.6 46.0 77.7 50 DEG ELBOW 4665 1/3 A351 CF8M 33701.4 70246.8 50.6 67.8 1/3A A351 CF8M 32279.4 70389.0 51.6 71.2 50 DEG ELBOW 4205 1/1 A351 CF8M 31568.4 76788.0 37.5 61.2 l

1/1A A351 CF8N 38962.8 75508.2 47.0 66.0 j

30 DEG ELBOW 64422 5/1 A351 CF8M 40669.2 81907.2 43.3

'64.0 j

5/1A A351 CF8M 36403.2-71242.2 42.5 71.2 i

me on i.

l i

{

TABLE 4-4 TROJAN MATERIAL PROPERTIES AT 650*F Pipe Fittings Material SA351 CF8A SA351 CF8M.

y(ksi) u(ksi) y(ksi) u (ksi) o o

ASME Code minimum 21.0 65.2 18.5 67.0 a,c.e es O

e v u. w i. m i.u io 4 11 e

TABLE 4-5 PRESSURE AND TEMPERATURE CONDITIONS IN THE PRIMARY LOOP FOR TROJAN Hot Leg Temperature:

F.17 'F pressure:

223B psig Crossover Leg Temperature:

552*F Pressure:

2200 psig Cold Leg Temperature:

553*F Pressure:

2290 psig vu.co.<n.u in 4-12

TABLE 4-6 FRACTURE TOUGHNESS CRITERIA USED IN THE LEAK-BEFORE-BREAK EVALUATION Location g

T J,,x a

Ic mat 2

2 (in-lb/in )

(in-lb/in )

All locations except

~

noted below 3

b 7

9 11 L

a.

The locations are shown in figure 3-2.

b.

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

vu.wwen.u io 4 13

=

4 a,c,e

]

,j 1

a Figure 4-1.

True Stress Strain Curve for SA351 CF8A Stainless Steel at 617'F

=.4isim ie 4 14 i

a,c.e l

e l

I 1

e l

\\

i a

Figure 4-2.

True Stress Strain Curve for SA351 CF8A Stainless Steel at 553*F svu.m stu,o 41s

a,c.e I

l I

l I

i l

l l

i i

l I

l l

I

~

Figure 4-3.

True Stress Strain Curve for SA351 CF8N Stainless Steel at 617'F vu.m n u "

4-16 1

a,c.e t

i l

\\

Figure 4-4.

True Stress Strain Curve for SA351 CF8M Stainless Steel at 552*F nu,n nsu,o 4 1y

1

~

~

a,c.e Figure 4-5.

J vs Aa for SA351 CF8H Cast Stainless Steel at 600'F nu.am u io 4-18

..c..

l J vs. Aa at Different Temperatures for Aged Waterial Figure 4-6.

(

3a,c.e (7500 hours0.0868 days <br />2.083 hours <br />0.0124 weeks <br />0.00285 months <br /> at 400*C) nu.mmu io 4-19 l

SECTION 5.0 LEAK RATE PREDICTIONS 5.1. Introduction l

t 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.

5.2 General Considerations I

The flow of hot pressurized water through an opening to a lower back pressure l

causes flashing which can result in choking.

For long channels where the ratio of the channel length, L, to hydraulic diameter, D, (L/D ) is g

g greater than (

Ja.c.e, both (

(

L 3a,c.e,

5.3 Calculation Method k

The basic method used in the leak rate calculations is the method developed by t

[

ja,c.e,

The flow rate through a crack was calculated in the following manner.

Figure l

5-1 from reference 5-1 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pe was found for a 27HsS45sto224u to 5-1

given mass flow, the (

3a,c.e cac found from figure 5-2 taken from reference 5-1.

For all cases considered, since (

Ja,c.e Therefore, this method will yield the two phase pressure drop due to momentum effects as illustrated in figure 5-3.

Now using the assumed flow rate, G, the frictional pressure drop can be calculated using aPf=(

) **

(5-1) where the friction factor f is determined using the (

la,c.e The crack relative roughness, e, was obtained from fatigue crack data on stainless steel samples.

