ML20236C791

Nonproprietary Technical Bases for Eliminating RHR Line Rupture from Structural Design Basis for Vogtle Unit 2

ML20236C791
Person / Time
Site:	Vogtle
Issue date:	10/31/1987
From:	Chan A, Palusamy S, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
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ML19304B639	List: ML20236C730 ML20236C752 ML20236C791
References
WCAP-11600, NUDOCS 8710270335
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'

WESTINGM!USE PR!PRIETARY CLASS 3 -

WC' P-11600-A 1

y bJ 4

TECHNICAL BASES FOR ELIMINATING-RHR LINE RUPTURE FROM THE STRUCTURAL-DESIGN BASIS FOR V0GTLE UNIT 2 October 1987 A. C. Chan S. A. Swamy i

D. H. Roarty Verified byi K.

C.' Chang y

. Approved by:

M

/ 5. 5. A lusamy, Manager Struptural Materials Engineering Work Performed under Shop Order GHFJ6502F WESTINGHOUSE ELECTRIC CORPORATION

.j Generation-Technology Systems Division

'

• 1-P.O. Box 2728

.Pittsburgh, Pennsylvania 15230-2728 8710270335 871016 aut. ieimno j:

PDR ADOCK 05000425-A-

'i PDR; L

-- -- --------------------- --- - --- - ------ ~ ---- "

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[ --

.j TABLE OF ' CONTENTS 4

. Section Title-Page-I

1.0 INTRODUCTION

1-1.

l 1.1 Background-1-1 l'. 2 Scope and.0bjective 1-1 1-4 1.3 References 2.0 FAILURE CRITERIA FOR FLAWED PIPES 2-1 2.1 General Considerations 2-1 1

2.2 ' Global' Failure Mechanism 2-1.

2.3 Local Failure Mechanism 2-2' j

l 2.4 ' References 2-3 3.0 OPERATION AND STABILITY OF THE RHR LINES 3-1 1

3.1 Stress Corrosion Cracking-3-1

~

3.2 Water Hammer 3-3 3.3 ' Low Cycle and High Cycle Fatigue 3-4

)

3 '. 4 Potential Degradation During Service 3-4 3.5 ' Assessment of Pipe Degradation Failure from Indirect Causes 3 3.6 References 3-6 4.0 MATERIAL CHARACTERIZATION 4-1 l

4.1. Pipe, Fittings and Weld Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-3 4.4 References 4-4 j

I 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 5-1

^

5.1 Loads for Crack Stability Analysis 5-2 o

5.2 Loads for Leak Rate Evaluation 5-2 5'. 3 Summary of Loads Geometry and Materials 5-2 5.4 Governing Location 5-3 nu.aotamo 3$

)

aJ-

l TABLE OF CONTENTS (cont.)

r Section Title Page v

l l

6.0 FRACTURE MECHANICS EVALUATION 6-1 >

6.1 Global Failure Mechanism 6-1 6.2 Leak Rate Predictions 6-2 6.2.1 General Considerations 6-2 6.2.2 Calculation Method 6-3 6.2.3 Leak Rate Calculations 6-4 6.2.4 Leak Detection Capability, Administration 6-4 Procedures and Technical Specification i

Requirements 6.3 Stability Evaluation Using the "Z" Factor 6-6 Approach 6.4 Local Stability Analysis 6-7 6.4.1 Crack Extension Consideration 6-8 6.5 References 6-8

7.0 ASSESSMENT

OF FATIGUE CRACK GROWTH 7-1 7.1 Acceptability of Fatigue Crack Growth 7-2 7.2 References 7-3

8.0 ASSESSMENT

OF MARGINS 8-1

9.0 CONCLUSION

S 9-1 APPENDIX A Limit Moment A-1 APPENDIX B Fatigue Crack Growth Considerations B-1 B.1 Thermal Transient Stress Analysis B-1 B.1.1 Critical Location for Fatigue Crack Growth Analysis B-1 B.1.2 Design Transient B-2 a.->eim io

n,,

T TABLE OF' CONTENTS.(cont.)'

f

. Sectio'n '

Titie Page B.'1. 3 -

Simplified Stress Analysis B-2 B.1.4 OBE Loads B-4 B.1'. 5 Total-StressLfor. Fatigue Crack Growth B.B;2-Fatigue Crack Growth Analysis

'B-5 B.2.11 Analysis' Procedure B-5 l

B.'2.2 Res'ults B-8 B.3 References B-8 i

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LIST OF FIGURES' 1

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. a, Figure Tit 1e

.Page' 1-1.

Vogtle Unit 2 RHR Line - Scope of Evaluation 1-5

2-1 Schematic of Generalized Load Deformation Behavior 2-4 4-1 True' Stress. Strain Curve for SA376 TP316 Stainless-

'4-8L

~ Steel at.617'F-(min'imum properties)

L4-2 True Stress Strain Curve for SA376 TP316 Stainless-4-9 Steel at 617'F-(average properties) i 4 J Versus'T Curves for.SMAW. Welds Showing i

Extrapolation

'4-10

~

5-l' Schematic Layout of RHR Lines Loop 1 5-6 H

. 5' 2 '

. Schematic Layout.of RHR Lines Loop 4 5-7 6 - 1.

- ['

Ja,c.e Stress Distribution 6-9 j

6-2.

Analytical Predictions of Critical Flow Rates of 6-10

. Steam-Water Mixtures 6-3;

[

la,c e Pressure Ratio as a 6-11 Function of L/D 6-4' I'dealized Pressure Drop Profile through a Postulated 6-12

^

Crack

!t 6-5 Loads Acting on.the Pipe Model at the Governing 6-13 Location 2457s-101387.10-y

y'

i;.

1

-LIST OF FIGURES-(Cont'd.)

Figure-Title Page' 6-6

. Critical' Flaw Size Prediction for the Base Metal 6 !

Using Limit. Load Approach

'6 Z-Factor Calculation for SMAW Weld to Demonstrate-

'6-15 Margin on Flaw' Size

'6-8 W-Factor Calculation for'SMAW Weld to Demonstrate

.6-16 Margin on Flaw Size

A-1.

Pipe with a Through Wall Crack in Bending A-2 l

B-l' Schematic.of RHR Line at (

B-13

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a LIST OF TABLES.

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q Table No..

Tit 1e

'Page.

41-Mechanical Properties of_ the.12 inch RHR'Line 4-5 j

- at 70.'F' as Determined by; Testing y

)

.4-2 Typical Tensile Properties of SA376 TP316,. SA351 CFA-4.]

i and Welds of Such Material for the. Primary' Loop

'4-3 Fracture Toughness Properties Typical. of. RHR Line -

4-7 l

5-1 Summary of Envelope Loads for<12 Inch Pipe-5-4 q

~

5-2

. Loading Components at Governing' Locations For 5-5 12 Inch Line

8 Comparison of Results vs. Criteria 8-3) 1 a

B Thermal Transients Considered for Fatigue Crack B-9 Growth Evaluation

)

B-2 Transient Stresses for RHR Line B-10

~B-3 Envelope Normal Loads B-11 B-4 RHR Line Fatigue Crack Growth Results B-12 l

.i p-h (c e' l

' asste totut.'to yjj m..........

i SECTION 1.0-

' INTRODUCTION sr

1.1 Background

.The current str_uctural design basis'for the RHR line requires postulating non-pachanistic circumferential and longitudinal pipe breaks. -This results~in

-additional plant _ hardware (e.g. pipe whip restraints and jet shields) which would mitigate'the dynamic consequences of.the pipe' breaks.

It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for these -

1

' lines, and thereby eliminate the need for some of the plant hardware.

Presented in this report are the descriptions of-a mechanistic pipe break

.j evaluation method and the analytical results that are used for establishing that a circumferential type break will not occur. The' evaluations considering circumferentially ' oriented flaws cover longitudinal cases.

'1.2.

Scope and Objective j

'The purpose.of this investigation is to demonstrate leak-before-break for the I

high energy' portion ~of-.the RHR line. The piping covered by this evaluation is shown in figure'1-1, and includes the high energy piping from the primary loop q

ijunction up to the first isolation valve. Down stream of the isolation valve, j

the piping is classified as moderate energy piping. Schematic draw'ings of the piping system are shown in Section 5.0.

The recommendations and' criteria proposed in NUREG 1061 Volume 3 (1-1) are used in this evaluation. These criteria and resulting steps of the evaluation procedure can be briefly summarized as follows:

l 1)

Calculate the-applied loads.

Identify the location at which the lf' highest stress occurs.

i l-2)-

Identify the materials and the associated material properties.

~

. m.-ioiano 11

s, 1

t

' 3)i Postulate a' surface flaw at the governing location with the least.

favorable combination of stress and material properties.. Determine ~

$ 1 fatigue crack growth.. Show that a through-wall crack will not tasult.

l

. ostulate a; through-wall: flaw at' the. governing location..The size j

4)

P offthelflaw should be large enough soLthat the leakage is. assured!of.

