ML20205G710

Technical Bases for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for Prairie Island Unit 2

ML20205G710
Person / Time
Site:	Prairie Island
Issue date:	09/30/1985
From:	Lee Y, Swamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
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ML20205G675	List: ML20205G668 ML20205G710 ML20205G724
References
WCAP-10928, NUDOCS 8511130321
Download: ML20205G710 (58)

Text

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8511130321 851021 PDR ADOCK 05000282 P, PDR A-2

Dirsctor of NRR Octe;ber 21, 1985 WCAP 10928 Attachment 2 1

TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR PRAIRIE ISLAND UNIT 2 1

September,1985 F. J. Witt S. A. Swamy R. A. Holmes Y. S. Lee H. F. Clark, Jr.

APPRGVED: \. N C h n APPROVED: *h J.\ N. Chirigos, Manager E. R. Johnson, Manager Structural Materials Structural and Seismic Engineering Development APPROVED: w, AvC.I,fjAs,;

J. J. McInerney, Manag(t Mechanical Equipment and Systems Licensing WESTINGHOUSE ELECTRIC CORPORATION Nuclear Energy Systems

• P. O. Box 355

, Pittsburgh, Fennsylvania 15230

. o.

FOREWORD l This document contains Westinghouse Electric Corporation proprietary

- information and data which has been identified by brackets. Coding associated with the brackets sets forth the basis on which the information is considered proprietary. These codes are listed with their meanings in WCAP-7211.

The proprietary information and data contained in this report were obtained at considerable Westinghouse expense and its release could seriously affect our competitive position. This information is to be withheld from public disclosure in accordance with the Rules of Practice 10 CFR 2.790 and the information presented herein be safeguarded in accordance with 10 CFR 2.903.

Withholding of this information does not adversely affect the public interest.

This information has been provided for your internal use only and should not be released to persons or organizttions outside the Directorate of ReguIation-and the ACRS without the express written approval of Westinghouse Electric Corporation. Should it become necessary to release this infnrmation to such persons as part of the review procedure, please contact Westinghouse Electric Corporation, which will make the necessary arrangements required to protect the Corporation's proprietary interests.

The proprietary information is deleted in the unclassified version of this report (WCAP-10928).

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TA8LE OF CONTENTS Section Title Page

1.0 INTRODUCTION

1 -1 1.1 Purpose 1-1 1.2 Scope 1 -1 l 1.3 Objectives 1-1  !

1.4 Background Infonmation 1-2 l 2.0 OPERATION AND STABILITY OF THE PRIMARY SYSTEM 2-1 2.1 Stress Corrosion Cracking 2-1 2.2 Water Hammer. 2-2 2.3' Low Cycle and High Cycle Fatigue 2-3 3.0 PIPE GEOMETRY AND LOADING . 3-1 4.0 FRACTURE MECHANICS. EVALUATION 4-1 4.1 Global Failure Mechanism 4-1 4.2 Local Failure Mechanism 4-2 4.3 Material Properties 4-3 4.4 Results of Crack Stability Evaluation 4-5 5.0 LEAK RATE PREDICTIONS 5-1 5.1 Introduction 5-1 5.2 General Considerations 5-1 5.3 -Calculation Method 5-1 5.4 Leak Rate Calculations 5-2 6.0 FATIGUE CRACK GROWTH ANALYSIS 6-1 7.0 ASSESSMENT OF MARGINS 7-1 8.0 . CONCLUSIONS 8-1

9.0 REFERENCES

9-1 APPENDIX A - Limit Moment A-1 APPENDIX 8 - Altermate Toughness Criteria for the 8-1 Prairie Island Unit 2 Primary Loop Components 8.1 Introduction B-1 8.2 Chemistry and KCU Toughness 8-1 8.3 The,As-Built Prairie Island Unit 2 Loops 8-1 8.4 Alternate Toughness Criteria for the Prairie B-2 Island Unit 2 Primary Loops on a Component-by-Component Basis v

LIST OF TABLES Tgtdt Title Page 3-1 Prairie Island Primary Loop Data 3-3 6-1 Fatigue Crack Growth at [ ]"'C 6-3 8-1 Chemical and Physical Properties of Prairie Island Unit 2 Primary Loop Material - SA 351/CF8M 8-4 8-2 Fracture Toughness Criteria for the Primary Piping Components of the Prairie Island Unit 2 Nuclear Plant 8-5 b

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w h

. LIST OF FIGURES

. Eiggtg. Title h 3-1 . Reactor Coolant Pipe 3-4 3-2 Schematic Diagram of Primary Loop Showing Weld Locations 3-5

- Prairie Island Unit 2 4-1 [ ]a,c.e Stress Distribution 4-8 4-2 J vs &a for SA351-CFSM Material at 600*F Taken 4-9 From a 23 Inch Diameter Pipe 4-3 J v's &a'for SA351 CF8M Cast Stainless Steel at 600*F 4-9 4-4 J-&a Curves at Dif ferent Temperatures, Aged Material 4-11

[ ]a.c.e (7500 Hours at 400*C) 4-5 " Critical" Flaw Size Prediction - Hot Leg at Load 4-12 Critical Location

" Critical" Flaw Size Prediction - Cold Leg at

~

4-6 4-13 Toughness Critical Location 5-1 Analytical Predictions of Critical Flow Rates of 5-4 Steam-Water Mixtures 5-2 [ la,c.e Pressure Ratio as a 5-5 Function of L/D 5-3 Idealized Pressure Drop Profile Through a 5-6 Postulated Crack 6-1 Typical Cross-Section of [ la,c.e 6-4 6-2 Reference Fatigue Crack Growth Curves for 6-5

[ Ja c.e 6-3 Reference Fatigue Crack Growth Law for [ 6-6

]a c.e in a Water Environment at 600*F -

A-1 Pipe with a Through-Wall Crack in 8ending A-2 8-1 Typical Layout of the Primary Loops for a Westinghouse B-6 Two-Loop Plant 8-2 Identification of Heats with Location for Cold Leg B-7 8-3 Identification of Heats with Location for Hot Leg B-8 8-4 Identification of Heats with Location for Crossover Leg. B-9 ix

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• 1 i

1.0 INTRODUCTION

1.1 PurDost l

This report applies to the Prairie Island Unit 2 Reactor Coolant System (RCS) primary loop piping. It is intended to demonstrate that for the specific parameters of the Prairie Island plant, RCS primary loop pipe breaks need not be considered in the structural design basis. The approach taken has been

. accepted by the Nuclear Regulatory Commission'(NRC) (Reference 1).

