ML20247L184

Leak-Before-Break Evaluation for Stainless Steel Piping, Byron & Braidwood Nuclear Power Stations Units 1 & 2

ML20247L184
Person / Time
Site:	Byron, Braidwood, 05000000
Issue date:	05/12/1989
From:	Deboo G, Hoang P, Meister J SARGENT & LUNDY, INC.
To:
Shared Package
ML20247L175	List: ML20247L171 ML20247L184
References
SL-4518, NUDOCS 8906020103
Download: ML20247L184 (114)

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EXECUTIVE

SUMMARY

On October 27,1987, the Federal Register published a change to 10 CFR 50, General Design Criteria (GDC) 4, Requirements for Protection Against Dynamic Effects of Postulated Pipe Ruptures. This rulemaking modifies GDC 4 to the extent that dynamic effects of pipe ruptures in nuclear power units may be excluded from the design basis provided it is demonstrated that the probability of pipe rupture is extremely low under conditions consistant with the design for the piping. The amendment allows exclusion of the dynamic effects associated with high-energy pipe rupture by application of leak-before-break technology.

NUREG-1061, Volume 3, " Evaluation of Potential for Pipe Breaks" provides the NRC Pipe Break Task Group's recommendations for the application of the leak-before-break (LBB) approach to the NRC licensing process. This document provides a step-by-step process to develop a technical justification for application of the LBB approach to a particular project.

The Byron /Braidwood Leak-Before-Break Evaluation follows the step-by-step approach of NUREG-1061 to develop the basis for application of leak-before-break technology to the Byron /Braidwood units.

The implementation of leak-before-break technology to the Byron and Braidwood stations will allow the dynamic effects of high-energy line rupture to be excluded from the design.

The exclusion of these dynamic effects will allow removal of pipe whip restraints and deletion of snubbers with the following results.

DEL 5iTION OF PIPE WHIP RESTRAINTS There are currently 64 pipe whip restraints installed on the systems in the LBB scope.

I These whip restraints are installed in congested areas of the containment building and with minimal gaps between them and the piping system. The possibility exists for a whip restraint I

to interfere with the thermal expansion of the piping and potentially cause an unanalyzed st-4518 lI

SARGENT & LUNDY SL-4518 05-12-89 ES-2 Q

l condition. The integrated hot functional test included walkdowns to identify and resolve these types of interferences for all normal thermal modes. This potential exists, however, for any unidentified or unanalyzed thermal mode. In addition, the whip restraints take up space needed for maintenance and in-service inspection (151) activities. They cause an I

increase in the time required to perform these activities and, as a result, increase the radiation dose to the plant workers. The whip restraints are located such that they easily become contaminated during normal plant operation and require extensive decontamination.

Therefore, the deletion of the pipe whip restraints will improve maintenance and 151, reduce personnel dosages in containment, and reduce the amount of contaminated material in the containment.

DELETION OF SNUBBERS Currently, piping analysis for systems connected to the reactor coolant system (RCS) includes an evaluation due to the loss-of-coolant accident (LOCA) forcing function for the appropriate RCS nozzle. This evaluation eften produces the governing piping forces and movements for the connected piping. To accommodate these dynamic loads at locations with high thermal movements, snubbers are often used as dynamic pipe supports. The Station Technical Specifications require snubber surveillance and testing during each refueling outage. Performing these activities in the containment building results in high radiation doses to the inspection and testing personnel.

A goal of the LBB program is to document the acceptability of the LBB approach for all loop connected piping that creates a break forcing function significant enough to require con-sideration for the piping attached to the adjacent loops. Once this assumption has been proven, the LOCA forcing function can be eliminated from the piping analysis for the loop l

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connected pdping. Eliminating the LOCA forcing function will allow a reanalysis to optimize the number of snubbers on this piping, which will significantly reduce snubber inspection and f

testing requirements.

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SARGENT E LUNDY 05-12-89 1

ES-3 l-The remainder cf this report is broken into seven sections as follows:

Scope System Review Versus NUREG-1061 Limitations a

Material Property Data i

Applied Loading q

Leak Detection Leak Rate Crack Sizing Crack Stability Analysis j

A brief description of the contents of each of these sections is as follows:

Scope This section describes the systems within the scope of the LBB program and addresses their j

safety functions. Piping and pipe whip restraint data are provided along with an isometric i

drawing of each piping subsystem.

Sy:.c - Review Versus NUREG-1061 Limitations The applications of LBB technology is limited by six criteria specified in Volume 3 of NUREG-1061, Section 3.1. This sec; ion of the report evaluates each Byron /Braidwood LBB system against the limiting criteria and demonstrates that the requirements are met. The six criteria of NUREG-1061 are as follows:

The design criteria for emergency core cooling systems, containment, and other engineering safety features remain unchanged.

Each system was reviewed for its potential for damage from the effect of stress corrosion, water / steam hammer, thermal and mechanical fatigue, erosion / corrosion, and creep. The Byron /Braidwood LBB systems were found I

not to be susceptible to these failure mechanisms.

l Component and piping support structural design margins have been main-I tained without reduction.

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SARGENT E LUNDY SL-4518 05-12-89 I

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l The evaluation of the Byron /Braidwood LBB systems for the probability of l

degradation or failure due to indirect causes such as fires, missiles, and equipment, systems, or components in close proximity was performed during the design of Byron /Braidwood and found acceptably low.

Each of the Byron /Braidwood LBB systems has been verified against the 8

requirements of I&E Bulletin 79-14.

The stainless steel materials in the Byron /Braidwood LBB systems are not

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I susceptible to cleavage type fractures over the full range of their operating temperatures.

Material Property Data The material tensile and toughness properties used in sizing the leakage crack and evaluating crack stability were obtained from the industry data presented in the NRC's Piping Fracture Mechanics Data Base (PIFRAC), NUREG/CR-4894. These generic material properties were found to be applicable to the Byron /Braidwood LBB systems by establishing that the PIFRAC tensile properties bounded the range of tensile properties obtained from the installed materials at Byron and Braidwood. This section presents the chosen mean and lower bound tensile properties used in sizing the leakage crack and evaluating crack stability respec-tively. It also establishes the lower bound toughness properties used to evaluate crack I

stability.

Applied Loading The most limiting material properties established in the preceding section are applied to the highest stress locations in each Byron /Braidwood LBB system. The highest stress location is presented for each system using the I&E Bulletin 79-14 analysis results for each analytical model.

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SABOENTS LUNDY SL-4518 05-12-89 ES-3 I

I Leak Detection This section describes the Byron and Braidwood leak detection systems including the Technical Specification requirements.

The section provides the basis for using the assumption that a leakage rate of I gallon per minute (gpm) can be identified by the operators within one hour.

Leak Rate Crack Sizing The leakage size crack (LSC) was established for each Byron /Braidwood LBB system using the EPRI computer code PICEP (Pipe Crack Evaluation Program). The results presented in this section were based on the system's normal operating fluid conditions, the algebraic sum of the weight and normal operating thermal loads, the mean material tensile properties, and a leak rate of 10 gpm.

Crack Stability Analysis The crack stability evaluation was performed using master curve methodology based on a modified limit load approach which uses Z factors to conservatively evaluate crack stability by elastic-plastic fracture mechanics methods. Master curves were developed for base and weld metal cracks in each Byron /Braidwood LBB system. From these master curves, the I

critical crack length (CCL) associated with the normal operating load (NOL) absolutely combined with the safe shutdown earthquake (SSE) load was obtained and compared to the I

LSC demonstrating the required margin of 2.0 on stable crack size. The combined normal operating and SSE load was increased by the required factor of 1.4 and used to obtain the CCL from the master curve. This CCL was compared to the LSC demonstrating the required margin on load by being greater than the LSC. In addition, a 3-Tearing evaluation was performed on the cast stainless steel components using the EPRI computer code FLET, Flaw Evaluation by Tearing Instability. This evaluation confirmed the results of the master curve methodology.

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1 In conclusion, this report documents that the three systems in the Byron /Braidwood LBB pro-gram met the system requirements in NUREG-1061 Volume 3. The margin of 10 on leakage detection, the margin of 2 on postulated crack size, and the margin of 1.4 on loads as

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required by the October 27, 1987., Federal Register have been demonstrated. Therefore, the use of LBB to eliminate the dynamic effects of postulated pipe breaks in these systems is appropriate.

I SARGENT & LUNDY, I

Prepared by:

- 1 % (4 /, hh"/L 05-12-89 I

G. H. DeBob Supervisor Engineering Mechanics Division I

Reviewed by:

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05-12-89 I

P. H. Hoang~

Senior Engineering Specialist Engineering Mechanics Division I

Approved by:

jf k 05-12-89 J. R. M i' ster Projec Manager Project Management & Engineering Division

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SANGENT & LUNDY SL-4518 05-12-89 1

CONTENTS Section Page ES EXECUTIVE

SUMMARY

ES-1 1

SCOPE-1-1 Systems Within Scope 1-1 l

Reactor Coolant Bypass System - RC Bypass 1-1 Residual Heat Removal System - RHR l i Safety Injection System - SI 1-2 Piping Within Scope 1-2 l

2 SYSTEM REVIEW VERSUS NUREG-1061 LIMITATIONS 2-1 j

2-2 Design Criteria Susceptibility to Failure 2-2 I_

Stress Corrosion 2-2 Water / Steam Hammer 2-5 Thermal and/or Mechanical Cyclic Fatigue 2-6 Erosion / Corrosion 2-8 2-8 Creep 2-8 Component and Piping Supports Degradation or Failure from Indirect Causes 2-8 IE Bulletin 79-14 Verification 2-9 Cleavage-Type Fracture Susceptibility 2-9 I

3 MATERIAL PROPERTY DATA 3-1 Installed Material 3-1 I

LBB Material Properties 3-2 Test Data Curve Fitting 3-4 4

APPLIED LOADING 4-1 4-1 High Stress Locations 4-2 l

Applied Loads st-4518 I

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Nuclear Safety-Related I

Leak-Before-Break Evaluation l

for Stainless Steel Piping Byron and Braidwood Nuclear Power Stations Units 1 and 2 g

Prepared for I

Commonwealth Edison Company Report SL-4518 May 12,1989 I

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SL-4518 SARGENT S LUNDY 05-12-89 I

i CONTENTS, Cont.