~

. detection with margin using.theLinstalled. leak detection equ;pment.

ll when the pipe is' subjected to normal" operating; loads. 'A margin of.

10'is demonstrated b'etween the calculated leak rate and the-leak ~-

detection capability. _ The associated flaw is ' called the. leakage size flaw.

j I

4 5). Using normal plus-SSE loads,. demonstrate that there is a~eargin 'of at least'2 between the leakage size flaw and the critical size. flaw.

Review the operating history to ascertain-that operating experience 6) has indicated noLparticular susceptibility to failure from the.

,j

' effects of. corrosion, water hammer and low and high cycle fatigue.:

7)-

For;the base and weld metals actually in-the plant' provide' the material properties including toughness and tensile test data.-

f cJustify that the properties used in:the evaluation are 1

representative of the plant specific material.

Evaluate long term effects such' as thermal aging where' applicable.

8)

Demonstrate margin of at least 1.4 on applied load for the leakage size' flaw.

The flaw stability criteria used in this analysis address both the global and l

i f

tial flaw. The

-. local stability for a postulated through-wall c rcum eren global analysis is carried out using the [

]a,c.e method,

,j

. based on traditional plastic limit load concepts, but accounting for [

f Ja c,e and taking into account the presence of a flaw (1-2).

a wr. ioint.io 1-2 a

.i..

l The local stability analysis is carried out using ' elastic plastic fracture mechanic analysis procedures.

In this report the EPRI elastic plastic fracture handbook method was used for the local stability analysis.

The leak rate is calculated for the normal onerating condition. The leak rate l

prediction model used in this evaluation is an [

]a,c.e The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through wall flaw to normal operating loads.

Surface roughness is accounted for in determining the leak rate through the postulated flaw.

As stated earlier, the evaluations described above considering circumferen-tially oriented flaws cover longitudinal cases in pipes and elbows.

The likelihood of a split in the elbows is very low because of the fact that the l

elbows are [

]a,c,e and no flaws are actually anticipated.

The l

prediction methods for failure in elbows are virtually the same as those for

[

Ja,c,e Therefore, the probability of any longitudinal flaw existing in the RHR line is much smaller when compared with the circumferential direction.

Based on the above, it is j

judged that circumferential flaws are more limiting than longitudinal flaws in elbows and thi oughout the system.

Several computer codes are used in the evaluations. The main-frame computer programs are under Configuration Control which has requirements conforming to Standard Review Plan 3.9.1.

The fracture mechanics calculations are independently verified.

ik nu..ioinna 1-3

f '1.3

References:

1-1. Report <of the U.S. Nuclear Regulatory Commission' Piping Review Committee j

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

j l' NUREG/CR-3464, 1983, "The Application.of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated.

t Circumferential Through' Wall' Cracks."

j

1-3 'Begley,'J. A.. et, al.,'" Crack Propagation Investigation Related to the Leak-Before-Break Concept for LMFBR Piping" in Proceedings,-Conference on.

Elastic Plastic Fracture, Institution of Mechanical Engineers, London 1978.

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-ISOLATION. VALVES

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MODERATE ENERGY PIPING Figure 1-1 Vogtle Unit 2 RHR Line Scope of Evaluation mr. ioim.ie '

15

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SECTION 2.0 FAILURE. CRITERIA FOR FLAWED PIPES w

(

1 2.1 General Considerations

]

Ac'tive research is being carried out in industry and universities as well as

)

other research organizations to establish fracture criteria-for ductile materials.

Criteria being investigated include those based on'J-integral initiation toughness, equivalent energy, crack opening displacement, crack opening stretch, crack opening angle, net-section yield, tearing modulus and f

void nucleation. Several 'of'these criteria are discussed in an ASTM publication (2-1).

{

A practical approach based on the ability to obtain material properties and to make calculations using the' available tools was used in selecting the criteria

.q

.for this investigation. The ultimate objective is to show that the RHR line

' containing a conservatively assumed circumferential through-wall flaw is stable under the worst combination of postulated faulted and operating j

condition loads within acceptable engineering accuracy. With this viewpoint, two. mechanisms of failure, namely, local and global failure mechanisms are considered.

2.2 Global Failure Mechanism t

For a tough ductile material which is notch insensitive the global failure will be governed by plastic collapse.

Extensive literature is available on this subject. A Pressure Vessel Research Committee study (2-2) reviews the literature as well as data from several tests on piping components, and discusses the details of analytical methods, assumptions and methods of j

correlating experiments and analysis.

A schematic description of the' plastic behavior and the definition of plastic load is shown in Figure 2-1.

For a given geometry and loading, the plastic load is defined to be the peak load reached in a generalized load versus displacement plot and corresponds to the point of instability.

1 2657s-101387.10 2-1

\\

l A simplified version of this criterion, namely, net section yield criterion has been successfully used in the prediction of the load carrying capacity of pipes containing gross size through-wall flaws (2-3) and was found to correlate-well with experimont.

This criterion can be summarized by the following relationship:

Wa < Wp (2-1) where Wa = applied generalized load Wp = calculated generalized plastic load Wp represents the load carrying capacity of the cracked structure and it can be obtained by an elastic plastic finite element analysis or by empirical correlation which is based on the material flow properties as discussed in Section 6.1 2.3 Local Failure Mechanism The local mechanism of failure is primarily dominated by the crack tip behavior in terms of crack-tip blunting, initiation, extension and finally crack instability. The material properties and geometry of the pipe, flaw size, shape and loadings are parameters used in the eve.luation of local failure.

The stability will be assumed if the crack does not initiate at all.

It has been demonstrated that the initiation toughness, measured in terms of JIc from a J-integral resistance curve, is a material parameter defining the crack initiation.

If, for a given load, the calculated J-integral value is shown to be less than J f the material, then the crack will not initiate.

Ic If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:

T,pp = h (2-2) 22

i applied 't' earing - modulus

'i

~ where T,pp.

==

modulus of elasticity

.E

=

u)/2

.- of flow' stress-= (o +

=

y a

= '

crack length c,: ou '=.

yield and ultimate strength of the material y

respectively.-

i In summary, the local crack stability is established by the two-step criteria:

(2-3)

~J < Jg, or.

3 Ie (2-4)

J T,pp Tmat, if J 2.4 References 2-1 J.D. Landes, et al., Editors,. Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia, PA 19109, November 1977.

2-2 J. C.' Gerdeen, "A Critical Evaluation of Plastic Behavior. Data and a

. Unified Definition of Plastic Loads for. Pressure Components," Welding l

Research Council Bulletin No. 254.

l 2-3 Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks, EPRI-NP-192, September 1976.

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P W,= PLASTIC LOAD l

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AP GENERALIZED DISPLACEMENT Figure 2-1 Schematic of Generalized Load-Deformation Behavior ais..ioenuo 2-4

SECTION 3.0 OPERATION AND STABILITY 0F THE RHR LINE AND THE REACTOR. COOLANT SYSTEM i

3.1 Stress Corrosion Cracking The Westinghouse reactor coolant system primary loop and connecting Class 1 lines have an operating history that demonstrates the inherent stability characteristics of the design.

This includes a low susceptibility to cracking f ailure from the effects of corrosion (e.g., intergranular stress corrosion cracking). This operating history totals over 400 reactor years, including five plants each having over 15 years of operation and 15 other plants each with over 10 years of operation.

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

(The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pressurized Water Reactors (PWR's). The results of the study performed by the PCSG were presented in NUREG-0531 (Reference 3-1) entitled " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Light Water Reactor Plants."

In that report the PCSG stated:

"The PCSG hhs determined that the potential for stress-corrosion cracking in PWR primary system piping is extremely low because the ingredients that produce IGSCC are not all present.

The use of hydrazine additives and a hydrogen overpressure limit the oxygen in the coolant to very low levels. Other impur.ities that might cause stress-corrosion cracking, such as halides or caustic, are also rigidly controlled. Only for brief periods during reactor shutdown when the coolant is exposed to the air and during the subsequent startup are conditions even marginally capable m7.coner io 3_1

of producing stress-corrosion cracking in the primary systems of PWRs.

. Operating experience' in PWRs supports this determination. - To date, no "j

stress-corrosion cracking has been reported in the primary piping or safe

=

ends of any PWR."

L During 1979, several instances of cracking in PWR feedwater piping led to the establishment of the-third PCSG. The investigations of the PCM reported in 1

NUREG-0691 (Reference 3-2) further confirmed that no occurrences af IGSCC have been reported for PWR primary coolant systems.

As stated above, for the Westinghouse plants there is no history of cracking failure in the reactor coolant system loop or connecting Class 1 piping. The discussion below further qualifies the PCSG's findings.

i.