1.2 1E221 The structural design basis for the RCS primary loop requires that pipe breaks  ;

be postulated. In addition, protective measures for the dynamic effects associated with RCS primary loop pipe breaks have been incorporated in the Prairie Island Unit 2 plant design. However, Westinghouse has demonstrated on a generic basis that RCS primary loop pipe breaks are highly unlikely and should not be included in the structural design basis of Westinghouse plants (see Reference 2). In order to demonstrate this applicability of the generic evaluations to the Prairie Island plant, Westinghouse has performed a fracture

, mechanics evaluation, a determination of leak rates from a through-wall crack, a fatigue crack growth evaluation, and an assessment cf margins.

1.3 Obiectives In order to validate the elimination c/ RCS primary loop pipe breaks for the Prairie Island plant, the following objectives must be achieved:

a. Demonstrate that margin exists between the " critical" crack size and a postulated crack which yields a detectable leak rate.

b. Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the Prairie Island plant.

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c. Demonstrate that fatigue crack growth is negligible.

1.4 Backaround Inforination _

Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated from the structural design basis of all Westinghouse plants. The concept of eliminating pipe breaks in the RCS primary loop was first presented to the NRC in 1978 in WCAP-9283 (Reference 3). That Topical Report employed a deterministic fracture mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. That approach was then used as a means of addressing ' Generic Issue A-2 and Asynsnetric LOCA Loads.

Westinghouse performed additional testing and analysis to justify the elimination of RCS primary loop pipe breaks. This material was provided to the NRC along with Letter Report NS-EPR-2519 (Reference 4).

The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to address this same issue using a probabilistic approach. As part of the LLNL research effort, Westinghouse performed extensive evaluations of specific plant loads, material properties, transients, and system geometries to demonstrate that the analysis and testing previously performed by Westinghouse and the research performed by LLNL applied to all Westinghouse plants including Prairie Island (References 5 and 6). The results from the LLNL study were released at a March 28, 1983 ACRS Subcommittee meeting. These studies which are applicable to all Westinghouse plants east of the Rocky Mountains determined the mean probability of a direct LOCA (RCS primary loop

-10 per reactor year and the mean probability of an pipe break) to be 10

~

indirect LOCA to be 10 per reactor year. Thus, the results previously obtained by Westinghouse (Reference 3) were confirmed by an independent NRC research study.

Based on the studies by Westinghouse, LLNL, the ACRS, and the AIF, the NRC completed a safety review of the Westinghouse reports submitted to address asymmetric blowdown loads that result from'a number of discrete break locations on the PWR primary systems. The NRC Staff evaluation (Reference 1) l l

5 1-2 I

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c-k a concludes that an acceptabl? technical basis has been prc'Jided so that asymmetric blowdown loads need not be considered for those plants that can demonstrate the applicability of the modeling and conclusions contained in the Westinghouse response or can provide an equivalent fracture mechanics

. demonstration of the primary coolant loop integrity.

This report provides a fracture mechanics demonstration of primary loop integrity for the Prairie Island plant consistent with the NRC position for not considering asymmetric blowdown.

1-3

l l

2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM The Westinghouse reactor coolant system primary loop has an operating history which demonstrates the inherent stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking), water hammer, or fatigue (low and high cycle).- This operating history totals over 400 reactor-years, including five plants each having 15 years of operation and 15 other plants each with over 10 years of operation.

2.1 Stress Corrosion Crackina For the Westinghouse plants, there is no histury of cracking failure in the reactor coolant system loop piping. For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously:

high tensile stresses, a susceptible material, and a corrosive environment (Reference 7). Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized by proper material selection 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 materials in the system, applicable ASME Code rules, fracture toughness, welding, fabrication, and processing.

l The environments known to increase the susceptibilty of austenitic stainless l steel to stress corrosion are (Reference 7): oxygen, fluorides, chlorides, l hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, l- sulfites, and thionates). Strict pipe cleaning standards prior to operation j and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment. Prior to being put into l service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications. External cleaning for Class I stainless steel piping i includes patch tests to monitor and control chloride and fluoride levels. For 2-1 L

preoperational flushes, influent water chemistry is controlled. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.

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 is expected to be less than 0.005 ppm by controlling charging flow chemistry and maintaining hydrogen in the reactor coolant at specified concentrations. Halpgen concentrations are also stringently controlled by maintaining concentrations of chlorides and

! fluorides within the specified limits. This is assured by controlling charging flow chemistry and specifying proper wetted surface materials.

2.2 - Water Hammer .

Overall, there is a low potential for water hanner in the RCS since it is designed and operated to preclude the voiding condition in normally filled lines. The reactor coolant system, including piping and primary components, is designed for normal, upset, emergency, and faulted condition transients.

The design requirements are conservative 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 nar}ow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, nt..nely 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. Preoperational testing and 2-2

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operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer can occur.

2.3 Low cvele and High Cvele Fatique Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the rules of Section III of the ASME Code. A further evaluation of the low cycle fatigue loadings was carried out as part of this study in the form of a fatigue crack growth analysis, as discussed in Section 6.

High cycle fatigue loads in the system would result primarily from pump vibrations. These are minimized by restrictions placed on shaft vibrations curing hot functional testing and operation. During operation, an alarm signals the exceedance of the vibration limits. Field measurements have been made on a number of plants during hot functional testing, including plants similar to Prairie Island. Stresses in the elbow below the reactor coolant pump have been found to be very small, between 2 and 3 ksi at the highest.