The relative roughness value used in these calculations was (-

-Ja c.e The frictional pressure drop using Equation 5-1 is then calculated for the assumed flow and added to the [

Ja,c.e to obtain the total pressure drop from the primary system to the atmosphere.

That is, for the primary loop Absolute Pressure - 14.7 = (

)"'C(5-2) for a given assumed flow G.

If the right-hand side of Equation 5-2 does not agree with the pressure difference between the primary loop and the atmosphere, then the procedure is repeated until Equation 5-2 is satisfied to within an acceptable tolerance and this results in the flow value through the crack.

This calculational procedure has been recommended by (

Ja c.e for this type of (

Ja,c.e calculation.

a m, w s.na m ii.

5-2

5.4 Leak Rate Calculations Leak rate calculations were made as a function of crack length for all the five critical locations previously identified.

The normal operating loads of table 3-1 were applied in these calculations.

The crack opening area was estimated using the method of reference 5-3 and the leak rate was calculated using the two phase flow formulation described above.

The flaw sizes to yield a leak rate of 10 gpm at each of the locations were established and are summarized in table 5-1.

5.5 References 5-1 [:

,j c.e a

5-2 [

ja,c.e,

5-3 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983.

n u.m us.$ u.u to 5-3

TABLE 5-1 LEAK RATE RESULTS Location 10 gpm Flaw Size (in.)

a,c.e 3

3 7

9 11 vs.

o.4 mis io 5-4

- a c.e

=

"a=l I

t 8

N<a Analytical Predictions of Critical Flow Rates of Steam-Figure 5-1 Water Mixtures l

l 5-5 l

a,c.e h*8 e

Es w

Iw E

,(W tau Figure 5 [

3a c.e Pressure Ratio as a Function of L/D 5-6

1 1

_ a,c.e

[

=

h

/

Figure 5-3.

Idealized Pressure Drop Profile Through a Postulated Crack nu minuio 5-7 1

SECTION 6.0 FRACTURE MECHANICS EVALUATION 6.1. 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.

Depending on the material properties and geometry of the pipe, flaw size, shape and loading, the local failure mechanisms may or may not govern the ultimate failure.

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 Jic 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 J;c 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 E

T,pp = g, q

f where:

T

= applied tearing modulus app E = modulus of elasticity of = (

Ja.c e (flow stress) a = crack length

(

Ja,c.e Stability is said to exist when ductile tearing occurs i.e., T,pp is less than Tmat, the experimentally determined tearing modulus.

In summary, the local crack stability will be established by the two-step criteria:

vu, wn,mmu io 6-1

J<J ie or if J 3 J T,pp < Tmat 3e and J < Jmax 6.2 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 lead 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 6-1) 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:

[-

ja.c.e nu, w.m"" "

6-2

[

ja,c.e 1

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 6-1).

6.3 Results of Crack Stability Evaluation Stability analyses were performed at the load critical and toughness critical locations defined in section 3.

The elastic plastic fracture mechanics (EPFM) l J-integral analysos for through-wall circumferential cracks in a cylinder were performed using the procedure in the EPRI fracture mechanics handbook (reference 6-2).

The lower-bound material properties of section 4.0 were used.

The results of the EPFM analyses are given in tables 6-1 and 6-2.

The critical toughness criteria established in section 4.3 were used.

The normal plus SSE loads given in table 3-2 were used for these calculations.

Two margin conditions were evaluated.

First the leakage flaw (i.e. the flaw yielding a leakage of 10 gpm) was doubled and EPFM analyses were performed for normal plus SSE loadings. The applied tearing modulus was calculated from the EPFH results (elastic plastic load controlled J-T analysis) using dJ T,pp = g, E7 f

I 4

9 v u, w w a w u so 6-3

where E is the modulus of elasticity and of is the flow stress taken as the average of the yield and ultimate strength.

In table 6-1 the calculated values are seen to be well below the critical toughness values. J is Ic exceeded only at locations 1 and 11.

Similar analyses were made using the leakage flaws with the normal plus SSE i

loads (from table 3-2) increased by a factor of 1.4.

These results are provided in table 6-2.

