ML20247K301

Leak-Before-Break Evaluation for Carbon Steel Piping

ML20247K301
Person / Time
Site:	Byron, Braidwood, 05000000
Issue date:	05/12/1989
From:	SARGENT & LUNDY, INC.
To:
Shared Package
ML20247K294	List: ML20247K288 ML20247K301
References
SL-4519, NUDOCS 8906010267
Download: ML20247K301 (81)

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I Nuclear Safety-Related iI Leak-Before-Break Evaluation l

l for Carbon Steel Piping Byron and Braidwood Nuclear Power Stations Units 1 and 2 Prepared for I

Commonwealth Edison. Company Report SL-4519 May 12,1989 I

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SARGENT & LUNDY g

! I M2703.073 05-89

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j SARGENTS LUNDY SL_g319 0F12-89 I

ES-1 EXECUTIVE

SUMMARY

I 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 I

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.

I 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)

I approach to the NRC licensing process. This document provides a step-by-s.tep 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 I

deletion of snubbers with the following results.

DELETION OF PIPE WHIP RESTRAINTS There are currently 16 pipe whip restraints installed on the systems in the LBB scope.

These whip restraints are installed in congested areas of the containment building and with l

minimal gaps between them and the piping system. The possibility exists for a whip restraint to interfere with the thermal expansion of the piping and potentially cause an unanalyzed E

sL-4519 I

SARGENT E LUNDY SL_4319 05-12-89 ES-2 crndition. The integrated hot functional test included walkdowns to identify and resolve th;se 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 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 ISI, 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 often produces the governing piping forces and movements for the connected piping. To accommodate these dynamic loads at locations with high thermal movements, snubbers are of ten used as dynamic pipe supports. The Station

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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 allloop 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 l

proven, the LOCA forcing function can be eliminated from the piping analysis for the loop connected piping. 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 l

I testing requirements, sL-4519

SANSENT& LUNDY St_43g9 05-12-89 ES-3 The remainder of this report is broken into seven sections as follows:

Scope System Review Versus NUREG-1061 Limitations a'

Material Property Data Applied Loading Leak Detection Leak Rate Crack Sizing Crack Stability Analysis A brief description of the contents of each of these sections follows:

Scope

?

This section describes the system within the scope of the LBB program and addresses its safety functions. Piping and pipe whip restraint data are provided along with an isometric drawing of each piping subsystem.

System Review Versus NUREG-1061 Limitations The applications of LBB technology is limited by six criteria specified in Volume 3 of NUREG-1061, Section 5.1.

This section of the report evaluates the Byron /Braidwood LBB main steam 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, coa +ainment, and other engineering safety features remain unchanged.

The 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 main steam system wris found not to be susceptible to these failure mechanisms.

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

SL-4519

3 I

SARGENT& LUNDY sL_4319 03-12-89 I

ES-4 8

I The evaluation of the Byron /Braidwood LBB main steam system for the

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probability of degradation or failure due to indirect causes such as fires, I

missiles, and equipment, systems, or components in close proximity was performed during the design of Byron /Braidwood and found acceptably low.

I The Byron /Braidwood LBB main steam system has been verified against the

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requirements of I&E Bulletin 79-14.

The carbon steel materials in the Byron /Braidwood LBB main steam system I-are not susceptible to cleavage type fractures over the full range of their

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operating temperatures.

I 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 main steam system by establishing that the PlFRAC tensile properties bounded the range of tensile properties obtained from'the i

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 respectively, it also establishes the lower bound toughness properties used to evaluate crack stability.

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

analytical model.

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SARSENT& LUNDY SL-4519 05-12-89 g-n, i

Leak Detection This section describes the Byron and Braidwood leak detection systems including the l

Technical Specification requirements.

The section provides the basis for using the assumption that a leakage rate of I gallon per minute (gpm) can be identifi...cy the operators within one hour, Leak Rate Crack Sizing The leakage size crack (LSC) was established for tha main steam 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 I

10 gpm.

Crack Stability Analysis The crack stability evaluation was performed using master curve methodology based on a l

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modified limit load approach which uses Z factors to conservatively evaluate crack stability by elastic-plastic fracture mechanics methods. f& ster curves were developed for base and weld metal cracks. From these master curves, the critical crack length (CCL) associated with the normal operating load (NOL) absolutely combined with the safe shutoown earthquake (SSE) load was obtained and compared to the LSC demonstrating the required margin of 2.0 on stable crack size. The combined normal operaung and SSE losd was increased by the required factor of 1.4 and used to obtain the CCL from the master curve. This CCL was i

compared to the LSC demonstrating the required margin on load by being greater than the LSC. In addition, a 3-Tearing evaluation was perior teu on the carbon steel components using the EPRI computer code FLE7, Flaw Evaluation by TearLg 1 'tability. This evaluation confirmed the results of the master curve methodology.

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i I SMRGENT E LUNDY f3_$l9g, f

I In conclusion, this report docu:nents that the main steam piping in the Byron /Braidwood LBB prograrn 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 required by the October 27, 1987, Federal Register have been demonstrated. Therefore, the use of LB3 to eliminate the dynamic effects of postulated pipe breaks in these systems is gy appropriate.

