2ND

INTERNATlONAl
THE

WORKSHOP

INTEGRITYNUCLEARONOF

COMPONENTS

April20121，1998

主催

社団法人目本溶接協会

原子力研究委員会
Japan Welding Engineeing Society

The

AtomicEnergyReseach

Sa血ty

and

Structural

Sung

Kyun

Kwan

Committee

Integrity

University

ResearchCenter，Program of the Workshop

Session I

Design and Analysis

I- 1

Concept of Advanced Design Standards for DFBR
Masahiro Ueta, JAPC, Japan

CONCEPT OF ADVANCED DESIGN STANDARDS FOR DFB
Masahiro UETA

THE JAPAN ATOMIC POWER COMPANY
1-6-1, Ohtemachi, Chiyoda-ku, Tokyo 100, JAPAN

Phone : 3-3201-6631 Fax : 3-3285-0542
E-mail : masahiro-ueta@japc.co.jp

Abstract

The Japan Atomic Power Company (JAPC) has been developing thedemonstration fast breeder reactor (DFBR) under the sponsorship of nine

Japanese electric companies and the Electric Power Development Co.Ltd
since 1986. JAPC has drafted the structural design guide for the DFBR atelevated temperatures (DDS) based on that for the prototype FBR MONJU.
However, in the last December the round-table committee, established in the

Atomic Energy Commission, published the report that the construction of the

DFBR should be projected based on experiences of operation andmaintenance of MONJU and reflecting latest R&D achievement. As theresult the DFBR project may be somewhat postponed. Meanwhile, therequirement to the DFBR are becoming more and more strict toward higher
economy consistent with reliability. To cope with the requirement, JAPC has

started drastic revision of the DDS. The main issues in the revision are

rearrangement of code systems and inrprovement of estimation method ofload conditions, accumulated strains by ratchetting and creep-fatigue
damage. This paper describes the concept of the advanced DDS.

1 . Introduction

Worldwide energy demand has been increasing with population
increase and economic development especially in developing countries,
although many countries are making efforts to save energy consumption. If

the consumption of fossil energy increase with energy demand, energy
resources, such as oil and natural gas, may be run out of in the middle of the-1-

next century. And it is also forced to limit consump~tion of fossil energy fromthe view point of protecting global environment from greenhouse effect andacid rain caused by exhausted gas like carbon dioxide and other oxidant

gases from fossil power plants. Nuclear power generation is the mostpromising and effective measures to cope with both energy demand andenvironment protection. Nuclear people convince that nuclear power
generation will play the essential role in energy supply in the next century.

Fast breeder reactors (FBRS) have the potential to utilize uranium
resources much more effectively than light water reactors (LWRs) by morethan 50 times, and so it is expected that FBRs can secure worldwide energy

supply for more than thousands of years. FBRs also have fascinating
potential to be able to incinerate minor actinides which are radioactive

wastes of ultra-long half-lives originated from nuclear power plants.
Therefore, Japanese government and utilities have been promoting thedevelopment of FBR since the initial stage of nuclear energy development.

The Japan Atomic Power Company (JAPC) has been developing thedemonstration fast breeder reactor (DFBR) following the prototype FBRMONJU under the auspices of nine Japanese electric power companies andthe Electric Power Development Co.Ltd. since 1986. Main specifications of
the DFBR were settled in January 1994. In December 1995, the sodium leakincident occurred in MONJU. It showed that it is essentially important for
FBRs to prevent sodium fire which may result in long-term plant outage andlower plant availability of FBRs. And deregulation in electric powerindustries forces nuclear industries to reduce power generation cost as fossil

power generation of which cost is drastically reduced recently. Therefore,

JAPC has been conducting the conceptual design study of the DFBR aiming
at eliminating anxiety of sodium fire and reducing construction cost of the

DFBR far below 1.5 times that of LWRs on the 1000MWe basis. In this study

the DFBR would be provided with comprehensive countermeasures against
sodium leak and fire to eliminate instrumentation well through sodium
boundaries, to enclose whole sodium piping by guard pipes, and to install
auxiliary sodium systems as purification systems into vessels.

The MONJU incident induced nationwide discussion on FBRdevelopment in Japan, and the round-table committee was established in the

Atomic Energy Commission. This committee reported in December 1997 thatFBR is one of the promising options of future energy resources, that MONJU-2-

should be restarted for research and development (R&D) of FBR, and thatthe DFBR should be projected reflecting experience obtained in operationand maintenance of MONJU and incorporating latest R&D achievements.
Therefore, sufficient time is kept until construction of the DFBR enough toachieve innovative R&Ds for improving economy of the DFBR drastically.

The current structural design guide for the DFBR at elevatedtemperatures (DDS) was drafted based on that for MONJU and incorporating
R&D achievements after MONJU. It has been applied as the design standard todesign studies of the DFBR. However, it is required to improve accuracy ofdesign evaluation for realizing economical design consistent with reliability.

Therefore, JAPC has started the long-range study for the advanced DDS as a

typical embodiment of innovative R&Ds.
This paper describes the policy to revise the DDS.

2. Policy to Improve DFBR Design
2.1 Reducing Uncertainty in Design

The nuclear steam supply system (NSSS) of the DFBR should havereliability prior to others, because sufficient experiences have not been

accumulated in FBR operation and maintenance. Economy improvementcompatible with reliability can be realized by reducing uncertainty in whole
design process as shown in Fig. 1. The policies to reduce uncertainty are as
follows.

(1) Material

To reduce scattering of material strength, chemical compositions andheat treatment conditions should be specified rigorously as far as material
cost does not increase significantly.

To reduce inaccuracy in extrapolation of material strength, data of
10ng-term creep rupture, ultra-high cycle fatigue and creep-fatigue in low
strain range should be acquired.

To improve prediction of time-dependent effects, data of long-term
aging effect at elevated temperatures, neutron irradiation effect and sodium

environment effect on material strength should be acquired.

-3-

(2) Structure

To reduce deviation of strength from normal value and to eliminate
unforeseen trouble, requirements on quality assurance should be imposed on
design, fabrication, construction, installation and inspection considering cost
imp act.

To reduce uncertainty of structural strength caused by worker's skill
and discontinuities in material or geometry, simple and robust design should

be aimed at.
(3) Loading condition

Thermal transient load is critical in assuring structural integrity of
FBR components since it is much more difficult to predict thermal transient

10ad than pressure load. Improvement of thermal load prediction willcontribute to economical design consistent with reliability. Thermal
hydraulic model tests are useful to understand thermal hydraulic behaviors,

to improve accuracy of thermal load prediction and to verify computer codesfor design analysis.

Usually thermal stresses are classified into the secondary stress, butthey may act as the primary stress in case elastic follow-up is significant atelevated temperatures. Therefore, strain intensification due to elastic
follow-up should be evaluated carefully in elevated temperature design. The

current method employed in the DDS has shown to be too conservative inmany cases, To realize economical design consistent with reliability, newconcept considering the mechanism of elastic follow-up should be introduced

into the advanced DDS.
(4) Design criteria of failure

Design criteria of failure should be well correlating with actual failuremodes and should conservatively envelope actual failure limits as close as

possible. Creep damage evaluation is the key toward well-correlated design
criteria. The current method applies time fraction concept for creep damage
evaluation. To realize economical design consistent with reliability, creep

damage evaluation method should be improved through discussions taking

suffrcient time to reach consensus among the related specialists andengineers. New methods as ductility exhaustion concept will be investigatedin addition to modifying the current time fraction concept.

~4-2.2 Safety Margin
Design margins should be kept sufficiently in design criteria in DDS,
for example, 20 of number of cycles and 2 of strain range in fatigue design.

Different margins should be defined dependent upon frequencies andsignificance of failure events.

Evaluation methods applied in design should envelope failure limits on

a conservative side, but as close to the limits as possible. In other word,

design margin should not be shared to inaccuracy of evaluation method,while additional margin is not expected in evaluation methods.
2.3 Inelastic analysis

Computer analysis is making great advance in recent years. Inelastic

analysis will be employed widely to DFBR design to realize economical
design consistent with reliability. JAPC will employ inelastic analysis toestimate ratchetting strain in design, to find optimized design by numericalexperiments in place of model tests, to define appropriate elastic follow-up
parameters, and to confirm that sufficient design margin is kept.

The guideline of inelastic analysis should be prepared to assure
conservatism and to attain licensability of designs employing inelastic
analysis.

3. Policies for Advanced DDS

3.1 General
The advanced DDS will adopt the graded code ,system as shown in Fig.2.

The discussion has started to introduce the graded code system into thedesign of plants in non-nuclear fields with future prospect of introducing thatsystem also into nuclear fields.

The first grade is a performance-based code, such as MITI ordinance No.

62, which specifies only requirements from the view point of safety andreliability. The second grade is a general structural design code which
provides the limits of load, stress, strain, damage and other parameters fromthe viewpoint of assuring structural integrity.

The third grade is a

supplemental code which would provide local rules, for instance, seismic

design guideline under a site-dependent earthquake condition. Non--5-

mandatory rules for special configurations such as tube plates are also
included in the third grade. The guideline of inelastic analysis may beclassified into the third grade. The fourth grade is the document to explain

the basis and background of codes and standards. The document would
include data of tests, experiments and analyses to validate the code.

The advanced DDS will aim at the code which has merit to do detailed

design analysis. For this purpose the advanced DDS will be composed ofscreening rules using simplified analysis, standard rules using detailedelastic analysis, and general rules in which it would be permitted to validate

the design by experiment and detailed inelastic analysis.

3.2 Main items to be revised(1) Accumulated inelastic strain

While current codes provide constant values of accumulated inelasticstrain limit independent of material, the advanced DDS would define thelimit values material by material to take advantage of improved ductility.

Several modes of ratchetting may occur under the combination ofprimary loads and secondary thermal loads as shown in Fig.3. It is not easy

to estimate inelastic strain due to all ratchetting modes using elastic
analysis. On the contrary, inelastic analysis can conservatively estimate
whole ratchetting behaviors by one calculation using appropriate model andconsidering all loading conditions. Therefore, the advanced DDS would adopt

design using inelastic analysis to ratchetting as a standard method. The
guideline of inelastic analysis should be prepared to provide the procedure toensure conservatism of analysis results.
(2) Fatigue

High cycle fatigue is a important cause of cracking and failure ofstructures in FBR, because sodium flow with high velocity at different
temperatures may occur flow-induced vibration and thermal fluctuation. The
key issue is to establish the method to estimate loading conditions as forces,

stresses and frequencies. The guideline should be prepared for high cycle
fatigue design to enhance reliability.

Since ultra-high cycle fatigue data are very limited at elevatedtemperatures, current elevated temperature fatigue curves are drawn by
linear extrapolation to ultra-high cycles without fatigue limit, although it-6-

may be too conservative. High cycle fatigue data should be acquired to define

design fatigue curves with balanced safety margin.
(3) Creep-fatigue

It is thought that structures in FBRs may have huge margin in creep-

fatigue because creep damage evaluation methods have over-conservatism.

JAPC is studying to improve evaluation methods of creep-fatigue damage
based on the following policies.

a. The objective of improvement is to realize economical design consistent
with reliability.

b. The damage evaluation method should well correlate with actual failure
behavior in service conditions.

c. The damage evaluation method should not have so many material
constants. Material constants should be defined without any special
sophisticated tests.

d. It is recommended that damage can be evaluated using easily available

material data
e. The same concept should be applied to both base metal and weldment of
any type of austenitic stainless steels,(4) Strength of weldment
It is well known that weldment has much higher probability of cracking
and failure than base metal region has because of geometrical and material

discontinuities. Residual stress and aging deterioration may affect the
integrity of weldment. It is difficult to predict the effects of those factors on

weldment strength by computer analysis. It may be feasible as an evaluation

method for weldment to modify the method for base metal by introducing
coefficients which would consider reduction in strength due to factors as
above mentioned. Since the coefficients are critical to realize economical
structural design consistent with reliability, it should be defmed carefully

based on results of material tests, structure model tests and computer
analyses.

(5) Buckling

The DFBR will adopt horizontal seismic isolation design to the reactor

building with the prospect of commercial FBRs applying three dimensional-7-

seismic isolation to standardize plant design. Horizontal seismic isolationbuilding may intensify vertical floor response compared with aseismicbuildings. Therefore, the advanced DDS will cover buckling of shells undervertical seismic loads considering interaction with horizontal seismic loads.

4. Future Program

JAPC has started the DDS committee by entrusting to The Japan
Welding Engineering Society (JWES). The DDS committee has 3 working
groups, code system WG, creep-fatigue WG and ratchetting WG. In near
future material WG will join in.

Since globalization of code systems is progressing, it is important for

the DDS to be in line with overseas codes such as ASME in U.S.A, RCC-MR

in France, R5 in UK and so on. JAPC would welcome specialists andengineers in the world to join discussion of the advanced DDS, and JAPC

have the plan to do collaborative work in benchmark analysis and dataacquisition.

We have the perspective to establish the advanced DDS as follows.

-2005 : First draft ofthe advanced DDS-2010 : Discussion and approval of the advanced DDS in the codecommittee

5. Concluding Remarks

Japanese government and utilities have been promoting thedevelopment of FBR which will play the important role in energy supply inthe 2lst century. The Japan Atomic Power Company (JAPC) has been
conducting the development of the demonstration fast breeder reactor(DFBR) since 1986. The conceptual design study of the DFBR are ongoing
aiming at eliminating anxiety of sodium fire and reducing plant construction

cost below 1.5 times that of LWRs on 1000 MWe output basis. However,construction of the DFBR may be postponed through discussion triggeredby sodium leak incident in MONJU. We have somewhat long ti_me until
construction of the DFBR. Therefore, JAPC is challenging to the advanced
DDS which will contribute to enhance economy of the DFBR consistent withreliability.

-8-

The key issues in the advanced DDS are to reduce uncertainty instrength of materials and structures, to correlate design criteria well withactual failure behaviors, and to improve accuracy of evaluation methods
especially of ratchetting, high cycle fatigue and creep-fatigue.

JAPC has started the DDS committee to draft the advanced DDS byentrusting to The Japan Welding Engineering Society (JWES). The
advanced DDS will be embodied around 2010.

-9-

Average strength

Material strength

Time-dependent effects

Design margin

Design point
Material strength for design
tructura strengt

Failure criteri!~lDesign limit
Evaluation method
Elastic follow-up

Loading condition
Omlna COn I lOn

Figure 1

Uncertainty in Design

H

H'
I

Frgure 2

DDS code system to be aimed

l~ ~) Pnmary load
Secondary thermal stressl) by axial temperature distribution
sodium surtace ~

l

~.
N)
i

2) in wall thickness

Temperature distributiOn R+1) ~ pM & SM
in wall thickness------ ~)+2) ~ PM & SB (Bree)

Figure 3 Ratchetting modes

I-2

A Study of the Helical Effect on the Connection by Three

Dimensional FEM Stress Analyses
Jien-Jong Chen &Yan-Shin Shih, Institute of Nuclear Energy Research Atomic
Energy Council, Taiwan, Republic of China

A study of the helical effect on the thread connection

by three dimensional finite element analysis

Jien-Jong Chen~b and Yan-Shin Shihb'Nuclear Division, Institute ofNuclear Energy Research, P. O. Box 3-3,

Lung-Tan, Taoyuan, Taiwan, ROC
b Department of Mechanical Engineering, Chung-Yuan Christian University,

Chung-Li,32023 Taiwan. ROC

Abstract
The three dimensional finite element analysis of the bolted joints with finite sliding

deformable contact has been studied, and the helical and fiiction effect on the load

distnbution of each thread is analyzed. It shows the analyiical analysis by
Yamamoto's method reaches a lower value of load ratio than the finite element analysis
at the first thread. The load distribution on each thread between axisynunetrical modeland three dimensional model are provided. HenGe, although increasing the coefficientof friction and decreasing of the lead angle may improve the load distribution slightly,
for I "-16UNF bolt joints, the error of load ratio at the first thread in axisymmetricalfinite element model is 1 20/0 with respect to three dimensional analysis.

I. Introduction
The bolted joint is a typieal counection that is widely used for the construction of

stuctural Gomponents. Owing to the easy replacement and installation, the bolted

joint has been applied to many components of nuclear power plants. However, two

problems of bolted joints connection should be concemed. The first one is the-13-

functionality of the joints which threaded end olosure is still kept, and the second is theload bearing fraotion on each thread of. the bolt and nut. In the recent year, many
authors (Chaaban et. al. 1992, Grosse et. al. 1990, Wileman et. al 1991, Lehnhoff et. al

1 996) have focused on the study of threaded end olosure of the bolted j oint. But theload distribution analysis has bcen seldom mentioned durirmg last few decades. The
investigation of the load distribution in the threaded couneetion has been studied sincel 940's. The Sopwith theory (Sopwith 1 948) for predicting the load distribution of thethreaded fasteners is a well known analyiical model. The action of a number of strains

is formuated by the axial extension of the bolt and eompression of the nut in Sopwith
theory. These strains include the bending defleetion of the thread, an axial recession
due to radial compression of the threads, and an axial reeession due to axial contractionof the bolt and expansion of the nut caused by radial pressure of the joints.
Altematively, Yamamoto (Yamamoto 1 980) proposed a procedure for calculating thedefleotion due to bending moment, shear loading and radial contraction and expansion

on the bolt and nut. The assumption ofplane strain has also been made by Yamamotofor dealing with thread region. Hence, the three dimensional bolt-nut assembly can bc

simply analyzed by axisymmetrical model to caleulate the load distribution along theaxial direction ofbolt. The analyiical and axisymmetrical fmite element analysis havebeen studied in the literature (Chaaban et. al, 1 992, Grosse et. al. 1 990, Wileman et. al

1991, Lehnhoff et. al 1996). However, the lead angle due to helical effect can not bemodeled by axisymmetrical model, such that, the helical effeet to the load distribution isstill lacking on the literature. Due to the progress ofthe modem fmite element method,modeling of the fmite sliding, and deformable to deformable contact problem becomes

possible by usmg ABAQUS code( Hibbitt et.al 1 998). The cunosrty rs nsen for- 14 -

exploring the differenee among theoretioal, axisymmetrical and three dimensional fmite
element models, and that is the purpose of the present study.

II. Modeling and Assumption.
A standard one ineh bolt (8UNC) and two fine threaded bolt (12UNF and 1 6UNF) areused to study the helical effect ofthe bolts. Four axisymmetrical finite element models

with the different mesh on the threads are shown in Figure I . In the axisymmetricalmodel, the radial symmetric is assumed on the center line of bolt. The three
dimensional model ofthe standard bolt assembly is shown in Figure 2, and the model ofthe fine thread bolt is shown in Figure 3 . Due to the difficulty ofthe three dimensionalmodeling, a small hole is assumed at the center of the bolt, this will slightly reduces thearea of the applied load. The elastie material is used through out this work, where theYoung's modulus(F) is assumed as 30E+6 psi, Poisson ratio( v) as 0.3, and thecoeffieient of friction( p) as 0.1 in this analysis. The uniform pressure loading (p=l OOOO psi) is applied on the top root surface of the bolt in this study.

III. Analytical Analysis
The load disinbution analysis has been calculated by means of Yamamoto's method.
The total defleetion of the bolt( 6b )and nut( 6~ ) can be expressed as

6;b = Kbwoosa
5 = K~wc06a (1)andwhere K is the stiffness of the bolt or nut, w is the unit applied load, the sub index b andn represents the bolt and nut. From the geometrieal parameters of this analysis model,
the stifiness of bolt and nut can be obtained as follows,

Kb = 3.57 and K 4 70 (2)The load distribution equation can be expressed as,- 15 -

F = F dd,(Ax) (3)b dnb(AL) 'where Fb is the load on the flfSt thread of the bolt. L is the engaged length of the bolted
joints, and ;t is a Gharacteristic length and expressed as

1 +l

;t=
~ A.E. _ (4)K)[ Kb: AbEb
0.0847

Eb+EtanP

where A is the crossseGtion area, p is the lead angle Hence the result of the load
disinbution on eaGh thread can be shown on Table I .

I.

Finite clement analysis.
From the literature(Chaaban et. al. 1 992, Grosse et. al. 1 990, Wileman et. al 1 991),

the extremely high density of fmite element mesh of the bolted j oint is used for studying

the bolt. However, these high dense mesh not only expends a lot of computer resouree

on calculating but also increases the difficulty of three dimensional fmite element

analysis. Hence, the mesh density verifioation is studied for axisymmeincal model.

firstly. Four models of the axisymmeincal finite element analysis with different
meshing density are shown in Figure I . The load distribution analysis for thesemodels of fmite element meshing are shown as Table 2. For the I "-8UNC bolt, thetotal force (Fb) due to applied pressure is 5 1 60 Ibf.

F =P;T'di2
loooo*'r(0.8ro557)2 5leo(Il/ )4where di is the root diameter of the bolt. The maximum error of total contaot forGewith respect to total applied loads on Table 2 is I .55E-4, that is shown a good-16-

agreement with four types of axisymmetrical meshed model. From the Figure 4, themaximum deviation is happened at frrst thread of the bolted j oints, and the maximum

differential is 60/0 between GOarse mesh model and fine mesh model. Therefore, thecoarse meshed model used in this study can be proposed for studying helical andfriction effects on the three dimensional fmite element analysis. Figure 5 shows theload distribution of analytical, coarse meshed axisymmetrical and three dimensional
finite element analysis results. The load distribution with analytiGal prediction has
shown somewhat lower than the result of fmite element analysis.

II. Result and Discussion
A. Helical effect on the loed distribution.

The simplified axisymmetrical model and three dimensional model of I inch 8UNC,
l 2UNF and 1 6UNF of threaded connections are analyzed in the present work. The
load disinbution on each thread of the axisymmetrical model and three dimensionalmodel are shown on Table 3 and Table 4. Through the comparison of load ratio of

axisymmeincal model with three dimensional model from Figure 5-7, The results
show the loading ratio quite similar on eaoh thread between axisymmetrical and three

dimensional models exGept 1 6UNF threaded joints with a little difference. HenGe, it
can be concluded that the helical effect does not influence the load distribution, and theaxisymmeincal model can give a well estimating on load disinbutions of 8UNC and

12UNF joints. For 16UNF, Ioad ratio of the first thread has 120/0 error of
axisymmetrical model with respect to three dimensional model. Besides, the other two
contaGt forces orthogonal to axial direction can be calculated in the three dimensionalload distribution analysis. It shows that these two contact forces are only the small
fraction of the total applied load, and these can not be obtained from axisymmetrical-17-

ftnite element analysis, because of the orthogonal contact force is the response of

opposite reaction force on the axisymmetrical model. Comparison of the load
distribution among 8UNC, 1 2UNF and 16UNF threads, shown in Figure 8, one canobtains that the load distribution at the flfst thread is improved significantly with a finer

thread configuration.
B.

Effect on the coefficient offriction to the loed distribution

There are four friotion coefficient have been used in this section, which is O, O. I . 0.3

and 0.5. The results ofload distribution with different coefficient of friction have beenshown on Table 5. It is shown that increasing the coefficient of friction improves theload distribution very slightly, but inereasing the coefficient of friction reduces slightly
the load bearing on the first thread,

Owing to the lead angles of the one inch 8UNC, 1 2UNF and 1 6UNF are the values of

2.66, 2.02 and I .39 degree, these angle are such small and may not alter the normal
contact force signifioantly. However, the results of load ratio on the frrst thread in thethree dimensional fmite element analysis exist somewhat lower than the axisymmetricalanalysis. Hence, the load distribution of the bolted joints assembly can be modeledwith axisymmetrical models for I " 8UNC and 1 2UNF bolted j oints.

I. Conclusion
The three dimensional load distribution on the finite sliding contact analysis between

two deformable elastic bolted joint assembly with 8 active threads have been studied.
The results oan be coneluded as follows.

l . The analytical calculation of load distribution by Yamamoto's method on the flfSt
thread is lower than the results from axisymmetrical and three dimensional finite
element analysis of one inch 8UNC standard thread bolted j oints .

-18-

2 The load distnbution of the first thread is improved more significantly by using

l 6UNF instead of 8UNC bolted joints for axisymmetrical and three dimensional
contact analysis.

3 . The effeGt of coefficient of friction on the load distribution of bolted joint assemblyis not evident, although increasing of friction will improve the load distribution
slightly.

4. The deviation of load ratio at the frrst thread are small for I " 8UNC and 1 2UNF,
but the error approaches to 1 20/0 for 1 6UNF bolted j oint.

Acknowledgnrents
This work is supported by Institute of Nuclear Energy Research. The authors wish

to aeknowledge the help from our colleagues at INER. The authors also wish to
express our appreciation to Dr. Kuen Ting encouragement .

Reference
Chaaban A., Jutras M., 1 992. Statie analysis of buttress threads using the fmite element

method" Joumal ofPressure Vessel Teohnology Vol. I 14, pp. 209-212.

Grosse I R Mitchell L. D., 1990 Nonlmear axral stifiness charactenstic of
axisymmetric bolted joint. Transaction of the ASME, Jounral of Mechanical Design,Vol. I 12, pp. 442 449

Hibbitt , Karlsson, Sorenson 1 998. ABAQUS Finite Element Code version 5 .7.
Lehnhoff T. F., Wistehuff W. E., 1 996, Nonlinear effects on the stress and deformationsof bolted joints. Transaction of the ASME, Journal of Mechanical Design, Vol. I 1 8,

pp. 54-58.
Sopwith D. G., 1 948. The distribution of load in screw threads. Institute of Mechanieal

Engineering, Applied Mechanics Proceedings 1 59, pp. 373 383-19-

Wileman J.. Choudhury M., Green I., 1 991. Computation of member stiffness in bolted
comeGtions. Transaction of the ASME, Journal ofMechanical Design, Vol. 1 1 3, pp.

432437.
Yamamoto A 1 980 The Theory and Computation of Thread ConneGtion Pressed
Yonkendo, Tokyo pp. 39-54.(in Japanese)- 20 -

Table I .

The load distribution of each thread by Yamamoto method.

Table 2. The oontact force on each threads of the different meshdensity of axisymmetrical finite element model.

Table 3 .

The load distribution of axisymmetrical model with three
types of bolt joints.

-21-

Table 4. The load distribution of three dimensional model withthree types of bolt joints.

Table 5 .

The load disinbution with effect of frietion coefficients.

- 22 -

Coars e DenBe Mesh

High Dense

Fig l(C):Fine meshed model

Fig I (a) :Coarse meshed model

Middl e Dense

Extremely High Dense Mesh

Fig l(b):Middle meshed model
Fig I (a) :Extremely high meshed model

Fig I : Four types of the finite element axisymmetrical models.