As noted in the table a margin of 1.4 on riormal plus SSE loads for leakage flaws is demonstrated at all locations except location.

where only a margin of 1.2 is demonstrated. To additionally evaluate lor cion 1, the loads were combined using the absolute load combination method (:ee section 3.3) as outlined in SRP 3.6.3.

The absolute load combination rasulted in F = 2235 kips and Mb = 34,716 in-kips. Based on these loads, the x

J was calculated to be [

la,c.e and T was applied applied calculated to be (

Ja,c.e. Thus the results of table 6-2 demonstrate I

margins on applied loads.

SRP 3.6.3. requires the same load combination procedure be used when addressing margin on flaw size and margin on load.

Thus the leads obtained by the absolute load combination methcd were used to evaluate margin on flaw size, i

The results are also given in table 6-1 where it is seen that a margin of 1.5 on the leakage size flaw is obtained.

For a factor of two on flaw size, the

[

load based on the absolute load combination method is met at the 94 percent level as indicated.

From table 6-1 it is clear that the limiting location is location 1 (the critical flaw size is the smallest at this location). Global stability j

analysis was performed at this location as described in section 6.2.

Figure 6-2 shows a plot of the plastic limit moment as a function of through-wall circumferential flaw length in the hot leg (location 1).

The maximum applied bending moment can be plotted on thir figure and used to determine a critical flaw length, which is shown to be [

Ja c.e inches.

In summary it is seen that large margins exist for both flaw and load at all locations except location 1.

Tne margin on flaw size based on normal plus SSE is well demonstrated, however the corresponding margin on load is 1.2, not me,*o,422a*

• 6-4 l

wEsTINSHOust PROPRIETARY Ct.AS: 2 1.4.

Conversely, using the absolute load combination method, the margin on load is well demonstrated with the margin on flaw size being 1.5, not 2.

The adequacy of the margins demonstrated at location 1 is discussed below.

First the reason why the sought after margins at location 1 were not demonstrated is discussed. This is perhaps best simply demonstrated by referring to figure 4-1.

Taking the applied stress as the effective stress, the normal plus SSE load (stress of 26.98 ksi) produces an unflawed outside surface strain of about 1.25 percent.

Also,1.4 times 26.98 ksi is 37.78 ksi which yields a strain of about 6%. That is, at the load level in question, half an order-of-magnitude increase in the applied strain results from only a 40 percent increase in load.

For a factor of 1.2 on normal plus SSE load the strain is 3.7 percent. For the loads from the absolute load combination method the stress is 30.1 ksi, a 11 percent increase over the normal plus SSE load. The corresponding strain is 2.6%.

At 94% of this load the strain is 2.0 percent (see footnotes of table 6-1). Overall the applied fracture toughness results discussed above are not particularly surprising.

Next the nature of the loads themselves is exsmined. A very large percentage of the bending moment is due to thermal expansion, a displacement controlled load. Likewise the seismic moment is displacement controlled considering damping. Consequently, displacement or strain is a better overall stability I

criterion for displacement controlled stresses.

Of course in the elastic i

region, load control and displacement control stress criteria are equivalent.

It appears plausible then to examine margins in terms of strain at Location 1.

The established margin on load wherein the margin is 1.2 (see table 6-2),

gives a margin of almost three (~3.7/1.25) in strain. For the 6.8 inch long flaw of table 6-1 subjected to 94 percent of the load obtained by the absolute load combination method, the strain is 2.0 percent or 60 percent greater than the strain for normal plus SSE load.

The stability for the flaw size with a i

60 percent increase in strain over that for normal plus SSE load contrasts with the 11% differences in stress for the two loads under discussion.

An alternative to meeting precisely the margins has been discussed above.

1 s

vu,wnuu io 6-5

.. J

For the Trojan plant, it is believed that a sufficient basis has been established for using engineering judgment and flexibility in assessing these results.

This is consistent with NUREG-1061, Volume 3, page 5-21.

Supplementing the above discussion concerning location 1 it should be kept in mind that conservative or benchmarked procedures have been applied in every step of the evaluation.

