ML20087G335

Nonproprietary Technical Basis for Eliminating Pressurizer Surge Line Ruptures as Structural Design Basis for Catawba Units 1 & 2

ML20087G335
Person / Time
Site:	Catawba
Issue date:	02/10/1984
From:	Sane A, Swamy S, Yang C WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
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ML19292D068	List: ML20087G310 ML20087G322 ML20087G335
References
WCAP-10488-(NP), NUDOCS 8403190404
Download: ML20087G335 (107)

Text

. -

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r ENC'.0SURE B 4

WCAP-10488 TECHNICAL BASIS FOR ELIMINATING PRESSURIZER SURGE LINE RUPTURES AS THE STRUCTURAL DESIGN BASIS FOR CATAWBA UNITS 1 & 2 S. A. Swamy C. Y. Yang A. D. Sane Y. S. Lee February 10, 1984

-APPROVED:

k- A C APPROVE 8:

JIN.'Chirigos,Mana'ged E. R. Johnson, Manager Structural Materials Structural and Seismic Engineering Development i

8403190404 840229 -

PDR ADOCK 05000

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TABLE OF CONTENTS SECTION TITLE PAGE

1.0 INTRODUCTION

1-1 1.1 Background 1-1 1.2 Scope and Objective 1-1 1.3 References 1-2 2.0: FAILURE CRITERIA FOR FLAWED PIPES 2-1 2.1 General Considerations 2-1 2.2 Global Failure Mechanism 2-1 2.3 Local Failure Mechanism 2-2 2.4 Corrosion 2-3 2.5 References 2-4 3.0' LOADS FOR CRACK STABILITY ANALYSIS 3-1

4. 0' CRITICAL FLAW SIZE CALCULATION 4-1 5.0 [. ] ANALYSIS FOR CRACK STABILITY 5-1 +a,c.e CALCULATIONS 5.1 The [ ] Model and the Material 5-1 +a,c.e Properties 5.2 Boundary Conditions and Method of Loading 5-2 5.3 Method Of Analysis 5-2 5.4 [ ] Results 5-3 +a,c,e 5.5 References 5-3 6.0 LEAK RATE PREDICTIONS 6-1 6.1 Introduction 5-1 6.2 General Considerations 6-1 6.3 Calculation Method 6-1 6.4 Crack Opening Areas 6-3 6.5 Leak Rate Results 6-3

- 6.6 References iii

,;i*  ;

TABLE OF CONTENTS (Cont'd.)

SECTION TITLE l PAGE 7.0 THERMAL TRANSIENT STRESS ANALYSIS

. 7.1 Critical Location for Fatigue Crack Growth 7-1 Analysis 7.2 Design Transients 7-2 7.3 Simplified Stress Analysis 7-2 7.4 [ ] Distribution for Severe 7-4 +a,c,e Transients 7.5 OBE Loads 7-5 7.6 Total Stress for Fatigue Crack Growth 7-6 7.7 Referenc s -

77 S.0 FATIGUE CRACK GROWTH ANALYSIS 8-1 8.1 Analysis Procedure 8-1 8.2 Results' 8-3 8.3 References 8-3 9.0 -CONCLUSIONS 9-1 APPENDIX A A-1 APPENDIX B B-1 d'

4 1' V e

" ^ ' ' ~ ~ ~ ~ ~ '

. , . . _ . , , _ . . . , - -- - - - - - - - - - ~ ' - - - " ' - -

r t

LIST OF FIGURES FIGURE TITLE PAGE 2-1 Typical Load-Deformation Behavior 2-5 3-1 -

Catawba Surge Line Piping Analysis Model 3-3 4-1 [ ] Stress Distribution 4-3 +a,c,e 4-2 Comparisonof[ ] Predictions 4-4 +a,c,e with Experimental Results 4-3 Critical Flaw Size for Pressurizer Surge Line 4-5 5-1 Loads Acting on the Pipe 5-6 5-2 The [ ] Model. [ 5-7 +a,e,e

]

5-3 The [ ] model of the pipe showing [ 5-6 +a,c,e

] [ ]

5-4 A close-up view of the [ 5-9 +a,c,e

]

5-5 The [ ] pattern in the vicinity of the 5-10 +a , c . e crack front.

c 5-6 [ ] on the middle of the crack surface 5-11 +a,c.e 5-7 [ ]atthe 5-12 +a,c,e pipe end which is subjected to the applied axial and bending stresses.

5-8 True stress-strain curve and the [ ] 5-13 +a,c e approximation 5-9 Schematic of the boundary conditions 5-14 5-10 Loading schedule for the internal pressure aoplied 5-15 to the inside surface of the pipe.

5-11 Loading schedule for the uniform axial stress amplied 5-16 to the pipe end.

5-12 Loading schedule for the bend moment applied to the 5-17 pipe end. .

- ~ -

5-13 5-18 +a,c,e V

I ., . .

l_i- ,

LIST OF FIGURES (Cont'd.)

FIGURE- TITLE PAGE 6-1 Analytical Predictions of Critical Flow Rates 6-6 of Steam-Water Mixtures 6-2 [ ] Pressure Ra'tio as a Function 6-7 +a,c,e

. of L/D 6-3 Idealized Pressure Drop Profile Through a Postulated 6-8 Crack 6-4 Crack surface profile under [ 6-9 +a,c,e

]

6-5 -

Crack surface profile under [ 6-10 +a , c , e

]

7-1 Comparison of Typical Maximum and Minimum Stress 7-11 Profile Computed by Simplified [ +a,c,e

]

7-2 Schematic of Surge Line at [ ] 7-12 +a,c,e 7 Schematicof[ ] Without 7-13 +a,c,e Discontinuity 7-4 Maximum and Minimum Stress Profile at Critical Location 7-14 for Unit Loading Transient 7-5 Maximum and Minimum Stress Profile at critical Location 7-15 for Steady State Fluctuation [- +a,c,e

] Transient 7-6 Maximum and Minimum Stress Profile at Critical Location 7-16 for Random Fluctuation [ ] Transient +a,c,e 7-7 Maximum and Minimum Stress Profile at Critical Location 7-17 for[ ] Cycling Transient +a,c,e 7-8 Effect of Discontinuity on Through-Wall Stress Profiles 7-18

.A-1 Equilibrium of Horizontal Forces A-3 B-1 Auxiliary Diagram for Derivation of Equation B-6 B-6 vi

g- I.

1 LIST OF TABLES TABLE TITLE PAGE

~

3-1 A Summary of Catawba Surge Line Locations with High 3-4 Loads and Stresses 5-1 [ 5-5 +a,c.e 3

6-1 Crack Surface Displacement Data 6-5 .

7-1 Thermal Transients Considered for Fatigue Crack 7-8 Growth Evaluation 7-2 Thermal Transient Stresses by Simplified Analysis 7-9 7-3 Pressure, Deadweight and Thermal Expansion Stresses 7-10 for Fatigue Crack Growth 8-1~ Fatigue Crack Growth Results 8.-4 yii et'-- r -

-w -ww 7 *7

• re-*y--W -m==y- -------ww-

- * ' - - - - - *

. i

1.1 BACKGROUND

The current structural design basis for the pressurizer surge line requires postulating nonechanistic circumfsrential (guillotine) breaks in which the pipe is assumed to rupture along the full circumference of the pipe. This results in overly conservative estimates of support loads. It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for the pressurizer surge line. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that can '

be used for establishing that a guillotine type break will not occur within the pressurizer surge line. The evaluations considering circumferential1y oriented flaws cover longitudinal cases.

1.2 SCOPE AND 08.1ECTIVE f

The general purpose of this investigation is to show that a circumferential flaw which is larger'than any flaw that would be present in the surge line will remain stable when subjected to the worst combination of plant loadings.

The flaw stability criteria proposed f.or the analysis will examine both the global and local stability. The global analysis is carried out using the

[ ] method, based on traditional [ ] +a,c,e concepts, but accounting for [ ] and taking into account the +a c.e i,resence of a flaw. This analysis using faulted loading conditions enables determination'of the critical flaw size. The leakage flaw is conservatively j selected with a length equal to half the critical length. The local stability enelytit it carried out by performing a [ +a,c,e

! ] of a straight piece of the surge line pipe containing a through-wall

! circumferential flaw subjected to internal pressure and external loadings (faulted conditions). The objective of the local analysis is to show that j ,

unstable crack extension will not result for a flaw [ .,,,,,

] calculated by the global analysis.

l l

l L

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. -. = ._ -. . _ . . . . . . - . _ - . .. . _

,o  : '

The leak rate is calculated for the [ ] condition. [ + ,c,e

] The crack +a c.e

-f opening area resulting from [ ] loads is determined from an +a,c e assumed through-wall flaw of [

] +a,c,e i

[ ] is accounted for in determining the leak rate through this +a r c e e crack. The leak rate is compared with the detection criterion of I gpm (Reg. Guide 1.45). The leak rate prediction model is an ( +a,c,e

] This method was used earlier to estimate the leak rates through postulated cracks in the PWR primary coolant loop. [1-1)

1.3 REFERENCES

1-1 Palusamy, S. S. and Hartmann, A. J.,

" Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack" WCAP-9570 Rev. 2, Class 3 June 1981, Westinghouse Nuclear Energy Systems.

5640Q:10/021384 1-2

_ _ _ _ -. . . -  :::~~ ::: - L--- - ~

, s 2.0 FAILURE CRITERIA FOR FLAWED PIPES 2.1 GENERAL CONSIDERATIONS '

1 Active 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 void nucleation. Several of these criteria are discussed in a recent ASTM publication (2-1). i 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 for this investigation. The ultimate objective is to show that the pressurizer surge line containing a conservatively assumed circumferential through-wall flaw is stable under the worst combination of postulated faulted and operating 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 For a tough ductile material if one assumes that the material is notch insensitive then the global failure will be governed by plastic collapse.

j Extensive literature is available on this subject. The recent PVRC study l [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 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.

i l

c. 5640Q:10/021384 2-1

.. - ,L .--.-,: : L - .J -.__ -..-..~~ -...--.L::  :.LL:L ::- -.

. o 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 experiment. This criterion can be suiiriiarized by the following relationship:

Wa < Wp (2-1) where ' We = applied generalized load Wp = calculated generalized plastic load

, In this report Wp will be obtained by an ( +a,c.e 3

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 evaluation of local failure.

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

been accepted that the initiation toughness, measured in terms of J gy 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 Jg , of the material, then the crack will not initiate.

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

T app

=N1 da a 2 (2-2)

g l

l 1

~

l 56400:1D/021384 2-2 l

l

.,:= .

r . := :. ._ , = - _ _ _ : r z : : : : z _. - _ - . .

