ML20215N167

Rev 1 to, Leak-Before-Break Evaluation of Reactor Coolant Loop

ML20215N167
Person / Time
Site:	Trojan File:Portland General Electric icon.png
Issue date:	10/31/1986
From:	Emerson R ABB IMPELL CORP. (FORMERLY IMPELL CORP.)
To:
Shared Package
ML20215N163	List: ML20215N162 ML20215N167
References
01-0300-1395, 01-0300-1395-R01, 1-300-1395, 1-300-1395-R1, TAC-06596, TAC-6596, NUDOCS 8611040390
Download: ML20215N167 (51)

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I Leak-Before-Break Evaluation of the Reactor Coolant Loop Final Report I

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l Submitted to:

l PORTLAND GENERAL ELECTRIC COMPANY Prepared by:

Impell Corporation 350 Lennon Lane Walnut Creek, California 94598 l

Impell Report No. 01-0300-1395 l

Revision 1 October, 1986 I

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Page 2 of 49

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IMPELL CORPORATION REPORT APPROYAL COVER SHEET

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Client:

Portland General Electric Company

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Project: Trojan Nuclear Plant Job Number: 0300-023-1356 Report

Title:

Leak-Before-Break' Evaluation of the Reactor Coolant Loop

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Report Number: 01-0300-1395 Revision 0

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The wor'r described in this Report wa's perforined in accordance with the Impe11 Quality Assurance Program. The signatures below verify the accuracy of this Report and its compliance with applicable quality assurance requirements.

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Date:

Prepared By:

Robert Emerson, Supervising Engineer hhk Reviewed By:

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Walter R. Bak, M r,

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Applied Mechanic tion Approved By:

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Date:/fr d9 I

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' Robert L. Grubb, Manager, Advanced Engineering Division

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REVISION RECORD Rev.

No.

Prepared Reviewed Approved Date Revision

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incorporated.

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Report No. 01-0300-1395 Revision 1 Page 3 of 49 I

TABLE OF CONTENTS I

Page 1.0 Introduction 5

2.0 Methodology and Numerical Results 7

2.1 Criteria 7

2.2 Description of Piping System 7

2.3 Crack Detectability Under Normal 8

Operating Conditions 2.3.1 Crack Leakage Analysis I

2.3.2 Leak Detection Systems 2.3.3 Detectable Crack Size 2.4 Fracture Mechanics Analysis 11 2.4.1 Loading Conditions and Critical Locations 2.4.2 Local Stability of Postulated k

I Crack 2.4.3 Limits of Applicability 2.5 Material Properties 16 2.5.1 J-Integral and Tearing Modulus Deta for Base Metal 2.5.2 J-Resistance Curves for Welds 2.5.3 Lower Bound J-Resistance Curves for Analysis 2.6 Net Section Evaluation 18 2.7 Summary of Results 18 3.0 Conclusions 19 Tables 20 Figures 28 Peferences 36 List of Symools 38 i

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Repor2 No. 01-0300-1395 Revision 1 Page 4 of 49 I

TABLE OF CONTENTS (Continued)

Page Appendix A Description of Computer Programs '

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Appendix B Sensitivity of Leakage Detection 45 Systems Appendix C Fatigue Crack Growth Analysis 47 I

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Portland Gen:ral Electric Report No. 01-0300-1395 Revision 1 Page 5 of 49

1.0 INTRODUCTION

A leak-before-break (LBB) evaluation has been perfomed to define the potential for pipe rupture on the reactor coolant loop (RCL) piping of the Trojan Nuclear Plant.

The approach is based on demonstrat-ing that rather than suddenly breaking in a double-ended guillotine fashion, as has been previously postulated, the RCL piping would develop stable, detectable cracks prior to rupture.

The approach follows a defense-in-depth philosophy by showing substantial margin in each of the analytical steps.

The approach uses a four-step evaluation based on g

deterministic fracture mechanics to show that a double-ended rupture of the RCL piping at Trojan is I

not a credible event.

First, the detectability of a large through-wall flaw under normal operating condi-tions is established by comparing calculations of the fluid leakage through the crack with the detection capabilities of five existing leak detection systems.

Second, the local stability of the crack under faulted loading conditions is determined using linear elastic fracture mechanics techniques.

Third, the local stability of the crack is evaluated under bending conditions assuming a fully-plastic pipe

.I section.

For both fracture mechanics evaluations, the demonstration of stability is based upon conservative, lower-bound fracture properties of the RCL piping and weld materials. Fourth, the global stability of the RCL is evaluated by showing that the reduced section at the postulated crack location is I

capable of carrying the piping loads for faulted conditions. The LBB evaluation described in this report is documented in References 10 and 11.

The leak-before-break approach is considered an acceptable method of evaluation for the Trojan d

reactor coolant loop because the RCL configuration is I

verified by as-built documentation and because the RCL does not experience any degradation effects which would invalidate the LBB criteria and methodology.

The significant PCL documentation aspects are:

(1 ) The design bases and margins of the RCL piping and components are not changed by the L8B

'I evaluation for pipe breaks.

(2) The as-built condition of the line is verified E

by field walkdown data and corresponds to the 5

as-designed condition.

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Report No. 01-0300-1395 Revision 1 Page 6 of 49

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(3)

In-service inspection of the RCL (in accordance with ASME Code Section XI) does not reveal any known flaws in the system.

The RCL piping is not subject to critical degradation mechanisms which limit the use of LBB analyses.

Specifically, the RCL is designed to provide the I

following:

(1 ) The piping material is not susceptible to intergranular stress corrosion cracking.

(2) The RCL does not experience water hammer k

loadings.

(3) The -iping is designe.d considering operating tran.,ients such that failure by high or low

nj cycle fatigue will not occur.

(4) The RCL piping material is not susceptible to cleavage-type fracture in its service temperature range.

(5) The RCL piping is' protected from failure by lJ indirect effects (missiles and fires).

3 Therefore, the LBB evaluation is a viable means of assessing the potential for rupture of the RCL piping and for resolving the asymmetric LOCA loads issue.

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Report No. 01-0300-1395 Revision 1 Page 7 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS 2.1 Criteria The leak-before-break evaluation was performed in I

accordance with the NRC criteria outlined in Generic Letter 84-04 which addressed the requirements for RCL leak-before-break evaluations.

In addition, the I

analysis was subsequently reviewed to ensure A\\

conformance with technical guidance provided in NUREG C

1061, Volume 3.

The seven specific criteria used are:

1) The loading considered in the analysis should include the static loads (pressure, deadweight I

and thermal) and the loads associated with safe shutdown earthquake (SSE) conditions.

2) The evaluation is performed considering that a large circumferential throughwall flaw is postulated to exist in the pipe wall. The length of the flaw is to be the larger of either (a) twice the wall thickness or (b) the flaw length that corresponds to a calculated leak rate of 10 gallons per minute (gpm) at normal operating conditions.