'For stress corrosion. cracking (SCC).to occur in piping, the following three j

conditions must exist simultaneously:

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

degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized by properly selecting a material immune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other material'in the system, applicable ASME Code rules, fracture toughness, welding, f'abrication, and processing.

i The~ elements of a water environment known to increase the susceptibility of j

austenitic stainless steel to stress corrosion are: oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g.,

sulfides, sulfites,andthionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment.

Prior to being put into service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.

av..mno 3-2

During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the major water chemistry control standards being included in the plant operating procedures as a condition for plant operation.

For example, during normal power operation, oxygen concentration in the RCS and connecting Class 1 lines is expceted to be in the ppb range by controlling charging flow chem-istry and maintaining hydrogen in the reactor coolant at specified concentra-tions. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and fluorides within the specified limits. Thus during plant operation, the likelihood of stress corrosion cracking is minimized.

3.2 Water Hammer Overall, there is a low potential for water hammer in the RCS and connecting RHR lines since they are designed and operated to preclude the voiding condition in normally filled lines.

The RCS and connecting RHR lines including piping and components, are designed for normal, upset, emergency, and faulted condition transients.

The design requirements are conservati'e v

relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design.

Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads.

To ensure dynamic system stability, reactor coolant parameters are -

stringently controlled.

Temperature during normal operation is maintained within a narrow range by control rod position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process.

Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system and connecting RHR lines.

Preoperational testing and an.uow to 3-3

I operating'experienceshave verified the Westinghouse approach. The operating.

transients of_the RCS primary-piping and connected RHR lines are such that no

_ significant water hammer can-occur.

3.3 Low Cycle and High= Cycle Fatigue

- Low ^ cycle fatigue considerations are accounted lforLin'tho' design of?tha piping system through' the l fatigue usage factor evaluation' to' 'show compliance with the

~

rules of Section III of the ASME Code. A further. evaluation of the low cycle-l fatigue loading is discussed:in Chapter 7'as part_'of this_ study in the form.of'

a fatigue crack growth analysis.

4 High cycle fatigue loads in the~ system would. result primarily from pump -

vibrations during operation. During operation, an' alarm signals the exceda'nce

- of the RC pump shaft vibration limits.

Field measurements have been made on-the reactor coolant-loop piping of a' number of plants during hot functional

- testing. Stresses in the elbow below the RC pump have been found to be very small,.between 2 and 3.ksi at'the. highest.. When translated to the. connecting i

RHR: lines, these stresses are even. lower, well below the fatigue enduran'ce l

limit for the RHR line material'and would result in an applied stress

. intensity factor below the threshold-for fatigue crack growth.

'34 Potential Degradation During Service l

In the Westinghouse PWR design there has never been any service cracking identified in the RHR piping. Only one incident of wall thinning has been identified in RHR lines of Westinghouse PWR design. However, this is of no concern in the present application as described ~later in this section.

Sources;of such degradation are mitigated by the design, construction.

inspection, and operation of the RHR lines.

Based on a' review of references 3-3 through 3-6 only one incident of water hammer has been reported in a PWR RHR system.

This incident was a result of incorrect valve line-up preceeding a pump start. The only damage sustained l

+

-=>.-ioim.io 3-4

was to several pipe supports.

Therefore we conclude that water hammer in the RHR system is unlikely to affect piping integrity or to cause pipe system degradation.

1 Wall thinning by erosion and erosion-corrosion effects will not occur in the RHR line due to the low velocity, typically less than 10 ft/sec and the I

material, austenitic stainless steel, which is highly resistant to these degradation mechanisms.

Per NUREG-0691 (3-2), a study of pipe cracking in PWR piping, only two incidents of wall thinning in stainless steel pipe were reported. One incident was related to the RHR system.

However, this occurred j

in the pump recirculation path which has higher flow velocity and is more susceptible to other contributing factors such as cavitation, than the RHR piping near the primary loop. Therefore, wall thinning is not a significant concern in the portion of the system being addressed in this evaluation.

l i

Flow stratification, where low flow conditions permit cold and hot water to separate into distinct layers, can cause significant thermal fatigue loadings. This was an important issue in PWR feedwater piping where temperature differences of 300*F were not uncommon under certain operational conditions. Stratification is believed to be important where low flow conditions and a temperature differential exist. This is not an issue in the RHR line, where typically there is no flow during normal plant operation.

During RHR operation the flow causes sufficient mixing to eliminate stratification.

The normal operating temperature of the RHR piping is about 617'F.

This is well below the temperature which would cause any creep damage in stainless steel piping.

3.5 Assessment of Pipe Degradation or Failure from Indirect Causes Appropriate protection against the potential of pipe degradation or failure from indirect causes is provided by plant design features and by the implementation of structure, system, and component design, fabrication and inspection requirements as specified in the design basis.

These features and nu..ioisano 3-5

l.

3..^

i requirements are consistent with those specified in the Standard. Review Plan as discussed in the Plant Vogtle FSAR as follows:

Flood protection is discussed in FSAR Section 3.4.1.

J

. FSAR Section 3.5 describes how protection is provided against internally l

generated missiles both inside the outside containment.

l Fire protection is discussed in Section 9.5.1 of the FSAR.

FSAR Sections 3.9.B.3 and 5.4 discuss in detail the design and fabrication requirements of Class ~1, 2 and 3 components and component supports.

In-service inspection and testing of the reactor coolant pressure boundary is covered in FSAR Section 5.2.4 and general discussions of inspection are

- provided in FSAR Sections 17.1.14 and 17.2.10.

It can be concluded by review of these sections of the Plant Vogtle FSAR that

- the required measures are taken to preclude the degradation or failure from 3

outside. sources of piping in the plant and that the methods utilized are consistent with those given in the' Standard Review Plan.

4 l

3.6 References

/

3-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of

]

. Light Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979.

! 2 Investigation and Evaluation of Cracking Incidents in Piping in i

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

3-3 Utter, R. A., et. al., " Evaluation of Water Hammer Events in Light Water l

Reactor Plants," NUREG/CR-2781, published July 1982.

an. ioinuo 3-6 l

1

,i 3-4

" Report of the U.S. Nuclear Regulatory Commission Piping Review Committee,. Evaluation of Other Dynamic Loads'and Load Combinations,"

NUREG-1061 Volume 4, Published December 1984.

l i

3-5. Chapman, R. L., et. al., " Compilation of Data Concerning Known and-Suspected Water Hammer Events in Nuclear Power Plants, CY'1969-May 1981,"

NUREG/CR-2059, Published April'1982.

3-6 " Evaluation of Water' Hammer Occurrence in Nuclear Power Plants,"

NUREG-0929 Revision 1, Published March 1984.

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. MATERIAL CHARACTERIZATION' a

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4.1 Pipe, Fittindi'and Weld Materials.

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,y The' pipe matorialj of: the RHR :line'is' S'A' 376-TP316, - a wrought " product form of:

ithetype~used3$r'.theprimky'looppipingofseveral.PWRp1 ants.- The RHR lineG' is: connected to 'the primary loop nozz1'e3 kThe other and of the RHR line:is

~

'1 '

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connected to 'an1 isolation valve. The piping layout includes fittings such as - ik d

elbows.tTheseelbowsareSA4d3:WP316steelwhichiswroughtandformedpipe offSA182-TP316 material. LThe' weld wire used in the shop fabrication ~is

~

generally ofilow carbon 316L.

In"some. casgs-316 weld wires were also used.

7 j

l, LThe field l weld ~used:308L weld wir'..The welding. processes used are gas q

e tungsten (arc.(GTAW), submerged arc (SAW).an'd shielded metal are'(SMAW).

l u

v.

.f

. w c.

In,the.following sections the ' tensile and fracture Ltoughness~ properties 'of.l[

these; materials:are' presented and criteria for'use in the leak-before-break 4 3p a

l

. analyses'are. defined, 4q 3 "", ~, o y i_4 y

n m

.<r

~4.2 Tensile Properties h

n LThe material certifications for the RHR line'were used to w&o'iirh the '

tensile properties for the piping, fittings and welds.-. Thhye prcperties are

/ 0 P

F

/

• 1 given in Table 4-1.

The properties in this table are at hoop temperature.

In i

4 I

the leak-before-break evaluation presented in tids re,p,orto the minimum s-properties at. operating temperature (617'F) are used fc+ the flaw stability-L'

' evaluation a'nd' average properties are used for the leak rate predictions.- The viability lof using such properties for the RHf linC, is p.'esented below.

(

^ As' noted in Table 4-1, _the specific ~ room temtje tature propertNs of the RHR f,

'line. heats compare favorably with the properiies of similar material of the c.:

primary loops- (see. table 2).

{

p,c.e

'l l

M5?s 101M7.10 41

' 5 - 2

______..____j

,,,i#

.q_.;.. f* t_ '

-$ l?,

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s i g 'y ~

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l[

. f..