These stresses are well below the fatigue endurance limit for the material and would also result in an applied stress intensity factor below the threshold  ;

for fatigue crack growth.

2-3 1

s .

3.0 PIPE GEOMETRY ANO LOAOING A segment of the primary coolant hot leg pipe shown below to be limiting in terms of stresses is sketched in Figure 3-1. This segment is postulated to contain a circumferential through-wall flaw. The inside diameter and wall thickness of the pipe are 29.2 and 2.69 inches, respectively. The pipe is subjected to a normal operating pressure of 2235 psi. Figure 3-2 identifies the loop weld locations. The material properties and the loads at these locations resulting from deadweight, thermal expansion, and Safe Shutdown Earthquake are indicated in Table 3-1. As seen from this table, the junction of the hot leg and the reactor vessel outlet nozzle is the worst location for crack stability analysis based on the highest stress due to combined pressure, dead weight, thermal expansion, and SSE (Safe Shutdown Earthquake) loading.

At this location, the axial load (F,) and the bending moment (Mb) are 1623 kips (including axial force due to pressure) and 28,422 in-kips, respectively. This location will be referred to as the load critical location. However, as seen later, the lowest end-of-service life fracture toughness occurs in the pipe defined by weld locations 10A and 11 of Loop A.

Since the loads at weld location 10A are higher than at location 11, location 10A will be referred as the touchness critical location. The associated heat of material will be called the touchness critical material.

The loads of Table 3-1 are calculated as followst The axial force F and transverse bending moments, My and Mg , are chosen for each static load (pressure, deadweight, and thermal) based on elastic-static analyses for each of these load cases. These pipe load components are combined algebraically to define the equivalent pipe static loads F,, Mg , and Mg ,. Based on elastic SSE response spectra analyses, The maximum amplified pipe seismic loads, Fd ' "yd' "Zd, are obtained.

pipe loads are obtained by combining the static and dynamic load components a*.

follows:

F, = + F lF 3 d M, . A,2.M,2 where:

i M y" lMysl +

lMyd N

Z" lNzsl  !"zdl 3-1

. e l

l The normal operating loads (i.e., algebraic sum of pressure, deadweight, and 100 percent power thermal expansion loading) at the load critical location and the toughness critical location, i.e., the junction of the hot leg and the reactor vessel outlet nozzle and a mid-location weld of the cold leg pipe, are as follows:

l Load Critical Location F = 1453 kips (including internal pressure)

M = 23,433 in-kips Touchness Critical location F = 1426 kips (including internal pressure)

M = 4228 in-kips The calculated and allowable stresses for A5ME III M8-3600 equation 9 (faulted i.e., pressure, deadweight, and SSE) and. equation 12 (normal operating thermal stress) at the load critical location are as follows:

Calculated Allowable Ratio of Equation Stress Stress calculated /

Number (ksi) (ksi) Allowable 9F 10.4 50.1 0.21 12 12.4 50.1 0.25 At the toughness critical location the calculated stresses and ratios are even less.

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3-2

TA8tE 3-1 P8A181E ISLAN8 P81NA8V LOOP 8ATA l

Faulted Loadsa Inside Wall Yield Ultimate Sending 81 rect Stress Weld 8adius Thickness Stress Stress Flow Stress Axial Load Moment. (ksi)-

Locations (in) (in) e e 1 a (Kips) (in-Kips) y u [s g- 2I 'y**uIl ,c.e F M F K s= E I (ksi) (ksi) ggggg x 'D x A +Z lb 14.6 2.69 18.8 67.0 42.9 1623 28422 2 14.6 2.69 20.3 18.8 67.0 42.9 1621 9554 10.8 1 3 15.6 2.87 18.8 67.0 42.9 1476 4 15.6 2.87 16843 11.1-19.5 67.0 43.3 1682 5717 1. 0 -

5 15.6 2.87 19.5 67.0 43.3 1688

) 6 15.6 2.87 4317 7.3 19.5 67.0 43.3 1677 25274 15.9

'i 7 15.6 2.87 19.5 67.0 43.3 1761 6598 8 15.6 8.4 2.87 19.5 67.0 43.3 1762 18332 13.3 i 9 15.6 2.87 19.5 67.0 43.3 1830 17051 13.0 l

10 13.85 2.55 19.5 67.0 43.3 1529 9432 11.9 l oa 10AC 13.85 2.55 19.5 67.0 43.4 1529 8487 i d, 13.85 11.3 11 2.55 19.5 67.0 43.3 1530 4997 9.3 j 12 '13.85 2.55 19.5 67.0 43.3 1510 5587 9.5 I

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Tou9 h ness critical location i

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$ A 07 HOT LEG Temperature: 599*F; Pressure: 2235 psi CROSSOVER LEG Temperature: 536*F; Pressure: 2190 psi COLD LEG Temperature: 536*F; Pressure: 2290 psi Figure 3-2 Schematic Diagram of Primary Loop Showing Weld Locations -

Prairie Island Unit 2 3-5

4.0 FRACTURE MECHANICS EVALUATION 4.1 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 deforination accompanying fracture. One method for predicting the failure of ductile material is the plastic instability method, based on traditional plastic limit load concepts, but accounting for strain hardening 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 termed as the flow stress. The flow stress is generally taken as the average of the yield and ultimate tensile strength of the material at the temperature of interest.

This methodology has been shown to be applicable to ductile piping through a large number of experiments and will be used here to predict the critical flaw size in the primary coolant piping. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 4-1) 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:

g

. a,c.e

)

where:

[

j a,c.e I

4 -1

[

j a.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 8).

4.2 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. Depending on the material properties and geometry of the pipe, flaw size, shape and loading, the local failure mechanisms may or may not govern the ultimate failure.

The stability will be assumed if the crack does not initiate at all. .It has beenacceptedthattheinitiationtoughnessmeasuredintermsofJ gg 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 the J gg of the material, then the crack will not initiate. 'If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:

T U 5-app da 2 .