I (3) The material resistance to fracture should be based on a reasonable estimate of lower bound properties as measured by the materials resistance (J-R) curve. The effect of thermal

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I aging and the sensitivity of the piping and welding materials to cleavage rupture over the full range of operating temperatures should be addressed.

(4) The applicability of J-integral and tearing lg modulus computations should be addressed to ensure that the results are acceptable for g

engineering application.

(5) The flawed piping should be evaluated to

.l demonstrate that the code limits for faulted l,a conditions are not exceeded for the uncracked section.

!g (6) The potential failure of the RCL piping due to l3 indirect causes should be considered.

(7) Tne potential for development of a through-wall crack should be' evaluated in accordance with the I

- rules of IWB-3600 and Appendix A of Section XI of the ASME Boiler and Pressure Vessel Code.

(This l

analysis is described in Appendix C.)

2.2 Description of Piping The Reactor Coolant System (RCS) consists of four System similar heat transfer loops connected in parallel to the reactor pressure vessel (RPV) and other compo-I nents such as the pressurizer and connecting piping.

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Report No. 01-0300-1395 Revision 1 Page 8 of 49 Figure 1 shows a. typical loop; the locations considered in the fracture mechanics analysis are indicated. Each loop contains a reactor coolant pump, steam generator and three segments of large-diameter piping. The three segments of loop piping are the hot leg, which connects the RPV to the steam generator; the crossover leg, which connects I

the steam generator to the reactor coolant pump; and the cold leg, which connects the reactor coolant pump to the RPV. The diameters and wall thicknesses of the three segments of piping are given in Table 1.

The reactor coolant loop piping and fittings are austenitic stainless steel and the standards I

governing the materials used in the piping, fittings, and welds are given in Table 2.

Smaller piping which comprises part of the RCS boundary, such as the I

pressurizer surge line, spray and relief line, loop drains and connecting lines to other systems, is also austenitic stainless steel.

The design pressure of the RCL is 2485 psig and the normal operating pressure is 2235 psig. The design temperature is 650 F and the normal full power I

operating temperature is 616.8*F for the hot leg, 552.3*F for the crossover leg, and 552.5*F for the cold leg.

The restrained thermal expansion that the RCL has undergone has been considered as a potential indirect cause of piping degradation. The analysis of the I

piping under extreme bending loads, which is described in Section 2.4, provides assurance that unstable crack extension will not occur even under I

fully plastic section conditions. The causes and effects of restrained thermal expansion have been addressed in a separate effort.

2.3 Crack Detectability The size of the postulated flaw evaluated in the Under Normal Operating leak-before-break analysis is defined by comparing Conditions the leak rates from postulated cracks under nomal I.

operating conditions with the sensitivity of existing systems available to detect leakage from the primary system. The analyses performed to determine the I

detectable crack size are described in this section.

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Report No. 01-0300-1395 Revision 1 Page 9 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS 2.3.1 Crack Leakage Analysis Two calculations are performed to determine the leak I

rate through a postulated crack. First, the crack opening area of a postulated crack under the applied normal operating loads is determined. Second, the I

fluid flow through the crack opening area is cal culated. These calculations were performed using the computer codes CRACK and IMLEAK, respectively.

CRACK is based on Reference 1 and is described in Appendix A.

CRACK calculates the crack opening area using linear elastic fracture mechanics, including

- I corrections for plastic zones at the crack tips, and considers internal pressure, axial force and bending moment contributions to crack opening.

IMLE.AK is based on Reference 2 and is described in Appendix A.

IMLEAK evaluates flow through thin cracks in piping given the upstream thermodynamic

.I conditions and the geometric configuration of the

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crack. The program has been correlated with experimental data from measurement of two-phase flow through simulated and actual cracks in piping. The flow model accounts for two-phase flow and includes pressure drop terms due to entrance geometry, friction, and acceleration.

For this analysis, the leak rate calculations were performed for several postulated circumferential I

crack lengths, ranging from two to nine inches. Two possible loads due to normal operating conditions were considered. First, axial loads due to I

self-balancing pressure stress were used to determine lower bound leak rates. Second, bending moments due to thermal loads were added to the equivalent pressure loading to yielo an estimate of the upper I

bound leak rate for normal operating loads.

The leak rate analyses were performed using the pipe I

section geometry and operating conditions of the hot leg. The bending mcment applied to determine the leak rate in the second analysis described above was the moment at the junction of the hot leg and the RPV.

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Report No. 01-0300-1395 Revision 1 Page 10 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS I

The leak rates calculated for the postulated circumferential cracks are tabulated in Table 3.

The results considering the pressure only provide a lower bound estimate of the leak rate. The results which I

include the thermal bending moment provide an estimate of the leakage that would result from a crack at a location of high stress. The leak rate will lie between these bounds depending upon the loads at the location of the postulated crack. For locations of high stress, where a crack would be more likely to develop, the leakage would be high.

For I

locations of lower stress, which have less potential for crack development, the leakage would be closer to the lower bound leak rates.

2.3.2 Leak Detection Systems I

There are five methods of detecting leakage from the reactor coolant loop piping. The sensitivity and response time required to detect a one gpm leak rate have been calculated previously and are discussed in I

Section 5.2.5.4 of the Updated Final Safety Analysis Report, Reference 3.

The results are presented in i

UFSAR Table 5.2-9 which is reproduced in Appendix B of this report. The results are summarized in Table 4.

2.3.3 Detectable Crack Size The leak rates calculated for several lengths of postulated circumferential cracks are presented in I

Figure 2.

The leak rates calculated using pressure stress only and using pressure and thermal bending are graphed separately. Also shown in the figure are the sensitivities of the five existing leak detection systems.

The lower bcund leak rate curve shows that a crack I

slightly longer than seven inches can be detected by all five systems. The leak rate calculation which includes bending shows that leakage from a 3.5-inch I

crack can be detected by all five systems and that the leak rate from a seven-inch crack is in excess of eight times the flow rate detectable by the least sensitive system.

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Portland General Electric Report No. 01-0300-1395 Revision 1 Page 11 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS As discussed previously, the actual leak rate will depend upon the loads at the location of the crack.

The lower bound leak rate results show that a seven-inch crack can be detected.

The upper bound I

results show that a margin of two on detectable crack length exists for a seven-inch crack and that the leakage from a seven-inch crack would be three times the 10 gpm required by the criteria.

Based upon these results, a seven-inch-long crack is sufficiently large such that it has a very high I

probability of detection under normal operating conditions and is selected as the crack length for evaluation under faulted conditions.

I (Note that evaluation of a seven-inch long crack is consistent with Generic Letter 84-04. The leak rate calculations show that the leakage from a seven-inch I

crack is greater than a factor of ten over the detection system's sensitivity.