The loads are conservatively calculated and combined. The fracture toughness criteria at location 1 are taken equal to that of (

]a,c.e while in fact the fracture toughnesses would be considerably higher based on KCU values. Thus margins are in fact greater than those established in the above discussion.

6.4 References 6-1. Kanninen, M. F., et. al., "Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks," EPRI NP-192, September 1976.

6-2, 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, me,wwouwe 6-6

TABLE 6-1 RESULTS OF STABILITY ANALYSES - MARGIN ON FLAW SIZE J

T U

Crack Length J,pp T,pp yc mat max 2

2 2

Location (in-lb/in )

(in-lb/in)

(in)

(in-lb/in )

Normal Plus SSE Loading 1

a,c.e 3

l 7

9 11 Lead by Absolute load Combination Method a,c.e 1

1 a

NA - not applicable b

This crack length is 1.5 times that of the leakage flaw, c

For this crack depth, the load used is 94% of the load obtained by the absolute load combination method and 4% greater than normal plus SSE load, noting that the stresses for the two loads in question differ by only 11 percent.

vusewenwo g.y

TABLE 6-2 RESULTS OF STABILITY ANALYSES - MARGIN ON LOADS J

T O

Crack Length J,pp T,pp ge mat max 2

2 2

Location (in-Ib/in )

(in-lb/in )

(in)

(in-lb/in) i Factor of 1.4 on Load Obtained by SRSS Combination a,c.e

~.

ya 3

7 P

9 11 Load by Absolute Load Combination Method a,c.e l

I I

i These results are for a lead of 1.2 times the normal plus SSE load not 1.4 times the normal plus SSE load.

b h

NA - not applicable I

(

i vu.w.ar.= ie 6-8

,w

a,c.e

,,,/,'!!Lilli j,

2.

=== Newttti A a el

\\N ll V

a Figure 6-1.

[

3 'C Stress Distribution 8

aru. main in 6-9

a,c.e 00 = 33.86 in.

t = 2.33 in.

F = 1918 kips

= 20.353 ksi o

y

= 60.00 ksi u

op = 40.18 ksi 0

T = 617 F 4

4 s

i I

Figure 6-2.

"Critical' Flaw Size Prediction - Het Leg at Critical Location 1

=.meui 6-10 i

I SECTION 7.0 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 [

.)a,c.e region of a typical system (see Location

[

Ja,c.e of figure 3-2). This region was selected because creek growth calculated here will be typical of that in the entire primary loop. Crack growths calculated at other locations can be expectea to show less than 10%

variation.

A[

Ja,c.e of a plant typical in geometry and operational characteristics to any Westinghouse PWR System.

[

, J a, c. e. All normal,

.. a r.

upset, and test conditions were' considered.

A suamary of the applied transients is provided in table 7-1.

Circumferentially oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in figure 7-1.

Specifically, these were:

Cross Section A:

[

Ja,c.e Cross Section B:

[

]a,c.e Cross Section C:

[

Ja,c,e Fatigue crack growth rate laws were used [

Ja c.e The law for stainless steel i

was derived from reference 7-1, with a very conservative correction for the R ratio, which is the ratio of minimum to maximum stress during a transient.

vu,4us, o:2.u io 71

-=. - - - _. -.

For stainless steel, the fatigue crack growth formula is:

h=(5.4x10-12) g ff inches / cycle 4.48 where K,ff = K,,x (1-R)0.5 R=Kmin/Emax

[

a,c.e

-)

l wher,e: [

Ja,c.e The calculated fatigue crack growth for semi-elliptic surface flaws of cir:umferential orientation and various depths is sumarized in table 7-2, and shows that the crack growth is very small, [.

Ja,c.e 27Ms 4445s/C224H 10 7-2

7.1 References l

7-1 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.

