ML20214P869

Rev 1 to Technical Bases for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for Prairie Island Unit 2

ML20214P869
Person / Time
Site:	Prairie Island
Issue date:	07/31/1986
From:	Lee Y, Swamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19292F920	List: ML20214P780 ML20214P830 ML20214P850 ML20214P869
References
WCAP-10928, WCAP-10928-R01, WCAP-10928-R1, NUDOCS 8609240083
Download: ML20214P869 (62)

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WEST!NGHOUSE CLASS 3 WCAP 10928 Revision 1 TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR PRAIRIE ISLAND UNIT 2 September 1985

. Revision 1: July 1986 F. J. Witt S. A. Swamy R. A. Holmes Y. S. Lee H. F. Clark, Jr.

A APPROVED: / APPROVED:

S. S. Palusamy, Manager E. R. Johnson, Manager Structural Materials Structural and Seismic Engineering Development APPROVED: 0,A. C A 'L/w4..

J J. Inerney, Man r Mechanical Equipment and Systems Licensing WESTINGHOUSE ELECTRIC CORPORATION.

Nuclear Energy Systems P. O. Box 355 Pittsburgh, Pennsylvania 15230 o

hh[ ADO k $$2 P

, FOREWORG The original version of this report was dated September 1985- Since that time significant revisions were made to the prime reference document WCAP-10931 3

(now WCAP-10931, Revision 1 and Reference 21 of this document) wherefrom revised end-of-service life toughness criteria were established. These new criteria required that additional analyses be performed and evaluations made, hence prompting the revision of this report. The revisions are~ identified by vertical lines in the column.

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TABLE OF CONTENTS Title Page Section '

1-1

1.0 INTRODUCTION

1-1 1.1 Purpose 1.2 Scope 1-1 1.3 Objectives 1-1 1.4 Background Information 'l-2 2-1 2.0 OPERATION AND STABILITY OF THE PRIMARY SYSTEM 2-1 2.1 Stress Corrosion Cracking 2-2 2.2 Water Hammer 2-3 2.3 Low Cycle and High Cycle Fatigue 3-1 3.0 PIPE GEOMETRY AND LOADING 4-1 4.0 FRACTURE MECHANICS EVALUATION 4-1 4.1 Global Failure Mechanism 4-2 4.2 Local Failure Mechanism 4-3 4.3 Material Properties 4.4 Results of Crack Stability Evaluation 4-5 5-1 5.0 LEAK RATE PREDICTIONS 5-1 5.1 Introduction 5-1

- 5.2 General Considerations 5-1 5.3 Calculation Method 5-2 5.4 Leak Rate Calculations 6-1 l 6.0 FATIGUE CRACK GROWTH ANALYSIS 7-1 7.0 ASSESSMENT OF MARGINS B-1 l

8.0 CONCLUSION

S 9-1

9.0 REFERENCES

A-1 APPENDIX A - Limit Moment APPENDIX B - Alternate Toughness Criteria for the B-1 Prairie Island Unit 2 Primary Loop Components B-1 B.1 Introduction 8-1 B.2 Chemistry and KCU Toughness B.3 The As-Built Prairie Island Unit 2 Loops B-1 B.4 Alternate Toughness Criteria for the Prairie B-2 Island Unit 2 Primary Loops on a Component-by-

- Component Basis 1

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LIST OF TABLES Table Title Page 3-1 Prairie Island Unit 2 Primary Loop Data Including Faulted Loading Conditions 3-3 3-2 Normal Condition (Dead Weight + Pressure + Thermal)

Loads for Prairie Island Unit 2 3-4 4-1 Fracture Toughness Criteria Used in the Leak-Before- l Break Evaluation ~4-8 6-1 Fatigue Crack Growth at [ la.c.e 6-3 (40 Years) ,

7-1 7-3 Summary of J,pp and Leak Rate Results as a Function of Crack Length at the Five Critical Locations 8-1 Chemical and Physical Properties of Prairie Island Unit 2 Primary Loop Material - SA 351/CFBM 8-3 8-2 Fracture Toughness Criteria for the Primary Piping Components of the Prairie Island Unit 2 Nuclear Plant 8-4 l

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LIST OF FIGURES Figure Title Page l

3-1 Reactor Coolant Pipe 3-5 3-2 Schematic Diagram of Primary Loop Showing Weld Locations 3-6

- Prairie Island Unit 2 -

4-1 [ la.c.e Stress Distribution 4-9 4-2 J vs Aa for SA351 CF8M Cast Stainless Steel at 600*F 4-10 4-3 4-11 J-Aa Curveg,gt la, Different Temperatures, Aged Material

[ (7500 Hours at 400*C) 4-4 Critical Flaw Size Prediction Based on Limit Load Methodology - hot Leg at Load Critical Location 4-12 4-5 Critical Flaw Size Predictions Based on Limit Load Methodology - Cold Leg at Toughness Critical Locations 4-13 4-6 Critical Flaw Size Fredictions Based on Limit Load.