The specific guidance provided by NUREG 1961 defines separate b

I leakage and critical crack sizes which are not separately defined in Generic Letter 84-04 criteria.

The approach applied in this evaluation is consistent I

with the general guidance of NUREG 1061 that through-wall cracks be detectable with substantial margin and by multiple detection systems.)

2.4 Fracture Mechanics The postulated through-wall crack which was shown to Analysis be detectable under. normal operating conditions is evaluated to determine the potential for unstable I

growth under faulted loading conditions. The fre.cture mechanics analysis performed to determine the stability of tre postulated crack is described in this section.

2.4.1

_ Loading Conditions and Critical Locations The stress analysis of the RCL piping was performed previously and is described in Reference 4.

The analysis was performed for Loop 2 which includes the I

14-inch-diameter piping that connects the pressurizer to the RCL. Six load cases were analyzed in the stress report: deadweight of piping, wet weight _ of I

primary equipment, thermal, pressure, operating basis earthquake (0BE) X+Y, and OBE Z+Y.

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2.0 METHODOLOGY AND NUMERICAL RESULTS The separate load case-results were combined to provide upper bounds on the faulted condition loads.

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Deadweight, equipment wet weight, thermal,. pressure, and safe shutdown earthquake (SSE) results were conservatively combined to give upper bound estimates of axial force and bending moment for the RCL

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piping.

(The SSE load is 1.67 times the maximum of OBE X+Y and OBE Z+Y load.)

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The load combination method used in this evaluation differs slightly from the combination method specified in NUREG 1061. This evaluation uses a more

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conservative load combination method. Specifically, in this evaluation the static load components are g

combined vectorially (i.e. SRSS) and then added to the vectorially combined dynamic load components.

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This combination method results in higher loads than the NUREG 1061 method where the static and dynamic components are added in each direction and then

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combined vectorially.

The upper bounds on faulted loads were determined for the terminal ends of each of the three piping

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segments.

In. addition, upper bounds on the loads at the intermediate locations were calculated by enveloping the load case results for several

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intermediate locations on each segment. The enveloping procedure results in intermediate loads which are higher than the terminal end loads and

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which ensure that conservative upper bound faulted loads are considered in the analysis.

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The faulted condition loads used in the fracture L

mechanics analysis of the postulated seven-inch crack are listed in Table 5.

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2.4.2 Local Stability of Postulated Crack The local stability of a seven-inch-long postulated

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crack was determined by calculating the stress intensity factor at each of the nine locations listed

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Report No. 01-0300-1395 Revision 1 Page 13 of 49 I

2.0- METHODOLOGY AND NUMERICAL RESULTS in Table 5.

For cases of purely elastic or g

I small-scale contained plasticity, the J-integral may be derived from the stress intensity factor using the relations:

2 J = K /E (for plane stress)

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2 J = K (1-v )/E (for plane strain)

(Definitions of the terms used in this report are I

provided in the List of Symbols on page 38.)

The conversion to the J-integral was performed using I

the formula for plane stress conditions which results d in upper bound J-integral values.

The stress intensity factors were calculated using the program CRACK which is based on Reference 1 and is described in Appendix A.

CRACK includes the standard correction for effective crack length due to yielding at the crack tip. Minimum material properties from the ASME BPVC (yield stress, ultimate stress and elastic modulus) for the piping material at the operating temperature were used to ensure upper bound estimates of the effective crack length i

and the stress intensity factor.

The approach used for the faulted condition evaluation essentially calculates the elastic portion of the J-integral. This accurately represents the I

total J-integral if only small scale yielding occurs at the crack tips.

As shown in Table 6, the crack tip yielding is small for this evaluation.

I Therefore, the elastic portion of the J-integral is a 500d estimate of the total J-integral.

,g lne results of the J-integral calculations are l3 presented in {able 7..The maximum value is 595 in-lbf/in which was calculated for the upper

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bound intermediate loads en the hot leg.

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Portland General Electric Report No. 01-0300.1395 Revision 1 Page 14 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS NUREG 1061 provides several options for calculating of the applied J-integral considering plasticity effects.

For the evaluation of the RCL, plasticity I

could be important for loads in excess of the calculated design loads.

To provide a margin on l

loads, an analysis considering a fully-plastic section and a displacement controlled J-integral formulation was performed in accordance with the approach provided in Appendix A of NUREG 1061, Volume 3 and Reference 5.

This analysis provides assurance that unstable crack extension will not occur even under conditions of extreme bending where [

the section is fully yielded. This analysis provides

!E an assessment of pipe stability under conditions of

'E higher loading than considered under the modified NUREG 1061 approach which considers elastic-plastic fracture behavior and accounts for strain hardening.

,I An applied J-integral versus tearing modulus curve was determined for the RCL using the displacement-I controlled method outlined in Appendix A of NUREG 1061, Volume 3.

The analysis considered application of end rotations to the piping which would result in fully-plastic yielding of the cracked pipe section.

I For the fully-plastic analysis, a 90 throughwall crack was postulated. The crack length is 25 inches which provides substantial margin over the detectable

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crack length.

The tearing modulus values for J =

1000 and J = 2000 are 12.3 and 12.1, respectively.

2.4.3 Limits of Applicability When the yielding associated with a given crack and loading condition is small, the J-integral based on I

the elastic stress intensity factor can be used to assess stable crack growth using the resistance curve concept.

That is, for cases where small-scale yielding occurs, the stress intensity factor and J-integral based on it uniquely measure the intensity of the stress field at the crack tip. However, for cases where crack growth occurs under conditions of I

large-scale yielding, additional requirements must be met for application of J-resistance curve approach.

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Report No. 01-0300-1395 Revision 1 Page 15 of 49 I

2.0 METHODOLOGY AND NUMERICAL RESULTS The plastic yielding zones at the crack tips for the analyses presented here are small relative to the tots 1 section of the pipe. Table 6 shows the size of d the plastic zone at the crack tip and tabulates the I

size of the zone as a percentage of the uncracked portion of the pipe section. This comparison demonstrates that the yielding is small relative to I

the total section. Therefore the use of the J-integral based on the stress intensity factor using plastic zone corre::tions to the crack length is appropriate. To further demonstrate the adequacy of I

the approach the requirements for fully-yielded sections are also applied. However, it is to be expected that the margins of applicability will be large since the yielding is small.

Several requirements for assessing the applicability of J-integral controlled fracture have been proposed.

(See, for instance, Reference 6.) Two of g

the more widely used are evaluated here. The first requirement applies to both initiation and growth phases of cracking. The requirement ensures that the crack opening displacement is small relative to the wall thickness and uncracked ligament section. The requirement is:

b > 25 J/ ag The uncracked ligament, in the case of our analysis, is in excess of 40 inches. The pipe wall thickness I

is conservatively used as the structural dimension.