.. o where- T,pp -

ipp. led tearing modulus E - = modalus of elasticity

• f

=

flow stress = [ .

] +a,c,e a = ' crack length

[ ,

+a,c e 3

/ In summary,' the local crach stability will be established by the two step 7

criteria:-

. J<Jgy, or' (2-3)

. ~

T,,, < Tu t' 1I J # JIN

, (2-4) 2.4 CORR 0SI W t

The Westinghouse reactor coolant system primary loop has an operating history i- (over 400 reactor years) which demonstrates its inherent stability

, characteristics. Additionally, there is no history of cracking in RCS primary loop piping. . In, addition to the fracture resistant materials used in the piping system, the chemistry of the reactor coolant is tightly controlled.

/ .

~

As stats %4bove, the r'eactor coolant chemist'ry is maintained within very specificlimits. For example, during normal operation, oxygen in the coolant i is' , limited to less than [r ] This stringent oxygen limit is achieved +a,c.e by controliing charging flo.' chemistry and maintaining hydrogen in the reactor coplant at a concentration of [ ] The oxygen concentrntion +a, c.e

in' the reactor coolant is verified by routine sampling and chemical analysis.

Halcger.. concentrations are also stringently controlled by maintaining cencqntrations of chlorides and fluorides at or below [ ). This +a.c,e w ,

' concentration is assured by controlling charging flow chemistry and specifying L - proper wetted surface materials. Halogen concentrations are also verified by routine sheraical sampling and analysis.

! 2 l ~

,N l

5640Q:10/02'1384 2-3 i

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. e

2.5 REFERENCES

2-1 J. D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia, PA 19109, November 1977.

2-2 J. C. Gardeen, "A Critical Evaluation of Plastic Behavior Data and a Unified Definition of Plastic Loads for Pressure Components," Welding Researc'h Council Bulletin No. 254.

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

5640Q:10/021384 2-4

. e s ,

W?

AL W'

I m.

O a

i m t <>

ad

= .

b U

.C t U l e f

i L, ' r

~_ .

Y GENERALIZED 0:5 PLACE. MENT FIGURE 2-1 Typical Load - Defomation Behavior 4

7 e

2-5

. n 3.0 LOADS FOR CRACK STABILITY ANALYSIS The surge line stress report [3-1] was reviewed to identify locations with high ASME N8 3600 faulted (Eq. 9) stresses. These locations are identified en I the surge line computer model in Figure 3-1. The loads at each of these locations were tabulated from the computer runs of [3-1] for the [ +a,c e '

] .

loading cases. The axial load, bending moment and stress at these locations were calculated from the tabulated loads as follows:

F=[ ] (3.1) +a,c,e My=[

] (3.2) +a,c.e M

g =[

] (3.3) +a,c.e 2 2 M= My,g 7 (3.4) a=f+f (3.5) where, subscript [DW, TH, SSE, AM) indicate the loading cases, F,

= axial load due to normal operating pressure M = Y component of moment y

M = Z component of moment j F = total axial load at the location M = bending ~ moment at the location A = metal cross-sectional area of piping Z = sectional modulus of pipe

(

i l

I i

l 5640Q:1D/021384 3-1 j _ .. - - - - - - ----

. s The wrought piping material is the same, (i.e., [ ] for the +a,c.e entire surge line and hence the location with the highest stress calculated by equation (3.5) was identified as the worst location for the global and the loca'l crack stability analysis of Section 4.0 and '

5.0, respectively. The (  :-c,e

] was selected as the critical section based on this criteria (sea Table 3-1). The calculated axial load, bending moment and longitudinal. stress at this location are:

t _.

+a c,e Axial Force Bending Moment Longitudinal Stress

+a,c.e 3

The operating transients of the surge line are such that no ( +a,c.e 3

REFERENCES 3-1 EDS Report No. 03-0093-1025, Revision 0, 'ASNE Boiler and Pressure Vessel Code Section III Class 1 Stress Report for the RCS Pressurizer Surge Line, Catawba Nuclear Station Unit 1."

l l

l l

?

i 1

1 i

5640Q:lD/021384 3-2

+ 0 , 0 **O Ik

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i 1

i i

i i

w U'

i i

i i

I J

i 5

f i

'I i

,l li I'

FIGURE 3-1: CATAWBA SURGE LINE PIPING ANALYSIS MODEL 9

S TABLE 3-1 A

SUMMARY

OF CATAWBA SURGE LINE LOCATIONS WITH HIGH LOADS AND STRESSES Axial Bending Stress Node Force Moment a No . . (k) (ft.-k) '

(ksi) +a,c e

~

4 e

l L

1 m

e 3-4

.,,,,g.w--r .-m==e--=+e '

y  %

• 4.0 CRITICAL FLAW SIZE CALCULATION The conditions which lead to failure in stainless steel must be determined using plastic fracture methodology because of the large amount of deformation accompanying fracture. A conservative method for predicting the failure of ductile material is the [ +a,c.e

] The flawed pipe is predicted to fail when the [ +a,c,e

] This methodology has been shown to be applicable to ductile piping through a large numbe of experiments, and will be used here to predict the critical flaw size in the pressurizer surge line. The failure criterion has been obtained by

[ +a,c,e

] The detailed development is provided in Appendix A, for a through-wall circumferential flaw in a pipe with [ +a,c,e

] The [ ] for these conditions is:

W_ +a,c,e e

WW 5640Q:10/021384 4-1

The analytical model described above accurately accounts for the piping internal pressure as we'.1 as imposed axial force as they affect the [ +1i,c e

] In order to validate the model, analytical predictions were compared with the experimental results [4-1] as _ shown in Figure 4-2. Good agreement was found.

In order to calculate the critical flaw size, a plot of the [

~]

+a,c.e versus crack length is generated as shown in Figure 4-3. The critical flaw size corresponds to the intersection of this curve and the maximum load line.

i The critical flaw size is [ ] using ASME Code [4-2] [ +1tec e

] stainless steel.

Since [ ] for crack smaller than [ ] dnd [ +1a cre

] the global stability criterion of Section 2.0 is satisfied.

Reference 4-1 Kanninen, M. F., et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with-Circumferential Cracks" EPRI NP-192, September 1976.

4-2 ASME Section III, Division 1-Appendices,1983 Edition, July 1,1983.

t i

i 5640Q:10/021384 4-2

~ - -- -

p ** O e .

C e * .

+ U

+

I h

i I

=

=

N W

D 3

w E e O

$o N $

.Q 5

a M

e,-

Q M

T, L

M m

M

% F1 N w T

e O

L 3

W 43

' ' o e ... -- = = + - - - - - .e.ww---%..--. . --,-- - , - - , , -

+a,c,e C

I

\

l FLAW GECMETRY

+

E 1

I l

l t

FIG!JRE 4-2 Comparison of [ Limit yc. ment: Dredi:tions +a,c.e With Experimer.tal Results l

4-4 a

, - . . . , ..~.--- .--.we-.--------~---w -w -~-.-e---- - - - - - - - - - --

PRESSURIZEP. SURGE LINE USING ASME CODE MINIMUM PROPERTIES

_ +a c,e

_a-1 FLAW GE0?iETRY

+a,c.e I

i f

Figure 4-3 Critical Flaw Size for Pressurizer Surge Line 4-5

)

5.0 FINITE ELEMENT ANALYSIS FOR CRACK STABILITY CALCULATIONS  !

Using the [ ] computer program, a [ +a,c,e

] crack was analyzed to determine the local l

stability." The loadings consist of [ +a,c,e 1

5.1 THE IFINITE ELENENTl MODEL AND THE MATERIAL PROPERTIES Figure 5-1 identifies all the loads acting on the pipe. The pipe thickness is

[ ], based on the thinnest location of the surge line under +a,c,e investigation. The outer diameter is [ ). Due to symmetry only one half +a ces of the circumference, i.e.,180-degree, is modeled. The length of the model is [ ] which is sufficiently long to attenuate the +a,c.e

-effect of the cract for t.orrect boundary load input from the pipe end.

Figures 5-2 through 5-7 all show the [ ] used for + ,c,,

analysis. The [ ] are identified in Figure 5-2 through 5-5.

The [ -] of interest for later leak rate predictions are shown in +a,c.e detail on Figure 5-6. The [ ] and their Z-coordinates required +a,c,e fo'r the application of the axial loads and the bending moment are shown in Figure 5-7.

I +a,c,e 3

The true stress-strain curve of the material is shown in Figures 5-8. The data are taken from the " Nuclear System Materials Handbook (5-2] for the stainless steel [ ] The stress-strain curve is ( +a,c,e

] It has been shown that the

[bi-linear] approximation gives good agreement with the experimental results

~a( +a,c,e

]

5640Q:10/021384 5-1

~

~ ~ ~ ~ ~ '

~ ~.T X : X ~ X X : - _.- .. - . . -.- L - =

[5-3]. The material properties used in the present analysis are [ +a,c,e

]

l 5.2 BOUNDARY CONDITIONS AND NETHOD OF LOADING The boundary conditions are described in Figure 5-9. The pipe is subjected to the internal pressure of [ ] and an axial load of [ ] A +a,c e bending moment of [ ] is then superposed to the pipe while the +a,c e pressure and the axial. loads are held constant. Due to non-linear material behavior, the loads are added to the pipe [ +a,c.e

]

Figures 5-10, 5-11 and 5-12 show the sequence of applying the loads to the

[ ] model of the pipe. Figure 5-10 shows [ +a,c,e

] after which it is held steady. As shown in Figure 5-11, the axial load due to [ +a,c,e

] Figure 5-12 shows application of the moment, starting at load step 3, where [ +a,c,e

( )

l 5.3 NETHOD OF ANALYSIS

As mentioned in Section 2 of the present report the local instability criterion is based on the Information of the J-integral, J, and tne tearing modulus, T. The tearing modulus is defined in Equation 2-2. When the J-value due to the applied loads equals and exceeds a critical value, J rJ IC IN of the material, the crack initiation will occur. When the T-value due to the applied loads equals and exceeds the T-value of the material T , the int.tability will occur.

[ +a,c.e i ) This method has been successfully used to analyze a cracked pipe under a combined axial load and bending moment [5-5].

! 5640Q:1D/021384 5-2 3-3- _ _ _ _ _

, . l I

The VCE method has been incorporated in the [ +a,cee

] to calculate the average [ ] of a crack as well as the +a,c,e

[ ] along the crack front for both (

) unalyses. The [ ] at each load level can be computed by way +a,c.e l of.the [ ] solution strategy. +a c.e ,

5.4 IFINITE ELEMENTl RESULTS ,

It should be noted that JIN ( Ic} "" mat refer to two stages of a fracture process of a material. If the applied J-value is less than J TN' the crack will not advance. Under this condition the tearing modulus need nct be evaluated. Figure 5-13 shows how the calculated value of the J integral increases up to and beyond the maximum operating loading at 653*F.

