a risk-informed approach to fatigue break criterion for

A RISK-INFORMED APPROACH TO FATIGUE BREAK CRITERION FOR

ASME CLASS 1 HIGH ENERGY PIPING

S. R. Gosselin, P.E.

Lucius Pitkin Inc. Richland, WA. USA

Phone : +1 509 420 7684 [email protected]

F. A. Simonen, Ph.D. Lucius Pitkin Inc.

Richland, WA. USA Phone : +1 509 420 7684

[email protected]

ABSTRACT This paper presents a risk-informed assessment of

fatigue pipe rupture (break) criterion with an application to a 12-inch Reactor Recirculation (RRC) System Class 1 pipe. In this case, a deterministic fatigue usage break criterion is established that will ensure: 1) low break probability, 2) adequate damage tolerance, and 3) insignificant risk. A bounding high stress-low cycle fatigue loading over a 60-year design life is assumed and the environmental fatigue life correlations in NUREG/CR-6909 are applied. Deterministic damage tolerance calculations examine the time for multiple cracks to initiate and grow through the pipe thickness. Probabilistic fracture mechanics models, coupled with industry service experience, relate small and large break loss of coolant accident (LOCA) probabilities and frequencies with cumulative fatigue usage factor (CUF). Risk impact calculations quantify small and large break core damage frequency contributions associated with alternative risk-informed CUF break criterion.

Sample calculations in this paper show that when component fatigue damage is limited to a 60-year environmental CUF<1.0 through-wall crack damage tolerance greater than 60 year design life, extremely low LOCA probabilities and negligible risk impacts can be expected. Additional reductions in the break probabilities and risk impacts can be obtained if 10-year inservice inspections are accounted for. These results suggest that a new criterion based on risk-informed principles can be used to support increasing the fatigue break criterion from 0.1 to the ASME Code design fatigue limit (CUF<1.0). Such criterion, if adopted could save the industry millions of dollars in subsequent dynamic analyses and the elimination of unnecessary pipe whip restraints.

INTRODUCTION The U.S. Code of Federal Regulations General

Design Criterion 4 [USNRC 2010] requires that all nuclear power plant structures, systems and components important to safety be protected from environmental and dynamic effects associated with postulated pipe ruptures. The U.S. Nuclear Regulatory Commission (NRC) requirements concerning the identification of postulated pipe rupture (break) locations and the subsequent dynamic/pipe whip analyses are specified in Section 3.6.2 of NUREG-0800 and Branch Technical Position 3-4 [USNRC 2007a]. With regard to fatigue, the NRC requires that a pipe rupture must be postulated at any reactor pressure boundary location where the cumulative fatigue usage factor (CUF) exceeds 0.1 as calculated by the design-by-analyses procedures in the Section III of the American Society of Mechanical Engineers’ (ASME) Boiler and Pressure Vessel Code (hereafter ASME Code) [ASME 2007].

The present fatigue break criterion (CUF<0.1) was originally established in 1975 and was intended to utilize piping design information to identify pipe locations having relatively higher break potential [USNRC 1975]. The criterion was based the qualitative assumption that actual piping failures generally occur at high stress and fatigue locations (e.g., terminal ends of a piping system at its connection to the nozzles of components).

In 2007 the NRC published Regulatory Guide 1.207 [USNRC 2007b] requiring designer's to account for the effects of light water reactor environments in the design fatigue analyses of reactor coolant pressure boundary (RCPB) components. In doing so the designer applies an environmental correction factor (Fen), as described in NUREG/CR-6909 [Chopra and Shack 2007], to the fatigue usage in air derived from the ASME Code analysis procedures. These Fen formulations can result in large penalties. Depending on the component location

1 Copyright © 2012 by ASME

Proceedings of the 2012 20th International Conference on Nuclear Engineering collocated with the

ASME 2012 Power Conference ICONE20-POWER2012

July 30 - August 3, 2012, Anaheim, California, USA

ICONE20-POWER2012-54534ICONE20-POWER2012-54534

and the extent of computational rigor employed, the original design usage factors have increased by factors as low as 2 and as high as 10.

Since the original fatigue break criteria of CUF<0.1 remains unchanged, the number of postulated break locations in the reactor pressure boundary will significantly increase. The impact of these additional NRC requirements, coupled with the absence of any quantitative technical basis, has highlighted the need to re-examine present fatigue usage break criteria.