-23-

~~~~)~~~ ~~~ ~ ~Fig 2 : The three dimensional model mesh of standard bolted j oints
assembly(upper) and bolt itself (Iower)- 24 -

~~A~~~ ~l ~Fig 3 : The three dimensional models of the I inch 1 2UNF and 1 6UNF bolted

joint assembly- 25 -

2000
Meshing density

Q 1500
JD
~.

~

=~
~~
~-

coarse meshed
middle meshedflne meshedextremely fine meshed~o 1000
~S

8
500

O
3
4 5nut flxed 6surtace7
1 Thread
2 number
from
Fig .

4 : The contact force on each thread for different
axisymmetrical flnite element model- 26 -

8

0.4
Yamamoto's (1 9eo)

-+,

~
~!
c9

Axisymmetrical modelH~~o 0.3Three dimensional model

li

CL
cs

c9

1.

t'o0.2

o
d*

o
~=
c5L

U
J

c:,o 0.1

0.0

3
5
7
4
1 Thread
2 numbertrom nut flxed surface
Fig. 5 : The load ratio on each thread for I " 8UNC bolted
joints

0.4

1:I

CU

-F
=~

~2
Uo 0.3~QL

12UNF, Axisymmetrical model

12UNF, 3-D model

CL
cU.
-(U

'H
o 0.2
4o
d'

o
~=
~

U
J

CDO 0.1

0.0

123

45678

Thread number from nut fixed surface

Fig 6

The load ratio on each thread of 1"-12UNF boit.
- 27 -

0.4

~
c9

o

~-~

~0 0.3~GL

16UNF Axisymmetrical model

16UNF 3-D model

co

c9

.~

o 0.2
'~

o
.~

o
c9
L

:,=

U
J

CDO 0.1

0.0

123

78

4
5
6
Thread number from nut flxed surface

Fig. 7 : The load ratio oon each thread of I " -16UNF bolt.

- 28 -

o. 4

Helical eFfect on 3-D analysis,~c9~2
Uo 0,3l~~-~ 8UNC~~ 12UNF~- 16UNF

c9

J~

t'o0,2
o
1.

o

:5:5
c9L

U
g 0,1
J
0.0

67

Thread number from nut fixed surface
Fig. 8The contact force ratio on each thread of three typicalbolt model with three dimensional analysis- 29 -

Session II

Structural Integrity and Life Prediction

II- lProbabilistic Fracture Mechanics of Nuclear Structural Components
Genki Yagawa, University of Tokyo, JapanPROBABILISTIC FRACTURE MECHANICS OF NUCLEARSTRUCTURAL COMPONENTS = CONSIDERATION OF
TRANSITION FROM EMBEDDED CRACK TO SURFACE CRA
Genki Yagawal), Yasuhrro Kant02) and Shinobu Yoshimural)l ) School of Engineering, University of Tokyo,

7-3-1 Hongo, Bunkyo, Tokyo I 13-0033, Japan
2) Department of Mechanical Engineering,
Toyohashi University of Science and Technology,
Tempaku, Toyohashi 441-8122, Japan

ABSTRACTThis paper describes a probabilistic fracture mechanics (PFM) analysis ofaged nuclear reactor pressure vessel (RPV) material. New interpolation
formulas are first derived for both embedded elliptical surface cracks andsemi-elliptical surface cracks. To investigate effects of transition fromembedded crack to surface crack in PFM analyses, one of PFM
round-robin problems set by JSME-RC111 commrttee I e "aged RPV
under normal and upset operating conditions" is solved, employing theinterpolation formulas .

1 . INTRODUCTION
In design and integrity evaluation processes for nuclear structuralcomponents, deterministic approaches have mainly been employed. All kinds of
uncertainty related to operating history, material property change and damage
mechanisms are taken into account in so-called safety factors. It is easily expectedthat the results obtained are too conservative to perform a rational evaluation ofplant safety and to make judgement of life extension because of the accumulationof conservatisms of all related factors.

Probabilistic Fracture Mechanics (PFM) has become an important tool [1, 2] .

The PFM approach is regarded as an appropriate method in rationally evaluating
plant life and in risk-based decision making such as risk-informed inspection [3]since it can consider various uncertainties such as sizes and distributions of cracks,

degradation of material strength due to aging effects, fluctuating loading histories,

accuracy and frequency of pre- and in-service inspections. Thus, various PFM-31-

computer programs have been developed and applied in practical situations in thelast two decades.

In Japan, one research activity on PFM approaches to the integrity studies ofnuclear pressure vessels and piping (PV&P) was initiated in 1 987 by the LE-PFM

subcommittee organized within the Japan Welding Engineering Society (JWES)
under a subcontract of the Japan Atomic Energy Research Institute (JAERI), andcontinued for three years [4]. The activity was followed by the RC I I I researchcommittee organized in the Japan Society of Mechanical Engineers (JSME) in
1991, and fmished in March, 1995 [5]. Succeeding it, a new PFM subcommitteeorganized in JWES started again in May, 1996 [6], and is scheduled to continue
until March, 1999. The purpose of the continuous activity is to establish standard

procedures for evaluating failure probabilities of Japanese nuclear structuralcomponents such as PV&P and stearn generator tube, combining the state-of-theart knowledge on structural integrity of nuclear structural components andmodern computer technology such as parallel processing. Within the LE-PFM
and JSME-RC I I I activities, we have set up the following three kinds of PFM
roud-robin problems on(a) primary piping under normal operating conditions,(b) aged reactor pressure vessel (RPV) under normal & upset operating
conditions, and(c) aged RPV under pressurized thennal shock (PTS) events.

The basic part of the last PTS problems is taken from some of US benchmark
problems [7, 8]. For these round-robin problems, various sensitivity analyses
were performed to quantify effects of uncertainty of data on failure probabilities.

Some of the detailed analysis results can be found in elsewhere [9-15].

In those analyses, we always assumed semi-elliptical surface cracks as initial

cracks for the purpose of convenience. However, as well known, embedded
cracks seem more practical and probable as initial cracks. Thus as one of maintasks in the JWES-PFM subcommittee, we are developing a new PFM model
considering transition from embedded to surface cracks during crack growth.
This paper describes its latest results. Here new interpolation forrnulas are frrst
derived for both embedded elliptical cracks and semi-elliptical surface cracks. To
investigate effects of transition from embedded to surface cracks in PFM analyses,

one of PFM round-robin problems set by JSME-RC111 committee, i.e. "aged
RPV under normal and upset operating conditions" is solved employing theinterpolation formulas.

- 32 -

2. OUTLINE OF PFM ANALYSIS
Figure I illustrates the flow of a typical PFM analysis. Firstly, somerandom variables are selected according to an analysis model employed. Randomvariables to be considered include initial crack sizes (crack depth and crack aspectratio) and location, accuracy and frequency of nondestructive tests, i.e. PSI and
ISI, material properties, cycles and amplitudes of applied loads. Next, crack

growth simulations are performed. Fracture mechanics models employed are
based on the linear elastic fracture mechanics, i.e. Newman-Raju solutions [16],influence functions [ 1 7-20] and others [2l], and the nonlinear fracture mechanics,
i.e. fully plastic solutions [22, 23] or their combination [24, 25]. Besides, creep
crack growth can be simulated based on the nonlinear fracture mechanics [26, 27].

During the crack growth simulation, PSI and ISI are considered, and failure
judgements of leakage and break are performed. Cumulative failure probabilities
are calculated as functions of operation time in round-robin problems (a) and (b),while failure probabilities for one PTS event are calculated in round-robin
problem (c). To effectively calculate accurate failure probabilities, the Stratified

sampling Monte Carlo (SMC) algorithm is offen employed. Its parallel version is
also very efficient [1 I].

3. INTERPOLATION FORMULAS OF STRESS INTENSITY FACTORS FO
EMBEDDED ELLIPTICAL CRACK & SEMI-ELLIPTICAL SURFACECRACK
3.1 Stress Intensity Factors for Embedded Elliptical Cracks

Soneda and Fujioka calculated stress intensity factors (SIFs) of various
elliptical cracks embedded in thick-wall cylinder subjected to arbitrarily
distributed loadings, and constructed an influence function database [28]. They
employed the finite element alternating method developed by Nishioka and Atluri[29]. The analysis process is schematically illustrated in Figure 2 andsummarized as follows :l) Solve the uncracked fmite body under the given external loads by using
the finite element method. The uncracked body has the same geometry
with the given problem except the crack.
2) Compute the stresses a.,.,k at the location of the original crack, using thefinite element solution.

3) Compare the residual stresses a.,,,k calculated in 2) with a permissible

stress magnitude. In the present study, 0.1 % of the initial a value
was used for the permissible stress magnitude.
4) Calculate theoretically the SIF, KI(i)i~fi~it*- 33 -.***kof the crack in infinite body,applying the stresses a,,*,k Onto the crack surface.

5) Solve the cracked infmite body, applying the reversed stresses - aonto the crack surface.
6) Compute the stresses a*uf*** at the boundary of the original finite body,****kusing the finite element solution.

7) Solve the uncracked body under the 0=*ta,* onto its boundary by using thefinite element method.
Repeat 2)-7) steps in the iteration process until the residual stresses a**k On thecrack surface become negligible (Step 3). To obtain the final SIF solution KI ' add
the stress intensity factors of all iterations as :
( i)

i=0

where n is the number of iterations.

Here we consider only the axial elliptical crack embedded in cylinder as in
Figure 3. Following a typical influence function method, a stress distribution in

the thikness direction of the cylinder is first approximated by 4th-orderpolynomial in a normalized coordinate. Then the SIF values are expressed withthe coefficients Ai Of the polynomial as follows := (a~sin2eKl(e)S + a~cos2e '4~ Ai=0Gl
l/ swhere Gi is the influence function for thesquare of elliptic integral of second kind

(2)

j-th

order stress distribution, Q is theand is approximated by the following

equatron :Q = I + 1.464 (a2 / al)1'65(3)

where al and a2 are half lengths of the crack in the axial and thickness directions,

respectively. Gi Values are tabulated in Ref. [28]. In the present study, Gi Vaiues
are evaluated at the three points of elliptic angle 0=7c/2, O and -1c/2. In the threecases, (*)1/4 term in Eq. (2) reduces to (al)1/4, (a2)1/4 and (al)1/4, respectively.

When only tensile stress o:T and bending stress aB are considered, thecoefficients Ai in Eq. (2) are simplified as follows :Ao = OT ~ aB

Al = 20B(4)

Ak = O, for k ~: 2

- 34 -

Table I shows parameter ranges for which influence functions Gi are given
those mfluence functron soluttons in thein Ref. [28]. We then interpolate
follOwing.
Gi = ~ ~ ~ ci=Pq,~Pnq~(5)

p=0 q=0 '=0

where

_ ert
)[( + ~~(ett)~'~
y
e/t
~=
e/t
j (e/t).

( -~" ~."

.2

e/t -

0.90 - 2adt

_ I-((-ao/a)I I2y mm '2
a'/an=a /al max
(2) 2-a /al)mln+ (a /a )aa/a I - 2/3

max

~=

.y
. - [2a~t) '~~ ~'"~~*(2a~ t)

~'"

(6b)

2/3

(

) + rL2a~t) ' 2
2adt - [[2adt( 6 a)

l 12

2ajt - 0.4
0.6

(6c)

In the present study, only Gi Values of axial elliptical cracks for linear
components of stress distributions are interpolated, although more other solutions
are given in Ref. [28]. The coefficients ci;pq' in Eq. (5) are tabulated in Table 2.

Figures 4(a) and 4(b) show some comparison of original and interpolated Gi
values. These figures confirm that Eq. (5) provides accurate interpolation for all
the solutions. It should be also noted here that, as shown in Figure 4(b), Eq. (5)exrapolates in the range of a2/al=0- 1/3. Its validity in the range is not proven yet.

3.2 Stress Intensity Factors for Semi-elliptical Surface Cracks

In PFM analyses considering the embedded elliptical cracks as described in
3.1, one can take into account uncracked ligament as small as 0.05t, where t iscylinder thickness. On the other hand, original finite element K solutions of

Newman-Raju [16], which have been employed in the PFM round-robin problems(a) and (b) are obtained till crack depth a =0.8t, i.e. uncracked ligament of 0.2t.

Miyoshi et al. performed precise finite element analyses for much deeper cracks
than a=0.9t [30]. To make K solutions for surface cracks consistant with those for
embedded cracks, a new interpolation formula is also derived in this study using
both Newman-Raju's finite element solutions [16] and Miyoshi's [30].
Fundamental policy for this task is as follows :1) Since K solutions increase rapidly over a/t > 0.95, the presentinterpolation certifies its validity until a/t=0.95.
- 35 -

3) The interpolation is performed only for K solutions at elliptic angle ofc =0 and lcl2.

The interpolation formula obtained is expressed as follows :+ ab)~~~rFl
KI at=((7)

where
Q - I o + 1 464 a/c 165( 8 a)

Fl = ~ ~ G'j~i(~)j(8b)- . . ( J.
i=0j=0(8c)~=1-fr~~T

where Gij are given in Table 3. The validity range of this interpolation issummarized in Table 4. Beyond the validity range in Table 4, Newman-Raju's
equation is adopted except a/t > 0.95. Figures 4(a) and 4(b) show the comparison
between finite element solutions and the interpolation of Eq. (7) for the purposeof convenience.

4. ANALYSIS PROBLEM = AGED RPV UNDER NORMAL AND UPSETOPERATING CONDITIONA PFM model analyzed here is basically the same as the PFM round-robin
problem for an aged RPV under normal and upset operating condition, which wasgiven by JSME-RC 1 1 1 committee [5] except consideration of embedded cracks.

4.1 Analysis Models
Two kinds of cracked models are assumed here. The one is a plate with asemi-elliptical surface crack, while the other is a plate with an axial embedded
elliptical crack. These are illustrated in Figures 6(a) and 6(b). The embedded
elliptical crack may be converted to the semi-elliptical surface crack, according toa certain conversion criterion as described later. In both models, plate thickness

and width are taken to be t = 0.2m and 2b = 12.6m, considering the beltline
portion of PWR pressure vessels. The plates are assumed to be subjected tovarious magnitudes of remote uniform tensile and bending stresses. For the
purpose of simplicity, we ignore curvature effects of actual pressure vessels in
evaluating 3-dimensional SIFs. Cumulative failure probabilities of one existing
crack, whose unit is l/crack, are calculated as functions of operation years.
- 36 -

4.2 Probabilistic Density Distributions for Ssemi-elliptical Surface Crack
As for the semi-elliptical surface crack, an initial crack depth (a) and a crack
aspect ratio (P=c/a) are assumed to be random variables. As probabilistic density

functions, we employ the Marshall distribution for (a) [3l], and a log-normal
distribution for (P) [32] as follows :exp --11a//[1~xp ( _I/ ~jjp(p)a {Inexp[p,'fu~)}c (9b)
p yl~~
ap
2Y

where oc = 1.035, ~= 1 336 Y O 5382 u 6 248 x 10 m and t O 2 m
4.3 Probabilistic Density Distributions for Embedded elliptical crack

As for embedded elliptical cracks, half length of longer axis (al)' that ofshorter axis (a2) and depth (e) are assumed to be random variables. The
probabilistic density distributions of al and p = al/a2 are simply taken to be thesame as Eqs. (9a) and (9b), respectively. (e) is assumed to be a uniform randomvariable. It should be noted here that this assumption might lead larger averagecrack sizes for embedded cracks.

4.4 Conversion Criterion from Embedded Crack to Surface Crack
When uncracked ligament of the embedded elliptical crack reaches 0.05t, theembedded crack (al' a2, e) is converted to a semi-elliptical surface crack (a, c) by
the replacement of al ~ c and depth a2 + e ~> a.

4.5 Cyclic Loads
Nineteen kinds of cyclic tensile and bending stresses given in Table 5 arechosen from the design loading conditions of Level A (normal operation
condition) and Level B (upset condition) which occur in the beltline portion of
PWR pressure vessels listed in the Marshall report [3l]. The nineteen loads areapplied in the order as listed there.

4.6 Failure Criteria
The failure modes considered here are leakage and break, whose criteria aresimply defined as follows :~***- tc - 37 -

Leakage criterion :a / t ~ 0.95

( I Ob)where I(~*. is taken to be the largest K value along

crack front.

For simplicity,

break phenomenon after leakage is not considered.

4.7 Fatigue Crack Growth
Only fatigue crack growth based on the Paris' Iaw is assumed, whose
coefficients are taken from the fatigue crack growth rate of nuclear pressure
vessel steels in water given in the ASME Code Section XI, Appendix A [33] :da/dN (m/cycle) =1.738 x 10 ~13 (AK) 5.95
( I I a)da/dN (m/cycle) = 5.325 x 10~9 (AK) 1'95

( for AK ~ 13.2 MPa~~~: )(1 Ib)The elliptical or semi-elliptical cracks keep their shape during crack growth,
except the case that the convergence criteiron in 4.4 is satisfied. As for ellipticalcrack, an amount of crack growth per one loading cycle is evaluated in all longer
and shorter axis directions, independently.

4.8 Fracture Toughness
Referring to the reduction of the upper shelf Charpy absorbed energy due toneutron irradiation [34] and some experimental data conducted by the JWES-LE
subcommittee [4], the time variation of KI* at 300 "C due to neutron irradiation isformulated as follows :300 'C :Ktc (t) =135.0 ( MPa~fl )(for F ~ 0.361)( 1 2)= 3.29 + l 18.71 x F ~o.l02(for F > 0.361)

where F is a neutron fluence (l019 n/cm2 ). The neutron fluence after a 40 year
operation is assumed to be 3 x I 019 n/cm2 according to that of the beltline portionof the PWR pressure vessels in the Marshall report [3l]. According to Eq. (12),the Ktc value reduces by about 19 (~o due to neutron irradiation during a 40 year
operatron.
In reality, since the neutron fluence varies along the thickness direction, theneutron fluence F is formulated as a function of the distance d(mm) from theinside surface of plate assumming Fo to be the neutron fluence at the insidesurface as [8] :- 38 -

F = Fo x exp (- 0.00945 x d)

According to Eq. (13), the neutron(13)
fluence at the outside surface is about 1 5 % ofthat of the inside surface.

5. RESULTS AND DISCUSSIONS
Figure 7 shows time variations of cumulative break and leakage probabilities

considering embedded elliptical surface cracks. The total number of samples is
168,384. Leakage probability is by two orders of magnitude smaller than break
probability. This means that before reaching leakage criterion of a/t = 0.95, most
samples satisfy the Kb criterion. Leakage probability is almost constant, while thebreak probability increases gradually. In the case of considering only surface
cracks, the break probability was in the order of 10~8 , and the leakage probability
12
was in the order of 10~ [13].
The present analysis results are one to two orders
greater than such a surface crack case. This seems strange, compared with ourintuition such that embedded cracks might have more resistance to failure thansurface cracks. As noted in 4.3, the probabilistic density distribution of initial
crack size for the embedded cracks was determined referring to that of surface
cracks, i.e. the Marshall distribution. According to some preliminary evaluation[5], this means that the former analysis case employs the initial distribution of
crack depth with three times larger average value than the surface crack case.
This may be the main reason of higher failure probabilities in the embedded crack
analysis case. More rational probabilistic density distributions for embedded

cracks should be investigated in the further study.

6. CONCLUSION
In the present study, we investigated effects of consideration of embeded
cracks in PFM analyses. At first, we obtained a new SIF interpolation forrnula
for elliptical crack embedded in cylinder, and also newly derived that forsemi-elliptical surface crack which is valid for deeper crack with a/t =0.95.

Implementing those formulas, we analyzed one of the PFM round-robin
problems, i.e. aged RPV under normal and upset operating condition. Leakage
probability is by two orders of magnitude smaller than break probability. The
leakage probability is almost constant, while the break probability increases
gradually. The present analysis results are one to two orders greater than thesurface crack case analyzed in the previous study. This result may be caused dueto the fact that an average crack size of embedded crack tends to be greater thanthat of the surface crack when the same probabilistic distribution of crack size is- 39 -

employed for both embedded crack and surface crack. More rational
probabilistic density distributions for embedded cracks should be investigated in
the further study.

ACKNOWLEDGEMENTS
The authors wish to thank all the members of the PFM subcommittee in theJapan Welding Engineering Society (JWES) for their efforts to perform thiswork, and also thank the Japan Atomic Energy Research Institute (JAERI) for
financial support.

REFERENCES[1] G. Yagawa, Structural integrity assessment of nuclear power plants using
probabilistic fracture mechanics, Nuclear Eng., 34 (1988) 19-30 (in Japanese).
[2] D. O. Harris and K. R. Balkey, Probabilistic considerations in life extension

and aging, Technology for the '90s, ASME PVP Division, pp.243-269, (1993).

[3] K.R. Balkey, N.B. Closky and M.K. Vermaut, Application of risk-based
methods to inservice inspection and testing, Trans. 14th SMiRT, Lyon, M05/2,( 1 997).

[4] JWES, Report of LE Subcommittee, Japan Welding Eng. Society (JWES),
JWES-AE-9003, (1990) (in Japanese).
[5] Final Report of RC111 Committee, Japan Society of Mechanical Engineers(JSME), (1995).

[6] JWES Report of PFM Subcommittee, JWES, JWES-AE-9705, (1997) (inJapanese) .

[7] D. L. Selby et al., PTS evaluation of the H.B. Robinson Unit 2 nuclear power

plant, NUREG/CR-4183, Vols. l, 2, (1985).
[8] B. A. Bishop, Benchmarking of PFM analyses of reactor vessels subjected toPTS Ioading, EPRI Research Project 2975-5, Final Report, (1993).
[9] S. Yoshimura, G. Yagawa et al., PFM analyses of LWR's pressure vessels
and piping : Quantitative study on influences of distributions of initial crack sizes,

Proc. 20th MPA Seminar, Vol.2, pp.44.1-10, (1994).

[10] G. Yagawa, S. Yoshimura, et al., Study on life extension of aged RPVmaterial based on PFM -Japanese round robin, Trans. ASME, J. Pressure Vessel
Technology, 1 17(1995) 7-13.

[1l] S. Yoshimura, M.-Y. Zhang and G. Yagawa, Life extension simulation ofaged reactor pressure vessel material using PFM analysis on a massively parallel

computer, Nuclear Eng. & Design, 158 (1995) 341-350.
[12] G. Yagawa, S. Yoshimura et al., PFM anaiyses of nuclear pressure vessels
under PTS events, Nuclear Eng. & Design, 174 (1997) 91-lOO.
- 40 -

[13] G. Yagawa and S. Yoshimura, A study on PFM for nuclear pressure vessels
and piping, Int. J. Pressure Vessels & Piping, 73 (1997) 97-l07.

[14] S. Yoshimura, G. Yagawa et al., PFM analysis for LBB evaluation of Light

Water Reactor's piping, J. of Atomic Energy Society of Japan, 39 (1997)
777-787.
[15] G. Yagawa and S. Yoshimura, Probabilistic fracture mechanics analyses ofnuclear pressure vessels and piping : A review of Japanese activity, Trans. 14th

SMiRT, Lyon, GMW/1, pp.541-552, (1997).
[16] J. C. Newman, Jr. and L. S. Raju, Stress intensity factor equations for cracks

in 3D finite bodies subjected to tension and bending loads, NASA -TM- 85793,
1984.
[17] I. S. Raju and J. C. Newman, Stress intensity factors for internal andexternal surface cracks in cylindrical vessels, Trans. ASME, J. Pressure Vessel
Technology, 104 (1982) 293-298.
[18] D. L. Stevens et al., VISA, NUREG/CR-3384, (1983).
[19] F. A. Simonen et al.,VISA-II, NUREG/CR-4486, PNL-5775 RF, R5, (1986).
[20] M. Shiratori, Analysis of stress intensity factors for surface cracks subjected

to arbitrarily distributed stresses, Bulletin of Faculty of Eng., Yokohama National

University, 35 (1986) 1-25.

[2l] S. Yoshimura, J.-S. Lee and G. Yagawa, Automated system for analyzing
stress intensity factors of 3D cracks : Its applications to analyses of two dissimilarsemi-elliptical surface cracks in plate, Trans. ASME, J. Pressure Vessel
Technology, 1 19 (1996) 18-26.

[22] G. Yagawa, Y. Kitajima and H. Ueda, 3D fully plastic solutions forsemi-elliptical surface cracks, Int. J. Pressure Vessels & Piping, 53 (1993)
457-510.
[23] G. Yagawa, Y. Kitajima and H. Ueda, Erratum of [22], Int. J. Pressure
Vessels & Piping, 74 (1997) 77-80.

[24] G.-W. Ye, G. Yagawa and S. Yoshimura, PFM analysis based on
3DJ-integral database, Eng. Fracture Mechanics, 44 (1993) 887-893.

[25] G. Yagawa and G.-W. Ye, A PFM analysis for cracked pipe using 3-Dmodel, Reliability Eng. & System Safety, 41 (1993) 189-196.

[26] G. Yagawa, S. Yoshimura and H. Ueda, An estimation scheme for J- andC*- integrals of surface cracks, Proc. Int. Symp. Pressure Vessel Technology &Nuclear Codes & Standards, Seoul, pp.8. 15-8.22, (1989).

[27] K. Hojo, G. Yagawa, S. Yoshimura et al., Applications of PFM to FBR
components, Nuclear Eng. & Design, 42 (1993) 43-49.
[28] N. Soneda and T. Fujioka, Stress intensity factor influence coefficients ofembedded cracks in cylinder, Komae Research Laboratory Report of Central
Research Institute of Electric Power Industry (CRIEPI), No. T94037, (1995).
[29] T. Nishioka and S.N. Atluri, Analytical solution for embedded elliptical

-41-

\

cracks and finite element alternating method for elliptical surface cracks subjected

to arbitrary loadings. Eng. Fracture Mechanics, 17 (1983) 247-268.
[30] T. Miyoshi, K. Ishii and Y. Yoshida, Database of stress intensity factors forsurface cracks in prelpost penetration, Trans. JSME, 56A (1990) 1563-1569.

[3l] W. Marshall, An assessment of the integrity of PWR PVs. UKAEA, (1982).
[32] T. Y. Lo et al., Probability of pipe failure of the reactor coolant loop of

Combustion Eng. PWR plants, NUREG/CR-3663, UCRL-53500, (1985).
[33] ASME, ASME Boiler & Pressure Vessel Code. Section XI, Appendix A,(1973).
[34] P. Tipping, Study of a possible plant life extension by intermediate annealingof the reactor pressure vessel 11 : Confirmation by further annealing schedules,

Proc.NEA-UNIPEDE Specialist Meeting on Life-Limiting and Regulatory
Aspects of Reactor Core Internals and Pressure Vessels, Sweden, ( 1 987).

Table I Ranges of parameters for SIF
calculations of embedded elliptical cracks
a2/al

2a2/t

e/ t

O. 1

0.1, 0.3, 0.5, 0.7, 0.9

0.3

0.2, 0.35, 0.5, 0.65, 0.8

0.5

0.3, 0.4, 0.5, 0.6, 0.7

0.7

0.4, 0.45, 0.5, 0.55, 0.6

1, 2/3, 1/2, 1/3

Input of basic data

• Initial crack size

• Pre-service inspection

~

• Ktc

Sampling of probabillstic variables• Cyclic loads...

• Fully plastic so]utions• In-service inspection

Crack growth simulation
during operation• Newman-Raju's solutions

Judgement of leakage and break probabilities~
Evaluation of leakage and break probabi]itles~
r~~~TloutputFig. 1

Flow of typical PFM analysis- 42 -

Table 2 Coefficients of interpolation functions ofinfluence coefficients for axial embedded cracks

Table 3 Coefficients for interpolation
for semi-elliptical surface cracks- 43 -

Table 4 Validity range of interpolationt~!

t-- -*----**---------1~
1

,

t
N
N

,

bt ~~
T~i~lf:
~ l
l

,

,
I

,

I

,

l

t

l Vtt ~b
N,

I

Uncracked Finbe Body,

1 Cracked Infintte Bodyb~t~.- -.'~~~~~

hera tio,'

i~
~~
i~
i~

K/ i ==o ~K/ (i)~ infinile: number of iterations

n

K, .
(f)

KI-i,,fi,,ile :Suess-intensity factor of finite bodyStress-intensity factor of infinite bodyat i-th iteration

Uncracked Finae Bod),tVVVVFig. 2 Algorithm of finite element alternating method- 44 -

2a2-~ i~~~ l
2a l

'l

-~
e

t

Fig. 3 Axial elliptical crack embedded in cylinder

Table 5 Loading conditions- 45 -

1 .2

C"

H:h- -H~:
e•~"
•t~i

- CI--H:H--H:l1a2/al = 1 2a2/t O 1

0.8

' $'

Uniform, +7T:/2

Unifonn, O(A::;••t'

~
C:

,t4

(L)

,.

'C;

~O

1-11'd-Cha'l-, *,*~

ldF .L

0.6

1,, t'---O---

Uniform, -7t/2

ll~l

Linear, +7c/2'f~3 tcl'] I'l Is'at

Linear, O

eP'1'1 I11'1

Linear, -7cl2-' ~ ~D- -"-Uniform, +7d2

'I'-1C~-

Uniform, O--'1-0--'-'

Uniform, -7d2~l-Linear, +7~/2"t' ", I~l,,1,I'I

Linear, O

Ic

Linear, -rd2

a;••"

O
(J

. ~*

0.4

.S.:••"

0.2

J(. ~•/'..*'

O

0.4 0.6 0.8

0.2

o

1

Depth, e/t

Fig. 4(a) Gi VS. e/t

1 .2
j I-_'i~_..;-._:_T; ~I:'~ 1lb 'I"_' :":1-''~Lg: ='1'--"'

2a2/t = O t5' e/t = O 25

0.8
c,D

~
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2.5

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3

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Fig 5(d) Fl vs a/c (a/t = 0.8)

- 47 -

M: bending moment
Gb = 3M / bt 2 : bending stressot : tensile stress

M (;t : tensile stress

M

M

t

M

M

Fig. 6(a) Semi-elliptical surfacetFig. 6(b) Elliptical crack embedded

crack in platein plate

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Fig. 7 Failure probabilities vs.
- 48 -

30

40

operation years

II-2

Remaining Life Prediction Method Using Operating Data andKnowledge on Mechanisms
Bom Soon Lee, Materials and Corrosion Research Laboratory,
Korea Electric Power Research Institute, Korea Electric Power

Corporation, Korea

Remaining Life Ptediction Methods Using Operating Data

and Knowledge on Mechanisms*

Bom Soon Lee Han Sub Chung, and Ki-Tae l~mMaterials and Corrosion Research Labcratory
Korea Electric Power Research Institute

Korea Electric Power Corporation, Korea

ABSTRACT
It is very important to bc able to predict the remaining life of components andstructures in a power plant, both for nuclear and fossil units. The information needed
can be obtained from the controlled laboratory experiments and the plant operating data.

On materials degradation, we have accumulated a large amount of data from bcth
sources. However, it is essential to formulate the best methodology to utilize theseinformation so that our needs can be met. In this paper, the methods currently used forremaining life prediction are discussed with typical results. Also discussed are thelirnitations and the benefits of different approaches along with suggestions for the future

R&D drections in this area.

INTRODUCTION
There are 478 nuclear power plants operating or being built in 32 countries
around the world supplying about 17% of the global need for electricity [Ref. I]. As
the nuclear power plants get'~l~er, aging bccomes an issue, because aging degradation
can affect the structural soundness , of systems and components. All the nuclear plants
around the world should face this )issue sooner or later. For instance, United States

have the largest number of nuclear power plants in the world, and by the year 2014, 48
commercial nuclear power plants in the United States are projected to reach 40 years ofoperation [Ref. 2] . The distribution by plant age for the operable USA plants is shownin Figure I . In managing aging of these plants, good maintenance practices including
corrective and predictive ones are required. At the same time, it is very important andnecessary to be able to predict the remaining life of a component so that a rightdecision can be made, in time, among the possible choices of taking no action, repair or

replacement. Also, the remaining life estimation will play an important role in the life
extension and license renewal processes for the existing nuclear plants.

* To be presented at the International Workshop on Integrity of Nuclear Components,
Tokyo, Japan (April, 1998)- 49 -

When a new system or component is designed, the candidate materials aretested under simulated environments so that the right material is selected to assure thatthe component will last the design life. In this process, the hrst step is identifying theimportant parameters that affects the material's perforrnance. Usually, temperature,stress levels, and other environmental parameters are identified in this process. The

next step is conducting accelerated tests varying only one parameter at a time while

keeping others constant. This experimental approach provides knowledge on the
degradation mechanisms even though it is not always successful to understand themwith a high level of confidence. This is especially true when multiple parameters are
involved .

The predictions based on experimental mechanistic approaches have failed inmany cases with the most notable example of Alloy 600, which was believed to last
well past the 40 years of design life as a steam generator tube material. During actual

operation of nuclear plants, it was discovered that this material was showing several
different types of failures modes. The reason for this failure to predict the performanceof this material was the lack of knowledge on all the parameters involved in the
degradation. Also, the synergetic effects between different parameters are difficult totest in the laboratory.

The operation of a nuclear plant can bc considered as a real time expenment
under realistic conditions. In fact, we should treat all the nuclear plant operating data

as real time (not accelerated) experimental data because we can pull out treasures offailure infonnation from these operating data. If we use these data effectively, we canget complementary information to the experimental results. In this paper, someexamples of utilizing operating data are given, and the interrelation between theoperating data and experimental data are discussed in the case of steam generator forpressurized water reactors (PWRs).
In utilizing the operating data, we encounter different limitations in predicting

remaining life depending on the type of a component, whether it is a component withseveral subcomponents or rather a simple one. In this paper, to illustrate thedifferences, globe valves in the containment systems for the Westinghouse PWRs and asteam generator are selected as examples. Globe valves have multi-subcomponents andfailure of each different subcomponents cause the failure of the valves. On the other

hand, a steam generator has many subeomponents as shown in Figure 2, but thedegradation of the tubes is always the life limiting factor for a steam generator. The
methods we use to utilize the plant operating data to understand the failure causes andmechanisms are discussed with these two different components.

- 50 -

AGlNG RESEARCH UTILIZlNG OPERATlNG DATAThe United States Nuclear Regulatory Comrnission (US NRO, Office of Nuclear

Regulatory Research has conducted the most extensive research on aging on systems,
structures and components (SSCs) of nuclear power plants. This prograrn, entitled"Nuclear Plant Aging Research (NPAR), seeks to improve the operational readiness of

systems and components that are vital to nuclear power plants and their safety byunderstanding and managing aging degradation. Since its inception, the NPAR program

has produced a wealth of knowledge on the aging of 19 systems and 29 components.
The following were the technical objectives of the program [Ref. 3] :O

identify and characterize aging effects which, if unmitigated, could cause
degradation of SSCs and thereby impak safety,Odevelop supporting data and information to facilitate management of age-relateddegradation,oidentify methods of inspecting, surveillance, and monitoring of SSCs, or ofevaluating their remaining life to ensure the timely detection of signiflcant
effects before safety function is lostoevaluate the effectiveness of storage, maintenance, repair, and replacement
practices in mitigating the effects of aging and dirninishing the rate and extent
of aging degradation,oprovide technical bases and support for the License Renewal Rule and thelicense renewal process, and develop a regulatory guide on the format andtechnical infonnation content for renewal applications .

Usually, for these aging studies, a detailed analysis was perfonned of thefollowing data bases summarizing the actual operating experience of SSCs:

Nuclear Plant Reliability Data System (NPRDS)

Licensee Event Reports (LERS)
Plant Specific Failure Data.

Analyzing these plant operating data provides information on degradation processes,

trend, failure modes, failure causes, and falure mechanisms. The main advantages ofusing the plant operating data are the large number of data and the fact that these are
realistic data. However, since these data are not from the controlled environment, it is
not usually possible to identify the exact failure mechanisms. Also, extra caution should- 51 -

be used in analyzing the data to obtain valid results. For instance, when analyzing
valve failure data, sirnilar types of valves should be grouped together.

Figure 3 shows the normalized aging-related failure frequency as a function ofage at failure for the globe valves in Westinghouse PWR Containment Isolation function[Ref. 4] . This kind of plot provides information on trend of failure rate as a function ofage, but this information cannot be used to make decision on when to replace all thevalves. The analysis of faiilure data provides how and why the valves fail as shown in

Figure 4 and 5, respectively. Figure 5 shows that the failures caused by "wornpacking " and "worn gasket" mcreased during the later years, which explains themcrease of the failure mode "external leak" shown in Figure 4. Worn or outof-adjustment packing and worn or damaged gasket result m external leak The failure
due to "aged seat/disk" also increased as the valves got older, which may explain theincrease of the failure mode "internal leak".

To obtain more detailed information on the effects of agmg on failure causes

the number of failures caused by three major causes were normalizod and plotted as afunction of age at failure (Figure 6). The frequency curve for the failures caused byworn seats/discs has two peaks at 12 years and 21 years, which explains the peaks at
the same ages in the plot for globe valve failures (Figure 3). The failure caused bycorrosion product/dirt buildup has a high peak at 3 years, after which the frequency

stayed low until it began to increase at 13 years. At 16 to 18 years, many valve
failures were caused by worn packing, which explains the peak in the frequency of