At the maximum loading, the [ ] has a corresponding value of ( +a,c e

] as shown on the figure. Since this value is smaller than the

[ +a,c e

] the condition for crack stability is automatically fulfilled. [ +a,c.e l

] The [ ] as a function of +a,c,e loads is shown in Figure 5-13 and Table 5.1. The verification of the analysis is shown in Appendix B.

I

5.5 REFERENCES

~

~

+"'

5-1 l

l L 5-2 Nuclear System Materials Handbook, Volume I Design Data, Revision 1, 10/1/76.

l 5-3 Palusamy, S. S., Hartmann, A. J., " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through Wall Crack," WCAP 9558, Revision 2, dated May 1982.

l 5640Q:10/021384 5-3

p ..

5-4 Parks, D. M.,

"The Virtual Crack Extension Method for Nonlinear Material Behavior," Computer Methods in Applied Mechanics a"4 Engineering," Vol.

12, 1977, pp. 353-364.

5 Yang, C.' Y. and S. S. Palusamy "VCE Method of J Determination for a Pressurized Pipe Under Bending," J. of Pressure Vessel Technology Trans.

ASME..Vol. 105,1983, pp.16-22.  ;

5-6 Bamford, W. H. and A. J. Bush, " Fracture Behatior of Stainless Steel,"

ASTM STP668, 1979, pp. 553-577.

i l

l l

l l

l 5640Q:10/021384 5-4

TABLE 5-1 [

]

+BsC Solution Step. AxialStress*(ksi) Bending Moment (in-kip) Jr(in-lb/in;2 5-5

. _ , , ..,,--..--..w-e- ----+e.-~. . . * + - - - - u.*- + a- - - - - ---

. )

(

+  ?.

=

) -

_ .P " -

N

_ e .

e c

e

= a

+

b M

E

= P I

P

= E H

T N

O G

N I

T C

A S

D A

. O L

1

. 5 E

R U

G I

F M

e rm

- , ;I l

, ! l l1 1'

4 t Il  ! i:iI.

, -

• 1

. _ l

\

+a,c :

i

+a,c.e

, Figure 5-2 The[ ] model . [

-1 t

5-7

=

+a,c,o I

.l

' ~ i J

Figure 5-3 The[ ] model af the pipe showing [ ] +a,c,e

[. ]

5-8

+a,c.e t

i l'

i l

+a , c . e -

Figure 5-4 A close-up view of the [ ]

5-9

e

+a,c c mi Figure 5-5 The [ "'"

] pattern in the vicinity of the crack front.

5-10

._a

l l ,

d ab

~

e c

a f

r u

s k

c a

r

- c e

h t

f o

e l

d d

i m

e h

t n

o

. ]

[

6 5

e r

. u g

i F

L mL..

, !L,!i; i, !1i 1 i!  ! I  ;

~

-- +a.c e t

i b

Figure 5-7 [ ] at the pipe +a,c,e end which is subjected to the applied axial and bending stresses.

5-12 ,

~ . . . ..'w-+-=.%.-w3. . .-- ===.e -. . ' . =

TRUE STRESS-STRAIN CURVE

\ -

ll

i 1

1,:

i

.t

I

'i

!i m

L to

i. if l'

i i

4 4

4

+

i Figure 5-8 True stress-strain curve and the[ ] approximation f l o

+ ,

h

.a- c% , s Y

0 j 4. '

!F,

'l4

$(

=

m, J

t!

I 7

4 I

ai

! l .

11 ut 1

b 4

e l l' 1 ',

1:

'a l

SuuW N

Figure 5-9 Schematic of the boundary conditions.

4

'l I +

9e O

+B, f '

~

e h

t t

o d

e i

l p

p a

e r

u s

s e

r p

l a

n r

e t

n .

ie p

ei hp t

e rh ot f

f l

eo ue dc ea hf cr su s

g ne id di as on Li 0

1 5

e r

u g

i F

. mb tl  :! :I  ! i; l. . i' l ,

L

a .>

s 9.

9 .

. +m~ .

d n

e p

e i

p e

h t

o t

d i

e l

p p

a s

s e

r t

s l

i a

x a

n n

f o

i

. n u

e h

t r

o f

e l

u d

e h

c s

g i

n d

a o

_ L 1

1 5

e r

u g

i F

mlm

-  ;' I ItIi!II\i;,,Ii t ! , , !

=-. - -

..  ?

-. . r, _.

.

~

\

)-' -

(

./

-c ;_ _ : -- ,

. t .

. ~

+

/ ^

% s a

'4\ -4r

. t s

i

/

.. 1 -

it -

il

1 .

n ji s m

i u

h

'j,

.I i

l.

f it

i il o

Ji if 4i

[

if lJ J.

.; i Figure 5-12 Loading schedule for the bend moment applied to the pipe end.

s lI .

+

li m

O e =

O 1

, s .

+ a , c ,-

~

+a,c,e Figure 5-13 -

M 5-18

. _, .-,-**--*""'N'"

6.0 LEAK RATE PREDICTIO Q

6.1 INTRODUCTION

,0etailed fracture rsechanics analysis has shown that through-wall cracks in the surge line would remain stable and not cause a gross failure of this RCS component. If such a through-wall crack did exist, it would be desirable to detect the ' leak rate such that the plant could be brought to a safe shutdown condition. The purpose of this section is to discuss the method which will be used to predict the flow through such postulated cracks and present the leak rate calculation results for a [ +a,c,e

] long through wall circumferential crack using the finite element method. The mechanical stability of a slightly larger crack [ ] has been +a, c , e shown in Section 5.0 using a different [ ). +a c.e 6.2 ' _6%ERAL CONSIDERATIONS The flow of hot pressurized water through an opening to a lower back pressure causes flashing which.can result in choking. For long channels where the ratic cf the. channel length, L, to hydraulic otameter, DH , (L/DH ) I5 greater than [ ), both [ ] must be considered. +a,c.e 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.

At_ this point, the flow begins to flash and choking occurs. Pressure losses due t5 momentum changes will dominate for [ ] However, for large +a,c e L/DH values, friction pressure drop will become important and must be considered along with the' momentum losses due to flashing.

6.3 CALCULATION METHOD The basic method used in the leak rate calculations is the method developed by

[ +a c.e

]

J 5640Q:10/021384 6-1

The flow rate through a crack was calculated in the following manner. Figure 6-1 f rom ( ]+ was used to estimate the critical pressure, Pc, for the +a cee surge line enthalpy condition and an essumed flow. Once Pc was found for a given mass flow, the [ ] was +a,c,e found from Figure 6-2 of ( .] For all cases considered, since +a,c.e

( ), ( ) Therefore, this method will yield the two-phase +a,c,e pressure drop due to momentum effects as illustrated in Figure 6-3. Now using the assumed flow rate G, the frictional pressure drop can be calculated using

&Pf=[ ] (6-1) +a,c,e where the friction factor f is determined from the [ ] for which +a c.e che crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ ] RMS as taken from Reference (6-3). +a,c.e The frictional pressure drop using Equation (6-1) is then calculated for the

+a,c,e assumed flow and added to the [

] to obtain the total pressure ,1 rop from the primary system to the atmosphere. That is Surge Line Pressure - 14.7 = [ +a,c e

] (6-2) for a given assumed flow G. If the right-hand-side of Equation (6-2) does not agree with the pressure difference between the surge line and atmosphere, then the procedure is repeated untti Equation (6-2) is satisfied to within an acceptable tolerance and this then results in the flow value through the crack. This calculational procedure has been recomended by [ ] +a,c,e for this type of ( ) calculation. The +a,c.e leak rates obtained by this method have been compared in Reference [ ] +a.c.e with experimental results. The comparison indicated that the method predicts

-leak rate with acceptable accuracy [ **'

).

I 5640Q:10/021384 6-7

L .. .

6.4 CRACK OPENING AREAS Figure 6-4 shows the shai,e of one quarter of the opened cract at the mean radius of the pipe, when the pressure and axial loadings reach their full valuas of [ ),respectively. Figure 6-5 in a similar plot +a,c.e when a moment of [ ] is superposed upon it. Table 6-1 presents +a , c . e :

the coordinates and displacements of the [ ] used to generate the two +a,c,e figures. The area under each curve is evaluated by numerical integration.

Multiplying each of the areas by 4 gives the total areas of the cracks at the mean radius of the pipe, for the two loading conditions. Two leak rates will be calculated basGd cn these areas. These are:

(a) Load A: the leak rate for the loading condition where there is

[ ] +a r c , e ,

(b) Load 8: the leak rata for the loading condition where there is

[

] + ,c.e For load A, the crack area is:

Ag=[ l .

+a,c.e For load 8, the crack area is:

i A8"E I

! +a,C,e 6.5 k[AK RATE RESULTS l

l l Using the [FHG [6-2)) method gives [ ] leak rate for Load Case A +a,c.e

! [' j For Load Case 8, the method gives [ +a,c e 1

Case 8 is considered more realistic since it [ +a,c.e l ] Both calculated leak rates are significantly higher than the leak detection criterion of 1 gpm (Regulatory Guide 1.45).

l 5640Q:10/021384 6-3

6.6 References 5640Q:1D/021384 6-4

a TABLE 6-1 CRACK SURFACE DISPLACEMENT DATA +a,c,e e

O m

9 e

6-5 t_ _ _ . _ _ _ - _ _ _ _ _ . . . . _ _ _ _ _ _ _ _ _ _ _ - _ - _ . _ - - - - - . . .

istes.i

+aac,o 5

5 i

1 t

b a

w 3 .

~

STAGNATION ENTHALPY (102 Stu/lb)

Figurs 6-1 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures l

6-6

- _J

totas 2

+a,c,e o

4 I

e s

w s

w E

a W

!l::

e u

i l ._

~

l LENGTH / DIAMETER RATIO (L/D) i i

I l

l

[

l l

r Figure 6-2 [ ] Pressure Ratio as a Function +8'C

L of L/D 6-7

• +4e '=--.w..- , , , , , , , , , , , , , _ _ _ ,. ,_ , __

-. , . . . , - - - - , . , , , - - - - . . _ , . .-,.,--_.,...,--....-~,-n,- -w--._,,-n

+arcre

+a,c e

- / ,-

~

_ ,- = = = = ^

Figure 6*3 Idealized Pressure Drop Profile Through a Postulatec Crack e

6-8

- . . . , l 4 .

e o.

c a ,

+ c

- + -

1

[

r e

d n

u .

e l

i f

o r

p e

c a

f r

u s

k

. c a

r C

4 6

e r

u g

i F

=

m,e

-  ; i: I .I ; tiit i -

,J44 em i-, .

c.,

O.

m N N s l

+ N

+

. . I

. l S

0 bJ L

G "O

C 3

W W

Q L

Il2.