This paper investigates the application of a risk-informed fatigue break criterion that rests on a technical basis of 1) low break probability, 2) high damage tolerance, and 3) insignificant risk to core damage or large early release.

INSIGHTS FROM SERVICE EXPERIENCE Significant attention has been placed on establishing

comprehensive databases that compile all reported service induced degradation in operating nuclear power plants. The most comprehensive work in this area is contained in OECD [Mathet et al. 2004] and PIPExp [Lydell 2010] databases. These databases are continuously updated and maintained to cover pipe failures in commercial nuclear power plants worldwide. The PIPExp data presented in this paper covers the period 1970 to November 2010, encompassing over 12,000 reactor operating years of experience, and includes 8113 records on pipe failure events plus an additional 561 records involving water hammer events that challenged or degraded the structural integrity of a piping pressure boundary component.

Fig. 1 ‐ Pipe Failure Data by Degradation Mechanisms [Lydell 2010]

Unlike the assumption that formed the basis for the present break criteria, service failures (cracks, leaks, and breaks) typically result from degradation mechanisms and loading conditions typically not anticipated in the original design. Depending on the degradation mechanism present, failures are not necessarily limited to weld locations. Fig. 1 shows that over 94% of the reported piping failures result from mechanisms not addressed in the ASME Code [ASME 2007] design fatigue analyses. These include flow

accelerated corrosion, stress corrosion cracking, vibration fatigue, corrosion (including microbiologically induced corrosion, crevice corrosion and pitting corrosion), water hammer, and fabrication defects.

Approximately 3% of all reported piping failures were caused by thermal fatigue mechanisms and most all of these resulted from loading conditions not considered in the original design (e.g., thermal stratification, cycling, and striping and thermal mixing conditions). Many of the thermal fatigue failures observed in the field, resulted from the growth of multiple fatigue crack initiation locations on the inside circumference of the pipe [Khaleel, et al. 2000]

In Fig. 2 pipe failure data are presented by system grouping and leak size categories. The majority of RCPB failures were reported as either cracks or leaks less than 1 gallon/minute. Approximately 3% of these were due to thermal fatigue (thermal stratification) mechanisms not considered in the original design analyses. The remainder of the RCPB cracks/leaks is primarily associated with stress corrosion cracking mechanisms.

In summary, service experience indicates that the scope of pipe break criteria should be expanded to address a wider range of degradation mechanisms and failure modes. With regard to thermal fatigue, service experience suggest that RCPB break probability is low even when crack initiation has been predicted (i.e., CUF=1).

Fig. 2 ‐ Failure Data by System Category and Leak Class [Lydell 2010]

CUMULATIVE USAGE FACTOR AND LEAK PROBABILITY

The ASME Code design by analysis procedures were intended to ensure reasonable confidence that the plant will provide reliable service throughout its design life and fatigue usage values in the plant’s Design Report were never intended to provide an accurate indicator of failure potential [Cooper 1992]. Fig. 3 summarizes results of probabilistic fracture mechanics calculations [Khaleel, et al. 2000] that show calculated probabilities of through-wall cracks (leaks) as a function of calculated fatigue usage factors for the selected piping and vessel


components that were addressed by Ware et al. [Ware et al. 1995].

The conservative nature of the calculated usage factors is confirmed by the prediction for many components with high probabilities of leaks for usage factors of 1.0 or even less. In addition, for any given value of usage factor there is a large range in the corresponding calculated values of leak probabilities. The results therefore show a poor overall correlation of usage factors with failure probabilities. Ideally there would at least be some relationship between calculated values of fatigue usage factor and probability of crack initiation.

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08

1.E-07

1.E-06

1.E-05

1.E-04

1.E-03

1.E-02

1.E-01

1.E+00

0.001 0.010 0.100 1.000 10.000

Cu

mu

lati

ve P

rob

abili

ty o

f T

hro

ug

h-W

all

Cra

ck a

t 40

Yea

rs

Fatigue Usage Factor at 40 Years(from INEL NUREG/CR-5999)

Solid Symbols - Water EnvirornmentOpen Symbols - Air Envirornment

Fig. 3 ‐ Calculated Probabilities of Through‐Wall Crack as Function

of Calculated Usage Factor for Both Water and Air Environments [NUREG/CR‐6674]

In Fig. 4 probabilistic fracture mechanics calculations were performed to help explain the trends observed in Fig. 3. Each curve in this plot addresses fatigue crack initiation in a specific piping component subject to stress cycling at specific stress amplitudes. The results show that for any combination of stress amplitude and pipe diameter there is indeed a unique relationship between usage factors and crack initiation probabilities. While stress amplitude appears to be a more dominant factor, the pipe diameter as treated in the probabilistic fracture mechanics model is also a factor.