I SARGENT & LUNDY, I

Prepared by:

/ cm (d, bVS 05-12-89 I

G.H.DeBqo Supervisor Engineering Mechanics Division E

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05-12-89 Reviewed by:

E P. H. Hoang 5

Senior Engineering Specialist Engineering Mechanics Division f

Approved by:

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05-12-89 I

J. R. Meister Project Manager Project Management & Engineering Division p%

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1 CONTENTS Section Page ES EXECUTIVE

SUMMARY

ES-1 1

SCOPE l-1 Description of Main Steam System - MS 1-1 Piping Within Scope 1-2 2

SYSTEM REVIEW VERSUS NUREG-1061 LIMITATIONS 2-1 4

Design Criteria 2-2 Susceptibility to Failure 2-2 Stress Corrosion 2-2 Water / Steam Hammer 2-2 Thermal and/or Mechanical Cyclic Fatigue 2-4 Erosion / Corrosion 2-5 Creep 2-6 8

Component and Piping Supports 2-6 Degradation or Failure from Indirect Causes 2-6 I

IE Bulletin 79-14 Verification 2-7 Cleavage-Type Fracture Susceptibility 2-7 3

MATERIAL PROPERTY DATA 3-1 Installed Material 3-1 LBB Material Properties 3-2 Test Data Curve Fitting 3-4 4

APPLIED LOADING 4-1 4-1 High Stress Locations 4-2 Applied Loads 5

LEAK DETECTION 5-1 5-1 Major Leak Detection System Components 5-2 Component Design I

sL-4519 i

I

SARGENT E LUNDY SL-4519 03-12-89 il CONTENTS, Cont.

Section P_ age Containment Cavity and Floor Drain Sumps 5-2 Containment Particulate and Gaseous Radioactive Monitoring Systems 5-3 Reactor Containment Fan Cooler 5-3 Conclusions 5-3 6

LEAK RATE CRACK SIZING 6-1 7

CRACK STABILITY ANALYSIS 7-1 Master Curve Construction 7-1 Master Curve Methodology 7-3 Master Curve Results 7-4

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3-T Crack Stability Evaluation 7-5 7-6 Margin for Axial Cracks 8

REFERENCES 8-l l'

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sL-4519

1 SARGENTS LUNDY SL-4519

~05-12-89 i

ill TABLES Table 1-1 Byron /Braidwood Leak-Before-Break 1.ine Data 1-2 Byron /Braidwood Leak-Before-Break Pipe Whip Restraints 3-1 Byron /Braidwood LBB Installed Material 3-2 Byron /Braidwood LBB Weld Prncesses 3-3 Installed Base Metal Tensile Property Boundaries at Room Temperature 3-4 Installed Weld Metal Tensile Property Boundaries at Room Temperature 3-5 Material Test Data in PIFRAC 3-6 PIFRAC Base Metal Tensile Properties at Room Temperature 3-7 PIFRAC Carbon Steel Weld Metal Tensile Properties at Operating Tempr 'ture 3-8 PICEP Material Properties 3-9 FLET Tensile Material Properties 3-10 FLET Fracture Toughness Material Properties 4-1 Main Steam System High Stress Point Loads 6 Leakage Size Crack Length 7-1 Circumferential Crack Size Margins 7-2 Circumferential Crack Load Margins 7-3 3-T Results for Main Steam 7-4 Axial Crack Size Margins 7-5 Axial Crack Load Margins I

I sL-4519 i

I SANGENT E LUNDY gt_g3g9 03-12-89 5

iv FIGURES Figure 1-1 Main Steam Subsystem MS05 1-2 Main Steam Subsystem MS06 1-3 Main Steam Subsystem MS07 1-4 Main Steam Subsystem M508 31 Carbon Steel True Stress-Stram Base Metal 3-2 Carbon Steel Weld and Base Metal 3-R Curves 3-3 Carbon Steel Mean Stress-Strain Curve Fit 3-4 Carbon Steel Lower Bound Stress-Strain Curve Fit 3-5 Carbon Steel Lower Bound 3-R Curve Fit 6-1 Leak - MS (OD=30.25 '..) Crack Length Curve (Carbon Steel) 6-2 Leak - MS (OD=32.75 in) Crack Length Curve (Carbon Steel) 7-1 Master Curve (MS System) Base Material Crack Margin (OD=30.25 in) 7-2 Master Curve (MS System) SAW Weld Crack Margin (OD=30.25 in) 7-3 Master Curve (MS System) Base Material Load Margin (OD=30.25 in) 7-4 Master Curve (MS System) SAW Weld Load Margin (OD=30.25 in) 7-5 Master Curve (MS System) Base Material Crack Margin (OD=32.75 in) 7-6 Master Curve (MS System) SAW Weld Crack Margin (OD=32.75 in) 7-7 Master Curve (MS System) Base Material Load Margin (OD=32.75 in) 7-8 Master Curve (MS System) SAW Weld Load Margin (OD=32.75 in) 7-9 3-T Curve for Carbon Steel MS Piping (OD=30.25 in) Crack Margin I

7-10 3-T Curve for Carbon Steel MS Piping (OD=32.75 in) Crack Margin st-4519 5

I ~ SARGENT& LUNDY SL-4519 05-12-89 5

y FIGURES, Cont.

Figure 7-11 3-T Curve for Ca %n Steel MS Piping (OD=30.25 in) Load Margin 7-12 3-T Curve for Carbon Steel MS Piping (OD=32.75 in) Load Margin i

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Section i SCOPE I

The leak-before-break (LBB) program for the Byron and Braidwood stations includes several high-energy piping systems that have postulated high-energy line break locations in the current design. The main steam piping inside the containment building is in the program.

DESCRIPTION OF MAIN STEAM SYSTEM - MS I

The main steam (MS) system's primary function is to contain and transport steam from the steam generators to the high-pressure turbine to drive the main generator.

I The safety-related portion of the MS system runc from the steam generator nozzle to the main steam isolation valves (MSIVs) outside containment in the valve rooms. This portion of the system has three safety functions.

In the event of a reactor trip, the steam generator acts as a heat sink to prevent heatup of the reactor coolant system. The MS system removes heat generated from the reactor through the PORVs, the safety valves, or the steam dump valves when the turbines are not available.

In the event of high-high containment pressure, low steamline pressure, or I

high steam pressure rate, the engineered safety feature (ESF) system provides signals for closure of the MSIVs and the MSIV bypass valves.

In the event of tube leaks in one of the steam generators, the MSIVs and the I

MSIV bypass valves provide an isolation boundary to limit contamination of the rest of the secondary system.