The inequality then becomes, for the maximum value of the J-integral calculated in this analysis:

2.43 > 0.37 The first condition is met.

The second requirement is based on considerations for the applicability of the deformation theory upon which J is based. The requirement is based on the nondimensional parameter w:

w = h h >> 1

Portland General Electric Report No. 01-0300-1395 R: vision 1 Page 16 of 49 2.0 METHODOLOGY AND NUMERICAL RESULTS In evaluating w, we again conservatively use the pipe wall thickness for b.

The lowest value of dd/da for the two weld materials considered is 7800.

Then:

w = 32 Generally a value of 10 is considered adequate, and therefore this condition is met.

2.5 Material Properties 2.5.1 J-Integral and Tearing Modulus Data for Base Metal The austenitic stainless steels used in the manufacture of reactor coolant piping exhibit extensive ductility. Quantitative characterization of the resistance to initial crack growth and the stability of crack growth is based on measurement of the J-integral and the tearing modulus.

Determination of these fracture toughness measures for the RCL piping materials are provided in References 6 through 9 ar.d tre summarized in I

Table 8.

Reference 6 provides J-R curves for forged 304N stainless steel at 600*F. This material and operating temperature corresponds to the RCL material and operating temperature, Figure 3 reproduces the J-R curve from Reference 6.

The test results for 304 stainless pipes, plates, and forgings, which is presented in EPRI Report NP-2261, Reference 7, was compared to the data presented in Figure 3.

The EPRI data is a lower bound on data covering a range of test specimen temperatures. The EPRI data indicates greater fracture resistance than the Reference 6 I

curve for 600*F and thus the curve for 600 F in g

Figure 3 is a lower-bound. J-T determinations for Type 316 cast stainless steel (at 600*F) from Reference 6 are provided in Figure 4.

Recent research presented in Reference 8 addresses the effect of thermal aging effects of 316 (CF8M) castings. JMC determination tests were conducted on CF8M spec mens that had been subjected.to accelerated thermal aging (7,500 to 10,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> at 600 F).

The J-R curve from these tests is presented I

in Figure 5.

The thermally-aged material was found to have high toughr.ess values of 571 lbs/in at room temperature and 856 lbs/in at 600*F. These values are comparable to those for submerged arc welds (SAW) as discussed in the following section.

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Report No. 01-0300-1395 Revision 1 Page 17 of 49 I

2.0 METHODOLOGY AND NUMERICAL RESULTS 2.5.2 J-Resistance Curve for Weld Material Data on the fracture resistance of stainless steel weld material was recently published in NUREG-1061, Report of USNRC Piping Review Committee, Reference g

9.

Data was presented for submerged arc welds (SAW) and inert gas tungsten arc welds (GTAW) in the fonn of applied J-integral versus measured crack extension. From this data, J-integral versus tearing modulus curves were derived for the two types of wel ds. The data from Reference 9 is reproduced in A

Figures 6, 7 and 8.

O To determine which welding process was more similiar to those used in the Trojan RCL, four typical weld procedures used on large diameter RCS piping were reviewed. The procedures include shielded metal arc welding (SMAW) and gas - tungsten arc welding (GTAW). These two welding procedures were used for the entire weld, or, alternately, the intitial GTAW root passes were completed using SMAW passes. No other welding procedures are allowed by the speci fications.

1I Consequently, the weld procedures used on the RCL are (for GTAW) identical or (for SMAW) closer to the l

J-resistance curve for GTAW welds. The lower

,E fracture resistance curve for SAW which produces l5 coarser grain stucture with lower ductility, provides i

a definite lower bound for fracture properties of the welding procedures used on the RCL.

2.5.3 Lower-Bound J-T Curves for Analysis lI Based on the review and comparison of applicable J-R curves for the RCL piping system materials and welds, the J-T curves for weld from NUREG 1061 provide lower bound fracture toughness data. The curves provide definite lower bound estimate of fracture resistance for the piping materials and welds and for thermally-aged castings.

d J-T curve for GTAW welds envelopes all referenced J-R l

test data for forgings and their welds (SMAW, GTAW).

The J-T curve for SAW welds is a lower bound for the l

J-R curves for CF8M castings after accelerated thermal aging. For conservatism, the lower bound SAW curve is used for the analysis of crack stability.

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Report No. 01-0300-1395 Revision 1 Page 18 of 49 2.0 METHODOLOGY.AND NUMERICAL RESULTS 2.6 Net Section Evaluation A large flaw such as the one postulated to produce detectable leakage from the RCL will reduce the load carrying capacity of the piping. To determine whether this effect is significant for the seven-inch-long flaw postulated for this analysis of the Trojan RCL, the net section stresses were evaluated using the ASME Code procedures and limits for faulted conditions. The primary stress for faulted conditions, calculated in accordance with Subsection NB-3656, is 27.9 ksi which is below the allowable of 37.6 ksi. Therefore, there is margin against failure of the flawed piping due to net section plasticity.

2.7 Summary of Results The results of the fracture mechanics evaluation are summarized in Figure 8.

The value of the applied J-integral calculated for the highest loads on the RCL piping are plotted with the material J-T curves for two stainless steel weld materials. Also, a J-T A

curve derived considering a fully plastic cracked LD section is indicated. Considering only the LEFM-based J-integral and the lower bound J-T curve, some crack extension may occur. However based on the applied J-T curve there is a wide margin on the tearing modulus which ensures that crack extension will occur in a stable manner.

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3.0 CONCLUSION

S The analyses presented in this report demonstrate I

that large margins against unstable crack extension exist for RCL piping postulated to have a long throughwall crack. Lower bound leak rate calculations show that a seven-inch circumferential crack is detectable by all five existing leak detection systems. Consideration of thermal bending I

results in higher leak rates and, at highly stressed locations, the leakage from a seven-inch crack is 30 gpm. Therefore, it is concluded that substantial margin to detect a seven-inch throughwall crack under normal operating conditions exists.

Local stability was shown using conservative load combinations of plant operating loads and safe shutdown earthquake loads and lower bound fracture mechanics properties of stainless steel weld metal.

Review of the weld procedures used on the RCL confirmed that the data for submerged arc welding provides a conservative lower bound on the fracture resistance of the RCL welds. For this lower bound f

data, a wide margin for the applied tearing modulus

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versus material tearing modulus was shown.

It is concluded that a wide margin against unstable crack extension exists under faulted conditions for the RCL piping.

The margin against net section failure of the cracked section was evasuated using upper bound loads for the faulted condition. The allowable stress is 33%

higher than the upper bound calculated stress and therefore substantial margin exists against net section failure.

The potential for development of a through-wall crack l

was evaluated.