W U

4 6

3 m

.hd U

M L

U m

t W

L 3

m N

I i

6-10

l 7.0 THERMAL TRANSIENT STRESS ANALYSIS H - The thermal transient stress analysis was performed to obtain the through wall stress profiles for use in the fatigue crack growth analysis of Section 8.0.

The through wall stress distribution for each transient was calculated for e-

1) the time' corresponding to the maximum inside surface stress and, 11) 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 [ +a,c e 1 l loadings were scperimposed on the through wall cyclical stresses to chtain ths l total maxirwum and minimum stress profile for each transient. Linear through well stress distributions were calculated by conservative simplified methods for all minor transients. More accurate nonlinear through wall stress distributions were developed for severe transients. These nonlinear distributions were developed from the [ ]'of similar +a,c.e geometry and transients.

7.1 CRITICAL LOCATION FOR FATIGUE CRACK GROWTH ANALYSIS The surge line stress report (3-1), design thermal transients (Section 7.2),

1-0 analysis data on surge line thermal transient stresses (based on ASNE Section III N83600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis. :The [ +a,c.e  :

) was determined to be the most critical location for the fatigue crack growth evaluation. This location is selected as the worst location (same as determined in Table 3-1) based on the following considerations:

+a,c.e 1) 11) 111) iv) i 56400:10/021384 7-1 a

f , ,

7.2 DESIGN TRANSIENTS The transient conditions selected for this evaluction are based on conservative estimates of the magnitude and the frequency of the temperature fluctuations resulting from various operating conditions in the plant. These

)

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 f::r its application over the

. design life of the plant. A total of ( ) was +a,cee

_ developed for the surge line by considering all the normal operating and upset

. transients in accordance with design specifications (7-1) and the applicable i

system standard design criteria documents (7-2). Some of the data of the  :

applicable criteria documents were refined to more closely represent the ,

'~

transients and to reduce the conservatism in fluid temperature fluctuations ,

and the rate of change of fluid temperature. The thermal transients considered for the fatigue crack growth evaluation are listed in Table 7-1.

7.3 SIMPLIFIED STRESS ANALYSIS i

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

In this method, a 1-D computer program was used to perform the thermal analysis to determine the through wall temperature gradients as a function of time. The inside surface stress was calculated by the following equation which is similar to the transient portion of ASME Section III NB3600, Eq.11:

'5640Q:10/021384 7-2

- -- :: = _ ,_ ...~.;;:u.- -.__-, - - _.,.-._._..: = = = = - m .z ::  :

l .,

, +0,c,0 F --

.. _ =

[ +a,c,(

) The maximum and minimum inside surface stresses were searched from the S g valves calculated for wach time step of the transient solution.

The outside surface stresses corresponding to maximum and minimum inside stresses were calculated by the following equations: +a,c,e f

l l

l 5640Q:10/021384 7-3

--=-_=__:_=_____ . _ _ -

The following material properties at room temperature were conservatively used from the ASME Section III 1974 appendices for the pipe [ ] and +1a , c , e nozzle [ ] safe end in the simplified calculations: +a,c.e

+a,c.e The maximum and minimum lineer through wall stress distribution for each thermal transient was obtained by joining the corresponding inside and outside surface stresses by [ ] The simplified analysis discussed in +a c e this section was performed for all minor thermal transients of Table 7-1

[ ] [ ] through wall +a.c e stress profiles were developed for the remaining severe transients as explained in Section 7.4.

The inside and outside surface stresses calculated by simplified methods for the minor transients are shown in Table 7-2. The comparison of the through wall stress profile, computed for a typical transient by the simi t-d method and and that based on the [

l +a,c.e

] is shown in Figure 7-1. This figure shows that the simplified method will provide more conservative crack growth.

7.4 I 1 DISTRIBUTION FOR SEVERE TRANSIENTS +a c.e The [ ] distributions were developed from the past similar +8'c

analyses for the severe transients, i.e., transient numbers [ +a,c e

] As mentioned earlier, the [ +1i,c e

] is the worst location for fatigue crack growth analysis. A schematic of the surge line and the nozzle geometry at this location [7-3, 7-4) is illustrated in Figure 7-2. In ordar to develop the through wall stress distribution for the Catawba surge iina, the [ +a,c.e

] analysis results of the pressur zer surge nozzles similar to that shown in Figure 7-2, and the results of [ ] (without +a,c e discontinuity) shown in Figure 7-3, were evaluated. These results indicated that the [ ), as shown in Figure 7-2, gives the +a,c.e 5640Q:10/021384 7-4

- - - _ _ _ _ _ _ _ _3 -- _;; _ _ _ _ _ _ _ _ _ _

, .?

worst through wall stress profile for the fatigut N k growth. The Catawba surge line stresses were therefore developed for this critical section which l

~'"

accounts for the discontinuity effects. The comparison of [ +a,c,e

] showed that the discontinuity at the pressurizer surge nozzle increased the [ ] by about 20 +a,c.e

' ~

percent. It was also observed that a reduction of wall thickness by 12 percent caused the stresses to decrease by about 4 percent. These and other similar parametric evaluations f rom the past analyses were used to obtain '

appropriate factors for computing the [ ] stress profiles +a,c e for the critical location of the Catawba surge line. These factors accounted for 1) the effect of thickness differer.ces, ii) the discontinuity effects and j iii) the effect of differences in the fluid temperature fluctuations of the l

severe transients. The through wall stress distributions for the Catawba l surge line were obtained by applying these factors to the finite element stress results of similar nozzle geometry. The stress profiles developed for the critical section by this method are shown in Figures 7-4, 7-5, 7-6 and 7-7 for transients [ ],respectively. The steady state fluctuation +a,c e transients [ ] of Table 7-1 are identical to transient [ ] except +a,c.e

~

for the maximum fluid temperature fluctuation (AT). Similarly, transients

[ ] are identical to transient [ ] except for AT. Hence, the +a,c e stress profiles for transient [ ] were obtained by appropriate +a,c e ratioing of the stress distributions of transients [ ] +a,c,e A comparison of through-wall stress profiles at the pressurizer nozzle with j the profiles at a section with no discontinuities is shown in Figure 7-8 to illustrate the effect of the discontinuity.

i 7.5 QBE LQADS l In addition to thermal transients, cyclical stresses due to 08E event were l also used for the fatigue crack grcwth evaluation. The maximum and minimum inside and outside 08E stresses were calculated from the 08E loading case computer run of the stress report [3-1]. A total of [ +a,c,e f ] were considered for fatigue crack growth. The OBE l

stresses are as follows:

i

[_

1 56400:10/021384 7-5

Inside Surface outside Surface Stress (ksi) Stress (ksi)

_ +arcae Maximum Minimum 7.6 TOTAL STRESS FOR FATIGUE CRACK GROWTH l

l The total through wall stress at a section was obtained by superimposing the pressure load stresses and the stresses due to [

+a,c e

] Thus, the total stress for fatigue crack growth at any point is given by the following equation:

+a,c,e To,tal for Fatigue =

Crack Growth

_ _I The thermal expansion moments were revised to account for the difference in the coefficients of thermal expansion between the Stress Report and the current ASME code for transients [ ] (which are the most severe) +a,c e by the following equation to reduce conservatism.

i M, = [ ] (7.8) +a , c , e l

where. .

+a,c e i.

l l

i 5640Q:10/021384 7-6

+c.c o

u. . . . .

[ +a,c e

]

7.7) for calculating the total stresses, are sununarized in Table 7-3.

7.7 REFERENCES

+a,c.e 7-1 7-2 7-3 l

7-4

i. 7-5 J

l l-l l

l l

5640Q:10/021384 7-7 i

a,. -..-----,s. . - ~ - ~ - -.- . - ~ , . . - . - --- . . - - - -. .

r

.q ' f'

, s: ' e-6 e TABLE 7-1 Jhr- ^~.O '

~r - -

- THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH

+r,c.e

. Trans. No. of'

.e-: .

Trans. No. of

- lNo lestriction-. ' Occurrences 'No. Description Occurrences I

'4' e

m f*

k 4

}

L t

l t

4

-s

. . 55570:10/0210841 . 7 s. .

8unt d e t ' " * ' '

'M'

k t TABLE 7-2 -

i TilERNAL TRANSIENT STRESSES BY SIMPLIFIED ANALYSIS i

Max Peak Stress (Sp ) Min Peak Stress.(Sp )

Corre- Corresponding Number Inside sponding Inside Dutside

.. Tran- of Max Dutside Min

!! sient Min Occur- Time S S Time S S

,,No. Description rences (ksi)

(sec) (k!1) (sec) (k!!) (k!i) ---

+a,cee i

?

m

'i

TA3LE 7-3 PRESSURE. DEADWEIGHT AND THERMAL EXPANSION STRESSES FOR FATI6UE CRACK GROWTH Axial Bending Force Moment Transient F,1 BM Stress (ksi)

No. Loadina (k) Inside (ft-k) - Outside

_ +arcre

_ ~

$6400:10/021384 7-10

---y ..

+a,c e

'+ -

]

' AXIAL

. STRESS

.(ksi) ,

1 I

t 4

t I' i i

- I f

J FIGURE 7-1: COMPARISON OF TYPICAL MAXIMUM AND MINIKJM STRESS PROFILE COMPUTEDBYSIMPLIFIED[ ] .

l 7-11 L _ - . - - _ .____ .. .--. -. - - _ . . - . _a

.s5-

.e .

'f 4l t.

q .

i .

+a c.e i,

l' _

i i

i -

4 1

-l 1

-m il

_3 N

9 4

+

ii 4

i A

2 b

p

'{ f 4

ir.

4 s

[

j i l

i FIG'JRE 7-2: SCHEMATIC 0F SURGE LINE AT [ ] +a,c.e l

r l-- . l

. e- .

s 'e l l

%c,- .s= .

M@

f. .

W

• ErCr(

M a

4 9

4 i

.- -e e

O FIGURE 7-3: SCHEMATIC OF [ ]

.WITHOUT DISCONTINUITY

~

7-13

. = __ . . . _ . ._.

%,.; , .? _

&&3(;e

L'.. , .

+a.c,e

~: AXIAL

~ ! STRESS' (ksi) m. .

I.

>^-

L 4

.y

r. .