Methodology - Crossover Leg at Toughness Critical - -

Location 4 4-14 4-7 Critical Flaw Size Predictions Based on Limit Load Methodology - Crossover Leg at Toughness Critical -

Location 8 4-15 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 ReferenceFatigueCrackGrowthCurves{og, 6-5 6-3 6-6 ReferencgFatigueCrackGrowthLawfor[.

la.c. in a Water Environment at 600*F

. A-1 Pipe with a Through-Wall Crack in Bending A-2 1x

- LIST OF FIGURES (Cont'd.)

Title Page .

Figure Typical Layout of the Primary Loops for a Westinghouse B-5 B-1 -

Two-Loop Plant Identification of Heats with Location for Cold Leg 8-6

. B-2 Identification of Heats with Location for Hot Leg B-7 B-3 Identification of Heats with Location for Crossover Leg B-8 B-4 O g O I

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1.0 INTRODUCTION

1.1 Purpose 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 Scope 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 f

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

1.3 Objectives l

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

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a. Demonstrate that margin exists between the " critical" crack size and a i 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.

i 1.4 Background Information Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated from the structural design f 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 l )

mechanics evaluation and a probabilistic analysis to support the elimination 7

of RCS primary loop pipe breaks. That approach was then used as a means of l addressing Generic Issue A-2 and Asymmetric LOCA Loads.

)

1 Westinghouse performed additional testing and analysis to justify the f l

elimination of RCS primary loop pipe breaks. This material was provided to l l

the NRC along with Letter Report NS-EPR-2519 (Reference 4). l l

1 The NRC funded research through Lawrence Livermore National Laboratory (LLNL) l l l

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

indirect LOCA to be 10-7 per reactor year. Thus, the results previously obtained by Westinghouse (Reference 3) were confirmed by an independent NRC research study.

1-2

-.,,*_,.v. -_,__,,._____,_,.__.,___,__._,_.__--_-__._-.._______-,_.-,,..-,,_..------,w-,...w,, ,w---w-

N 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) concludes that an acceptable technical basis has been provided 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 censidering asymmetric blowdown.

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t 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 1

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.,intergranularstresscorrosioncracking),waterhammer,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 Cracking

\

For the Westinghouse plants, there is no history 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 irmnune 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.

4 The environments known to increase the susceptibilty of austenitic stainless steeltostresscorrosionare(Reference 7): oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment. Prior to being put into f'

service, the piping is cleaned internally and externally. During flushes and i preoperational testing, water chemistry is controlled in accordance with written specifications. External cleaning for Class 1 stainless steel piping includes patch tests to monitor and control chloride and fluoride levels. For

- 2-1 1

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 j

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. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and i

fluorides within the specified limits. This is assured by controlling

! charging flow chemistry and specifying proper wetted surface materials.

I 2.2 Mater Hammer i ~

- Overall, there is a low potential for water hamer in the RCS since it is designed and operated to preclude the voiding condition in normally filled l

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 narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the 2-2 l

• ' ~ - ,e--r-em,,_ , . _ _ _ ,

.s.' '

flow and vibration characteristics of the system. Preoperational testing and 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 Cycle and High Cycle Fatigue l

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 l

fatigue crack growth analysis, as discussed in Section 6.

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

vibrations. These are minimized by restrictions placed on shaft vibrations during hot functional testing and operation. During operation, an alarm l

signals the exceedance of the vibration limits. Field measurements have been l

i made on a number of plants during hot functional testing, including plants similar to Prairie Island.. Stresses in the elbow below the reactor coolant l pump have been found to be very small, between 2 and 3 ksi at the highest.

l 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.

e 2-3

3.0 PIPE GEOMETRY AND LOADING

. The general analytical approach is discussed first. 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 4 reactorvesseloutletnozzle(Location 1)istheworstlocationforcrack 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 (Fx) and the bending moment (M b) 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, significant degradations of end-of-service life fracture ,

toughnesses due to thermal aging occur in several pipe segments and fittings.

The highest stressed weld location for which a pipe or fitting suffers such

! - degradation will be referred to as a toughness critical location. The

' associated heat of material will be called the toughness critical material.