7-2 [

ja,c,c 7-3 (. '

)n,c.e nu,m<wownso 7-3 v-

--e-

+-,

g mer-3

-m>as

,--w 9

--w-

'w-y w

y

+rw-

-wi 9*=w

TABLE 7-1

SUMMARY

OF REACTOR VESSEL TRANSIENTS NUMBER TYPICAL TRANSIENT IDENTIFICATION NUMBER OF CYCLES Normal Conditions 1

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

of full power / min) 3 Step load increase and decrease 2000 4

Large step lead decrease, with steam dump 200 6

5 Steady state fluctuations 10 Upset Conditions 6

Loss of load, without immediate turbine 80 or reactor trip 7

Loss of power (blackout with natural circulation 40 in the Reactor Coolant System) 8 Loss of Flow (partial loss of flow, one pump only) 80 9

Reactor trip from full power 400 Test Conditions 10 Turbine roll test 10 11 Primary Side Hydrostatic test conditions 50 12 Cold Hydrostatic test 10 am. o.m m ue 7_4

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

TABLE 7-2 TYPICAL FATIGUE CRACK GROWTH AT

[

.)a,c.e (40 YEARS)

FINAL FLAW (in)

[

INITIALFLAW(IN)

Ja,c.e

[

3a,c.e

[

3a,c.e 0.292 0.31097 0.30107 0.30698 0.300 0.31949 0.30953 0.31626 0.375 0.39940 0.38948 0.40763 0.425 0.45271 0.4435 0.47421 27H t/0845s/02244410 7-5

i l

i

.I I

I 4, cit i

I 1

Figure 7-1.

Typical Cross-Section of (

Ja,c.e n u, m v u so 7-6

- e,c, e

~

i!T ao>us M

Wzua oa2 E

Es 4

d<s Ioeowu<zo i

i Figure 7-2. Reference Fatigue Crack Growth Curves for [

3a.c.e 7-7 1

.u.--.

BM22 i

e.c. e

-l

• l l

i l

l i

1 P

Figure 7-3.

Reference Fatigue Crack Growth Law for ('

Ja c.e in a Water Environment at 600'F 7-8

WESTINGHOUSE PR!PRIETARY CLASS 2 SECTION

8.0 ASSESSMENT

OF MARGINS The results of the leak rates of section 5.4 and the corresponding fracture toughness evaluations of section 6.3 are used in performing the assessment of margins.

At the load critical location (location 1), J,pp and T,pp for a (

ja,c,e long through-wall circumferential flaw are found to be [

]a,c.e and I ',c'.e

[

Ja c e, respectively. The T,pp is shown to be less than T fI mt

.]a and J,pp is shown to be less than J,,x of (

.]a,c.e in-lb/in. A[

inch flaw yields a leak rate of 10 gpm.

Thus at the load critical location, there is a margin of at least two between the flaw size yielding a leak rate of 10 gpm and the critical flaw size [

]a,c e In addition, the leakage size flaw of [

]a,c.e is shown to be stable when subjected to maximum loads obtained by the absolute sum load combination method as sug-gested by SRP 3.6.3.

Stability for the leakage size flaw was not established using a factor of 1.4 on the normal plus SSE load. A factor of 1.2 was established however.

Stability of a flaw twice the size of the leakage size flaw was demonstrated at 94 percent of the load obtained by the absolute sum load combination method. Stability was also established for this load for a flaw 1.5 times the leakage flaw size.

Because most of the normal plus SSE bending moment is displacement controlled an evaluation based on strain was made.

Using strain as a criterion, a f actor of three on strain exists for the factor of 1.2 on load.

Similarly a factor of 1.60 on strain exists for a flaw twice the leakage size flaw which was stable at 94 percent of the load obtained by the absolute load combination method. Overall it is judged that adequate margins on flaw size and load have been demonstrated at location 1 consistent with the philosophy of NUREG 1061, Vol. 3.

At the remaining four toughness critical locations a flaw up to at least

[

'la,c.e inches long is stable. A margin of at least 2 is demonstrated between the flaw size yielding a leakage of 10 gpm and the critical flaw size. The normal plus SSE loads are increased by a factor of /2 vu,mm.-men to 8-1

WEsTINGH1usE PRoPRitTARY class 2 and crack stability is demonstrated for the leakage size flaws (i.e. flaws yielding a leakage of 10 gpm).

Thus the margin on loads is demonstrated.

In section 6.3 the "maximum" flaw size at the load critical location (limiting location) using the limit load method is found to be at least [

Ja,c,e inches. Thus, based on the above paragraphs, the critical flaw sizes at these locations would exceed (

]a,c.e respectively.