,f 4-2

,-_.,n--, - - - m--,m, --n---,

e , , , - ,,-- ,- r ----,,--,--__,---------n- , _ _ _ . ,--.e--

. s l l

where:

, T,,, = applied tearing modulus E = modulus of elasticity o g=( ]a,c.e (flow stress) a = crack length g j a.c.e In summary, the local crack stability will be established by the two-step criteria:

J<J, g or T,,, <T 1 Ic mat In this analysis, a hypothesized circumferential through-wall 7.5-inch long flaw is taken as a reference f_13 and is used as a basis for evaluation.

4.3 Material Procerties The primary loop piping material for Prairie Island Unit 2 is ASTM SA351-CF8M, a cast product form. The material for the primary loop fittings is also SA351-CFOM. Welds of interest are indicated in Figure 3-2.

The tensile and flow properties of the load critical location, the hot leg and the reactor vessel nozzle junction, and the toughness critical location, the mid-section weld of the Loop A cold leg pipe, are given in Table 3-1.

The pre-service fracture toughness of cast materials in terms of J have been found to be very high at 600*F. Typical results are given in Figures 4-2 and 4-3 taken from References g and 10. J gg is observed to be over 5000 2

in-lbs/in . However, cast stainless steels are subject to thermal aging during service. This thermal aging causes an elevation in the yield strength

, of the material and a degradation of the fracture toughness, the degree of degradation being proportional to the level of ferrite in the material. To 4-3

i l

t determine the offacts of thermal aging on piping integrity, a detailed study l was carried out in Reference 10. In that report, fracture toughness results were presented for a material representative of ( i

. ]' ' ' d Toughness results ve e provided for the material ,

in the full service life condition and these properties are also presented in Figure 4-4 of this report for information. The J gg value for this material at operating temperature was approximately [ ,] # and the maximum value of J obtained in the tests was in excess of (

~

. ]' ' ' d The tests of this material were conducted on small si;ccinens and therefore rather short crack extensions, (maximum extension 4.3 um) so it is expected that higher J values would be sustained for larger specimens. T mat was [ ] at operating temperature. The effects of the aging process on the end-of-service life fracture toughness is discussed in Appendix 8. The end-of-service life toughness is seen to exceed that of ( ] with two exceptions. One heat, falling slightly below 4

that of ( ] d , is in Loop A at the load critical location (i.e., the junction of the hot leg and the reactor vessel nozzle). In Appendix 8 the end-of-service life toughnesses for this heat are established using the alternate toughness criteria methodology. Since the load critical location is coincident with a location exhibiting toughnesses less than that of

( ]C, the toughnesses associated with this Loop A hot leg heat are conservatively taken as representative of the toughnesses of all the remaining heat with the exception of the toughness critical material. The end-of-service life toughnesses for the toughness critical material are also determined in Appendix 8.

The fracture toughness criteria to be used in the fracture mechanics evaluation, based on the alternate toughness methodology of Appendix 8, are:

(

O 6

{ j a.c.e 4-4 i

-. _ . . . . . - - - _ . _ _ - - - . . . - - - . _. _-.--_,c... . - . - _ , - - - . - - - - _ _ _ _ _ _ . - _ ~ . . -

Available data on aged stainless steel welds (Reference 10 and 11) indicate the J g, values for'the worst case welds are of the same order as the aged material, but the slope of the J-R curve is steeper, and higher J-values have

' 2 been obtained from fracture tests (in excess of 3000 in-lb/in ). The applied value of the J-integral for a flaw in the weld regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at temperature. Therefore, weld regions are less limiting than the cast material.

In the fracture mechanics analyses that follow, the fracture toughness l properties tabulated above will be used as the criteria against which the applied fracture toughness values will be compared.

l 4.4 Results of Crack Stability Evaluation l

Figure 4-5 shows a plot of the plastic limit moment as a function of through-wall circumferential flaw length in the hot leg of the main coolant l piping. This limit moment was calculated for Prairie Island Unit 2 from data for a pressurized pipe at 2235 psi with an axial force of 1623 kips, operating at 599'F with ASME Code minimum tensile properties. The maximum applied bending moment of 28,422 in-kips can be plotted on this figure, and used to detemine a critical flaw length, which is shown to be [ ]"'C inches.

In Figure 4-6, a plot of the plastic limit moment as a function of through-wall circumferential flaw length in the toughness cr.itical material portion of the cold , leg of the main coolant pipe is given. This limit moment was calculated for a pressurized pipe at 2290 psi with an axial force of 1529 kips, operating at 536*F with ASME Code minimum properties. The maximum applied bending moment of 8487 in-kips is plotted on the figure and determines a

a citical flaw length of ( l .c.e inches.

In Table 3-1, the outer surface axial stress (a,) at the load critical location is seen to be 20.3 ksi. Stresses due to the internal pressure of 2235 psi are as fcilms (see Reference 12):

4-5

s e (circumferential stress): 11.1 ksi e radial stress: 0 r

The von Mises effective stress, e,ff, (see Reference 13) is given by e.gg = (2 I'a ~ 'r)

• I'c ~ 'rI

• I'a ~ 'c) and is [ }"'**'.

Thus the offactive stress is less than the yield stress and by the Von Mises plasticity theory yielding does not occur. Also, similar consideration at the toughness critical location confirms that yielding does not occur there.

Hence, linear elastic fracture mechanics is' applicable for analyzing the pipes with hypothesized flaws. The analytical method used for the local stability evaluation is sununarized below.