I ^

f

- FIGURE;7-4: MAXDtVM AND MINIMUM STRESS PROFILE AT CRITICAL LOCATION;FOR UNIT l.0ADING TRANSIENT o

, 7-14

. , . em uu.?pyw N --w m .-- & a=4 , "##'

- _.___.____e_.____ ' _ _ _ _ _ " _ _ ~ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ . **'".'*_'.r'*'_

..-v--:#,'

, +a,c,o

)

\

l l

t A

AXIAL .

STRESS (ksi)

L FIGURE 7-5: MAXIMUM AND MINIMUM STRESS PROFILE AT CRITICAL LOCATION FOR STEADY STATE FLUCTUATION [ ] TRANSIENT +a , c , e 7 15 - -- -- - -

--e e -

-ry - v-y- my -

% w e----,- --- - m t --y++- e y --- y.--t*m-

• M"<*T

n .

i.-L_ _ . - . -

+a,c, AXIAL STRESS (ksi) l .

l FIGURE 7-6: MAXIMUM AND MINIMUM STRESS PROFILE AT CRITICAL LOCATION FOR RANDOM FLUCTUATION [ ] TRANSIENT +a,c,e 7-16

_ = - . . . - . . - - . . . - - - - - -

, - . . . . . _ =-

.g. +a,c e

~'

- a:

b -

AXIAL' .

STRESS (ksi) e-s

- r l

! s FIGURE 7-7: MAXIMJM AND MINIMUM STRESS PROFILE AT CRITICAL LOCATION FOR [ ] CYCLING TRANSIENT +a,c.e ,

. .___. _ _ . _ . , . _ _ . 7-17 ---

-. , . - - . - , _ - - = _ . .

>> * - 1

-a c,e t

AXIAL STRESS (KSI) i FIGURE 7-8: EFFECT OF DISCONTINUITY ON THROUGH-WALL STRET.S PROFILES i

7-18

y .

3:n <

3, F #

y 8.0 FATIGUE CRACK GROWTH ANALYSIS 1

~x r o

~

l' The fatig e crack growth analysis was performed to determine the effect Lof. the design th'ermal transients in Table'-7-l'along with the OBE load transient. The analysis was perfomed for the critical cross section

~

of the:medel which is identified in Figure 7-2. A range of crack depths

[ ;was postulated, and each was subjected to the transients in Table 7-1 as wellias the OBE Load-) transient. ,

8,1 ' ANALYSIS PROCEDURE-The[ fatigue crack growth analyses _ presented herein were conducte pisame manner as suggested by Section XI, Appendix A of the ASME Boiler iand Pressure; Vessel 5 Code.

' The~ analysis procedure involves assuming an

^

initial flaw ex ist s 'at some~ point and predicting the growth of that flaw due to an imposed. hieries of stress trar.sients.

~

The growt,h of a crack per -

loading cycle is dependent on the range. of applied stress intensity factor AK'y ,-Q 2e following relaMon: -

k=CoaK" y (8-1) where "Co" and the exponent "n'? ~ are materi'ai proper;ies, and X is defined later, in Equatton (8-3). For inert environments these material7 properties are constants, but for some water environments they are.cependent on the level of mean stress present during the cycle. This can be accounted for by: adjusting .the 'value of "Co" and "n" by a function of the ratio of mini-rnur to maximum stress for.any given transient, as will be discussed later.

Tatigue crack igrowth properties .of stainless steel in a, pressurized water

.environmer,t'have been used in the analysis.

'The input recpired=for a fatigue crack growth analysis is basically the

.information 3ecessary to calculate the parameter 1.rsy , which depends on cracx andItruSture geometry-fand the range of applied stresses in the , area wher the crack exists..Oncef 2X7 is calculated, the "ov .h due to that particular cycle can be calculated by Equation (3-lb 'b- increment of growth is then

. iadded to the criginal crack -size, the ^ el. -ad, and the analysis proceeds

_.. .wto the next tr%csient.

- .j ' The procedure is corunued in this manner until all

. the transients have been analyced.

1 o

. . _ i z ara. in c +- - _

J .

] . .

The crack tip stress intensity factors (K )g to be used in the crack growth analysis were calculated using an expression which applies for a semi-ellip-tic surface flaw in a cylindrical geometry [8-1].

The stress intensity factor expression was taken from Reference 8-1 and was calc'ulated using the actual stress profiles at the critical s,ection.

The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial:

c (x) = A0+Aj f+A2 ( )2 + (k) (B-2) she stress intensity factor K y(e) was calculated at the deepest point of the crack using the following expression: +anc *

.(8-3)

+a,e,

~

where'

.J L.

Ca'culation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law detemined fr::m considera' tion of the available data for stainless steel in a pressurized water environment. Thds law allows for the effect of mean stress or R ratio (K I dn IK' m) on the growth rates.

8-2

^~ ^

__ _-- _ : T. ^

~

,iL ,. '

I ..

The reference crack growth law for stainless steel in a pressurized water environment was~ taken from a collection of data [S-2 code curve is available, and it is defined by the following ectation: J t

h=(0.0054x10~3)(X,ff)4.48 (8-4) where -X,ff = (Xg ,,x) (1-R)M K

I min "I max h = crack growth rate in micro-inches / cycle 8.2 _RESULTS Fatigue crack. growth analyses were carried out for the critical cross section.

Analysis was completed for a range of postulated flaw sizes oriented circum-ferentially, and the results are presented in Table 8-1. The postulated flaws are assumed to be six times as long as they are deep. Even for the largest postulated flaw of [

] the result shows that flaw growth through the wall will not occur during the 40 year de-sign life of the plant.

For smaller flaws, the flaw growth is significantly lower. For example, a postulated [

] inch deep flaw will grow to less than 1/2 the wall thickness. These results also confim operating plant experience.

There have been no leaks observed in Westinghouse PWR surge lines in over 400 reactor years of operation. .

8.3 REFERENCES

8-1 McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for Internal Longitudinal Semi-Elliptical Surface Flaws in a Cylinder Under Arbitrary Loadings", Fracture Mechanics ASTM STP 677, 1979, pp. 365-380.

8-2 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.

8-3

--, ,...-._.__,.--+_,,c , ....v_,,,.. _-,__,._,,_.__m., . _ , _ . - . _ - - , . . - . - - - . . - - . , , _ . - . . . - - , - . . . . - - -

TABLE 8-1 FATIGUE CRACK GROWTH RESULTS THICKNESS = C ] +"'*

INITIAL CRACK LENGTH AFTER YEAR CRACK LENGTH +a,c,e (IN.) 10

( 20 30 40

_J

• This is conservatively taken as the minimum thickness of the counterbore region.

_ . P. ..t . . - - - - - - - .-

9eO CONCLUSIONS A mechanistic fracture evaluation of the Catawba Units 1 and 2 pressurizer surge line was perfomed. The worst location in the pressurizer surge line was identified [ +a,c,d

] The critical crack lengtn at this location was calculated

]. Finite element elastic-plastic analysis was per- +a,c, as'[ '

fomed using a th' rough-wall flaw [ +a,c,

. ] The applied [ ] +a , c ,

was calculated corresponding to the maximum applied load including the Safe Shutdown Earthquake load. The applied [ ] is thu'suless than +a,c,

[ ] for the material. These results demonstrate - +a,c, that_[ ] crack will remain stable when subjected to maximum load.ing +a,c, conditions considering both global and local failure mechanisms.

The leakage through a crack [ ] long'[ ] was calculated as [ ] +a,e gpm under the nomal operating loads. The Catawba Plant has an RCS pressure boundary leak detection system which is consistent with the requirements of Regulatory Guide 1.45 ano can detect leakage of 1 gpm in one hour. Thus, there is a factor of at least [ ] between-the calculated leak rate and the Catawba +a, plant leak detection systems capability. .

Fatigue crack growth was detemined for postulated inside surface flaws using plant design transients. Crack growth results indicated that even a postulatec surface flaw which is [ ] .of the wall thickness in depth will not penetrate +a,e the wall over the plant life. Thus, there is no known mechanism which could cause a through-wall crack of the type assumed in the stability calculations.

The incidence of stress corrosion cracking is eliminated by appropriate water i chemistry control. Furthermore, operational occurrences will not create water hansner in the surge line.

Based on the above, it is concluded that guillotine breaks in the pressurizer surge line should not be considered as a part of the structural design basis-of the Catawba plant.

9-1

4 5

9 5

\

l.

l APPENDIX A sEQUILIBRIUM OF THE SECTION 1

l A-1 . _ . . . . _

m ,

APPENDIX A

.The internal stress- system at the crack plane has to be in equilibrium with q

the applied. loading i.e. the hydrostatic pressure P, axial forc2 F and the bending moment-M .

b The angle B which identifies the point of stress inversion follows from the equilibrium of horizontal forces (Sea Figure A-1).

This is 1 t- .

1- +a,c,e Solving fdr 8, t- -

+a,c,e The ex'ternal bending moment at the instant of failure follows from the equilibri~um of moments, which is most easily taken around the axis 1-1. Thus Mb can be determined from

_, +a,c,e 55070:10/011284 A-2

i

,G __

'

• Q N

+

1 7

O I

N k

me su E

m O

w C

O N

b u

i (22 o n

.3 7-egne W

• pm 7

N W E

c L

l3

.e b-N Y -

b

\

l l

A-3 i-i

- . . ~ . . - - - - - . - - _ . - . . _ _ . . . _ . .___,,,_ _. ,

4 t .1

APPENDIX B VERIFICATION OF THE FINITE ELF. MENT RESULTS e

h 5640Q:10/021384 B-1 J

- ' ' - ^

r

~

g

, o The purpose of the verification presented herein is to assure the correctness of'the fracture mechani:s analysis for the pipe. Both the K -values g due to the pure axial stretching and the pure bending are investigated. The outer fiber stresses corresponding to the maximum applied bending moment are investigated also.

(1) It!g.Jg' for a circumferential1v cracked nine subiected to oure 'bendina.

The elastic solution for this problem has been studied by Folias (B-1]

and others. Under the present geometrical and loading conditions, the n is given by I +a,c,e O

m m

5640Q:10/021384 __

8-2

er

[ +a c r e, 3

+a,c,e Substituting [ ]

ksi/in. The difference between the results by Eq. (B-3) and the VCE method is 4.5 percent. Better agreement would have been obtained if the bending effect,'i.e., the c factor, had been included in the calculation. However, the 4.5 percent errror is' acceptable for the engineering analyses.