A postulated flaw with the maximum j

permissible dimensions for intervice examinations was I

shown to grow a negligible amount (less than 1%)

under all design and postulated plant loadings.

b Therefore, based on analyses performed in accordance with the criteria of Generic Letter 84-04 and following the technical guidance of NUREG 1061, Volume 3, it is concluded that pipe rupture of the I

reactor coolant loop at Trojan Nuclear Plant in a double-ended guillotine manner is not a credible event.

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Portland General Electric Report No. 01-0300-1395 Revision 1 L_

Page 20 of 49 TABLE 1 REACTOR COOLANT LOOP PIPING DIMENSIONS Piping Segment Inside Diameter (inch)

Wall Thickness (inch)

Hot Leg 29.0 2.43 Crossover Leg 31.0 2.57 Cold Leg 27.5 2.30 l

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TABLE 2 REACTOR COOLANT PIPING MATERIALS Component Product Material Spec Type Grade Reactor Coolant Pipe Forging Code Case 1423 F304N & 316N Casting SA 351 CF8A & CF8M b

Reactor Coolant Fittings Casting SA 351 CF8A & CF8M

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Welding Materials Electrodes SFA 5.4 E308 & E308L

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Rods SFA 5.9 ER308 & ER308L

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[

[

m--_m

Portland General Electric Report No. 01-0300-1395 Revision 1 Page 22 of 49 TABLE 3 LEAK RATE UNDER NORMAL OPERATING CONDITIONS Leak Rate (gpm)

Crack Length (inch)

Pressure Stress Only' Pressure plus Thermal Bending 2

0.11 0.98 3

0.34 2.91 4

C.75 6.26 5

1.40 11.3 6

2.33 19.7 7

3.57 30.0 8

5.18 43.4 9

7.22 60.3 l

1 - Axial stress equal to pr/2t.

2 - Axial stress equal to pr/2t plus thermal bending moment at RPV outlet nozzle / hot leg.

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E Portland General Electric 5

Report No. 01-0300-1395 Revision 1 Page 23 of 49 TABLE 4 SENSITIVITY AND RESPONSE TIME OF LEAK DETECTION SYSTEMS Sensitivity Response Time (leak rate detectable (Time required to detect Leakage Detection System in one hour) 1 gpm leak rate)

Containment Air Particulate Monitor 0.6 gpm 45 min Containment Gaseous Activity Monitor 2.5 gpm 100 min Containment Humidity Detector System 2.0 gpm Not Sufficiently Sensitive Containment Condensate Monitoring System 0.8 gpm 50 min Containment Sump Level 3.7 gpm 220 min I

From Reference 3.

See Appendix B for applicable notes concerning methodology and assumptions.

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E Portland General Electric -

5 Report No. 01-0300-1395 Revision 1 Page 24 of 49

I TABLE 5 LOADS FOR FAULTED CONDITION EVALUATION Location and j

I Piping Segment Stress Analysis Node F (KIPS)

M (KIP-IN)

~

RPY Outlet (Node 405) 2085 30,406 2

Hot Leg Intermediate 2586 29,264 SG Inlet (Node 420) 2187 30,014 SG Outlet (Node 438) 1778 10,953 2

Crossover Leg Intermediate 1938 22,069 RCP Inlet (Node 459) 1968 20,424 RCP Outlet (Node 468) 1542 12,307 2

Cold Leg Intermediate 1560 16,404 RPV Inlet (Node 483) 1551 15,404 I

1 - Locations and stress analysis nodes are shown in Figure 1.

2 - Upper bound estimate of maximum loads on intermediate piping segment.

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Pcrtland General Electric R: port No. 01-0300-1395

(-

R: vision 1 Page 25 of 49

[

h TABLE 6 A

I C

PLASTIC ZONE SIZE UNDER FAULTED LOADING CONDITIONS

[

Piping Segment Stress Analysis Node r

(in)

% Plasticity RPY Outlet (Node 405) 1.19 2.60 Hot Leg Intermediate 1.42 3.10 SG Inlet (Node 420) 1.22 2.66 b

SG Outlet (Node 438) 0.172 0.35-Crossover Leg Intermediate 0.376 0.76

[

RCP Inlet (Node 459) 0.349 0.71 RCP Outlet (Node 468) 0.301 0.69

{

Cold Leg Intermediate 0.417 0.96 RPV Inlet (Node 483) 0.386 0.89

[

[

l - Results are for 7-inch circumferential crack 2 - Locations and stress analysis nodes are shown in Figure 1.

{

3 - Radius of plastic zone correction at crack tips 4 - Percentage of the uncracked piping section included in the plastic zones at the tips of the postulated crack

[

[

[

[

j Portland General Electric 5

Report No. 01-0300-1395 Revision 1 Page 26 of 49 I

TABLE 7 g

J-INTEGRAL VALUES FOR FAULTED LOADING CONDITIONS Location and 1

2 Piping Segment Stress Analysis Node J (in-lbf/in )

RPV Outlet (Node 405) 495 Hot Leg Intermediate 595 SG Inlet (Node 420) 509 SG Outlet (Node 438) 74 g

m Crossover Leg Intermediate 173 RCP Inlet (Node 459) 159 RCP Outlet (Node 468) 136 Cold Leg Intermediate 195 RPV Inlet (Node 483) 179 lI

!I 1 - Locations and stress analysis nodes are shown in Figure 1.

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E Portland General Electric 5

Rep;rt No. 01-0300-1395 Revision 1 Page 27 of 49 I

TABLE 8 J-INTEGRAL AND TEARING MODULUS DATA I

NATERIAL PRODUCT Jrc TEST Je dJ/da TEARING REFERENCE t

GRADE SPECIMEN TEMP.

Ib/in I b/in2 MODULUS (THICrNESS) 304 PIPE 4P8 RT 6,000 100 EPRI NP-2261 I

SA-312 (0.35 in)

Ref. 7 304 PLATE CT 600T 1,500 24,300 333.8 BAMFORD SA-240 (1.0 in)

& BUSH Ref 6 CT RT 4,450 28,600 188.1 304N FORGING (CASE-1423-1 )

(2 ini 600*F 2.569 20.700 2 31. 1 BAMFORD 4 BUSH (3PB)

RT 4,000 107,400 706.3 Ref. 6 2 in 600T 2,737 49,000 547.0 GT RT 4,293 47,400 36 4.1 CF9t(316)

CASTING SA 351 (2 in) 600*F 1,933 19,000 225.1 BAMFORD 4 BUSH 3P8 RT 4,000 81,700 627.5 Ref. 6 (1.75 in) 600T 1,200 51,700 612.6 I

CF9t(316)

Not RT 571 5,800 AGED: 7,500 hr CASTING Indicated A. VIGNES 9 750 T 608 T 856 5,800 Ref.,*'tS

(

WELOS SAW CT 550T 5 56 Figures 788 NUREG ER-308-STA5.9 (420) 7,812 Volume 1 Appendix F GTAW CT

$50'F 960 25,800 Figures 688 Ref. 9 (1 f n)

PIPES LOWER BOUNO FORGING 550Y 960 25,800 Figure 8 NUREG 1061 DATA FOR GTAW Volume 1 ANALYSIS CA5 TING Appendfx F 550 Y 420 7,812 Figure 8 Ref. 9 AGED 4 SAW I

Portland General Electric Report No. 01-0300-1395 I

Revision 1 Page 28 of 49 I

FIGURE 1 REACTOR COOLANT LOOP GE0 METRY (TYPICAL)

I I

~

I kN f

na N

g

.n

',I

.,rg A

FM".[ f i

I C

f I

/

=

us

,s m

/

,../

(

[

ss

[

N L.