As seen in Table 3-1, the toughness critical locations are 1, 4, 8, 10, and

! 10A(seeFigure3-2),notingthatLocation1isalsotheloadcritical f location.

I The loads of Table 3-1 are calculated as follows:

The axial force F and transverse bending moments, My and Mz, 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 Fs, My s, and Mrs. Based on elastic SSE response spectra analyses, amplified pipe

. seismic Joads F d Myde Mzd, are obtained. The maximum pipe loads are obtained by combining the static and dynamic load components as follows:

3-1

-r - - - - ,-,_m-,,w-,.,-

._ _ . . . _ . .=

F, = lFsl + lFd ! .

H #M 2 Mb" y .

where:

M y" lMys! + !Myd M, = lM,3l + lM zd The nomal operating loads (i.e., algebraic sum of pressure, deadweight, and 100 percent power thermal expansion loading) at the locations identified in Figure 3-2 are given in Table 3-2. The loads were determined as described

! above.

The calculated and allowable stresses for ASME III NS-3600 equation 9 (faulted 1.e., pressure, deadweight, and SSE) and equation 12 (normal operating thermal stress) at the load critical location are as follows:

Calculated Allowable Ratio of Stress Calculated /

Equation Stress Allowable Number (ksi) (ksi) i 50.1 0.21 .

- 9F 10.4 50.1 0.25 12 12.4 At the other locations the calculated stresses and ratios are even less.

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3-2

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3 TABLE 3 PRAIRIE ISLAle UNIT 2 PRIMARY LOOP DATA INCLUDING FAULTED LOADING CONDITIONS i

Faulted Loads a 1

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! Inside Wall Yield Ultimate 8en Fng otrect Stress j Weld Radius Thickness Stress Stress Flow Stress Axial Load Moment (ksi) -

Locations (in) (in) o go g,, [ ,

))a.c.e (Kips) (in-Kips) F, M b

] (k 1) (ksi) gg, F

x 4 o, = y+T i

! '1b.c 14.6 2.69 18.8 67.0 42.9 1623 28422 20.3

! 2 14.6 2.69 18.8 67.0 42.9 1621 9554 10.8

) 3 15.6 2.87 18.8 67.0 42.9 1476 16843 11.7 I 4c 15.6 2.87 19.5 67.0 43.3 1682 5717 7.8 I 5 15.6 -2.87 19.5 67.0 43.3

• 1688 4317 7.3 I Y 6 15.6 2.87 19.5 67.0 43.3 1677 25274 15.9 7 15.6 2.87 19.5 67.0 43.3 1761 6598 8.4 8c 15.6 2.87 19.5 67.0 43.3 1762 18332 13.3

, 9 15.6 2.87 19.5 67.0 43.3 1830 17051 13.0 I 10C 13.85 2.55 19.5 67.0 43.3 1529 9432 11.9

! 10Ac 13.85 2.55 19.5 67.0 43.4 1529 8487 11.3 11 13.85 2.55 19.5 67.0 43.3 1530 4997 9.3 12 13.85 2.55 19.5 67.0 43.3 1510 5587 9.5 1

• Includes internal pressure b

c Load criticai location Toughness critical locations

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N TABLE 3 NORMAL CONDITION (DEAD WEIGHT + PRESSURE + THERMAL) LOADS FOR PRAIRIE ISLAND UNIT 2 ,

Weld Axial Load Bending Moment Location Fxi (Kips)a _

y (gn.gggs) l b'C 1453- 23433 2 1453 6910 3 1340 12219 4C 1640 2465 2153 5 1668 6 1656 22803 7 1728 4753 8C 1728 14899 9 1792 11906 10 1426 4326 10Ac 1426 4228 11 1426 3868 12 1422 3703

' Includes internal pressure bLoad critical location C Toughness critical locations 3-4

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1 Crack f '

M

! 2.69 s

~ '

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_______5 _______

"4 @  % '

+ P + 1 h _______.________ M d

h29.2 M i

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I P = 2,235 psi F = 1,623 kips M = 28,422 in-kips 1

FIGURE 3-1 Reactor Coolant Pipe l

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REACTOR PRESSURE VESSEL f  :

HOT LEG COLD LEG 0

R

'.7.n REACTOR COOLANT PUMP STEAM GENERATOR CROSSOVER LEG s ,

O L u O O 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

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4.0 FRACTURE MECHANICS EVALUATION 4.1 Global Failure Mechanism Determination of the ronditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the 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 i

force, and imposed bending moments. The limit moment for such a pipe is given i by: ,

I a,c.e

( )

where:

[

3a,c.e 4-1 1

-Q . .