The portion of the MS system piping under this evaluation is between the steam generator and the containment penetration assembly that forms an analytical boundary.

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I sL-4519 5

I SARGENT & LUNDY SL-4519 05-12-89 3

1-2 g

I PIPING WITHIN SCOPE

!sometric drawings of the piping systems in this scope are shown in Figures 1-1 through 1-4.

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

B Braidwood units. Minor differences in routing between the units and the actual pipe support 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.

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

Table 1-2 presents the pipe whip restraints on the piping that was evaluated as a part of the Byron /Braidwood LBB main steam program. The table includes the appropriate loop number, line number, and whip restraint number.

There are a total of 32 postulated break locations in the piping within the scope of the LBB program and a total of 16 pipe whip restraints.

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sL-4519 l

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SARGENT S LUNDY Byron /Breidwood Lcak-Table 1-1 Before-Break Line Data SL-4519 05-12-89 E

Operating

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' MS01 AA-30.25 MS05 30.25 1.250 SA155 KC65 960 539 M501AB-32.75 MS06 32.75 1.343 SA155 KC65 960 539 l'" '

M501AC-32.75 M507 32.75 1.343 SA155KC65 960 539 MS01AD-30.25 MS08 30.25 1.250 SA155 KC65 960 539 I

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SARGENTS LUND1f Byron /Braidwood Le:k-Table 1-2 Before-Break Pipe Whip SL-4519 Restraints (1) 05-12-89 I

Pipe Whip Restraint Subsystem Line No.

Loop No.

MS-P1 MS05 MS01AA30.25 1

MS-P16 M506 MS01AB32.75 2

MS-P23 M507 MS01AC32.75 3

MS-P8 MS08 MS01AD30.25 4

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Note 1: These are the pipe whip restraints in a single unit.

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1 Section 2 SYSTEM REVIEW VERSUS NUREG-1061 LIMITATIONS NURr G 1061 provides recommendations for application of the LBB approach in the NRC licensin~g 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 I

in accordance with existing regulations. The evaluation of environmental effects should be considered case-by-case.

I-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-granular stress cerrosion cracking), water hammer, or low-and high-cycle I

(i.e., thermal, mechanical) fatigue.

Component and piping support structural integrity should be maintained with I

no reduction in margin for the final safety analysis report (FSAR) loading combination that governs their design.

The LBB approach should not be considered applicable if there is a high I

probability of degradation or failure of the piping from more indirect causes such as fires, missiles, 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, 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 I

the full range of system operating temperatures where pipe rupture could have significant adverse consequences.

The six limitations are addressed relative to the Byron /Braidwood units in following subsections. These subsections demonstrate that these limitations do not apply to the main steam system addressed in this report.

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Ii SARGENT & LUNDY SL-4519 3]!

05-12-89 2-2 3!

I, DESIGN CRITERIA s

The containment was designed to accommodate a LOCA resulting from breaks in the reactor Ei coolant pressure boundary up to and including a break equivalent in size to the double-ended 5j rupture of the largest pipe in the reactor coolant system. It is not proposed to use the LBB l

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.

l SUSCEPTIBILITY TO FAILURE I

This section addresses susceptibility of the systems to failure from the effects of stress corrosion, water / steam hammer, thermal and mechanical cyclic fatigue, erosion / corrosion, g

and creep.

W I

Stress Corrosion The susceptibility of the main steam system to stress corrosion cracking has been reviewed, and it has been determined that carbon steel piping in pressurized water reactors is not susceptible to stress corrosion cracking.

I Water / Steam Hammer The main steam system 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 following phenomena: flow into empty or voided lines; rapid valve opening, closing, or instability; water entrainment in steam lines; column separation; steam bubble collapse; slug irnpact 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.

sL-4519 I

SARGENT& LUNDY SL-4519 05-12-89 I

2-3 I

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 events were improbable and the severity of the repcrted incidents was low. It should be noted that all reported water / steam hammer events resulted in only support damage. There were no events in PWRs that resulted in a loss of pressure boundary integrity. Additional 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 I

3).

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

I NUREG 2781 (Reference 2) identified six water / steam hammer events in the main steam system. Of the six events evaluated, four were anticipated steam hammers resulting from f

valve closure; one event reported damage to two hydraulic snubbers on a relief line, which l

were probably caused by relief valve actuation; and one was caused by inadvertently opening the main steam isolation valves (MSIVs) on a partially warmed main steam line, which caused steam to condense and created a water slug that impacted on the closed turbine stop valve.

The damage caused by these events has been limited to pipe supports outside the reactor building. The events caused by valve closures and the relief valve discharge are postulated events, and the main steam system has been designed for them. The one unanticipated event where the MSIV was inadvertently opened before adequately heating the main steam line did not cause any damage to piping or supports.

As discussed below, it is unlikely-that water / steam hammer could affect the safety-related portion of MS system piping. In the design of the MS system, piping transient loads from steam hammer due to valve closure and relief valve discharge are included in pipe support -

l and component design. These are the only water-hammer-type loads postulated to affect the portion of the main steam system within the scope of this report. During integrated hot functional testing of Braidwood Unit 1, water / steam hammer did occur between the MSIV and l

the main steam dump valves. The main steam dump valves to the condenser were opened sL-4519 i

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SARGENT & LUNDY SL4519 05-12-89 24 I

causing depressurization of the main steam header. This sudden depressurization resulted in flashing and acceleration of condensate which had accumulated down stream of the closed MSIV. In another related event, Byron Unit i experienced a water hammer downstream of the B MSIV when restarting after the first refueling outage. A quantity of water that collected in a low point in the system was accelerated by steam pressure down the pipeline causing hammering at the downstream elbows when the MSIV bypass valve was open.