. p

',y' E

's

\\

N

'f

'.. [

g MD i

14 I

I

$ denotes break location analyzed:

A - Reactor Pressure Vessel Outlet Nozzle (Node 405)

I B - Steam Generator Inlet Nozzle (Node 420)

C - Steam Generator Outlet Nozzle (Node 438)

D - Reactor Coolant Pump Inlet Nozzle (Node 459)

I E - Reactor Coolant Pump Outlet Nozzle (Node 468)

F - Reactor Pressure Vessel Inlet Nozzle (Node 483)

(Bounding loads for intermediate locations on each leg were also analyzed.)

This figure is based on Figure 3.6-2 of the Trojan UFSAR (Reference 3).

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Portland General Electric I

Report No. 01-0300-1395 Revision 1 Page 29 of 49 FIGURE 2 CALCULATED LEAK RATES FOR NORMAL OPERATING CONDITIONS AND LEAK DETECTION SYSTEM SENSITIVITIES l

25 re nding 20 15 1

I i

3 10 l

Pressure Only

~

sf vt

~"

,/

,/

b d _

==T 1

2 3

4 5

6 7 8 9

10 cuct sin (tach)

Leak Detection Systems:

a - Sump Level b - Gaseous Activity Monitor c - Humidity Detector System d - Condensate Monitoring e - Air Particulate Monitor

Portland General Electric Report No. 01-0300-1395 Revision 1 Page 30 of 49 I

I FIGURE 3 JIC DETERMINATION FOR 304 FORGED STAINLESS STEEL (COMPACT SPECIMENS)

J (In. Lb./in!)

22000 LEGEND:

S CircumferentialOrientation RT g p,7,ygi;r:p,il ior;ytanoa.6oo r 2mm a AslalOrientation 6oo*F g

16000 Blunting Line

(

• **DI Blunting Line 14000 (600*F) 12000 I

A i==

8000 I

.0.

O a

40

• O 2000 f

O I

l i

i 1

0 0.10 0.20 0.30 0.40 0.50 aa (Inches)

I I

From Reference 6 - W. H. Bamford and A. J. Bush, " Fracture Behavior of Stainless Steel s. "

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Portland General Electric Report Noo 01-0300-1395 Revision 1 Page 31 of 49 I

FIGURE 4 JIC DETERMINATION FOR 316 CAST STAINLESS STEEL (COMPACT SPECIMENS)

I J (In. Lb./in!)

l Blunting Line (Room Temp.)

Blunting Line

\\

18000 16000

g l

12000

~

l 0

0 8000 3

84 6000 0

4000 I

o LEGEND:

2000 E circumierential Orientation. nT O Circumferential Orientation,6oo*F A Atlet Orientation, RT 0

a Asiat Orientation,6oo*F I

1 0

0.10 0.20 0.30 0.40 0.50 Aa (Inches)

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From Reference 6 - W. H. Bamford and A. J. Bush, " Fracture Behavior of Stainless Steel s. "

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Portland General Electric

~

Report No. 01-0300-1395 Revision 1 Page 32 of 49 FIGURE 5 J-R CURVE FOR AGED CF8M (7,500 hr at 750*F)

I

!I i

i i

f

- 3000 "j

500 g

l

=

I?

320 c (608 F) e e LOC uy g

3

-s

- 2000 3

=

100 RT A

. a.

200

~

e : es

- 1000 0

p" A 829 %

p" O at I teos seced e

, E ICO O 3394l SPECmtal i

l i

f f

0 s

i 2

3 5

,)

4

!I From Reference 8 - A. Vignes 4

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Portland General Electric Report No. 01-0300-1395 Revision 1 Page 33 of 49 I

FIGURE 6 g

J-RESISTANCE CURVE FOR TUNGSTEN INERT GAS WELD (550*F)

I

!I 4

14000

/

ros sia, croov..

A IT COMPACT SPECIMEN

~D IT COMPACT SPECIMEN O

O 2T COMPACT SPECIMEN 4

I "'*

o 2T COMPACT SPECI N D

o I

o 0

a O seee.

/

O o

a 0

o o

kesee.

/

h

'o 0

n 4ees_

U On " a 2eee-00 e

e.eea e.e30 e.ese e.ese e.12e e.tse e.tse e.2te e.248 CRACK EXTENSION in INCHES I

I From NUREG 1061 Volume 1, Report of the USNRC Piping Review Committee, Refercnce 9.

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Portland Gen:ral Electric Report No. 01-0300-1395 R3 vision 1 Page 34 of 49 b

FIGURE 7 J-RESISTANCE CURVE FOR SUBMERGED ARC WELD (550*F) 1I a

f, f,

m,u. e.

y...e r

y<.e.

q

.se p

1:

...e

/

/

2..e

/

/

I i 2.e

/

/

~

li.se

/

/

. ce.. -

-o2 lI

=.

=: '

1

...e

/

/

M. I !.!!'.d..?.

I I/

/

.e sic

..se

~

e eee.ee e.2e e.4e e.ee e.se i.ee i.2e i.4e i.ee s.se 2.ee enAcmenowTH czw) e-ei I

I I

From NUREG 1061, Volume 1, Report of the USNRC Piping Review Comittee, Reference 9.

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3 Portland General Ele.ctric 3

Report No. 01-0300-1395 R vision 1 Page 35 of 49 I

. FIGURE 8 APPLIED J-T CURVE COMPARED TO LOWER B0UND J-T CURVES I

I 7

I 6

I 5

TUNGSTEN INERT GAS WELD

.E LOWER BOUND FOR A

)4 PIPES. PLATES &

O/

7 FORGINGS - TYPE 316

.c GTAW & TYPE 308 3

J-T Applied f

I 2

SUBMERGED ARC WELD LOWER 80lmD FOR CF8 & CF8M i

CASTINGS & SAW TYPE 308

-- --- J = 595 0O 50 100 150 200 250 300 TEARING MODULUS. T i

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l

I Portland Gantral Electric Report No. 01-0300-1395 Revision 1 Page 36 of 49 REFERENCES 1 - H. Tada and P. Paris, " Estimation of Stress Intensity Factors and the Crack Opening Area of a Circumferential and a Longitudinal Through-Crack in a Pipe," Del Research I

Corporation, St. Louis, Missouri.