,a

-p'

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fly, Y

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)

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' ar i q :

.f,g

-: J g

(( V Ci;

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) O f-t ' ;

,\\,.

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't\\.

.s t g1 Q. -V

.e j ' '.

(ff p'

i H

.(

y ja,c.e-4 Jl.

..f m,...

In bp}[of the following materiali/roperEih nre use'd in the t

fort in this report.

-. ?

3

.u,

.,' ((

c A.

e

~

yinift. m Properties for Flaw Stability Analysis'(617'F) k - }

9.f a,c.e

a..

9

~,

.j b /,

I

..,f n

\\

e v

.,v, y

a 1

/

s s I

i

-1

AveragePropertiesforLeakRateCgu'Jations-(617'F) t i

s a,c.e 1

i.

(

,) -

0,I d/,',7,.

/

j

? r.f

'r

. i[.

f d

ll

((

'l:

, ;t > t

's'L j

<,/

il f"

f-

. nir.ooman 4-2 y

1

' i-b

%nwg,g M.C M

-1

>J f

m 9 ; ;. ;a.,

>3 w~

J

'Q t

<i t

f<

k.

ij

'I

?.h[k@&,j!

l ry NN 4

k _ ;14 L

3 1

, (,

er n 9

g

4.3 LFracture joyghn p Proper}ies

+

y f

T?

].

)

t l

Series ofitacturftbughness tests on' SA376-TP316 p'ipe nshrial and E308 welds t

t garp repor ud:in references 4-2 and 4-3.

Data from thespitsts at 600'F are o

~.

value for summ{riye; 'in table ~4;3. As seen from t!is-tabl,a, the lowest JIe g

\\

o ti thdwehkmaterials;observedfromthe-J-Rcurvesias'[.

pc.e 2

' j in-lb in.. The Tgy corresponding to this value,'of J;c.was~ Inter, !\\

.[' 7] % % The sample yielding the lowest siode of the J-it curve had a

T f( :.JM .

The corresponding J was found to be' Ocund.

at Ic 2

[

']a,c.e in-lb/in..While the data on welds cited' above Show superior.

s toughne'ss pr6 peri.ies,. they were 'not-used in the' leak-beforedeqak evaluation..

Instead,Tiower.beund toughness data for the SMAR welds (noting that the. weld.

q y

3 ist the gover:Mg locytion'was 'SMAW) were thed

'itilows:

3

,(-

t

.4 -

j

=

s The data on SMAW:dids based on testing IT specimer n,reasented in reference a

3 3

g

+

4-4 can be c]nsidered on loner bosno thughness date. The resulting test data

,jj$reasfollows-

~

(

A 2

jp t

/

3 Ic 959in-}b/in

^

J

=

y

\\

\\

\\

2 3000 in-lb/in J

~

max i

f Tmat = 140(correspondingtoJgp 9

/

s{\\

t.

t The J versus T for the data is, plotted"in figure 4-3.

Thesedlatacanbe-Basedonextrapolatibsf(aluesofJasdescribedinNUREG1061 extrapolated to higher v 2

lf J value of 6000 in-lb/in is found to be 88.

The maximum recommended extrapolation is 2't$mes the highest J level where valid data are available.

l The lower bound SMAW weld fracture toughness for crack stability evaluation j p be summarized from figure 4-3 as follows:

(

o s

2 N'

J

= 959 in-lb/in

i IC g

3

\\

\\g

\\

i b

"'f**f g

r

)'

x 4-3,

. m i

2 mat = 140 corresponding..to J of 3000 in-lb/in T

2 Tmat = 118 corresponding to J of 4000 in-lb/in 2

mat = 102 corresponding to J of 5000 in-lb/in T

The nozzles connecting.the RHR Lines with the primary loop are made of forged stainless steel. Forged stainless steel is considered not susceptible to thermal aging for applications at hand.

4.4 References F. J. Witt et a,1., " Integrity of the Primary Piping System of 4-1 Westinghouse Nuclear Power Plants During Postulated Seismic Events,"

WCAP-9283, March 1978.

L 4-2 S. S. Palusamy, " Tensile and Toughness Properties of Primary Piping Weld Metal for Use in Mechanistic Fracture Evaluation," WCAP 9787,'

.May, 1981 (Westinghouse Proprietary Class 2).

4-3 S. S. Palusamy, et al., " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack," WCAP-9558, Rev. 2, May 1982,. (Westinghouse Proprietary Class 2).

t 4-4 Toughness of Austenitic Stainless Steel Pipe Welds, EPRI-NP-4768, Electric Power Research Institute, October 1986.

E e

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.1 TABLE 4-2

'I TYPICAL' TENSILE PROPERTIES OF SA376 TP316, SA351 CF8A and WELDS OF-i SUCH MATERIAL FOR' THE PRIMARY LOOP '

j Test Temperature _

Average Tensile' Properties f

Plant-Material

('F)

Yield (psi)

Ultimate (psi)

J A

SA376 TP316 70 40,900 (48)"

83,200(48) 650 23,500 (19) 67,900(19)

-E 308: Weld' 70 63,900 (3) 87,600(3) l B'

.SA376 TP316 70 47,100 (40) ~

88,300(40) 650 26,900 (22)

_69,100 (25)

~'

E 308 Weld 70 59,600 (8) 87,200 (8).

650 31,500 (1) 68,800_(1)

C SA376 TP316 70

'46,600 (36) 87,300 (36) 650 24,200'(18) 66,800 (19)-

E 308 Weld 70 61,900 (4) 85,400 (4)

D SA351 CF8A 70 47,300 (14) 84,500(14) 650 26,000 (4) 70,500 (4)

Weld 70 61,200 (31) 84,500(32)

a. (

) indicates the number of test results averaged.

i seu.noenne 4-6 j

e.

c, e.

a c,

a s

ta

~ m T

s

~

)

2 n i

c /

I b

J l

E N

n I

i A

L

(

RHR E

)

H i

e b

T s

t 00 0

p a

00 0

]

m 5,1, 2, -

F

(

O i

s t

50 1

L e

l 66 6

A i

U C

t I

r P

e Y

p 3

T o

r 4

S P

E b

E I

e 00 0

L T

l 00 0

B R

i d

7, 5, 0, -

A F.

s l

T P

n e

10 5

O e

i 22 4

R T

Y P

SSEN H

p G

m U

e

)

O T

F 00 00 0

T 00 00 0

t

(

66 66 6

E s

R e

U T

TCA R

F s

)

t 6

s i

e 3

t E

6 d

n f

a o

66 h

11 8

A t

33 0

8 s

PP 3

F e

la TT E

C w

i

(

o r

66 f

1 L

e 77 dd 5

t 33 ll 3

a AA ee A

ab M

SS WW S

[

?u

N

/

a.c.e 1

s 4

Y Figure 4-1.

True Stress Strain Curve for SA376 TP316 Stainless Steel at 617'F (minimum properties) nir.noint ie 4.g

I t

.q

i. [

[

.c

.c e

N u-

7.-

~

a,c.e

-1 i

l a

i i

't l

J I

/

Figure 4-2.

True Stress Strain Curve for SA376 TP316 Stainless Steel at 617'F (average properties)

)

I l

' mostanotut.io 4.g 4

.l 4'

i

~1 J

'6000'

\\

\\

)

\\

l

\\

I

\\

5000

\\.

.\\

\\\\'

o 4000

\\

\\

N

\\

g

\\

z

\\

3000 2000 i

1000 88

-l O

0 100 200 300 T

Figure 4-3.

J versus T Curve for SMAW Welds Showing Extrapolation mr.namno 4 10

,g.,..

b

(

1

.m LSECTION 5.0 v

, LOADS FOR: FRACTURE MECHANICS ANALYSIS' Figures.5-1 Land 5-2'are. schematic layouts of the RHR lines attached to the

reactor coolant loops l'and 4.-

o The itresses due to axial 1.oads and bending moments were calculated by the following equation:

o=[+{

(5.1)

where, o

=

stress axial load.

F

=

bending moment j

M

=-

metal cross-sectional area 1

A

=

2

.=

section modulus I

The bending moments for the desired loading combinations were calculated by-'

j the following equation:

]

M. = / h

+M

(*}

y Z

where,-

l'!

M bending moment for required loading

=

Y component of bending moment M

=

y 2 component of bending moment-

- ~-

.M

=

Z C

The axial load and bending moments for crack stability analysis and leak rate predictions were computed by the methods to be explained in Sections 5.1 and 5.2.