The larger effect of stress amplitude comes from differences between the fatigue curves in ASME Code Section III and the more recently developed curves used to calculate probabilities of crack initiation. There were differences in both the curves for mean fatigue life and in the methods used to address uncertainties in fatigue life. The code fatigue curves were developed from data available in the 1960s and addressed uncertainties using factors on a mean fatigue life curve (i.e. factor of 20 on fatigue life and factor of 2 on stress amplitude). In the probabilistic fracture mechanics calculations the probabilities of crack initiation were based on more recent fatigue data [Majumdar et al. 1993] and addressed uncertainties with statistical distributions that characterize the scatter in the fatigue data. The net

effect of the different treatments is that the probabilistic model tends to predict longer fatigue lives at high stress amplitudes than at low stress amplitudes.

With regard to size effects the probabilistic models treat each 2 inch length of weld as corresponding to one of the small specimens used to generate the fatigue data. Thus (for example) for a given stress amplitude the probability of crack initiation in a 30-inch pipe is predicted to 30 times greater than for a 1-inch pipe.

1.0E-06

1.0E-05

1.0E-04

1.0E-03

1.0E-02

1.0E-01

1.0E+00

0.001 0.010 0.100 1.000 10.000C

um

ula

tive

Pro

bab

ility

of

Fat

igu

e C

rack

In

itia

tio

n a

t 40

Yea

rsFatigue Usage Factor at 40 Year

Sa= 20 ksi

Sa = 200 Sa = 60

D = 1"

30"

5"

10

20"

1"

1"

30"

30"

CU

F =

1.0

Fig. 4 – Trajectories of Crack Initiation Probabilities vs. Fatigue

Usage Factors for Stress Amplitudes 20, 60 and 100 ksi and Pipe Diameters 1‐30 inches

CUMULATIVE USAGE FACTOR AND LEAK TOLERANCE

Generally, the fatigue usage values tend to be dominated by those selected load set pair combinations containing extremely rapid (step change) thermal transient loading conditions. These conditions result in large intensified peak stresses at the inside surface and significant through-thickness stress gradients. These intensified peak stresses define the time to crack initiation and the fatigue usage. Because of the large through-wall stress gradient (i.e., the ratio of membrane-to-gradient stress is small), the thermal stresses drop-off rapidly at locations away from the inside surface and the driving force for crack growth is significantly reduced. Consequently, the time to grow a crack through-wall (i.e., damage tolerance) can be very long [Gosselin, et al. 2007].

RISK INFORMED BREAK CRITERIA KEY ELEMENTS Consistent with the original objective of NUREG-1800

break criteria are expressed in terms of deterministic CUF parameters calculated as part of the ASME design practices. However, a risk-informed basis for a break criterion can provide confidence that: 1) component fatigue damage will be limited to an amount required for crack initiation (CUF<1), 2) components are able to tolerate fatigue accumulation and growth of cracks that escape detection during initial or final inspections, 3) failure (e.g., through-wall crack, leak, and small and large break loss of coolant) probability is below accepted


thresholds, and 4) system and plant level risk impacts associated with the break criterion are insignificant.

EXAMPLE CALCULATIONS This example calculates a fatigue usage risk-

informed break criterion for 12-inch Schedule 100 Type 313NG stainless steel piping in a BWR/5 Reactor Recirculation System (RRC). Deterministic damage tolerance calculations examine the time for multiple cracks to initiate and grow through the pipe thickness. Probabilistic fracture mechanics models, coupled with industry service experience, relate small and large break LOCA probabilities and frequencies with CUF. Risk impact calculations quantify small and large break core damage frequency contributions associated with alternative risk-informed CUF break criterion.