[

9 f

ja.c.e The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect the Ifmit 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.

i 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 JIe 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 JI c 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:

E

• dJ 7

- app Ta'

}f i

4-2

, , , .-.--,--.---,w-,n--,r

s where:-

Tapp = applied tearing modulus E = modulus of elasticity or = [ la.c.e (flow stress) a = crack length ja.c.e

, [

In summary, the local crack stability will be established by the two-step criteria:

J<J Ic or T,,, < Tmat if J 1 ge 4.3 Material Properties The primary loop piping material for Prairie Island Unit 2 is ASTM SA351-CF8M, a cast product form. The material for the prim'ry a loop fittings is also

! SA351-CF8M. Welds of interest are indicated in Figure 3-2.

l The tensile and flow properties of the load critical location and the toughness critical locations 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 Figure 4-2 taken from Reference 9. J Ic is observed to be somewhat over 5000 in-lbs/in 2. 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

O

• determine the effects of thermal aging on piping integrity, a detailed study was carried out in Reference 10. In that report, fracture toughness results ,

were presented for a material representative of (

( .]a,c.e Toughness results were provided for the material ,

in the full service life condition and these properties are also presenteo in Figure 4 3 of this report for information. The J te value for this material at operating temperature was approximately [ ]a,c.e and the maximum value of J obtained in the tests was in excess of ( ']a c.e The tests of this material were conducted on small specimens and therefore J rather short crack extensions, (maximum extension 4.3 m) so it is expected that higher J values would be sustained for larger specimens. Tmat was (

I la.c.e at operating temperature. The effects of the aging process on the end-of-service life fracture toughness is discussed in Appendix 8.

There end-of-service life toughnesses for the heats are established using the alternate toughness criteria methodology. By that methodology a heat of i material is said to be as good as ( Ja.c.e if it can be demonstrated that its end-of-service fracture toughnesses equal or exceed those of (

]a.c.e. Of the twenty-one heats examined in ,

Appendix B, seven are seen to,be not as good as [ la,c.e,. All the pipe 4

segments of both cold leg loops are among the seven. In the hot leg, only the piping segment of Loop A welded to the reactor vessel outlet nozzle is so i degraded. This particular location is also the most highly stressed (the load criticallocation). In the crossover leg the elbow closest to the steam

generator is so affected as are the piping segments closest to the pump for both loops. Since the load critical location is coincident with a location exhibiting toughnesses less than that of ( la.c.e. the toughnesses associated with this Loop A hot leg heat are conservatively taken as represen-I tative of the toughnesses of all the remaining heats with the exception of the other toughness critical materials. The fracture toughness criteria to be used in the fracture mechanics evaluation, based on the alternate toughness methodology of Appendix B, are given in Table 4-1.

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, 4-4

~

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-- .- ------w-.. -----3 -.,_,.-----,e .-.-,.m.-, -.-.-- -.-c. . . . . __m,- - . , , - . . . . - ---*------e--w*,w.-,-- ~e-

. . - . . -- -- - e._,.-- . - , - .,n ..w.. -

...-a . -

f Available data on aged stainless steel welds (Reference 10 and 11) indicate the J Ic 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 ifmiting than i

the cast material.

In the fracture mechanics analyses that follow, the fracture toughness properties given in Table 4-1 will be used as the criteria against which the l j applied fracture toughness values will be compared.

4 4.4 Results of Crack Stability Evaluation 1

Figure 4-4 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 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 i 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

! determine a critical flaw length, which is shown to be ( la.c.e inches.

In Figures 4-5 through 4-7. plots of the plastic limit moments as a function

! of through-wall circumferential flaw lengths in the remaining four toughness f

critict.1 materials of the main coolant pipe are given. These limit moments were calculated as above using the appropriate pressure, forces, and

dimensions as given either in Table 3-1 or Figure 3-2 with bending moment as a i parameter. The ASME Code minimum properties at 536'F were used. Critical l flaw lengths were determined as in Figure 4-4 by use of the maximum applied

! bending moment. The critical flaw length in Figures 4-5 through 4-7 are all seen to exceed the [ ]a.c.e inches established for the hot leg pipe.

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

4-5 -

l

i'h, . .

. _ . . , 1 j .

o c(circumferentialstress): 11.1 ksi or radial stress: 0 .

The von Mises effective stress, ce ff (see Reference 13) is given by

((ca ~S * (' c - ' r) + ("a - 'c )

.ce ff = t.,

b, and is [ Ja.c.e, ,

Thus the effective stress is less than the yield stress and by the Von Mises plasticity theory yielding does not occur. Also, similar considerations at

! the other toughness critical locations confirm that yielding does not occur there. Hence, linear elastic fracture mechanics is applicable for analyzing the pipes with hypothesiz'ed flaws. The analytical methdo used for the local stability evaluation is sussiarized below.