globe valve failure at 16 - 18 years (Figure 3).

As shown m this example rt rs not easy to predict the remaining life of acomponent which has many subcomponents. However, this kind of operating data
analysis provides an important information on when to conduct preventive maintenanceto prevent failures. In this example, it was shown that many reported failures werecaused by regular maintenance items, such as valve packings. These aging-related
failures can be, and should be, prevented by conducting proper scheduled maintenances,which utilizes operating data.

Prediction of Remaining Life of A Steam Generator
Korea is in the process of replacing steam generators (SGs) at its oldest
nuclear plant, Kori 1, which went into commerciai operation in 1978. The tubes ofthese SGS are mill annealed alloy 600, and similar tubes world wide suffered fromseveral forms of degradation over the years. By changing the water chemistry andextensive effort to limit the ingress of aggressive agents, most of the old problems were

- 52 -

solved except the intergranular attack/stress conosion cracking (IGA/SCO.
Degradation of heat transfer tubes is the main issue for the nuclear steam generator life
prediction.

As discussed above, several different types of corrosion cause the degradationand failures of steam generator tubes, and both statistical and mechanistic approaches
are required in order to predict rernaining life of a steam generator with a high level ofconfidence. Which one is more useful depends on type of degradation, extent ofexperimental and operational data available, and the extent of mechanistic understanding
for the relevant degradation type.

Degradation of steam generator tubes can be considered as an ideal subj ect for
statistical evaluation. Thousands of identical tubes in a steam generator are exposed tosame operation condition so that degradation of each tube can be treated as randomvariable in a statistical sense for a specific type of degradation. In service inspection(ISD produces tube degradation data during each outage. It is not unusual that asignificant amount of defected tubes are identfied at every inspection at the scheduled
outage.

Nevertheless, it should be noted that the extent of applicability of statistical

approach be limited and be reviewed with care. There have been series of steam
generator design improvement by different vendors, and many different models are underoperation presently. Any steam generators of a model may not be identical to one
another bccause details of operational conditions, especially water chemistry, vary fromplant to plant. Once a life threatening degradation form is identified, each plant start tostruggle to control and mitigate the degradation by corrective, preventive, and predictive

maintenance efforts, which include for example chemical cleaning, shot peening, corrosion
inhibitor, and sludge lancing . Even without the maintenance effort, a steam generator ishardly operated under an identical condition throughout the whole life. Secondary sidecorrosion of tubes attributed to occasional impurity intrusion into water chemistry is notuncommon. There have bcen significant improvement of non destructive examination(NDE) techniques for ISI. Improved ISI induces a transient of record by uncovering
defects not detected before. In a few instances, it was too late when significant amountof tube degradation data enough to define a statistic trend were available because steam

generator replacement or a maj or sleeve campaign was needed already.
Even though factors listed above lirnit the applicability of statistical approach,there are still cases where statistical analysis of the operating data plays a key roll insteam generator life prediction. Primary water stress corrosion crack (PWSCO in rolltransition and secondary side inter granular attack / inter granular stress corrosion

crack (IGA/IGSCO in a steam generator with a well controlled water chemistry history- 53 -

can be a good subject of statistical analysis. PWSCC occurs under a well controlled

chemistry condition, almost constant temperature, narrow range of residual stress
distribution caused dominantly by manufacturing tolerance. Secondary side IGA/IGSCC
is induced by localized impurity concentration inside crevices. The localized impurity

concentration may be considered as a part of aging process for steam generators
without serious chemistry transients .

Weibull Function
It has bcen verified that the Weibull probability distribution function (Ixif) is themost appropriate one for interpreting failure trends of steam generator tubes among
various pdf's [Ref. 5, 6]. The Weibull pdf is expressed as following:b-lf(t) = 6~(b t-toto e~
to 6~ to) exp[-( t~to )]b(1),where t ; operation timeto ; defect initiation timeb ; Weibull slope
6 ; characteristic time

f~can be rearranged as follows,
The cumulative Weibull function F(t) ( = f(x) dr)
1

In In = A + b In (t- t ) (2),where A = -bln(a-to)Eqation (1) and (2) denote three parameter Weibull function. If to is set to bezero, which is to say the degradation can occur at any time from the begiuning ofoperation, equation ( I ) and (2) can be modified to two parameter Weibull function.
Details of the Weibull function can be found elsewhere [Ref. 5, 6].

Figure 7 shows a Weibull proj ection of axial PWSCC in the roll transition near

top of tubesheet in three steam generators of one of Korean nuclear power plants. 100% of total 99eo tubes have been inspected every ISI since the fifth regular outage. A
tube was defined as cracked when the eddy current inspection detected crack indication,

whereas the crack initiation point may be defined as the time when a microscopicallysmall crack has grown to a size big enough to be detected by the current inspection

method. An aiternative tube repair criteria based on crack length has been adapted
rather than the depth based 40%; wall thickness criteria for this particular type of- 54 -

degradation in the unit. Statisticai approach can be a valuable tool when the remaining
life of these steam generators is to bc evaluated.

Crack growih rate data were required to evaluate the remaining life since crack
length based repair criteria were adapted. The plant specific crack growth rate could be

defined with a high level of confidence by using thousands of crack length datameasured by ISI. Each crack growih rate datum was defined as the difference between
crack lengths divided by length of operation period between successive ISls. Figure 8
shows another Weibull proj ection for which crack initiation point was defined when thecrack had grown longer than 5.5 mm. A two parameter Weibull proj ection fitted wellto the data set. Another projection line for tubes with cracks longer than 10 mm,
which is a tentative length criterion, could be drawn by using the plant specific averagecrack growih rate at the crack length of 5.5 mm. This projection contains conservatism

since crack growih rate decreases as a crack grows. The curve can be assumed as a
future proj ection of plug ratio when 10 mm length criterion is applied. The life of thesesteam generators can be estimated against the axial PWSCC in the roll transition whenthe maximum allowable plug ratio is defined in the unit.

MECHANISTICAL APPROACHES
While "statistical approaches" are used to estimate the probability of occurrenceof an event (e,g., number of SG tubes which has cracks longer than some criticallength) by simple extrapolation from credible existing data, "mechanistical approaches"are evaluating the parameters which can affect an event (e.g., crack growih rate in SG
tubes) individually and totally, and thereby, preferably setting up a proper model for the

event. By using this kind of model with the actual engineering conditions, one canevaluate or predict the possibility of an event.

One of the typical example of mechanistic approaches is the prediction of thestress corrosion crack growih rate of alloy 600 in PWR environment. Of the several

models suggested for the SC crack growih mechanism [Ref. 7], following threemechanisms are widely accepted: I ) slip / film rupture / dissolution mechanism, 2) film

induced cleavage, 3) hydrogen embrittlement. Different mechanisms , however, may
operate in one alloy / environment system with a dominant mode bcing determined byperhaps small changes in the material, environment, or stressing condition.

According to the slip / film rupture / dissolution mechanism, which is believed

most suitable to explain the SCC of SG tubes in PWR, the crack growih rate can be
estimated from the relation between the oxidation rates and the stress / strain
conditions at the crack tip. These are supported by a good correlation in a number ofsystems bctween the oxidation current density (e.g., obtained on a straining surface) andthe crack growth rate [Ref. 8. 9] .

- 55 -

For a given crack tip environment, corrosion potential, and material condition,

the crack growih will be controlled by the change in oxidation charge density with time

and the frequency of film rupture at the strained crack tip. As the crack grows, thebare new surface becomes passivated (oxidized). The next crack growih can happen
only when the passivated layer (oxide film) at the crack tip ruptures. Therefore, theaverage crack growih rate, Vt, can be related to the oxidation charge density passed
between film rupture events, Qf, and the strain rate at the crack tip, 8ct [Ref. 10]Vt--

M Qf 8ct(D

ZPF 8fwhere: M, p = atomic weight and density of the crack tip metal

F = Faraday's constantz = electrical valence of a metal atom
8f = fracture strain of the film

The oxidation current transient generally shows an initially high bare surface dissolution

current density, io, for a short time, to, and thereafter decreases according to a powerlaw relationship as the following:(2)

lt = io [t/to]~nwhere:

it = oxidation current at time tt=timen = constantBecause of this power law relationship, equation (1) can be reformulated as following

equation (3). This equation means that the crack growih rate behaviour can beevaluated by the repassivation response "n" value.
Vt = f(n) [8ct]~

(3)

,, ,,
The effect of n on
predicted crack growih rate vs. stress intensity isillustrated in Figure 9. There were much efforts to estimate or measure an appropnate,, n,, value under given temperature, environment, and alloy / microstructure conditions[Ref. 11 - 14]. Figure 10 shows that the range in crack growih data in PWR
environments are reasonably predicted by this modeling methodology.

- 56 -

Another example of mechanistic approaches is the evaluation of plugging criteriaof a SG tube to prevent tube burst. The plugging criteria of a tube can be fixed by amaxirnum permissible crack length, ao, whic.h can be obtained according to the following
equation:

ao = ele Aa Aa(4)

where: ac = the minimal critical size which causes a fast burst of the tube

Details of ac assessment can be found elsewhere [Ref. 15].

Aal = the maximum uncertainty from the dimensioning of thecracks of significant size by means of the inspection

technique used

Aa2 = crack growih between the two subsequent inspections
By using this criteria with the help of continuous in-service monitoring of primary to
secondary leak, a maximum permissible in-service leak criterion also can be set up.

As discussed above, the mechanistic approaches are very successful toformulate the working hypothesis and thereby quantify and validate against real
observation for some corrosion problems. However, the diversity of working parameters
and yet unknown synergistic effects between them limits the validity of the mechanistica pproaches .

CONCLUSIONS
Experimental data and plant operating data are the two main sources of
information that can be used to predict the remaining life of a component. However,
only recently, it has been realized that plant operating data can bc treated as long terrn,

real time experimental data except that the experimental conditions vary a few times
during the real time experiments. In this paper, a few examples of utilizing the bothtypes of data to predict the remaining life of a component were shown and discussed,
and the following conclusions may be drawn.

It is not easy to predict the remaining life of a component which has manyosubcomponents that fail independantly using the operaing data. However, the anaysis ofoperating data can provide an irnportant infonnation on when and how to conduct
preventive maintenance to prevent failures.

Weibull statistics is an usuful tool in predicting the remaining life of aocomponent utilizing the operating data as long as it can be assumed that the failure
mechanism stay the same.

The mechanistic approaches are necessary in formulating the workingohypothesis, which can be validated by experiments and in some c ases by utilizing the- 57 -

operating data.

More research should be focused on improving the methods to utilize theooperating data to predict the remaining life of a nuclear component, since plenty ofoperaing data are available, and are still being produced.

REFERENC E S

1. IAEA, International Atomic Energy Agency, IAEA-SM-353, 1998.
2. Blabnik. D.E., et al., "Insights Gained From Aging Research," NUREG/CR-5643,

BNL-NUREG 52323 March 1992
3. Vora, J.P., and Vagins, M., "Nuclear Plant Aging Research Program Plan,"
NUREG/CR-1144, Rev.2, 1991.

4. Lee B S et al "Agmg Assessment of Westinghouse PWR and General Electric

BWR Containment

Isolation

Functions," NUREG/CR-63339,

BNL-NUREG-52462, Brookhaven National Laboratory, March 1996.
5. EPRI, NP 7493, "Statistical Analysis of Steam Generator Tube Degradation,"
1991.

6. EPRI TR 103566 "Statrstical Analysrs of Steam Generator Tube Degradation:
Additional Topics," 1994.

7. Parkins, R., British Corr. J., 7, 15, 1972.

8. Ford, F.P., "Stress Corrosion Cracking", in Corrosion Processes,
Ed. R.N. Parkins, Applied Science, 1982.

9. Parkins, R., '~nvironment Sensitive Fracture - Controlling Parameters",
Proc., Third Int. Conf. on Mechanical Behaviour of Materials, Vol. 1,
p.139-164, 1980.

10. Ford, F.P. and Anderson, P.L., "Practical Application of Knowledge of
Mechanisms of EAC in LWRs", Proc., Int. Symp. on Steam Generator
Technology Update, KEPRI, Taejon, Korea, December 1997.
11. Thompson, C., et al, "Analysis of A600 and X-750 SCC Cracks", Proc.,
Alloy 600 Experts Meeting, Arlie, VA, EPRI, Palo Alto, April 6-9, 1993.

12. Thompson C et al, "Prediction of PWSCC in Nickel Base Alloys Using
Crack Growih Models" Proc Seventh Int Symp. on Environmental
Degradation of Materials in Nuclear Power Systems - Water Reactors,
NACE, p.867-880, 1995.

13. Carey D M et al, "Alloy 600 SCC Growih Measurements in Primary
Coolant" Prod Workshop on PWSCC of Alloy 600 m PWRs, EPRI,
November 1994.
14. Angeliu, T, et al, "Repassivation and Crack Propagation of Alloy 600 in
288C Water" Paper 96086 Corroslon/96, NACE, 1996.

15. Flesch, B., et al, 'Tailure Analysis Module for the EPM Code for PWR

Steam Generator Tube Mamtenance" Proc SMIRT 11 Tokyo 1991

- 58 -

Flgure 1

Age distribution of operable plants in the USA

Figure 2

Schematic diagram of a Westinghouse steam generator

Figure 3

Normalized failure frequency for globe valves in the Westinghouse CI
function

Figure 4

Failure modes for globe valves in Westinghouse PWR CI function
during two periods

Figure 5

Failure causes for globe valves in Westinghouse PWR CI function

Flgure 6

Frequency of globe valve failures caused by major failure causes

Figure 7

Weibull Projection, Fraction of Tubes Cracked vs. EFPY

Flgure 8

Plug Ratio Projection Using Weibull Projection and Plant Speclfic Crack

Growih Rate Data
Figure 9,, ,, predicted crack growth rate for ductile austenitic
The effect of n on,, ,,alloys (Fe and Ni-based) over the range of n values
consideredapplicable to 250-300C water. (after [10])Figure 10

Compilation of Alloy 600 crack growih rate data vs. stress intensity inhigh temperature deaerated pure water. Model predictions are shown for
n values of 0.5 and 0.7. (after [10]),, ,,- 59 -

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Figure lAge distribution of operable plants in- 60 -

the USAll

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-61-

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Figure 3

Nonnalized failure frequency for globe valvesin the Westinghouse CI function- 62 -

: NPRDS (1989 . 93)
SOURCE~~~
INTERNAL LEAK

1 - 11 YRS~TERNAL LEAK

1 2 - 24 YRSFAIL TO CLOSE
FAIL TO OPEN
FAIL TO OPER AS REQ

FOUND DURING I&S

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Figure 4

Pailure modes for globc valves in Westinghous~e

PWR CI function during two periods- 63 -

SOURCE : NPRDS (1 989 • 93)~i~

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AGE AT FAILURE (YRS)-1- DIRT BUILDUP - WORN PACKING - WORN sEAT/olsc

Figure 6

Frequency of globe valve failures caused by major failure causes- 65 -

0.9

0.632

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Pig ure 7

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Weibull projection, fi-action of tubes cracked vs.

- 66 -

32 50

EFPY

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weibull projectien

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crack tength:,

above 5 5mmb- .
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2

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Flgure 8

Plug ratio

projection using Weibull projection

and plant

specific crack growth rate data

- 67 -

50

r~

Andresen/Ford PLEDGE Predictions vs. 'n'

Ductile Fe- & Ni-base Alloys _ _

L 104

1 0~5

250 - 3000c Water _ _ - - ~_ ~_ - -

>

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.~

.(~
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ol~.((\~~ ~ -

o~ 103~ca/rll
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E

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,,
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Figure q.

The effect

1 0~9

20
30
40
Stress Intensity, ks~/inalloys (Feof n,, ,,on prec]icted crack growthand Ni-based) over the range ofapplicable

to 250-300C water. (after [lO])- 68 ~

50

60

ra te

for' ductile austenitic" nuvalues considered

1 02

>

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~ 101
u)

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20 Intensity,
30 40
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60

Compilation of Alloy 600 crack growth rate data vs. stress [ntensity

in high temperature deaerated pure water. Model predictions are shownfor n,, ,,values of 0.5 and 07 (after [10])- 69 -70

II-3

Structural Integrity Assessment of the Reactor Pressure Vessel

under the External Cooling Condition

Jong Sung Kim, Yoon Snk Chang, Tae Eun Jin, Korea PowerEngineering Company, Inc., Korea

Structural Integrity Assessment of the Reactor Pressure Vessel

under the External Reactor Vessel Cooling Condition
Jong Sung Kim' and Tae Eun Jin

Key Words : In Vessel Retention, Extemal Reactor Vessel Cooling, Core Melting
Accident, Transient State, Finite Element Method, Elastic-Perfect Plastic

Behavior, Larson-Miller Curve, Damage RuleAb stract

Since the suggestion of External Reactor Vessel Cooling(ERVO, the effects ofmelting and cooling on the response of structural integrity of the Reactor Pressure
Vessel(RPV) under core melting accident conditions have been investigated. This paper

describes the vessel response according to the ERVC condition and analysis method.

The steady state and transient analysis for the temperature and stress field were
performed using ABAQUS. Especially, transient analyses were studied for the variableboundary conditions. To obtain an analogy with real phenomena, the material properties

were determined by combining and modifying the existing results considering phasetransformation and temperature dependency. The results show that the vessel can bemelted if there is no external cooling. Finally, the potential for vessel damage isdiscussed using the Larson-Miller curve and damage rule.

l . Introduction
ERVC(extemal reactor vessel cooling) has recently been studied and developed as the(4 5)management strategy for severe accidents(l-3). In previous studies ' for IVR(in-vessel
retention), the extrapolation method was used for the determination of material properties(6)

beyond the temperature range presented in the ASME code or the material propertiesof RPV material not to consider phase transformation and change are used.
But the real material properties are considered to be different from those found by the

extrapolation method because material properties, especially thermophysical material
properties, are abruptly changed with temperature due to phase transformation. Previous
studies(4,5) also utilized the steady state analysis to maintain a constant value of theboundary condition variables for time. However, optimal ERVC time can not be determined* KOPEC(Korea Power Engineering Company, Inc.)- 71 -

using steady state analyses. Therefore, the ambiguous conjecture that the reactor cavity

would be completely submerged before molten debris reaches the RPV Iower head, ispresented by those studies.

In this study, however, the material properties for temperature are established like real

phenomenon, the boundary conditions are determined to consider the transient state thatboundary condition variables - the levels of molten debris/external vessel water, and thetemperatures/pressures of molten debris, external vessel water and gas in core - aresubjected to change over time. The database and the user subroutine program areconstructed in order to determine the material properties of molten debris and thespecification of boundary conditions according to transient state. The temperature andthe stress analysis are performed for the core melting accident due to large break
LOCA by using ABAQUS(7). Finally, the damage and the stuctural integrity for RPV
are evaluated by using the Larson-Miller curve and damage rule(8). In addition, theeffects of analysis conditions on structural integrity - molten debris quantity, ERVC

condition, analysrs method and connection condition with RPV upper part - arereviewed.

2 . Analysis model
2.1 Model configuration

The RPV configuration is shown for the case where molten debris and metallic layer

are located in the RPV Iower head in Fig. 1. The materials of RPV and cladding are
SA508 Gr.2 Cl.1(9) and SB166(ro) respectively. The thickness of cladding, t.lad, rs
3. 175mm.

2.2 Chemical composition and material properties

The chemical composition of SA508 Gr.2 Cl.1 is presented in Table 1(9). The liquidus

and solidus temperatures of SA508 Gr.2 Cl.1 are determined to be 1501.04iC and 1461.0
9~C respectively by substituting the chemical composition of Table I into Howe's(n) andKubachewski's(12) correlations. Density pRPV rs determined from the correlation by(13)

Jablonka et al. as shown in Fig. 2. The specific heat, c p. RPv and the thermal

conductivity, kRPv, for temperature are detemuned from ASME Code(6) and Thomas- 72 -

(14)results as shown in Fig. 3 and 4, respectively. The elastic modulus, ERPV and theyield strength, 6 ys. RPv, for temperature are determined from ASME Code(6) and Grill'sresults(15) as shown in Fig. 5 and 6. Poisson's ratio vRPv rs determined to be 0.3. The

thermal expansion coefficient, aRPv, for temperature is determined from ASME Code(6)(13)

and Jablonka's results as shown in Fig. 7.
(16)
The Larson-Miller curve forSA508 Gr.2 Cl.1 is shown in Fig. 8. LM~ = O . OOIT

(20 + Iog t.)) is Larson-Miller parameter. t, is rupture time( hour).
( ro)

The chemical composition of SB166 is presented in Table 2 . The liquidus
temperature of SB166 is determined to be 1427 iC (6). The thermal conductivity of SB166 fortemperature, h*/~, is detennined as shown in Fig. 9(6). The emissivities of SA508 Gr.2
Cl.1 and SB166, eRPv, are determined as shown in Fig. 10(17)Molten debris is assumed to be composed of 72.25~(weight percentage) U02 and27.75~6 Zt02. The solidus and liquidus temperatures of molten debris are 2536~C and 254
2~C(18) respectively. The density of molten debris, pl' versus temperature is determined

as shown in Fig. 1 1(18,19). The specific heat c p, i and the thermal conductivity k(rs)versus temperature are shown in Fig. 12(18) and 13 , respectively.
The volumetric

expansion coefficient, ~,, for temperature is determined as shown in Fig. 14(18,19). The
viscosity, p /' for temperature is as shown in Fig. 15(18,19)

The metallic layer is assumed to be composed of stainless steel. Therefore, emissivityof metallic layer, et, is determined to be 0.43(17)3 . Boundary condition
3.1 Accident scenario

The significant event stage and the time frame for a core melting accident occurs due
to large break LOCA and SI(safety injection) failure are presented in Table 3(20)3.2 Boundary condition variable

Boundary condition variables - the levels of molten debris/metallic layer/external
vessel water, the temperatures/pressures of molten debris, metallic layer, external vessel

water and gas in core, and the decay heat of molten debris - change over time.

- 73 -, J,

Therefore, the applicable boundary condition region and the magnitude of heat transfer
coefficient also change according to time.

The levels of molten debris and metallic layer over time, H, and Ht, are shownin Fig. 16(20). The temperatures of molten debris and metallic layer over time, T, and(20)

Tt, are shown in Fig. 17 . The temperature and the pressure of the gas in coreversus time, Tg" and Pg", are shown in Fig. 18 and 19(20) respectively. The level, thetemperature and the pressure of external vessel water versus time, H/' Tl and(20)

Pt, are shown in Fig. 20, 21 and 22 , respectively. The decay heat of molten debris

for time, Q, is shown in Fig. 23(20). The Air temperature outside of RPV, T.i., issupposed to be maintained at a constant 1271C.

3.3 Thermal boundary condition
In transient analysis, the thermal boundary condition of RPV is determined using thefollowing assumptions:o The heat transfer is assumed to be the one for a quasi-static state.

o The therrnal conduction through debris crust and the gap between molten debris
and the inner wall of RPV Iower head are ignored.

o The temperature distributions of molten debris, metallic layer and extemal vessel
water are uniform over time.

o The heat convection on the RPV inner wall by the metallic layer is assumed to be

the one from molten debris. The heat transfer on the RPV inner wall during theperiod when molten debris and metallic layer exist in the solid state is assumed tobe the heat convection from the molten debris.

o The water exists in RPV Iower head until 6528sec after the accident but the heat
convection is due to the gas in the core conservatively.

o The effect of insulation on the thermal behavior of RPV is not considered.

o The effect of cladding and gas in gap on the heat convection coefficients forthe inner wall of RPV Iower head are considered by using an equivalent film
coef f icient(21)The initial and the thermal boundary conditions for RPV are shown in Fig . 24. The
heat transfer mechanism change for time is presented in Table 4. In Table 5, the heattransfer coefficients on the RPV inner and outer walls are presented for the heattransfer except radiation. In Table 6, the radiation heat fluxes on the RPV inner wallare presented according to the magnitudes of RPV inner wall temperature TRPV and~ 74 -

Tmet during the transient state.

In steady state analysis, the thermal boundary condition of RPV is determined byusing the assumptions that the lower head is fully submerged in cavity water prior to

the deposition of molten debris on lower head and molten debris is located on lower
head from the accident initiation time. The boundary condition variables are fixed overtime in steady state analysis.

3.4 Mechanical boundary condition

The mechanical boundary condition for the RPV is deterrnined by considering the deadweight effects of molten debris and external vessel water.

The mechanical boundary conditions for the RPV are shown in Fig. 25 where zl andz2 are the depth from the metallic layer surface and water level, respectively. Unit forpressure is MPa.

4 _ Analysis and Results
4.1 Analysis method

The analysis method for temperature and stress analysis is shown in Fig. 26. As
shown in Fig. 26, the analysis system consists of a database, an ABAQUS Input File(7)and an user subroutine program(7). The database presents the material properties ofmolten debris with temperature. These material properties are presented in section 2.2.2.

The ABAQUS Input File(7) assigns the finite element model and configuration, material

properties of RPV, boundary conditions, analysis type and output data. The user
subroutine prograrn(7) assigns the boundary conditions associated with the time variationof boundary condition variables and links between the database and the ABAQUS Input
File.

4.2 Finite element model

The finite element for an analysis model is shown in Fig. 27. The numbers of anelement and a node are 157 and 5059, respectively. The element property is an 8-nodequadratic axis-symmetric element(7). The finite element model is simplified by thedeletion of nozzles because the effects of nozzles on the integrity of RPV Iower headare minor. The model is an axis-symmetric model due to the axis-symmetric boundary- 75 -

conditions and model configuration.

4.3 Temperature analysis

Temperature analysis is performed considering molten debris quantity, ERVC condition
and analysis method.
The conditions for temperature analysis are presented in Table 7.

The NLT(nondimensional ligament thickness) w at MTGP(maximum temperature
generation part) versus time are shown for conditions @ and (~) in Fig. 28. The NLT

w means the nondimensional value calculated to divide ligament thickness w~ byinitial thickness w, at MTGP. As shown in Fig. 28, the value of NLT is zero andmolten debris penetrates through the thickness of RPV Iower head at 5.5xl03sec for thecase that ERVC is not applied and steady state analysis is performed. However, molten
debris does not penetrate in the case that ERVC is applied and steady state analysis isperformed .

In cases of condition (~) - @, the NLT w at MTGP are shown according to time in
Fig. 29. As shown in Fig. 29, the effects of molten debris quantity on NLT at MTGP
are minor in the case that ERVC is applied and steady state analysis is performed.

In cases of condition ~) and ~), the NLT w at MTGP are shown according to timein Fig. 30. The time for transient analysis when the NLT is maintained with constantvalue is later than the one for steady state analysis. When this steady state is reached,

the NLT by transient analysis is less than the one by steady state analysis.

The NLT w at MTGP versus time are shown for conditions ~) - O in Fig. 31. The
effect of ERVC completion time on NLT is minor for the case that the one is less than
10000sec and the NLT decreases with the increase of one.
4.4. Stress analysis

Stress analysis is performed considering the following variables : molten debris
quantity, ERVC condition, connection condition with RPV part and analysis method.
The conditions for stress analysis are presented in Table 8.

The hoop stresses at the internal points of MTGP versus time are shown for
condition ~) in Fig. 32. The hoop stress at MTGP increases with the radial coordinate- 76 -

r and the maximum value of hoop stress occurs on the outer wall surface of MTGP.

The hoop stresses at the outer wall of MTGP versus time are shown for conditions~) - R in Fig. 33. The hoop stress at the outer wall of MTGP increases with molten

debris quantity in the case where ERVC is applied and steady state analysis isperformed .

The hoop stresses at the outer wall of MTGP versus time are shown for conditions
R and C in Fig. 34. The hoop stress at the outer wall of MTGP has almost a constantvalue irrespective of the connection condition with RPV upper part.

The hoop stresses at the outer wall of MTGP versus time are shown for conditions

R and ~) in Fig. 35. The hoop stress at the outer wall of MTGP caused by thetransient analysis is smaller than the one by the steady state analysis.

The hoop stresses at the outer wall of MTGP versus time are shown for conditions~) - ~) in Fig. 36. The hoop stress at the outer wall of MTGP versus time decre4ses
with an increase in ERVC completion times.
4.5 Damage and integrity evaluation
The conditions for damage evaluation are equal to the ones for stress analysis.

Damage factors are calculated by using the temperature and von-Mises effecpive
stress distributions of RPV, and the Larson-Miller parameter and curve for SA508 Gr.2
Cl.1.

To calculate the cumulative damage factor 1), the following life-fraction rule(8) is
used:

t'

D=~tl =1

(1)

where t, is the duration time for specific condition i and t~ is the rupture time for
specific condition i. The satisfaction of equation (1) means rupture.

The failure times when the average value of the cumulative damage factors of allinternal points in the maximum temperature generation part is equal to I are presented

for various conditions in Table 9. From Table 9, the failure time decreases with anincrease of molten debris quantity for the case where ERVC is applied and steady state
analysis is perfonned. The failure time by steady state analysis is smaller than the one

by transient analysis as shown in the comparison between the results of conditions ~)- 77 -

and ~. The failure

time decreases with an increase of ERVC completion time .

5 . Concluslons
The following conclusions are obtained from the structural integrity evaluation
performed for RPV subjected to a core melting accident due to large break LOCA:(1) Molten debris penetrates through the thickness of RPV Iower head for the casewhere ERVC is not applied and steady state analysis is performed.
(2) The failure time decreases with the increase quantity of molten debris for the casewhere ERVC is applied and steady state analysis is performed.

(3) The failure time by steady state analysis is smaller than the one by transient
analy sis.

(4) The failure time decreases with an increase of ERVC completion time.

References(1)

O. Kymaelaeinen, H. Tuomisto andT.G. Theofanous, In-vessel

melt retention

as a

accident management strategy for

theLoviisa nuclear power

plant, in the

24th

and retention

of a

melt,

ldaho Operations

Of fice,

WRSM, Bethesda, Maryland, 1996.
(2)

T.G. Theofanous et

In-vessel coolabilityal.,

DOE/ID-104eo, Vol.1, U.S. Department of Energy,(3)

R.E. Henry et al.,

Cooling of core debris within

Nuclear Technology,

Vol.101 ,(4) J.L. Rempeat al.

,

vessel1995.

lower head,

1992, pp.385-399.

Light water reactor

EGG-2618,ldaho National Engineering(5)

the reactor

c orelower head f ailureLab.,

analysis, NUREG/CR-5642,

ldaho, Inc., 1993.
EG&G

D. Azordi and P. Gruner,

Numerical simulation ofpenetration subjected to a

postulated core damage accident, Transaction of the 14th

International Conferenceon Structure Mechanics

a BWR vessel lower head within Reactor Technology, Lyon,

France, 1997, pp.443-450.

(6)ASME bojler and pressure vessel code,

Section 111,

- 78 -

Division 1,

Appendices, 1995.

(7) ABAQUS user's manual, Ver.5.6, HKS Inc., 1996.

(8) R. Viswanathan, Damage mechanisms and life assessment of high-temperaturecomponents. ASM International, 1989.

(9) ASME boiler and pressure vessel code, Section II, Part A-Ferrous, 1995

(10) ASME boiler and pressure vessel code, Section 11 Part B Nonferrous 1995

(11) A.A. Howe, Estimation of liquidus temperatures for steels, Ironmaking andSteelmaking, Vol.15, N0.3, 1988, pp.134-142.

(12) O. Kubachewski, Iron binary phase diagrams, Berlin Springer Verlag, 1982.

(13) A. Jablpnka, K. Harste and K. Schwerdtfeger, Thermomechanical properties of ironand iron-carbon alloys: density and thermal contraction, Steel Research, Vol.62,
No.1, 1991, pp.24-33.

(14) B.G. Thomas, I.V. Samarasekera and J.K. Brimacombe, Mathematical model of thethermal processing of steel ingots: part 1. heat flow model, MetallurgicalTransaction B, Vol.18B, 1987, pp.119-130.

(15) A. Grill, J.K. Brimacombe and F. Weinberg, Mathematical analysis of stresses incontinuous casting of steel, Ironmaking and Steelmaking, No.1, 1976, pp.38-47.

(16) G.L. Thinnes. G.E. Korth, S.A. Chavez and T.J. Walker, High-temperature creep

and tensile data for pressure vessel steels SA533B1 and SA508-C12, Nuclear
Engineering and Design, Vol.148, 1994, pp.343-350.

(17) T.G. Theofanous, G. Wang and X. Chen, The emissivity of the steel layer, T.G.

Theofanous et al., In-Vessel Coolability and Retention of a Core Melt, DOEl
ID-10460, Vol.1, 1995, 1-3-8.

(18) J.K. Hohorst et al., SCDAP/RELAP5/MOD2 code manual, volume 4: MATRO-alibrary of materials properties for light-water-reactor accident analysis, NUREG/

CR-5273, EGG-2555, Vol.4, R3, EG& G Idaho, Inc., 1990.

- 79 -

(19) C.C. Chu et al. Uncertainty analysis for therrnophysical properties used in in-vessel

retention analyses, T.G. Theofanous et al., In-Vessel Coolability and Retention of aCore Melt, Appendix L, DOE/ID-10460, Vol.1, 1993, L.1-28.

(20) Korea Power Engineering Company, Inc., Severe accident evaluation report, 1996.

(21) J.P. Holman, Heat transfer Fourth edition Tower Press 1976(22) R.J. Witt, Local creep rupture failure modes on a corium-loaded lower head,
Nuclear Engineering and Design, Vol.148, 1994, pp.385-411.

(23) M. Jahn and A.A. Emara, High Rayleigh number convection in enclosed
fluid layers with internal heat sources, U.S. NRC, NUREG-75/065, 1975.

(24) M. Jahn and H.H. Reineke, Free convection heat transfer with internal heat sources:

calculations and measurements," Proc. 5th Int. Heat Transfer Conf., Tokyo, Vol.3,