Section Pg l

5 LEAK DETECTION 5-1 Limiting Condition for Operation 5-1 Major Leak Detection System Components 3-2 Component Design 5-3 Containment Cavity and Floor Drain Sumps 5-3 Containment Particulate and Gaseous Radioactive Monitoring Systems 5-3 Conclusions 5-4 6

LEAK RATE CRACK SIZING 6-1 7

CRACK STABILITY ANALYSIS 7-1 Master Curve Construction 7-2 Master Curve Methodology 7-4 I

Master Curve Results 7-5 J-T Crack Stability Evaluation 7-5 7-6 Margin for Axial Cracks 8

REFERENCES 8-1 I

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TABLES Table l

1-1 Byron /Braidwood Leak-Before-Break Line Data 1-2 Byron /Braidwood Leak-Before-Break Pipe Whip Restraints 3-1 Byron /Braidwood LBB Installed Material

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3-2 Byron /Braidwood LBB Weld Processes 3-3 Installed Base Metal Tensile Property Boundaries at Room Temperature 3-4 Installed Weld Metal Tensile Property Boundaries at Room Temperature l

3-5 Material Test Data in PIFRAC 3-6 PIFRAC Base Metal Tensile Properties at Room Temperature 3-7 PIFRAC Cast Stainless Steel Tensile Properties at Operating Temperature i

3-8 PICEP Material Properties i

3-9 FLET Tensile Material Properties 3-10 FLET Fracture Toughness Material Properties 4-1 RHR System High-Stress Point Loads 4-2 51 System High-Stress Point Loads 4-3 RC Bypass System High-Stress Point Loads 1 1 Leakage Size Crack Length 7-1 Circumferential Crack Size Margins 7-2 Circumferential Crack Load Margins 7-3 3-T Results for Cast Stainless Steel I

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SARGENT& LUNDY SL-4518 05-12-89 IV TABLES, Cont.

Tables

.I 7-4 Axlal Crack Size Margins j

7-5 Axial Crack Load Margins I

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FIGURES

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Figure 1-1 Residual Heat Removal Subsystem RH02 Loop 1 1-2 Residual Heat Removal Subsystem RH02 Loop 3 1-3 Accumulator Safety Injection Subsystem 5101 1-4 Accumulator Safety injection Subsystem 5104 l

l-5 Accumulator Safety injection Subsystem 5109 l-6 Accumu!ator Safety Injection Subsystem 5103 1-7 RC Bypass Subsystem RC01 1-8 RC Bypass Subsystem RCO2 1-9 RC Bypass Subsystem RC03 RC Bypass Subsystem RC04 1-10 3-1 Stainless Steel True Stress-Strain Base Metal 3-2 Cast Stainless Steel True Stress-Strain 3-3 Cast Stainless Steel 3-R PIFRAC 3-4 Stainless Steel Base Metal Mean Stress-Strain Curve Fit 3-5 Cast Stainless Steel Mean Stress-Strain Curve Fit 3-6 Cast Stainless Steel Lower Bound Stress-Strain Curve Fit j

3-7 Cast Stainless Steel PIFRAC 3-R Curve Fit 6-1 Leak - RH Crack Length Curve (Stainless Steel) 6-2 Leak - SI(t=1.00 in) Crack Length Curve (Stainless Steel) 6-3 Leak - S!(t=1.00 in) Crack Length Curve (Cast Stainless Steel) 6-4 Leak - Sl(t=0.365 in) Crack Length Curve (Stainless Steel) sL-4518 j

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i FIGURES, Cont.

l Figure 6-5 Leak - RC Bypass Crack Length Curve (Stainless Steel) i 7-1 Master Curve (RH System) Base Material Crack Margin 7-2 Master Curve (RH System) SAW Weld Crack Margin 7-3 Master Curve (RH System) Base Material Load Margin 7-4 Master Curve (RH System) SAW Weld Load Margin l

7-5 Master Curve (SI System) Base Material Crack Margin (t = 1.00 in) 7-6 Master Curve (S! System) SAW Weld Crack Margin (t = 1.00 in) 7-7 Master Curve (51 System) Cast Stainless Steel Crack Margin (t=1.00 in) 7-8 Master Curve (S1 System) Base Material Load Margin (t = 1.00 in) 7-9 Master Curve (Si System) SAW Weld Load Margin (t = 1.00 in) 7-10 Master Curve (51 System) Cast Stainless Steel Load Margin (t=1.00 in) 7-11 Master Curve (Si System) Base Material Crack Margin (t = 0.365 in) 7-12 Master Curve (SI System) SAW Weld Crack Margin (t = 0.365 in)

I 7-13 Master Curve (S1 System) Base Material Load Margin (t = 0.365 in) 7-14 Master Curve (Si System) SAW Weld Load Margin (t r 0.365 in) 7-15 Master Curve (RC Bypass) Base Material Crack Margin 7-16 Master Curve (RC Bypass) SAW Weld Crack Margin 7-17 Master Curve (RC Bypass) Base Material Load Margin l

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7-18 Master Curve (RC Bypass) SAW Weld Load Margin I

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1 Figure 7-19 3-T Curve for PIFRAC Cast Stainless Steel Crack Margin 7-20 3-T Curve for PIFRAC Cast Stainless Steel Load Margin I

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. SARGENTS LUNDY SL-4518 05-12-89 1-1

. Project 812F29 Section 1 SCOPE The leak-before-break (LBB) program for the Byron and Braidwood stations includes several

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high-energy piping systems that have postulated high-energy line break locations in the current design. Portions of three piping systems inside the containment building are in the program.

SYSTEMS WITHIN SCOPE A brief description of the three systems are as fo!!ows:

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l Reactor Coolant Bypass System - RC Bypass The Byron /Braidwood reactor coolant system is a four-loop system with one pump and steam generator per loop. The pressurizer is connected to loop 4. Each loop contains two loop-stop valves: one between the RC pump and the reactor and one between the reactor and the steam generator.

The LBB analysis for the Byron /Braidwood reactor coolant loop has been performed by Westinghouse and accepted by the NRC. The RC Bypass system piping is being evaluated under the Byron /Braidwood LBB program.

The RC loop bypass lines, which run between the hot and cold legs of each loop, were designed to provide a recirculation path to allow a loop isolated from the reactor to be brought back into service while the plant is operating. The Byron /Braidwood units are not licensed for "N-1" operation and thus this loop bypass line is not used.

Residual Heat Removal System - RHR The residual heat removal (RHR) system removes heat from the reactor coolant system and transfers it to the component cooling water (CC) system. Heat removal occurs during the I

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second phase of normal plant cooldown when heat removal by the main steam system (first phase of normal cooldown) is no longer practical.

The RHR system is also used during the postulated loss-of-coolant accident (LOCA). At this time it acts as part of the emergency core cooling system (ECCS) by providing a low-pressure, high-volume water source to the reactor coolant system during the injection and recirculation phases. The high pressure portion of the RHR piping included in the Byron /

Braidwood LBB program connects to the hot legs of loops 1 and 3. This piping provides suction water to the residual heat removal pumps during shutdown cooling.

Safety injection System - SI The piping that provides emergency core cooling inlet flow is included in the safety injection l

system. This system injects borated water into the reactor coolant system in the event of the postulated LOCA. The SI system piping consists of injection lines to both the cold and j

hot reactor coolant legs. The cold leg injection piping consists of a high-pressure (CV) header to each of the cold legs, a medium-pressure line (SI accumulator) to each cold leg, and a low-J pressure line (RH) tying to each medium-pressure line. The hot leg injection piping consists of a medium-pressure line (SI) to RC loops 1 and 3 and a low-pressure line (RH) tying to each medium-pressure line. This piping provides coolant injection for a LO' A throughout the C

pressure range of the reactor coolant system. The medium-pressure Si lines to each cold leg (Si accumulator lines) are included in the LBB program.

PIPING WITHIN SCOPE I!

Isometric drawings of the piping systems in this scope are shown in Figures 1-1 through 1-10.

l These isometric drawings represent the as-analyzed pipe routings for the four Byron /

l Braidwood units. Minor differences in routing between the units and the actual ploe support I

locations have been included in the detailed piping analysis for the four units. Each isometric drawing identifies pipe whip restraints, pipe break locations, line number, valve number and analytical boundary (anchor).

The last numeric characters in the line number give the nominal pipe outside diameter.

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Table 1-1 provides the piping materials, the analytical model (subsystem designation, line sizes, normal operating fluid parameters) for this scope.

1 Table 1-2 presents the pipe whip restraints on the piping that was evaluated as a part of the l

Byron /Braidwood LBB program. The table includes the appropriate loop number, line number, and whip restraint number.

I There are a total of 92 postulated break locations in the piping within the scope of the LBB program and a total of 64 pipe whip restraints.

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SARGENT & LUNDY Byron /Breidwood Leak-Tabla 1-1 Before-Break Line Data SL-4518 03-12-89 I

1 Operating I

Temp ('erature OD Tr PDT Pressure F)

Line No.