The stress intensity factors corresponding to tension and bending are expressed, respectively, by (see Reference 14)

K *#

t t /IT F{a)

K b " 'b /Ta Fga) where F t (a) and bF (a) are stress intensity calibration factors corresponding to tension and bending, respectively, a is the half-crack length, a is the half-crack angle, e is the remote uniform tensile stress, and ab is the remote fiber stress due to pure bending. Data for Ft (s) and F (a) are given in Reference 14. The offact of the yielding b

near the crack tip can be incorporated by Irwin's plastic zone correction method (see Reference 15) in which the half-cro.k length, a, in these formulas is replaced by the offactive crack length, ag ,, defined by a,gg = a + b "

2v s y 2 4-6

s

~ ^'

for plane stress plastic corrections, where ye is the yield strength of

!the material and K is the total stress intensity due to combined tensile and bending loads. Finally, the J g-value is determin'ed by the relation Jg=

K/E,ishereE=

2 Young'sModulus(25.3x106 ,,g), -

J was calculated for a through-wall flaw'7.5 inches long (the reference flaw) 2 in the hot-leg pipe and found to be [ ]C in-1bs/in . The flaw was increased by ( ]C d .

Both of these values are less than J gg for all the cast material described

- above with the exception of the toughness critical material. Thus, based on '

this analysis, crack initiation is not expected to occur in the Prairie Island piping even [or circumferential through-wall flaws up to ( ]' d

'long under the loading conditions given with the exception noted.

n ,

i '

l For the toughness critical mat'erial J was calculated for a through-wall flaw

,7.5 inches long using the appropriate stresses defined by the' loads at 2

location 10A of Table 3-1. J was found to be [ ]"' C d in-lb/in . The flaw size was increased by (

~

]a c.e in-lb/in twhich is still less~than J Ic for the material. For a flaw

[ l.c.e a

in-lb/in .2 Thus, crack initiation is not expected to occur in the toughness critical material pipin.g.

In shMnery, crack initiation is not expected to occur in the Prairie Island '

Unit 2 piping even for circumferential flaws up to ( ]"'C long under the loading conditions given with even greater flaw sizes existing for the toughness critical material piping.

n a

Ng k

/

4-7 k_ _ - _ - _ _- - - - -

~

1

a. .

~

~

&,c,e lllI/////1,,

la

~ wruz.,,

tt /

V a,c,e FIGURE 4 1 ( 3 STRESS DISTRIBUTION 4-8

e .

I

' - a,c.e  ;

i i

i i

t

! l i

i t

i i

2 i

l i

~

Figure 4-2 J vs. aa for SA351-CF8M naterial at 600'F taken from a 23 inch diameter pipe 4-9 4.. _

e _..-

ra

~ '

a,c.e 4

+

.T'

.: Figure 4-3 J vs aa for SA351-CF8M Cast Stainless Steel at 600*F 4-10

I

• b q , ,- - - -

= . ., .

A .- .

M>

e w

3'

, u.e 4

r 1 -

l' ,

t g ..

1 d

i t 1.. --- ,

k

- FIGURE.4-4 l J-Aa Curves at Different Temperatures-.for Aged Material [--- ]a,c.e'

/' (7500 Hours at 400*C) -

i . -, .

l S; 4

4-11 -

1 m% v ---eb-myy y %- 9-m-,, e e-ym.. _ ,,r , ,,-

1

• a v"- r ~

N' r - w + -e - - -

t- w9-ye-gg r,

- . - a,c.e I

L l u I FLAW GEOMETRY  :

00 = 34.58 in. i t = 2.69 in.

p = 2235 psi F = 1623 kips -

e y

= 18.8 ksi ,

a u

= 67.0 ksi .

a f

= 42.9 ksi Temp = 599*F Figure 4-5 " Critical" Flaw Size Prediction - Hot leg at Load Critical Location 4-12

,a,c,e

~

t i v i FLAW GEOMETRY OD = 32.80 in.

t = 2.55 in.

P = 2290 osi F = 1529 kips (inc. P) e y

= 19.5 ksi e

u

= 67.0 ksi ep = 43.25 ksi Temp = 536*F i

\

l l

?

l Figure 4-6. " Critical" Flaw Size Prediction - Cold Leg at Toughr.ess l' Critical Location 4-13 l

l

- - ~ . -. -

e s -

.- 5.0 LEAK RATE PREDICTIONS

, '5.1 Introduction Fracture mechanics analysis has shown that postulated through-wall cracks 'in the primary loop would remain stable and not cause a gross failure of this component.' If such a through-wall crack did exist, it would be desirable to detect the leakage 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 a postulated crack and prasent the leak rate calculation results for through-wall circumferential cracks.

5.2 General Considerations i

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 ratio of the channel length, L, to hydraulic diameter, D , ( D) 3 '

H H a

greater than [' J c.e, bcth [ Ja , ,e must be

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

the fluid. At this point, the flow begins to flash and choking occurs.

Pressure' losses due to momentum changes will dominate for [ .]a,c.e However, for large L/D H values, riction pressure drop will become important and must be considered 41ong with the momentum losses due to flashing.

5.3 Calculation Method '

i The basic method used inLthe leak rate calculations is the method developed by

[

l-i

,a.C,e J

• The flow rate through a crack was calculated in the following manner. Figure r

5-1 f rom Reference 16 was used to estimate the critical pressure, Pc, for the

primary loop enthalpy condition and an assumed flow. Once Pc was found for a I given mass aflow, the [ a l .c.e 5-1 t

I' - - . _ . . . . _ , . . _ . _ . . _ , . . _ . . _ _ - _

was found from Figure 5-2 taken from Reference 16. For all cases considered, since [ .] Therefore, this method will yield the two-phase pressure drop due to momentum offacts as illustrated in Figure 5-3.

Now using the assumed flow rate 6, tho frictional pressure drop can be calculated using a p, = [ ]a,c.e ( 5-1) where the friction factor f is determined using the [ ]

The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ ]a,c.e ,,3, The frictional pressure drop using Equation 5-1 is then calculated for the assumed flow and added to the [

l a.c.e to obtain the total pressure drop from the primary system i to the atmosphere. That.is, for the primary loop Absolute Pressure - 14.7 = ( ]*** (5-2) for a given assumed flow 6. If the right-hand side of Equation 5-2 does not agree with the pressure difference between the primary loop and the atmosphere, then the procedure is repeated until Equation 5-2 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. This calculational procedure has been recommended by [

l a.c.e for this type of (

]C calculation.