NC-2.8, 1974.

(25) M.N. Ozisik. Heat transfer: a basic approach, McGraw-Hill, Inc., 1985.

- 80 -

Table 1. Chemical composition of SA508 Gr.2 Cl.1(9)Table 2. Chemical composition of SB166(ro)Table 3. Significant

event and time of core melting accident due to LB LOCA(20)

Table 4. Heat~:Note) B

bottom U : uppertransf er

mechanismchange according to

m.1. : metallic

layer m.d.

- 81 -

molten debri stime.

Table 5. Heat convection coefficients applied for various parts of RPV wall.(unit : W/m2 IC )* : the boiling convection coefficient of water on Leidenfrost point under latm.

Table 6. Radiation heat flux for transient state and various temperature conditions.(unit : W/m2)~ Note) o : Stefan-Boltzmann constant(=5.669xl0~8 w/m2ic4)
TRPV. Tmet. Tg~~ : Absolute temperatures(K)R-met : Heat flux from RPV inner wall to m.1.
met-R : Heat flux from m.1. to RPV inner wall
R-gas : Heat flux from RPV inner wall to gas- 82 -

Table 7. Conditions for temperature analysis.

Table 8 Conditions for stress and damage analysis.

*

The ERVC initiation time is 1800sec.

Table 9. The penetration and the failure times for each condition (umt xlO sec)

- 83 -

~A~08 cr ~i~ Ci l

8QOO,'Llj"f olr f~*l~/Cy!;nc~'*' cll

7eoo -

sheli

\

/

E

406n 19l

OD,~ 720O

,

O
680o

l

\

~
,,-••-,/:L~\; ,~e40ooFig. 1.

RPV configuration and composition.

2. Density vS.

Fig .

8.oEtg600
900 1200
Tem peratu re( 'C )

300

15ao

18OO

tem p.

4400o
*

O

rl

L)

o
o

6.0E•9

~E

Q
o
co

E

i'l

o)

E~

E

x:~

4 oE~g
It

CD
q,

!:

CL

o
::

U
c

\

\

!

,

f

\

~

\

\

i

\\

j

32000
\

o
o

2.oE~g\,U

E

I
40o 8oo 120o

a)

240oo

16oo

T e m pe ra tu re ( tC )Specificheat Of SA508 vS. temp .

280ac

c
h

O.OE~O

Fig. 3.

360oo

:~

!

o
o
a,
(,)

400oo

,,,

r-\

Fig.4.

- 84 -

\* *~

/

4oo
80o 1200
Temperature('c)

Thermal COnductiVity Of SA508 vs.

160o

temp .

400OOO

2'5Et8~:\2'oEt8 'oo,oEE 1'5E~8 ~O,~-

300OOO
,p

\

\

,L
e,:

\

,,

*\

\

\
\

OD

o,

r
,
l 1 oE'8 L
~O
E
l
o
o
c9
UJ

\

\

\\

c
o

\

20OOOO

\

\

\.

O~

,,

5'oEt7 h\\Uo> 100OOO

\

\

i

[

_+_L]

o oEto4oo
12oo
Tem80operature('O

Elastic

Fig. 5.

O

1200
400
80o
Tem
perature(C)16oomodulus of SA508 vs. temp. Fig. 6.

Yielding

iaO
2.OE-5looo '\

,o

c
,I

c9

o

~
=

a
X
c-

1 SE-5

~ ,\\
[

O

loo

O

1-

f

O
o
o
o
o
o
O,

o

1 OE-5

\

~r \\
\

o
,:
c:

VS. temp .

loooo

2 5E-5costrength of SA508

1600

\

\.

cL

a

\

10 L

5 OE-6c,o:-

IL=LJ-O OE~O2oo 4oo 60o 80o looo 12oo 14oo

16oo

Tem peraturecc)

Fig .

7.

Average linear thermal expansionof Flg 8

of SA508 coefficient vs. temp.

- 85 -

25 3o 35

LMP

Larson-Miller curve

fOr

40

'5

50

55

SA508Gr.2 Cl.1.

o9

5o

o
o
CD

~'

E

~

~
~

::

CO 4o
~

O
-

ll

a:
U,

ll

'l

(O

~

//

~O
3o
~>

o
c
c

>
o-

>
O

a~
.*

1:,

O
O

:-

E
O

o .6

a .5

E

UJ

J:

Fig .

,

0.7

o

~>

2o

15

h

o.8

co

10

o .4

600
Tem 12ooperatu re (K)

1 200

IOOO

1 800

16OO

140o

Tem perature( C)

9. Thermal conductivity of SB166 vs.

tem p .

Fig .

10. Emissivity of RPV inner wall vs.
tem p .

1 ooo

100oo

X
aD

E

x 80O
,

O,

~ 90OO

,,

o

Q
Q

Q 6OO

,D

L

O 800O

o

=

=
o 40O

:~ 700O

r

O

CD

~O

o
o
:; 2DO
o
o

co

=
O

a

Q

co

o

60oo
700

140o

210o 280o

700

3500

11. Density of molten debris vs.

14oo

21OQ 280o

3soo

T e m pe ratu re (K)Te m peratu re(K)

Fig .

./L

e,

co

temp. Fig.

- 86 -

12. Specific heat of molten debris vs.

tem p .

X
E
~
~

10

o

o

=

6

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- 87 -

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- 88 -

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- 90 -

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part vs. time for condition ~).

- 92 -

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for various- 93 -ERVC completion times.

Session 111

Plant Experiences

III- lThe Evaluation of IGSCC Problems of Stainless Steel Piping

in Taiwan BWR-6 Nuclear Power Plant
Kuen Ting, Department of Nuclear Regulation Atomic Energy
Council,Taiwan, Republic of China

The Evaluation of IGSCC Problems of Stainless Steel Piping

in Taiwan BWR-6 Nuclear Power Plant

Kuen Ting
Department of Nuclear Regulation
Atomic Energy Council
Republic of China

ABSTRACT
Taiwan BWR-6 Kuosheng Nuclear Power Plant Unit I implemented the inspeotion of the IGSCCsusoeptible wekhnents. of stainless steel piping in the reactor recirculation, reactor water clean-up,
residual heat removal, core spray and feedwater systems. The purpose of this paper is to present thestatus of the fraoture problems in the weldments. The crack growth analysis due to IGSCC and thestandard weld overlay design based on the ASME Code Section XI and NUREG-03 1 3 Rev. 2 for thefracture weldments are discussed in detail. Then, the contingent programs including the inspection
program, fracture evaluation, the standard weld overlay are completely established to prevent the pipe
break during the reactor operation.

INTRODUCTION
The objeetive of this paper is to present the intergranular stress corrosion cracking GGSCC) statusofthe primary coolant system at BWR-6 Nuclear Power Station (NPS) unit I in Taiwan. This work wasaehieved by identifying all the IGSCC-susceptible systems and the associated weldments. Following theprocedures of inspeotion, evaluation and repair are based on the requirements of the United States

Nuclear Regulatory Commission (USNRC) Report NUREG-0313. Rev. 2[l] and Generic Letter 88O I [2] .

An inspections have been performed on the various IGSCC-susceptible weldments at BWR-6 NPS
unit I in Taiwan since 1 983. The inspections, since 1 987 (EOC-4), have been performed to therequirements of NUREG-03 13, Rev.2. The total of 184 weldments which are contained in reactor
recireulation, reaetor water clean-up, residual heat removal, core spray and feedwater systems are
categorized as five groups. However, the inspection frequencies for the various weldment categories
were more than the NUREG-03 1 3 requirements during the past I O years of operation in Taiwan. Whena flawed weld is identified during inservice inspection, the flaw evaluation may be performed to meetthe requirements of ASME Code Section XI in lieu of repair to justify continued operation for a
specified evaluation period. This evaluation is based on the linear elastio fracture mechanics to Galculate

the IGSCC growih. If the crack growth exceeds the allowable critical length and depth, identified
IGSCC in stainless steel may be used standard weld overlay for long-tenn repair. The design of thestandard weld overlay is based on the net section collapse conoept. Significant field, and experimental

evidenGe has been assembled to meet ASME Code Case 504-1 [3] for the long term repairs of weld

- 95 ~

overlays. The effeGts of weld overlay on the sustained stress, global mass and global stiffness were also
evaluated after each outage repair based on the Generic Letter 88~)1 with Supplement I [4].

During the past I O years of the operation of Taiwan BWR~5 Unit I , total of 24 weldments at therecirculation and RHR systems implemented the standard weld overlays. A Gomplete proceduresconsisting of in-service inspection, flaw evaluation, standard weld overlay and quality control for the

IGSCC-susceptible weldments were established in Taiwan. The inspection for these overlayed
weldments is continuously conducted during each refueling outage. Due to the long-term inspection, theprevention of the pipe fracture due to IGSCC during the operation can be achieved to maintain the
nuclear safety.

INSPECTION PROGRAM OF IGSCC-SUSCEPTIBLE WELDMENTS
Five systems considered IGSCC-susceptible weldments in in-service inspection program during
each refueling outage. These systems are :l . Reactor Recirculation (RR:)

2 . Reactor Water Clean-up (RWCU)
3 . Residual Heat Removal (RHR;)
4. Core Spray (CS)
5. Feedwater (FW)A total of 1 84 IGSCC-susceptible weldments were identified for these systems. Figure I and 2
present the susGeptible weldments on the reoirculation, IuJR suction and RWCU systems. Other than
the nozzle and thermal sleeve welds (KB, KC and TSW welds) all welds comprise of either Type 304 or
3 1 6 stamless steel weldments. Some of the weldments in these systems have been solution heat treated.
Details of the materials of the Nl (recirculation outlet) and N2 (recirculation inlet) nozzles and safe-end

weldments are provided in Figure 3 and 4. The N2 safe-end has a welded thermal sleeve which
considered the crevice in this evaluation. Most of the piping of the RHR, core spray and feedwater
systems are fabricated with Garbon steel and are, therefore, not susceptible to IGSCC. In these systems

only, the nozzle welds are considered susceptible since they were partly welded with Inconel 1 82
material. Details of the materials of the N4 nozzle (feedwater nozzle), N5 (core spray nozzle) and N6(RHR nozzle) weldments are shown in Figure. 5-7.

Inspection of the various IGSCC-susceptible weldments at BWR-6 unit I has bcen performed
since 1 983. However, the inspections performed to meet EPRI/USNRC requirements for IGSCC
detection and sizing started at 1 987 (EOC-4). The inspection frequencies for these weldrnents weremore than the requirements ofNUREG-03 1 3, Rev.2, as listed in Table I .

INSPEcrION RESIJIJTS
No indications were reported at Tawain BWR-6 unit I for the flfst frve operating cycles (EOC-lthrough EOC-5). However, 10 flawed welds were found during 1 989 refueling outage (EOC-6). Morethan half of fracture welds located at the recirculation loop B (6 of I O). After the fracture analysis, AS-l 4, BS-J2 and N2G-J2 welds cannot operate during the next fuel cycle. Then, the standard weld overlay
repairs (as shown in Figure 8) were applied to 9 welds except BS-J4.

There are ten mlet nozzle-to-safe end weldrnents between the RR system and Reactor Pressure
Vessel. A thermal sleeve is welded to the inside of the safe end and each nozzle. These thermal sleeve-

to-safe end weldments within the inservice inspection (ISI) program, are identified as
N2A/B/C/D/E/F/G/H/J/K-TSW. During 1 991 outage (EOC-7) cracks were detected in four of theseweldments (N2A/D/G/H-TSW) and conducted weld overlay repairs. In 1 992 outage (EOC-8), six
cracks in weldnwnts N2E-TSW were discovered as shown in Figure 9, but a flawed pipe evaluation~ 96 -

determined that this weldment was suitable for GOntinued operation without repair for at least one
additional fuel Gycle. These Gracks in N2E-TSW were observed to grow (as list in Table 2) requinng aweld overlay repair in 1 993 outage (EOC-9). Thus, Iicensee decided to conduGt the weld overlay repair

for the weldments N2B/C/E/F/J/K-TSW. Table 2 provides a summary of all disGOVered flawed
weldments during each refueling outage and the disposition of these flaws. ReGent the 1 2th refuelingoutage (1997), four welds N2A-J4, N2C-J4. AD-J5 and AS-J7 of Loop A were still found IGSCC andoverlay repairs were conducted. Up to now, only one indication at B S-J4 of RHR system which was
identified as the GOunterbore signal was continuously inspected. All the overlay welds were inspectedduring each outage. There is no indication to find again.

The stress improvement technique IHSI considered in NUREG-03 1 3, Rev.2 has been applied to 75
welds in the reGirculation and RHR system in 1 984. Excluded 4 safe-end-to thennal sleeve welds with
Inoonel 600 creviced design, there were 1 6 welds with IHSI to discover IGSCC.
The inspection was conduoted by the ultrasonic examination. The 1/2 V technique was used during
the early stage of the in-service inspection. However, the weld surfaces often were concave due to theshrinkage and grinding. Thus the contact between the sensor and the surface was not fit well to reduce
the possibility of the crack detection (see Figure I O(a)). The I .5V technique as shown in Figure lO(b)was used to extend the inspection scope and increase the reliability of the crack deteGtion.

FLAW EVALUATION
When the cracks were discovered at the inspection weldments, the fracture evaluation based on the

NUREG-03 1 3 Rev. 2 and ASME Code Section XI must be performed to determine the repair works or
not .

Flaw evaluations are generally performed using linear elastic fracture mechanios teohniques. The
applied intemal pressure, dead weight, thermal expansion, weld overlay shrinkage and residual stresses
were found in the stress reports. With these load combinations and a linear elastic fracture mechanics

model for the flaw weldment, the crack tip stress intensity KI can be detennined by the Buchalet-

Bamford polynomial fit method[5]. Flaw growth can be determined using the upper bound weld
sensitized crack growih data shown in NUREG-03 1 3 . Rev.2:

da 2 161
dt = 3'59 x l0-8KI ' 'where da /dt denotes the craok growth rate, inches/br.

The crack was conservatively assumed to have an initial depth ao equal to the maximum reported
depth and the initial length lo equal to the sum of the individual lengths. Crack growth analysis was

performed to determine the final crack depth af after 1 8 months of operation (or 1 2000hours). The
aspect ratio of the crack depth and length a /1 was assumed as the same during the crack growth. Then,the final crack length can be calculated as

If = af lo

ao- 97 -

Then the allowable flaw size is determined using the criteria in ASME Section XI IWB-3640.

For the case of the inspection data of N2E-TSW as listed in Table 2, the input parameters wereexpressed asao = 0.14", Io = 8. l" ao /lo = 57.85
Using the crack growih analysis, the fmal crack depth af reached 0.475" after the assumption of

l 2000 hours operation. Then the fmal crack length lf was 27.5". Corresponding to the parameters in

IWB-3640, four parameters lo /rd, If /rd, ao /t and af /t were used to determine the allowablecrack size as shown in Figure I I . The orack size can be tolerate to operate at the next fuel cycle without

the repair. However, the significant growth of the fraoture weldrnent was observed during the next
refueling outage as displayed in Table 2. Using the fracture analysis, the life prediGtion of the cracks

was only 2000 hours. Thus, the repair work must be neGessary for the safe operation during the next
fuel cycles.

Weld overlay induced shrinkage stresses were analyzed by ANSYS finite element computer
program[6] . The actual weld overlay shrinkages measured after application of each weld overlay wereequivalent as the thermal shrinkages at the nodes corresponding to repair welds. This approach is used
to simulate the mechanical shrinkage observed in the field.

Ea~TI = 5
where I is overlay length, 6 is the as-built weld overlay shrinkage. E is Young's modulus and a isthermal expansion coefficient. The equivalent temperature drop AT will input at the nodes
corresponding to repair welds .

WELD OVERLAY REPAIR
Presently, there are 24 welds with weld overlay repairs. All these welds are m the reclfculation
system. The overlays were designed as " standard" overlays meeting the requirements of ASME Section
XI IWB-3640 for overlay sizing. This repair program is to define the managerial and administrative

eontrols for the implementation of Weld Overlay Repairs during outage which are covered the
document preparation and review, preparation work, implementation and final reports. The detailed
work items are listed as shown in Figure 1 2. The scope of work addressed by the repair program shallmeet the requirements of ASME Section XI and Code Case 504-1 , In addition, the scope of work shallcomply with the requirements of NUREG 03 1 3 Rev. 2 and Generic Letter 88-0 1 . All piping from thereactor pressure vessel to the frrst isolation valve was originally constructed in aceordanGe with the
ASME SeGtion 111 identified in the piping system stress report. The design, welder, filler metal

procurement, and welding repair shall be performed in accordance with ASME SeGtion XI.
Nondestructive examination and testing work shall be performed in aooordance with ASME Section V
and XI.

The standard overlay design considers the determinations of the thickness and the width. The net
section GOllapse criterion is used to calculate the minimum thickness [7]. 360" circumferential tbroughwall~rack and safety factor 3 are assumed. The overlay thickness oan be GalGulated by the iterations of
the following equations.

- 98 -

p = PT(1 -a/t - P~ Iaf )2 - a/t

p = 2af (2 _a/t)sinp;T

where af rs the material flow stress, a is the crack depth whioh equals to pipe wall thiGkness and t is

the overlay thickness. An iteration scheme is performed using these equations until the minimum
required weld overlay thickness is determined.

The width of the overlay must consider the following conditions:(1). The width is at least I .51!~(2). InspeGtability(3). Cover the crack extension(4). To minirnize possible damage to base material.

The effects of weld overlay on the sustained stress, global mass and global stiffaess were also
evaluated after each outage repair based on the Generio Letter 88-0 1 with Supplement I .

CONCLUSION
Reported fracture weldments information survey results show that IGSCC is probably the mostcommon aging degradation meehanism for Taiwan BWR-6 austenitie stainless steels piping. The
strategy for controlling or mitigating IGSCC is necessary to manage in the future.

Due to the in-suffilcient of supply of the high quality Class I pipe components, the replacement of
the fracture pipe is impractical in Taiwan. Therefore, the fracture analysis and the standard weld overlay
repair are the important contingent plant during each refueling outage. These evaluation and the repair

works can demonstrate that there is substantial margin for each of these cracked welds underconservative, bounding conditions to allow for continued operation for a minimum of one additionall 8month operation eycle. A complete procedures inoluding the inspeetion program, flaw evaluation,
repair and quality control have been established in Taiwan to augrnent the safety of long-term nuclear
operation.

REFERENCE
l. United States Nuclear Regulatory Commission (USNRC) NUREG 0313, "Technical Report onMaterial Selection and Processing Guidelines for BWR Coolant Pressure Boundary Piping", Rev. 2,
publishing January 1 988.

2. United States Nuclear Regulatory Commission (USNRC) Generic Letter 88-0 1 , Subject: NRC
Position on IGSCC in BWR Austenitic Steel Piping, dated January 25, 1 988, with Supplement I ,dated February 4, 1992.

3. Taiwan Power Company, "Generic Letter 88-0 1 . Supplement I Weld Overlay Shrinkage Effect
Assessment Report for Second Nuclear Power Station Unit 1 993 Outage, ZTP046 02 10 1 993- 99 -

4. American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code Section XI,"Rules for Inservice Inspection ofNuclear Power Plant Components", 1 989 Edition.

5. Buchalet C.B. and Bamford W H "Stress Intensity Factor Solution for Contmuous Surface Flaws
in Reactor Pressure Vessels", Mechonics ofCrack Growth, ASTM: STP 590, 1976.

6. Desalvo G. and Gormin, User Information Manual and Theoretica/ Manual. ANSYS Engineering
System, 1 991 .

7. Smith E. ,"Theoretical Justification ofthe Association ofa Critical Net-Section Stress with Fracture
Initiation at a Crack Tip", International Journal of Pressure Vessels and Piping, Vol. 8, 1 980, pp.
303-3 1 1 .

- 100 -

Table 2 The Data of Cracks of N2E-TSW Weldment at Two Inspections

Table 3 IGSCC Flaw Records of Taiwan BWR-6 Piping Weldments- 101 -

N18~~

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Figure. I Loop A welds[3]Figure 2 weldments in Loop B- 102 -

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Figure 5 N4 Nozzle Weldments- 103 -

Cb~,• G8lec,

RPV b~,SA8NGL,OTh~da-- Ednbl
aold 88leQT~~ef~i h-~,~,,~~A~a.Fe,

Figure 6. N5 Nozzle Weldrnents
RPV Nordo(sA~eGtn)

8~ E~la~•, 8ilee)

bQ,.' Ie2

8u.c

,aw~

a,u•' ~tn,

,cCw•U(C•tn s,•eO
OruJ ul~)
J1 W~'~'ff;B'/\ \~,,',,(a'lc'taRrg~O 5d.E,eE'al~'

T~~ S~-~!•D(S~e ct Fa]
11•md Slw•(StJdExD•ao(~Go a. O(t~" 88le6,

Figure 7. N6 Nozzle Weldments
PIPE WAL L/

WELD OVERLAY

\

., ,

i
~

ar

S_

Figure 8 standard Weld Overlay- 104 -

f\

BUTT WELD

No. 1

N0.7

FLOW

O 6"

No 6

o. 7"

No. 2
1 .411

Indication Area

1 7,l

1 .711SAFE WELD THERMAL
END

No. 3

1 .5"

SLEEVE

0.3,,

No. 5

No 4

Figure 9 N2E-TSW Weldment Cracks

CRACK
l 12

(a) I /2 V Ultrasonic Examination Technique

CRACK
l .5V

(b) I . 5 V Ultrasonic Examination TechniqueFigure I O I /2 and I .5 Ultrasonic Examination Technique- 105 -

'wa eao Ldd'T

IO

\

f

/

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IJ)

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i '3 '4 s G P' 'e 1'ccRAcF~lF'wcTHla"ciJ"FFRE'Nc'E t L "~' d 'Figure I I Fracture Mechanios Analysis ofN2E-TSW

e NCD DOCUMENT e SITE WALK DOWN O pLANT CONDrrloN
O REMO VAL
EQUD?MENTO REPAIR PROGRAM ' e INTERFERENCE CHECKAND DECONO SPECIAL PROCEDURE e plplNG HANGER e LEAK SEALePIPlNG HANGER CHECK

CHECK

CHECK

e DESIGN Doc e plpING SUPPORT O WORo PIPING SUPPORT CHECK
e pROCEDUIU3 O EQUIPMEN SET UP O FERRJTE CHECKo CLEANINGe MA113RIAL CERTIFICArl3D e WORK SCHEDULE O SHRINKAGE
e PAClaNG
CHECKMEASURe waLDER QUALIFICATION e MOCK UP TE ST e N DE

e CALn3RArION ' e WELDERoee

GL 88~)1 Sl EVALUAnoN
PACKAGE TUl~~1 OVERO QA PROORAMe

ISI PLAN RBVISEDe GL88~1 EVALUAIION e ALARAe HP SURVEYQUALIFICATIONO NDE REPORTe QC/QA HOLD POn~T
Figure 1 2 Weld Overlay Repair (WOR) Work Items- 106 -HYDRO TESTlll-2 ISI NDE in Korean Nuclear Power Plants
Yi-Hwan (Peter) Jeong, Korea Electric Power ResearchInstitute, Korea

ISI NDE in KOrean NuClear POWer Plants
Yi-Hwan (Peter) Jeong

Hee-Jong Lee

Yong-Sik Kim
Materials & Corrosion Laboratory
Korea Electric Power Research InstituteTaej on, Korea

Abstract
Structural integrity of nuclear components is important for a safe

operation of nuclear power plants. Therefore, the structural integrity
should always be maintained by reliable, periodic inservice inspections.

Korea Electric Power Company(KEPCO) operates the entire Korean
nuclear power plants. Since the safe operation and the inservice
inspection(ISD of the nuclear power plants are under the plant owner's
responsibility, Korea Electric Power Research Institute(KE;PRD, the R&D

division of KEPCO, has established an integrated ISI NDE support
system(called, ISI NDE Total Support System: TSS), primarily focused on

the inservice inspection of nuclear power plants. And the KEPRI has
initiated a long -term plan for the development of key ISI NDE technologythat is necessary for an efficient on-site support. This paper describes

the KEPRI's TSS and the long-terrn plan for the key ISI NDE
technology development.

Introduction
Since Kori Unit I , the first nuclear power plant in Korea, started

to generate electricity in 1978, Korea has built twelve nuclear power
plants and produces lOgigaWatts of electricity (25~ of the total electricity- 107 -

in Korea, and 9th in the nuclear power generation worldwide). According
to the latest survey (Table I ), a total of 336 units at 81 power plants

nationwide (including hydraulic, f ossil and nuclear) generated over4lgigaWatts of electricity in 1997(17th in the overall power generationworldwide). Over the past years, Korean nuclear power plants showedremarkable improvements in the perforrnance such as drop of unplanned

shutdowns and reduction of outages. The nuclear power generation is
now a wing of the main source of electricity in Korea. In a long-term,
the share of nuclear electricity production will increase to over 45%

nationally with the completion of additional nuclear power plants(23nuclear power plants by the year 2006).
In addition to

supplying a quantity of electricity to users, it is also

important to supply

high-quality, stable(interruption-free) electricity. To

supply high-quality,

stable electricity, the structural integrity of thenuclear components should be maintained by reliable, periodic inservice
inspections.

Tab le 1 . Electricity

generated by Korean Power Plants
As of January of 1998 (Unit

Notes: (1) 9th in the

nuclear and 17th in the total power generations worldwide.

- 108 -MW )

PSI/ISI
During the plant operation, plant components are under severe

environmental conditions such as high temperature, high pressure,corrosion, mechanical stress and vibration. Therefore, various types ofinservice defects or premature failures can occur in the plant components.
Such structural failures can cause catastrophic accident. In particular, thecomponent failures in the primary system in the nuclear power plant cancause serious damages to the properties and human lives by radiation
exposure. Therefore, periodic inservice inspections should be conductedduring the plant operation in order to maintain the structural integrity of

nuclear components.

Plant component inspections are classified as preservice
inspection(PSD and inservice inspection(ISD depending upon whether theinspection is perforrned at the completion of plant construction or during

the plant operation. The PSI is perforrned to confinn that all structuralmembers of the plant have been constructed safely and properly, and also

to acquire initial inspection data that can be compared with the future
inservice inspection results. The ISI of nuclear components is generally

perfonned in 10-year period during the plant operation. The PSI and ISI

in Korean nuclear power plants are perforrned according to ASME code

Section XI: "Rules for Inservice Inspection of Nuclear Power Plant

Components" Because the ASME code mandates that PSVISI be
perforrned by nondestructive exarnination(NDE), the NDE for nuclear
component examination is particularly important.

Status of ISI NDE
During the last several decades, the nondestructive
evaluation(NDE) technology for inservice inspection(ISD in nuclear power

plant has been greatly improved. Particularly, with the advancement ofcomputer technology in recent years, the NDE equipment for reactor
pressure vessel(RPV), steam generator(SG), turbine, pressurizer, reactor- 109 -

coolant pump(RCP) and pipings were fully 2lutomated and sophisticated.

Consequently, the ilccuracy, the speed and the reliability of the ISl
perforrn~mce \vere impro\'ed accordingly.

Ilo\\'e\'er, the KEPCO has been he'avily relying on foreigncountries in conduc ting the inservice inspections of the nuclear
components due to lack of key ISI NDE technology. Consequently, the ISl
operation costs KEPCO a great deal of financial burden. Furthermore, due

to lack of the key ISI NDE technology, the KEPCO has not been able torespond to site problems in an efficient and timely manner. Suchdifficulties of Korean nuclear power plants provided the needed impetus

for KEPCO to develop a long-term plan for the ISI NDE technology
development and to prepare for the independent operation of inservice
inspection of nuclear power plant components.

ISI NDE Long-Term Plan
The KE;PCO, as a utility, is primarily responsible for the safe andreliable operation of the nuclear power plants. Therefore, the KEPCO
monitors the periodic inspections of nuclear components performed bydomestic and foreign companies. The mission of KEPRI, the R&D
division of KEPCO, is to perform the research and development of ISI
N~DE technology as well as the associated technology transfer. To meetsuch mission, the KEPRI has established an on-site support system,
called ISI NDE total support system(TSS), that can respond to various
plant ISI problems in an efficient and timely manner. Furthermore, inorder to successfully support the plant ISI operation, the KEPRI has

established a long-term plan for the development of key ISI NDE
technology(Figure I ). The key ISI NDE technology areas were chosen
among the most important technical areas in conducting the on-site ISI
support tasks.

The TSS and the key IS'I NDE technology development programs
contain the following three stages: preparation, independent operation, and- I Io -technology transfer stages. The preparation stage(Years 1997-1999) is forthe initial development of the essential NDE and the ISI engineering
methods that will become the basis of the independent ISI operation. The

major NDE equipment will be automated and computerized to improve theefficiency and the speed of the nondestructive examination. The
independent operation stage(Years 2000-2002) is for the application of the

major ISI NDE technology for the nuclear component examination solely
by the domestic teams. Moreover, the plant-required ISI inspectors willbe trained and qualified by the domestic training and qualification

system ( Korean Perforrnance Demonstration Initiative: KPDD. The
technology transfer stage(beyond 2003) is for exporting the key ISI NDE
technology as well as the services to foreign countries as the opportunity
arises.

ISI

Tot a lKev. Techno I ogy

NDE Area1)Transducer2)Automated System
3)Signal Processing4)Pat tern Recognit ion

EN 'G Area

1) FMA
2) RBI

Suppor t

S~. ,s t em

E\G Area\.'DE Area

1)Special Insp.

1)ISI Plauning_9)QA_? ) FMA

3)\'DE Val idat ion

3)Repair & Rep.

4)NDE Consul t ing~)KINS Appro\~a l5)Flawed Sp. Fab.
ONDE Model ing5)C.ode Interpretat ion6)Training & Qual .

7 )We I d i ng7)Dat abase \.fanagement- \.'DE Data

- F~A Stress Data

Figure 1. ISI NDE Technology Tree-111-

ISI NDE Total Su pport System
The ISI of nuclear components typically include such tasks as ISI

planning, nondestructive testing(NDT), quality assurance(QA), fracture

mechanics analysis(FMA), and repair or replacement of defective
components. And these tasks are being perfonned according to the rules
and the regulations of ASME Boiler And Pressure Vessel Code: e.g., theNDE concepts and the criteria(Section XD, the NDE procedures(Section
V), the repair and replacement(Sections 111, IX, and XD, etc.. The ISI
planning is to plan the overall scope of the ISI operation, the sequence ofcomponent examinations, and the associated time schedule. The NDT isfor the actual examination of nuclear components using various NDTtechniques such as ultrasonics, Eddy current, X-ray, dye penetrant,
magnetic particle and visual examination. The quality assurance(QA) is toconfirm the accuracy and reliability of the NDT performed by otherinspection tearns. The FMA is to determine whether the detected flaws ofcertain sizes could influence the component integrity during the operation.

If the flaws could indeed influence the component integrity, then, theFMA determines how long the defective components could be safely
operated without repair or replacement. The repair and replacement tasks

in TSS are to recommend the repair and replacement methodology fordefective components that can satisfy the relevant code requirements.

Since the component defects are being found during the ISI, thenecessary repair or replacement tasks can not be incorporated into the ISI

planning. Therefore, it is necessary to make a prompt, sound judgement

at the site for whether the component should be either repaired or
replaced depending on the applicable code requirements. And if it has tobe repaired or replaced, then, it should be determined which repair or
replacement methods should be applied.

the above ISI tasks require the knowledge andthe experience

on NDEtechnology and NDE code requirements in orderto perform thetasks inan accurate and proper way. Nevertheless, in thepast, some of

the tasks

were conducted by non-NDE personnel resulted

m unnecessary

In fact,- I12 -

time delays and expenses to the power plant. Furtherrnore, the recent

development of risk-based ISI technology for nuclear power plantencourages the NDE to establish a new strategy for the ISI application.

To meet such demand, the KEPRI has developed the ISI NDE total
support system(TSS) that can support various ISI tasks in an efficient,reliable and timely manner(Figure 2). Under the TSS, the NDE experts
provide an extended service on the above ISI NDE tasks. Furthermore,
the overall ISI tasks can be performed consistently without the conflicts

in the data or code interpretations amongst many different groups of ISI

personnel. It should be noted, however, the KPERI does not intend to

perform the routinely repeated ISI NDE tasks, but to perform a
supporting role by conducting advanced or special ISI NDE tasks that
require special knowledge and experience on NDE.

Under the TSS, a number of additional NDE-related services are also
provided which include NDE valiadation, NDE consulting, approval from
regulatory agencies(e.g., KINS in Korea), code interpretation, training &qualification, and database management(NDE data and FMA stress data) .

-113-

Figure 2. ISI NDE Total Support System

Key ISI NDE Technology
The ISI site support under the TSS contains a number of key
technologies that are essential for a successful application of the TSS toplant ISI. In the past, the KEPCO has been heavily relying on foreign ISI

service companies due to lack of the key ISI NDE technology.
Consequently, the KEPCO had to manage a great deal of financial
burden. In order to alleviate such difficulties of KEPCO, the KEPRI has

selected 9 technical areas to develop in parallel with the TSS which

include transducer manufacture, NDE systems automation, signal
processing, pattern recognition, fracture mechanics analysis(FMA),
risk-based inservice inspection(RBISD, flawed-specimen fabrication, NDE

modeling, and welding. These technical areas were chosen among thefundarnental NDE techniques that are needed for conducting each of the

major TSS tasks(Figure 3). Brief descriptions of each area were given
below :

-H4-

Specimen Fab.

Welding
NDE Model i ng

Training & Qual .(PDI)

Weldlng - - - -Figure 3. Sources of the

key IS I

NDEtechnology

Although the NDE procedure as shown in Figure 3(within dotted line)
might be more or less different depending on the testing conditions, theNDE procedure typically includes the following tasks: data acquisition by

NDE equipment, (signal or numeric) data processing, feature extraction,
graphical display, pattern recognition, and decision. Although the details ofeach step of the above procedure would be out of the scope of thispaper, we could easily recognize the following 4 essential technologies

contained in the whole procedure: transducer manufacture, NDE systems
automation, signal processing, and pattern recognition:(1) Transducer Manufacture: As the ISI NDE technology became
complex and sophisticated, varieties of transducer types and designs are
required to fabricate. To fulfill such needs, the transducer manufacturing
capability is an absolute necessity.

(2) NDE Systems Automation: The ever increasing complexlty the- 115 -

volume and the number of testing components in recent years require thespeed, the accuracy, and the reliability. of nondestructive examination. To

meet such demand, and to keep pace with the fast growing modern NDE

technology, the NDE systems automation for the nuclear component
examination should be seriously considered.

(3) Signal Processing: Answers for NDE problems are not always
contained in the raw data as acquired. For an example, some flaw
characteristics are extractable not directly from the ultrasonic flaw signal,

but from some other modified data types such as frequency spectrum ortransfer function. Such signal processing techniques as FFT, correlation,
convolution, filtering, to name but a few, are essential for analyzing thecomplicated data types.

(4) Pattern Recognition: When the signal processed NDE data is
insufficient to extract a reliable answer, the data has to be further relied

upon the pattern recognition techniques. Pattern rec ognitions aremathematical and statistical analysis methods that make decisions basedon multiple feature values extracted from the various data types(e.g., Iinear or nonlinear regression. Bayes decision theory, discriminantfunction, etc.). As the complexity of NDE data types grows, the need forpattern recognition becomes more absolute.

(5) Fracture Mechanics Analysis(FMA): the nuclear components having
defects larger than the code-allowable sizes(in terms of length, depth,
width, etc.) should basically be either repaired or replaced. However, thedefective components may continuously be operated for an extended
period of time only if the safe operation of the defective components istheoretically(using FMA) demonstrated. The KE;PRI is in the development

process for the stress analysis and FMA programs, and for establishing

the stress database for major nuclear components that is necessary to
perforrn the FMA.

(6) Risk-Based Inservice Inspection(RBISD: The ASME Code-based ISI-H6-

in nuclear power plant requires to inspect many components having noprevious records(or history) of component failure or defect occurring. The

risk-based ISI(RBISD is therefore to avoid such inefficiency by allowing

utility to inspect only the components whose probability of failure is

ranked high. The component ranking is detennined based on the risk
level calculated as follows:

Risk = (Core Melt PotentiaVPipe Rupture) x (Potential for Pipe Rupture) ----- (1)where, all t.erms indicate the probability of occurring .

The RBISI considerably reduces the components to inspect(A 70~
reduction has been demonstrated by US plants), and hence, reduces theoverall ISI costs as well as the overhaul period. The RBISI is currently

being applied only for the piping s in the United States, and further
applications of the technique to other major components are currently

being considered. The KEPRI had initiated the RBISI development
prograrn in 1997(Figure 4). The KEPRI is planning to apply the RBISI to
actual Korean nuclear power plants in the year 2001.

Domest ic Joint

Research

Col laborat ion wi th

NDEC orWes t i nghouse

KlNS Approval

Note : Currently applied for pipings only in USA