These events did not affect any portion of safety-related piping, which is upstream of the MSIV. Some pipe supports in non-safety-related piping were damaged but no damage to the non-safety-related piping was identified. The configuration of the safety-related piping is such that any accumulated condensate will dra'in to the MSIV. A 12-inch-diameter drip leg upstream of the MSIV will collect any condensate when the valve is closed and allow for manual drainage of the line. As a result of the events at Byron and Braidwood, the following modifications were made to prevent recurrence of the main steam water / steam hammer in the non-safety-related portion of the system:

Operation of the main steam dump valves was revised to provide finer g

control of main steam header pressurization.

W Administrative procedures were established to ensure periodic draining of the g

drip leg upstream of the B MSIV during heatup.

g A new drip leg and associated drain piping was added downstream of the B MSIV to allow draining of the low point in the system during heatup.

Thermal and/or Mechanical Cyclic Fatigue Unidentified operating conditions in the main steam system that might contribute to thermal or mechanical fatigue are thermal stratification and system operating or flow induced vibration.

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 sL-4519 I'

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SARGENT & LUNDY SIA 519 05-12-89 2-5 I

I 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 Pipping;" 88-08, " Thermal Stresses in Piping Connected to Reactor Coolant Systems;" and most recently, 88-11, " Pressurizer Surge Line I

Thermal Stratification." The flow stratification or uneven thermal mixing identified in these documents is not postulated to occur in the main steam system. Also previous industry operating experience has not identified thermal stratification in this system. The main I

steam system has been reviewed for its susceptibility for thermally stratified flow. This review has determined that thermal stratification is not possible in this portion of the main steam system (Reference 4).

System operating flow induced vibration in the main steam system was monitored during the preoperational and startup vibration testing at Byron Units 1 & 2 and Braidwood Units 1 & 2.

This system was 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 vibrations, it should be I

noted that positive displacement pumps are not used by this system.

Erosion / Corrosion Commonwealth Edison has recognized the potential for erosion / corrosion in steam piping for I

several years and has a program in place to monitor piping for erosion / corrosion at Byron and l

Braidwood. This program evaluates the wall thickness of the piping at several points in the secondary system during each outage. The points monitored were selected based on pipe con-figuration and steam moisture content and are considered the points most susceptible to erosion / corrosion damage.' The portion of the main steam system in the LBB program is not a part of the sampling population because the steam moisture content is low. However, the erosion / corrosion program will identify any unusual degradation requiring the evaluation.of I

this piping.

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SL-4519 I

I SARGENT& LUNDY 3L,4319 05-12-89 2-6 I

As a part of the Sargent & Lundy Byron /Braidwood " Analysis of Secondary Loop Systems for Operation at Vessel Hot Leg Temperatures of 600'F and 594*F," SL-4442, the potential for erosion / corrosion of the main steam piping was evaluated. This report evaluated the main steam system based on the S&L-recommended flow velocities for steam piping and the recommendations of the EPRI report " Erosion / Corrosion in Nuclear Plant Steam Piping:

Causes and Inspection Program Guidelines." SL-4442 concludes that the steam moisture content and velocity is well below the values for which erosion / corrosion would be a concern.

Creep I

Creep and creep-fatigue is not a concern since all op rating temperatures are well below the temperature of 700*F for ferritic steel piping, at which creep and creep fatigue becomes a concern.

COMPONENT AND PIPING SUPPORTS All equipment, piping and supports within the scope of the main steam LBB program have been seismically designed per the rules of ASME Section 111. Performance of the LBB analysis does not reduce the margin of safety of any supports.

I 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).

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 l

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LI SARGENT & LUNDY St_4319 05-12-89 2-7 I

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

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 Cleavage-type fracture is not a concern for carbon steel at system operating temperatures ranging from 212*F to 557'F for the main steam system.

The piping installation specification required a minimum hydrostatic test temperature of 70'F for carbon steel material (Article 309.10a). Within this temperature range cleavage-type fracture is not a I

concern for carbon steel.

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SANGENT4 LUNDY SL_g319 I

0 F12-89 3-1

.I Section 3 MATERIAL PROPERTY DATA This section identifies the material properties of the piping, piping components, and welds installed on the Byron /Braidwood LBB main steam system. 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 materials are I

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 the mean and lower-bound true stress-true strain data for sizing the leakage crack and for evaNating 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 I

the lower bound fracture resistance (3-R) data.

INSTALLED MATERIAL The installed materials on the Byron /Braidwood LBB main steam system were identified by obtaining their certified material test reports (CMTRs). From these CMTRs the material specifications 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 specifications for the installed piping and fittings is presented in Table 3-1. For the main steam lines, the installed materialis predominantly A155 KC65.

Welding processes and welding materials were obtained from the fabricator's and installer's l

I weld process sheets. Table 3-2 presents a summary of the welding processes and filler metals I

sL-4519

SARGENT & LUNDY SL-4519 05-12-89 3

3-2 5,

- i 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 the main steam lines mest of the fabrication welds are predominantly SAW type using EM13K filler metal and the field welds are predominantly SMAW type using E7018.

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.

LBB MATERIAL PROPERTIES Room temperature UTS and YS were obtained for materials similar to those installed on the Byron /Braidwood LBB main steam system from the stress strain data given in NRC's P! FRAC data base. The 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 installed SA234, Grade WPB fittings were not available 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 l

represent the installed fittings.

The tensile properties at room temperature for test specimens from the chosen heats are presented in Table 3-6 for base metals. The comparison of these tensile properties 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 main steam system material properties.

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I SARGENT & LUNDY St_,319 05-12-89 3-3 I

I True stress-true strain data for test specimens from the heat numbers presented in Table 3-5 have been plotted in Figure 3-1 for carbon steel base metals. To coincide with the operating temperatures of the Byron /Braidwood LBB main steam system, the chosen carboa steel test specimens were tested at 550*F and 70*F. Since the size of the leakage crack could be underestimated by using the lowest stress-strain data, the mean stress-strain data are used to conservatively calculate the leakage size crack (LSC). By examining the stress-strain plots in Figure 3-1, the mean stress-strain data are determined to be from test specimen F26-5 of Heat No. 52 for carbon steel base metal. Lower bound stress-strain data are required to conservatively represent the material tensile properties in the crack stability evaluation.