5.4 Leak Rate Calculations For the load critical location leak rate estimates were performed by applying the normal operating-bending moment of 23,433 in-kips in addition to the normal operating axial force of 1453 kips. These loads were applied to the hot leg pipe containing the postulated reference (7.5 inches long) through-wall flaw and the crack opening area was estimated using the 5-2 l

4- .

method of Reference 14. The leak rate was calculated using the two-phase flow formulation described above. The computed leak rate was 80.8 gpm. In order to determine the sensitivity of leak rate to flaw size, a through-wall flaw [

Ja .c.e in length was postulated. The calculated leak rate was [

] C d flaw produced a leak rate of 3.6 gpm.

In a similar manner leak rates were also calculated for the toughness critical location (location 10A of Figure 3-2) using the normal operating bending moment of 4228 in-kips and the normal operating axial force of 1426 kips.

Crack lengths considered were [ ]C,7.5,[ Ja ,c.e inches.

The leak rate for the [ ]C and 7.5 inch flaws were 2 and 14.6 gpm, i-respectively. The [ ]*** flaws produced flow cates of 34.8 gpm and 153.5 gpe, respectively.

The Prairie Island plant has an RCS pressure boundary leak detection system which is consistent with the guidelines of Regulatory Guide 1.45 for detecting leakage of 1 gpm in one hour. Thus, for the 7.5 inch flaw at the load critical location in the hot-leg, a factor in excess of 80 exists between the calculated leak rate and the criteria of Regulatory Guide 1.45. Relative to the [ ]a,c.e. a factor of over 23 exists.

For the toughness critical location, a 7.5 inch flaw has a factor of over 14 above the criteria of Regulatory Guide 1.45. For a [ J a,c.e inch flaw the factor exceeds 150.

\

5-3

6 s ll -

l e a,c.e E

E I

c_

8 C

N 3

STAGNATION ENTHALPY (102 Stu/lb)

Figure 5-1 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 5-4

o . .

.. a,c.e o

, 4

-9 s

W s

E a

2 2

e u

LENGTH / DIAMETER RATIO (L/D) l l

i I

Figure 5-2 [ ]C'8 Pressure Ratio as a Function of L/D 5-5

. _ a,c.e a.c.e l

7

_- _ : =^

I i-l  % A Figure 5-3 Idealized Pressure Drop Profile Through a Postulated Crack 5-6

6.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the presence of small cracks, a fatigue crack growth analysis was carried out for the [

]C region of a typical system (see Location

[ ]C# of Figure 3-2) . This region was selected because crack growth calculated here will be typical of that in the entire primary loop. Crack growths calculated at other locations can be expected to show less than 10%

variation. Thermal aging has been shown not to impact fatigue crack growth (References 10 and 11).

A( ,

]C# of a plant typical in geometry and operational characteristics to any Westinghouse PWR System. [

j a,c.e All normal -upset, and test conditions were considered and circumferentially oriented surface flaws were postulated in the region, assuming the flaw was located in three dif ferent locations, as shown in Figure 6-1. Specifically, these were:

Cross Section A: - a,c.e Cross Section 8:

Cross Section C:

E Fatigue crack growth rate laws were used [

]C The law for stainless steel was derived from Reference 18, with a very conservative correction for the R ratio, which is the ratio of minimum to maximum stress during a transient.

For stainless steel, the fatigue crack growth formula is:

ff=(5.4x10-12) g ff 4.48 1nches/ cycle 6-1

where K,ff = K ,,, 0 -R) U min " max

[

j a.c.e a,c.e where:-[ ] a,c.e The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is sumarized in Table 6-;. and shows that the crack growth is very small, regardless [

a 3 .c.e 6-2

. e l

TABLE 6-1 FATIGUE CRACK GROWTH AT [. Ja.c.e (40 YEARS)

FINAL FLAW (in) a.c.e INITIAL FLAW (IN) _ _ [ Ja,c.e [ ]a,c.e 0.292 0.31097 0.30107 0.30698 0.300 0.31949 0.30953 0.31626 0.375 0.39940 0.38948 0.40763

-0.425 0.45271 0.4435 0.47421 4

6-3

c e O

- a,c.e Figure 6-1 Typical Cross-Section of [ ]a,c.e 6-4

& -E N'

-q r ~

S.C. 8 d

4

.f

-m a

o c

s a

w 2

0-a 2

3 n

q s

'4 .

a s .

!i o

+ a 0

-x o

4

.. E u

6 F

Figure 6-2 Reference' Fatigue Crack Growth Curves for [

Ja,c.e 1

6-5 J

W 4

a 4 4 m: ,--<-w. +-,- ,v -- ~ -- w a -

- ee .n.. en-~-- -,--,- ,--w,-n,-w-w,, - - - , . , , ,w, - - - --wv ,

1 .

, . ,y e

a.c.' e 9

.' d s

1.

' Figure 6-3 Reference Fatigue Crack Growth Law for[ Ja.c.e in a Water Environment at 6000F -

6-6 ..

O

l.

7.0 ASSESSMENT OF MARGINS For a through-wall circumferential flaw 7.5 inches long at the most highly loaded location, the calculated J is [ ]' # in-lb/in . This value is only about 60% of the lower bound JIc 'I E 3 I"~IU#i" established in Appendix B for this region and for all heats except one. For the lone exception (the toughness critical location), the calculated J for a 7.5 inch through-wall circumferential flaw is only [ ]a.c.e in-lb/in2 as compared to a J gg of [ 3a ,c.e in-lb/in 2which yields a margin of over 3.5. Thus, crack initiation would not be anticipated for these I situations. Increasing the hypothesized flaw by [ ]"'C inches increased the applied J values for the two cases above to [ ]a,c.e 2

in-lb/in2and [ ]C in-lb/in , respectively. These values are still less than the J s at the pertinent locations. A fTaw [ ]a,c.e Ic inches long at the toughness critical location produces an applied J of only 2

[ ]"'C # in-lb/in which is still less than the associated J Ic*

. As shown in Section 3.0, a margin of a factor of not less than 4 exists between calculated and ASME Code allowable faulted condition and thermal i

stresses. .