- nushamvo s.1

L 5.1 Loads for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the following equations:

F-lFDW + FTH + F l + IFSSE l (5.3)

=

p N

IIH )DW + (N )THI + IIN )SSEl (5.4)

Y Y

Y Y

I(N )DW * (N )TH I + IIN )SSE l (5.5) l M

2 Z

Z Z

l Where, the subscripts of the above equations represent the following loading

(

cases, DW deadweight i

=

TH normal thermal expansion

=

SSE SSE loading including seismic anchor motion

=

P load due to-internal pressure

=

5.2 Loads for Leak Rate Evaluation

~

The normal operating loads for leak rate predictions were calculated by the following equations:

F FDW.+ FTH + Fp (5.6)

=

N (N )DW + (N )TH (5.7)

Y Y

Y M.

(M )DW + IN )TH (5.8)

=

2 Z

2 5.3 Symmary of Loads, Geometry and Materials Table 5-1 provides a summary of envelope loads computed for fracture mechanics evaluations in accordance with the methods described in section 5.1, and 5.2.

j The cross-sectional dimensions and materials are also summarized. Load data

,j are tabulated at the highest stressed locatior. (node 2015, loop 4), and the I

second highest stressed location (node 2101, loop 4), for the 12 in, pipe.

The loading components are provided in table 5-2.

i 1

s.2 mwwuno i

t c

. 5.4'. Governing Location lThenormalplusSSEaxialstressesalongtheRHR'linestartingfromthe primary loop nozzle junction up to the first isolation valve were compared.

The maximum stress occurs at node 2015 on loop 4 RHR line. Based'on pipe load and materials properties, node 2015 was determined to be the governing location. The welding process at this location is-SMAW.

This location is identified in figure 5-2.. Detailed fracture mechanics analyses were performed at'this highest stressed location.

l

. e niwioinuo 5-3

k 4

4 4

0 8

9 8

n 2

1 1

1 i

(

)sp Fi 5

8 1

4 k

8 7

8 7

(

1 1

1 1

)

s E. e DAh 4

4 4

4 II c 7

7 7

7 SD n N'

i 0

0 0

0 I

(

1 1

1 1

)

N-s OLK e MLCh 5

5 5

5 I AI c 0

0 0

0 0

0 0

0 NWH n I

Ti M

(

1 1

1 1

)

L.

s E

ALK e N

NLCh 5

5 5

5 I

I AI c 2

2 2

2 L

MWH n 1

1 1

1 O

Ti R

N

(

1 1

1 1

HR E

H L

C U

N D

I E

H 0

0 0

0 2

C 4

4 4

4 1

S 1

1 1

1 RO 1

F

)

E s

5 S

D e

D I

h 5

5 5

5 E

A S

c 7

7 7

7 L

O tan B

L UI i 2

2 2

2 A

OD(

1 1

1 1

T E

PO L

E L

V A

6 n

N I

6 6

6 6

i E

R 61 61 61 61 E

73 73 73 73 5

F T

3P 3P 3P 3P 0

O A

AT AT AT AT 0

M S -

S -

S -

S -

Y 1

RA P

f O

o 4M O

L 4

4 4

4 s

lS sen k

E 5

5 1

1 c

0O 1

1 0

0 i

0N 0

0 1

1 h

2 2

2 2

t N

l la N

g g

w O

.n n

m I

d i

d i

T e

l t e

l t u

I t

aa t

aa m

mr mr i

D l

l N

u re u

re n

O a

op a

op i

C F

NO F

NO m

n N

n n

o O

t o

t o

I s

i s

i d

T e

t e

t e

A hda thda s

C gac x gac a

i oo eioo B

O L

HLL NHLL I

l

)

k

-n gi 6'

n(

3 8

0 i

dZ w

nN 1

1 8

e 1

1 3

)

Bt 1

4 n

e p

m do o

ao M

oL l

t1

)

s0 k

e1 h2 n

g gi 9

9 1

i -

n(

H -

i 8

7 7

n dY 1

3 1

E t o nN 8

1 N

xi e

1 I

et Bt L

Na n

- c e

H o

m C

L o

N

(

M I

2 1

)

E k

7 1

2 5

9 H

l(

T a

i e 0

0 3

6 O

Ar 2

9 R

xc 1

F o

F SN O

I T

)

A k

2 C

O n

5 L

gi n(

6 1

3 E

G i

L N

dZ B

I nM 6

3 4

A N

e 5

3 9

T R

Bt 1

E

)

n V

4 e

O m

G p

o o

M T

o A

L d

S a,

)

T o5 k

N l1 E

0 n

N t2 gi 3

5 1

O s

n(

P e-i 4

1 4

M h

dY 4

0 7

O gn nM 8

1 C

io e

1 Hi Bt G

t n

N a

e I

c m

D o

o A

L M

O

(

L

)

k 1

0 0

5 6

3 5

7 l(a ie 1

7 3

6 xc 1

9 Ar 1

o F

e o

t l

r M

t a

u h

m s

+

de d g r

s h

ap a i e

e E

c oy e e h

r S

n LT D W T

P S A

,d

~ f !.

{

.j N'

i t.

CONNECTS TO'RCL l

y N

/

SAFETY-INJECTION--

CONNECTION MODERATE ENERGY PIPING I-s Figure 5-1 Schematic Layout of the RHR Lines - Loop 1 mr.nconna 5.g

e i

.a N

MODERATE ENERGY PIPING b N/

CONNECT TO 3

REACTOR COOLANT l

LOOP 4-

,.l

. NODE 201 SAFETY-j INJECTION N0DE 2101-CONNECTION MODERATE ENERGY PIPING' Figure 5-2 Schematic Layout of the RHR Lines - Loop 4

. mr.noonno 5.y i

______ _ d

SECTION 6.0 FRACTURE MECHANICS EVALUATION O

l 6.1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the [

Ja,c.e method, based on traditional plastic limit load concepts, but accounting for (

,]a,c.e and taking into account the presence of a flaw.

The flawed pipe is predicted to fail when the remaining net section reaches a stress level at which a plastic hinge is formed. The stress level at which this occurs is called as the flow stress.

(,

Ja,c.e This methodology has been shown to be applicable to ductile piping through a large number of experiments and is used here to predict the critical flaw size in the RHR line. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 6-1) when loads are applied.

The detailed development is provided in Appendix A for a through-wall circumferential flaw in a pipe with internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:

[i

]

a,c.e (6.1)

[

Ja,c.e (6.2) where:

(

)a,c,e 6-1

[

l ja,c.e The analytical model described above accurately accounts for the piping

. internal pressure.as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental results-(reference 6-1).- Flaw stability evaluations using this

~'

'I

-analytical model, are presented in section 6.3.-

6.2 Leak Rate Predictions The purpose of this section is to discuss the method which will be 'used.to

~

predict the flow through such a postulated crack and present the leak rate' calculation results for postulated through-wall circumferential cracks in the RHR line.

6.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure (causing choking) is taken into account.

For long channels where the ratio of l

the channel length, L, to hydraulic diameter, D, (L/D ) is greater than g

H Ja,c.e must be considered.

[

Ja.c.e,-both [

In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturation pressure of the fluid.

l Pressure losses At'this~ point, the flow begins to flash and choking occurs.

due to momentum changes will dominate for [

Ja c.e However, for values, friction pressure drop will become important and must be l

large L/Dg l

considered along with the momentum losses due to flashing.

l-i uv nomno 6-2

[

}

s.

3. -

t 6.2.2' Calculation Method

'InLusing'the isentropic equilibrium model,'the basic method used in the leak rate calculations is the method developed by ("

)

Ja.c,e,

i The_ flow rate through a; crack was. calculated in the following manner.. Figure 6-2 from reference 6-2 was used to estimate the critical pressure, Pc,.for the j

primary loop enthalpy condition and an assumed flow. Once Pc was'found for 'a l

)"'Cd l

^

-given mass flow.-the~[

was..found from figure 6-3 taken'.from reference 6-2.

For all cases considered, since[

la,c.e Therefore, this method will' yield the two phase pressure drop due to momentum effects as illustrated in figure 6-4.

Now using the assumed flow rate, G, the frictional pressure drop

'can be calculated using APf=[

Ja,c,e (6.3) where the friction factor f is determined using the [

Ja,c.e The crack relative roughness, e,-was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these

. calculations'was(,

Ja.c.e pg3, The frictional pressure drop using Equation 6.3 is then calculated for the assumed flow and added to the [.

'Ja.c.e to obtain the total pressure drop from the primary system to'the atmosphere. Thus, Absolute Pressure - 14.7 = {

Ja,c.e (6.4) mr.nemua 6-3

_____ j

for a given assumed flow G.

If the right-hand side of equation 6.4 does not agree with the pressure difference between the piping under consideration and the atmosphere, then the procedure is repeated until equation 6.4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. This calculational procedure has been recommended by [

.)a,c.e for this type of (

.)a,c.e calculation.

6.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulated through-wall crack length for the critical locations previously identified. The crack opening area was estimated using the method of reference 6-4 and the leak rates were calculated using the calculation methods described above. The leak rates were calculated using average material properties with the normal operating loads of axial force F = 178 kips and bending moment M = 1848 in-kips.