I The lower bound stress-strain data are determined to be from test specimen F6-1 of Heat No.

48 for carbon steel base metal.

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 for carbon steel base metals. From Table 3-7 the lower bound flow stress for carbon steel weld metal was determined to be 62,278 psi from test specimen F40W2-59 of Heat No. 59.

Recent finite element analysis studies evaluating cracks in pipe weldtrients (Reference 6) have shown that elastic-plastic fracture mechanics solutions using only weld metal predict higher crack initiation and instability moments when calculated using the EPRI estimation scheme contained in the Flaw Evaluation by Tearing Instability (FLET). These studies recommend the use of lower bound base metal stress-strain with lower bound weld metal fracture toughness (3-R) in FLET. Deformation 3 and aa data for test specimen from the heats presented in Table 3-5 are plotted on Figure 3-2 for carbon steel base and weld metal.

I The chosen carbon steel test specimens were tested at 550*F. Also, only test specimens with a thickness equal t r greater than the pipe wall thickness were chosen. The lower bound l 53 3-R data were determined to be from test specimen W1 of Heat No. 62 for carbon steel l

SAWS.

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sL-4519 I

SARGENT & LUNDY SL-4519 05-12-89 R

3-4 3

I TEST DATA CURVE FITTING The EPRI Pipe Crack Evaluation Program (PICEP) used to determine the LSC and the FLET g

program used to evaluate crack stability require the true stress-true strain data to be curve 5

fitted by the Ramberg-Osgood relation.

L+a[L (3-1) 1 S

"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, o, 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 7, the Ramberg-Osgood coefficient and exponent were determined from the g

chosen test specimen true stress-true strain data in the yield strain to 1% strain range. The 3

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. Figure 3-3 presents the plot of the stress-strain test data with the Ramberg-Osgood curve fit for the chosen mean carbon steel test specimen. Figure 3-4 presents the plots of the stress-strain test data with the Ramberg-Osgood curve fit for the chosen lower-bound carbon steel base metal test specimen.

The FLET program uses a bifunctional curve fit for the fracture toughness test data. The bifunctional curve fit form is I,

J=c (aa)*, aa < aa (3-2) transition I

/ 3), aa > aa (3~3)

J=ag+a2e transition where J is the deformation 3 integral parameters, as is the crack extension and aatransition is the crack extension at which extrapolation of the test data is necessary to ensure g

3-controlled crack growth for large crack extensions. As specified in NUREG-1061 Volume E

is defined as the aa at w = 5.

3,aatransition sL-4519 i

SARGENT & LUNDY SL-4519 05-12-89 3-5 w is defined by w=hh (3-4) 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

. frccture toughness properties for use in FLET are presented in Table 3-10. Figure 3-5 shows thm plot of the 3,aa test data with the bifunctional curve fit for the chosen lower bound cirbon steel weld material.

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'sL-4519

,.J SARGENT&LUNDY Byron /Breldwood LBB Table 3-1 Installed Material SL-4519 05-12-89

.I Carbon Steel Piping A155 VsC65 A155 KC70 Fittings SA234 WPB Valves NONE 3

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st-4519 I

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SANGENT & LUNDY Byron /Braidwood LBB Tabie 3-2

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Weld Processes 519 l

3' carbon steel weids ggrxess ggetal

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SAEGENT E LUNDY Installed Base Metal Tensile

- Table 3-3

,E; Property Boundaries at Room SL-4519 3 -'

Temperature 05-12-89 I

M:terial

-UTS (psi)

YS (psi)

Carbon Steel:

Maximum 81,900 51,700 Minimum 66,000 35,100 I

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SARGENTS LUNDY Installed wald Metal Tensile Table 3-4 Property Boundaries at Room SL-4519 I

Temperature 0.5-12-89 Materls!

UTS (psi)

YS (psi)

Carbon Steel:

l Maximum 84,900 72,100

)

Minimum 64,750 48,500 I

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sL-4519 4

i SARGENT& LUNDY M:t: rial Test Data in PIFRAC Tabis 3-5 SL-4519 05-12-89 1

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1 Material Heat Weld Specification No.

Type Carbon Steel:

Base Metal A155 52 SA333 GR6 48 SA333 GR6 49 SA333 GR6 50 Weld Metal SA51C GR70 62 SAW SA516 GR70 59 SAW SA516 GR70 58 SAW 4

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.5ANGENT S LUNDY PlFRAC Base Mstal Tensile Table 3-6

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Properties at Room Temperature SL-4519 j

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05-12-89 g

Carbon Steel:

Heat No.

Spec. Id.

UTS (psi)

YS (psi)

II) 48 F6-1 75,509 37,775(I) 48 F6-2 65,674 49,773 I

49 F9-1 65,848 39,230 49 F9-2 70,170 39,470 50 Fil-3 73,071 38,682 I

50 Fil-4 76,778 49,417 52 F26-1 63,266 37,203 52-F26-2 76,074 38,481 5

(1)

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

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SARGENT S LUNDY PIFRAC Carbon Steel Weld Metal '

Table 3 E Tensile Properties at Opera 1ng SL-4519 3

Temperature 05-12-89 l

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Heat No.

Spec. Id.

UTS (psi)

YS (psi)

Flow (psi) 58 F40-6-5 99,178 65,886 82,532 58 F40-6-6 106,387 68,361 87,374 39 F40W2-58 101,934 56,251 79,093 I

59 F40W2-59 79,527 45,030.

62,278 62 CW-12 96,761 76,339 86,550 62 CW-M2 95,507 82,284 88,896 62 CW-02 95,507 72,732 84,120 t

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SARGENT& LUNDY PICEP Msterial Properties Table 3-8 SL-4519 03-12-89 I

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I Mean Carbon Steel Base Metal:

Test Specimen: F26-5 Heat No. 52 33,607 psi a

=

o 0.001245 c

=

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=

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n

=

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=

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=

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SARGENTS LUNDY FLET Tensile Materi:1 Properties Table 3-9 SL-4519 I'

05-12-89 I;.