Figure 4. The ApprOach fOr KRBISI-117-

The following 3 technical areas (flawed-specimen fabrication, NDE

modeling, and welding) are related to the KEPRI's Korean Performance
Demonstration Initiative(KPDD program initiated in 1997(Figure 5). In thepast, all ultrasonic inspectors were allowed to inspect nuclear components
if they have been qualified once in any certified training and qualification

institute. However, because detecting and sizing of flaws of complex
shapes(IGSCC, under-clad cracking, etc.) require special skills andexperience, the codes(ASME and CFR in the United States) mandated
that the performance demonstration of all ultrasonic inspectors be

conducted by an integrated qualification process combining the inspector,

the procedure, and the equipment, in order to improve the accuracy andthe reliability of NDE performance on specific ISI task. The KEPRI isplanning to apply the KPDI program to actual Korean nuclear power
plants in the year 2001. Brief descriptions of 3 additional technical areas

were provided next:(7) Flawed-Specimen Fabrication: In order to evaluate the ultrasonicinspector's performance, it is required to prepare test specimens

containing certain number of flaws of certain sizes as described in the

ASME Code Section XI, Appendix VIII. The fabrication offlawed-specimen requires special techniques, and it involves technicalknow-hows and therefore high fabrication cost. Considering the situation,
it would be beneficial to have the specimen fabrication capability in-houseto establish the training and qualification system for KPDI.

(8) NDE Modeling: For the application of performance demonstration

program(KPDD to ISI NDE inspectors, the test specimens that represent
all nuclear components should be prepared. To meet such requirement, agreat number of specimens of an extremely high cost has to befabricated. However, the use of NDE modeling can help reduce thenumber of the specimens by theoretically(NDE modeling) demonstrating
the equivalency of similar inspection situations. The NDE modeling canalso help avoid unnecessary re-training and qualification for the similar

inspection situations. In addition, the NDE modeling can be utilized for an- I18 -

effective education for the theoretical wave propagation(9) Welding: The welding techniques are continuously needed tofabricate the test specimens for the inspector's performance demonstrationand to develop optimum repair and replacement technology for defective
com ponents.

col laborat lon wi th col laborat ion with

European
EPRI
NDEC
Manutacturer
FabrSpec
i cat ionimen( Iowa state
Univ.
for
Mode I i ng)

KINS Approval
col laborat ion with
Foreign Ut i I ity

Note : currentfy applied for pipings, RPv, and studs & Bolts in USA.

Figure 5. The Approach for KPDI

Summary
Structural integrity of nuclear components is important for a safe
operation of nuclear power plant. Therefore, the structural integrity should

always be maintained by reliable, periodic inservice inspections. Since thesafe operation and the inservice inspection of the nuclear power plants

are under the plant owner's responsibility, KEPRI, the R&D division of

KE;PCO, has established an integrated ISI NDE total support
system(TSS), primarily focused on the inservice inspection of nuclear-119-power plants. And the KEPRI has initiated a long-tenn plan for thedevelopment of key ISI NDE technology that is necessary for an efficient
on-site support. Although KEPRI took the lead in this effort to establish
the TSS, the KEPRI intends to collaborate with industry, universities, andresearch institutes throughout the development process in order topromote the entire Korean ISI NDE technology.

Ref erences
1 . Plant Operatron Statistics Korea Electnc Power Company 1997

- 120 -

III-3

Core Shroud Replacement of Fukushima-Daiichi Unit #3

Jun Matsumoto, Nuclear Power Plant Management Dept.
Tokyo Electric Power Company, Japan

Core Shroud Replacement of Fnkushima-Daiichi Unit #3

J. MATSUMOTOTokyo Electric Power Companyl - I -3 Uchisaiwai-cho, Chiyoda-ku Tokyo, Japan

Abstract
The core shroud replacement of a Boiling Water Reactor (B WR) is now

in progress at Fukushima-Daiichi Unit # 3(1F3) of the Tolyo ElectricPower Company (TEPCO) in Japan. The core shroud and other coreinterna/ components made of type 304 stainless stee/ (SS) are replacedwith the ones made of low carbon ope 316L SS to improve Intergranular
Stress Corrosion Cracking (IGSCC) resistance. This project is the first

application of the replacement procedure developed for the welded coreshroud, and employs various advanced technologies. The outline of thecore shroud replacement project and applied technologies are discussed inthis paper.

1. Introduction
Since the core shroud cracking was found, the Japanese BWR owners and the plant

manufacturers have conducted several R&D programs to establish the countermeasuresfor core shroud cracking. These countermeasures are repair and mitigation
technologies including replacement, bracket reinforcement, tie rod reinforcement,

hydrogen water chemistry (HWC), water jet peening, shot peening, and noble metal

technologies. When the crack was found in Fukushima-Daiichi Unit #2 ( IF2), the coreshroud was reinforced by the brackets which had been studied as one of the repair

concepts.
The countermeasures for the actual plants should be carefully selected considering
effect and applicability of the countermeasures, uncertainty of the developed

technologies, estimated cost (tool fabrication and implementation cost, inspection costthrough the plant life etc.), estimated schedule, possibility of cracking, political- 121 -

environment, and so on. Based on the evaluation ofthese influential elements, TEPCO
decided to replace the core shrouds made of 304SS which has relatively higher potential

ofIGSCC. Among 1 7 BWR plants owned by TEPCO, four plants have the coreshrouds made of 304SS. In accordance with the long term outage schedule, I F3 wasselected as the first plant to replace the core shroud with the one made of 3 1 6L SS which

has much less potential of IGSCC.
In addition to the core shroud, the majority of the internal components made of 304SS
are replaced in the IF3 project. Those are (1) core shroud, (2) top guide, (3) core plate,(4) core spray spargers, (5) feed water spargers, (6) jet pumps, (7) DP/LC pipe, (8) ICM

guide tubes, and internal pipes and nozzle safe ends connected to these components asshown in Figure I . Prior to the application to IF3, integrity and reliability of thereplacement process were demonstrated by the full scale mock-up test, which was

conducted by the Nuclear Power Engineering Corporation (NUPEC).

2. Replacement Process
The replacement sequence is shown in Figure 2. Afier the removal of all fuels anddetachable components, inside of the pressure vessel and the remaining internal

components were chemically decontaminated. Subsequently, most ofthe removals are
conducted by means ofElectrical Discharge Machining (EDM) process under water.
Following the removal of all components to be replaced, the new components were
installed in reverse order ofthe removal. Most of the installation process including

welding was performed in a dry condition to acquire high quality of the weld. The
followings are the major technologies applied to the core shroud replacement process.

2.1 Chemical Decontamination
Prior to cutting the core shroud, the chemical decontamination was conducted. The

Chemical Oxidation Reduction Decontamination (CORD) process provides a highDecontamination Factor (DF), and generates a small amount ofwaste in comparison toother decontamination processes. Decontamination solution was circulated by thePrimary Loop Recirculation (PLR) pumps, and the dissolved active metal was captured
by ion exchange resins (Figure 3). In the core shroud replacement project for IF3, theoxidation and reduction process was repeated three times in a week. The volume ofwaste was approximately 5m3, and the DF measured under water was approximately 40
at the bottom of the RPV .

- 122 -

2.2 Core Shroud Cutting
The old core shroud was cut remotely under water by EDM process. Four EDMcutting carriages are mounted on the circular track assembly and run individually. Each

EDM-cutting carriage has a rotating graphite electrode as shown in Figure 4. A swarf
collection hood is attached to the EDM-cutting carriage. After this primary cutting, theold core shroud was transferred under water to the Dryer Separator Storage Pool (DSP)
for slicing. The height ofthe old core shroud in the DSP should be less than 3.5m toassure the shielding effect ofwater. The old core shroud was cut circunferentially at
two elevations so that the height of each old core shroud was less than 3 . 5m.

After removal of lower half of the old core shroud, the upper face of the remaining

shroud support is machined to acquire a sufficient flatness for the welding. Figure 5

shows the machining tool which uses a rotational milling head.

2.3 In-vessel Shielding
Afier the cutting and removal process under the water, the in-vessel shielding was set

up to reduce the dose from the radioactive RPV wall. Reduction of radiation dose level
during the installation process is important because the new components are welded in a

dry condition. Figure 6 shows the shielding panels hanging from the RPV flange. The
shielding panels have windows to allow welding the riser pipe and the riser brace without

removing the shielding panels.

2.4 Core Shroud Installation
Although the design of the new core shroud is improved in the material and welding
to increase IGSCC resistance, the basic structural geometry was maintained to assure thesame plant performance.
The overall structure of the core shroud in IF3 is shown in Figure 7. The left sideshows a cross section ofthe new core shroud, and the old core shroud is shown on theright side. The number of welds is reduced because it directly influences the probabilityof cracking. The new core shroud consists ofthree forged cylinders (3 1 6L SS) with novertical weld, and a lower ring (alloy 600) with two vertical welds. Totally, sevencircumferential welds are decreased to four welds, and twenty-four vertical welds are

reduced to two welds.
Figure 8 is the illustration ofthe weld joint between the lower face of the core shroud

and the upper face of the remaining shroud support. The lower face of the new coreshroud meets the upper face of the remaining shroud support at a narrow angle to reduce- 123 -

welding heat input and to shorten welding time in comparison to the old core shroud.
In the old core shroud, the joint was welded from inside with the backing ring which

formed a crevice. In order to avoid a crevice in the weld joint and to get a smoothconfiguration ofthe outside surface, the new core shroud was welded from both sides.

Figure 9 shows the bead sequence ofthe weld joint. Tungsten Inert Gas (TIG)
welding was applied. First, the outer diameter (OD) side was welded, and then the

inner diameter (ID) side was welded. Since the distance between the core shroud andthe jet pump difftLser is only 70mm at the joint elevation, the OD welding machine is thinto fit the space, and was mounted on the OD track prior to the shroud installation asshown in Figure I O. The filler wire was also attached on the outside ofthe new coreshroud prior to the shroud installation in order to save the space for wire feeding system.

The ID welding machines were attached manually after the new shroud installation, andran on the ID track remotely (Figure 1 1 ).

At the development stage of this welding process, the weld joint was evaluated byvisual test, penetration test, radiographic test, microstructure, metallography, side bend

test, crevice bent beam test (CBB), tensile bond strength test, and measurement ofhardness and residual stress. Based on the evaluation, the limit ofthe essential variables

for this welding method was established.

2.5 Alignment of the Components
It is important to align the core shroud, the core plate, and the top guide with the

CRD housings to secure appropriate motion ofthe control rods. Figure 12 shows thealignment tool for the new core shroud, core plate and top guide. A Iaser targetingsystem is set in the gas purged tube of the alignment tool. The new core plate wasaligned with respect to existing CRD housings as shown in Figure 1 3 . Bottom of thealignment tools were set on the top of CRD housings, and top ofthe alignment tools
were fixed in the core plate master holes. The new core plate was aligned checking thelocation of the core plate master holes and the CRD housings. The new core shroud
and the top guide are aligned similarly.

3. Waste Management
The removed core shroud and other components were sliced and put into thecontainers in the DSP. The underwater plasma arc cutting was employed for the slicing
process because of its fast cutting speed. The stuffed containers were decontaminated- 124 -

and transferred to the site bunker pool and the dry storage facility in the site.

Althoughall of the waste is low level radioactive waste, the site bunker pool stores moreradioactive waste.

4. Conclusion
TEPCO is replacing the core shroud of IF3 to improve IGSCC resistance. The
shroud replacement is considered to be one of the most comprehensive countermeasures
to improve IGSCC resistance for all the welds ofthe core shroud. As of April 1998,
the new core shroud has already been installed properly, and the project is underrestoration process of other internal components. Subsequently, the core shroud

replacement will be performed in several Japanese BWR plants.

- 125 -

Feed Water Sparger

Top GuideCore Shroud
Core Spray Sparger

Jet Pump
Core Plate

DP/LC Pipe

ICM Guide Tube

Figure 1

Replaced Components- 126 -

(D

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i

l

Ht~

~]

Remove Fuel,

Chemical Decontamination

Core Componcnts

Figure 2 ( I )Remove Upper Shroud,

Remove Lower Shroud,

Feed Water Spargers,

Core Plate, ICM Guide Tubcs,

Top Guide

DP/LC Pipe

Replacement Sequence

R

R

O

R

l

/

H'
t~D
co

/

l

'!.

li~ h

Remove Jet Pumps

Install Jet Pumps

Figure 2 (2)Install Core Shroud

Replacement Sequence

Weld Corc Shroud

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l

H'
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l

t

l

Instal] DP/LC Pipe,

Install Core Plate, Top Guide,

ICM Guidc Tubcs

Feed Water Spargcrs

Figure 2 (3)Replacement Scquence

Restore Fuel, Core Components

RPVCore Shroud

Gas Treatment L.

UV System

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Spray Nozzlel

EDM Cutting Machine

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EDM Electrode

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Ion Exchange Resin

Figure 3

PLR Pump

Heater COOler

Chemical Decontamination System

Figure 4 EDM Cutting Machine

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Figure 5 Shroud Support Machining Tool

Figure 6

In-Vessel Shielding

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RPV

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Figure 8 Old and New Design of Joint Between
Weld Line

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Figure 9 Weld Sequence
Figure 7

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Figure 1 2 Alignment TOOl

Figure 1 3Core Plate Alignment

III-4The Evaluation of Erosion-Corrosion Problems of PWR Carbon
Steel Piping in Taiwan

Kuen Ting, Department of Nuclear Regulation Atomic Energy
Council, Taiwan, Republic of China

Yin Pang Ma, Institute of Nuclear Energy Research Atomic

Energy Council, Taiwan, Republic of China

THE EVALUATION OF EROSION/CORROSION PROBLEMSIN TAIWAN PWR CARBON STEEL PIPING

Kuen Ting
Department of Nuclear Regulation
Atomic Energy Council
Republic of China

Yin Pang Ma
Institute of Nuclear Energy Research

Atomic Energy Council
Republic of China

ABSTRACT
Taiwan PWR Nuclear Power Plant Units I and 2 implemented themeasurements of the wall thinning of the carbon steel piping under the request of
regulation authority to prevent the events due to the erosion/corrosion since 1 989. At

the frrst, the licensee established the comprehensive inspection program by itself.

Over 2000 components were inspected per each unit and 300-500 pipe components
were examined by ultrasonio testing per soheduled outage. The simple predictive
methodology determined the operability of each individual piping oomponent in thenext fuel Gycle. Based on the inspection results, the susceptible pipe components

were established. The implementation of the effective correction management andimproved inspection program can improve the safety, as well as the efficiency, of
long-tenn reactor operations.

INTRODUCTION
Erosion/corrosion is a piping degradation process that induces the pipe wall
thiming. The oxide layer on the carbon steel pipe is dissolved by single-phase or two-phase flow in the nuclear power plant[1]. In 1 989, the regulatory authority requested

the licensee of Taiwan nuclear power plants to develop their own inspection program

to prevent the failure of piping due to erosion/corrosion damage. The licensee
established a full scope of pipe wall thickness database through "Pipe Wall Thiming

Measurement•Program"[2]. Over 2000 pipe component
and 300-500 pipe components were examined by ultrasonic testing per scheduled
outage. Based on the inspection results and the analytical method, the evaluation of

the susceptible pipe components in the licensee inspection program was establishedduring each scheduled outage. This paper will discuss the implementation of this- 135 -

effeGtive management and the inspection program. The efforts can improve the safety,
as well as the efficiency, of long-term reactor operations.

INSPECTION PROGRAM
In 1 989, after a severe piping rupture at Surry nuclear plant, the regulatory
authority requested the licensee to develop the plant specific-monitoring program toprevent the failure of piping due to erosion-corrosion damage. The licensee, Taiwan

Power Company (TPC), established the comprehensive inspection program tomeasure the thickness of most of the carbon steel piping components in the balance ofthe plant. The inspeoted oomponents consist of straight pipes, elbows, tees andexpanders/reducers in the feedwater, condensate, extraction, heater drain and reheaterdrain systems. The detailed configurations for measured points in each typicalcomponent are displayed as Figure I .

The evaluation method of the inspection program is based on the measured data.
The thinning due to erosionfcorrosion wear T., oan be caloulated as
T,. = Tl ~ T,orT. - T. (ifno inspection data for reference)

T,. = (previous measured thickness)-(present measured thickness)

In which Tl and T. are the maxirnum and minimum measured thickness,
respeotively. T~ is the nominal thickness. The wear rate can be defined as

W. = T.. IH

where H is the operating hours between two inspection intervals. The allowable
wear thickness T. is

T. =T. -T~
where T~ is the required thickness from the stress report or by analysis.
Thus, the remaining time of operation , L. ,can be estimated as

L. = T. /W.
The allowable number of remaining fuel cycles, I,al , can then be predieted as
lrd = L. l(KO.)

next fuel Gycle.

The tbtal number of the selected pipe components was greater than two
thousand per PWR unit. However, the oapacity ofrealistic ultrasonic testing was only

three to five hundred pipe components per outage. So, one PWR plant needs at least
seven outages to examine all selected oomponents and build up the complete databaseof the wall thiming.

- 136 -OBSERVATIONS OF THE INSPECTION
The measured data from four outages have been studied according to the type of

the pipe components, such as 90" elbow, 45" elbow, reducer, tee and straight pipe.
Figure 2 displays the occurrences ofthe thinnest thickness versus measured points for

all inspeGted 90" elbows. The position B5 that is the outward bent of the elbow had
the majority of the occurrences of the thinnest thiGkness. Ahnost 700/0 of all inspected

elbows show that this region is easily erosive owing to the change of the flow
direction. This phenomenon is also found in the 45" elbow inspection as shown in
Figure 3 . The majority of the worst thinning occurs at the location B3 which is theoutward bent of the elbow.
For the erosion/corrosion of the reducer, the thinnest thickness locates at the twoends Al, A4, Dl and D4 as shown in Figure 4. The fraction of these locations isapproximately 90010.
The distribution of the thiunest thickness is uniform for tee components in Figure

5. According to the experience, B and C sectors are not easily subjected toerosion/corrosion. However, Figure 5 revealed that B and C sectors with 380/0 are thepositions of the thinnest thickness.

From Figure 6 presents the statistical distribution of the measured data forstraight pipe. The inlet position is the susceptible location for erosionfcorrosion as
antioipated.

THE EVALUATION OF THE POTENTIAL RISK OF THE CARBON STEELPIPING UNDER EROSION/CORROSION
The ultrasonic inspection program of earbon steel piping due to erosionfcorrosion was

oomprehensive. An analytic method to evaluate the potential risk of the carbon steel

piping subjected to erosion/corrosion is necessary. Berge's model[3] assumed thesoluble iron species production and mass transfer effect on the erosion corrosion rate.
Thus, the total erosion/corrosion rate Vc is expressed as:

V KIK2 (C,q - C.)
c ~ Kl +K2

where C~ is the soluble ferrous ion concentration at the bulk water; Kl is thereaotion constant which depends on fluid velocity and water temperature; K2 is masstransfer coeffilcient which depends on the Sherwood number Sh = al Re"' Sc"= , Re

and Sc are the Renolds number and Schmidt number, respectively; al ' a2 and a3- 137 -

are constants which depend on the pipe component geometry. C,q is the equilibriurn
soluble ferrous ion conoentration at the oxide layer which depends on the pH value
and the coolant temperature. In present work, C,q is expressed as

C.q a exp[~(T -a5)2 J +a7
a6
where these constants are dependent of pH values.

The calculation results of the present model are in good agreement with thereferenced data from the previous experimental works [4-6] as shown in Figure 7, 8
and 9 . The parameters studied here are simulation to the realistic environments of
PWR nuolear power plants. The values of pH, the fluid temperature and fluid velocity

in PWR are fanged from 8.9 to 9.5, 300"C to 500"C and Im/sec to 4m/sec. The
presented model also shows the good relations with the measured data of the PWR
realistic oonditions ranging from 120"C to 230"C and the velocity over Im/s.

Based on the predicted analysis, the carbon steel piping system, such ascondensate system, feedwater system, main steam system, extraction steam systemand drain system, are oategorized into 36 groups to evaluate the potential relative riskof the pipe systems of the PWR plant under the attack of erosion/corrosion. The frrst

column of Table I shows the categorized groups which are divided by the criticalcomponent. For example, the critical components of the feedwater system consider as

feedwater pump, HP heater I and 2, steam generator. Thus, the feedwater system isoategorized as three groups to evaluate the behaviors with erosion/oorrosion. The pipe

diameter, flow rate, steam quality nominal thickness and allowable thiekness are theinput parameter for the prediction of the rate of erosion/corrosion and the life of the

carbon steel pipe components as listed in the last two column of Table I . The
superscripts "*" notes "the not-used-normally" pipe components which are not usedduring normal operation. The pipes from the 3'd stage feedwater heater dump to thecondenser are the most dangerous pipes for not-used-normally pipes, if the upstream
level oontrol valves of the pipe components are leaking, the water flashing in the level

control valves due to the high pressure gradient across the valves may cause the wall

thickness of the down stream pipe components to thin and exceed the limitation
within one fuel cycle[7].

RESULTS AND DISCUSSIONS
Based on the prediction as mentioned above, the damage of the two-phase pipe

- 138 -

components is worse than the single-phase pipe components. The extraction lines
from the high pressure turbine to the frrst stage feedwater heaters in which the steam
quality is ranges from 880/0 to 950/0 were the high risk of the erosion/oorrosion. The
pipe components in the following pipes are also highly susceptible, suoh as:(1). from the 4th stage feedwater heaters to the 3'd stage feedwater heaters ,(2), from the '3'd stage feedwater heaters to feedwater pumps,G). from the drain pumps to the feedwater pumps,(4). from the 2~d stage feedwater heaters to the drain tank,(~)• from the drain tank to the drain pumps,(~). from the 2~d reheater to the first feedwater heaters.

Table 2 desGribes the records of the replacements of the pipe components during
each outage. The extraction lines and the exhaust lines are the highly replacement rate
whieh are consistent with the predicted work as shown in Table I .

CONCLUSION
The actual operating conditions in the nuclear power plants cannot be effectively

eontrolled as stringently as in the laboratory, and the analyiical predictions andinspection program are not so reliable as anticipated. To avoid the erosion/eorrosion

damage, the current inspection program should be improved by a proper analyiicalmethod and engineering j udgment.

The present work provides a pipe erosion/corrosion inspeetion program andanalytical method to assist the relevant engineers in making engineering judgnlent andto find out the susceptible pipe components and to evaluate the relative potential riskof the pipe components . The results are helpful for the engineers to renew and audit
the monitoring prioritization of the pipe components.

REFERENCE
l. U. S. Nuclear Regulatory Commission Bulletin 87-0 1 , "Thinning of pipe walls in
nuclear power plants," July 9, 1 987.

2. Taiwan Power Company, Nuclear power plant pipe wall monitoring program,

NED-NPS-OOOl, R2, TAIPOW, June, 1989.
3. J.P. Berge et al., Effect of chemistry on corrosion-erosion of steels in water andwet steam, Proceeding of 2~d Conferenee on Water Chemistry, BNES, 1 980.

4. ASM Intemational Handbook Committee, Metals Handbook, 9th edition, Vol.13,
Corrosion.

5. J.P. Gof~m. Thickness measurements of pipes submitted to erosion and corrosion

problems in the steam, feedwater and condensate systems of the Doel I and 2
power plants, Intemational Atomic Energy Agency (IAEA) Specialist Meeting on- 139 -

ErosionfCorrosion, September 1 2-14, Vienna, Austria, 1 988.
6.

M. Bouchacourt, Flow assisted corrosion in power plants-part I : The EDF

research program, Intemational Atomic Energy Agency (IAEA) Specialist
Meeting on Erosion/Corrosion. September 1 2-14, Vienna. Ausina, 1 988.
7.

H.J. Chao and Y.P. Ma, User's manual for nuclear power plant pipe
erosion/corrosion data base system, INER- 1 2 1 5, May, 1 994.

- 140 -

A
I~,

e DJ
t9
8 f :"

5A,, r ~/'~
i:r~~2i-'---:"'t"'

tI

' ih
-~i l' /'/It:'i l' :t~~: ,• ''L\~/\ I (\-~\*;:i4i//; .It'- :"'t.. ..
'f !; 9'/:~~~~~/\ \(/~~{~ll,'!~. / //7/'~~... : ( ~, '/;'

A(a) 90" elbow

~.

~'-

(b) 45" elbow

IA-- "t!Lc4 rt/' *',. , \"I)i~ ;c-~ ~

i

I.\ t

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I(~ ~." /\/t~4.
~__"~:L~] f~le~~:. /~' -'1:

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I~= ; JLi" '~i' 'AI _ y?r\:..~::'~~ !~"'~f~~1'~~~,\" /\fs-pr '\1- ' -(c) reduoer'~:'. ~!\.!
\~*l(d) tee

:1'
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oo

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Figure I . Detailed configurations for measured points in inspection components- 141 -

.LJ3. :

if~-. : - ;-?
l,.
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5i

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~ 40
o
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20

o A1 A2A3A4A5 ~ I B2B3B'4B5B6 B7 B8B9d I C2C3D I D2D3

POSITION
Figure 2.

Statistical analysis ofthe measured thinnest position for 90" elbow

30
25 ~

Oz 20 ~Lu :O
izuJ15•~,~
cc =
D ::
O
OIC1
O=

O

A1 A2 A3 ~l B2 B3 B4 B5 C1 C2 C3 D1 D2 D3•POSITION
Figure 3 .

Statistical analysis of the measured thimest position for- 142 -

45" elbow

25

20

d
zUJ 15 ~
o
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D10
o
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5-

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Figure 4.

A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4 D1 D2 D3 D4POSITION

Statistioal analysis of the measured thinnest position for reducer
35 -:

30 +

25 ~

c; ;

z:
z i :15 _•~
UJ
CCD
O :i
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A1 A2 A3 B1 B2 B3 C1 C2 C3 D1 D2 D3 D4 D5A1 IA12~l3

POSITION
Figure 5.

Statistical analysis for the measured thimest position for tee- 143 -

16
14
12
C;

Z 10
LLJ'

O

~8
86
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4

2
O A1 A2A3 A4 A5 ~l B2 B3B4 B5C1 C2d3C4C5D1 D2 D3D4D5POSITION
Figure 6.

Statistical analysis of the measured thinnest position for straight pipe

- 144 -

AB
RHTR
HP

CJ

Lp

RE:lR

LP

SPR
S/G

AF

AF

AD

AD

~

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CONDENSER

~

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J~
Ol

ssn

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l

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Fwr

AE)

AE

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DP

AB Main steam system
AC Main turbine system
AD Condensate system
AE Feedwater system

AD

3

AD

5

AD

6

sPE

~

AF

AF

AF

SPE Steam Packing Exilauster

DP Drain Pump

AF Feedwater heater extraction and drairL system

Figure 7.

AD 4

The sketch of the balance of plant for PWR nuclear power plant

AF

,.

MODELPREDICTIONOOODO V = 1m/s
m/s
AAAAA

x

oooooV = 3 m/s
x**x*V = 4 m/sx

pH = 9.. o

X

x
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Figure 8. Comparison of wear rate between calculation and experimental data versus temperature for low velocity(Experimental data from reference[6]).~II 'l.

400-~:30035 m/sV=
pH = 7.

PREDICTED MODEL

O
O
O
O
O

MEASUREMENT
OODOD

40 pp bo\=E200
E

~

D

LJ

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6Qo .

TEMPERATURE (F)Figure 9.

Comparison of wear rate between calculation and experimental data versus temperature for high velocity[Experimental dada from referenee(5)]

2000

~
Oo 1500~.

T

PREDICTED MODELpH

MEASUREMENT
DOOOD

O
O
O

02

180 C
7.0

5 ppb

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d~
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D

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10

20

30

40

FLOWVELOCITY(m/s)

Comparison of wear rate between calculation and experimental data versus flow velocity

50

60

Table l Risk Category of Piping due to Erosion/Corrosion

,-,

~

~o

Table 2 Replacement Records of Piping due to ErosionfCorrosion

l

~:;

o
I

AB : Main steam system
AC : Main steam drain system

AD : Condensate system
AE : Feedwater system
AF : Feedwater heater extraction and drain system

FA : Auxiliary main steam system

2 . EOC :End of CycleSession IV

Prpmg and LBBIV- lDevelopment of Modified Piping Evaluation Diagram for LBB
Ap plication

Y.J.Kim, Y.Z.Lee & N.S.Huh, Dept. of MechanicalEngineering, Sung Kyun Kwan Univ., Suwon, Korea
C.R.Pyo, Dept.of Mechanical Engineering, Indnk college, Seoul, Korea
J.S.Yang, Korea Electric Power Research Institute, Daejon, Korea

Development of Modified Piping Evaluation Diagramfor LBB Application to Korean Next Generation Reactor

Y J Kim', Y.Z. Lee', N.S. Huh', C.R Pyob J S Yang

'School of Mechanical Engineering, Sung Kyun Kwan Univ., Suwon, Korea
bDept. of Mechanical Engineering, Induk Inst. of Technology, Seoul, Korea'Korea Electric Power Research Institute, Daejon, Korea

Recently, the Piping Evaluation Diagram (PED) is accepted in nuclear industry for an
efficient application of Leak-Before-Break (LBB) concept to piping system at an initial
piping design stage. The objective of this paper is to develop the modified PED which canaccount for the variation of the material properties of the PED development stage and thoseof the assembly stage. For this purpose, a parametric study was performed to investigate theeffect of stress-strain curve on the detectable leakage crack length and the effect of fracture
resistance curve on the LBB allowable load. Finite element analyses were also performed toinvestigate the effect of stress-strain curve on the LBB allowable load. Finally a modified
PED was proposed as a function of crack length and the allowable safe shutdown earthquake
load. The LBB analyses based on the modified PED are in good agreement with those basedon the traditional PED. By adopting the modified PED, the variation of material properties
can be considered in the LBB analysis and the computing times required for the applicationof LBB during the design process can be considerably reduced.

1. INTRODUCTION
The Leak-Before-Break (LBB) concept has been accepted as a justifiable approach fordesigning nuclear piping systems with the elimination of massive protective measures due tothe pipe rupture design requirement. If the LBB concept is applicable to nuclear piping
design, the construction cost can be significantly reduced by eliminating unnecessary pipe
whip restraints and jet impingement devices. In order to apply the LBB concept to piping
system, the nuclear piping design should meet all the criteria specified in NUFLEG 1061 , Vol.

3Ll] and Standard Review Plan 3.6.3[2]. Current LBB analyses on nuclear piping system arebased on the plant-specific loads due to the geometry of the piping system and the operating
condition of concerned piping system. Therefore, if the analysis results do not meet the LBB
criteria, all the piping routing should be changed until the results are satisfactory. In
performing LBB evaluation, highly specialized knowledge in fracture mechanics is required
and lengthy period of time is required due to its complicated procedures and methods.
Recently, numerous studies have been reported on the applicability of LBB during theinitial piping design stage. The "Allowable Load Window" using the maximum and theminimum
moment of each piping was proposed by N
Evaluation Dragram)" usmg the Normal OPeratron (NOP) Ioad and the Safe Shutdown

Earthquake (SSE) Ioad was also proposed by Fabi et al.[4- 151 --

"BAC (Boundary Analysis Curve)" using the stress at the normal operating condition and the

maximum stress. Y.J.Kim et al.[6] developed the "PEDs" using the NOP Ioad and the SSE
load and applied for the safety injection and the shutdown cooling line in Korean Next
Generation Reactor (KNGR). These diagrams were developed prior to the piping design andcan be used for the evaluation of critical points for the LBB requirements in the piping
system. Thus the iterative process of design to modification regarding the LBB applicability
can be eliminated.
According to the current LBB criteria, the final LBB evaluation should be performed using

the archival material properties. However, the PED may have to be modified to account forthe difference between the material properties of the PED development stage and those of theassembly stage. The objective of this paper is to develop a modified PED which can accountfor the variation of the material properties and apply them to the safety injection line and the

shutdown cooling line in KNGR.

2. DEVELOPMENT OF NUCLEAR PIPING MATERIAL DATABASE
NUclear Plping Material database (NUPIM database) was constructed prior to thedevelopment of the modified PED. The objective ofNUPIM database is to obtain the material
properties such as stress-strain curves and fracture resistance curves required for the modified

PED. In designing nuclear piping systems, a plenty of material properties should beconsidered, and thus the development of NUPIM database which covers about 500 test data
was essential.

Searching a material property in NUPIM database can be easily done owing to the hybrid

database which is connected to the DBMS (DataBase Management System). This DBMS
searches the selected item from the stored material properties. The material properties stored

in database was obtained from domestic material test programs such as YGN 3&4[7], UCN
3&4[8] and the dynamic test for carbon steel[9]-Search index of database consists of material name, welding method, yield stress, pipe
thickness, temperature, pipe size et al. In order to provide various material properties, astatistic calculation module with plenty of raw data has been attached to the database.
Therefore, not only the material raw data but also pre-processed second hand material data
can be easily obtained using NUPIM database.

3. PROCEDURE FOR DEVELOPING MODIFIED PED
3 . I General procedure

The general methodology and criteria used for developing the modified PED are based on
NUREG I 061 , Vol. 3 and SRP 3.6.3. The material properties required for the development of
the modified PED were obtained from NUPIM database which consists of previous domestic
LBB test results. Generally, the worst case material properties (worst stress-strain curve andworst fracture toughness) are used in the analysis for ensuring the conservatism of the LBB
analysis. In this paper, an additional margin for the consideration of archival material
properties, which are usually obtained at the fmal stage of the design process, was added byapplying three different stress-strain curves in the LBB analysis. Different fracture resistancecurves considering archival material properties were also used in the analysis.