2 - R.P. Collier, et. al., " Study of Critical I

Two-Phase Flow Through Simulated Cracks,"

Interim Report BCL-EPRI-80-1, Battelle Columbus Laboratories, 1980.

3 - Portland General Electric Company, " Trojan Nuclear Plant - Updated Final Safety Analysis Report," July 1982 plus Amendments 1 and 2.

4 - Westinghouse Electric Company, " Stress Analysis of Reactor Coolant Loop / Support Systems for Trojan Nuclear Plant, Unit No.1," WCAP 8438, Revision 1, December 1975.

I 5 - H. Tada, P.C. Paris, and R. Gamble, " Stability Analysis of Circumferential Cracks in Reactor Piping Systems," NUREG/CR-0838, June 1979.

6 - W.H. Bamford and A.J. Bush, " Fracture Behavior of Stainless Steel," in ASTM 668, Elastic-Plastic Fracture, November 1977.

7 - K.H. Cotter, et. al., " Application of Tearing b

Modulus Stability Concepts to Nuclear Piping,"

EPRI NP-2261, February 1982.

8 - A. Vignes, " Understanding of the Phenomena of Materials Degradation by Aging and Embrittlement I

- Margin of Safety and Solutions to Limit or Avoid such Degradations" in " International Conference on Nuclear Power Plant Aging, I

Availability Factor and Reliability Analysis,"

San Diego, California, July 1985, pp.165-168.

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Portland General Electric Report No. 01-0300-1395 Revision 1 Page 37 of 49 REFERENCES (Continued) 9 - Report of the U.S. Nuclear Regulatory Commission Piping Review Committee, NUREG 1061

- Volume 1, " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Boiling Water Reactor Plants, August,1984.

- Volume 3

" Evaluation of Potential for Pipe d

!E Break," October,1984 Volume 5

" Piping Review Committee 5

Conclusions and Recommendations," April,1985.

l 10 - Impe11 Calculation RCL-1, Revision 1, Job 0300 - 023.

lE 11 - Impe11 Calculation RCL-2, Revision 0, l5 Job 0300 - 023.

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Report No. 0300-1395 R: vision 1 Page 38 of 49 I

I LIST OF SYMBOLS I

I a

Half-crack length a,ff Effective half-crack length, accounting for yielding A

Crack opening area due to bending b

A Total crack opening area total A

Crack opening area due to pressure p

A Crack opening area due to axial tension t

b Structural dimension:

uncracked ligament or wall thickness E

Young's Modulus J

J-integral J

Calculated J-integral y

J Material J value IC K

Stress intensity y

K Stress intensity due to bending b

K Stress intensity due to pressure p

K Stress intensity due to axial tension t

M Bending moment p

Internal pressure of pipe r

Inner radius of pipe or radius of plastic zone at crack tip jl l

R Mean radius of pipe

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g Portland General Electric 3

Report No. 01-0300-1395 R: vision 1 Page 39 of 49 I

LIST OF SYMBOLS (Continued)

I I

t Average wall thickness of pipe T

Tearing modulus T,

Axial load w

EPFM nondimensional parameter e

Half the subtended angle of the crack a

Longitudinal stress due to bending b

o Flow stress g

I a"

Longitudinal stress due to pressure Longitudinal stress due to axial tension t

A Non-dimensional crack parameter, A= a/WT y

Poisson's ratio I

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E Portland General Electric 5

Report No. 01-0300-1395 Revision 1 Page 40 of 49 APPENDIX A Description of Computer Programs _

I Descriptions of the two Impell computer programs, CRACK and IMLEAK, which were used in the leak-before-break evaluations are presented in this appendix.

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R3 port No. 01-0300-1395 Revision 1 l

Page 41 of 49 APPENDIX A DESCRIPTION OF COMPUTER PROGRAMS l

Description of CRACK The program CRACK is based on the method developed by Tada and Paris (Reference 1). The equations used to calculate KI and opening area for circumferential

'I cracks are presented below.

The circumferential through-crack may be subjected to I

a tensile axial load, Ta; a bending moment, M; and internal pressure, p.

The total Ky is composed of three components:

K otal " Kt+Kb+Kp t

where Kt"#t (nRG)

Ft Kb*#b (nR6)I/2 Fb p (nR 6 )l/2F K

=g p

p T

Ut" M

Ub",]

o-M g

p

= 1 + 7. 5 (j)

- 15 (f)2.5 +33(f)3.5 1.5 Ft I

b = 1 + 6.8 (f)i.5 - 13.6 (f)2.5 +20(f)3.5 F

M 2

F = (1 + 0.3225A )

, for 05Asl p = (0.9 + 0.25A), for 1<As5 il I

I

Portland General Electric Report No. 01-0300-1395 Revision 1 Page 42 of 49 APPENDIX A DESCRIPTION OF COMPUTER PROGRAMS The crack opening area for a circumferential through-I crack is a function of the crack length, internal pressure, axial tension, and bending moment. The total crack opening area is given by:

A otal = At+Ab+Ap t

where 2

+

nR I El +

(

)]

At*Ab=

t 4

= h (2nRt) G A

p p

a = T,/2nRt t

O = M/nR t b

o = pR/2t p

2 2 [1 + (f)1.5 f 8.6 - 13.3 (f) + 24 (j) f I = 26 t

+(f)3 f22.5-75(f)+205.7(f)2 -247.5(f)

+242(f)

], for 0<6<100*

2 4

G = A + 0.16 A, for OsAsl 2

3 4

G = 0.02 + 0.81 A + 0.30A + 0.03 A, for 1< A55

.P I

I The above formulae are based on linear elastic frac-lI ture mechanics. Usually the material near the crack tip yields, and this is accounted for by a plastic-l zone correction. An effective half-crack length which l

accounts for the plastic yielding is computed by:

II l I l

Portland General Electric Report No. 01-0300-1395 Revision 1 Page 43 of 49 APPENDIX A DESCRIPTION OF COMPUTER PROGRAMS 2

I a,7f = a +

2 2na n The a ff value thus computed is used to recalcu-e late K, and the procedure is repeated iteratively I

until the a ff values between two consecutive e

I computations are within 2%.

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Portland General Electric Report No. 01-0300-1395 Revision 1 Page 44 of 49 APPENDIX A DESCRIPTION OF COMPUTER PROGRAMS Description of IMLEAK IMLEAK is a computer code developed to evaluate the leak rate of pressurized fluids through narrow I

cracks. The code is based on the approach outlined in Reference 2.