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Lower Bound Carbon Steel Base Metal:

Test Specimen: F6-1 Heat No. 48 I

29,140.6 psi o

=

o j

0.001079 c

=

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=

2.632488 n

=

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=

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SARGENTS LUNDY.

FLET Fractura Toughness Mat::rlal Tabla 3-10

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Properties SL-4519

}

05-12-89 i

Lower Bound Carbon Steel Weld Metal:

Test Specimen: W!. Heat No. 62 f

2 1,530.32 in-lb/in JIC

=

7,425.950195 c

=

0.341156 m

=

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=

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=

0.398714 a3

=

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= 3.4 A

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==

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r SARGENT5 LUNDY gt_,339 05-12-89 I

4-1 I

Section 4 APPLIED LOADING This section identifies the highest stress location for the main steam system and summarizes the procedure used to identify the location. It specifies the loads and directional components to be obtained at this location and the load combination methods. The system thermalloads used in this evaluation correspond to the system ti ermal mode during normal plant operation.

HIGH STRESS LOCATIONS Since the highest stressed location is.o be considered with the most limiting material toughness properties, the highest stress location for the main steam system is identified without consideration of the local material properties.

The highest stress location is I

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 main steam 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 resultant moment for the individual weight, normal operating thermal (NOT), and safe I

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 I

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 sL-4519

I I

SARGENT & LUNDY SL_4319 05-12-89 4-2 s

APPLIED LOADS The axial force, torsional and bending moment components from the high stress point for the Byron /Braidwood LBB main steam system is presented in Table 4-i.

An additior,al review considering only the piping bending moments was conducted. The highest stress location

{

remained the same with and without including the torsional moment.

I 4

I I

iI Ii 1

IJ I

I

\\

sL-4519 I

t

I l

SARGEW& LUNDY Main Ste:m Syst:m High-Table 4-1 Stress Point 1.oads SL-4519 I

05-12-89 I

(kips)

(in-kips)

System Cond.

Fa Ma Mb Mc Mr MS 30.25 Weight 0

-2.

-6.9

-825.3 t = 1.25" Thermal

-6.4 34.5 156.8

-2701.7 n

SSE+ SAM 31.4 1451 6725 2268 Norm Op.

-6.4 32.5 149.9 3527 3530 Faulted 37.8 1483.5 6875 5795 9113 MS 32.75 Weight

.1 9.8 25.9

-1518.6 t = 1.343" Thermal

-8.7

-33.4

-215.2

-3832.0 n

SSE+ SAM 34.7 1975.7 2840.2 8545.6 Norm Op.

-8.8

-43.6

-189.3

-5350.6 5354 Faulted 43.5 2019.3 3029.5 13896.2 14366 E

4 l

I I

I

' I St-4519 I

l SARGENT& LUNDY SL_4319 05-12-89 l.E E

5-1 i

Section 5 1

1 l

LEAK DETECTION I'

j j

General Recommended Action 3 in the Executive Summary of NUREG-1061 Volume 3 (Reference 8) states:

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 j

of detection for the largest postulated leakage size crack used in the fracture i

mechanics analyses is greater than a factor of ten on unidentified leakage.

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.

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

leak.

MAJOR LEAK DETECTION SYSTEM COMPONENTS The design of the reactor containment 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 containment leakage are the reactor cavity and containment floor drain sumps, the containment atmosphere particulate radioactivity monitoring system, and the containment gaseous radioactivity monitoring system. In addition to these three primary systems, there 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 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.

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

I sL-4519 1

I SARGENT & LUNDY SL-4319 05-12-89 3

5-2 g

l the containment atmosphere particulate radioactivity monitoring system, the containment floor drain and reactor cavity flow monitoring system, and the containment gaseous radioactivity monitoring system.

The Technical Specification provides surveillance requirements to demonstrate that each of l

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 g'

atmosphere particulate radioactivity monitoring system, the containment gaseous a

radioactivity monitoring system, 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 that unidentified leakage is less than 1 gpm.

COMPONENT DESIGN This section provides additional details concerning the design of the Technical Specification leakage detection components.

Containment Cavity and Floor Drain Sumps l

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 g

leakage alarm with a setpoint of I gpm is provided in the main control room.

E The accuracy of these loops is 0.63% or 0.063 gpm for the reactor cavity sump and 0.093 gpm l

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 l

leakage rate of I gpm.

The calibration frequency for these loops is 18 months. All t

j mstrumentation for the reactor cav.ity and floor drain leakage collection sumps is seismically l

designed and supported.

l I

SL-4519

~ I SARGENT S LUNDY SL-4519.

0F12-89 I

5-3 I

I Containment Particulate and Gaseous Radioactive Monitoring Systems These systems are not designed to identify leakage from the main steam system because radiation levels in main steam would not be sufficient to exceed the radiation level setpoints of the systems.

l Reactor Containment Fan Cooler' Under normal conditions the radiation level in the main steam system is sufficiently low to preclude detection of a leak in a main steam line via the containment atmosphere particulate radioactivity monitoring system or the containment gaseous radioactivity monitoring system.

However, the containment floor drain sump will identify a i gpm leakage through a main steam line. The preceding discussion on containment cavity and floor drain sumps concerning this containment floor drain sump also applies to main steam leakage.

I Additional indication of leakage from a main steam line would be provided to the main control room by the reactor containment fan cooler (RCFC) inlet and outlet temperature and dewpoint monitors. The RCFC inlet and outlet temperature is also indicated on the analog computer. This information will identify to an operator that excessive steam has been released to the containment volume. The operator can then use the reactor building floor drain sump flow rate to quantify the leakage rate.