In Section 4,4, the " critical" flaw sizes at the load critical location and .

the toughness critical location are calculated using the limit load method to be [ ]a,c.e inches and [ ]a,c.e inches, respectively. Based on the above, the 'cri,tical" flaw sizes at these locations will, of course, exceed

[ -]**C inches, respectively. -

In Section 5.0, it is shown that at the load critical location a flaw of 7.5 inches would yield a leak rate in excess of 80 gpm while for a

[ ]a,c.e inch flaw, the leak rate is over 3 gpm. Thus, there is a margin of at least 4 between the flaw size that gives a leak rate well exceeding the

, criterion of Regulatory Guide 1.45 and the " critical" flaw size of (

ya c.e l

a At the toughness critical location, a flaw of [ l c.e inches would yield a leak rate of 2 gpm while flaws of 7.5 and [ ]a,c.e inches would produce 7-1

r o .

leakage exceeding 14 and 150 gpa, respectively. Thus, at the toughness critical location there is also a margin of at least 4 between the flaw size that gives a leak rate exceeding the criterion of Regulatory Guide 1.45 and the ' critical" flaw size of [ ]"' C d .

In summary, relative to

1. keidi

a. Prairie Island Unit 2 is enveloped by the J values established from testing of highly aged material.

b. Margins at the critical location of at least 4 on faulted conditions and thernal stresses exist relative to ASME Code allowable values.

2. Flaw Size

a. A margin of at least 4 exists between the " critical" flaw and the flaw yielding a leak rate of 1 gal / min.

b. If limit load is used as the basis for " critical" flaw size, the margin for global stability compared to the reference flaw would be near 5.

c. A margin of at least 33% exists between the reference flaw size (7.5

. inches long) and the established minimum " critical" flaw size. Flaw initiation is shown not to occur for either case, i.e., J for flaws of size of at least [ ]aiE?e

3. Leak Rate At the load critical location a margin in excess of 80 exists for the reference flaw (7.5 inches long) between calculated leak rates and the criteria of Regulatory Guide 1.45. For the reference flaw the margin is i in excess of 14 for the toughness critical location.

1 7-2 i

c

., e I

8.0 CONCLUSION

S

. This report justifies the elimination of RCS primary loop pipe breaks for the Prairie Island Unit 2 plant as follows:

a. Stress 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 should not occur in the RCS piping because of system design, testing, and operational considerations,

c. The effects of low and high cycle faligue on the integrity of the primary piping are negligible.

d. Adequate margin exists between the leak rate of the reference flaw and the criteria of Reg. Guide 1.45.

e. Ample margin exists between.the reference flaw chosen for leak detectability and the " critical" flaw.

f. Ample margin exists in the material properties used to demonstrate end-of-service life (relative to aging) stability of the reference flaw.

The reference flaw will be stable throughout reactor life because of the ample margins in d, e, and f above and will leak at a detectable rate which will assure a safe plant shutdown.

Based on the above, it is concluded that RCS primary loop pipe breaks need not be considered in the structural design basis of the Prairie Island Unit 2 plant.

~~

8-1

O .

3 x '

9.0 REFERENCES

1. USNRC Generic letter 84-04,

Subject:

" Safety Evaluation of Westinghouse

- t Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary Main Loops", February 1,1984.

2.- Letter from Westinghouse (E. P. Rahe) to NRC (R. N. Vollmer),

MS-EPR-2768, dated May 11, 1983.

- 3. WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events," March,1978.

n 4. Letter Report NS-EPR-2519. Westinghouse (E. P. Rahe) to NRC (D. G.

Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.

5. Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983.

6. Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.

7. NUREG-0691, " Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors", USNRC, September 1980.

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

9. Landes, J. D., et. al., Fracture Toughness of 316 Stainless Steel Piping Material at 600*F, Westinghouse R&D Report 79-703-PIPRE-R1, May 17, 1979. (Westinghouse Proprietary Class 2) m 9-1

4

10. SCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping For W NSSS,' W Proprietary Class 2, Noventer 1983.

11. Slama, G., Petrequin, P. , Masson, S. H., and Mager, T. R. , "Ef f act of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and Welds", presented at SMiRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel Reactor Pressure Soundary Components, August 29/30, 1983, Mor.terey, CA.

12. Durelli, A. J., et. al., Introduction to the Theoretical and Experimental Analysis of Stress andJtrain, McGraw Hill Book Company, New York, (1958), pp 233-236.

13. Johnson, W. and Mellor, P. B., Engineering Plasticity, Van Nostrand Relmhold Company, New York, (1973), pp 83-86.

14.- 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.

15. Irwin, G. R., " Plastic Zone Near a Crack and Fracture Toughness," Proc.

7th Sagamore Conference, P. IV-63 (1960).

16. [

ya .c.e

17. t j a.c.e 9-2

18. Bamf ord, W. H.. '" Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment " Trans. ASME Journal of Pressure Vessel Technology, Vol.101, Feb.1979.

19. [

j a.c.e

. 20. [

a 3 c.e

21. Witt, F. J., and Kim C. C., Toughness Criteria for : Thermally Aged Cast Stainless Steel, Westinghouse Proprietary Class 2 Report WCAP 10931, September 1985.

i 9-3

e 4 APPENDIX A LIMIT MOMENT

[

j a.c.e A-1

O

- a,c.e FIGURE A-1 PIPE WITH A TriROUGH-WALL CRACX IN SENDING A-2

s

, APPENDIX B i

' l ALTERNATE TOUGHNESS CRITERIA FOR THE PRAIRIE ISLAND UNIT 2 PRIMARY LOOP COMPONENTS

8.1 INTRODUCTION

Not all of the individual piping components of the Prairie Island Unit 2 primary loop piping satisfy the present [ Ja c.e criteria. In this appendix, the alternate toughness criteria developed in Reference 21 will be used to categorize the various individual piping cor,ponents thus establishing criteria based upon which the mechanistic pipe break evaluation may be performed. First the chemistry and calculated KCU values are given followed by an identification of each of the heats of material with a specific loop and location. The criteria for the various individual loop components are tabulated.