All calculations were performed using a modified version of the pcPRAISE computer code [USNRC 1995 and Khaleel et al. 2000] that incorporated the stainless steel initiation life model (Fig. 5) and the following environmental factor Fen formulation in NUREG/CR-6909 [Chopra and Shack, 2007]:

Fen exp(0.734 T O )

T , and O are transformed temperature, strain rate, and dissolved oxygen (DO) respectively, defined as follows:

o

o

o

T 0 (T 150 C)

T (T 150) /175 (150 T 325 C)

T 1 (T 325 C)

0 ( 0.4%/ s)

ln(

/ 0.04) (0.0004 0.4%/ s)

ln(0.001) ( 0.0004%/ s)

O 0.281 (All DO Levels)

As defined in NUREG/CR-6909, Fen is simply the ratio of the initiation fatigue life in air airN at room

temperature and the initiation fatigue life in water waterN

at normal operating temperature. The minimum Fen is equal toFen exp(0.734) 2.08 .

The fatigue loading assumed in this study is summarized in Table 1. The alternating stress intensity is representative the dominant cyclic load set pairs at the fatigue sensitive residual heat removal (RHR) return tee connection in NUREG/CR-6260 [Ware 1995]. At this alternating stress intensity, a 10 cycle per year frequency will result in a 60-year end-of-life CUF of 1.0 in air.

Table 1 – Example Fatigue Loading Assumptions

Load Value

Pressure Stess 8.8 ksiThermal Expansion Stress 20 ksiPeak Stress Intensity (Sp = 2Salt) 260 ksiAlternating Stress Intensity (Salt) 130 ksiMembrane-to-Gradient Stress Ratio 0.1

Fig. 5 – Austenitic Stainless Steel Air Fatigue Curve [Chopra and

Shack, 2007]

These cyclic stress levels were used to calculate both crack initiation and crack growth; however, they include effects of stress concentrations (stress indices) in a manner prescribed by the ASME Code. In many cases, the stress indices address very high local stresses (e.g., weld root stress concentrations) with values up to 2.0. Because surface stresses are not indicative of internal stress levels remote from the peak stresses, adjustments were made to account for absence of local stress concentrations on cyclic stresses assumed during the crack growth calculations.

DAMAGE TOLERANCE ASSESSMENT Deterministic damage (flaw) tolerance calculations

were performed to examine the time for cracks to initiate and grow through the pipe thickness. These calculations assume that multiple fatigue cracks will initiate about the entire inside surface. The equivalent single crack (ESC) models and fatigue flaw tolerance methodology developed in NUREG/CR-6963 [Gosselin, et al. 2007] and ASME Section XI Non-mandatory Appendix L [ASME 2010] were used to account for the initiation, growth, and coalescence of multiple fatigue cracks. By definition, an ESC is a hypothetical pre-existing crack that will result in the same cumulative through wall crack (TWC) probability as when multiple cracks initiate about the inside surface and grow through wall. In this example, a pre-existing ESC with an initial depth 18% through wall is assumed.

Fig. 6 shows the TWC (leak) damage tolerance for several 130 ksi cyclic loading frequencies expressed in terms of 60-year CUF values 0.1 to 1.0. At a 60-year CUF equal to the present fatigue break criterion (0.1), the 12-inch RRC piping would have to operate an additional 600 years before the crack penetrates the outside pipe surface. At a cyclic loading frequency resulting in a 60-year CUF of 1.0 the TWC damage tolerance is 70-years (10 years beyond the design life).

BREAK PROBABILITIES AND FREQUENCIES Probabilistic fracture mechanics coupled with industry

service experience, were used to determine loss of


coolant accident (LOCA) probabilities ( LOCAP ) and

frequencies (LOCA ) using the following relationships:

LOCA SBLOCA|TWC LBLOCA|TWC TWC

LOCA / WELD YR SBLOCA|TWC LBLOCA|TWC TWC / WELD YR

P (P P ) P

(P P )

0

100

200

300

400

500

600

700

800

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

ES

C L

EA

K D

AM

AG

E T

OL

ER

AN

CE

(Y

RS

)

60-YR CUF

12‐inch Reactor Recirc PipeType 316NG Stainless Steel

2Salt = 260 ksiESC Depth (a/t) = 0.18

ESC Aspect Ratio (a/2b) = 0.0265

Leak

No Leak

60-yr CUFTWC Damage

Tolerance (Years )