Subsystem (in)

(in)

Material (psla)

RC04AA-12 RH02 12.75 1.125 SA376 TP316 2287 618 RH01AA-12 RH02 12.75 1.125 SA376 TP316 2287 618 RC04AB-12 RH02 12.75 1.125 SA376 TP316 2287 618 RH01AB-12 RH02 12.75 1.125 SA376 TP316 2287 618 RH01BA-12 RH02 12.75 0.406 SA312 TP304 450 350 RH01BB-12 RH02 12.75 0.406 SA312 TP304 450 350 SIO9BA-10 S101 10.75 1.000 SA376 TP316 700 350 RC29AA-10 S101-10.75 1.000 SA376 TP316 2332 557 RC29AD-10 S103 10.75 1.000 SA376 TP316 2332 557 I

5109BD-10 S103 10.75 1.000 SA376 TP316 700 350 RC29AB-10 5104 10.75 1.000 SA376 TP316 2332 557 SIO9BB-10 5104 10.75 1.000 SA376 TP316 700 350 I

RC29AC-10 S109 10.75 1.000 SA376 TP316 2332 557 S109BC-10 5109 10.75 1.000 SA376 TP316 700 350 5109AA-10 510 1 10.75 0.365 SA312 TP304 700 120 SIO9AB-10 S104 10.75 0.365 SA312 TP304 700 120 I

S109AC-10 5109 10.75 0.365 SA312 TP304 700 120 S109AD-10 S103 10.75 0.365 SA312 TP304 700 120 RC21AA-8 RC01 8.625 0.906 SA376 TP304 2332 557 SA312 TP304 RC21BA-8 RC01 8.625 0.906 SA376 TP304 2287 618 SA312 TP304 I

RC21AB-8 RCO2 8.625 0.906 SA376 TP304 2332 557 SA312 TP304 RC21BB-8 RCO2 8.625 0.906 SA376 TP304 2287 618 I-SA312 TP304 RC21AC-8 RC03 8.625 0.906 SA376 TP304 2332 557 SA312 TP304 RC21BC-8 RC03 8.625 0.906 SA376 TP304 2287 618 I

SA312 TP304 RC21 AD-8 RC04 8.625 0.906 SA376 TP304 2332 557 SA312 TP304 I

RC21BD-8 RC04 8.625 0.906 SA376 TP304 2287 618 SA312 TP304 I

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SARGENT& LUNDY Byron /Brcidwood Leak-Table 1-2 Before-Break Pipe Whip SL-4518 Restraints 05-12-89 I

I-Pipe Whip Restraint Subsystem Line No.

Loop No.

RCl-1 RC01 RC21AA8 1

RCl-4 RC01 RC21AA8 1

RC2-1 RC02 RC21AB8 2

RC2-4 RCO2 RC21AB8 2

RC3-1 RC03 RC21AC8 3

RC3-4 RC03 RC21AC8 3

i RC4-1 RC04 RC21AD8 4

RC4-4 RC04 RC21AD8 4

SilR-10B S101 RC29AA10 1

SilR-30 S101 5109BA10 1

514R-15B S104 RC29AB10 2

SI4R-35 S104 5109BB10 2

S19R-475B S109 RC29AC10 3

S19R-495 SIO9 S109BC10 3

513R-640A S103 RC29AD10 4

S13R-655 S103 S109BD10 4

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I SARSEI;TELUNDY SL-4518 0F12-89 2-1 g

-l Section 2 SYSTEM REVIEW VERSUS NUREG-1061 LIMITATIONS l

'l NURE G-1061 provides recommendations for application of the LBB approach in the NRC I

licensing process. In Volume 3 of NUREG-1061, Section 5.1 provides six limitations to be applied to the mechanistic evaluation of pipe breaks in high-energy fluid system piping:

For specifying design criteria for emergency core coolant systems, contain-ments, and other engineered safety features, loss of coolant shall be assumed in accordance with existing regulations. The evaluation of environmental I

effects should be considered case-by-case.

The LBB approach should not be considered applicable to high-energy fluid system piping, or portions thereof, that operating experience has indicated particular susceptibility to failure from the effects of corrosion (e.g., inter-ranular stress corrosion cracking), water hammer, or low-and high-cycle 3

i.e., thermal, mechanical) fatigue.

Component and piping support structural integrity should be maintained with no reduction in margin for the final safety analysis report (FSAR) loading I

combination that governs their design.

The LBB approach should not be considered applicable if there is a high probability of degradation or failure of the piping from more indirect causes I

such as fires, inissiles, and damage from equipment failures (e.g., cranes),

and failures of systems or components in close proximity.

The LBB approach should not be considered applicable to high-energy piping, j

or portions thereof, for which verification has not been provided that the requirements of I&E Bulletin 79-14 have been met.

The LBB approach described in this report is limited in application to piping systems where the material is not susceptible to cleavage-type fracture over the full range of system operating temperatures where pipe rupture could I

have significant adverse consequences.

The six limitations are addressed relative to the Byron /Braidwood units in following E

l 5 subsections.

These subsections demonstrate that these limitations do not apply to the l

systems addressed in this report.

I sL-4M B LI L__________

SARGENTE LUNDY SIA 518 05-12-89 2-2 DESIGN CRITERIA The containment was designed to accommodate a LOCA resulting from breaks in the reactor coolant pressure boundary up to and including a break equivalent in size to the double-ended rupture of the largest pipe in the reactor coolant system. It is not proposed to use the LBB results to modify these design criteria. Also, the functional design for the emergency core cooling system is unchanged from the original design and remains a nonmechanistic pipe rupture. Application of LBB for environmental qualification of electrical and mechanical equipment is not considered in the scope of this report but may be considered on a case-by-case basis at a later date.

I SUSCEPTIBILITY TO FAILURE This section addresses susceptibility of the systems to failure from the effects of stress l

corrosion, water / steam hammer, thermal and mechanical cyclic fatigue, erosion / corrosion, and creep.

Stress Corrosion The susceptibility of the systems addressed in this report to stress corrosion cracking has been reviewed, and it has been determined that the installed piping is not, susceptible for the following reasons.

The piping materials used for the RC,51, and RH systems are ASME SA-376 type 304 or 316 or SA-312 type 304 austenitic stainless steel. Unstabilized austenitic stainless steels when used in other than the solution-annealed condition are suoject to stress corrosion cracking provided that these three conditions are present simultaneously (Reference NUREG-0313, Revision 2):

1 an aggressive environment, e.g., an acidic aqueous medium containing chlorides or oxygen; 1

sensitized steel; and high stress, 3

s<-.,.

I

SARGENT 51. UNDY SL-4518 05-12-89 2-3 1

I

'I I

If any one of the three conditions described above is not present, intergranular stress J

corrosion attack is not considered likely.

I To prevent intergranular stress corrosion attack of austenitic stainless steel piping and I

components the following programs have provided at Byron and Braidwood:

Control of Primary Water Chemistry to Ensure a Benign Environment. The water chemistry is the primary control in preventing an aggressive environ-I ment for the subject stainless steel in the reactor coolant system and auxiliary lines that are connected to it.

In particular, the oxygen and chloride concentrations are controlled below the maximum permissible level

-I specified in Technical Specification 3.4.7. The Technical Specification limits are as follows:

Steady-State Transient Dissolved oxygen s 0.10 ppm s 1.0 ppm Chloride s 0.15 ppm s 1.5 ppm The control of oxygen content is accomplished by using hydrazine to limit the initial oxygen content to the Technical Specification steady-state limit and I

by using a 15-psig hydrogen blanket on the volume control tank. Chloride content was controlled during fabrication, shipping, and storage by the precautions taken to prevent the intrusion of chlorides into the system as specified in Article 115.2 of the Piping Installation Specification. During I

operation the primary water (reactor coolant system makeup water) is maintained at a similar chemistry content as the Technical Specification steady-state limits to prevent contamination of the reactor coolant system.

The stations' chemistry control programs restrict the use of noncategorized chemicals in and on stainless steel piping exposed to temperatures greater I

than 200*F. These programs significantly limit the amounts of chlorine, fluorine, bromine, sulfer, and heavy metals used in and on stainless steel systems.

Control of Pipe Manufacture and Fabrication. Control was provided to avoid the use of sensitized stainless steels and to prevent sensitivity during fabrication.

I.

Material selection for piping and fittings is controlled by specifying the materials per ASME Section II, Material Specifications.

The material specifications utilized are ASME SA-376 type 304 or 316, or SA-312 type 304 g

s <..,.

I

i SARGENT E LUNDY -

SL-4518 05-12-89 2-4 p

stainless steel. All stainless steels were supplied in the solution annealed El condition in accordance with Article 3.3.13.a of Standard Form 278-A.

E No bends were used in the piping within the scope of this report. The only 3

cast stainless steel fittings used for piping within the scope of the LBB g,

program are in the reactor coolant loop connections of the SI Accumulator system. Forged seamless fittings are the fittingt used for all other piping in this program. All forged. fittings (elbows and tees) are wrought stainless steel in accordance with ASME SA-403, Gr. WP304. Required radiographic, ultrasonic, or liquid penetrant examinations were performed on these fittings by the manufacturer in accordance with ASME Section Ill, NB-2550 for Class 3

1 and NC-2550 for Class 2 seamless materials to ensure material soundness.

3 The SI accumulator nozzle to the reactor coolant piping was fabricated to ASME SA-351 CF8A specifications.

Stainless steel castings are not susceptible to intergranular stress corrosion cracking (IGSCC) because they

)

are fast-cooled resulting in significant delta ferrite content. There have been no reports of IGSCC in weld metal or castings from oxygenated water.

The sensitivity of stainless steel can increase during the fabrication and installation process (as a result of bending, cutting, and welding). In order to prevent increased sensitivity during fabrication and installation, several measures are taken.

Control of welding ) processes and procedures to avoid increasing heat affected zone (HAZ sensitivity: All weld procedure qualification tests fcr ASME Section III austenitic stainless steel welds included an inter-granular corrosion test as specified in Article 303.7 of the Piping j

installation Specification. HAZ sensitivity is controlled by limiting the 3 J weld interpass temperature to 300'F maximum and weld heat input to El 50,000 Joules per inch maximum as required in the Piping Fabrication and Installation Specifications.