The stress intensity factors corr'esponding to tension and banding are expressed, respectively, by (see Reference 14)

Kt " 't /Ta F(o) t Eb " 'b /Ta Fb (') ,

, where tF (o) and bF (a) are stress intensity calibration factors correspondingtotensionandbending,respectively,aisthehalf-cracklenhth, f a is the half-crack angle, at is the remote uniform tensile stress, and , s

'I

! ob is the remote fiber stress due to pure bending. Data for Ft (0) and Fb (c) are given in Reference 14. The effect of the yielding near the crack tip can be incorporated by Irwin's plastic zone correction method (see Reference 15)inwhichthehalf-cracklength,a,intheseformula is replaced by the effective crack length, ae ff, defined by i 2

+

1 K a,ff = a + q 2w g j .

L 4-6 i l!

- --- . _ _ , .... .-_,_ ,,,_...__. .. _ _ .,___ _ - m , _ - . _ . , . , , _ . ,

for plane stress plastic corrections, where ey is the yield strength of the material and K is the total stress intensity due to combined tensile and bending loads. Finally, the J I-value is determined by the relation JI=

6 K2 /E, where E =. Young's Modulus (25.3 x 10 psi for 599'F and 25.6 x 106 for 536*F per the ASME Code).

Japp was calculated for the five critical locations using crack length as a parameter. The results are presented in Table 7-1 of Chapter 7 wherein Japp values and leak rates are examined in assessing margin.

At the load critical locatio. (Location 1 in the hot leg) a (

3-a,c.e through wall circumferential flaw is seen to produce a Japp of

( Ja c.e in-lb/in 2 which is below JI c for the toughness critical material located thert.

In the crossover leg the Japp for 7.5 inch long flaws at the toughness critical locations are less than 25% of J Ic- Japp is still well less than JI c for a flaw [ . _,

la,c.e. The same ,is true for toughness critical location 10 of the cold leg.

At toughness critical locaticn 10A in the cold leg where JI c is the smallest, Japp is less than 40% of JI c for a 7.5 inch long flaw. For a

( la.c.e inch long flaw, Japp is still well less than JI c.

In summary, crack initiation is not expected to occur in the Prairie Island l

Unit 2 piping for circumferential flaws at least up to [ la.c.e inches long with the exception being for the toughness critical material in the hot leg.

A flaw at least [ la.c.e inches long is shown to be stable there.

4-7

TA8LE 4 FRACTURE TOUGHNESS CRITERIA USED IN THE LEAX-8EFORE-BREtx-EVALUATION Location or s JeI Jeax 2

s '

2 Description (in-lb/in) Tmat (in-1b/in) r

-.a.c.e e

\

l aincluded in this grouping are the heats at location'l Dhelowerofthevaluesforbothloopsaregivenhere.

6 k

4-8

9 e o

- 8,C,4

/,/////

2.

Neutral Asis

/

E I

8,C,8

( .

FIGUPE 4-1 [ ] STRESS DISTRIBUTION

~

l t

4-9

. - _ _ _ _ - . , . . . . , . _ _ . _ _ . , _ . _ _ _ . . . -______._.._._.,-_,..-7, . _ _ _ _ , , , _ _ . , - , , _ . . _ . . . .r_... _,

~

a,c.e .

t i

l l

l l

i L

Figure 4-2 J vs na for SA351-CFBM Cast Stainless Steel at 600*F l

4-10 L-

3a,c.e FIGURE 4-3 J-aa Curves at Different Temperatures for Agnd Material [

(7500 Hours at 400*C) 4-11

..y . .

- a,c.e L l v i FLAW GEOMETRY OD = 34.58 in.

t = 2.69 in, p = 2235 psi F= 1623 kips e

y

= 18.8 ksi ey = 67.0 ksi of

= 42.9 ksi Temp = 599'F Figure 4-4 Critical Flaw Size Prediction Based on Limit Load Methodology - Hot Leg at Load Critical Location 4-12

a,c.e L l v i FLAW GEOMETRY OD = 32.80 in.

t = 2.55 in.

P. = 2290 osi F = 1529 kips (inc. P) oy = 19.5 ksi u

= 67.0 ksi op = 43.25 ksi Temp = 536*F t

Figure 4-5 Critical Flaw Size Prediction Based on Limit Load Methodology - Cold Leg at Toughness Critical Locations 4-13

s t i a,c.e l

4 FLAW GEOMETRY 00 = 36.94 in.

t = 2.87 in.