Fig. I illustrates the analysis procedure applied for the modified PED development. In the- 152 -

first step of the analysis, each piping was grouped by system, material, size, and geometry. In

the second step, the NOP Ioad range was established. In the third step, the Detectable
Leakage Crack (DLC) Iength was calculated. In the fourth step, the applied J-integral for the

DLC and twice the DLC Iength was computed. In the fifih step, from the applied J-integral
and the material J-integral, J/T analyses were performed to find the instability point of thecrack. In the sixth step, the SSE Ioad range was calculated in accordance with the NUREG
l 061, Vol. 3. Finally, the modified PED was developed as a filnction of crack length (theDLC and twice the DLC) and the allowable SSE Ioad.

Fig. I Procedure for development of the modified PED

3.2 Grouping for LBB analysis

The LBB analyses were performed on the safety injection line and the shutdown coolingline in KNGR, which are classified by material, size and operating temperature as listed inTable I . The minimum and the maximum NOP Ioad ranges were set as wide as possible toapply to the KNGR, based on the actual load observed from the currently operating nuclear
power plants.
3.3 Analysis of detectable leakage crack- 153 -

For a given NOP Ioad level, the DLC Iength which produces the I O gpm leak rate wascalculated by PICEP code[1 O] based on the NOP Ioad and three different stress-strain curves;

namely upper bound, best-frt, and lower bound. In order to assess the crack stability, two
steps were involved. One is to assess the stability of a crack length equal to twice the DLC
under (NOP+SSE) Ioad, and the other is to assess the stability of a crack length equal to the

DLC under ~/~ (NOP+SSE) Ioad. Three dimensional elastic=Plastic finite element analyses

were carried out for both the DLC and twice the DLC using the commercial fmite element
program, ABAQUS[1 1]. Considering the symmetric condition, only one quarter of the pipe
was modeled as shown in Fig. 2. A four-point bending condition was applied to create tensile
crack opening mode at the crack tip area.

Table lGrouping for characteristics of each system

1 st Symmetry

)~

~c

ll

2nc sym~ety

4

./ \,

= C~aeX2~e Symmetry \\Fig. 2 Three dimensional mesh design for finite element analysis- 154 -

3 .4 Crack stability evaluation

For the crack stability evaluation, the J-integral / Tearing modulus (J/T) method was used.

For a stable crack growih, the material's tearing modulus (Tmat. ) must be greater than the
tearing modulus (Tapp. ) obtained from the applied load. The instability point may be foundby plotting Japp. versus Tapp, and Jmat. versus Tmat. on a single figure. The material J-R
curves can be obtained as following from the J-R data for weld metals.

Jmat' =Cl( )Aa2 C(1)And differentiating Eqn. ( I ) with the crack length, the material tearing modulus can beobtained as follows:

dJ= CIC2 (Aa) C2 -l
da mat .

(2)

In this paper, various Cl and C2 Values were used for the development of the modified
PED. The J/T curves for applied load were obtained as follows. In the first step, from theresult of fmite element analysis, the applied J-integrals for crack lengths of a - 6 , a , anda + 6 as well as 2a - 6 , 2a , and 2a + 6 were determined for the applied load. The crack
length increment, 6 , was assumed as I Oo/o of crack length. In the second step, an equation for

the J was postulated as a function of crack length a as following.
()

J = C a2 + c2a + c3

app. l(3)

And also by differentiating Eqn. (3) with crack length, the applied tearing modulus was
obtained as shown in the following equation.

dJ
da app.

2c a + c2(4)

Finally, the J/T analyses were performed to find the instability point of the crack by applying

following equation.

dJapp dJmat (= Tmat )Fig. 3 shows a typical J/T evaluation diagram.

3.5 SSE Ioad
The SSE Ioad for the modified PED was calculated in accordance with NUREG I 061 , Vol.
3[1]. For the DLC Iength, safety factor of J~ was applied on the load as following.

Mtota/ =~/~[
M + MSSE + Metc I 'NOP- 155 -

(5)

( Mtotal)
SSE Load = MSSE + Metc
= 1/~ ~MNOP(6)

And for twice the DLC, safety factor was not applied on the load as following.

Mtotal-[~ MNOP + MSSE + Metc]'

SSE Load -(- MSSE + Metc)= Mtotal ~ MNOP

Material J/T~,;1,

Applied J/T,
J inst.

Tcaling Modutus
Fig. 3 J/T evaluation diagram

4. VERIFICATION OF ANALYSIS METHOD AND FE MODEL
The proposed modified PED method depends on the finite element analysis. Therefore, thefinite element analysis procedure involved in this method should be appropriate for piping
analysis. In order to verify the finite element analysis procedure, the modified PED method
was applied to a piping system and compared with the experimental data.
4. I Fracture test for through-wall circumferential cracked pipe

An extensive experimental program for the fracture evaluation of stainless steel flux
welded pipe was conducted by BattelleL12]. Seven full-scale pipes (four pipes with throughwall circumferential cracks and three pipes with circumferential internal surface cracks) were

tested at 288'C considering the nuclear piping operational condition. The experiment covered

three different pipe outer diameters of 168.3mm, 413mm, and 71 1 .2mm. For each pipe, thecrack initiation load and the maximum load were measured. One of the seven experimental
data was selected for the verification of the proposed modified PED method. Fig. 4 shows theschematic illustration of the piping test, and the specific dimensions are listed in Table 2. The

loading condition was four point bending, and the internal pressure was not considered.
A finite element model was designed to simulate the Battelle pipe test. The crack initiation- 156 -

(7)

load and the maximum load were calculated from the fmite element analysis, and comparedwith those from the experimental data. Fig. 5 and Fig. 6 show the stress-strain curve and thefracture resistance curve for the pipe used in the fracture test, respectively. In order toevaluate the effect of planform size on the analysis result, the fracture resistance curves forthe three different planform sizes of IT, 3T and 9.5T CT specimen were applied, respectively.
The crack initiation load was predicted based on the Jlc values which are listed in Table 3 .

Table 2 Summary of pipe geometry and experimental result

l 95

380
Table 3 Results of J-R tests for SA358 TP304 stainless steel(a) The specimen thickness.
(b) The specimen width.

(c)The side-grooving percentage.

PiPe s~~~~~ef~~~s~~~tess$t~~~eti~~~Hlj
S fQet,Fig. 4 Schematic of test frame used in pipe fracture experiment

500

Weld Metal
400

Base Weld

cO

CL

300

~
(D

200

~

Cl)

1 OO

O

.OO 0.10 0.200.300.40

O

StrainFig. 5 Stress-strain curves for TP304 stainless steel base metal and weld metal

3000
9.5T

2500
~E 2000~.
~ 1500~~ IoooIT

500

o I o 2030 40 50 60
crack Extension, mm
Fig. 6 J-R curves for TP304 stainless steel

4.2 Analysis method
In order to evaluate the effect of the pipe weldment on the crack behavior, three different
material properties were considered in the finite element analysis as follows. In the firstmodel, the full pipe was modeled as the all-base metal. In the second model, the full pipe wasmodeled as the all-weld metal. The final model was designed with combination of the weld

metal and the base metal considering the actual weldment observed from the pipe as
illustrated in Fig. 7. The crack initiation load was deterrnined from the point satisfying thefollowing equation.

Japp (a P) = Jlc (8)- 158 -

In calculating the maximum load, the J/T analysis was used.
ALL BASE METAL~(a )ALL WELD METAL~(b)

VVELD METAL~~BASE METAL

Fig. 7 Schematics of pipe with circumferential through-wall crack, for (a) all base metal (b)all-weld metal, and (c) weldment pipe
4.3 Analysis results

Fig. 8 shows the resulting J values with increasing moment for three different cases. While
the variation ofJ values was similar for the all-base metal and the weldment analyses, the all-weld metal analysis produced much different applied moment versus J-integral curve. That is,even the crack was embedded on the weldment, the deformation behavior of the cracked pipe
was governed by the base metal. This implies that the base metal has a greater effect on theelastic-plastic crack behavior than the weld metal[1 3] .

80000
All base

E

60000

~_

1
~
~CaL,

CD

40000

O
d.

E

I

i)

Weldment

20000~~~]All~Jld.'., , .l

O
O

20

40

.

60

80

Moment, xE6 kN-m
Fig. 8 Comparison of applied moment vs. J-integral curves for all-base metal, all-weld metal,

and weldment pipe

- 159 -

Fig. 9 shows the calculated crack initiation loads for three different cases using Eqn. (8).
Three different Jlc values obtained from the I T, 3T and 9.5T CT specimens were applied andcompared with the experimental data. As shown in Fig. 9, the analysis results based on the
Jlc values from the I T CT specimen showed the best agreement with the test results. The
analyses based on the all-base metal and the weldment pipe showed better agreement with theexperimental results than the all-weld metal analysis.

8• IT 1 3T L 9.5T
='25
a' '

~
1'

dL

~2
coc"

A

o

L

'\=1.5

l

ID

CD

~

U
1
~o

--.---.S2l9o
Q~ 0.5O

AII base All weld Weldment
Fig. 9 Comparison of crack initiation loads for all-base metal, all-weld metal, and weldment

pipe

Fig. I O shows the resulting maximum loads for three different analyses. Three different
fracture resistance curves were used for respective case. As shown in Fig. lO, the analysis

results for the all-base metal and the weldment pipe showed a good agreement with theexperimental results. For the all-base metal model, the maximum load predicted based on thefracture resistance curve from the 9.5T CT specimen produced a better agreement than thatfrom the I T CT specimen. It can be seen that the pipe fracture behavior can be simulated
more realistically by adopting the fracture resistance curve for the 9.5T CT specimen due tothe size effect, as shown in Fig. I I . The analysis based on the all-weld metal showed muchhigher values than the experimental data for all cases. As mentioned above, the base metal
produces more significant effect on the elastic-plastic crack behavior, and thus this all-weld
metal model is not appropriate. As a result, the finite element analyses with the all-base metal

and the weldment piping provided a good agreement with the experimental data. In particular,
the analyses with the all-base metal produced slightly more conservative results than thosewith the weldment piping. Considering the conservatism involved in LBB application, thecurrent LBB analysis based on the all-base metal was proved to be appropriate.

- 160 -

2

• IT 1 3T JL9.5T

1 .8

1'

c"

o 1.6
~o

2 1.4

=
CQ

,

c"o 1.2

:~

~
1
~

L

I~

co

1)

o 0.8

1:,

~006
o.

1:

2 0.4

o_

0.2

O
All base

AII weld

Weldment

Fig. l O Comparison of maximum loads for all-base metal, all-weld metal, and weldment pipe

5000

x 16•~lch Pip,I

45co
400c

x

3500

C] r; D r_ J ,_~CC!

O D 9.5T

Cl
:i

>~l

E 300O,~

D

X-~ 2500[X~~
E 2000C!

X
ri~*A~~~~' Ai~. ~~ ~~~~ ~

,
1 500

L~] d!~~~ *~C~ ~)(A 3T
lOOO

(~~(
r

eoo

'',\

~

IT

o
o

20

40 OO OO

lOO

1 20

Crack Extension, mm

Fig. l I Comparison of TP304 stainless steel J-R curves from various plane size CT

specimens & pipe experiments

5. RESULTS AND DISCUSSIONS
Fig. 12 - Fig. 1 7 are some ofthe analysis results for the safety injection and the shutdown

cooling line. Fig. 12 and Fig. 1 3 show the modified PEDS for Group I in case ofthe DLC andtwice the DLC, respectively with varying Cl values. Fig. 14 and Fig. 1 5 show the modified

PED of the DLC and twice the DLC, respectively with varying C2 Values. Fig. 1 6 and Fig.

17 also show the modified PED ofthe DLC with varying Cl and C2 Values, respectively for- 161 -

Group 5 . Here, the Cl and C2 Values are the power law constants obtained from thematerial tests as shown in Eqn. ( I ). In order to extend the applicability of the modified PED,

the Cl and C2 Values were ranged i 500/0 in these figures. Also the actual LBB design
point for previously designed plant[14] is indicated by " I " in the modified PEDs. This pointis placed below the lower bound curve in both the DLC case and twice the DLC case. A
considerable amount of margin between the actual load and the LBB allowable load was
obtained. From these figures, the analysis result obtained by using the lower bound stressstrain curve was the lowest one, and thus proved that the current LBB analysis procedure isconservative. It can be seen from Fig. 12 - Fig. 1 5 (Group 1) that the Cl value has the mostsignificant effect on the allowable SSE Ioad while the C2 Value has negligible effect. It wasalso observed from Fig. 16 and Fig. 17 (Group 5) that the C2 Value has some effect on theallowable SSE Ioad while the effect of the C2 Value is less than that of the Cl value.
Therefore, the Cl value has more significant effect on the allowable SSE Ioad than the C2

value. In addition, broader allowable SSE Ioad region was obtained for Group 5 which haslarger outer diameter and thickness. It was observed that the applicability of LBB can beextended with increasing pipe size.

300~hupper bound -~ Best-f't- C1 =1 1 55l

I~ Lower bound- C1=385

C I =770Ez~: 200
~ca

~

.,.

LLJ

,E]

'/ : ~:. ' / 1; ,d:"

cl)
cl)

/'~/' , IA

~.1

-JQ

co 100

~

er .'
~r

/ ' /~l' // "/ ///LBB Desi9n point (YGN 3,4)

l

/

O

3

4 Crack
5 length(DLO, cm
6

7

8

Fig. 1 2 Modified piping evaluation diagram for Group l(DLC, C2=0.3)- 162 -

9

300
-e-Upper bound-~Best-fit- .CI =1 1 55

I

ihLower bound_ C1=385

C I =770

E

Z
x 200
~:)

- - -ElE!- - -

cV

_--O

J

O- -

LJJ
Cl)
Cl)

. -...-.--~ . - . ~

Er--

~

.~. .~.~~) 1 OO
cU.,..,.

.,.

~r~
LBB Design polnt (YGN 3,4)

I

/

O

7 8 9 10 11 12 13 14 15 16
Crack length(2xDLC), cm

Fig. l 3 Modified piping evaluation diagram for Group l(2XDLC, C2=0.3)

300-B-Uppef bound-~ Best-fit
C2=0.3- - - - C2=0.45

I~-Lower bound_ C2=0. 1 5

EZ~: 200
~
~
UJ

,,:

l:

,,

,:

,,
.g::

Cl)
Cl)

,,

,,

d

d

d

'!'
,~~

.~l

JO

cUO~

1 OO

d
LBB Design point (YGN 3,4)

l

/

o

345

7

Crack length(DLC), cm

Fig. l 4 Modified piping evaluation diagram for Group l(DLC, Cl=770)- 163 --

300
H~Upper bound -~ Best -fit

I

I~ Lower boundC2=0.3 - - - C2=0.15- . C2 =0 . 45

Ez~_ 200
1)

~
UJ

cl)
cl)

d~

d~

de

~ca 100~o

LBB Desian point (YGN 3.4)

I

/

O

7 8 9 10 11 12 13 14 15 16
Crack length(2xDLC), cm

Fig. 1 5 Modified piping evaluation diagram for Group l(2XDLC, Cl=770)

800-E-Upper bound- C I =846-~ Best-fit
C I =564

IF LoweT bound- C1 =282

El 600

O

z
!~
~
~

.lo

' / ' / 'l

UJ 400

cl:)

EI'

a:)

O'

~

cu~o

- 200
~

./'/'// /'/'/'/'/'/ ' / ;1"./'/ // / l'l IA//////\/ ' LBB Desian point (YGN 3,4)
/ /ll
~
of /

g

o
5

9

Crack length(DLC), cm
Fig. 1 6 Modified piping evaluation diagram for Group 5(DLC, C2=0.6)- 164 -

800
HE-Upper bound -~ 8est-fit- C2=0.9~rLowor boundC2=0.6 _ _ _ C2=0_3

EI 600

o

z
~
ID

cu

o

JuJ 400cl)
cl)

E' . '

o' .

o

~:)

cv

~

~ 200

g

/

l' /

/

/

/
l

/

/

/

/

/

A

\

LBB Desion point (YGN 3,4)

O
3

13

Crack length(DLC), cm
Fig. 1 7 Modified piping evaluation diagram for Group 5(DLC, Cl=564)

6. CONCLUSION
Following conclusions were obtained from the development of the modified PEDS for thesafety injection piping and the shutdown cooling piping in Korean Next Generation Reactor.
(1) The modified PED which can account for the variation of material properties in assessing
applicability of LBB during the initial piping design process was proposed, and developed forthe safety injection and the shutdown cooling piping in KNGR.
(2) The most conservative results were obtained with the lower bound stress-strain curve.

(3) The Cl value produced more significant effect on the allowable SSE Ioad than the C2
value .

(4) The range of LBB application was extended with an increase in diameter and thickness ofa pipe.

7. REFERENCES[l] "Evaluation of Potential for Pipe Break," NUREG 1061, Vol. 3, USNRC, 1 984[2] "Leak Before Break Evaluation Procedure," SRP 3. 6.3, USNRC, 1 987[3] A.D. Nana, K.K. Yoon, "Leak-Before-Break Allowable Load Windows Approach," 1994
ASME Conference, PVP-Vol. 287, pp. 35 41 1994
[4] R.J. Fabi, D.A. Peck, "Leak-Before-Break Piping Evaluation Diagram," 1994 ASME
Conference, PVP-Vol. 283, pp. I 1 1 1 15 1994

[5] S A Swamy P R Mandava D C Bhowmrck D E Prager "LBB Consideration for A- 165 -

New Plant Desrgn " NEA/CSNIIR(95)18 Vol 1, pp. 199-205, 1996
[6] N.S. Huh, J.H. Lee, Y.J. Kim, C.R. Pyo, S.D. Park, "Development of Piping Evaluation
Diagram for LBB Application to Safety Injection and Shutdown Cooling Piping," Proc. of
the Korean Nuclear Sociee/Autumn Meeting, Vol. II, pp. 697-702, 1996
[7] Y.J. Kim, C.S. Seok, "Stress-Strain and Fracture Resistance Characteristics of Primary

Piping Weld Materials for Younggwang 3/4 Nuclear Power Plants," S.K.K.U. Research
Report, 1 992[8] Y.J. Kim, C.S. Seok, "LBB Test Evaluation of Secondary Piping Materials for Ulchin 3/4
Nuclear Power Plants " S K K U Research Report, 1996
[9] Y.Y. Earmme et al., "Study of Effect of Dynamic Strain Aging on Fracture Toughness for

Main Steam Line Piping ofNuclear Power Plant," KAIST Research Report 1996

LIO] D M Noms B Chexal "PICEP : Plpe Crack Evaluation Program," EPRI Report NP
3596-SR, 1987
[1 I] ABAQUS User's manual, Hibbitt, Karlsson & Sorensen, Inc., 1 996[12] G. Wilkowski et al., "Analysis of Experiments on Stainless Steel Flux Welds,"

NUREGICR-4878, 1987
[13] B.R. Ganta, D.J. Ayres "Analysrs of Cracked Prpe Weldment " Palo Alto CA EPRI
NP-5057, 1987
[14] "Fmal Safety Analysrs Report (YGN 3/4) " Technrcal Report, KOPEC, 1993

- 166 -

IV-2

Fracture Mechanics Evaluation for The Cast Duplex
Stainless Steel after Thermal Aging

Shigeru Urata, Manager of Nuclear Safety Engineering
Section Kansai Electric Power Co. Inc., Japan

kIRACTURE MECHANICS EVNiUA nON roR THE CAST DUPLEX Sl:AINI.ESS STEEL AFIER
Tr{ERMAl. AGINIG

Shigeru Urata
The Kansal F.1ectric Po\~'er Company. Osaka. Japan.

Abstract

For the primary coolant piping of PWRS in Japan. cast duplex stainless steel \\'hich is excellent in terms ofstrength, corrosion resistance. and v~~eldability has con\'entionally been used. The cast duple~c stainless steel

contains the ferrite phase in the austenite matrix and thermal aging after iong term service is knovb'n tochange its material characteristics.

It is considered appropriate to apply the methodology ofelastic plastic t~racture mechanlcs for an e\'aluation
oi' the inte_~rity of the primary coolant piping after thermal aglng. Therefore. \\e e\'aluated the integrity ofthe primary coolant piping fbr an initial PWR plant in Japan by means of elastic plastic fracture mechanics.

The evaluation results sho\~~ that the crack \\ill not gro\\ into an unstable f~racture and the integrity ofthe

piping \\~ill be secured. even \~'hen such through \\-all crack length is assumed to equal the fatigue crack
gro\\th len_~th tbr a service period of~up to 60 years.

l . n~lTRODUCTION
For the primary coolant plping of PWRs in Japan. cast duple\ stainless steel (ASME SA_~5 1 Gr CF8M orequivalent) has been applied. This cast duple\ stainless steel contains a ferrite phase of about 5 to 250 o in
the austenite matrix in order to impro\'e its characteristics ofcorrosion resistance. strength. \\-eldabillty. etc.

In the cast duple\ stainless steel. the f~errite phase is gradually separated \\'ithin the operatlng period at thepWR operation temperature (at about _300 C). and that causes changes in its material characteristics.

An aging oflhe material tends to increase its tensile strength but reduces its fi-acture toughness. The dc_gree

ofthe toughness reduction becomes more si_gnificant \\'ith increasing ferrite contcnt and the rate of- 167 -

tou_ghness reduction increases ',vith rising temperature. Both ofthese tendencies are mitigated vL~ith thc

aging progress.

The present paper summarizes evaluation results on integri~' oi~the primar) ' coolant piping afier thermalaglng ibr initial PWR piant in Japan.

2. CRACK STABII_ITY ANAJ_YSIS
2 . I Evaluation Method

Fracture toughness of the primary coolant piping decreases due to thennal aging under long-term
plant operation. but this fracture type is considered to be a ductile fracture because ductile crack
gro\\th is recognized In the material test data after thermal aging. Therefbre. it is appropriate to apply

elastlc plastic fracture mechanics ibr an evaluation ofthe stabillty ofthe assumed crack.
The e\'aluation tlo\\~ is sho\\'n in Fi~:. I .

2.2 Selection ofEvaluatin Location

Material toughness. \\~hich represents fracture resistance. ~\~ill be reduced \vith increasing ferrlte

content and also \~~ith increasing operating temperature. In addltion. a high temperature causes a large

fi'acture force due to thermal bending moment. Thereibre. the hot leg piping at the reactor vessel

outlet nozzle \\as selected as the location to be evaluated since both temperature and imposed load are
se\'erer and ferrite content is relati\'ely high in the evaluating plant.

2 ,_3 Evaluation of Crack Gro\\th

The size ofthe fatigue crack is calculated under the condition that the assumed inltial defect on thepiping inside surface gro\\s duc to the stress cycles applied b~' thc plant operation.

( I ) F.\'aluation Conditions(a) Size oflnitlal DefectThe size ofthe initial defect ~\'as conser\'ati\"ely assumed. v\'ith a sufi~rcient margin. to be about

tw~ice the size ofthe detectable single defect. Namely. the initial defect ~\'as assumed to be a

- 168 -

seml-eiliptical. circumferential det~ect on the piping inside surface and its size to be 0.2t (depth)
l .Ot (surface length). \~'here "t" is the \\~all thickness.

(b)

Stress Cycle Used in Crack Gro\\th Analysis
The stress cycle \~'as produced on the basis of the transient conditions w~ith consideration of theactual operating status of' the plant.

(c)

Fatigue Crack Growih I.avL~
The fatigue crack growth la\~' is represented by the follow~ing equations provided on the basis of

the test data under the PWR prlmary coolant envlronment :l,da dN = CKefilTl . . . . ( I )Kei~f = Kmax ( I -R)1) ) (2....)

R == Kmm Kmax " " (_3)
Where.

da dN : Fatigue crack gro~\th rate Lmm cycle]

C : Constant = 7.0 l0~1()

m : Constant = 4.0
Keff : Effective stress intensity factor range [MPafm]
Kma~ Kmin : Ma)(imum. minimum stress intensity factors [MPafm]

R : Stress ratio
2)

E_\'aluation Results
The results oi'the fatigue crack _gro\\th analysis are sho\\'n belo\\~. It re\'eals that the fatigue crack

gro\\th rate is lo\\~ and the crack does not penetrate the pipe vL~all (67.4mm thick) even after long-term plant operation.

1 ) Assumed initial crack sizc

Depth : 1 3 .6 mm

I.ength : 67.4 mm- 169 -

2) Crack size afier plant operation of60 years

Depth : 26.5 mm
Length : 108.8 mm
2.4 Crack Stabili ' Evaluation( I ) Evaluation Method
Fracture toughness ofthe primary coolant pipe decreases after thermal aging. but this fracture
ty'pe can be considered to be a ductile fracture because a ductile crack growih is recognized in its
material tests. Therefore. elastlc plastic fracture mechanics are to be applied ibr an evaluation of
the stability ofthe assumed crack. Concretely, the crack stabillty is evaluated \\-ith the J jntegral
value (Japp). \\~hlch sho\\'s a fi'acture ibrce in comparison to the fracture toughness \'alue (Jmat).
\~'hich sho\\'s the fractttre resistance of the material after thermal aging.

(2) J Integral Value (Japp)
J jntegral \'alue is calculated \\-ith the stress analysis by using the f~mite element method on thebasis ofthe design load and the crack length tbr evaluation.

(a) Design Load
Loads imposed on the piping are those initiated under the normal operating condition and also

those caused by S I earthquake (maximum design earthquake).

(b) Analysis Model
For the analysis. the f~mite element model is applied to the elbo\~~ on the outlet side (hightemperature side) of the reactor \'essel \\-ith a circumt~erential crack on the inside surface.

Fig. 2 sho\\'s the fmite element model used fbr the analysis.

The crack gro\~th analysis \\'as conducted \~~ith the assumption ofan initial crack on the basis

ofthe number oftranslents durin_~ the operating period oi' 60 years. Then. the calculated
surface crack length \~~as conservatively con\'erted into the through \\'all crack ~\'hich length isassumed I .6t as sho~\~n in Fig. I for the crack stability e\'aluation.

- 170 -

(c) Analysls ResLtlts
Fig. 3 shov\'s the analysis results fbr the Japp \'alue at the evaluating location.

(_~) The Fracture Toughness Value (Jmat)
The Jmat value w~as determined by the lo\\'er bound cur\'e oftoughness (-2 a lo\\er boundcurve) obtained \~lth the toughness prediction model (H3T Model : Hyperbollc-Time
Temperature Toughness) on the basis of the ferrite content at the evaluating location.

(a) The H_3T model
The H_3T model expresses the relation oi' Jlc and J6 to the aging time vL~ith the fbllov,'ing
hyperbolic i~unction as sho\\'n in Flg. 4.

M = A ~ B . . . . . (4)t -~ C~\'hereM : Tou_g:hness ( Jlc or J6) kJ m~

t : Aging tirne. Hours
A : Fully aged toughness kJ m~
B : Constant relating to aging temperature
C : Constant relating to agin_g time

No\\ rt Is assumed that B exp (D I * Co) and D = Q R. \\-here T is the aglng temperature(K). Co is a constant and R is gas constant. When T is Ti. B is Bi. When T = T_j. B is B_j.

Then.

Bi B_j-exp[D(1 Ti-1 'lli)] """"" """"""' (5)Bi B_j-=e\p[Q R(1 Ti-1 1:i)1 The acti\'ation ener_gy Q (k-J mole) obtained l'rom temperature dependence of B is del~Ined to be

the embrittlement activation energy.
On the other hand. the initiation actl\'ation energy F is dei~med fi-om temperatLtre dependence ofinitiation time of'tertiary embrittlement tF. namelytFl tl'i ~ e\p [F R (1 'I'i - 1 1:i)1 " " " " " " " " " " " ' (7)- 171 -

The Q and F values obtained from the J6 changes fbr the test materials aged at 290. 325, _350 and
400 C are about 100 to 1 80 kJ mol. From this. it Is considered that the result w~ill beconservative \\'hen I OOkJ mol is used.
Fi_g. 5 sho\\~s the Flo\\' chart ofthe tou_ghness prediction model.

First. the data fbr the aging temperature of400 'C is fitted vL'ith the H3T model to obtain theconstant A (the i~ully aged toughness).
With this constant A fiY. ed. the data fbr agin_~: at 300 and 350 C are f~Itted \\'ith the H_3T model toobtain the constants B and C at the agin_~ temperatures of_300 and _350'~'C.

By obtaining relation ofthe above-obtained constants I~ B. and C to chemical composition. thepredlction e\pression can be obtained.
Table I show's e)camples ofthe prediction ibrmulas obtained for Jlc and J6 at _~25 'C \\-hich areconstant A ibr the prediction model. We have been modifying these ibrmulas by Increaseing data
further.

Table I Predictlon lbrmulas oi'constant A for the toughness prediction model(Temp.: -~25~C )Fo o : Ferrite content (o o) by ASTM A800

In e\'aluating the integrity of duple\ stainless steel components after thermal aging. It is

important to predlct the J- A a curve. When the Jlc and J6 values at an optional aging condition
are obtained. the J- A a cur\'e can be obtained by the fbllo\ving expression \\'ith the above
prediction model by giving the flo\~~ stress :c- 172 -

(b)Jmat value for evaluation
The i'errite content of the evaluating location \~~as calculated b~_' the method described in ASTM

A800 \~~ith the chemical compositions on the material certificate.
Fig. 6 sho\\~s the Jmat value (-2 a lo\\'er bound value) obtained by the ferrlte content at thee\'aluating location.

(4)

Crack Stabilit~' F_valuation Results

The circumferential surface crack obtained on the basis ofthe fatigue crack grovbth analysis after
the service period of 60 years \\~as conservatively converted to the circumferential through \~-allcrack ofthe same length. In order to estimate the stability ofthis assumed crack, the Japp andJmat values obtained in Items (2) and (3) above \~'ere compared in Fig. 7.
Fig. 7 sho\\~s that Jlc l" Japp even after the service period of60 ~_'ears.

Therefore. e\'en \\~hen fatigue crack gro\\th Is assumed afier the service period of60 years. theevaluation sho\\'s that any ductile crack \\-ill not be initiated under the design load conditions.

and the integrity of the primary coolant piping is coni"rrmed to be secured \\'ith sufficient margin.

3 . CONC I_US ION
Inte_~rit~' e\'aluation \\~as conducted by means ofthe elastic plastic fracture mechanics fbr the primarycoolant pipin_~ after thermal agin_~ jn an initial PWR plant.

The results sho\\~ that. e\'en \\~hen fatl_~:ue crack gro\\th is assumed after a service period of 60 ~_ 'ears. thecrack is estimated to be stable and the inte_~ri~' ofthe piping is conf~Inned to be maintained in consideration
ofthe thermal agin_~ oi'cast duple\.' stainless steel after lon_~-term plant operation.

RF.FF.RE.NCF_Sl ) Kanasaki

H. et.al. "Fracture toughness and fatigue crack gro\\ih of PWR materlals in Japan"2)Su7uki. I.

et.al. "I.on_~ term thernlal a_~in_~ oi~ cast duplex stainless steels" ICON-4.

- 173 -ICONE- I .

l.Ot

0.2t

Initial defect

1.6t

0.4t

Crack growth analysis

Stress cycle

1.6t

Assumed crack forevaluation

Elastic-plastic

fracture mechanics
Stability evaluation

Design load

Unstable fracture

Yes

No
integrity

Fig. l

Evaluation Flo\v for Thcrmal Aging- 174 -Countermeasure

':tb

\,

~~

\

Fig. 2

\

\

\

-~Y
\1b

Elemental Division Diagram(Crack length : 2c = 2t)- 175 -

200
'f,

~a

f::

~,
*l
!~:

~l
'F,

~
~
{:3

'1

'~
~)
::

lOO

>
c:s

eJ)
(:)

F,

lOO '-OO
Crack length (mm)Fig. 3J Integral Value at Evaluation Position

Toughness:M(CV, Jlc, J6)
ll

ll

ll

ll
It\

lh

IL

I\

l\

l\

B

M=A+t+C

l\
l

lI
[l
ll
ll
Ill II
ll l

A~:F

Il
ll
ll
1-J
Il

-C tF Aging Time, t
Fig. 4

Schematic illustration of the toughness prediction model- 176 -

300

(L)

~:)

O
fl
r'

'-f,

O
,~
O
,~
O
h,
*~,

n~J1

~~,

O

,-,

~:

O
f,
rl
fl
,-,(1)

~

JD
c~

'~
cl)

Colculation of constant A, B, C
r7)
cl)

fJ
rl
f,

'-4

'-,

CU

JO

13

T[ : or .T~= A +

O

t+c

r'
P~

'~_,

"~
-d

O

.'-1

'~*
.J

J

C._ Ibr .T- A

calL;ulntlon ot Cl'a cur\'c using .T[(] and .T6 \",lluc

'-4
i

A/

ll

'1~ r~~~~~~~ J=CI L~ ac

/

/

Tl

' ic~

6 nl Tll

O . 2 Tl~ rll~~aFig. 5

FIO\\ Chal't Ol' IhC tOU~lllICSS pTCCllctiOll lllOdCl- 177 -

800
700
rl

600

\

!4

l~
,~
::;

5 oo

fC::

~,

~J
(L)

:3
~-t

400

>
eJ)
fJ

C

300

,,

200

lOO

4

8 12 16 20
Crack growth A a (mm)Fig. 6

J- A a Curve (Hot Leg Pipe Material)

Jmat

400
Qlf~
f'C:

.\
'1
.~'

'V
'l,

C~

200

C~~
C:5

'1

Japp

JIC

o

l OO

2OO

CTack length (mm)Crack length forevaluation after service

period of 60 years
Fig. 7

Crack Stability Evaluation Diagram- 178 -30O

IV-3 Fracture Behavior of Carbon Steel Pipe with Local WallThinning Subjected to Bending Load

Katsumasa Miyazaki & Satoshi Kanno, Mechanical Engineering
Research Laboratory, Hitachi Ltd., Japan

Masayuki Ishiwata & Kunio Hasegawa, Hitachi Works, Hitachi Ltd., JapanFracture Behavior of Carbon Steel Pipe with Local Wall Thinning

Subjected to Bending Load

Katsumasa Miyazaki, Satoshi Kanno,
Mechanical Engineering Research Laboratory, Hitachi Ltd.
lbaraki, Japan

Masayuki Ishiwata and Kunio Hasegawa
Hitachi Works, Hitachi Ltd.
lbaraki, Japan

1-1, Saiwal cho 3 chome Hrtachi shi lbarakl ken 317 Japan

ABSTRACTTo evaluate the structural integrity of power plant piping, monotonic bending tests wereconducted on 4-inch diameter full-scale carbon steel pipe specimens with local wall thinning.
The local wall thinning was simulated as a eroded part. Eroded part was subjected to tensile
or compressive stress caused by applied bending load. The deformation or fracture behavior

at maximum moments was classified into three types. When eroded part was subjected totensile stress, ovalization and crack initiation was occurred. When eroded part was subjected
to compressive stress, Iocal buckling occurred. When fracture behavior was ovalization or
local buckling, the applied load decreased slightly form the maximum load. On the other hand,in the case of crack initiation, the applied load was abruptly decreased from the maximum
10ad, because of circumferential crack propagation. The relationships between applied load
and load line displacement for eroded pipes were compared with those for cracked pipes. The

maximum loads and total displacement for the eroded pipes were higher than those forcircunferentially cracked pipes. Maximum moments predicted using a net section stressapproach gave overly conservative evaluation for pipe specimens with local wall thinning.
Safety margin for locally thinned pipes was discussed compared with the design maximum
load. The safety margin in this paper were roughly more than 2 times.

1. INTRODUCTION
Carbon steel is used extensively in piping systems in power plants. High temperature andhigh pressure water and steam flow at high velocities through these pipes. Erosion / corrosion
has caused wall thinning in high-energy carbon steel piping. It is important to evaluate thestrength of piping with local wall thinning to maintain the integrity of the piping systems.
Burst and bending tests were performed on 6-inch diameter carbon steel pipes with local
wall thinning by JAERI (Japan Atomic Energy Research Institute). From the experiment, it

was found that fracture load was high, when locally thinned thickness of a pipe was theminimum required wall thickness defined in terms of design pressure (JAERI Report, 1993).
ASME (American Society of Mechanical Engineers) has considered the need to provide
appropriate guidance to system engineers on the subject on erosion and erosion/corrosion in
ferritic piping (Deardolff A. F., 1990). New criteria for acceptability of localized wall

- 179 -

thinning of piping components have been developed for inclusion in the ASME Boiler &Pressure Vessel Code, Section XI (ASME B&PV code, 1997a).
This study was undertaken in order to clarify the fracture behavior and strength ofcarbon steel pipes containing simulated local wall thinning to erosion/corrosion. Deformationand fracture load for locally thinned pipes were compared with those for cracked pipes. The

maximum moments were compared with the plastic collapse moments based on net section
criteria. In addition, safety margins were discussed using design maximum load.

2. EXPERIMENTAL PROCEDURE
2.1 Material
The materials used in the experiments were carbon steel pipe called "Carbon steel pipe

for hrgh Pressure Servrce STS 410" m JIS (Japanese mdustnal Standards) , which arecommonly used in coolant piping systems of nuclear power plants in Japan. The carbon steel,
STS410 is similar to ASME A333 Gr.6. The yield strength and ultimate strength at ambient
temperature were 326 MPa and 490 MPa, respectively for STS410.

2.2 Pipe Specimens
Full-scale experiments were performed on 4-inch diameter schedule 80 carbon steel
pipes made of STS410 with local wall thinning. The pipe geometry and dimensions of local
wall thinning are shown in Figure I and Table 2. The local wall thinning was machined on
the inside of STS410 pipe. All local wall thinning to simulate erosion/corrosion thinning wasdone by grinder machining. The eroded depth means metal loss depth in the thickness
direction and the eroded angle was a circumferential wall thinning angle. The eroded length
was an axial local wall thinning length which was roughly equal to the diameter of the 4-inch
pipe .

2.3 Load Conditions
A static bending load without internal pressure was applied by four-point loading until

maximum load appeared. The bending test apparatus was shown in Figure 2. The major spanand minor span, respectively, were 960 mm and 300 mm. Displacement velocity at cross head
was O. I mm/s. The total elapsed time from start of loading until fracture ranged from 10 to 20

minutes. The experiments were conducted at ambient temperature.
The eroded area of the local wall thinning was located such that its center coincidedwith the plane of maximum bending moment. There were two types of bending loads; for the
majority of specimens, bending tensile stresses were applied to the locally eroded areas,
bending were applied to the locally eroded areas of the other pipe specimens. Figure 2
showed an example of a pipe in which the tensile stress occurred at the eroded thinned area.

3. TEST RESULTS
3.1 Load - Displacement Curves
The maximum moments for all specimens are shown in Table 2. Eight pipe specimens
applied tensile stress at locally eroded area, and two others applied compression stresses byfour-point bending. In Table 2, the plastic collapse moment calculated based on net section- 180 -

stress criteria and fracture behavior, described in detail later, were also listed.

The load-displacement curves for STS410 pipe specimens with local wall thinningsubjected to tensile stress were shown in Figure 3. The fracture behavior was discussed later,

the eroded pipes subjected to tensile stress at eroded area showed two types of fracture
behavior. One was ovalization, and the other was crack initiation with local necking at theeroded part. For Figure 3 (a), fracture behavior was ovalization (Specimens No. TP-5, TP-6
and TP-7). When fracture behavior was ovalization, the relationship between applied load

and load line displacement remained the same in spite of difference of eroded area
configuration. The maximum load was observed at about 80 mm in load line displacement.
For Figure 3 (b), fracture behavior was crack initiation (Specimens No. TP-3, TP-4 andTP-8). When the crack was initiated, the applied load dropped abruptly. In this case, thedisplacements of these specimens, after maximum load was applied, were very small, because