I Control of heat treatment processes: Materials are used in the solution annealed condition and subsequent heat treatments in the 800'F to 1500'F temperature range are prohibited (Reference Standard Form E

278-A, Section 3.3.13).

5 i

Control of cutting processes: Flame cutting of stainless steel pipe is e

prohibited as specified in Articles 303.6c of the Byron /Braidwood Piping g

Fabrication and Installation Specifications.

Control of cleanliness: Pipe inner and outer wall cleanliness during fabrication, shipment, storage, and installation are controlled in accordance with the recommendations of ANSI 45.2.1, 45.2.2, and 45.2.3, I'

l l

l st-4518

SARGENTS LUNDY SL-4518 05-12-89 2-5 l

.I l3 which were specified in Article 4.1 of Standard Form 278-A.

The

,E metallic thermal insulation installed on the piping will help to prevent surface contamination during operation.

I Water / Steam Hammer Each system in the Byron /Braidwood LBB scope has been reviewed to determine its potential for water / steam hammer events to occur. As described in~ NUREG-0582 (Reference 1),

water / steam hammer events may be caused by the fo!!owing phenomena: flow into empty or I

voided lines; rapid valve opening, closing, or instability; water entrainment in steam lines; column separation; steam bubble collapse; slug impact due to rapid condensation; and pump startup, stopping, or seizure. System designs and operating procedures combined with plant operating experience as well as industry experience formed the basis of this evaluation.

NUREG/CR-2781 (Reference 2) evaluated 40 reported water hammer incidents in pressurized water reactors. The evaluations for the systems in this program demonstrated that although the systems were susceptible to the generic causes of water / steam hammer, water hammer l

events were improbable and the severity of the reported incidents was low. It should be noted that all reported water / steam hammer events resulted in only support damage. There I

were no events in PWRs that resulted in a loss of pressure boundary integrity. Additional I

studies based on industry experience have reached the same conclusions and have provided recommendations for mitigating the causes of water / steam hammer in these systems (NUREG-0927, Reference 3).

I The following subsections discuss the potential for water / steam hammer in each system and provides details of the operating procedures used to minimize this potential.

Reactor Coolant Bypass System. As identified in the NUREG 2781 (Reference 2) the only I

water hammer events to occur in the reactor coolant system (RCS) involved the pressurizer relief system. The design and operation of this system eliminates the possibility of void formation in the RCS except in the pressurizer. The effects of a water / steam hammer in the pressurizer relief system on the RCS and in particular on the bypass line are isolated by the pressurizer. Other causes of water hammer in the RCS such as pump startup or trip are I

' ' - > =

I o

SL-4518 SARGENT E LUNDY 05-12-89 2-6 I

I l

insignificant, and stop valve opening and closure during system operation is not possible.

Therefore, there will not be flow in the bypass line to induce a water hammer.

Safety Injection (SI) System. NUREG 2781 identified four water hammer events involving the SI system. Three of the four events were caused by voids that were the result of improper venting while the lines were being filled. The fourth event was caused by steam bubble col-lapse in the 51 accumulator line. The steam bubble formation occurred during leak testing as a result of inadequate testing procedures that allowed the line pressure to drop below the I

saturation pressure. Procedures for initial fill and venting ensure voids will not occur in the safety injection piping. High point vents in the SI piping are provided for proper venting of g

lines and pumps. The head of water provided by the refueling water storage tank (RWST) and E

the low temperature of the SI system further ensures the lines will remain full and steam bubbles will not develop. Valve operating and pump startup or trip times have been deter-mined to produce negligible effects on 51 system.

Residual Heat Removal (RHR) System. NUREG 2781 (Reference 2) identified a single water hammer event in the RHR system. The probable cause was reported as pump startup into a voided line during a refueling shutdown.

The elimination of voids in the RHR system is achieved through required venting procedures during initial line fill. During normal plant operation parts of the RHR are pressurized by the RC and 51 system and are at building ambient or low temperature, therefore a reduction in pressure will not result in the formation of a steam bubble. Valve opening and closing times as well as pump startup or trip times are determined to be of sufficient duration to produce negligible effects on the portion of the RHR system in the LBB scope.

Thermal and/or Mechanical Cyclic Fatigue For all ASME Class 1 piping within the scope of the LBB program, normal and upset thermal and seismic fatigue was evaluated as part of the ASME piping stress analysis. Unidentified operating conditions in the RHR,51, and RC Bypass systems that might contribute to thermal I

s<--

l I

l

___.__._-______a

SARGENT& LUNDY SL-4518 0F12-89 2-7 I

1 l

l or mechanical fatigue are thermal stratification and system operating or flow induced vibration.

E Thermal stratification usually occurs in horizontal pipe segments when fluid at a significantly different temperature than the fluid in the piping is introduced at low flow velocities. When the fluid's bouyancy forces are greater then the inertial forces acting on the fluid, the flow tends to separate into hot and cold layers. Thermal stratification has been identified in NRC Notice 84-87, " Piping Thermal Deflection Induced by Stratified Flow," and NRC Bulletins 79-13, " Cracking in Feedwater System Piping;" 88-08, " Thermal Stresses in Piping Connected to Reactor Coolant Systems;" and most recently, 88-11, " Pressurizer Surge Line Thermal I

Stratification." The flow stratification or uneven thermal mixing identified in these docu-ments is not postulated to occur in the Byron /Braidwood LBB systems. Also previous industry I

operating experience has not identified thermal stratification in these systems. Each system in the Byron /Braidwood LBB scope has been reviewed for its susceptibility for thermally stra-tified flow. This review has determined that thermal stratification is improbable for these systems. As required by NRC Bulletin No. 88-08, including Supplements 1 and 2, a review of piping systems connected to the RCS was performed. The review (Reference 4) determined that the Byron /Braidwood LBB systems were not susceptible to thermal stratification or temperature oscillations induced by leaking valves.

System operating flow induced vibration in the Byron /Braidwood LBB systems was monitored during the preoperational and startup vibration testing at Byron Units 1 & 2 and Braidwood Units 1 & 2.

These systems were tested for several normal modes of operation, and the measured vibration levels were found to be within acceptance levels consistent with ANSI /

ASME OM-3 in all cases. Since positive displacement pumps can produce flow induced vibra-tions, it should be noted that positive displacement pumps are not used by the Byron /Braidwood LBB systems.

I I

s'- -

I

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

J

SARGENTE LUNDY SL-4518 05-12-89 2-8 I

Erosion / Corrosion Use of high-quality stainless steel for RC Bypass, RHR, and 51 systems will prevent erosion / corrosion problems. Control of water chemistry per Technical Specification 3.4.7 l

will provide an additional protection from erosion / corrosion for the systems connected to the RCS.

Creep Creep and creep-fatigue is not a concern since all operating temperatures are well below the temperature of 800*F for austenitic stainless steel, at which creep and creep fatigue becomes a concern.

COMPOf4ENT AND PIPING SUPPORTS All eqaipment, piping and supports within the scope of Byron /Braidwood LBB program have been seismically designed per the rules of ASME Section III. Performance of the LBB analysis does not reduce the margin of safety of any supports.

DEGRADATION OR FAILURE FROM INDIRECT CAUSES Application of LBB on the pipe lines in this scope would not increase probability of degra-dation or failure of the piping from indirect causes. Effects on piping systems and compon-ents from the indirect causes such as fires, missiles, and damage from equipment failures, and failures of systems or components in close proximity were evaluated and dispositioned during the design of the Byron /Braidwood stations (see the Byron /Braidwood UFSAR).

I All piping, components, and supports inside containment are designed seismically to prevent failures of safety-related piping and components as a result of failures of non-safety-related items. Byron /Braidwood Technical Specification 3/4.7.8 provides the surveillance require-ments for safety-related snubbers.

Adherence to the requirements of this Technical j

Specification will ensure that the installed snubbers are adequately maintained and will perform their intended functions.

sL-4518 ll

I SARGENT& LUNDY SL-4518 05-12-89 2-9 I

I IE BULLETIN 79-14 VERIFICATION Safety-related piping systems at Byron /Braidwood have been verified to the requirements of I&E Bulletin 79-14 during construction of the stations. Subsequent to 79-14 verification, modifications to safety-related piping systems are reviewed to verify that dimensional verifi-cation is applicable. If dimensional verification is applicable, it is specified in the engi-neering modification approval letter.

CLEAVAGE-TYPE FRACTURE SUSCEPTIBILITY The installed piping systems listed in Table 1-1 are exposed to the normal system operating temperatures ranging from 120*F to 618*F for the RC Bypass, RH, and 51 systems.

Cleavage-type fracture is not a concern for the piping material at these operating temperatures.

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g SARGENTELUN0Y SL-4518 05-12-89 3-1 I

Section 3 MATERIAL PROPERTY DATA I

This section identifies the material properties of the piping, piping components, and welds installed on the Byron /Braidwood LBB systems. It presents the tensile properties and ductile fracture toughness data obtained from the NRC's Piping Fracture Mechanics Data Base (PIFRAC) (Reference 5), which most represents the properties of the installed materials.

The ranges of the CMTR tensile properties for the installed materiais are enveloped and compared to the tensile properties for identical and similar materials from PIFRAC. The PIFRAC materials having similar ranges of tensile properties are used in this evaluation. The true stress-true strain data of the materials chosen from PIFRAC are compared to establish I

the mean and lower-bound true stress-true strain data for sizing the leakage crack and for evaluating crack stability, respectively. The chosen PIFRAC 3-Resistance (3-R) data are compared to establish the most limiting material toughness for use in the crack stability evaluation. Finally, this section describes the methodology used to calculate the Ramberg-Osgood coefficients and exponents for the mean and lower-bound true stress-true strain data and the coefficients and exponents for the bifunctional curve fit of the lower bound fracture resistance (3-R) data.