P = 2190 psi F = 1682 kips (inc. P) o y = 19.5 ksi a = 67 ksi op = 43.25 ksi Temp = 536*F l

1 l

4 Figure 4-6 Critical Flaw Size Prediction Based on Limit Load Methodology - Crossover Leg at Toughness Critical: Location 4 4-14

o .

L l

- a,c.e I

FLAW GEOMETRY OD = 36.94 in.

t = 2.87 in.

~

P = 2190 psi F = 1762 kips (Inc. P) o y = 19.5 ksi o, = 67 ksi op = 43.25,ksi Temp = 536*F l

l i

I gure 4-7 Critical Flaw Size Prediction Based on Limit Load Methodology - Crossover Leg at Toughness Critical Location 8 .

4-15 l

i

o ,

l l

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 present 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 DH , (L/0H) 15

~

greater than [ 3a.c.e both ( la.c.e must be considered. In this situation 1!he 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 to momentum changes will dominate for ( la.c.e However, for large L/0H values, friction pressure drop will become important and must be considered along with the momentum losses due to flashing.

5.3 Calculation Method The basic method used in the leak rate calculations is the method developed by

(

t ja.c.e,

~

The flow rate through.a crack was calculated in the following manner. Figure 5-1 from 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 given mass flow, the [ 3a.c.e 5-1 l

l d

g was found from Figure 5-2 taken from Reference 16. For all cases considered, since[ la.c.e Therefore, this method will yield _

the two-phase pressure drop due to momentum effects as illustrated in Figure 5-3. Now using the assumed flow rate G, the frictional pressure drop can be calculated using APg=[ la.c.e (5-1) where the friction factor f is determined using the [ Ja.c.e

' The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ Ja.c.e gg3, The frictional pressure drop using Equation 5-1 is then calculated for the assumed flow and added to the (

' la.c.e to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary loop Absolute Pressure - 14.7 = [ Ja.c.e (5-2) for a given assumed flow G. If the right-hand side of Equation 5-2 does not agree with the pressure diff*3rence between the primary loop and the atmcsphere, tnen 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 reconmiended by [

la.c.e for this type of [

la.c.e calculation.

I l

5.4 Leak Rate Calculations Leak rate calculations were made as a function of a crack length for the five critical locations previously identified. The normal operating loads of Table I 3-2 were applied in these calculations. The results are tabulated in Table

! 7-1 of Chapter 7 wherein Japp values and leak rates are examined in assessing margin.

5-2

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

At the load critical location (Location 1)a a 7.0 inch long through wall circumferential flaw is seen to leak in excess of 60 gpm while a flaw half this length gave almost ( la.c.e gpm, At the remaining four toughness critical locations a 7.5 inch long flaw produced leak rates ranging from 8 gpm to 26 gpm. Flaw sizes from (

la.c.e larger produced leak rates in excess of ( lC#qpm at all four locations.

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 critical locations, a factor of at least 15 exists between the calculated leak rate and the criteria of Regulatory Guide 1.45 with two exceptions. A value of 8 gpm for the 7.5 inch long flaw exists at one toughness critical location (Location

4) where the normal operating loads are quite small. A factor of ( la.c.e exists for a flaw ( la.c.e that length. At the load critical location Location 1 a 7-inch long flaw produces over 60 gps.

l l

l l

k l

f 5-3 l

s e a,c.e l

l ,=

. a l =

5 1

i t l

B s

5 a

I l

I

~ 2 STAGNATION ENTHALPY (10 31m/lb)

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

a,c.e

~

, e 4

Y 9

z w

I w

a

' M 2

e u

LENGTM/ DIAMETER MATIO (L/D)

Figure 5-2 [ ]"'C Pressure Ratio as a Function -

of L/D I

5-5 l

I

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

- l l

l

_ _= - :=

b _.

i Figure 5-3 Idealized Pressure Drop Profile Through a Postulated Crack

\

i 5-6 l

i 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 [

la.c.e region of a typical system (see Location

( la.c.e 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(

]a,c.e of a plant typical in geometry and operational characteristics to any Westinghouse PWR System. (

3 ja.c.e All normal, upset, and test conditions were considered and circumferential1y oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in Figure 6-1. Specifically, these were:

.- a,c.e Cross Section A:

Cross Section B:

Cross Section C: ,

Fatigue crack growth rate laws were used [

la c.e 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:

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

~

where X,ff = Kaax (1-R)0.5 min l max i

4 ja,c.e a,c.e a,c.e where: (

]

The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 6-1, and shows that the crack growth is very small, regardless (

ja.c.e 6-2

.~_

l TABLE 6-1 i

FATIGUE CRACK GROWTH AT ( la.c.e (40 YEARS)

FINAL FLAW (in) a,c.e

[ ]a,c.e [ ja.c.e INITIAL FLAW (IN) , ,

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 l

l 9

m 6-3

i

- a,c.e -

i i

l l

1 l

t ,

i s

I

[

Figure 6-1 Typical Cross-section of [ Ja,c,e 6-4 I

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

~

- a.c. e i!T U

u.