crack propagated in the circumferential direction.

Load-displacement curves for pipe specimens (Specimens No. TP-9 and TP-ll)
subjected to bending compressive stresses were shown in Figure 4, compared to the curves
for the same local wall thinning geometry when tensile stress caused by bending moment wasapplied. In Figure 4 (a), normalized eroded depth, d/t was 0.5 and eroded angle, 20 was
On the other hand, normalized eroded depth, d/t was 0.8, erode angle, 2e was 180' ,180" .

in Figure 4 (b). The maximum loads for TP-9 and TP-11, where compressive stress wasapplied to the eroded area, were less than those for the specimens subjected to tensile stress.
However, applied loads decreased slightly with increasing the displacement.

3.2 Fracture Behavior
When a local eroded area in a pipe was located on the tensile stress side, two types offracture behaviors occurred at the maximum moments. When the eroded depth was shallow,
ovalization occurred at the maximum moment. The fracture behavior of such an eroded pipe
(specimen No. TP-6), occurring ovalization was shown in Figure 5. Crack initiation andgrowth at the eroded area were not observed at or after maximum moment. The load slowly

decreased after maximum point accompanying deformation of ovalization. Total
displacement was significantly large, when ovalization occurred.

When the eroded depth was deep, Iocal necking was observed first, then a crack
initiated around the maximum load. The fracture behavior of the pipe (specimen No. TP-4)
with deep eroded depth was shown in Figure 6. Local necking appeared in the local wall
thinning area around the maximum load as shown in Figure 6 (a), and a crack initiated in thelocal necking area and propagated circurnferentially as shown in Figure 6 (b). Total
displacement after crack initiation was relatively small, and the load abruptly decreased. The
crack initiation part after pipe fracture test was shown in Figure 7. The crack initiation point

agreed with the maximum eroded part. When the applied load increased, the maximum
eroded part was necked. The crack initiation occurred, because of locally rupture at the

maximum eroded part.
When the local eroded area was subjected compressive stress, Iocal buckling occurredin the locally thinned area. Figure 8 is a photograph of local buckling (specimen No. TP-1 1),

which was taken after the un-loaded.

- 181 -

4. ANALYSIS OF EXPERIMENTS
4.1 Fracture Behavior at Maximum Load
Fracture behavior of pipes with local wall thinning can be classified into three types at

maximum load in case of receiving a monotonic bending moment; ovalization, Iocal neckingcrack occurrence and local buckling .
When the eroded area in a pipe was subjected to tensile stress, ovalization or crack
initiation occurred at the maximum load. The ovalization or crack initiation depended on theinitial geometry of local wall thinning. The boundary between the ovalization and the crack
initiation was illustrated in Figure 9. When eroded angle was large and depth was deep, Iocal

necking and crack initiation was observed. Crack was initiated at the maximum moment andpropagated circumferentially after the maximum moment was reached. When the eroded area
was short and thin, the pipe occurred ovalization at maximum moment, and showed a large
degree of deformation.
Pipes containing a local erosion area with short angle and shallow depth showed highstrength and a large degree of deformation. It seems reasonable to accept such eroded sizes
below the straight line in Figure 9. It is conceivable that those were allowable local wall
thinning sizes caused by erosion corrosion.

When the eroded area in a pipe was subjected to compressive stress, Iocal buckling
occurred at and after the maximum moment, as shown in Figure 8. Although pipes subjected
to compressive stress showed slightly less strength than with those subjected to tensile stress,
the degree of pipe deformation was large.

4.2 Counparison of Cracked Pipes
Four-point monotonic bending experiments were have been done on the same diametercarbon steel pipes with circunferential cracks at ambient temperature (Miyazaki, et al., 1996).

Testing machine and bending apparatus used for these experiments were the same as thoseshown in Figure 2. Also, the cracked pipes were made of the same material as the eroded pipe
discussed in this paper.

The load-displacement curves for pipes with cracks, compared with the curves for pipes
with local wall thinning, called eroded pipes, were shown in Figure 10. The cracks were
through wall and part-through wall cracks. The crack angle and the eroded angle for both
pipes were 2e = 60' . The maximum loads and total displacement for the eroded pipes werehigher than those for circumferentially cracked pipes.

4.3 Net Section Stress Approach
Fracture strength for a circumferentially cracked carbon steel pipes could be predictedby net-section stress approach (Kanno, et al., 1993). When the crack depth is d, and crack
angle is 20, the plastic collapse moment Pc based on net section stress criteria is estimated by,

M 20 fR t(2sln P - d sinO)

(1)

t

1 dO)

P = ~(1C -(2)t

pc = Mc(3)

4(Lout ~ Lin )- 182 -

where M* is the plastic collapse moment, t is the nominal vyall thickness, O is the eroded half
angle, R is the pipe mean radius, P is the neutral angle of the pipe, of rs the flow stress, Iout

was the major span and li~ was the minor span of the test apparatus. The flow stress wasusually used as an average value of yield strength oy and ultimate strength o~. That is, theflow stress is defined as,

Of

Ov +0u(4)

2

Maximum load Pwa* for circumferentially cracked carbon steel pipes can be accurately
estimated from the plastic collapse moment calculated by Equations from (1) to (3).

Maximum moment, P~~* obtained in cracked pipe test and plastic collapse moment, P*
calculated by Equations from (1) to (3) was compared in Figure 1 1 . The base material,
STS410, used for cracked specimens was the same as that of the eroded pipes. The monotonic
four-point bending tests using cracked pipe were performed at ambient temperature and at
265~C (Miyazaki, et al., 1996). The maximum load obtained from cracked pipe test wasslightly larger than plastic collapse load. It means that maximum load for a cracked pipe
subjected to a bending moment can be predicted conservatively by Equations from ( 1) to (3).
The same technique was applied for locally eroded pipes, using Equations from (1) to(3). Plastic collapse load calculated from Equation from (1) to (3) was also listed in Table 2.

The relationship between maximum moment obtained from eroded pipe fracture tests and theplastic collapse moment was shown in Figure 12. When fracture behavior was ovalization,
the ratio of maximum load to plastic collapse load, Pwa/P, was about 1.2. On the other hand,in case of crack initiation at eroded part subjected tension and local buckling at eroded partsubjected compression, P~~/P. was about from I .25 to I .85.

The ratio of maximum load to plastic collapse load, P~*/P* , when occurring crack
initiation, was higher than that, when occurring ovalization. It means that maximum moments
will be overly conservative, when applying the Equations from (1) to (3) for locally eroded
thinned pipes.

4.4 Safety Margin
The primary stress for a pipe caused by a bending moment is designed to be less than
1 .5 S~ where S~ is the design stress given by ASME Boiler & Pressure Vessel Code Section

III (ASME B&PV Code, 1997b) and MITI (Ministry of International Trade and Industry)
Notification No. 501. The design stress S~ was 1 37N/mm2 for STS410. Then, I .5S~ = 206
N/mm2 for STS410 and 188 N/mm2 for STS370. The maximum design load, PD is given by,
MD = 1.5S~ X Z(5)PD = MD(6)

4(Lout ~ Lin )where MD is the design moment, Z is the section modules of the pipe. The maximum design
load for 4-inch diameter carbon steel, STS410 were calculated using Equation from (5) to (7)as,PD = 88 kN for STS410 (4B, Schedule 80)- 183 -

(7)

The maximum moments in Table 2, the safety margins for eroded pipes tested in thispaper were shown to be roughly more than 2 times.

5. CONCLUSIONS
Monotonic bending tests were conducted on 4-inch diameter full-scale carbon steel pipe
specimens with local wall thinning at ambient temperature. It is concluded as follows;
1) Fracture behavior and sequence of fracture process for locally eroded pipes could
be classified into three types; deformation of ovalization, Iocal necking - crack initiation -propagation when tensile stress is applied to the eroded area, or local buckling duringloading in case of compressive stress at the eroded area. It is conceivable that the locallyeroded sizes caused ovalization are allowable.
2) Fracture strength and total deformation of pipes with local wall thinning pipes
were higher than for the cracked pipes with the same depth and angle.

3) Net-section stress approach gives an overly conservative evaluation for pipes
containing locally wall thinning, as eroded area. The ratio of maximum load to plastic
collapse load, P~~/P* , when occurring crack initiation, was higher than that, whenoccurring ovalization. It means that maximum moments will be overly conservative, when
applying the Equations from (1) to (3) for locally eroded thinned pipes.

4) When comparing the maximum design moment, pipes with locally wall thinninghave more than twice the safety margin compared with design load, even when eroded
depth is 80 % of pipe wall thickness.

REFERENCES

ASME 1997a "ASME Boiler & Pressure Vessel Code Sec. XI, Rules for Inservice

Inspection of Nuclear Power Plant Components".
ASME, 1997b, " ASME Boiler & Pressure Vessel Code Sec. 111, Rules for Constructionof Nuclear Power Plant Components".
Deardorff, A. F. and Bush, S. H., 1990, "Development of ASME Section XI Criteria for

Erosion-Corrosion Thinning of Carbon Steel Piping", ASME PVP PVP Vol 186 NDE Vol 7
pp.7 1 -75 .

Japan Atomic Energy Research Institute, 1993, "Technical Report on the Piping
Reliability Tests at the Japan Atomic Energy Research Institute (Japanese)", JAERI-M, 93076, pp, 104-1 15, May.

Kanno, K., Kimoto, H., Hayashi, M., Ishiwata M., Gotoh N Mrura N FuJroka T
and Kahima K., 1993 "Low Cycle Fattgue Crack Growth and Ductile Fracture Under
Dynamic /Cyclic Loadings for Japanese Carbon Steel Piping : Part I - Experimental Study",
ASME PVP, PVP-Vol. 266, pp. 17 1-174.
Miyazaki, K., Kanno, S., Ishiwata, M., Gotoh, N., Miura, N., and Kashima K., 1996,"Fracture behavior under Monotonic and Low Cycle Loading in Carbon Steel Pipes andWelded Pipe Joints with a Defect" , ASME PVP, PVP-2 Fatigue and Fracture Vol. I , Book No.
HO I 05 1 - 1 996.

- 184 -

Table 1

Table 2

Geometry of local wall thinning.

Test results for eroded carbon steel pipes subjected to monotonic bendingtl Pmax 'Maximum Load, *2 Pc : Plastc Collapse Load, *3 PD :- 185 -

Design Load, *4 PmaxIPD : Safety Margin1 200Outer Diameter : 1 14.3 mm

Maximum Wall Thinning Part Inner Diameter :Wall Thickness :
.

l

i

I
l
i

-. -[ -. i

l

2e

~
Detall of Wall Thlnnlng Part

Fig. lGeometry of pipe specimen with local wall thinning.

Monotonic

Loading
RollerSupporting~7

RollerLoad ing
Pipe Specimen

(~

e

e

@

:O!

ono
oVo

@
wau Thinning~~~lFig. 2300 mm~
960 mmMonotonic four-point bending test machine.

- 186 -

97.1 mm
8.6 mm300-TP-5(alt=0.35, 2e= 180" )

250

~
Z
~ 200
O_

TP-6

~icc

(alt=0.5, 2 e= 90 ' )o 1 50TP-7

J~1(alt=0.5, 2e= 60 ' )o 1 OO,~~
50

Ovalization

OO

20
40 60 80 1 OO
Load Line Displacement, 6 (mm)1 20(a) 0+ah,ation

300
TP-8

250

~
Z
~
~ 200
O_
1:ca

o

(alt=0.8, 2 e= 60 ' )~1'.,~. TP_4
~ (alt=0. 8,
f'

1 50

J
~ 1 OO~
CL50'••••L/. 2 e = 90 " )TP~3 crack
Initiation.)(alt=0.8, 2 e= 180~: Crack initiation was observed
O

O 20LOad
40Line6080 1 OO 1 20
Displacement, 6 (mm)(b) Crack initiation

Fig. 3Load-displacement curves for locally thinned pipes subjected to tensile stresses.

- 187 -

300
TP-2

d/t = 0.5, 2 O =1 80'

CTension)

250

~
Z
~ 200
~

I Il / I Ib' II'

I'll 'IIIIIIII'llllllllllIIll"I'lllllIIIlll'O_~ 1 50Jo

I 11 Ic~l

TP-9(COmpresSlon)

1:o

~
Q'

1 OO

50
I

Material : Carbon Steel, STS41 O

O

O 20Load40
60 80 1 OO 1 20
Line Displacement, 6 (mm)(a) Eroded depth d/It = 0.5. Eroded angle 20 = 180'300
d/t = 0.8, 2 e =1 80-

250

~
Z
~ 200
O_

U 1 50oJ~o 1 OO
CU

TP-1 1••_• •••_...... / (COmpresslon)..

TP-3(Tension)

Q.
Q.

50

Material : Carbon Steel, STS41 O

O

80 1 OO
Load Line Displacement, 6 (mm)(b) Eroded depth d/It = 0.8. Eroded angle 20 = 180'
Fig. 4Load-displacement curves for locally thinned pipes subjected to compression stresses
compared to those for pipes subjected to tensile stresses.

- 188 -

Eroded Part

~ S~:.~

t'~~-~ r~~: .:~~~ ~;f- ::T~ ~~=~ 'Maximum Eroded Section

Fig. 5Fracture behavior of a pipe with a locally thinned area subjected to tensile stress(Fracture behavior : Ovalization, Specimen No. TP-6, Eroded depth d = 4.3 mm,
Eroded angle 20 = 90' , Eroded axial length I = 100 mm).

- 189 -

Eroded Part~ ~*11~]

M

M

dLocal Necking(a) Local necking (at P = 20 1 kN)

',
M

M
o)

(b) Crack propagation (at P = 1 61 kN)
Fig. 6 Fracture behavior of a pipe with a locally thinned area subjected to tensile stress.

(Fracture behavior : Crack initiation, Specimen No. TP-4, Eroded depth d = 6.9 mm,

Eroded angle 20 = 90' , Eroded axial length I = 100 mm).

- 190 -

Eroded Part

Fig. 7

Crack initiation part of a pipe with a locally thinned area sbbjected to tensile stress.

(Fracture behavior : Crack initiation, Specimen No. TP-4, Eroded depth d = 6.9 mm,

Eroded angle 2e = 90' , Eroded axial length I = 100 mm).

Eroded Partl;:
Local Buckling

M
o~

M

~

Fig.

8 Fracture behavi or of a pipe witha locally

stress'- 191 -

thinned area subjected tocom pressive

1 .O

Crack Initiation Region~~~

E- 0.8d~'loOIDo 0.6
1~

o

L 0.4
LUU

oo

o

oo

e

STS41 O
Eroded Length, Q

1 OO mm~
Bound ary

o

ON
c~

EL 0.2oZoOvalization(NO Crack Initiation Region)

0.0
1 80

Eroded Angle, 2 6 (deg.)
Fig. 9270 360

Boundary between crack occurrence and ovalization for pipes with locally thinned
areas subjected to tensile stress.

- 192 -

300

Eroded Pipe

2e O = 600 (TP-7, dlt=0.5)

250

~
Z
~ 200
CL

f"

Eroded Pipe

lj

~lc:f

o

1 50

with Part- Through ',Wall Crack (dlt=0.5) '\f

J~o 1 OO(TP-8, dlt=0. 8)

Pipe ".

l
l

l

,
I

l
l
I
I
I
l

Q.

,

~

l
I
I
I

50

I
I

Pipe
with Through Wall Cra ck(dlt-- I .O)

O

O

Fig. 1020
40 60 80
Load Line Displacement, 6 (mm)
1 oo

1 20

Comparison of load-displacement curves between cracked pipes and locally thinned

pipes. (Eroded angle, 2e = 60' )- 193 -

300
Cracked Pipe

~
Z 250
~

Material ::~Carbon Steel , STS41 O

o
g

as

E

O_

[l

200

Pmax = Pc~
S 1 50
c~f

E

.=E_

Temp.

~

~

o
e

1 OO

RT

Norm. Crack Depth
dlt

0.5
1 .o

265lc

0.5
1 .O

50
50

1 OO 1 50 200 250

300

Plastic Collapse Load, Pc (kN)
Fig. 1 1

Comparison of the maxrmum load, P, with the plastic

colla pse

load,

Pmax f or

cracked pipes.

300

~
Z
~
~
~
E

O_

Eroded Pipe
Material :250

Carbon Steel, STS41 O

e
C]

200

10
c~;

(~~
Pmax = Pc~oJE 1 50.E_
Applied LoadatEroded Part

FractureOoTensionOvalization[]

Compression~~~ 1 oO

50

Behavior

Crack I nitiation

Local Buckling50 1 OO 1 50 200 250 300
Plastic Collapse LOad, Pc (kN)

Fig. 12Comparison of the maximum load. P. wrth the plastrc collapse load P. for erodedpi pes.

- 194 --

I V-4

Leak-Before-Break Assessment of CANDU Pressure Tube
Consid~ring Leak Detection Capability

Youn-Won Park and Yeon-Ki Chung, Korea Institute of \_ ~uclear Safe~', Korea

Leak-Before-Break Assessment of CANDU Pressure Tube Considering Leak
Detection Capability

Youn-Won Park and Yeon-Ki Chung
Korea Institute of Nuclear Safety

19 Kusung-dong Yousung-ku Taejon, Korea, 305-338

Abstract
The LBB(Leak Before Break) concept is employed as defense in depth in CANDU reactors
in order to avoid an unstable failure in pressure tubes. The LBB in pressure tubes can besatisfied when detection, confirmation, and loeation of the leak are canied out and the reactor
is placed in a depressurized condition before the crack exeeeds critical crack length. Leakdetection and location is provided by the Annulus Gas System. Therefore, the input
parameters for LBB analysis, such as leak detection and leak location eapabilities, should bemade available through AGS in-situ test. Recently, a simulated moisture inj ection test has

been performed in one of the CANDU reactors in Korea, and the LBB assessment was
perfonned using the test results. Only leak detection eapability is considered for LBB
analysis; the leak location is ruled out beeause of lack of information in the AGS response
vis-a-vis leak location procedure. In this paper LBB assessment is presented using both theleak detection and leak location eapability,

1. Introduction
Since several ineidents of leaking in the pressure tubes in the CANDU reactors have been
experienced, significant efforts have been made to improve the pressure tube integrity
through design, material, and fabrieation upgrades during last deeade.

However, delayed hydride craeking (DHC) is still believed to be the major potential threat topressure tube integrity. To cope with the likelihood of loss of pressure tube integrity due tosueh DHC, Leak-Before-Break (LBB) is used as a defense-in-depth. The flaw acceptanceeriteria for pressure tubes, provided by the Fitness For Service Guidelines, (FFSG)[1]requires demonstration that DHC will not oceur from flaws in pressure tubes when thehydrogen equivalent concentration exceeds the Terminal Solid Solubility (TSS) for
dissolution at full power operating condition, and that if the flaw was propagated by DHC,
unstable rupture of the pressure tube is avoided by LBB.

The LBB coneept requires timely detection, confirmation of leak, and appropriate operator
aetion to shut the reactor down to a cold depressurized state before the growing craek exceeds

the critieal erack' Iength (CCL). Since the typical CANDU reaetor consists of 380 to 480
pressure tubes, even though a leak is detected in time and the reactor is cooled down, it is not
so easy to locate the tube eoncerned. This is beeause once the reaetor is in a depressurized
condition, the crack closes and the leak is near impossible to fmd. Unless the leaking tube isidentified, an entire core inspection is necessary at a cold state. This is very difficult, time

consuming, and expensive. Therefore, Ieak deteetion and leak location are important from an
integrity viewpoint as well as eeonomical standpoint. Normally, only the leak detection isconsidered in the LBB analysis because there is not enough data for leak location tests[2].
During the pre-operational test of Wolsung unit 2, a complete set of tests for leak detection- 195 -

and leak location using simulated moisture inj ection was conducted; these results aGCOunted
for LBB analysis for Wolsung units 2, 3 and 4 in Korea.

2. Concept of Leak-Before-Break Assesslnent of CANDU Pressure Tubes

CANDU Reactor
A CANDU reaetor consists of a large tank, called calandria, oontaining D20 moderator at70"C, and is penetrated by 380 horizontal fuel channels each 6 m long. Each channel consistsof a pressure tube containing fuel and coolant D20 at a pressure of I OMpa and at atemperature ranging from 260~C at inlet to 300~C at outlet. The pressure tubes surrounded
and insulated from the cold moderator by a ealandria tube. The spaee between the pressure
tube and the calandria tube is filled with recirculating C02 gas, whieh is called Annulus Gas

System (AGS). The pressure tubes are made from cold-worked Zr-2.5Nb with a wall
thickness of 4 mm and an inside diameter of I 03 mm. The calandria tubes are made fromannealed Zircaloy-2 with a wall thickness of I .4 mm and an inside diameter of 1 29 nun. The
pressure tubes are rolled into the end fittings at each end of the fuel ehannel. The residual
stress produced by the rolled j oint fabrication process is still considered as a potential causeof developing DHC, even though it has been remarkably reduced by using a zero elearance
rolled j oint[3 1 -FEEOERS

FEEDEAS

CALA~l DP IA

TVBE
,,

J

FI~Ird(5AhNULUS F UE L
Ol.]TI FTEUNDLESt'lICS.ct,L]RE

TUSE

C12)

Fig• 1 CANDU Fuel channel
Delayed Hydrided Cracking
While the fatigue crack growth which is considered as a principal cracking mechanism for
PWR reactors is an order of I 0-8 m/cycle, the erack growth by DHC is an order of I 04
m/cyele. Thus the contribution of DHC to crack growih is mueh more dominant, so that only
DHC is taken into account for LBB analysis of pressure tubes. The principal mechanism of
DHC is as follows [2] :Once nucleated, the crack grew in the radial-axial plane under the hoop tensile stress when

hydrides were present as the reactor temperature was below Terminal Solution Solubility(TSS) of hydrogen. When the reactor was hot, the hydrides dissolved and crack growth
stoppedj but, the crack surfaces were oxidized by heat transport of D20. During later
shutdowns the hydrides reprecipitated and the cracking recommenced. When the reactor wasagain returned to power the hydndes redissolved, the eracking stopped and the crack- 196 -

oxidation GOntinued, with the newest part of the crack having the oxide with the smallest
thickness. Eventually the crack growth in the radial direction led to penetration of the tube
wall so that the heavy water of the pressurized heat transport water leaked into the annulus
gas system.

Reactor Operating Condition
Because the Reactor Outlet Header (ROH) pressure is the PHTS pressure parameter forreactor operation, all the reference pressure values are expressed in terms of the ROH
pressure taking pressure drop into account. The reactor operating conditions at FPH, ZPH,
and ZPC are presented in table 3 . For pressure and temperature to be used in LBB analysis,
the pressure drops from the pressure tube inlet and outlet to the ROH were derived for every
5'C increments from 40'C to operating temperature, by using linear interpolation[4]-Table I : Reactor Operating Conditions and Design Pressures at Different Locations[4]Parameters for Leak-Before-Break Analysis
The LBB analys'is means that when an unlikely leak through-wall DHC crack occurs,
throughout the period of time needed for leak detection, eonfirmation, Iocation and finallyplaeing the reactor in a cold depressurized state, the growing crack length is always less thanthe critical crack length in order to assure that the probability of an unstable pressure tube
rupture is extremely low.

Therefore, the LBB analysis integrates the AGS response to leakage from an assumed
through-wall DHC crack with the change in crack velocity and critical crack length as thereactor goes from full power to a cold depressurized condition. The actual station operatingprocedures should be used to analytieally simulate the sequence of events (SOE) following
the crack penetration of pressure tube. The parameters needed for the LBB analysis are as
follows [ I J :l ) Maximum crack length at penetration (Lp)
From the LBB view point, the crack dimensions at onset ofleakage is very important. This isestimated to be around four times the normal thickness [ I J . Lp of 20 mm is used in thisanalysis based on the recommendation of FFSG.

2) Delayed Hydride Cracking Velocity (DHCV)
Conservatively, upper bound values of DHCV in axial direction are used:

V 5 2xlO exp(-414451RT) (1)
3) Critical Crack Length (CCL)- 197 -

The half-critical crack length (c), as a function of temperature, stress, fluence, and hydrogenconcentration, can be determined from the following equation:
K~fT

(2)

C - 2Ma
8af2 In[sec( h )]2 cr f(Ty + au(Tf~ 2ah = P(L + l)w

M =[1+1.255(c2 /(r~w))-(0.0135(c /r~w) (O 0135(c /(r~w)2]1/2

4) Leak Rate
A correlation of D20 Ieak rate (kg/hr) with eraek length (mm) is given:

Q = -1 1 .2 + 0.0014(2C)3 (3)where, 2C~~ = maximum axial crack length (mm)This equation is based on PHTS operating conditions of 6MPa and 250~C. Thus, specificconditions other than 6MPa and 250~C, Ieak rates ean be determined using sealingcoefficients.

5) Station Response to Moisture in AGS
The time needed to detect, confirm, and loeate the leak should be provided. Following thefirst leak at penetration, the time necessary for initiating the rate of rise alarm is taken fromthe AGS performance test carried out at Wolsung unit 2 in May 1 997. The AGS performanee
test and its results will be described later on.

6) Pressure tube dimension
For a conservative approach, pressure tube dimensions were taken from expected values after

l4.51 9yearsof operation [4]• Installed thickness is based on the mi
mm, subtraeting both the corrosion allowance of O. 1 40 mm and wear allowance of 0.064
mm. Dimensions after 1 5 years of operation are made by considering an additional wall
thinning of O. 1 49 mm and a diametral creep strain of 2. I O/o. The wall thickness is also madeby considering a local wear of 0.028 mm due to the annulus spacer rolling. At the end of 1 5-

year serviee, the wall thickness is 3.83 mm, the maximum imer radius is 53.37 mm, and themean radius is 57.20 mm.

Assumptions used for LBB analysis[5]l ) The most likely area for DHC crack to oceur is inboard of the rolled j oint; because of
relatively high residual stress. The crack growtll in one direction should be stopped due torestraint when the crack arrives at the rolled joint. However, to be conservative, the DHC
craek growih is assumed in both directions throughout the event.
2) The crack is to propagate at the upper bound crack velocity
3) A Iower bound estimate of critical craek length and fracture toughness is used
4) The pressure tube leak oeeurs in the longest AGS ehaunel and the AGS is operated in very
Iow recirculation flow. This leads to a longer moisture detection time.
5) There are sufficient hydrides at power for crack growth to oecur.
6) The pressure tube dimension of 1 5 years in operation is used for accounting for the wear
and corrosion allowances, and the diametral creep and wall thinning.
- 198 -

3. Leak Detection and Location Capability Test
Leak detection in a pressure tube is provided by the Aunulus Gas System (AGS). The AGS
circulates dry C02 gas in the aunular space between the calandria tubes and pressure tubes.
The fuel channel annuli are connected in a series/parallel configuration to fonn a closed loop
as shown in Figure 2. There are 44 inlets conneeted, in parallel, to the recirculating gas
supply system and 44 outlets connected, in parallel, to the drain and the suction side of therecirculating compressors.

A total of ten tests were completed[6]. Seven were moisture injection tests to determine

system dew point versus time response with continuous injection until the dew point
increased to - I O'C. Three of the ten tests were to measure the change in dew point
response to leaking valve in and out proeess. The main purpose of the leak location test was
to qualify the AGS abnormal operating proeedure to be used for identifying leak channels.

1 1 Stnng

44 Inlet Tubes
Reactor
380 Channel

4 Bank3 to 11channelAnnuli in
SeTies/ParallelOO
Dew Point

O Compressor

M ete rBeetleFig. 2 : Aknnulus Gas System Configuration

Leak Detection Test
Of the seven moisture injection tests, three tests were conducted with the Primary Heat
Transport System (PHTS) cold and four tests with PHTS Zero Power Hot (ZPH). Table lshows the test matrix for leak detection and leak location.
Test 1, 2, and 3

The first injection was carried out in the main circuit at the downstream of the compressors
and upstream of the gas distribution manifolds. The injected D20 and C02 gas was therefore
distributed to all the 44 strings. The purpose ofthese tests was to investigate dew point versus- 199 -

time response and the dew point's rise rate under several conditions including flow rate,
string length, operating modes, and so on. The test results show that there are several peaks
which are related to the first wave of moisture from three channel, five channel, sevenchannel, and so on. Interestingly, using these results, a relative flow resistance from different

numbers of channels can be drawn out. The flow rate of each string was calibrated to be thesame and the lengths from the inj ection point to the first channel inlet of each string were
almost the same. Therefore, the differences in time in Fig. 3 come from the elapse timeneGessary for passage through different numbers of fuel channels, the following relation Gan
be derived. Note this relation is applicable to a hot reactor condition because it is only
response ratio for geometrical lengths.

(3C + L)N = 0.6 for 3 ehannels(5C + L)N = 0.92 for 5 chaunels
where C = equivalent length of one fuel channel. L = Iength from injection point to firstchannel inlet of each string plus length from outlet of last ehannel to dew point meter, and
V=gas velocity.

Therefore, the time required to pass through one fuel channel is equivalent to the time,
multiplied by I . 3 3 , neeessary for traveling path L. Using this conversion factor the responseof an I I channel string can be drawn out to be I .88 hr and eomparing this value with the test
result of I . 89 hr shows that such eonversion is acceptable.

As shown in table I , in test 2 and 3 , injection was into one of eleven channel strings which
are the longest strings at Wolsung unit 2. Test 2 was carried out with a flow rate of 2.6 L/sand test 3 with 5. I L/s. The resulting Rate of Rise alarm was initiated at 2 hours and at I . 5hour respectively.

Table I : Mainx ofAGS Performance Tests- 200 -

Rale of Rise (Test 1)

35. OO

30. OO~; 5 channel response ;- 25.00~\oo 20.00

11 channel response j~i~~o 15.00~oaz 1000
5. OO

O. OO
O. OO

50. OO I OO. OO 1 50. OO

200. OO

Elapsed Time (minutes)

Fig. 3 Rate of Rise responses from strings having different number of chaunels
Test 4This test was similar to test 3 with an inject + recirculate phase followed by an inject + purge

phase. The reactor was hot and the recirculation flow was adjusted by throttling thecompressor inlet and outlet valve to simulate very low flow AGS operation, which is I .65 L/s.
The flrst wave of moisture came through with a peak of about 6'C/hr at about 2.6 hours after
starting injection. During the inject + purge period, the dew point dropped to around -30'C inabout 2 hours .

Test 5This test was done with two compressors and with an addition of 0.30/0 oxygen. In
comparison with test 3 which has similar starting dew points and similar injection rates
shows that the rate of rise was higher for a hot reactor than a cold reactor. The first wave
appeared at one hour and the peak rate of rise of dew point was 23~C/hr.

Test 5B
This test was carried out under very similar conditions to test 5 ; very high flow rate with two
eompressors. The frrst signal appeared about one hour after the start of injection. The peak of
the rate of rise of dew point was about 3 5~C/hr. As an example, the dew point trend obtained
from test 5B is shown in figure 3 .

Test 6A
This test was with a compressor and an inj ection into the 3 channels string. The first signal

appeared about 25 minutes after the start of injection. Since a response corresponding to- 201 -

normal operating conditions with I I channel strings was not available, test 6A results wereconverted to I I channel data using the conversion factor obtained from test I .

Dew Point Trend and Rate of Rise

40
35
30

O
-5

~
O
~
.5
4'

; Rate of Rise i

25 ~

-1 O

20 (DO-1 5

15 ;a~~5'
10 ;~~ -2 Oa -2 5
(D~

5~*
O~

-30

-5

-35

-1 O

-1

o

12345

6

7

Elapsed Time (Hr)Fig. 4 Dew Point trend and Rate of Rise for Test 5B
The test results are summarized in table 2. When eomparing the rates of rise for a cold reactor
with those for a hot reactor, the responses at cold eonditions are slower than those at hot
conditions. Therefore, four data are used for determining the relationship between initiation
time for rate of rise alarm and AGS fiow rate, as shown in figure 4. It is valuable to note that,since the time for rate of rise alann is the time required to pass through I I channel plus
conneoting tubing length L, the relationship between time for rate of rise and flow rate should

be a form of y = C/x, where r time for Rate of Rise, C = constant, and x = Flow Rate. The
curve fit, shown in figure 4, is very close to the expected form.

Table 2 : Test Sununary for AGS Leak Deteetion* data converted from 3 channels to I I channels- 202 -

Rate of Rise Versus Flow Rate
3
2 .5

2
J~

8
if

1 .5

~o
~
co

OC

O.5

O
o

1

2

3

4

5

Flow Rate (L/s)

Fig. 5 Rate of Rise versus AGS flow rate for hot reactor eondition

Leak Location Test
Current Leak Location Procedure
A moisture level, showing a continuing tendency to increase, indicates a leak into the systemfrom a calandria tube or pressure tube. To locate the leak source, the purge flow is passed
through each of the four banks of I I inlets in turn. The recirculation mode is changed topurge mode. Two out of the four headers valves are closed at one time based on the
dichotomization and the moisture level is observed on a dew point analyzer. This will narrow
the location down to one of the four banks.

In the bank giving the high moisture reading, the purge flow is passed through eaeh of I linlets by closing the rotameter outlet needle valves in five of the lines. Based on the
dichotomization process, Iines with no channel leakage will be isolated and it will narrow the
10cation down to one line which feeds three to eleven chaunels depending on the nuniber ofchannels in series. Therefore, to locate the leaking string requires 6 diehotomization
processes.

TestResults (Test 3A, 5A, 6B)
This test is similar to leak deteetion test 3 , except that leaking and non-leaking bank andstring valving operations were carried out after the frrst wave appeared. As shown in Fig. 4,
the first wave appeared about I 05 minutes after the inj ection.

When two non-leaking banks were valved out, a reduction in total flow was produced so thatan increase in dew point was observed. Valving back in two non-leaking banks resulted in a
reduction of dew point; but, the initial value before the valving out operation was notachieved beeaus~ of continuing moisture injection. Valving out one non-leaking bank andone leaking bank resulted in drastically reducing the leaking string flow rate, thus a large

reduction in dew point occurred. The opposite effect was seen on when valving in theconcerned banks or strings. Valving out and in the leaking string showed a large change indew point, whereas valving out and in six non-leaking strings produced only small effects.
- 203 -

In general, a significant reduotion in the dew point was obtained when the leaking string or abank containing the leaking string was valved out, and a significant increase in the dew point
was obtained in the opposite ease.

Two other tests, test 5A and test 6B, performed with a hot reactor, showed a similar trend.
Three tests for leak location are summarized in table 3 .

Dew Point Trend Crest 3A)
O

-5

-10
-15

o
~~
8 -20
~

o -25
-30
-35
-40
0.0

3.0

1 .O

5.0

Elapsed Tlrne (Hr)Fig. 6 Leak Location Test (Test 3A)

Table 3 Summary ofLeak Location Test

( I ) time required for dew point trend to roll off

As shown in table 3 , one operation of valving out or in takes only 7 to 8 minutes. When using
the leak location proeedure, the narrowing down from 4 banks to a leaking string, at least six

dichotomization processes are needed and will therefore take less than 50 minutes. It isvaluable to note two things: firstly, the dew point response was much faster than expeeted,and secondly, the rate of rise produced from a leaking string or a bank containing a leaking
string is always higher than very high rate of rise alarm set point which is 4.5'C/hr. Before
leak location test it was predicted that using the recireulation mode the second wave, affectedby valving out or in operation, must appear at the time when the time as much as required forfirst wave elapses since the first wave appears. However, as shown in table 3 and Fig. 4, the- 204 -

6.0

corresponding wave appeared immediately and the dew point trend rolled off in 7 to 8
minutes after valve operation. It means that the waves produced by valving out or in process

were not made by moisture transportation. Therefore, even though the exact mechanism ofsuch wave apparition is not yet clearly defmed, it can be presumably drawn out, that the same
trend can be obtained when the AGS is in a stagnant mode.

4. LBB Analysis and Its Application
Based on the current operating procedures a complete sequenoe of events are establishedfrom leak detection to the time of the reactor being placed in a oold depressurized state.

Sequence of Events based on Current Operating Procedure[7]The Sequence of Event (SOE) including the leak detection, confrrmation, Ieak location andreactor shutdown should be developed based on station operating procedure. Leak detection
can be made by 'a rate of rise alarm and/or beetle alarm. Normally the rate of rise alarmprovides an earlier response to moisture so that in the SOE, Ieak detection is supposed to bemade by rate of rise alarm. Based on the AGS test results (Fig. 3), a rate of rise alarm is to be
initiated at 2.6 hours after the start of a leak, when the AGS is under very low flow operation.

Operators should reduce power from Full Power Hot (FPH) to Zero Power Hot (ZPH : 26 1 'C
and 8MPa). Reactor temperature drop with a maximum allowable cooldown rate, 2.8'C/min,is followed by a pressure drop with a maximum rate of I bar/min. Since the beneficial effectof thermal maneuvering on DHC velocity is presented by Moan et al [7], the following
thermal maneuvering is taken: reduce PHTS temperature from FPH to a temperature lower