(2) Eg_due to oure bendina The Kg for a circumferential1y cracked pipe subjected to bending may be estimated by taking the average of that produced by the tensile outer fiber stress, a b, and by the fiber stress at the location of the crack tip, s' . These stresses are shown in Figure B-1. The relation between a and a' is given by b

a' = a bcos n (B-5) where a= crack angle .(see Figure B-1). Therefore the Kg due to bending is K

g

= [ ] 'B-6) +arere

. Inserting Eq. B-5 into Eq. 8-6 and taking [

l

! ], one obtains:

(s-7) +a,c K

g,

= [ ]  :

i=

l A pure bending load with [

] was + arc, used for the VCE analysis and the [ ] produced [ +a,c,

] of

+a,c,

] This [ ] is converted to the [

+a,c,

! [- ]

l 5640Q:10/021384__ B-3 - .

• • -w s- *w--g-1-w- -$i- pg mat,+-,-,.mm*-,wswim-n-.--a-g--.i-+-----+,,emv-e+me pW

..,.s-.---i-.g.meew-gw -m-ag.m-w,-y-ww'

.c -

C7 Substituting [ +a,c,e

] The difference in this case is about 3 percent.

It need be noted that Eqs. B-3 and B-7 are valid only for the elastic deforination, When loads increase the linear elastic theory underestimates _the [ ] The deviation is considerable when +a,c,e large plastic zone in the crack tip region is developed. However, these equations can be used for reference purpose. This means that the actual

[ ] should be always greater than those given by Eqs. B-1 and +a,cre B-6. This condition or requirement is met for the present analysis.

4 (3) Check on the Outer fiber stress In addition to examining the [ ] values, the axial stress +a,c,e which directly relates to the open mode of fracture is examined herein.

Only the outer fiber stress on the tension side is checked. Since there is no plastic deformation in the region remote from the crack up to the maximwm applied loads, the bending stress can be computed by a z (B-8) b=

where M = bending moment I = moment of initia z = distance from the neutral axis.

l l: Based on the geometrical data employed in the present analysis, [ +a,c,

] For [ +a,c, L

] In addition to l

l I

i 56400:10/021384 8

_ __ _, _ _ _ _ _ _ -4__ _ _ __ _ _

g T-the bending stress, there is an axial stress, e,, of ( ) ksi +a,c,e constantly acting on the pipe. Therefore, the combined fiber stress at the Guassian point investigated is

' " ' + '

tot a b

[

• 1 +a,c .

The corresponding stress given by [ ] is [ ] ksi. The error is +a,c e 0.4 percent.

Reference

. B-1 Folias, E. S., "Ori the Effect of Initial Curvature on Cracked Flat Sheets," Int. J. of Frac. Mech., Vol. 5, 1969, pp. 327-346.

=

0 5640Q:1D/021384 8-5 7 e -m---,w-- e-- e eeev m- - - , m+- ---

e-y-e ,w-w---*yein+- w---+3w+ --.,e#-ppp -+wwwn-e w -wg- + w-+p-g---wwgrw-,--w----r-------r-seg--e e++v--ty-NvN

g .. ,

s.

l -l .;; e i ; , ;

,. r.

, ' . ., : -. t f

I 4

'Z l.

t

=

/:b

• ~

/..-. .

I s _,

j. .

e a

. y.-

1 .

R n l-lg l-y , ,/

/ /

i M

e-=I.O b

R

~ '

M cosa o' = 1 (RO cos2) = :b a =: crack ar.gle s '

9 m

- Eicyr9 3-l'. Auxiliary didgPIT # 0F C9fiVE~iCri Of 3 GE!iOr. 3-6.

1 B .. .

' [ 3, , , n.

, ..--c, ~..w-~~~-"~~~k -

_ _ _ _ _m..= ._

ENCLOSURE C CNS 3.6 PROTECTION AGAINST DYNAMIC EFFECTS ASSOCIATED WITH THE POSTULATED RUPTURE OF PIPING General Design Criterion 4 of Appendix A to 10CFR50 required that structures, systems, and components important to safety be protected from the dynamic effects of pipe failure. This section oescribes the design bases and design

- measures to ensure that the containment vessel and all essential equipment in-side or outside the containment, including components of the reactor coolant pressure boundary, have been adequately protected against the effects of blow-down jet and reactive forces and pipe whip resulting from postulated rupture of piping.

l Except where specified otherwise in Table 3.6.1-3, criteria presented herein regarding break size, shape, orientation, and loca' tion are in accordance with

.the guidelines established by NRC Regulatory Guide 1.46, and include considera-

~ tions which are further clarified in NRC Branch Technical Position MEB 3-1 and APCSB 3-1 where appropriate. These criteria are intended to be conservative and allcw a high margin of safety. For those pipe failures where portions of l these criteria lead to unacceptable. consequences, further analyses will be performed. However, any alternative criteria will be adequately justified and fully documented.

3.6.1 POSTULATED PIPING FAILURES'IN FLUID SYSTEMS INSIDE AND OUTSIDE CONTAINMENT 3.6.1.1 Design Bases 3.6.1.1.1 Reactor Coolant System The Reactor Coolant System, as used in Section 3.6 of the Safety Analysis Report, is limited to the main coolant loop piping and all branch connection nozzles out ,

l to the first butt weld. Dynamic effects are only considered for pipe breaks l postulated at branch connections. The particular arrangement of the Reactor Coolant System, building structures, and mechanical restraints preclude the formaticn of

_ l plastic hinges for breaks postulated to occur at the branch connections. Con-

.,MEBL sequently, pipe whip ~ar.d jet impingement effects of the postulated pipe break Q102 l components.atThis

- these locations resiraint will not result configuration, in unacceptable along consequences with the particular arrangement to essentia of the Reactor Coolant System and building structures, mitigates the effects of the jet from the given break such that no unacceptable consequences to essential components are experienced.

The application of criteria for protection against the effects of postulated breaks at tne branch connections results in a system response which can be

-accommodated directly by the supporting structures of the reactor vessel, the steam generator, and the reactor coolant pumps. The design bases for postulated breaks in the Reactor Coolant System are discussed in Section 3.6.2.1.

P 3.6-1 Rev. __

o .

[- ~

CNS Systems which do not contain mechanical pressurization equipment are excluded from moderate-energy classification (e.g. , systems without pumps, pressurizing tanks, boilers, or tho:: which operate only from gravity flow or storage tank water head), however, limited failures are assumed to occur for the purpose of considering the effects of flooding, spray, and wetting of equipment in the station. analysis.

TN identification of piping failure locations will be performed in accordance with Section 3.6.2.

3.6.1.1.2.1 Interaction Criteria

'The following criteria define how interactions shall be evaluated. The safety evaluation of each interaction is described in Sections 3.6.1.3 and 3.6.1.1.5.

.a) Environmental Interaction An active component (electrical, mechanical, and instrumec n and control) is assumed incapable of performing its function u r periencing environmental conditions exceeding any of its environmental ra ngs.

b) Jet Impingement. Interactions Active components (electrical, mechanical, and instrumentation and control) subjected to a jet are assumed failed unless the active component is en-closed in a qualified enclosure, the component is known to be insensitive to:such an environment, or unless shown by analysis that the active function

'will not be impaired.

c)- Pipe Whip Interaction l A whipping pipe is not to be considered to inflict unacceptable damage to other pipes of equal or greater size and wall thickness.

.A whipping pipe is only considered capable of developing through-wall leakage cracks in other pipes of equal or greater size with smaller wall thickness.

An active component (electrical, mechanical, and instrumentation and control)'is assumed incapable of performing its active function following impact by any whipping pipe unless an analysis or test is conducted to show otherwise.

3.6.1.1.3 Protective Measures 3.6.1.1.3.1 Reactor Coolant System l' The fluid discharged from postulated pipe breaks at branch connections will produce reaction and thrust forces in branch line piping. The effects of these 3.6-3 Rev. _

CNS c) Safety-related portions of the Main Steam and Feedwater Systems are Duke Class B. Class B system materials, fabrication, nondestructive examinations and documentation'are'in accordance with ASME Section III, Class 2.

d) The guard pipe on the main steam piping inside-containment is extended over the length of the vertical portion of the piping to prevent jet impingement on the _ ice condenser doors following a postulated rupture.

e) As a result of a Duke-NRC meeting in May 1976, guard pipe was removed from the main steam and main feedwater piping in the doghouse which was 3

originally designed with guard pipe. Breaks in the Main Steam piping will be postulated based on consequence except for the break exclusion region

, inside the doghouse. For the main steam piping which is part of the break

(~ ' exclusion region an augmented inservice inspection will be performed, as ,

, discussed in Section 6.6, for welds where no break is selected. Breaks in the main feedwater will be selected as outlined in Section 3.6.2.

2 3.6.1.1.3.4 Control Room Protection from Postulated Piping Breaks The control room is located on the top floor of the Auxiliary Building and is

. bounded on the north'and south sides by electrical penetration rooms which contain no piping. The east side of the control room is bounded by the equip-ment area housing the control room ventilation equipment. Piping in this area consists of low pressure. low volume chilled water and low pressure, low volume heating steam. 01 the west side, the control room is bounded by the computer room and supporting areas. Piping in this area consists of sanitary waste and vent piping, drinking water and instrument air; none of which are high energy systems. Immediately t,elow the control room is the cable room containing no p hing. .

Based'on the above physical parameters, the control rcom is structurally isolated from areas containing high energy systems; therefore, there are no unacceptable consequences to the control room from the postulated break of high-energy piping systems.

3.6.1.1.4 Acceptability Criteria The-capability to eventually. achieve a cold shutdown condition is not jeopardized even if the pipe failure is followed by a single active failure. The system requirements and available redundancy are determined on a shutdown logic diagram, or a required equipment list for mitigating the effects of the postulated failure.

Repair of failures is considered to assure achievement of the cold shutdown condition where such repairs can be shown to be practical and timely, and pro-vided the unit can be held in a safe shutdown state during the time required for the repair.

3.6.1.1.5 Leak-Before-Break Analyses for High Energy Piping Other Than Reactor Coolant System Developments in fracture mechanics make it possible to prove applicability of a " leak before break" pipe failure concept for certain piping systems in addition to the Reactor Coolant System. Under this concept the size flaw 3.6-6 Rev. _

1

CNS which results in detectable leakage is shown to be much smaller than the size flaw which would lead to a catastrophic acuble-ended pipe break (DEPB). It is shown that there is no mechanism for developing a large break without going through an extended period during which the crack would leak copiously and corrective action would be taken in accordance with plant Technical Specifications.

The " leak before break" analysis involves, in general, the following steps:

1. Compilation of data on piping loads, materials, geometry, and_ transients.

2. Determining piping stresses and initial flaw sizes and locations. -

3. Performing fatigue crack growth calculations using fracture mechanics techniques.

4. Evaluating crack stability and calculating leakage rate from a stable crack.

5. Comparing leak rate predictions with leak detection capability.

If the leakage rate from a stable crack is shown to be within the limits of leak detectability, then the applicability of the " leak before break" concept has been demonstrated.