N e

w z

N o

a W

E E

s 4

k a

i s

a w

u 4

s u

~ -

Figure 6-2 Reference Fatigue Crack Growth Curves for [

Ja.c.e 6-5

f -

8942 2 4

- a.c. e l

l l

l l

Figure 6-3 Reference Fatigue Crack Growth Law for[ 3a.c.e in a Water Environment at 6000F 6-6 9

7.0 ASSESSMENT OF MARGINS The results of the fracture toughness evaluations of Section 4.4 and the corresponding leak rates of Section 5.4 are summarized in Table 7-1.

j At the load and also toughness critical location (Location 1) Japp for a 7.0 inch long through-wall circumferential flaw is well less than J Ic and exhibits a leak rate in excess of 60 gpm. A crack about ( la.c.e yields 23 gpm while a flaw only [ la.c.e inches long produces 9 gpm, a factor of 9 greater than the 1 gpm of Regulatory Guide 1.45. Thus at the load _

critical location there is a margin of at least two between the flaw size that gives a leak rate well exceeding that of the regulatory guide and the critical flaw size ofC la.c.e, At the remaining four toughness critical locations flaws up to at least

[ la.c.e inches long are stable and leak rates in excess of 60 gpm are exhibited. For flaws ( la.c.e the critical flaw size, the minimum leak rate is at least 8 gpm, taken on an individual basis. Thus again there is a margin of at least two between the flaw size that gives a leak rate well exceeding that of the regulatory guide and the critical flaw size of ( _.

ja.c.e, I- As shcwn 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 stresses.

(

In Section 4.4, the " maximum" flaw sizes at the load critical location and the toughness critical locations are shown using the limit load method to be at least ( la.c.e inches. Thus, based on the above paragraph, the critical flaw sizes at these locations, exceed [ la.c.e inches, j

respectively.

I i

In suunary, relative to

. 7-1

~

1. Loads

a. The Japp values for Prairie Island Unit 2 are anytloped 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 thermal stresses exist relative to ASME Code allowable values.

2. Flaw Size

a. A margin of at least 2 exists between the critical flaw and the flaw yielding a leak rate well in excess of I gal / min.

b. If limit load is used as the basis for critical flaw size, the margin for global stability well exceeds that based on fracture mechanics.

3. Leak Rate -

For all the critical locations the largest flaw sizes shown to be stable with margin exhibit a leak rate in excess of 60 gpm. Flaw sizes half the largest flaws shown to be stable exhibit leak rates well in excess of the criteria of Regulatory Guide 1.45.

l l

7-2 l

TABLE 7-1

SUMMARY

OF Japp AND LEAK RATE RESULTS AS A FUNCTION OF CRACK LENGTH AT THE FIVE CRITICAL LOCATIONS Crack Length Leak Rate Jc I Japp 2 (qpm)-

Location a (in-1b/in1 2

(in) (in-lb/in1 a,c.e 9

23 64c B

d 4

63 8 26d 376 d

10 15

• 153 2

10A 8

d 15 35 61 t

l

c. For this location the flaw size for this leak rate is 7.0 inches.

d. For these locations the flaw size for the leak rates is 7.5 inches.

l l

7-3

= .

k

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 fatigue on the integrity of the primary piping are negligible.

d. Adequate margin exists between the leak rates of small stable flaws and the criteria of Reg. Guide 1.45.

e. Ample margin exists between the small flaws of Item d and larger stable flaws.

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

- For each critical location a fisw is identified which 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 Y F' " " -T"'" -- ' - -.v.-.../,-,__, _, _ _ _

~. .

l

9.0 REFERENCES

1. USNRC Generic letter 84-04,

Subject:

" Safety Evaluation of Westinghouse

- 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. H. 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.

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

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

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

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.

l

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 79-7D3-PIPRE-R1, May 17,1979 (Westinghouse Proprietary Class 2).

t l

l 9-1

0

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

11. Slama, G., Petrequin, P., Masson, S. H., and Mager. T. R., "Effect 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 Boundary Components, August 29/30, 1983, Montersy, CA.