The code is used in leak-before-break analyses to estimate leak rates from postulated

,gg cracks to determine whether the crack is detectable.

The analytical model in IMLEAK is a modified verision of Henry's non-equilibrium two-phase flow model. The model accounts for non-equilibrium effects in the flow due to flashing within the flow path. The model also includes pressure drops due to entrance losses, I

friction, and fluid acceleration.

The model handles complex crack geometry, including turns in the flow path, variable flow area, and crack surface roughness. The program also contains its own steam properties subroutines. The program has been benchmarked against test results.

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Portland General Electric I

Report No. 01-0300-1395 Revision 1 Page 45 of 49 APPENDIX B Sensitivity of Leak Detection Systems l

Table 5.2-9 from the Updated Final Safety Analysis I

Report, Reference 3, is reproduced in this appendix.

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l Portland General Electric Report No. 01-0300-1395 Revision 1 Page 46 of 49 APPENDIX B SENSITIVITY OF LEAK DETECTION SYSTEMS TAB u 5.2 9 SENSITIVITY AND RESPONSE TIME OF REACTOR C001 ANT FRESSURE BOUNDARY LEAEAGE DETECTION STSTEM sensitivity Responsa Time I

14akage Detection (leak rate detectable (Time required to detect System in one hour) 1 mon leak rete)

I*

Containment Air 7 articulate Monitor 0.6 spa 45 min I

I Containment Geseous Activity Monitor 2.5 spa -

100 min containment Bumidity Detector System

• 2.0 spa Not sufficiently Sensitive Contai pt Condenaste Monitoring I

System s 0.8 spa 50 min Containment Sump Leve1 *I 3.7 spa 220 min E

[a] (1) Sensitivity threshold for identifying RCF3 leakage is 7.0 x 10" WCi/cc.

(2) Reactor coolant corrosion product activity level is 0.005 bCi/c1.

(3) All other assumpticus are as described in Section 5.2.4.4 for the Containment air particulate monitor evaluation.

(4) Detector delay time is 20 min.

[b] (1) Sensitivity threshold for identifying RCFB leakage is 1.0 x 10*' 9Ci/cc.

(2) Reactorcoolantgaseousactivit{81evel is 0.15 pCi/ce.

(3) Containment volume is 5.60 x 10 cc.

(4) Containment leakage is sero. Normal recirculation flow is 855.500 cfm and is suf ficient to ensure complete mixing.

(5) The time of activity buildup to detectable levels is much shorter than the half-life of elements of concern.

[c] (1) 30 percent of RCFB leakage flashes to vapor at the point of leakage.

(2) Containment leakege is sero. Normal recirculation flow of Containment air coolers is sufficient to ensure complete mixing.

(3) Initial Containment conditions are 120*F and 53 percent RB.

(4) Busidity detector sensitivity threshold for identifying RCFB leakage is

.I 3 percent RH.

[d] (1) 30 percent of RCFB leakage flashes to unter vapor at the point of leakage.

(2) Containment leakage is sero. Normal recirculation flow of Containment air coolers is 855,500 cfm and is sufficient to ensure complete mixing.

I (3) Air coolers condense all leakage vbich has flashed to vapor.

(4) Sensitivity threshold for identifying RCFB leakage is three-esive drainage cycles within one hour.

[e] (1) 50 percent of RCFB leakage drains as liquid directly to Containment Building sep with least eensitive leahage detection capability.

(2) Sump is initially empty and has 10-gallon / inch capacity.

(3) Sump level indicators are provided as described in Table 11.2-9.

(4) Sensitivity threshold for identifying RCFB 1eakage is receipt of seccad sump level light.

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Portland General Electric I

Report No. 01-0300-1395 Revision 1 Page 47 of 49 APPENDIX C Fatigue Crack Growth Analysis

'E The approach and results of a crack growth analysis 3

performed to evaluate the potential for development of a through-wall crack is presented in this appendix.

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L Portland General Electric p

Report No. 01-0300-1395 L

Revision 1 Page 48 of 49 r

L APPENDIX C FATIGUE CRACK GROWTH ANALYSIS L

A calculation was performed to determine the growth r

of a postulated flaw under design and postulated L

plant loadings. The calculation was done using the fracture mechanics procedure requirements of NUREG 1061 Vol. 3, Paragraph 5.6, Reference 9.

The plant loadings include dead weight, pressure, thermal expansion (including the effects of postulated steam generator snubber lockup), Operating Basis Earthquake (0.B.E.), and Safe Shutdown Earthquake (S.S.E.).

{

From review of the plant loadings, the locations chosen for analysis were the hot leg welds at the

[

reactor vessel outlet and steam generator inlet. The loading at these two locations envelope loading conditions throughout the reactor coolant loop. The

{

cross section used for the analysis was the nominal pipe dimensions for the hot leg.

The fracture d

toughness material properties were based on those of a submerged arc weld which represents a lower bound on these properties.

Initial flaws were postulated as inside surface,

[

semi-elliptical circumferential flaws. The depth of the flaws used was the maximum permissible dimension for inservice examinations as specified in Table IWB-3514-2 of Section XI of the ASME code.

The

[

allowable flaw depth is a function of the aspect ratio, depth to length ratio, and an aspect ratio of 1/6 was assumed. Using these criteria and

[

assumptions, the postulated initial flaw depth is 0.25 inch and length is 1.50 inches.

[

Using the criteria previously identified and conservatively applying loadings in compliance with NUREG 1061, final growth at the end of RCL design life time was.00154 inch.

This is less than 1% of

[

the initial flaw size and the growth is essentially negligibl e.

[

[

[

[

Portland General Electric Report No. 01-0300-1395 Revision 1 Page 49 of 49 APPENDIX C FATIGUE CRACK GROWTH ANALYSIS I

These results indicate tnat although there could be some flaw growth, the growth would be negligible.

Therefore, the structural integrity of the RCL piping i

is not expected to be jeopardized by fatigue crack penetration during the design life of the piping.

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Domestic Offices 350 Lennon Lane Walnut Creek, CA 94598 j

2'345 Waukegan Road Bennockburn, IL 60015 225 Broad Hollow Road l

4 Melville, NY 11747 g

800 South Street The Watermill Center Waltham, MA 02154 333 Research Court Technology Park / Atlanta Norcross, GA 30092 International Offices Impell Corporation A Division of Combustion Er,gineering Ltd.

Genesis Centre, Garrett Field Birchwood, Warrington g

WA3 7BH g

United Kingdom Impell-France S A.R.L.

10, Rue du Colise'e 75008 Paris France Affiliate Company Impell Pacific 920 Southwest Sixth Avenue Portland, OR 97204 2201 Dwight Way Berkeley, CA 94704 Progest S.p.A. (Fiat TTG)

Via Cuneo 21 10152 Torino Italy

. -.... ~ - -

-. - - -. - -.. -