For high energy piping where applicability of the " leak before break" concept has been demonstrated, the postulation of DEPB's per Sections 3.6.2.1.2.1 and 3.6.2.1.2.3 is replaced for certain consequence analyses by postulation of cracks. The crack size and orientation is defined by the detailed fracture mechanics analysis. The dynamic effects of pipe rupture described in Sections 3.6.2.2 and 3.6.2.3 are not applicable. Compartment pressurization effects are limited to consideration of pressurization due to the crack. This elimination of DEPB also eliminated the need to perform the consequence analysis for pipe whip, jet impingement, and compartment pressurization associated with these DEPB's. For flooding analyses and environmental effects (temperature, humidity, and water spray), the DEPB is retained and Sections 3.6.2.1.2.1 and 3.6.2.1.2.3 apply. Table 3.6.1-4 lists the lines for which

" leak before break" concept applicability is demonstrated.

3.6.3.2 Description of Piping System Arrangement Separation is the primary consideration in piping system layout and arrangement.

Where physical separation is not feasible, protective devices are provided to protect essential components.

3.6-6a Rev. _

New Page

L I

P i

CNS F

E 3.6.2.1.1.1 Postulated Piping Break Locations and Orientations k Reference 1 defines the original basis for postulating pipe breaks in the IF reactor coolant system primary loop. Reference 1.a provides the basis for MEB eliminating from certain aspects of design consideration previously postulated Q13 reactor coolant system pipe breaks, with the exception of those breaks at i

branch connections. See Table 3.6.2-1 and Figure 3.6.2-2.

E 3.6.2.1.1.2 Postulated Piping Break Sizes

,h For a circumferential break, the break area is the cross-sectional area of the pipe at the break location, unless pipe displacement is shown to be limited by -

? MEB analysis, experiment or physical restraint.

- Q105 I

3.6.2.1.1.3 Line Size Considerations for Postulated Piping Breaks b Sranch lines connected to the Reactor Coolant System are defined as "large" for the purpose of this criteria as having an inside diameter greater than 4 inches c

l up to the largest connecting line. Where postulated, pipe break of these lines cesults in a rapid blowdown of the Reactor Coolant System and protection is

, basically provided by the accumulators and the low head safety injection pumps E (residual heat removal pumps).

E l 3.6.2.1.2 General Design Criteria for Postulated Piping Breaks Other Than p Reactor Coolant System

" a) Station design considers and accommodates the effects of postulated pipe breaks with respect to pipe whip, jet imp:agement and resulting reactive forces for piping both inside and outside Containment except as specified

- in section 3.6.1.1.5. The analytical nathods utilized to assure that l- concurrent single active component failure and pipe break effects do not jeopardize the safe shutdown of the reactor are outlined in Section

_ 3.6.2.3.

- b) Station general arrangement and layout design of high-energy systems

utilize the possible combination of physical separation, pipe bends, E pipe whip restraints and encased or jacketed piping for the most practical b design of the station. These possible design combinations decrease postu-F lated piping break consequences to minimum and acceptable levels. In all E cases, the design is of a nature to mitigate the consequences of the break E

- so that the reactor can be shutdown safely and eventually maintained in a cold shutdown condition.

b y c) The environmental effects of pressure, temperature and flooding are con-R trolled to acceptable levels utilizing restraints, level alarms and/or other warning devices, and vent openings.

t R~-

3.6-9 Rev. _

L

k. ... __ - , . .

7 . _

&,. s

'CNS.

i p). -Minimum essential component and systems performance is provided as required for the type;of break.

7q) .The effects of pipe ruptures are.not allowed to result in offsite doses in excess of 10CFR100 allowable limits.

- r) ' Operability in an environment is assured for all equipment required to mitigate the break by the equipment specification requirements based on -

' conservative design conditions.

Js) Emergency procedures are prepared that are to be followed after a postulated

^

• piping break for high-energy systems as required. .

-3.6.2.1.2.1 Postulated Piring Break Locations For High-Energy Piping Systems

. Systems identified as containing high-energy piping are examined by detailed idesign drawing review for postulated pipe breaks as defined below. ' Systems inalyzed for consequences-of postulated piping breaks are listed in Table.

L -3.6.1-1. Refer to Section 3.6.1.1.5 for exceptions to the following criteria

~

J due' to application' of- the " Leak Before Break" concept.

' Terminal ends 2are considered as piping originating at strtctures or components

. (such as~ vessel and equipment nozzles and structural piping anchors) that act

't as rigid constraint. to.the piping thermal. expansion. Typically, the anchors o ' assumed.for the piping code stress analysis would be terminal ends. The branch

connection to the main run is one of the terminal ends of a branch run, except intersections of runs of comparable size and fixity which have a significant effect'on the' main run need not be considered terminal ends when the stress

. analysis model includes both the run~and branch piping and the intersection is

=not. rigidly constrained to the building structure.

a) (Breaks in Duke" Class A piping are postulated at the following locations (see Table 3.2.2-3 for class correlations):

1) :The terminal ends of the pressurized portions of the run.

2).. I At intermediate locations selected by either one of the following methods

1. ~

.i) -At each weld location of potential high stress or fatigue, such as pipe fittings (elbows, tees, reducers, etc.), valves, flanges, and: Welded attach.nents, or ii) At all intermediate locations between terminal ends where the

'following stress and fatigue limits are exceeded,

a. The maximum stress range shall not exceed 2.4 S, except as noted below.

3.6-13 Rev. _

l

CNS l 3.6.2.1.3 Failure Consequences ,' ,

-[ The. interactions that are evaluated to determine t*w failure consequences associated with postulated pipe breaks are dependent on the energy level of the contained fluid. They are as follows:

a) High-Energy Piping

1) Circumferential Breaks and Longitudinal Splits a) Pipe Whip (displacement)

MEB- b) Jet Impingement

'Q23 c) Compartment Pressurization d) Flooding e) Environmental Effects (Temperature, humidity, water spray)

2) Througnwall leakage cracks a) Environmental Effects (Temperature, Humidity) b) Flooding lb) Moderate-Energy Piping

1) Through-wall leakage cracks MEB a) Flooding Q23 b) Environmental Effects (Temperature, humidity, water spray) c) Water Spray For high energy piping there are certain exceptions as detailed in Reference I la'for the reactor coolant loop and Section 3.6.1.1.5 for all other piping systems.

3.6.2.2 Analytical Methods to Define Forcing Functions and Response Models 3.6.2.2.1 Reactor Coolant System Dynamic Analysis This section summarizes the dynamic analysis as it applies to the LOCA result-ing from the postulated design basis pipe breaks at main reactor coolant branch line connections.Further discussion of the dynamic analysis methods used to verify the design adequacy of the reactor coolant loop piping, equipment and supports is given in Reference 1 as~it pertains to postulated breaks at Ibranchconnections.

The particular arrangement of the Reactor Coclant System for the Catawba Nuclear Station is accurately modeled by the' standard layout used in Reference 1 and l -

presented in Reference 1.the postulated branch connection break locations do not cha In additicn, an analysis is performed to demonstrate that at each postulated l branch ' connection break location the motion of the pipe ends is limited so as to preclude unacceptable damage due to the effects of pipe whip or large motion of any major components. The loads employed in the analysis are based on full pipe r.rea discharge except where limited by major structures.

3.6-18 Rev. _

Tabla 3.6.1-3 (Prge.3) to this closed valve. ~ ting at structure or components that act as rigid constraint to the piping thermal expan-sion. . Typically, the anchors assumed for the code stress analysis would be terminal ends.

Stresses in the system either side of the closed valve will be about the same; therefore, terminal end classification based on constraint and high stresses are not applicable. Duke reviews these closed valve locations to assure high stresses are not developed as-a result of rigid constraint from nearby anchors of com-ponent connections in the non pressurized portion of the piping.

APCSB 3-1, Appendix B and C SAR Section 3.6.2.1.2.1 In Appendix B, pipe break locations are specified Duke criteria specifies that if the f.hreshold for ASME Section III Code Class,1, 2, and 3 piping stress levels.are not exceeded, then no inter-such that a minimum of two intermediate breaks are mediate breaks are postulated.

selected per run although threshold limits are not exceeded (for ASME Section III Code Class 1, 2, and 3 piping). In Appendix C, a minimum of either two or one intermediate breaks within the boundary of each compartment is specified.

APCSB 3-1, Appendix B SAR Section 3.6.1.1.5 The criteria used to determine the design basis In specified cases the " leak before break" concept piping break locations involves the assumption of is employed, as explained in Section 3.6.1.1.5.

dnuble-ended pipe break locations selected according to stress and/or fatigue analyses, or consequence analysis.

MEB 3-1, Section B.1 SAR Section 3.6.1.1.5 The criteria used to determine the design basis In specified cases the " leak before break" concept piping break locations involves the assumption of is employed, as explained in Section 3.6.1.1.5.

double-ended pipe break locations selected according to stress and/or fatigue analyses, or consequence analysis.

Rev. __

t l_

, Tcble 3;6.1-3-(PIga 4)

MEB 3-1, Section B.1.b(6)- -SAR'Section 3.6.2.4 ,

Section B.1.b(6) requires'that guard pipe ass'emblies ' Duke criteria is different from NRC criteria as between containment-isolation. valves meet the follow- described and: justified below:

.ing requirements: ,

. Guard pipes provided between containment isola-

a. The design pressure and temperature should not tion' valves are designed in accordance with SAR be less than the maximum operating temperature Section:3.6.2.4. Guard pipes are subjected to-

. -and pressure of the enclosed pipe under normal plant conditions.

a pressure. test as required by the material

specification before welding to'the penetration-assembly.

f b .' .The design ~ stress limits'of Paragraph NE-3131(c)- ..

should not be exceeded under.the-loading ~asso-It'is impractical to test guard pipes.in the ciated with design pressure and temperature in finished penetration assembly due to the con-l combination with the safe. shutdown earthquakes. . figuration and potential damage to internal i

process pipe and associated insulation. Inde-

1. c. Guard pipe assemblies should be subjected to a pendent design analysis have been conducted to a

single pressure test at a pressure equal to de- provide assurance that Duke penetration designs sign pressure. .are acceptable. In addition, the' extent of NDT.

conducted on guard pipes to flued head butt weld is such to assure integrity.of design.

I MEB 3-1, Section 8.1.c(1)

St.R Section- 3.6.2.1.2.1.