12. Durelli, A. J., et. al., Introduction to the Theoretical and Experimental l Analysis of Stress and Strain, McGraw Hill Book Company, New York,

! (1958),pp233-236.

! 13. Johnson, W. and Mellor, P. B., Engineering Plasticity, Van Nostrand I I 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 II-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).

l l

16. [

js,c.e

17. [

j ja c.e l

l l

! 9-2 L '.

18. Bamford, 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. [

3a c.e

20. [

ja.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 Revision 1. July 1986.

t l

t i

l 9-3

"h &

APPENDIX A LIMIT MOMENT

[

O

.e e 8,C,e

)

A-1

. r-l e l s,

l

.I s - , 1 l

L

& 5 I

l

- - a,c.e s

t l

s t

• l l

FIGURE A-1 PIPE WITH A THROUGH-WALL CRACX IN BENDING

-s A-2 ,

1 l

L

APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE

~

PRAIRIE ISLAND UNIT 2 PRIMARY LOOP COMPONENTS B.1 INTRODUCTION Not all of the individual piping components of the Prairie Island Unit 2 primary loop piping satisfy the original [ la.c.e criter.ia (Reference 10). In this appendix, the alternate toughness criteria developed in

• Reference 21 will be used to categorize the various individual piping components thus establishing criteria based upon which the mechanistic pipe break evaluation may be performed. First the chemistry and calculated KCU values are givet 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.

B.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 l- chemistry and end-of-service life KCU toughness values are given in Table B-1. Of the twenty-one heats, seven fail to qualify as having the toughness properties of ( la.c.e. Both loops are invobfed. Only the hot leg of Loop B is seen to be free of material which is not as tough as that of (

ja.c.e, l B.3 THE AS-BUILT PRAIRIE ISLAND UNIT 2 LOOPS 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 8 in Prairie Island Unit 2.

Sketches,for associating piping component with specific locations and loop are

. given in Fighres B-2 through 8-4. The individual components are identified by heat numbers. The heat numbers above the sketches refer to Loon B; those below, to Loop A. The components which have toughnesses less than that of

( 3a.c.e are identified.

i 1

B-1

B.4 ALTERNATE TOUGHNESS CRITERIA FOR THE PRAIRIE ISLAND UNIT 2 PRIMARY LOOPS ON A COMPONENT-BY-COMPONENT BASIS The alternate toughness criteria for the Prairie Island Unit 2 primary loops may be obtained by applying tha methodology of Reference 21 to Table B-1.

First, it is observed that fourteen of the 21 heats fall into (

la.c.e, i.e., they'are as tough as [ la.c.e. The remaining heats fall into ( Ja.c.e, An example calculation for a [ Ja,c.e heat is given below. Similar a

calculations for the remaining six [ J c.e heats were made. The toughness criteria for the twenty-one heats are given in Table B-2.

The example calculation will be made for Heat [ la.c.e. The ferrite content is [ la.c.e and the end-of-service life KCU is [ la.c.e 2

da3/cm . The [ .

(

.]a.c,e, Since the end-of-service life KCU exceeds the fully aged KCU, the heat falls into ( Ja,c.e. Thus:

Je=[

I 3a.c.e Tmat " [

ja.c.e and Jmax " (* .

ja,c,e B-2

h M e

TABLE B-1 Chemical and Physical Properties of Prairie Island Unit 2 Primary 14op Material - SA351/CF8M a

%<.tn (ti . 8,C,e W

l l

l l

l l

O l

l l

l l

l l

l B-3

--n-a-+=

H  %

f e

U e .

6 I

O M

W E

w I,

L

- E o

E A aC

=

a. 5 a U

N E

b3 a -.

w g

=

W M ==

M y M

W ==E E;

$fm W E p 5

C m

i N

e M

W GJ d e

- 5 a

CL se -

sa O

E

• e d

i L

' B-4

w 1 1

. Steam Generator pg.

ter**

Crossover Leg Hot Leg 4..

i,,/ Loop B Pump Reactor Vessel Cold Leg j ,

I

's Loop A ,/ ~ ',

A..f

+

40

.I i

l l

i l

Figure B-1 Typical Layout of the Primary Loops for a Westinghouse

. Two-Loop Plant t

I

(

l B-5 1

a,c.e Figure B-2 Identification of Heats with Location for Cold Leg 3

B-6

O Y e e'

W

_ a.C.*

[I .

1 I

Figure 8-3 Identification of Heats with Location for Hot Leg l'

e i

5-7 s.

. ,b .o.'

- a,c.e Figure B-4 Identification of Heats with Location for Crossover Leg B-8 n 1