1-j INSTALLED MATERIAL The installed materials on the Byron /Braidwood I.BB systems were identified by obtaining their certified material test reports (CMTRs). From these CMTRs the material specifica-g E

tions and types were obtained along with the ultimate tensile strength (UTS) and the yield strength (YS) of each heat. A summary of the material types and specificat' ions for the installed piping, fittings, valves, and nozzle safe ends is presented in Table 3-1. Most of the installed material is SA376 TP316, except for the 8-inch reactor coolant loop (RCL) bypass E

I

SARGENT& LUNDY SL-4518 05-12-89 3-2 4

I lines, which are SA376 TP304.

The only cast stainless steel, SA351 CF8A, in the Byron /Braidwood LBB scope is the SI accumulator nozzle to the reactor coolant loop piping.

Welding process and welding materials were obtained from the fabricator's and installer's weld process sheets. Table 3-2 presents a summary of the welding processes and filler metals used. Generally these welds were made using a gas tungsten arc weld (GTAW) process for the root passes with shielded metal arc weld (SMAW) and/or submerged arc weld (SAW) processes for the remainder of the weld. Therefore, the predominant weld metal is deposited using either a SMAW or SAW process for these welds.

For these lines, most of the welds by the piping fabricator are predominantly SAW types using ER-316L filler metal except for the RC Bypass welds, which are predominantly SAW types using ER308 filler metal. The field welds for these lines are predominantly SMAW types using E316-16 filler metal except for the RC Bypass field welds, which are pre-dominantly SMAW types using E308-16 filler metal.

The UTS and the YS for the installed material at room temperature were obtained from each of the CMTRs. A summary of the maximum and minimum values for these properties is pre-sented in Table 3-3 for each material category. This table presents the high and low boun-daries for the tensile properties. These same material properties are obtained for the installed weld filler metals and presented in Table 3-4 for each material category. Since the magnitude of embrittlement from thermal aging in cast stainless steel is related to high ferrite content, the ferrite content for the cast stainless steel nozzles was also obtained.

The ferrite content from the CMTR for each nozzle ranged from 12.4% to 17%

LBB MATERIAL PROPERTIES Room temperature UTS and YS were obtained for materials similar to those installed on the Byron /Braidwood LBB systerns from the stress-strain data given in NRC's PIFRAC data base. The matching and similar material specifications from the PIFRAC data base are listed in Table 3-5. It should be noted that specifications for fittings similar to those of the sL-4518 I

o

Ea SARGENTa LUNDY SL-4318 03-12-89 3-3 installed SA234, Grade WPB fittings were not availaole in PIFRAC; therefore, based upon the similarities in chemical compositions and tensile properties, three different heats of SA333

" Specifications for Seamless and Welded Steel Pipe for Low Temperature Service" Grade 6 was chosen to represent the installed fittings. In addition, since PIFRAC did not contain specifications for type 316 stainless steel, SA358 type 316L was selected to represent the installed type 316 stainless steel components. The only SA351 cast stainless steel provided in PIFRAC that was thermally aged was SA351 CF3 in Heat No. 8.

This test specimen was thermally aged at 662*F for 10,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />. The cast stainless steel in Heat No. 7 is from the same casting as Heat No. 8; however, it was not thermally aged. The ferrite content of these cast stainless steel heats is 17.7%, which is greater than the highest ferrite content of the installed nozzles. The tensile properties at room temperature for test specimens from these heats are presented in Table 3-6 for base metals. The comparison of these tensile properties I

to those of the installed materials establishes a close correspondence between their upper and lower bounds.

Based on this correspondence, the PIFRAC material properties are deemed to be representative of the Byron /Braidwood LBB system material properties.

True stress-true strain data for test specimens from the heat numbers presented in Table 3-5 have been plotted in Figures 3-1 for the stainless steel base metals and in Figure 3-2 for cast stainless steel. To coincide with the operating temperatures of the Byron /Braidwood LBB systems, the chosen stainless steel test specimens were tested at 640*F,550'F, and 70'F.

Since the size of the leakage crack could be underestimated by using the lowest stress-strain I

data, the mean stress-strain data are used to conservatively calculate the leakage size crack (LSC). By examining the stress-strain plots in Figures 3-1 and 3-2, the mean stress-strain data are determined to be irom test specimen A5-5 of Heat No. 34 for stainless steel base metal and test specimen Il-29L from Heat No. 8 for cast stainless steel. Lower bound stress-strain data are required to conservatively represent the material tensile properties in the crack stability evaluation. The lower bound stress-strain data are determined to be from test specimen A23-1 of Heat No. 37 for stainless steel base metal and test specimen 12-20L of j

Heat No. 8 for cast stainless steel.

g s<..s.

I

SARGENT 5 LUNDY SL-4518 05-12-89 3-4 The required tensile property for a modified limit load evaluation is the flow stress and is defined as the average of the yield and ultimate tensile stress at operating temperature. The required flow stress was obtained from the previously identified lower bound test specimen g,

for stainless steel base metals. The flow stress for stainless steel SMAW and SAW was E

defined as 51000 psi as required by (Reference 6). From Table 3-7 the lower bound flow stress for cast stainless steel was determined to be 44,406 psi from test specimen 12-20L of Heat No. 8.

Deformation 3 and aa data for test specimen from the heats presented in Table 3-5 are plotted on Figure 3-3 for cast stainless steel. The chosen cast stainless steel test specimens were tested at 550*F. Also, only test specimens with a thickness equal to or greater than the pipe wall thickness were chosen. The lower bound 3-R data were determined to be from test specimen Il-5LC-ID of Heat No. 8.

TEST DATA CURVE FITTING The EPRI Pipe Crack Evaluation Program (PICEP) used to determine the LSC and the FLET program used to evaluate crack stability require the true stress-true strain data to be curve fitted by the Ramberg-Osgood relation.

n

• -. =L+a(L (3-1)

O o

O where a is the 0.2% offset yield stress, the yield strain co = c /E, and the Ramberg-Osgood o

o coefficient, n, and exponent, n.

Recent studies have determined that the curve fit in low strain ranges (i.e., less than 1% strain) is preferred.

Therefore, as recommended in Reference 8, the Ramberg-Osgood coefficient and exponent were determined from the chosen test specimen true stress-true strain data in the yield strain to 1% strain range. The Ramberg-Osgood constants and material tensile properties for use in PICEP are presented in Table 3-8, and for use in FLET in Table 3-9. Figures 3-4 and 3-5 present the plots of the j

stress-strain test data with the Ramberg-Osgood curve fit for the chosen mean stainless steel base metal and cast stainless steel test specimen, respectively. Figure 3-6 presents the plots st.-4518 1

I SARGENT & LUNDY sL-4518 j

05-12-89 3-5 I

of the stress-strain test data with the Ramberg-Osgood curve fit for the chosen lower-bound cast stainless steel.

I The FLET program uses a bifunctional curve fit for the fracture toughness test data. The bifunctional curve fit form is J=c (aa)*, aa < aa (3-2) transition E

J=ag+a2e(- / 3), aa > aa (3-3) transition I

where 3 is the deformation 3 integral parameters, aa is the crack extension and oatransition is the crack extension at which extrapolation of the test data is necessary to ensure J-controlled crack growth for large crack extensions. As specified in NUREG-1061 Volume 3, oatransition is defined as the aa at w = 5.

u is defined by I

w=hh (3-4)

I where b is defined as the uncracked ligament length of the test specimen and da is the crack increment. A 3-versus-tearing modulus (T) plot is used to determine the bifunctional curve fit constants. The tearing modulus is defined by T = hh (3-5)

"f where E is the Modulus of Elasticity and of s the flow stress. Using these relations the i

bifunctional curve fit constants are determined. The bifunctional curve fit constants and fracture toughness properties for use in FLET are presented in Table 3-10. Figure 3-7 shows I

the plots of the 3,aa test data with the bifunctional curve fit for the PIFRAC lower bound cast stainless steel.

g s<..sie

SARGENT & LUNDY Byron /Braidwood LBB Table 3-1 Installed Material SL-4518 I'

05-12-89 I

Stainless Steel Piping SA376 TP316 I

SA376 TP304 SA312 TP316 SA312 TP304 Fittings SA403 WP316 SA403 WP304 Valves '

SA182 F316N Nozzle Safe Ends I

SA182 F316N SA351 CF8A I

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SL-4516 I

- SANGENT & LUNDY Byron /Braidwood LBB Table 3-2 Weld Processes SL-4518 I

05-12-89 i

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i Stainless Steel Welds Weld Process Weld Metal GTAW ER308 ER308L SMAW E308-15 l

I E308-16 I

E309-16 E316 I

E316L E316L-16 E316-16 I

SAW ER308 ER316L I

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SL-4518 I

a-_

SARGENT & LUNDY-Installed Base Metal Tensile Table 3-3 Property Boundaries at Room SL-4518

. I Temperature 05-12-89 t -.

Material UTS (psi)

YS (psi)

L Stainless Steel:

i' l~ -

Maximum 94,300 56,572 Minimum 75,000 30,000 Cast Stainless Steel:

Maximum 85,050 42,050 I

Minimum 78,650 35,500 I

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$L-4518 I

1-4.

SARGENT & LUNDY Installed weld Metal Tensile Table 3-4 Property Boundaries at Room SL-4518 Temperature 05-12-89 I

Material UTS (psi)

YS (psi)

Stainless Steel:

j I

Maximum 92,183 71,139 Minimum 79,300 42,000 I

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l SARGENT E LUNDY Meterial Test Dsta in PIFRAC Table 3-5 SL-4518 I.

05-12-89 Material Heat Specification No.

Stainless Steel:

Base Metal SA182 TP304 3

I SA376 TP304 4

SA376 TP304 5

SA358 TP316L 34 SA376 TP304 37 I

SA376 TP304 39 Cast Metal SA351 CF3 7

SA351CF3 8

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L SARGENT& LUNDY PIFRAC Base Matal Tensil2 Tabla 3-6 i

Properties at Room Temperature SL-4518 I

05-12-89 Stainless Steel:

Heat No.