8.2 CHEMISTRY AND KCU TOUGHNESS The correlation of Reference 11 which is based on the chemistry of the cast stainless steel piping was used to calculate the associated KCU value. The chemistry and end-of-service life KCU toughness values are given in Table 8-1. Of the twenty one heats, only two fail to meet the current [

3"'"# criteria. Both heats are in Loop A, one in the hot leg connecting to the reactor vessel at the load critical loca' tion and the other in the mid-section of the cold leg.

l 8.3 THE AS-8UILT PRAIRIE ISLAND UNIT 2 LOOPS i

Prairie Island Unit 2 is a two-loop Westinghouse type pressurized water reactor plant. A typical 2-loop primary system is sketched in Figure B-1.

. The two loops are identified as Loop A and loop B in Prairie Island Unit 2.

Sketches for associating piping component with specific locations and loop are given in Figures B-2 through B-4. The individual components are identified by heat numbers. The heat numbers above the sketches refer to Loop B; those below, to Loop A. The components which do not meet the current criteria are identified (see Figures B-2 and 8-3).

B-1

+ .

'iv,

8. 4' ALTERNATE TOUGHNESS CRITERIA FOR THE PRAIRIE ISLANO UNIT 2 ~

, PRIMARY LOOPS ON A COMPONENT-8Y-COMPONENT BASIS d' .

The alternate toughness criteria for the Prairie Island Unit 2 primary loops may biobtained by applying the methodology of Reference 21 to Table 8,,1.

First, it is observed that all except two heats fall into [ .,, )"',

i.e.,theysatisfytiie{surrent[ ]# criteria.

Not leg heat No. [ ]' d has a calculated end-of-service life dCU at room temperature o.'( I Ja ,c.e daj/cm which falls below that of (

]" ' ' # . The 6-ferrita content is [ ]a.c.e By Reference 21, the

(

)*' ,

Thus, for full embrittlement .

J gg = [ ]W -

T mat

= [ ]C .

J =

[ 3a c,e max /_ - ,

Ja ,c.e KCU =

(-

Since the end-of-service life ECU value exceeds the full-embrittlement KCU a

value. Heat No. [ ]"'C is a [ J .c.e and the end-of-service life fracture toughness may be found by using the

[ ] # (Equations 2-20 through 2-22) of Reference

21. The results are given in Table 8-2.

s' ~

FY

~

s

, _ e s 8-2 D - g g

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

v. , , , , . _ . _ _ _ _ , _ _ . _ _ _ . , _ _ . , __ _ _ _ _ _ _ , _ . . . - .. _ _ , . _ , . . . . , , . , . . _ . . - _ _ . _ _ . _ _ _ _ . _ _ , _ _ _ . .

Cold leg Heat No. [ ]a,c.e has a calculated end-of-service life KCU value at room temperature of [ ]* ' C # daJ/cm which falls below that of

[ ]. The 4-ferrite content is [ ]" # By Reference 21, the [

3a ,c,e

- Thus for full embrittlement a

J =[ J .c.e T

mat "E3' J,,, =[. ]*****

KCU =[ ]

Since the en'd-of-service life KCU value exceeds the full-embrittlement KCU value. Heat No. [ ]I'E is a [ Ja ,c.e and the end-of-service life fracture toughness may also be found by using the

[ ]a,c.e (Equations 2-20 through 2-22) of Reference

21. The results are given in Table B-2.

4

>=

B-3

. .._. .,_ . 4 .m. .. . .

-r , -

-qw L . p -

- .. + 1

_ 3 .

~

- - se ._..

s_ 'q L -r '; ~n

,. .} 1

--?-' -,

y:: c _ g-c 1, 4

L

(

4 H

4 i

s

~

/e

~ '

}j'. N b1 Chemical'and Physical Properties of Prairie Island Unit 2- '

% Idop Meterial - 34351/CF8M 4

' a mm a

-" ' . 4,C,4 M r

] r tf-s-

,s -

/

4 .

T '-

• ,e i,

',' ',(

s 5

4-e E T U

S k, 'N l

E

+

- p.

f-f

.c) .

. '. . i

, ,>:.' g ;j

..t d k

., e

?

J r

m--

e f

it

'4.' L l

?-

,+,

4* [, 2 t.

w~ l t ,

y I.f r' -

I T

O 4

B-4 5

4-

+1 s TI T "-

P WrM 1"M TYm+ 3We sm-7 e ** D 4 5%b- & cw *t- 4 4 ev+w+ - ---e- - - -Wm*+e++e'e -.m-e,----sva-w--+w+--w=- tW en v ww -ym-g N-W** rv e

, . -a TA8LE 8-2. I FRACTURE TOUGHNESS CRITERIA FOR THE PRIMARY PIPING COMPONENTS OF THE PRAIRIE ISLAND UNIT 2 NUCLEAR PLANT a,c.e 9

b

St'eam Generator

~

s%

'**~

Crossover Leg Hot Leg i..

i,) Loop B Pump Reactor Vessel

^

\

Cold Leg

[ '

s

~

'.._./

Loop A

. .'f,( ,

.I Figure B-1 Typical Layout of the Primary Loops for a Westingnouse Two-Loop Plant l

l l

l l

B-6 l

9 0

- - a,c.e Figure B-2 Identification of Heats with Location for Cold Leg B-7

l

_ a,c.e 4

Figure B-3 Identification of Heats with Location for Hot Leg

, B-8

R9 -.,. '~ g . . .

.- i ..

s E3 _

4

- 7 l p .'^-

/

,e, l

s a,c.e

.. 1 1, .

F Figure B-4

. Identification of Heats with Location for Crossover. Leg c

Wh w

b 9

B-9 1 , .