In sumary, relative to 1.

Flaw Size a.

A margin of at least 2 exists between the critical flaw and the flaw yielding a leak rate of 10 gpm.

b.

If limit load is used as the basis for critical flaw size, the margin for global stability well exceeds that based on local stability fracture mechanics evaluation.

)

2.

Leak Rate A margin of 10 exists between the calculated leak rate from the "leakage-size flaws" and a leak detection capability of 1 gpm.

The capability of each of the pressure boundary leak detection syctems are given in table 5.2-9 of the Trojan FSAR.

3.

Loads The J,pp values for the Trojan plant are enveloped by the a.

toughness allowables established for thermally aged material, b.

At load critical location the leakage size flaw was shown to be stable when subjected to normal plus SSE loads obtained by absolute sunnation of individual components.

This method of combination results in maximum load at the location of interest.

vu.m.-mm e 8-2

wsSTINGHouSE PROPRIETARY class 2 l

At toughness critical locations the leakage-size flaws were shown to c.

be stable when subjected to '.2 times the normal plus SSE loads.

4.

General (LoadCriticalLocation) a)

Margin criteria on both flaw size and load were not met at the load critical location using a specific procedure for calculating normal plus SSE loads.

b)

For the normal plus SSE load obtained by the absolute sum combination method, a factor of 1.5 on the leakage size flaw was established as compared to a target of 2.

For a factor of 2 on flaw size, stability was established at 94 percent of the load, c)

A factor of 1.2 on load was established as compared to a target of 1.4.

d)

Using strain as a criterion, a factor of 1.6 was established for the 94 percent load of item 4b). For the 1.2 factor of item 4c), a factor of 3 on strain was established.

e)

The margins established at the load critical locations are judged adequate consistent with the philosophy of NUREG 1061, Volume 3.

nu.w.-mme.ie 8-3

WESTIN2 house PRIPRIETARY CLASS 2 SECTION 9.0 CONCLUSIONS This report justifies the elimination of RCS primary loop pipe breaks for the Trojan plant as follows:

Stress corrosion cracking is precluded by use of fracture resistant a.

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.

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

primary piping are negligible, d.

Adequate margin exists between the leak rate of small stable flaws and the capability of the Trojan plant's reactor coolant system pressure boundary Leakage Detection System.

Amp 4 margin exists between the small stable flaw sizes of item d and e.

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.

i For each critical location a flaw is 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 Trojan plant.

27Ns/CMSe-0303H 10 g-1

v vw.,-

m,

1 l

l APPENDIX A LIMIT M0HENT l

I 8,C,9

)

me..ar " "

A-1

~

- a,c.e FIGURE A-1 PIPE WITH A THROUGH-WALL CRACK IN BENDING A-2

1

~

APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE TROJAN CAST PRIMARY LOOP COMPONENTS B.1 INTRODUCTION Not all of the individual cast piping components of the Trojan primary loop piping satisfy the original (

Ja,c,e criteria (reference 4-3).

In this appendix, the alternate toughness criteria for thermally aged cast stainless steel developed in reference 4-5 will be used to categorize the various individual cast piping components thus establishing criteria based upon which the mechanistic pipe break evaluation may be performed.

Reference 4-5 has been reviewed by the NRC wherein the NRC concluded that reference 4-5 may be utilized for establishing the fracture criteria for thermally aged cast stainless piping applicable for the leak-before-break analyses (reference B-1).

First the chemistry and calculated room temperature Charpy U-notch energy (KCU), values are given followed by an identification of each of the heats of material which did not meet the [

Ja,c.e criteria and their location.

B.2 CHEMISTRY AND KCU TOUGHNESS The correlation of reference 4-4 which is based on the chemistry of the cast stainless steel piping was used to calculate the associated KCU value. The chemistry and end-of-service life KCU toughness values are given in table B-1.

Of the forty-seven heats of cast stainless steel, five fail to meet the current [

Ja,c.e criteria.

These heats occur in the fittings of the hot, and crossover legs and in the cold leg pipe.