Spec. Id.

UTS (psi)

YS (psi) 3 ZP6-lC 70,304 29,661 3

ZP6-1L 71,648 30,127 i

.g 4

ZPl7-lC 88,721 36,564

' g

'4 ZPl7-IL 86,734 35,954 4

ZPl7-8C 86,734 36,467 5

ZP12-lC 84,399 34,915 I-5 ZP12-9L 85,695 27,491 5

ZP12-14L 86,054 34,413 34 AS-1 89,577 34,757

.3g 34 A5-2 74,638 26,4864g) 37 A23-1 76,540 28,784 43) 37 A23-2 112,582 57,164 39 A35-1 91,114 34,084 I

39 A35-2 70,577 25,342 Cast Stainless Steel:

Heat No.

Spec. Id.

UTS (psi)

YS (psi) 7 Il-IL 86,632 38,834 I

7 Il-2L 84,281 34,553 7

12-1L 83,987 36,153 7

12-2L 83,365 37,278 7

'13C-IL 74,912 34,626 I

8 Il-26L 88,875 40,389 8

Il-27L 93,381 44,717 8

12-19L 92,624 45,376 I

(1)

Extrapolated to 70*F since only specimen tested at 550*F are available.

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I SARGENTE LUNDY PIFRAC Cast Stainless Steel Tcble 3-7 Tensile Properties at SL-4518 I

Operating Temperature 05-12-89 Cast Stainless Steel:

Heat No.

Spec.Id.

UTS (psi)

YS (psi)

Flow (psi)

I 7

12-3L 59,043 24,486 41,765 7

12-6L 58,250 25,594 41,922 7

13C-2L 56,094 22,990 39,542

_I 8

11-28L (1)

(1)

(1) 8 11-29L 64,112 27,554 45,833 8

12-20L 63,279 25,533 44,406 Note (1) Test data are incomplete I

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SL-4518

SARGENT& LUNDY PICEP Mettrial Propertiss Tabla 3-8 SL-4518 I.

05-12-89 Mean Stainless Steel Base Metal:

i 1

1 Test Specimen: A5-5 Heat No. 34 21,830 psi a

=

n 0.000854 c

=

I o

2.809049 a

=

3.596975 n

=

6 25.55x10 psi E

=

67,966 psi 8

=

u 0.3 v

=

Mean Cast Stainless Steel:

Test Specimen: Il-29L Heat No. 8 27,554 psi o

=

o 0.001081 c

=

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=

7.861181 n

=

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=

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=

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SARGENT& LUNDY FLET Tensilm Mnterial Propertiss Tabla 3-9 SL-4518 05-12-89 l

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Lower Bound Cast Stainless Steel:

Test Specimen: 12-20L Heat No. 8 I

25,533 psi c

=

o 0.001001 c

=

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=

6.498142 n

=

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=

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"u

=

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=

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SARGENT5 LUNDY FLET Fracture Toughness Material Table 3-10 Properties SL-4518 05-12-89 I

I Lower Bound Cast Stainless Steel from PIFRAC:

Test Specimen: Il-5LC-ID Heat No. 8 2

f 2,026 in-lb/in J

=

IC 12,515.1 c

=

0.540934 m

=

7,825.833 ai

=

-6,699.651 a2

=

0.216832 a3

=

aatransition = 0.099536 I

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6 2

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0. =3 i(

0a 0

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=

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$8 O $,"

c M

5 ARGENT & LUNDY SL-4518 03-12-89 4-1 Section 4 APPLIED LOADING This section identifies the highest stress locations for each functional system in the Byron /Braidwood LBB program, and summarizes the procedure used to identify the locations, it specifies the loads and directional components to be obtained at these locations and the load combination methods.

The system thermal loads used in this evaluation correspond to the system thermal mode during normal plant operation.

HIGH STRESS LOCATIONS Since the highest stressed locations are to be considered with the most limiting material toughness properties, the highest stress location for each functional system is identified without consideration of the local material properties. The highest stress locations are considered to be coincident with the worst material toughness properties developed in Section 3.

The current piping stress analysis for the installed piping and support configuration of each Byron /Braidwood LBB system stress analysis is used to identify the highest stress point in it.

The high-stress locations were identified by first locating the points with the largest s

I resultant moment for the individual weight, normal operating thermal (NOT), and safe shutdown earthquake (SSE) loads. Then, at each of these points the axial force, the bending and torsional moments in local coordinates were obtained for the weight, NOT, SSE, and seismic anchor motion (SAM) loads. The highest stress point is determined by combining the weight, NOT, SSE, and SAM moments for each component direction. The weight and NOT loads are combined algebraically and then absolutely summed with the SSE and SAM loads.

The point with the largest combined resultant moment is the highest stress location.

I I

s-

1 SAR8ENT& LUNDY SL-4518 05-12-89 4-2 APPLIED LOADS The axial force, torsional and bending moment components from the high stress point for each Byron /Braidwood LBB system is presented in Tables 4-1 through 4-3.

An additional review considering only the piping bending moments was conducted. The highest stress location for each system remained the same with and without including the torsional moment.

I:

Il I

I l

I I

sL-4518

I.

SAN 8ENT &LUNDY SI System High-Stress Tabla 4-2 Point Loads SL-4518

(

'I 05-12-89 i

(kips)

(in-kips)

System Cond.

F5 Ma Mb Mc Mr SI10" Weight 1

-7 3

2 t = 1.0" Thermal

-16 213

-407

-880 n

SSE+ SAM 5

136 92 232 Norm Op.

-15 206

-404

-878 988 Faulted 20 342 496 1110 1263 SI10" Weight 1

0

-3 2

I t =0.365" Thermal

-5 2

67

-24 n

SSE+ SAM i

18 231 128 Norm Op.

-4 2

64

-22 68 Faulted 5

20 295 150 332 i

I I

(I I

)I g

I l

SL-4518

T..

l-SARGENT S LUNDY RHR System High-Stress Tabla 4-1 Point Loads SL-4518 I-.

05-12-89 (kips)

(in-kips)

System Cond.

Fa Ma Mb Mc Mr l

RHR 12 Weight 0

45

-4 369 t =1.125" Thermal 6

-24 161 111 n

SSE+ SAM 10 160 361 1054

}

Norm Op.

6 21 157 480 505 Faulted 16 181 518 1534 1629 I

I I

I I

I SL-4518

I l

SARGENTE LUNDY RC Bypass System High-Stress Tabla 4-3 l

Point Loads SL-4518 05-12-89 I

I (kips)

(in-kips)

System Cond.

Fa Ma Mb Mc Mr I

RC Bypass 8" Weight

-1 8

-12 106 t =0.906" Thermal 4

-67 100 303 n

SSE+ SAM 4

108 303 234 Norm Op.

3

-59 88 409 423 Faulted 7

166 391 643 771 I

I I

I I

\\I I

SL-4518 I

SARGENTH LUNDY SIA 518 05-12-89

~I-5-1 I

l I-Section 5 LEAK DETECTION i

I l

General Recommended Action 3 in the Executive Summary of NUREG-1061 Volume 3 (Reference 9) states:

I Leak detection systems in existing nuclear plants should be examined on a case-by-case basis to ensure that suitable detection margins exist so that the margin of detection for the largest postulated leakage size crack used in the fracture mechanics analyses is greater than a factor of ten on unidentified leakage.

I Licensees and applicants have the option of requesting a decrease in leakage margin provided they can confirm that their leakage detection systems are sufficiently reliable, redundant, diverse, and sensitive.

I i

This chapter addresses the Byron and Braidwood leakage detection system design for the purpose of demonstrating the acceptability of utilizing 1 gpm as the minimum detectable I

leak.

LIMITING CONDITION FOR OPERATION Technical Specification 3/4.6 addresses the reactor coolant system (RCS) leakage detection system. The Byron /Braidwood limiting condition for operation for RCS, leakage is specified in Technical Specification 3.6.2.b as 1 gpm unidentified leakage. Unidentified leakage is defined as all leakage that is not -

seal water flow supplied to the reactor coolant pump seals (controlled leakage);

leakage into closed systems, such as pump seal and valve packing leaks that are captured and conducted to a sump or collecting tank; leakage into the containment atmosphere from sources that are both specifically located and known either not to interfer with the operation of l

I.

leakage detection systems or not to be leakage through a non-isolable fault in an RCS component body, pipe wall or vessel wall; or RCS leakage through a steam generator to the secondary coolant system.

l I

sL-4518 I

SARGENT& LUNDY SL-4518 05-12-89 5-2 l

\\

4I i

I This LBB analysis is based on the minimum detectable through wall leakage for the applicable piping systems. Therefore this report uses the 1 gpm limiting condition for operation upper

)

limit for RCS leakage for the LBB analysis.

MAJOR LEAK DETECTION SYSTEM COMPONENTS The design of the reactor coolant leakage detection system defined in Section 5.2.5 of the Byron /Braidwood Updated Final Safety Analysis Report (UFSAR). The three primary systems for monitoring RCS leakage are the reactcr cavity and containment floor drain sumps, the containment atmosphere particulate radioactivity monitoring system, and the containment E!

gaseous radioactivity monitoring system. In addition to these three primary systems, there Ei are several other systems that would give an indication that leakage is occuring in containment. These systems are the containment area radiation monitors, the containment i

pressure indicator, and the reactor containment fan cooler inlet and outlet dew point and dry bulb temperature indicators. All of these except the area radiation monitors provide direct indication in the main control room.

Technical Specification 3.6.1 requires that the following reactor coolant leakage detection systems be operable:

the containment atmosphere particulate radioactivity monitoring system, the containment floor drain and reactor cavity flow monitoring system, and a

the containment gaseous radioactivity monitoring system.