Intermediate breaks in Class 1 piping are postulated Duke criteria states that if.there are no

! at the two highest stress locations based on intermediate-locations where S exceeds 2.4 S l Equation (10) if two intermediate locations or U exceeds 0.1, no intermediate breaks are" l

cannot te determined by application of Equations . postulated.

l (10), (12), and (13) or U>0.1.

j MEB MEB 3-1, Section B.1.c(2)

Q35 SAR Section 3.6.2.1.2.1 Intermediate breaks in Class 2 and 3 piping are Duke criteria specifies that if the threshold postulated where the stresses exceed 0.8 (1.2S stress. levels are not exceeded, then no inter-l S)butatnotlessthantwolocationsbasedoN+

3 mediate breaks are postulated.

l' hTghest stress. Where the piping consists of a straight run without fittings, welded attachments, and valves, and all stresses are less than 0.8 (1.2Sh + SA ), a minimum of one location should be chosen based on highest stress.

Rev. __

Tablo'3.6.1-3_(Page 5)l ' ~

MEB'3-1. Sections'B.I.c(3)

--SAR Section 3.6.2.1.2.I' > yl a

. Breaks in'non-nuclear piping.should be postulated.

at the following_ location: '

. Duke criteria'is generally equivalent to NRC'cri-

.teria'as described and justified below:

k W

'. Q35 ' a. . Terminal _. ends ,'

,  ; Breaks in Duke' Class F piping (non-nuclear,

- ' seismic) are' postulated at terminal ends and at-

b. At each intermediate.pipeifitting,: welded intermediate 1ocations based on the use of ASME

-attachment, and valve.

3 l Section III analysis techniques,- the same as' ._

Duke Class'B and C piping. Duke Class F piping' t

.is' constructed in accordance with ANSI B31.1 and Lis dynamically analyzed and restrained for seis- .

micl loadings similar to ASME Section III piping. t Materials are specified, procured, received, j stored, and issued under Duke's QA program simi-Llar to:ASME-Section III materials except'that'

! certificate.of compliance in lieu of mill test

reports are acceptable on minor. components, and

. construction documentation for erected' materials-

.is not-uniquely' maintained. Construction docu-l mentation for erected. materials is generically

[ '-

maintained. MTR are required for the bulk of  ;

piping materials.

i '

MEB 3-1,'Section B.2.e SAR Section 3.6.1;1.2 j Through-wall cracks may be postulated-instead of Duke criteria.is generally equivalent'to NRC cri-

breaks in those fluid systems that qualify as teria as clarified below

l high energy fluid systems for short operational j periods. This operational period is defined as The operational period that classifies such sys-

about 2 percent of the time that the system oper- tems as moderate energy 'is either

} ates as a moderate energy fluid system. >

j- a. One percent of the normal operating life- .

j span of the plant, or ,

a J

j

b. Two percent of the time period required to accomplish the system design function.

i i,

Rev. _

1 I

4

W Tcble 3.6.1-3 (Pige"Sa)

.o Regulatory Guide 1.46: :SAR Section'3.6.2.1.2.~1; d' Longitudinal breaks are postulated in. piping.~ runs:

if

4: inches: nominal pipe size and larger. Circum- -Duke criteria is'the same as NRC Branch Techni-cal Position'APCSB'3-1 and roughly equivalent ferential breaks.are postulated in piping runs to Regulatory Guide 1.46 with . expansion of def-

' exceeding 1 inch nominal pipe' size. inition as described below:

As a minimum,ithere should be two. intermediate Longitudinal breaks are postulated in piping break locations for each pipingErun or. branch runs.4 inches' nominal pipe size and larger ex-ru a, cept that longitudinal breaks are'not postulated at. terminal ends where the piping has no longi-tudinal welds.

Duke criterie specifies that if the threshold.

stress levels are not exceeded, then no inter-mediate breaks are postulated.

_ Regulatory Guide 1.46 SAR Section 3.6.2.1.2 A whipping pipe should be considered capable of Duke criteria is the'same as NRC Branch Techni-rupturing an impacted pipe of smaller nominal cal Position APCSB 3-1 and roughly equivalent pipe size and lighter wall thickness. to Regulatory Guide 1.46 with expansion of-def-inition as described below:

The energy associated with a whipping pipe is-considered ~ capable of (a) rupturing impacted:

pipes of smaller nominal pipe sizes, and (b) developing through-wall cracks in larger nominal pipe sizes with thinner wall thicknesses.

Rev. .

u $

'.Tsblo 3.6.1-3'(Page 6)

Regulatory Guide l' 46.

SAR Section'3.6.1.1.2' '0

]

1.

Messures for restraint against pipe whipping Duke criteria is generally. equivalent to as' a result of the design basis breaks postulated... Regulatory Guide 1.46 with expansion of

.need not be provided for piping where ... the fol- definition as described below:

' lowing applies:

High energy piping is reviewed for pipe Both of the following piping system conditions -whipping and is defined as those systems

, are. met: that during normal plant conditions are either in operation of maintained pres-4 (1) the design temperature is 200*F or less, and surized under conditions where'either i

or both of the following are met:

(2) the design pressure is 275 psig.or less.

a. ' maximum temperature exceeds ~200'F, or

b. maximum pressure exceeds 275 psig,.except that (1) non-liquid piping system with a maximum pressure less than or equal to 275'

.psig are not considered high energy regard-

' less of the temperature, and (2) for liquid systems other than water, the atmospheric boiling temperature can be applied.

Systems are classified as moderate energy if the total time that either of the above conditions

! are met is less than either:

i

a. one (1) percent of the normal operating lifespan of the plant, or

b. two (2) percent of the time period required to accomplish its system design function i

i 1

1

, j

-=

Tablo 3.6.1-3 (Pcge 7)-

~ .

Regulatory Guide 1.46 ESAR Section 3.6.1.1.5 The criteria used to determine the design basis- .

In specified cases the " leak before break" concept F piping break locations ~ involves the assumption of is employed, as expla'ined in Section 3.6.1.1.5.

double-ended pipe. break. locations selected _according ; .[.

,.y.' , to stress and/or fatigue analysis, or consequence ,

analysis. ., +

11' s

'l ft P

j 2 x Rev. _

New Page

~

af, .

Table 3.6.1-4 Piping Systems For Application of

" Leak Before Break" Concept

- System Math Model Line Description Reactor Coolant- --

Reactor Coolant Loops Reactor Coolant' NC-201 Pressurizer Surge Line r-G s

5 9

I Rev. __

New Page l-

~

.n . .

%% ,w, y v 4WS A y Q(~ - a 2

p %;w

- c

> cNS  :

.;r:"r 3~

M,y 2,a. .

r *-

,f_ <-

Oh{W; jf 1. ~

J A typical' piping analysis problem, with representative math model, is shown in V S Figure 3.9.3-1. The rosults'of this analysis.are given in Table,3.9.3-10.

. ~.

e. 3.9.3.1.4 Design Leading Combinations and Design Stress Limits for M Mechanical Equipeont' Furnished by Duke

<~

f s y, +, c. w 4 T'LThe21oadcombinationsandcorrespondingstresscriteriaforDukeMechanical 7 equipment and:' valves are presented in Tabic 3.9.3-9.

.. Y., _ , .

3. 9.13. L 5 > Piping Supports and Restraints

,y y ,

% i The 9esign loading combinationLassociated with each cokpon1nt operating con-

dition .is given in. Table 3.9.3-ll~ fcF supportr., restraints, and anchors and N hin Table 3.9.3-12'for mechanical orhydraulic snubbers.

f i Loads fo'r each loading combination are combined algebraically except that com-

^

c s ponents which contain positive and, negative values are combined to assemble

+ the worst' case load combination.

Design' stress limits for each co o ent.operiting condition are in accordance with Subsection NF-~of the ASME Boiler and Pressure Vessel Code for those por-

' tions of sup'p orts.and restraints within the NF jurisdictional boundary. Stress

~

11mits for Normal'and Upset Conditions are in accordance with Article XVII-2000.

For Condition, design stress limits for manufacturer's. standard support 4 JfhcomraFaulted: onents are in' accordance with the' requirements of Appendix F. Emergency Con-

' dition' stress limits, as specified in Article 1,VII-2000.are used for the design offall '6ther components for Faulted Condition. ' Stresses for those portions of 1

. , supports and restraints outside the NF jurisdictional boundary are limited to r W p W the' allowable values in Table 3.9.3-1L

-Snubbers are used at locations where-restraints a_re necessa'ry. based on piping stress analysis, but thermal movement of the pipe must not be constrained.

... Performance selection is-based on1 manufacturer's load capacity data and the a requirement that the allowable travel of.the snubber exceed the calculated

-pipe thermal. travel. ; The midpoint"of pipe therm'a1. travel is set at the mid-

[g ' ipoint of the snubber travel range witn hot and cold settings: established ac-

g Mcordinglyi. If snubbers are used tafmitigate effects of operational vibration,

• '/Ar v 7the analytical and' design methadology utilized as well as design specifica-

. ME8 . .

• tion' requirements to assure?that structural and mechanical performance char-Q89D f acteristics.and product' quality are achieved will be developed and available

,g,,- for9. revie.w. g

- N E.

S/ 'Each snubber assemblyjis accessible after i~n sta11ation and all adjustment features are unobstructed and visable where possible. The manufacturer's

(([, gfigure number, sizeAstroke', and load rating is mounted on each snubber.

um a 1 - ', .

((M The loading combinations for Westinghouse items are given on Table 3.9.1-2.

Vy"a x .m . . .

3. 9'. 3. 2 . Pump and Valve *Uperability Progra_m.. . -

pf y

~

?3.9.3.-2.1 . Westinghouse Pump and Valve Operability Program 37% f

• p d b y Mechanical. equipment classified as safety related gsfunctionunderpostulatedp,lantconditions. must be with Equipment capable of condition faulted performing

.m

> YD o 3.9-43 Rev.

-A .

4 y ' .e k y f &f *Q_

l? "

p

+

. Table 3.9.1-6 (Page 24)

Computer Programs Used in Analysis l Application: SIMPHIP A.. Author: - EDS Nuclear, Inc.

220 Montgomery Street San Francisco, California 94104 Source: 'EDS Nuclear, Inc.

Version: August- 15, 1979

. Facility: EDS Nuclear, Inc. and UCC Dallas S

8.

Description:

SIMPWIP is a computer program for the simplified time-l history analysis of pipe whip problems. The program provides a con-servative estimate of the restraint forces and deformations, during

'the initial _ impact phase up to the time the piping system initially l: . stops after impact with the restraints. SIMPWIP can handle both Ecircumferential and longitudinal pipe rupture events. The program-assumes the. motions of-the system to be two-dimensional and in the small-deflection' range.

l -Extent and Limitations of its application: . All routines of the SIMPWIP program are used as detailed in the description.

'C. Serification: The program has been verified by comparison with the

.results of the hand calculated problem.

1 9

Rev. __

~ .-