The crack length yielding a leak rate of 10 gpm (10 times the leak detection requirement of 1.0 gpm) is found to be [

_]a,c.e, Thus the reference flaw size of (

3a,c.e is established.

6.2.4 Leak Detection Capability, Administrative Procedures and Technical Specification Requirements The VEGP leakage detection criterion includes a detectable unidentified leak rate of 1.0 gpm and, in accordance with NUREG-1061, Volume 3, a margin of 30 was applied to the leak rate to define the RHR line leakage size flaw used in the stability analysis. The basis for the 1.0 gpm leak rate is the presence (inside containment) of diverse and redundant leakage detection systems to measure containment gaseous radioactivity, airborne particulate radioactivity, f

a containment air cooler condensate flow and containment sump level. The sensitivity and response time of the detection equipment for unidentified 2estant)M7 to 6-4

leakage is such that a leakage rate, or its equivalent,-of 1 gpm can be detected in approximately one" hour as shown in FSAR figure 5.2.5-1.

In addition, humidity, temperature', and pressure monitoring of the containment

o atmosphere are used for alarms and indirect indication of leakage to the containment. These methods are in compliance with Regulatory Guide 1.45 as discussed in FSAR Subsection 5.2.5.

The above methods are supplemented by visual and ultrasonic inspection of the reactor. coolant pressure boundary during plant shutdown periods, in accordance with the Inservice Inspection Program (FSAR Subsection 5.2.4)'.

In addition, technical specification 4.4.6.2.1 requires monitoring of containment gaseous or particulate radioactivity and normal sump inventory and discharge at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br />. This section of the technical specification also requires performance of a reactor coolant system inventory balance at least once per 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.

Reactor coolant inventory monitoring provides an indication of system leakage. Operators perform a RCS leakage calculation (VEGP Nuclear Operations Procedure 14905) at least once every 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.

If the limits of Technical Specification 3.4.6.2 are determined to have been exceeded then'the operators are directed to the Abnormal Operating Procedure, Reactor Coolant System

{

Leakage (VEGP 18004).

j i

The RHR Line is integral to the RCS. Any leakage from the RHR line would be unidentified leakage until the exact location was determined. At that time the leak would have to be designated as reactor coolant pressure boundary leakage. Detected leaks will be repaired within the system limiting conditions for operation established in either technical specifications or l

administrative procedures. When leakage is detected in reactor coolant

' pressure boundary piping, Technical Specifications 3.4.6.2 requires that the l

plant be in. hot standby within six hours and in cold shutdown within the next thirty hours. Repair would be required before restart, I

nu.netuno 6-5 o

Experience at other We.stinghouse plants indicate a normal background unidentified average leakage rate of between 0.1 gpm and 0.3 gpm, and it has also been demonstrated with pressurized pipe tests that leak rates above 0.1 gpm at one location can be readily detected visually. The undefined leakage rate at VEGP is expected to be similar to other plants. Experience at similar plants and the results of these tests indicated that a 1.0 gpm leak rate can be reliably detected and located during plant operation.

6.3 Stability Evaluation Using the "2" Factor Approach A typical segment of the pipe under maximum loads of axial force F and bending moment M is schematically illustrated as shown in figure 6-5.

In order to calculate the critical flaw size, a plot of the limit moment versus crack length is generated as shown in figure 6-6.

The critical flaw size The corresponds to the intersection of this curve and the maximum load line.

critical flaw size is. calculated using the lower bound base metal tensile properties established in section 4.0.

From figure 6-6 the critical flaw size

~

is seen to be [

3a,c e for the base metal.

The weld at the location of interest (i.e. the governing location) is an SMAW weld. Therefore, a "2" factor correction for SMAW welds was applied (references 6-5 and 6-6) as follows:

2 = 1.15 [1 + 0.013 (0.D. - 4)]

(6.5) where OD is the outer diameter of the pipe in inches. ' Substituting OD = 12.75 inches, the 2 factor was calculated to be 1.2808.

The applied loads were increased by the 2 factor and the plot of limit load versus crack length was regenerated as shown in figure 6-7.

The lower bound base metal tensile properties (section 4.0) were used for this purpose. From figure 6-7, the critical flaw size is seen to be [

3a,c e long.

Noting that the flaw yielding a leakage of 10 gpm (i.e. leakage size flaw) was calculated to be ['

] long, a factor of 4.4 exists between the leakage size flaw.and the critical flaw. Thus, a margin of greater than 2 on flaw size is in evidence.

2"""""

6-6

i In. order to determine the' margin on applied loads-(normal _plus SSE), the l applied loads were' increased by a factor of 1.81 (i.e. /2 2) and the plot of

. limit load versus crack. length was generated as. shown in figure 6-8..

Again iA the lower bound base metal tensile properties were used for this purpose.

From figure 6-8 the-critical flaw size-is seen to be [

.)a c;e long which is larger than the (

-Ja.c.e inches long leakage size flaw. Thus a margin on. flaw size greater than 2.7 on'(normal plus SSE) loads.is demonstrated.

6.4. Local Stability Analysis 1

In'this section the local stability analysis is performed to show that unstable crack extension will not result when postulated.through wall flaws are subjected to maximum plant' loads.

At the critical location identified in section'5.0, the (normal plus SSE) outer surface axial stress; e,, is seen to be 25.2 ksi based on the minimum wall' thickness. The (normal plus SSE) axial force and bending moment are F = 185' kips and M = 2040 in-kips.

The minimum yield strength for' flaw stability analysis is [

Ja.c.e ksi i

(seesection4). The EPRI elastic plastic fracture handbook method is used to using the normal plus SSE loads. The Qapplied was-calculate the Japplied calculated for a [

Ja c.e long postulated through wall flaw (which l

is 2 times the reference flaw size) and was found to be (--

Ja,c,e 2

f 959 in-lb/in,

which is lower than JIc In addition, for a leakage size flaw, i.e. the reference flaw of [.

Ja,c,e long, the normal plus SSE load was increased by a factor of 6. The J-T analysis gave an applied J of (

Ja,c.e.

The applied tearing modulus J was calculated from the basic definition, namely, applied (6.6)

Tapplied

• 2

,o The h was obtained by calculating AJ corresponding to a small increment 4a in crack length using elastic plastic load-controlled analysis, mmnomno 6-7

4 The T was found to be (

Ja,c.e. The T I

applied mat

,3

'C is about 125. Thus the T is lower than T 8

l applied mat and, therefore, unstable crack propagation will not result.

i A

6.4.1 Crack Extension Considerations I I'

' The crack extension corresponding to the maximum calculated Japplied Ja,c.e would be about ('

)"'C at each crack tip.

If the

.J is calculated for a irrger crack length allowing for crack extension applied w uld be [

of ['

]a,c.e at each crack tip, the Japplied Ja,c.e The tearing modulus corresponding to this increased I

J w uld be about'(

. )a c.e which is again lower than the Tmat applied

70. Therefore consideration of crack extension does not change.the crack stability conclusions.

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

6-2 (

ya.c.e 6-3 (

ya,c.e,

6-4 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983.

6-5 NRC letter from M. A. Miller to Georgia Power Company, J. P. O'Reilly, dated September 9, 1987.

6-6 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.

m r.no,ur,o s.g

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4 Z-Factor Calculations for sgAW Welds to Demonstrate Margin r/

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on Flaw Size 6-15

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Figure 6-8 'W-Factor Calculations for SMAW Welds to Demonstrate Margin of Loads 6-16 nu.ne'="*

[,. r '

'1-J

.SECTION 7.0 3

M ASSESSMENT'Of FATIGUE CRACK. GROWTH

- T'he fatigue crack growth on the Vogtle Unit 2 RHR line was determined by.

~

comparison;with a~ generic. fatigue crack growth ~ analysis ~'of a 'similar-piping system.. The details;of the generic fatigue crack growth ' analysis'are presented in Appendix.B. By' comparing al.1 parameters critical to the fatigue

. crack growth analysis, between Vogtle' Unit 2 and generic, it was. concluded -

that;the generic analysis"would ' envelop the fatigue crack growth lof the Vogtle q

-Unit 2 RHR;1ine.

Due to similarities in ' Westinghouse PWR designs.it was possible. to perform' a

. generic fatigue crack. growth. calculation which would be applicable to many

-projects.

.A comparison was made..of stresses and number of cycles, material,

- geometry, and types of discontinuities.

l The~following summarizes the parameters which'were compared:

Generic. Hot Leg Nozzle

-Vogtle Unit'2 Hot Leg LCritical Location To Pipe Weld Nozzle to Pipe Weld Pipe Outer Diameter 12.75" 12.75" Thickness 1.005 1.005 l

Material Austenitic Stainless Steel Austenitic Stainless Steel

-Normal Temperature 617'F 617'F Normal Pressure' 2235 psig 2235 psig

, Normal Operating 20.3 ksi 23.1 ksi Stress'(Press,DW, Thermal Exp.