0.1 679

0.2 3480.4 1750.6 1170.8 871 70

Fig. 6 – 12‐inch RRC pipe TWC Damage Tolerance

Break Categories In this study the following bounding small and large

break sizes reported in NUREG/CR-6674 [USNRC 2000] were assumed: Small Break (SBLOCA) = ≥ 20 gpm and <300 gpm Large Break (LBLOCA) = ≥3000 gpm

Conditional Break Probabilities Conditional break probabilities given detected

through-wall leak rates ( LOCA|TWCP ) reported in

NUREG-1829 [USNRC 2008], are shown in Fig. 7. The LWR Class 1 curve covers all pipe diameters and is based on 1632 recorded events cracks, leaks and breaks.

Fig. 7 ‐ Likelihood of Structural Failure According to Service

Experience with Light Water Reactor Piping Reproduced from NUREG‐1829 Fig. D.23 [USNRC 2008]

LWR Class 1 data points from Fig. 7 were normalized using the conditional failure probability at >0 gpm. The leakage value at >0 gpm was plotted as having a leakage rate of 0.1 gpm and the following linear regression was developed to fit the resulting data points:

10 BREAK GPM|TWC 10Log (P ) 0.8766 Log (GPM) 0.0095

The resulting linear curve (on a log-log plot) as shown in Fig. 8 was used in to estimate the following small and large LOCA conditional break probabilities:

2SBLOCA|TWCP 6.4205 10

4LBLOCA|TWCP 8.7590 10

1.E-06

1.E-05

1.E-04

1.E-03

1.E-02

1.E-01

1.E+00

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06

Co

nd

itio

nal

Pro

bab

ilit

y G

iven

TW

C o

f L

eak

Rat

e >

x

x = Leak Rate, gpm

Curve 5 - LWR Class 1

Log10(Probablity) = -0.8766×Log10(Leak Rate) - 0.0095

Data from Lydell Figure D.23 of NUREG-1829 CURVE 5 LWR Class 1 (1632 records)

Fig. 8 ‐ Trend Line Curve for Conditional Probability of Failure Given a Through‐Wall Crack; based on LWR Class 1 Results in NUREG‐1829 [USNRC 2008].

Through-Wall Crack Probability Calculations were performed to estimate the through-

wall crack probabilities. These calculations focused on the contribution of initiated fatigue cracks. The first part calculates the probability that fatigue cracks will initiate as a function of time over the operating life of the plant. The second part evaluates the probability that these initiated cracks will grow to become through-wall cracks. The crack propagation was assumed to start from a 0.5 mm (0.020 in.) deep initiated flaw. This crack depth is near the start of the mechanically small crack region and large enough to ensure that the subsequent crack growth can be predicted by linear elastic fracture mechanics (LEFM).

Cumulative through-wall crack probabilities ( TWCP )

associated with the Table 1 cyclic loadings were calculated for several environmental cases (Fen = 1 to 10). For each case, the time to crack initiation air was corrected for environmental effects ( water airN N Fen ).

The results are shown in Fig. 9.


1.0E-08

1.0E-07

1.0E-06

1.0E-05

1.0E-04

1.0E-03

1.0E-02

1.0E-01

1.0E+00

0 10 20 30 40 50 60 70

TW

C P

RO

BA

BIL

ITY

TIME (YRS)

FEN = 10

FEN = 8

FEN = 6

FEN = 4

FEN = 2

FEN = 1

12-inch PipeTYP 316NG Stainless Steel2Salt = 260 ksi600-cycles/60-yearsInitiation Crack Depth = 0.020-inchsm/sg = 0.1

Fig. 9 ‐ Through‐wall crack probability (No inservice inspection)

LOCA Probability and Frequency In Fig.’s 10 and 11 weld LOCA probabilities ( LOCAP )

and frequencies ( LOCA ) are examined at 60-year

environmental fatigue usage values of 0.8 and 1.0.