- CONCLUSIONS I

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 I

Byron /Braidwood main steam system. Therefore a leakage rate of 10 gpm has been used for the LBB analysis of this report.

i l I

SL-4519 5

SARGENT S LUNDY SL_4319 05-12-89 6-1 I

i Section 6 j

i LEAK RATE CRACK SIZING q

I The Byron /Braidwood LBB main steam program used the EPRI computer code PICEP (Pipe j

Crack Evaluation Program) (Reference 9) 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 ca'culates the leak rates for given crack sizes to form a crack-size-versus-leak-rate I

curve. The LSC is the crack length at a 10 gpm leak rate.

The PICEP code calculates the crack-opening area and the corresponding flow rate for non-extending through-wall pipe cracks. The crack opening area is computed using the elastic-plastic estimation scheme validated in Reference 9. The leak rate computational thermo-hydraulic models were validated against the available measured test results, in Appendix C of the same reference.

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

The dead weight, normal operating thermal and pressure loads were combined report.

algebraically on a component basis.

The PICEP code also required the material true stress-true strain property, at the corresponding normal operating temperature expressed in the RAMBERG-OSGOOD format, to determine the crack opening area needed for the leak rate computations. The results of EPRI Report NP-5057 (Reference 6) 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 sinc.e 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 I

sL-4519 I

i w_

l I

SARGENT E LUNDY SL-4519 05-12-89 6-2 gi for the base metal, as determined in Section 3 of this report were used for a conservative l

estimation of the LSC.

l I'

l The results of the leakage rate calculations are presented in Figures 6-1 and 6-2 for

]

circumferential and longitudinal cracks.

These figures present the leakage rates as a function of the crack size. The LSC is determined from these figures as the crack size l

1 allowing a 10 gpm leak rate. Table 6-1 presents the LSC length for each system in the l

Byron /Braidwood LBB scope.

I I

I E

lI I

I I

I s<-.,,,

I

I SANGENT & LUNDY Leakage Size Crack Length e 6-1 05-12-89 I

l W

Circumferential Crack Orientation:

E Line Thickness LCS E

System Line Size (in)

(in)

I MS 30.25 1.250 14.245 I.

MS 32.75 1.343 14.041 Axial Crack Orientation:

MS 30.25 1.250 10.597 g

5 MS 32.75 1.343 10.944 I

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lll

SARGENT E LUNDY SL-4519 05-12-89 7-1 Section 7 CPACK STABILITY ANALYSIS The Byron /Braidwood LBB main steam 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 further crack extension. EPRI studies (Reference 10) have justified the use of Z-fictors modifying *the limit load approach to conservatively evaluate crack stability by clastic-plastic fracture mechanic methods.

The Z-factor is a reduction factor which cccounts for the reduced fracture toughness in carbon steel and flux welds.

For circumferential cracks, master curve methodology (Reference 11) 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 11) has been used by incorporating the Z-factors from the proposed draft for Evaluations of Flaws in Ferritic Piping, Appendix Z to ASME Code Section XI issued in January,1988 (Reference 12). In addition to the master curve evaluation, a 3-Tearing cvaluation was also performed using a procedure first suggested by Zahoor (Reference 13).

The EPRI computer code FLET (Reference 14) was used to perform this evaluation using the j

1:wer bound carbon steel material properties identified in Section 3.

MASTER CURVE CONSTRUCTION The master curve was constructed by plotting the stress index,51, given by 7

j SI = 5 +MP (7-l) m sL-4519

SARGENT& LUNDY SL-4519 05-12-89 k,,

7-2 L

g I

as a function of the crack length, L defined as I

L=2eR (7-2) wher'e S=2 f(7-3}

[2 sin s - sin 0],

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

=

the pipe mean radius in inches R

=

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 cf

=

if e + 8 is greater than w, then S= 2af (7-#)

[ sin s ]

with 8 defined as a = - 3 (P / f)

(7-6) m I

s<..,,,

I

I

{

SARGENTS LUNDY SL-4519 05-12-89 I

7-3 I

I The 51 values are dependent on the material flow stress, of, therefore a master curve was generated for base metal and weld metal using the applicable flow stress. The flow stress y + o ). The lower bound material used to construct the master curve is defined as of = } (o u

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

I Carbon Steel Base 52325 psi Carbon Steel Weld 62278 psi Because the Si values are also dependent on the load margin factor, M, a master curve was l

generated for M=1.0 and M=1.4.

The master curves constructed with Mel.0 are used to establish the margin on crack size while the master curve constructed with M=1.4 are used to establish the margin on load.. The P values used in the construction of the master m

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

I The master curves constructed for the Byron /Braidwood LBB program are presented in Figures 7-1 through 7-8.

MASTER CURVE METHODOLOGY After construction of the master curves, actual stress indexes were calculated for the base metal and weld metal of the Byron /Braidwood LBB main steam 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).

I For carbon steel base metal and weld metal, the ASI is m+P+P)2 (7-7)

ASI = M (P b

e l

I sL-4519 i

SARGENT & I.dNDY SL-4519 05-12-89 7-4 I

where 0.46

. 958M [1.0 + 0.021 A (NPS-4)]/o (7_g) g y

2 21050 in-lb/in,

for carbon steel base metal with 31C or I

0.46 Z = 2.566 M [i.0 + 0.0184A (NPS-4)]/o (7_9) g y

I for carbon steel SMAW/SAW with 350 s 3ic < 600, A = [0.125 (Rit) - 0.25] 0.25 for 5 s R/t s 10, or A = [0.4 (R/t) - 3.0]O.25 ior 10 < R/t s 20, and Mg = og/S. Z must be greater than or equal to M /2.4.

m g

NPS is the nominal pipe size and t is the nominal piping wall thickness. The yield stress used to calculate the carbon steel Z-factors was the lower of the actual o or 40 ksi. The Z-factor y

(Referer.ce 10) accounts for the reduced fracture toughness of carbon steel base metal and E

flux welds, which are not expected to reach full plasticity at failure. For this reason the E

secondary stresses, P,, may not be relieved as failure approaches, and therefore are included in the stress index for flux welds.