V Thermal Transients See Appendix B

' Thermal transient loadings are nearly identical for the two projects.

.i t

l aestincias?.to '

7.}

jj

I i

This comparison demonstrates the many similarities between the Vogtle RHR line l

and the generic RHR line evaluation. The only uncertainty in this comparison 1

is the higher level of pressure plus thermal expansion stress for Vogtle.

Since this is essentially a mean stress (steady state) this difference will have only a minor impact on the fatigue crack growth calculation by in:reasing the R ratio.

It is judged that this would increase the fatigue crack growth by less than [

]. ace Applying a [

la c.e increase to the ganeric fatigue crack growth data (from appendix B) results in a final flaw depth of approximately [

la,c.e for an initial flaw of [

).a,c,e These results demonstrate that no significant fatigue crack growth will occur over the 40 year plant design life even for the largest postulated flaw.

7.1 Acceptability of Fatigue Crack Growth A detailed discussion pertaining to the fatigue crack growth law used in the analysis described in Appendix B and the data used in defining the law are provided in reference (7-1). For the assessment of crack growth acceptabil-

?

ity, the crack growth results of the generic analysis presented in appendix B are used conservatively and are considered applicable to the Vogtle Unit 2 RHR lines. Detailed discussion in support of' this assumption has been provided in the previous section.

The maximum allowable preservice indication may have a depth of 0.1 in, per IWB-3514.3, Allowable Indication Standard for Austenitic Piping, ASME Code,Section XI - Division 1, 1986 edition. Estimated fatigue crack growth results are given in the previous section of this report.

[

Ja,c.e is conservatively chosen as a basis for examining the acceptability of fatiaue crack arowth.

[

ja,c.e Thus, the first criterion on flaw depth is satisfied.

l l

l l

nu.nemne 7-2 L

r l,c,e a

The worst case transient AK value for the maximum crack depth is [

The flow stress for the base metal at 617'F is 46.75 ksi which can be used to obtain a conservative estimate of the plastic zone size.

The expression for plastic zone size, r, calculation is:

[

3a,c.e p

= y (, flow)2 5

AK rp Thus, the plastic zne size is calculated to be [

]a,c,e The remaining ligament for the 0.182 in. deep end-of-fatigue-life flaw is 0.823 in. (i.e.

1.005 - 0.182). Thus, the plastic zone size is less than the remaining ligament.

Based on the above, it is concluded that for the Vogtle Unit 2 RHR Lines, the fatigue crack growth during service will not be significant.

! Ii) 7.2 References 7.1 Bamford, W. H. " Fatigue Crack Growth of Stainless Steel Reactor Coolant l

Piping in a Pressurized Water Reactor Environment," ASME Trans. Journal of Pressure Vessel Technology, February 1979.

7.2 Rice, J. R., ASTM STP,1967, Volume 415, p. 247.

m e

nir.noiw>.io 73

u m

o r,

sd'"

SECTION 8.0-e

' ASSESSMENT OF MARGINS-y In the preceding-sections,'the leak rate calculations, fracture mechanics:

~

4 Lanalysis and; fatigue crack growth assessment'were performed. Margins at.the; '

critical'. location are summarized below:

LIn'sectionI6.3 she " critical"' flaw size using. limit load' method is calculated-sto be {

Ja,c.e long. Using the IWB-3640: approach (i.e. "Z" factor ~-

approach), the critical flaw size'at the. governing location. weld is found.to.

be (

L]C'long.

In:section 6'4 it is demonstrated that a' postulated [L i)"'C long through'-wall flaw will L remain stable' when subjected to.. normal plus SSE loads. Based on the above, the' critical flaw

~

a -size will~ exceed ('

').a,c,e-In section 6.2:it is'shown that 'at the critical location, a flaw of [

n Ja,c.e would yield'a:1eak rate of 10 gpm.

Thus.nthere is a margin..of m

C at least 2 on flaw lsizeLand a' margin'of 10 with respect to.the plant leak detection. capability of 1 gpm.

'In"sectionsL6.3and 6'.4 it is shown that the reference flaw (

,)a,c.e. yielding' a leak rate of 10 gpm would be stable when subjected to a loa'd equal to I/2 (normal + SSE). Specifically, using the IWB-3640 approach'(section 6.3) a [

la.c.e long through-wall flaw was shown'to be-stable when subjected to [2 Z (normal + SSE) loads. Also, based on local

. stability analysis (section 6.4).the leakage size flaw of [

3a,c.e inches was shown to be stable when subjected to [2-(normal + SSE) loads.

i nu.homu*

8-1

1

' I'n sumary, relative to

- 1.

' Loads 4

'The 1eakage-size crack will not experience unstable crack extension even if very large loads of /Y (Normal plus SSE) are applied.

1 1

2.

Flaw Size A margin of at least 2 exists between the critical flaw and the flaw a.

yielding a leak rate of 10 gpm.

1 2.

If limit load is used as the basis f.or critical flaw size, larger margin for. global stability _would result.

3.

' Leak Rate A margin of 10 exists for the' reference flaw [

Ja c.e between calcu. ated. leak rate and the criteria of Regulatory Guide 1.45.

j l

A' summary comparison of criteria and analytical results is given in Table 8-1.

The criteria are seen to be met.

0 e

4 m7. nniur.io 8-2 j

L---

14 TABLE 8-l' COMPARISON OF RESULTS VS. CRITERIA

.p.

i CRITERION RESULT

'1.

NUREG1061 Volume 3

-Met l

Section 5.2(h) -

(Required margin ~ of 2 demonstrated).

Margin on Flaw. Size 2'

NUREG1061' Volume 3 Met Section5.2(i)-

(Required margin of /2 demonstrated)

Margin on Load I

3..

NUREG 1061 Volume 3 Met Section 5.7 -

(Margin of 10 on leak. rate 1.

Margin on Leak Rate demonstrated) 4.

NRC criteria on allowable Het fatigue crack growth (af < 60% wall thickness)

Plastic zone size < remaining ligament) i 3

l 1

4 4

i i

I

[ mwmar se 8-3 J

L 7

r

,.c

,I

' SECTION 9.0 L

m3 CONCLUSIONS =

or ThisireportIjustifies the' elimination.of RHR.line pipe breaks as the

' structural design basis for the Vogtle Unit 2 as follows:

a.

15 tress. corrosion cracking is precluded by use of fracture resistant ~

materials in the piping system and controls on reactor coolant

' chemistry,' temperature, pressure, and flow during normal operation.

b.

Water hammer sh'ould-not occur in the RCS piping'(primary loop and the attached auxiliary lines) because' of system design, testing, and

-operational considerations.

~

i c.

The ' effects of: low and high' cycle fatigue on the integrity of the RHR.line piping are negligible, 3

1 w

'd.

1 Ample margin exists between the leak rate of sma11' stable flaws and the~ requirements of Reg. Guide 1.45.

1 e.-

Ample margin exists between the small stable flaw sizes of item d and the' critical flaw.

y f.

With respect to stability of the reference flaw, ample margin exists between the maximum postulated loads and the plant specific faulted loads (i.e. Normal + SSE).

The reference flaws will be stable because of the ample margins in d, e, and f and will leak at a detectable rate which will assure a safe plant shutdown, I

a:

Based on the above, it is concluded that pipe breaks in the RHR piping need

]

not-be considered in the structural design basis of Vogtle Unit 2.

l l

I l

me,nomno 93

APPENDIX A LIMIT MOMENT l

s h

8,C,9

)

)

h l

l I

i f

I nu.ncixuo A-1

i l

i l

i '-

1

'l

),

(

)

1 4

i s l t

- a,c.e i

1 f

i 4

l

)

4 Figure A-1 Pipe with a Through-Wall Crack in Bending i

mwi,.,,.

32 1

I APPENDIX B FATIGUE CRACK GROWTH CONSIDERATION 1

B.1= Thermal ' ransient Stress Analysis T

1 The thermal transient stress analysis was performed for a typical PWR plant to

.obtain the through wall stress profiles for use in the fatigue crack growth analysis of Section B.2.

The through wall stress distribution for each transient was calculated for i) the time corresponding to the maximum inside surface stress.and, ii) the time corresponding-to the minimum inside surface stress. - These two stress profiles are called the maximum and minimum through wall stress. distribution, respectively, for convenience.

The constant stresses due to pressure, deadweight and thermal expansion (at normal operating temperature, 617'F) loadings were superimposed on the through wall cyclical stresses to obtain the total maximum and minimum stress profile for j

aach transient.

Linear through wall stress distributions were calculated by conservative simplified methods for all transients.

B.1.1 Critical Location for fatigue Crack Growth Analysis l

The RHR.line stress report design thermal transients (Section B.1.2),.1-D analysis data on RHR line thermal transient stresses (based on ASME Section 111 NB3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis.