1 a m,4 usinar m io B-1

B.3 THE AS-BUILT TROJAN LOOPS Trojan is a four-loop Westinghouse type pressurized water reactor plant. A typical four-loop primary system is sketched in figure B-1.

The loops are identified as Loops 1, 2, 3 and 4 in the Trojan plant.

Sketches for associating piping components which have toughnesses less than that of I Ja,c e are identified (see figures B-2 to B-4).

B.4 ALTERNATE TOUGHNESS CRITERIA FOR THE TROJAN PRIMARY LOOP MATERIAL ON A COMPONENT BY COMPONENT BASIS The alternate toughness criteria for the Trojan cast primary loop material may be obtained by apply',ng the methodology of reference 4-5 to table B-1.

First, it is observed that 42 of the 47 heats fall into category 1, i.e.,

they are at least as tough as (

Ja,c.e The remaining five heats fall into category 2.

Typical toughness calculations using the methodology of reference 4-5 are given below for a category 2 heat.

The 90 degree elbow on crossover leg (

.)a,c.e has the lowest calculated end-of-service life KCU at room temperature of (

Ja c.e 2

daJ/cm which falls below that of (

Ja,c.e The 6-ferrite content is [

Ja,c,e By reference 4-5, the (

1

).a,c.e Since the end-of-service life KCU exceeds the fully aged KCU, the heat falls into category 2.

Thus:

Jie = {

ja,c.e T

• I l** '*

mat me,muwowse,o g_g

and

• I max ja,c.e These toughness results and the results for the remaining heats which do not meet the [

Ja,c.e criteria are summarized in table B-2.

B.5 REFERENCES B-1 Letter: Dominic C., Dilanni, NRC to D. M. Muslof, Northern States Power Company, dated December 22, 1986, Docket Nos. 50-282 and 50-306.

l 1

m e.4.n, mas u in B-3

TABLE B-1 CHEMISTRY AND CALCULATED KCU VALUES FOR EACH PRIMARY LOOP PIPING OF THE TROJAN huCLEAR PLANT j

8,C,0 1

i

~

nu.an* in B-4

TABLE B-1 (cont.)

CHEMISTRY AND CALCULATED KCU VALUES FOR EACH PRIMARY LOI OF THE TROJAN NUCLEAR PLANT a.c.c f

e G

6

.p

=. anne ia B-5 m.3,-,--

7,-,.,.7-gy

,,,.--y 7

,y__

,.-,,em,ww,-,-mpmei,y.c rym,-wp--my ew c waw

TABLE B-1 (cont.)

CHEMISTRY AND CALCULATED KCU VALUES FOR EACH PRIMARY LOOP PIPING OF THE TROJAN NUCLEAR PLANT

~

a,c.e i

nm.an n is B-6

TABLE B-1 (cont.)

i CHEMISTRY AND CALCULATED KCU VALUES FOR EACH PRIMARY LOOP PIPING OF THE TROJAN NUCLEAR PLANT 8,C,e l

e 6

4

=. '" * "

B-7

TABLE B-2 FRACTURE TOUGHNESS CR1TERIA FOR THE CAST PRIMARY PIPING

~

COMPONENTS OF THE TROJAN NUCLEAR PLANT

- a,c.e' i

l l

All heats except noted here meet' the (

Ja,c.e criteria and therefore will have toughness at least equal to (-

ja,c.e I

vu.co,4n.u io B-8

l PUMP CROSSOVER LEG COLD LEG N

/

i HOT LEG.

/

\\

REACTOR i

STEAM VESSEL GENERATOR I

I i

Figure B-1 Typical layout of the Primary Loops for a Westinghouse Four-Loop Plant Without Isolation Valves 11Ms/Qtt1H ty B-9

1 a,c.e i

i Figure B-2 Identification of Heats with Location for Cold Leg i

i cu,miniu to B-10

l

- a,c.e Figure B-3 Identification of Heats with Location for Hot Leg n.a nsu,o g.$3

a,c.e l

l l

1 I

P i

Figure B-4 Identification of Heats with Location for Crossover Leg B-12 vu.mi.= in