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, I

s<..,,,

I

1 i

I SARGENT E LWDY st_4319

)

03-12-89 7-5

.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 main steam system.

Table 7-2 presents the ASI using M=1.4 to assess the margin on loads. The 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 main steam system.

I 3-T CRACK STABILITY EVALUATION Because the application of a modified limit load approach to carbon steel flaws is still under I

development, a 3-T analysis on the postulated carbon !

, cracks was performed to confirm the master curve resu'lts. The EPRI computer code FLET (Reference 14) was used to perform this evaluation. It uses the EPRI/GE 3-estimation scheme to compute the 3 parameter associated with the LSC length and the applied load. As defined by Zahoor (Reference 13),

the tearing modulus, T, is computed for the applied load and the LSC ngth by calculating the change in 3 caused by an incremental change in crack length using the following definition for T.

M" (7-10) 2 a

"f I

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

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 7-10. Once the material property 3-T curve was established, the intersection point I

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 I

sL-4519 IE

SARGENT 5 LUNDY SL-4519 05-12-89 7-6 I

assessment values of T that are less than the T at instability. As recommended (Reference 6), this procedure was conducted using the lower bound true stress-true s+ rain data for the base metal with the lower bound 3-Resistance data for weld metal. The loads used are the algebraically combined deadweight and normal operating thermal absolutely summed with the SSE and SAM loads from Section 4.

Figures 7-9 and 7-10 present the 3-T plots using the load margin factor of 1.0 on the applied g

load and the crack margin factor of 2.0 on the applied LSC for the 30.25-inch-diameter and a

32.75-inch-diameter main steam lines, respectively. Figures 7-!! and 7-12 present the 3-T l

plots using the crack margin factor of 1.0 on the applied LSC and the load margin factor of 1.4 on the applied load for the 30.25-inch-diameter and the 32.75-inch-diameter main steam lines, respectively. These plots demonstrate the stability of the postulated LSC for both the required margin on crack size and the required margin on loads.

Table 7-3 lists the 3-T assessment and the 3-T instability values for the main steam lines.

MARGIN FOR AXIAL CRACKS The axial crack stability evaluation was performed to assure that the required margin on size E

and load are met.

This evaluation was performed using the PICEP computer code to E

calculate the CCL. PICEP uses the empirical formulation for axial crack failure presented in (Reference 15).

In evaluating crack stability for an axial crack onentation, only pressure loads are significant and were the only loads used for this evaluation.

Lower bound elastic properties from Section 3, and 2-factors previously identified in Section 7 were used in PICEP to calculate the CCL.

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.

sL-4519 I

i SARGENT5 LUNDY circumferential crack Size Table 7-1.

Margins SL-4519 0F12-89 I-Margir.

System M

CCL LSC (CCL/LSC 2 2.00)

M5$:30 Base Metal 26,340 31.784 14.245 2.23 Weld 27,717 34.478 14.245 2.42 I

MS $-32 l

Base Metal 31,644 30.02 14.041 2.14 Weld 33,412 33.123 14.041 2.36 l

I-I I

lI LI I

SL-4519 I

I SARGENT & LUNDY Circumferential Crack Table 7-2 Load Margins SL-4519 I.

05-12-89 Margin System ASI CCL LSC (CCL/LSC d 1.00)

MS D=30 l

Base Metal 36,876 25.440 14.245 1.79 l

Weld 38,803 28.463 14.245 2.00 MS D=32 Base 44,302 21.866 14.041 1.56 Weld 46,777 25.447 14.041 1.81 lI

.I I

I I

I I

I I

I SL-4519

i-i SARGENT S LUNDY 3-T Results for M:in Ste m Table 7-3 SL-4519 03.12-89 I

Assessment Instability Crack Size Margin of 2:

3 (in-lb)

T 3(in-lb)

T g

2 in in OD = 32 ~

5500 13 6516 15.525 I

OD = 30 4500 12 6472 16.617 I

Load Margin of 1.4:

OD = 32 2100 6.8 6404 17.709 OD = 30 1500 5.1 6383 18.333 j!

I I

I 1

I SL-4519

SARGENTS LUNDY '

Axial Crack Size Margins Table 74 SL4519 ~

03-12-89 CCL LSC Margin System (inches)

(inches)

(CCL/LSC 2 2.00)

MS (30")

34.32 10.597 3.24 MS (32")

36.69 10.944 3.35 1

i i

i i

i I

SL-4519

~

SARGENTELUNDY Axial Creck Load Margins Table 7-5 SL-4519 05-12-89 I

l CCL LSC Margin System (inches)

(inches)

(CCL/LSC 21.00) l MS (30")

23,95 10.597 2.26 MS (32")

25.59 10.944 2.34 I

I I

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SARGENT & LUNDY SL-4519 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, i

Rev.1, March 1984 1

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

Subject:

Response to I

NRC Bulletin 88-08 and 88-08: Supplement I and 2 l

5.

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

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

6.

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

7.

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

8.

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

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

9.

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

10.

Evaluation of Flaws in Ferritic Piping, EPRI NP- 045, October 1988.

11.

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

12.

IWB-3650, Acceptance Criteria for Flaws In Ferritic Pipin t and Appendix Z, Evaluation of Flaws in Ferritic Piping; Blue Cover - Final 3 raft No. 2, January,1988, 13.

A. Zahoor and R. M. Gamble. Evaluation of Flawed Pi:>e Experiments. Palo Alto, Calif.: Electric Power Researcl Institute, November 1986. NP-4883M.

14.

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

sL-4519

SARGENT & LUNDY SL-4519 05-12-89 8-2 l

15.

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

International Journal of Pressure Vessels and Piping, Vol. 4 No. 4, October 1976, pp. 253-283.

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