[

Ja c.e This location is selected as the worst location based on the following considerations:

i) the fatigue usage factor is highest.

ii) the effect of discontinuity due to undercut at weld will tend to l

increase the cyclidal thermal transient loads.

iii) the review of data shows that the 1-D thermal transient stresses in the RHR line piping section are generally higher near the [

.)a,c.e i

l

.)

1 B'.1.2 Design Transients The transient conditions selected for this evaluation are based on conser-

~vative' estimates of the magnitude and the frequency of the temperature fluctuations _resulting from various operating conditions in the plant. These y

I are representative of. the conditions which are considered to occur during plant operation.. The' fatigue evaluation based on these transients provides confidence that the component is appropriate for its application'over the design life of the plant. All the normal operating and upset thermal transients, in accordance with design specification and the applicable system design criteria document'(B-1), were considered for this. evaluation. Out of

.these',(

3a,c.e i

B.1.3 Simplified Stress Analysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients.

^j

'I

Ja,c.e The inside surface stress was calculated by.the following equationLwhich is similar to the transient portion of ASME Section III NB3600,-

l.:

Eq. 11:

-i Ja,c.e S$ = [..

(B.1)

where, S$ = inside surface stress l

N l

+a,c.e l-(B.2)

(B.3) t

\\

/

i I

sostenous?.a g.2 1

J l

N

/

l

.(B.4) a,c e' l

\\

\\

1 i

'I i

e l

/

.l i

i

.y

[

Ja,c.e The maximum and minimum inside surface stresses were se' arched frein the S$ values. calculated for each time step of the transient solution.

The outside surface' stresses corresponding to maximum and minimum inside stresses j

were calcu atel ' d by the following equations:

1 I

)* ' (B.5)

S01 =-[

)+ ' ' ' (B.6) j 502 = {

i

where,

/

t i

i s

+a,c.e.

' All other parameters are as defined previously I

1 nu.noim.ie B-3

The material properties for the RHR pipe ((SA376 TP316)) and the RCL [

J a,' ',e The c

values of E and a, at the normal operating temperature, provide a conserva-tive estimation of the through wall thermal transient stresses as compared to room temperature properties. The following values were conservatively used, which represent the highest of the (piping and nozzle) materials:

-/

+a,c,e

/

\\

The maximum and minimum linear through. wall stress distribution for each thermal transient was obtained by [

la,c.e The simplified analysis discussed in this section was performed for all thermal transients of table B-1.

The inside and outside surface stresses calculated by simplified methods for the transients are shown in table B-2.

B.1.4 OBE Loads

~

The stresses due to OBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contribute significantly to crack growth due to small number of cycles.

B.1.5 Total Stress for Fatigue Crack Growth The total through wall stress at a section was obtained by superimposing the pressure load stresses and the stresses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of table B-2). Thus, the total stress for fatigue crack growth at any point is given by the following equatiori:

B-4

,d y 1

I 4

,g 1

-Tot'al Thermal-

Stress Due Stress

^

forx

-Transient to' Due to-

,g

=

+

DW +

+

Internal' (B.7)

Fatiguei Thermal Pressure

' Crack Growth

-Expansion

!Theenvelopthermalexpansion, deadweight ~andpressureloadsforcalculating the tota 1' stresses of. equation B.7 are summarized in table B-3.

B.2 Fatigue Crack Growth Analysis

'The fatigue-crack gr'owth analysis was performed to determine the effect'of the' design thermal transients, in tab 1e' B-1.

The analysis was performed'for-the_-

eritical cross section of the model which is identified in figure B-1.

A J

range of crack depths was postulated, and each was subjected to the transients

..in table B-1.

.B.2.1 ' Analysis-Procedure j

The fatigue crack growth analyses presented herein were conducted in the'same manner as suggested by Section'XI, Appendix A~of the ASME Boiler and Pressure Vessel Code. 'The analysis procedure involves assuming an initial flaw exists at some point and predicting the growth of that flaw due to an imposed series of stre'ss transients.

The growth of a crack per loading cycle is dependent on the range of applied stress intensity factor AK, by the following g

relation:

h=CoAK" (B.2.1) g is where "Co" and the exponent "n" are material properties, and AKg g

defined lata, in equation (B.2.3). For inert environments these material properties are constants; but for some water environments they are dependent on the level of mean stress present during the cycle. This can be accounted I

B-5

]

WEs71NGHOUst PROPRIETARY class 2 for by. adjusting the value of.."Co" and "n" by a function of th'e ratio of f-

'. minimum.to maximum stress for any given transient, as will be discussed

-later.-. Fatigue crack growth properties of stainless steel in a pressurized water environment have been used in the analysis.

The input required for a fatigue crack growth analysis is basically the.

information'necessary to calculate the parameter AK;, which depends on

~

crack and structure geometry and the range of applied stresses in the area l

where the crack exists. Once AK is calculated, the growth due to that g

particular cycle can be calculated by equation (B.2.1). This increment of growth is then added to the original crack size, the AK; adjusted, and the analysis proceeds to the next transient. The procedure is continued in this manner until all the transients have been analyzed.

The crack tip stress intensity factors (K;) to be used in the crack growth l

analysis were calculated using an expression which applies for a semi-elliptic i

surface flaw in a cylindrical geometry (reference B-3).

-The stress-intensity factor expression was taken from Reference B-3 and was calculated using the actual stress profiles at the critical section. The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial:

o (x) = A + A3{+A({}

A({}3 (B.2.2) 0 2

3 The stress intensity factor K (v) was calculated at the deepest point of y

the crack using the following expression:

\\

+a,c.e (B.2.3) s 4

s B-6

........,0...

hI

.o I,

I{

./

c.,

a c.e-i e

i I

I-l-

N

/

Calculation of the fatigue crack growth for each cycle was then carried out using the' reference fatigue crack growth rate law determined from considera-tion of. the available data for stainless steel in a pressurized water environment.- This law allows for the effect of mean stress or R ratio Imin/K,,x) on the growth rates.

(K y

The reference crack growth law for stainless steel in a pressurized water environment was taken from a' collection of data.(B-4) since no code curve is 7

available, and-it is defined by the following equation:

" I' 3 a,c.e (B.2.4) jff = (K,,x) (1-R)1/2 where K g

Kimin R=K Imax da g = crack growth rate in micro-inches / cycle

?

B-7

t f =.

8.2.2 Results.

Fatigue _ crack. growth analyses were carried out for the critical cross section. Analysis was completed for a range of postulated flaw sizes oriented circumferential1y.. and-the results are presented in Table B-4.

The postulated flaws are assumed to be six. times as long as they are deep..Even for the.

largest postulated flaw of (

la,c.e the result shows.that the flaw growth through the wall will ret occur during.the 40 year. design' life of the plant. For smaller flaws, the flaw growth is 'significantly lower. ' For example, a postulated i Ja,c.e inch deep flaw will grow less than [.

J. These results also confirm plant j

operating experience.-

B.3 REFERENCES j

B-1 [.

3a,c.e B-2 ASME Section III, Division 1-Appendices,1983 Edition, July 1,1983.

..L i

l B-3 McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for.

Internal Longitudinal Semi-Elliptical. Surface-Flaws in a Cylinder Under

'f i

Arbitrary Loadings",_ Fracture Mechanics ASTM STP-677,1979, pp. 365-380.

'B-4 Bamford

.W. H., " Fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment", ASME Trans. Journal of Pressure Vessel Technology, February 1979.

1 1

i c:mn*

• B-8

TABLE B-1 THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION p.

No. of Trans.

Occurrences No.

Description N-a,c.e

/

6 9

i e

~ ^ ^

. _ _ _ _._ ___ h o

j 4

i 1

i

. TABLE B-2 TRANSIENT' STRESSES FOR RHR LINE (psi)

Transient Maximum Corresponding Minimum Corresponding N o ~.

-Inside Stress Outside Stress Inside Stress Outside Stress i

s.

y a,c.e i

l 1

l i

.e

)

/

N i

i nir.noiur.io B-10

e TABLE B-3 ENVELOP NORMAL LOADS s

a,c.e Cond'< tion Normal Operating

/

s E

9 1

B-11 l

'. I 1: -

L J

E

<l

=

' TABLE B-4 RHR LINE FATIGUE CRACK GROWTH'RESULTS' i

Section. Thick' ness [

-),+a,c, 4

.. INITIAL CRACK DEPTH AFTER' YEAR CRACK DEPTH (in) 10 20 30 40 s.

a,c.e j

-)

F

. l

/

\\

i s

• This is conservatively taken as minimum thickness of the counter bore region i

B-12 i

r i

l

e i

o

~

l I

a,c.e

~.

RHR PIPE i

J Figure B-1 Schematic of RHR Line At [

3 "' **

mr.nwino B-13