The Technical Specification provides surveillance requirements to demonstrate that each of these leakage detection systems is operable and action statements to define the required response when one of these systems is determined to be inoperable.

Technical Specification Surveillance Requirement 4.6.2.1 requires that the containment atmosphere particulate radioactivity monitoring system, the containment gaseous radioactivity monitoring system, the reactor cavity sump discharge, and the containment floor drain sump discharge and inventory be monitored once every 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> to demonstrate st-4518

SARGENTE LUNDY SL-4518 05-12-89 I

5-3 4

I E

that unidentified RCS leakage is less than I gpm. In addition, an RCS water inventory balance must be performed once every 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />, COMPONENT DESIGN This section provides additional details concerning the design of the Technical Specification leakage detection components.

I Containment Cavity and Floor Drain Sumps Leakage that collects in the reactor floor drain sump and the reactor cavity sump is measured via weir boxes on these sumps with the level in the boxes converted to an equivalent leakage rate. This leakage is recorded in the main control room, and a high leakage alarm with a setpoint of I gpm is provided in the main control room.

I The accuracy of these loops is 0.63% or 0.063 gpm for the reactor cavity sump and 0.095 gpm for the floor drain sump. This calculated accuracy includes all components in the instrument loop. This accuracy ensures that the instrument loop is sufficiently sensitive to detect a leakage rate of I gpm.

The calibration frequency for these loops is 18 months. All instrumentation for the reactor cavity and floor drain leakage collection sumps is seismically designed and supported.

I Containment Particulate and Gaseous Radioactive Monitoring Systems i

These instruments are part of the microprocessor-based process radiation monitoring system. Containment particulate and gaseous samples are continuously analyzed. The main control room contains a CRT display, an operator's keyboard, and associated hardware that I

maintains an hourly record of radiation levels. When the set radiation levelis exceeded, the service, setpoint, and intensity level are displayed on the CRT. System failure also alarms in the main control room.

I l

SL-4518

SARGENT & LUNDY SL-4518 0F12-89 5-4 I

The accuracy of the system is 132% per channel. The calibration frequency for the detectors and microprocessor is 18 months. Instrumentation for this system is non-safety-related, seismically supported.

These monitoring systems will alarm and identify an abnormal level of gaseous radioactivity or radioactive particles in the containment atmosphere. Although the level of the reading will not be directly correlated to an RCS leakage rate, these monitors will indicate that g

further evaluation of leakage is required.

The reactor containment and cavity sump B

instruments or an RCS water inventory balance can then be used to quantify the leakage rate.

CONCLUSIONS Based on the design and Technical Specification requirements for the leakage detection equipment inside containment, a leakage rate of I gpm is identifiable for the Byron /Braidwood RC, RH, and 51 systems. Therefore a leakage rate of 10 gpm has been used for the LBB analysis of this report.

I Il I

l I!

l i

SARGENT& LUNDY SL-4518 05-12-89 6-1

~

}

f Section 6 LEAK RATE CRACK SIZING 1

I The Byron /Braidwood LBB program used the EPRI computer code PICEP (Pipe Crack Evaluation Program) (Reference 10) to establish the leakage size crack (LSC). The minimum credible detectable leak rate under the system normal operating conditions, established in Section 5, is adjusted by a factor of 10 to place a margin on that leak rate. PICEP calculates the leak rates for given crack sizes to form a crack-size-versus-leak-rate curve. The LSC is I

the crack length at a 10 gpm leak rate.

The PICEP code calculates the crack-opening area and the corresponding flow rate for non-i extending through-wall pipe cracks. The crack opening area is computed using the elastic-l I

plastic estimation scheme validated in Reference 10. The leak rate computational thermo-hydraulic models were validated against the available measured test results, in Appendix C of the same reference.

l The PICEP analyses utilized the normal operating loads established in Section 4 of this 2

I The dead weight, normal operating thermal and pressure loads yere combined l

report.

algebraically on a component basis.

l The PICEP code also required the material true stress-true strain property, at the corresponding normal operating temperature expressed in the RAMBERGOSGOOD format,

.to determine the crack opening area needed for the leak rate computations. The results of EPRI Report NP-5057 (Reference 7) indicated that the crack-opening area in the case of I

circumferential welds is governed by the base metal stress-strain properties. This would also be true for longitudinal welds since the crack opening deformations are expected to be I

influenced significantly by the base metal beyond the small width of weld region itself.

Generally, lower bound stress-strain properties would result in a larger crack opening area and consequently a higher leak rate. Accordingly, the mean true stress-true strain properties E

sL-4518 I

1

li SARGENT& LUNDY SL-4518 05-12-89 6-2 for the base rnetal, as determined in Section 3 of this report were used for a conservative estimation of the LSC.

I The results of the leakage rate calculations are presented in Figures 6-1 through 6-5 for circumferential and longitudinal crac s.

These figures present the leakage rates as a k

function of the crack size. The LSC is determined from these figures as the crack size allowing a 10 gpm leak rate. Table 6-1 presents the LSC length for each system in the g,

Byron /Braidwood LBB scope.

E I

I I

I I;

I I;

l:

I yl sL-4518

)

I)I

SAEGENT& LUNDY

!=J<ags Siza Crack Length Table 6-1 SL-4518 1

05-12-89 l

Circumferential Crack Orientation:

l Line Thickness LCS I

System Line Size (in)

(in)

RHR 12 1.125 3.652 51 10 1.00 3.450 SI (Cast) 10 1.000 3.901 SI 10 0.365 5.142 RC Bypass 8

0.906 3.6'19 I

Axial Crack Orientation:

RHR 12 1.125 4.946 SI 10 1.000 7.393 51 (Cast) 10 1.000 4.333 SI 10 0.365 3.644 RC Bypass 8

0.906 4.580 I

I E

I I

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=

l

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4,

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9,

s 4, 0)

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=

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,p)

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D=S C

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I SARGENT & LUNDY SL-4518 05-12-89 I'

7-1 I

Section 7 CRACK STABILITY ANALYSIS The Byron /Braidwood LBB program used a modified limit load approach similar to the crack stability evaluation of ASME Code Section XI. The limit load approach assumes that failure occurs when the stress in the uncracked pipe ligament has become fully plastic without I

further crack extension. EPRI studies (Reference 11) have justified the use of Z~ factors modifying the limit load approach to conservatively evaluate crack stability by elastic-plastic fracture mechanic methods. The Z-factor is a reduction factor which accounts for the reduced fracture toughness in flux welds. Although 2-factors have not been developed for thermally aged cast stainless steel, this analysis assumed the Z-factors applicable for stainless steel SAWS to be governing. To validate this assumption, a 3-Tearing evaluation of the cast stainless steel nozzle was also performed.

For circumferential cracks, master curve methodology (Reference 6) based on the modified limit load approach is used to establish the critical crack length (CCL) associated with the combined dead weight, normal operating thermal, SSE and SAM loads from Section 4. To establish the margin of 2 on crack size, the ratio of the CCL to the LSC must be greater than or equal to 2. To establish the margin of 1.4 on load, a different master curve is produced using the combined loads multiplied by the margin factor of 1.4. The ratio (CCL to the LSC) must be greater than or equal to 1. The master curve methodology presented in the proposed SRP 3.6.3 (Reference 6) has been used for the stainless steel piping. For the cast stainless I

steel nozzle, a 3-Tearing evaluation was also performed using a procedure first suggested by Zahocr (Reference 12). The EPRI computer code FLET (Reference 13) was used to perform this evaluation using the lower bound cast stainless steel material properties identified in Section 3.

I l I I

l i

SARGECT & LUNDY SL-4518 l

05-12-89 l

7-2

!ll MASTER CURVE CONSTRUCTION The master curve was constructed by plotting the stress index,51, given by I

Si = 5 +MP (7~I) m I

as a function of the crack length, L defined as L=2eR (7-2) where S = 2af (7_3)

[2 sin 8 - sin e],

s = [ (3 - 0) - w ( Pm/ f ) ]/2, (7-4) and the half angle of the crack size in radians, e

=

the pipe mean radius in inches R

=

I Pm= the combined membrane stress in psi I

load margin factor; 1.0 for crack size margin M

=

or 1.4 for margin on load

= flow stress of I

sL-4518 Io

-SARGENT5 LONDY SL-4518 05-12-89

'I.

7-3 l

l If e + s is greater than 3, then l.

S = 2e (7-5)

[ sin s ]

with a defined as l

8 = - 3 (P l'f)

(7-6) m The Si values are dependent on the material flow stress, af, therefore a master curve was generated for be.se metal and weld metal using the applicable flow stress. The flow stress used to construct the master curve is defined as og = { (ay + o ). The lower bound material u

properties identified in Section 3 produced the following of which were used to construct the master curves.

I Stainless Steel Base 41421 psi II)

Stainless Steel Weld 51000 psi Cast Stainless Steel 44406 psi I

(1) As required in Reference 6 I

Because the Si values are also dependent on the load margin factor, M, a master curve was generated for M=1.0 and M=1.4.

The master curves constructed with M=1.0 are used to I

establish the margin on crack size while the master curves constructed with M=1.4 are used to establish the margin on loads. The P values used in the construction of the master I

curves combined absolutely the dead weight, SSE and SAM loads presented in Section 4.

m The master curves constructed for the Byron /Braidwood LBB program are presented in Figures 7-1 through 7-18.

I I

sL-4518 I

SARGENT& LUNDY st A 518 05-12-89 7-4 I

MASTER CURVE METHODOLOGY After construction of the master curves, actual stress indexes were calculated for the base metal and weld metal of each Byron /Braidwood LBB system. These actual stress indexes are used to enter the applicable master curve and extract the associated crack length. This crack length is the critical crack length (CCL).