1.0E-11

1.0E-10

1.0E-09

1.0E-08

1.0E-07

1.0E-06

1.0E-05

1.0E-04

1.0E-03

1.0E-02

1 2 3 4 5 6 7 8 9 10Fen

We

ld L

OC

A P

roba

bili

ty

PLBLOCA + PSBLOCA at 60-yr CUFWATER = 1.0

Minimum Fen for Austenitic

PLBLOCA + PSBLOCA at 60-yr CUFWATER = 0.8

Fig. 10 ‐ Weld LOCA probability when 60‐yr environmental CUF

equals 0.8 and 1.0

1.0E-09

1.0E-08

1.0E-07

1.0E-06

1.0E-05

1.0E-04

1.0E-03

1 2 3 4 5 6 7 8 9 10Fen

LO

CA

F

reqe

ncy

pe

r W

eld

-Yr

SBLOCA + LBLOCA at 60-yr CUFWATER = 1.0

Minimum Fen for Austenitic SS

SBLOCA + LBLOCA at 60-yr CUFWATER = 0.8

Fig. 11 ‐ LOCA frequency when 60‐yr environmental CUF equals 0.8

and 1.0

The highest TWC probabilities and frequencies occur when environmental affects are minimal (Fen = 2.0). In this case the cyclic the specified 60-year CUF criterion is obtained with higher applied cyclic stresses and subsequent crack growth is faster. Similarly, when Fen is very high, alternating stresses are smaller and the time to grow an initiate crack through-wall is much longer.

RISK IMPACT CALCULATIONS Risk impact calculations quantify small and large

break core damage frequency (CDF) contributions associated with 60-year CUF break criteria examined in Fig.’s 10 and 11. The CDF estimates assume that all welds in the piping have the same bounding fatigue loading in Table 2. The results are presented in Fig.’s 12 and 13.

The CDF for small and large breaks is equal to:

SBLOCA SBLOCA SBLOCA WELDSCDF CCDP N

LBLOCA LBLOCA LBLOCA WELDSCDF CCDP N

Base on a review of weld counts for Reactor Recirculation System (RRC) piping in BWR5 plant, the following number of welds were assumed:

Total RRC System Piping Welds = 200

Total 12-inch RRC Piping Welds = 60

From Table 7.4 in NUREG/CR-6674 the following small and large break conditional core damage

probabilities are assumed: 6SBLOCACCDP 1 10 and

4LBLOCACCDP 5 10 .

1.0E-12

1.0E-11

1.0E-10

1.0E-09

1.0E-08

1.0E-07

1.0E-06

1 2 3 4 5 6 7 8 9 10

SB

LO

CA

+ L

BL

OC

A

CD

F p

er Y

r

FEN

60‐yr CUFWATER > 0.8

NO BREAK ZONE [CUF WATER < 0.8]

12‐inch Reactor Recirculation System Pipe

High Stress ‐ Low Cycle Fatigue Case2 Salt = 260 ksi

Sm/Sg = 0.1SBLOCA CCDP = 1E‐6

LBLOCA CCDP = 5E‐4Number Welds = 60

SBLOCA + LBLOCA CDFYRat 60‐yr CUFWATER = 0.8

Fen CDFWELD‐YR

1 1.075 e‐9

2 1.93e‐10

4 2.84e‐11

6 1.12 e‐11

8 4.44 e‐12

10 2.87 e‐12

SBLOCA + LBLOCA CDFYRat 60‐yr CUFWATER = 0.8

Fen CDFWELD‐YR

1 1.075 e‐9

2 1.93e‐10

4 2.84e‐11

6 1.12 e‐11

8 4.44 e‐12

10 2.87 e‐12

Fig. 12 – 12 inch Reactor Recirculation System piping SBLOCA +

LBLOCA core damage frequency at 60‐year environment fatigue CUF < 0.8


SBLOCA + LBLOCA CDFWELD‐‐YR

at 60‐yr CUFWATER = 1.0

Fen CDFWELD‐YR

1 4.48E-09

2 1.40E-09

4 1.59E-10

6 5.82E-11

8 3.20E-11

10 1.69E-11

SBLOCA + LBLOCA CDFWELD‐‐YR

at 60‐yr CUFWATER = 1.0

Fen CDFWELD‐YR

1 4.48E-09

2 1.40E-09

4 1.59E-10

6 5.82E-11

8 3.20E-11

10 1.69E-11

Fig. 13 – 12 inch Reactor Recirculation System piping SBLOCA +

LBLOCA core damage frequency at 60‐year environment fatigue CUF < 1.0

CONCLUSIONS AND RECOMMENDATIONS In this paper two alterative fatigue break criteria were

examined for a fatigue sensitive RHR return connection to the 12-inch Reactor Recirculation (RRC) System piping. This location was identified as critical location in the environmental fatigue evaluations in NUREG/CR-6260 [Ware et al. 1995] and was considered representative of other RRC piping welds.