For stainless steel base metal, the actual stress index (ASI) is ASI = M (Pm+P)

(7-7}

b where Pb = the combined primary bending stress, including dead weight, SSE, and SAM combined absolutely The ASI for stainless steel welds and cast stainless steel is g

5' ASI = M (Pm+Pb+P)Z (7-8)

E e

where I

Pe = combined expansion stress due to normal operating thermal and is combined algebraically, Z = 1.15 [1.0 + 0.013 (OD-4)] for SMAW, Z = 1.30 [1.0 + 0.010 (OD-4)] for SAW or cast stainless steel, and OD = pipe outside diameter in inches.

The Z-factor (Reference 11) accounts for the reduced fracture toughness of flux welds, which are not expected to reach full plasticity at failure. For this reason the secondary stresses, P, may not be relieved as failure approaches, and therefore are included in the stress index e

for flux welds. Since cast stainless steel is subject to reduced fracture toughness from l

,<..,,8 I

SARGENTS LUNDY St-4518 05-12-89 8

7-5 thermal aging, the Z-factor for SAW and the secondary stresses, P, were included in the e

stress index for cast stainless steel.

MASTER CURVE RESULTS Table 7-1 presents the actual stress indices (ASI), using M=1.0 to assess the margin on crack size.

It presents the CCL obtained from the applicable master curves and the ratio, CCL/LSC, as the crack size margin.

The required crack size margin of 2 has been demonstrated for circumferential cracks in base metal and welds of the RHR, 51, and RC Bypass systems. Table 7-2 presents the ASI using M=1.4 to assess the margin on loads. The I

CCL from the applicable master curves and the CCL/LSC ratio are presented. The required margin of 1.4 on loads has been demonstrated for circumferential cracks in base metal and welds of the RHR, SI, and RC Bypass systems.

3-T CRACK STABILITY EVALUATION Because the application of a modified limit load approach to thermally aged cast stainless steel flaws has not been developed, a 3-T analysis on the postulated cast stainless steel crack was pe.rformed to confirm the master curve results.

The EPRI computer code FLET (Reference 13) was used to perform this evaluation. It uses the EPRl/GE J-estimation scheme to compute the 3 parameter associated with the LSC length and the applied load. As defined by Zahoor (Reference 12), the tearing modulus, T, is computed for the applied load I

and the LSC length by calculating the change in 3 caused by an incremental change in crack length using the following definition for T.

E dJ M*

(7-9) 2 da E

E is the modulus of elasticity and og, the flow stress.

I sL-4518 I

SAF.6ENT & LUNDY SL-4518 05-12-89 7-6 I

This 3 and associated T were used to determine crack stability by locating it on a 3 versus T plot of the 3-Resistance (3-R) data.

The tearing modulus, T, is calculated from the bifunctional curve fit of the lower bound 3-R material data identified in Section 3 using equation 7-12. Once the material property 3-T curve was established, the intersection point of the line extending from the origin through the assessment 3, T point and the material property 3-T curve defines the T at which instability occurs. Crack stability is ensured for e.ssessment values of T that are less than the T at instability. The loads used are the gi algebraically combined deadweight and normal operating thermal absolutely summed with the E

SSE and SAM loads from Section 4.

I Using the cast stainless steel. lower bound toughness properties from PIFRAC, Figures 7-19 and 7-20 present the J-T plots for the margin on crack size and the margin on load evaluations, respectively. For margin on crack size, the required crack length of 2.0 times the LSC was found to be stable. For margin on load, the LSC was found to be stable when subjected to combined loads increased by the required factor of 1.4. The assessment values 2

of 3 for both crack size and load margin are well below the 3000 in-lb/in limit specified in Reference 14.

Table 7-3 lists the J-T assessment an i the J-T instability values for the cast stainless steel nozzles.

MARGIN FOR AXIAL CRACKS The axial crack stability evaluation was performed to assure that the required margin on size and load are met.

This evaluation was performed using the PICEP computer code to calculate the CCL. PICEP uses the empirical formulation for axial crack failure presented in (Reference 15).

Il Il sL-4518 I!

SARGENTE LUNDY SL-4518 l

OF12-89 7-7 I

In evaluating crack stability for an axial crack orientation, only pressure loads are significant and were the only loads used for this evaluation. Lower bound elastic properties from Section 3, and Z-factors previously identified in Section 7 were used in PICEP to calculate the CCL.

I Table 7-4 presents the axial crack size margins and Table 7-5 presents the axial crack load margins. These margins are well above the required values.

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sL-4518 I

l

E SARGENT& LUNDY Circumferential Crack Size Table 7-1

_g Margins SL-4518 05-12-89 5

l I

I Margin j

System ASI CCL LSC (CCL/LSC 2 2.00) l RHR Base Metal 17,455 13.990 5.652 2.48 Weld 26,010 12.642 5.652 2.24 i

SI "t= 1.000" Base Metal 8,632 15.210 3.450 4.41 Weld 29,129 9.780 3.450 2.84 Cast 29,129 8.650 3.901 2.22 51 "t=0.365" Base Metal 13,484 14.132 5.142 2.75 Weld 20,685 12.932 5.142 2.51 RC Bypass Base Metal 15,342 10.006 3.699 2.71 Weld 29,842 7.627 3.699 2.06 I

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!I SARGENT E LUNDY Circumferential Crack Table 7-2 Load Margins SL-4518 E.

05-12-89 Margin System ASI CCL LSC (CCL/LSC 21.00)

RHR Base Metal 24,436 12.029 5.652 2.13 Weld 36,423 10.044 5.652 1.78 l

51 "t= 1.000" I.

Base Metal 12,085 14.537 3.450 4.21 Weld 40,780 7.325 3.450 2.12 Cast 40,780 5.949 3.901 1.52 51 "t=0.365" Base Metal 18,738 12.771 5.142 2.48 Weld 28,958 11.012 5.142 2.14 RC Bypass Base Metal 21,479 8.734 3.699 2.36 SAW 41,778 5.548 3.699 1.50 I

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sL-4518

~

I SARGENTE LUNDY 3-T Results for Cast -

Table 7-3 Stainless Steel SL_4318 I

03-12-89 I

g Crack Size Margin of 2:

Assessment Instability (in-fb) 3(in-fb)

T T

3 in in 1800 26 6279 92 Load Margin of 1,4 l

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SL-4518 I

SARGENT& LUNDY Axla! Crack Size Margins Table 7-4 SL-4518 lI' 05-12-89 I

j l g..

lE CCL LSC Margin System (inches)

(inches)

(CCL/LSC 2 2.00) l RHR 16.96 4.946 3.43 51 (1")

15.46 7.393 2.09 SI (.365")

9.66 3.644 2.65

. RC Bypass 15.6 4.580 3.41 I

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I SL-4518

g:

SANGENTE LUNDY Axial Crack Load Margins Table 7-5 l

SL-4518 I

05-12-89 CCL LSC Margin 1

System (inches)

(inches)

(CCL/LSC 21.00) l RHR 11.76 4.946 2.38 51 (l")

10.75 7.393 1.45 51(.365")

6.69 3.644 1.84 RC Bypass 10.93 4.580 2.39 I

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SL-4518

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SARGENT E LUNDY

. Master Curve (RH System) Base Figure 71

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l

SARGENT S LUNDY SL-4518 05-12-89 8-1 Section 8 REFERENCES 1.

Water Hammer in Nuclear Power Plants, NUREG-0582, July 1979 2.

Evaluation of Water Hammer Events in Light Water Reactor Plant, NUREG/CR-2781, July 1982 3.

Evaluation of Water Hammer Occurrence in Nuclear Power Plants, NUREG-0927, Rev.1, March 1984 4.

Letter from W. E. Morgan of Commonwealth Edison to U. 5. Nuclear Regulatory Commission - Document Control Desk, dated October 3,1988,

Subject:

Response to NRC Bulletin 88-08 and 88-08: Supplement I and 2 5.

"A User's Guide to the NRC's Piping Fracture Mechanics Data Base (PIFRAC),"

NUREG/CR-4894, MEA-2210, May 1987.

6.

Proposed SRP 3.6.3 Leak-Before-Break Evaluation Procedures, September 1987.

7.

B. R. Ganta, D. 3. Ayres, " Analysis of Cracked Pipe Weldments", EPRI NP-5057, February 1987.

8.

W. Server, B. Beaudoin, D. Quinones, " Applying Leak-Before-Break to High-Energy Piping" NSAC-Il4, November 1987.

9.

Report of the U.S. Nuclear Regulatory Commission Piping Review Committee;

" Evaluation of Potential for Pipe Breaks," NUREG-1061 Volume 3, November 1984.

10.

D. M. Morris, B. Chexal, "PICEP; Pipe Crack Evaluation Program", EPRI NP-3596-SR, Revision 1, Special Report, December 1987.

11.

Evaluation of Flaws in Austenitic Steel Piping, EPRI NP-4690-SR, July 1986.

1 L

12.

A. Zahoor and R. M. Gamble. Evaluation of Flawed Pix Experiments. Palo Alto, Calif.: Electric Power Research Institute, November L986. NP-4883M.

13.

A. Okamoto, D. M. Morris, FLET: Pipe Crack Instability Program, EPRI, Revision 0, September 1988.

1 i

sL-4518

7 g

SARGENT E LUNDY SL-4518 l

05-12-89 8-2 I.

l 14.

Be 3. Youngblood, Chief, Licensing Branch No.1, Division of Licensing, U. S. Nuclear Regulatory Commission, Washington, D. C., letter to D. L. Farrar, Director of Nuclear Licensing, Commonwealth Edison Company, Chicago, Illinois, Pertaining to E

Docket Nos. STN 50-455, STN 50-456, and STN 50-457, October 28,1985.

E 15.

F. Erdogan. " Ductile Fracture Theories for Pressurized Pipes and Containers."

g International Journal of Pressure Vessels and Piping, Vol. 4 No. 4, October 1976, g'

pp. 253-283.

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