These results suggest that a new criterion based on risk-informed principles can be used to support increasing the fatigue break criterion from 0.1 to the ASME Code design fatigue limit (CUF<1.0). When fatigue damage is limited to the ASME Code limit:

Pipes are able to tolerate fatigue accumulation and growth of cracks that escape detection during initial fabrication or field inspections.

Leak damage tolerance significantly exceeds 60-year design life requirements for new reactors and present reactor plants seeking license renewal.

Leak and break probabilities and frequencies are low and below normally accepted thresholds.

System and plant risk impacts are significantly less than limits established for other risk-informed applications.

REFERENCES ASME, 2007. "ASME Boiler & Pressure Vessel Code,

Section III, Division 1, Subsection NB, Class 1 Components," 2007 Edition, American Society of Mechanical Engineers, N.Y.

Chopra, O. K. and W. J. Shack, 2007. "Effect of LWR Coolant Environments on the Fatigue Life of Reactor Materials," NUREG/CR-6909, U.S. Nuclear Regulatory Commission.

Cooper, W.E., 1992. "The initial Scope and intent of the Section III Design Procedures”, Paper presented at PVRC Workshop on Cyclic Life and Environmental Effects in Nuclear Applications, January 1992.

Gosselin, S.R., F.A. Simonen, P.G. Heasler, and S.R. Doctor, 2007. “Fatigue Crack Flaw Tolerance in Nuclear Power Plant Piping – A Basis for Improvements to ASME Code Section XI Appendix L”, NUREG/CR-6934, U.S. Nuclear Regulatory Commission, Washington, D.C.

Khaleel, M.A., F.A. Simonen, H.K. Phan, D.O. Harris, and D. Dedhia 2000. “Fatigue Analysis of Components for 60-Year Plant Life,” NUREG/CR-6674,” prepared by Pacific Northwest National Laboratory for U.S. Nuclear Regulatory Commission.

Lydell, B.O.Y. 2010. “PIPExp-2010 High Level Summary of Database Content as of March 31, 2007,” Scandpower Inc., Huston, TX, USA.

Mathet, E., K. Gott and B.O.Y. Lydell 2004. “OECD Pipe Failure Data Exchange Project – 2003 Status Report,” ICONE12-49217, Proc. 12th International Conference on Nuclear Engineering, Arlington (VA).

Majumdar, S., O.K. Chopra and W.J. Shack 1993. "Interim Fatigue Curves for Fatigue Crack Initiation and Propagation", NUREG/CR-5999, U.S. Nuclear Regulatory Commission.

Ware, A.G., D.K. Morton, and M.E. Nitzel 1995. "Application of NUREG/CR-5999 Interim Fatigue Curves to Selected Nuclear Power Plant Components," NUREG/CR-6260, U.S. Nuclear Regulatory Commission.

USNRC 1975. “Postulated Break and Leakage Locations in Fluid System Piping Outside Containment,” Branch Technical Position MEB 3-1, U.S. Nuclear Regulatory Commission.

USNRC 1987. "Determination of Break Locations and Dynamic Effects Associated with the Postulated Rupture of Piping," NUREG-75, Section 3.6.2, U.S. Nuclear Regulatory Commission.

USNRC 1995. "Theoretical and User's Manual for pc-PRAISE – A probabilistic Fracture Mechanics Code for Piping Reliability Analysis", NUREG/CR-5864, U.S. Nuclear Regulatory Commission.

USNRC 2007a. "Determination of Break Locations and Dynamic Effects Associated with the Postulated Rupture of Piping," NUREG-0800, Section 3.6.2, U.S. Nuclear Regulatory Commission.

USNRC 2007b. Regulatory Guideline 1.207, U.S. Nuclear Regulatory Commission.

USNRC 2008. “Estimating Loss-of-Coolant Accident (LOCA) Frequencies through the Elicitation Process”, NUREG-1829, Vol. 1 & 2, U.S. Nuclear Regulatory Commission.

USNRC 2010. “General Design Criteria for Nuclear Power Plants,” Title 10 Code of Federal Regulations Pat 50, Appendix A, U.S. Nuclear Regulatory Commission.


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