than ZPH by 40'C, hold this temperature for about 1 5 minutes and subsequently increase thetemperature to ZPH again. After reaching ZPH, the AGS should be put into the stagnant
mode and should be followed by the colleetion of leakage and the leak location process,according to the AGS abnormal operation procedure. As soon as the leaking source is locatedor total allowable duration time elapses, the reaotor should be cooled down to below I OO 'C

with a maximum allowable cooldown rate and followed by pressure drop down to 3 Mpa.
Under these conditions, the critieal crack length of the pressure tube is very large, so that
failure of the pressure tube is unlikely to occur. The cooldown should continue until thereactor is safely placed in a cold depressurized eondition, using station proeedure. To cool thereactor down from FPH to a condition below 90"C, maximum cooldown rates are assumed asfollows: 2.8 "C /min for 3 l0-260 'C, 2.4 'C /min for 260-150 'C, I .2 ~C /min for 150-90 'C, and

0.4 ~C /min for below 90 'C as presented in FSAR of Wolsung unit 2[8]. The maximum
pressure drop rate is assumed to be I bar/minute throughout the SOE.

LBB Analysis
The crack length at each time step (every O. I hour), when time=ti, has been calculated using
the following expression [ I J :(4)

Li = Li_1 + (Vi + Vi_1)(ti ~ ti-1 )Where. Li = Iength of erack at time ti, Vi = DHC velocity at time ti.

The critical crack length at each time step was also generated from Equation (2). The LBB
can be demonstrated by showing that the growing crack length is always below the criticalcrack length throughout the SOE. A cooldown path based on the foregoing SOE is shown in- 205 -

Fig. 7 where temperatures and I O times pressure at ROH are presented as a function of time
elapsed after the initiation of leak. When using the SOE. LBB could not be assured because
the crack length exceeds the CCL before leak deteetion as shown in Fig. 8.
Therefore, it was .hrstly required to reduce the time for leak detection. This could be achievedby inoreasing the AGS gas flow rate as shown in Fig. 5. The minimum gas flow rate of AGS
was set to be 2.3 L/s so that the leak detection could be done within I . 9 hour after the start ofleak. Secondly, pressure plateau at ZPH should be reduced from 8Mpa to 7Mpa to increase
CCL at ZPH condition. Finally, cooldown process from ZPH to depressurized condition
should be done in two steps: ZPH to 1 50~C and 1 50~C to cold state and pressure drop should

be made in-between at 1 50'C. As a result, the LBB can be satisfied using this modified
operating procedure as shown in Fig. 8.

Initial Cooldow n P8th

12

10

~
CL 8
=
~
~

6

c02

~ 4
=
O
az
2

o
o

300
50 1 OO 1 50 200 250
350

ROH Tempemturo (OFig. 7 Current cooldown path

5. Conclusion

A oomplete leak-before-break analysis was carried out for pressure tubes of a CANDU
reactor considering leak detection and leak location capabilities. The detection and location
capabilities for pressure tube leak were determined by a simulated moisture inj ection tests
perfonned at Wolsung unit 2. A relationship of leak detection time versus AGS gas flow rate

was derived and an appropriate minimum AGS gas flow rate meeting LBB was determinedfrom this relationship. Leak loeation tests were carried out based on the current
dichotomization process using valving out or in a leaking string or leaking bank containing aleaking string. The responses of dew point were very sensitive to valving out or in process
such that very high rate of rise alarm eould be initiated in 7 to 8 minutes following valving
out or in whatever a leaking string was concerned. When using a dichotomization procedure,
about 6 valving out or in operations are expected to loeate a leaking string, which require lessthan I hour. These test results were used as inputs for LBB analysis.

The upper bound values of DHCV and pressure tube dimension expected after 1 5 years ofoperation were used and all the assumptions were made very conservatively. When using- 206 -

current operation procedure with the foregoing AGS test results, LBB was not satisfied sothat five main improvements were neeessary to meet LBB: reducing leak detection time byincreasing AGS gas flow rate. Iowering pressure plateau at ZPH from 8Mpa to 7Mpa,
inserting a pressure drop process during reactor cooldown from ZPH to cold depressurized
condition. and providing a maximum allowable time for operator to take action for leaklocation procedure. From these LBB analyses results. reactor operating procedure and plant
teehnical specifieations were modified.
*"" ~.b** ''~-~"=,."'' ~~.

'. ''Fig. 8 LBB Analysis Results based on eurrent operating proeedure

LBB Analysis (fln.hy alcceptabte)'.
=.."'' '.~.

Fig. 9 LBB Analysis and corresponding cooldown procedure to meet LBB

Reference[ I J Fitness for Serviee Guidelines for Zr Alloy Pressure Tubes, COG-91-66, 1 991 May.
[2] N. Azer, D.H. Barber, P.J Boucher, P.J. Ellis, J.K. Mistry, V.P. Singh and R. Zaidi, "Leak

Detection Capability in CANDU Reactors", Speeialist Meeting on LBB, Lyon, Oct. 1 995[3] G.D. Moan, C.E. Coleman, E.G. Price, D.K. Rogers and S. Sagat, "Leak-Before-Break in
the Pressure Tubes of CANDU Reactors", Int. J. Pres. Ves. & Piping 43 (1990) I -2 1 .

[4] Design Report of Fuel Channel Stress Analysis for Wolsung NPP 2, 3 , and 4, AECL. May
l 995[5] Preliminary Generic LBB Analysis for Wolsung Units 2, 3, & 4 by AECL, Oct. 1 996[6] Moisture Injection Tests at Wolsung Unit 2 prepared by V.P. Singh of Ontario Hydro,

Aug. 1997[7] PHTS Operating Procedure for Wolsung Units. Oct. 1 997.

- 207 -

IV-5

Approximate Evaluation Method for Ductile Fracture Analysis

of Circumferentially Through-Wall-Cracked Pipe Subjected

to Combined Bending and Tension
Naoki Miura (CRJEPD , Materials Science Department, Central
Research Institute of Electric Power Industiy, Japan

Approximate evaluation method for ductile fracture analysis
of circumferentially through-wall-cracked pipe
subjected to combined bending and tension
N. Miura
Materials Science Department, Central Research Institute of Electric Power Industry
2- I I - I , Iwado Kita, Komae-shi, Tokyo 20 1 -85 1 1 JAPAN

Abstract
To estimate the structural integrity of the Light Water Reactor piping, combined loading

consists of tensile load due to internal pressure and bending load under seismic condition
should be considered as a basic loading mode. However, theoretical investigation on themethodology to evaluate ductile fracture behavior is not adequate to date. In this study, anapproximate evaluation method for ductile fracture analysis of circumferentially through-wallcracked pipe subjected to combined bending and tension was newly developed. This method
can explicitly incorporate the contribution of both tension and bending. The effect of growing
crack is also considered in the method. This method was then applied to the full-scale pipe
fracture tests. Based on the comparison with experimental results as well as finite element
calculations, it could be ascertained that the proposed method could well predict ductile fracture

behavior under combined loading. The effect of combined loading on ductile fracture wassensitivity-studied using the proposed method. As a result, it was quantitatively found that thesuperposition of internal pressure reduced the maximum load of cracked pipe.

1. Introduction
To substantiate that pipe whip restraint and jet impingement shields can be eliminated,piping for nuclear power plant is required to satisfy Leak-Before-Break (LBB) analyses forpostulated defects. One of the most important aspects in LBB analyses is to predict the ductile
fracture behavior and the load-carrying capacity of a circumferentially through-wall-cracked

pipe. Elastic-Plastic Fracture Mechanics (EPFM) methods can employ analytical approaches
such as elaborate numerical simulations or some approximate evaluation methods. Numericalsimulations are generally considered to give reliable predictions for a specific condition.
However, approximate evaluation methods seem to be more useful to investigate the effect ofvarious analytical situations systematically. Based on the J-Tearing theory, some methods havebeen proposed to evaluate the ductile fracture behavior of circumferentially through-wallcracked pipe subjected to pure bending. When actual plant piping conditions are considered,however, the effect of the internal pressure acting as a tensile load also occurs. Theoreticalinvestigation on the methodology to evaluate ductile fracture behavior under combined loading
is not adequate to date. If combined loading effects are not properly considered, then the actual
fracture resistance of the pipe could be in error.

In this study, an approximate evaluation method for ductile fracture analysis of
circumferentially through-wall-cracked pipe subjected to combined bending and tension is
newly developed. This method is intended to incorporate the contribution of both tension andbending. The method is then applied to the full-scale pipe fracture tests. The effect of
combined loading on ductile fracture is sensitivity-studied using the proposed method.

- 209 -

2. Development of new evaluation method
2.1 Existing consideration on combined loading
Kumar and German ( 1988) have developed the GE/EPRI Estimation scheme for acircumferentially through-wall-cracked pipe subjected to combined bending and tension. Fully
plastic solutions for power-law material were computed under proportional loading condition.
Brust, et al. (1995) have presented more reliable solutions using elaborate three-dimensional
finite element calculations under constant internal pressure condition. These solutions tabulated

as functions of pipe dimension, crack dimension, strain hardening exponent, etc, are limited

and discrete at this time. For example, the former solutions are given for only two
radius/thickness ratios, two crack angles, and two or three hardening exponents. For the latter
solutions, internal pressure is fixed to 15.5 MPa, which is equal to PWR operating condition,and cannot be a variable. Therefore, theses solutions are not always adequate to analyze ductile

fracture behavior when analytical parameters such as crack angle or internal pressure
significantly change during loading.

Paris/Tada method (Paris, 1983), LBB.NRC method (Klecker 1986), LBB.GE method(Brust, 1987), and LBB.ENG2 method (Brust, 1987 and Gilles, 1994) are other approximate
evaluation methods which may have a potential to evaluate ductile fracture under combined
loading. These methods were originally developed to apply to pure bending or pure tension.
They are often applied to combined loading by substituting an internal pressure for anequivalent bending by some means. However, these approaches tend to give conservative
prediction especially for the case with larger pressure (Brust, 1995). This might be because theinteraction between tension and bending cannot be properly taken account in these cases.
Attempts to evaluate J-integral under combined loading by adding J-integral due to pure tension

and that due to pure bending were also seen in some instances, which have naturally no
theoretical base since the principle of superposition 6annot be available in nonlinear region.

2.2 Evaluation of moment-rotation relation
Consider a circumferentially through-wall-cracked pipe subjected to combined fourpoint bending and tension due to internal pressure shown in Fig. I . Assume that the total loadpoint rotation, c may be written as

c=cc+cNC(1)

where cc and cNC are c due to crack and c without crack, respectively. cNC can be easily
obtained as_ M Linwhere M is the applied moment, Lin is the inner span, E is the Young's modulus, and I is thesecond moment of area for uncracked pipe. cc may be separated into its elastic and plastic

components.

cC=cEL+cpL(3)

The cEL and cpL under combined loading are evaluated as the following procedures.

When a pipe is subjected to bending moment, M and tensile load, T, bending stress, aband tensile stress, at areo:b*
Mt 2=7CRT t
7cR2,at(4)- 210 -

where R is the mean radius and t is the pipe thickness. The stress intensity factor, Kl isrepresented as the sum of its bending and tensile components, Kb and Kt

K K + Kt = ab ~~~Fb (O) + at ~r~Ft (O) (5)where e is the half crack angle, and the compliance functions, Fb and Ft are provided asfunctions of 6/7T and R/t based on Sanders' solutions (Klecker 1986).

Fb (e) = I + Ab (~)1'5 + Bb (~)2.5 + Cb (~)3'5 (6)1

Ab -3.26543 1.52784 -0.072698 0.0016011 R/t

Bb = Il.36322 -3.91412 0.18619 -0.004099 (7)Cb -3 . 1 8609 3 . 84763 -O. 1 8304 0.00403 (Rlt)2(R/t)3
2.5
F (6) I + At (~)1'5 + Bt(~) + Ct (~)3'5 (8)A t -2.029 1 7 1 .67763 -0.07987 0.00 1 76 Rlt

Bt = 7.09987 -4.42394 0.21036 -0.00463 (9)Ct 7 .7966 1 5 . 1 6676 -0.24577 0.0054 1 (R/t)2(R/t)3cEL are related with the strain energy due to crack, U using the Castigliano's theorem.

While the energy release rate, G may be written as

where A is the crack area. From Eqs. (10) and (1 1),A/dA

aM o E

Since M and T are independent each other, substituting Eq. (5) into Eq. ( 1 2), thencEL_
~E ab lb (O) + ~ I (6) (13)w herelb (e) = 4 OF~ (O) dO (14)-211-

I, (6) = 4 OFb (6) F (O) dO( 15)

Thus the cEL can be explicitly calculated.

As for the evaluation of cpL, the beam theory can be enforced to the pipe of nonlinear
elastic material under the deformation theory of plasticity. This hypothesis is the same as the

supposition used in the LBB.ENG2 method (Brust, 1987). Note that the LBB.ENG2 method
considered the compatibility for pure bending or pure tension, while the new evaluation methodattempts to take account of both bending and tension loading simultaneously.
The stress-strain relation is defined as the following power law.

ao

~o ao
Based on the beam theory with the assumptions that plane sections remain plane, the deflectionof the beam and the related strain and stress distributions are illustrated in Fig. 2. The stress
distribution has to be in equilibrium with the applied tensile load and the bending moment.
This follows

R T = MO h kTgT p(17)

M = Mo h kM - [kT gT - K'~M gM] p2(18)

where p identifies the location of the neutral axis shown in Fig. 2, Mo is the yield moment for
uncracked pipe given as

M0= 4010R2t(19)

and the parameter, h is

h=( R )l

(20)

Ce80p

~~

with p as the radius of curvature. Functions KM , KT , gM , and gT are introduced toapproximate integral equations appeared in the process of the equilibrium calculations, which
are given by

7c/2 F( I +1)

l+1 2 r( I +~)kM= (sin e)~ d6_~ 2n(2 1 )

2n 2

kT= IT/2_~~r(2ln(sinO)~de2 F( +~)
I +1)
n+1 2

(22)

2n
(23)

gM= n 7T
gT - I + ~

(24)

n7c

The Gamma function, Fin Eqs. (21) and (22) can be approximated in the range of interest hereby the Stirling's formula- 212 -

F(Z) = e~ z zzl 39
0.5V~~xljl+
I
+
I
_
_
571 }
12z 288z2 51840 z3 2488320 z4(25)

The process to lead the equilibrium equations with the above related functions are provided inAppendix A.
By solving Eqs. ( 17) and ( 1 8), the relation between moment (or load) and curvature canbe written as follows. For the case of M ~ O ,_ I _ I _1CM In
xd2y~~~~
[
oceo

(26)

d 2 p R KM MOJ

where the parameter, Ic is given by
16 K"M (kT gT - K'~M gM) (o:~a:)lc-i 1+

-2

l+

2

k~ g~

(27)

Alternatively, for the case of M = O ,dx2=-=p R aeo[ -KRTJn

KM Mo~ ~(KT gT - KM gM) (29),c ~KM
kT gT

where x and y are components of the coordinate shown in Fig. 3.
Again, after the manner of the LBB.ENG2 method for pure bending, considering thatthe existence of circumferential through-wall crack can be replaced by a pipe section with
reduced thickness, te extending for a distance a at the cracked section .as shown in Fig. 3. The
reduced thickness section is expected to simulate the reduction of the compliance due to theexistence of the crack. Eqs. (26) and (28) can be analytically solved for all sections in Fig. 3
by applying appropriate boundary conditions. The detail calculations led from Eqs. (26) or(28) are omitted here, however, they can be accomplished in the similar way as the steps by
Brust (1987). It is noted that ,cdefined by Eq. (27) includes the stress components, so that Eq.
(26) should be recursively computed as a function of x. Although, the beam of interest shows
nonlinear elastic response and is independent of loading history. Therefore one may determine
an arbitrary loading path up to a specific loading state. If the proportional loading is supposed,
then lc defined in Eq. (27) is not affected by the stress components, and is independent of x.
For the case of M ~ O , the solution of Eq. (26) at the reduced thickness section isl (~)"acO( - )7CK( ) x" Cbn

c=k te4 CO
KM
cPL corresponds to c at x = a / 2

,

(30)

then

cPL ~ i (L)n oc ejO ( 7C K )n (O:b)n ~

(31)

AM O
2
~Re t/V~~O1
cEL is obtained by taking a = l

and n = l

- 213 -

cEL_ iiaba80 'cl olo ~ (32)
~R te

where

From Eqs. (31) and (32), cpL can be evaluated asn I ,cn olbn-lcPL=(~)- olo( ~ )na ( ) cEL (34)
te 4KM Icl
Similarly, for the case of M = O , the solution of Eq. (28) at the reduced thickness
section is

1R t.t 7c
2~MKao)n (at)n x (35)cPL = - (-)" o~ co (cPL corresponds to c at x = a / 2 , thencPL _ i (~)n a so ( 7c lc )n (ol~a;:)n ~ (36)~R~oe 2t KM

cEL is obtained by taking a = I and n = l

iL~02 Icl

Rte ol02
From Eqs. (36) and (37), cpL can be evaluated ascPL_(t)_ _n I _- ( lc )na 'cn (o:t)n-lcEL (39)
te 2KM 2 Icl ao

The equivalent reduced thickness in Eqs. (34) and (39) may be taken so that the NetSection collapse moment of the cracked pipe under bending condition is identical to that of the
reduced thickness section.

,1 2[
4 afR2 t[cos~ ~ ~ ~ af (40)J

sin 6] = 4 o:fR te[coswhere af is the flow stress. The term in the bracket in right side indicates the effect ofsuperposing tensile stress, and is equal to the ratio of the plastic collapse moment of cracked

pipe subjected to combined bending and tension, 4 afR2 te [cos (2[ o I l~+~at) _l~ sin 6 to that

2 2af

subjected to pure bending, 4 cyR te [cos (~) ~ ~ sin e] at 6 = O . Since Eq. (40) is premisedon the rigid perfectly plastic behavior, it gives better results for small cracks. While it tends tooverpredict cpL for larger cracks (Gilles, 199 1 ). In the present evaluation method, thecorrection factor of rd:4 deterrnined by Gilles ( 1 99 1 ) is used. That is,-214-sin 6J '_t cos= ~![a~o:IfJJt / [cos ~ ~ ~el

1[- - t6

te cos ~ a_o:'JJ / [cos ~ - i Sin OJ
lc

4

22 '

(4 1 )

tor0>~ !
4and are linearly interpolated between rd6 and 7d:4.

2.3 Evaluation of J-integral
When an experimental or an analytical relation between load and displacement is
known, the n-factor approach can be applied to evaluate the J-integral, J for a circumferentiallythrough-wall-cracked pipe subjected to four-point bending (Zahoor, 198 l, Wilkowski, 1987).

The existing evaluation methods mentioned above (i.e., Paris/Tada method, LBB.NRC
method, LBB.GE method, and LBB.ENG2 method) that predict maximum load capability havethe n-factor analyses embedded in them. The earlier n-factor solutions have been developed
for cracked pipes under either pure bending or pure tension. J is defined as the release rate ofpotential energy of nonlinear elastic material, therefore, the n-factor solutions can be extended
for combined loading if the energy release rate for cracked pipe subjected to multiple loads isgenerally described. Pan ( 1 984) has investigated the energy release rate for stationary cracked
body under multiple loads, and has derived a generalized J evaluation equation. Further, Miuraand Wilkowski (1995) have developed the combined-load n-factor solution which wasapplicable to combined bending and tension taken account of growing crack. In this study, thiscombined-load n-factor solution is expected to be included in the present evaluation method.

When a circumferentially through-wall-cracked pipe shown in Fig. I is subjected to
combined four-point bending and tension due to internal pressure, J can be given as follows inspite of loading paths.

IT, M dM +J aA aA
Ic=0, TdT

(42)

where 6 is the axial displacement shown in Figure I . In the above equation, conditions that

Here J, c, and 6 are assumed to be divided into the elastic and plastic components,

J=JEL+JpL(43)

where subscripts EL and pL denote elastic and plastic component, respectively.

The elastic J,

JEL is directly obtained from KI in Eq. (5).

JEL=K12/E(44)

And for the plastic J, JpL,

JPL = acpL IT, M dM + aa8A~~L Ic=0, TdTaA- 215 -

(45)

The 6pL due to axial load is considered to be small, so that the second integration of the rightside in Eq. (45) is negligible. By using the geometrical relations as,

A=2Rt6(46)

cPL/2=ApL/Lz

(47)
(48)

M=Lz/P
where ApL is the plastic load-point displacement, and P is the bending load shown in Fig. I ,then Eq. (45) can become

JPL= acpL IT MdM= I aApL Is,PdP

aA ' R t ae

(49)

Note that the non-dimensional tensile stress, s (defined as atlaf) is used instead of T.
Assuming that ApL is a function only of the ratio of P to the Net-Section collapse load, then 1lpL

can be expressed as

ApL = ef 7c(Pl / hp (O, s)) (50)h (6 s) = cos (-+ - s) - - sin 6 (51)p'222

where hp rs equal to the ratio of the plastic collapse load of cracked pipe subjected to combined

Ioading to that of uncracked pipe subjected to pure bending. Substituting Eq. (50) into Eq.
(49), then gives

ApLo~2RthpThis equation gives JpL for a stationary (non-growing) crack under bending and constanttension loading.

For a generalized case considered growing crack and tension changing, Eq. (50) is
written in its inverse form,

P = hp F (ApL)(54)

In this equation, ApL, e, and s are assumed to be independent of each other, and since JpL is afunction of ApL, e, and s, the differential of JpL is

dJpL- aJpL de+ aJpL

aO

a ApL

a JpL

dApL +

ds

(55)

aS

Substituting Eqs. (52) and (54) into Eq. (55), and are given by-216-

JPL= 2 n P dApL + 2 Yp JPL d0+ 2 pp JPL ds

(56)

oo

where,

yp- 2ahp/aO

pp = - a2hp / as a6 (58)
2 ahp I ae

The second term of the right side of Eq. (56) corresponds to the contribution of the crack
growth. The functions n and yin the original n-factor approach are replaced with functions np

and yp, respectively. The third term of the right side of Eq, (56) corresponds to thecontribution of the tension changing. If an axial tension is constant, this term is equal to zero.
For the case of s = O , Eq. (56) is naturally identical to the original n-factor solution. Figs. 4 to

6 show the values of np, Yp, and /lp as a function of O, respectively. It can be seen that thesefunctions are affected by s as well as e.

2.4 Evaluation method for ductile fracture analysis
By integrating the moment-rotation evaluation in Section 2.2 and the J-integral
evaluation in Section 2.3 together with the criterion of crack extension, an evaluation methodfor ductile fracture analysis taken account of both combined loading and growing crack was
newly established. The developed evaluation method has an ability to consider an arbitrary
combined loading history by nature, however, constant tensile load due to internal pressure
followed by external monotonic bending load is supposed in this paper.

Fig. 7 shows the schematic of the developed evaluation method. As shown in thisfigure, the unique relation between J and M is provided for a specific crack extension Aa. J
corresponding to the Aa is determined by the J-R curve, then M can be fixed at the point on the

J-M curve. The relation between M and c is also unique for the specific Aa, so c is given by
the point on the M-c curve. Repeating this process for a set of Aa, and taking an envelope of

the points, then the prediction of the ductile fracture behavior with crack extension isaccomplished. The flow of the evaluation procedure is sununarized in Fig. 8.

3. Application to pipe fracture tests

3.1 Analytical object and condition
The developed evaluation method was applied to analyze full-scale pipe fracture tests
under combined bending and internal pressure. The predicted ductile fracture behavior wascompared with experimental results as well as finite element calculations to validate the
propriety of the method.

Table I shows the matrix of six pipe fracture tests analyzed. These tests were
performed as a part of the NRC's Degraded Piping Program - Phase 11 (Wilkowski, 1989) andthe IPIRG-2 Program (Scott, 1997). The materials used for the tests were AI06 Grade B and
SA333 Grade 6 carbon steels, and Type 304 stainless steel. Tensile properties necessary to
perform a fracture analysis are also listed in Table I . All tests were performed at 288"C. Fourtests were conducted under tension loading due to internal pressure followed by bending, andtwo tests were conducted under pure bending condition for comparison. Experiments 413 1-7,-217-

4 1 3 1-3, and Experiments 4 1 3 1 -5, 4 1 3 1 - I have almost the same dimensions and test conditions

except for the loading condition. Using these cases, the effect of combined loading can beevaluated. Experiment I .9 was performed under dynamic loading (time to crack initiation =0.32 seconds). All the other experiments were conducted under quasi-static loading.
Two J-R curves were provided for the predictions for each material. One was the J-R
curve obtained from compact tension (CT) specimen which had been cut from the test pipe.
The other was the J-R curve obtained from the pipe fracture test using the combined-load nfactor approach (Miura, 1 995). The use of the pipe J-R curve is desirable for the purpose of
the validation of the proposed method, however, small specimen J-R data are typically applied
because they are often the only data available, and because predicted results are generally
considered to be lower bounds. The J-R curves used for the analysis are shown in Figs. 9(a)to (d). These J-R curves were extrapolated according to the equations shown in these figures.
A three-dimensional finite element analysis was carried out for Experiments 4 1 3 1 -7,
4 1 3 1-3, 4 1 3 1 -5, and 4 1 3 1 - I to confirm that the experimental results were appropriate.

The analysis was conducted using the computer code "MARC" with twenty-noded solid
elements. Fig. 10 shows an example of the finite element mesh for the analytical model.
Tensile properties in Table 2 were used for the analysis. In the analysis, the internal pressure

was applied first, then the bending moment was applied with the constant internal pressure.

The experimental relation between A and 1la was numerically followed for the crack growth
condition.

3.2 Results of analysis
Figs. I l(a) to (O show the predicted relations between bending load and load-point
displacement compared with the experimental (and the analytical if available) relation(s).
Fig. I I (a) shows the load-displacement relations for Experiment I .9 (carbon steel 6-

inch diameter pipe under bending and pressure). The proposed method using the CT J-R curvesignificantly underestimates the load-displacement relation. The prediction using the pipe J-R
curve is still lower than the experimental result. The CT J-R curve for this material was
obtained using sidegrooved specimens which lowered the toughness by constraining plasticity,and resulted in large difference. One reason of the poor predictability might be due to thecomplex crack growth aspects such as angled, asymmetric, and fast crack growth for thisspecific experiment.
Fig. I I (b) shows the load-displacement relations for Experiment I .8 (carbon steel 1 6-

inch diameter pipe under bending and pressure). The drop of the displacement at about 275
mm in the experimental results just at maximum load is due to an unstable crack jump at thispoint. Although the proposed method using the both J-R curves gives slightly lower
displacement than the experimental result, the predicted maximum moments are in good
agreement with the experimental value .

Figs. I l(c) and (d) show the load-displacement predictions for Experiments 413 1-7(carbon steel 10-inch diameter pipe under pure bending) and Experiment 4 1 3 1 -3 (carbon steel

lO-inch diameter pipe under bending and pressure), respectively. It can be seen that the
proposed method realizes fine predictions independent of the presence of pressure bycomparing with experimental and FEM results. The relations from the CT J-R curve are lower
than that from the pipe J-R curve. However, the differences are small for both experiments.
Figs. 1 1(e) and (D show the load-displacement predictions for Experiments 4131-5(stainless steel 6-inch diameter pipe under pure bending) and Experiment 4 1 3 1 - I (stainless steel

6-inch diameter pipe under bending and pressure), respectively. The similar trends as the caseof Figs. (c) and (d) can be observed, and the good predictability is secured both for combined
loading and for pure bending. There seems to be a slight difference between the experimental
and the FEM Ioad-displacement relation for Experiment 4 1 3 1 -5. The prediction based on the
proposed method is closer to the FEM result rather than the experimental one.

Fig. 12 shows the predicted divided by the experimental maximum loads. The
predictions by the proposed method using both the CT and pipe J-R curves and the prediction
by FEM (if available) are provided for each experiment. The proposed method has an enoughcapability to predict maximum moment for various materials, pipe size and crack size. The useof the pipe J-R curves gives better predictions than the use of the CT J-R curves compared with- 218 -

the experimental maximum loads. The predictions using the CT J-R curves are relatively
conservative, however, the difference of the predicted maximum loads caused by the selectionof the used J-R curve is not remarkable.

4. Sensitivity study on combined loading effect
4.1 Effect of internal pressure on ductile fracture
The effect of the superposing pressure on ductile fracture was investigated using the

proposed method. An analytical object was SA333 Grade 6 carbon steel 4B x Schedule 80
cracked pipe at 288'C. Dimensions of the pipe were 52.85 mm in mean radius and 8.60 mm inthickness. A crack with the total angle of 30' was supposed. The constant internal pressure
preceding bending moment was considered to be a parameter, so that the ratio of the tensilestress to the reference stress, atlolo was equal to 0.00, 0.05, 0,lO, 0.15, 0.20, and 0.25. The
tensile properties listed in Table I and the pipe J-R curve presented in Fig. 9(c) were used forthe analysis.

Fig. 13 shows the relations between moment and rota

relations have similar configuration, and the maximum moment decreases with increasing
tensile stress. For the case of at/ao = 0.25 , the maximum moment is about 12(~o lower thanthat without tensile stress. It can be also seen that the rotation at the maximum moment
decreases with increasing tensile stress.

Fig. 14 shows the relations between J-integral and momeat the same J-integral decreases with increasing tensile stress, as is the case of Fig. 1 3 . While
the value of J-integral at the maximum moment is not affected by tensile stress.

4.2 Dependence of dimensional parameters on maxrmum moment
Next, a sensitivity analysis was performed to study how the effect of the superposingpressure was affected by typical dimensional parameters represented by crack angle and pipe

diameter. Again, an analytical object was SA333 Grade 6 carbon steel Schedule 80 (pipethickness in 8.60 mm) cracked pipe at 288'C. The same tensile properties as well as J-R curveas the above-mentioned analysis were applied. Table 2 shows the matrix of the analyticalcondition. For the case that the dependence of crack angle was to be seen, nominal diameter
was fixed to 4B (mean radius in 52. 85 nun) and total crack angle was varied from 30' to 1 80'.
On the other hand, for the case that the dependence of pipe diameter was to be analyzed, crack
angle was fixed to 30' and nominal diameter was varied from 4B (mean radius in 52.85 mm) to24B (mean radius in 289.3 mm). The constant internal pressure preceding bending moment
was considered to be a parameter, so that the ratio of the tensile stress to the reference stress,

atlolo was equal to 0.00, 0.05. 0.lO, 0.15, 0.20, and 0.25 for both cases.

Fig. 1 5 shows the dependence of crack angle on maximum moment. The maximum
moment gradually decreases with increasing tensile stress, and this trend becomes significant as
the crack angle increases.

Fig. 1 6 shows the dependence of pipe diameter on maximum moment. The maximum
moment again decreases with increasing tensile stress, and this tendency is clear for the larger

diameter pipe. Thus, it is numerically found that the reduction of the maximum moment withincreasing internal pressure becomes notable for larger crack angle and for larger pipe diameter.

5. Conclusions
In this study, an approximate evaluation method for ductile fracture analysis of
circumferentially through-wall-cracked pipe subjected to combined bending and tension was
newly developed. Existing evaluation methods are limited and are inadequate for practical use
to analyze the effect of combined loading, while the proposed method can explicitly incorporate
the contribution of both tension and bending. The effect of growing crack is also considered in

the method. The method was then applied to the full-scale pipe fracture tests. It could beascertained that the proposed method could well predict ductile fracture behavior under- 219 -

combined loading. The effect of combined loading on ductile fracture was then sensitivitystudied using the proposed method. As a result, the effect of superposition of internal pressure
as well as the dependency of crack angle and pipe diameter could be quantitatively evaluated.

Appendix A Equilibrium equations and related functions
The equilibrium of the tensile load is expressed as'cl 2

R T= adA = 2 R2 t ad0= ~MO ~lT(A- I )

A - Icl2
where Mo and h are given in Eqs. (19) and (20), and IT is

IT= - (- sin 6 - sin p)n dO + (sm O + sm p)n dO(A-2)- ;T/ 2 - p
IT Can be written using the areas ST~

and St+ in Fig. Al.

IT ST~ + ST+(A-3)

Here, assuming that the curves surrounding ST~
f(6) = (sin 6)~ in the range of [O, 7z/2], ST- and ST~

and ST~ are similar to thefunctionare given by'c/2 ~_ pST~= (sin 6)~dOx 27c x(1 -sinp)l

(A-4)

~

+p
+_'r/2
2~~~x(1+sinp)i

ST -(sin 6)~ de x

(A-5)

2
Substituting Eqs. (A-4) and (A-5) into Eq. (A-3) then gives-J

IT=KT[(1 + 7c ) (1 + sin p);T- (1 - 7c ) (1 - sin p)n~ is given in Eq.
where KT(22).

If p is enough(A-6)

small,(1 :!: sin p)nl *1:!:'sinp isapproximately satisfied, then

IT = kT [~ sin p + J
(A-7)
7c

Further, supposing sin p * p, andIT 2KT[[~+2-7c]p=2kTgTp(A-8)

Substituting Eq. (A-8) into Eq. (A- I ), one may write- 220 -

~
which is identical to Eq. ( 17).

Similar procedure is also employed for the equilibrium of the bending moment asfollows'T12

M = c(R sm 6) dA 2 R t asin 6d6= ~MO hlM (A-10)
where IM is

IM = (- sin O - sin p)~ (- sin 6) de + (sm O + sm p)n sm O dO(- sin 6 - sin p)i+ I de + (sin e + sin p)1+ I de - IT sin p- 'T/ 2 - p(A- I I )IM can be written using the areas SM- and SM+ in Fig. A2.

IM = SM- + SM IT Sm p (A-12)
Assuming that the curves surrounding SM- and SM+ are similar to the function
f(6) = (sin e)~ + I in the range of [O, n/2], SM- and SM+ are given by~_p
27c x(1-sinp)l
SM- = (sin e)~+I d6 x (A-13)
ITI 2

+1

~

71:/2 ~ + p + l

SM+ = (sin e)~+ I dO X 2 ~ x (1 + sin p)L (A-14)
2
Substituting Eqs. (A-13) and (A-14) into Eq. (A-12) then gives
l +1
2p)(1 +sinp)~ +(1IM=KM (1 + 7c
- 7c )(12p-sinp)nJ

l+1

where kM is given in Eq. (21). Again, if p is enough small,(1 :!: sin p)~ + I = I :!: (~ + l) sin p is approximately satisfied, then- -+ l) p sin plj ~ ITsin p

Further, supposing sin p * p , and- 221 -

~ n ~[ I ~ p2J - 2 kT[[1l + 2-JI p2 = 2 (K"M - [kT gT - kM gM] p2}

IM=2KM[1+nlc(A- 17)

Substituting Eq, (A- 17) into Eq. (A- lO), one may write

M = Mo ~ {K"M - [kT gT KM gMIJ p
which is identical to Eq. ( 1 8).

The above derivation is based on the assumption that the p is enough small. For
example, when p is less than 0.1lz:, it was numerically ascertained that the difference of IT
computed by Eqs. (A-2) and (A-8), and the difference of IM computed by Eqs. (A-1 1 ) and (A1 7) are both less than 6(~o. p is a function of strain hardening exponent, n and the tensile tobending load ratio, RT/M. n usually ranges between I to 10 for typical structural materials,

a significant ductile fracture behavior occurs at much higher value of RT/M than 0.4, theassumption on p is considered to be appropriate for the loading mode of interest.

References
Brust, F. W. Approxnnate Methods for Fracture Analysrs of Through Wall Cracked
Pipes, NUREG/CR-'4853, February 1987.
Brust, F. W., et al., Assessment of Short Through-Wall Circumferential Cracks in

Pipes, NUREG/CR-6235, April 1995.
Gilles, P., Brust. F. W., 1994. Approximate Methods for Fracture Analysis of Tubular

Members Subjected to Combined Tensile and Bending Loads, J. Offshore Mechanics andArctic Engineering (OMAE), 1 1 6, 22 1 -227.

Gilles, P., Brust, F. W., 1991. Approximate Fracture Methods for Pipes - Part I:
Theory, Nucl. Eng. Des., 127, 1-17.

Klecker, R., et al., NRC Leak-Before-Break (LBB.NRC) Analysis Method for
Circumferentially Through-Wall Cracked Pipes Under Axial Plus Bending Loads,
NUREG/CR-4572, May 1986.
Kumar, V., German, M. D., Elastic-Plastic Fracture Analysis of Through-Wall andSurface Flaws in Cylinders, EPRI NP-5596, Final Report, January 1 988.

Miura, N., Wilkowski, G., 1995. J-R Curves From Through-Wall-Cracked Pipe Tests- Combined Loading and Short Crack Considerations -, 1995 ASME Pressure Vessels andPiping Conf., 304, 171-179.

Pan, J., 1984. Some Considerations on Estimation of Energy Release Rates for 'Circumferentially Cracked Pipes, Trans. ASME J. Pressure Vessel Technol., 106, 391-398.

Paris, P. C.. Tada, H., The Application of Fracture Proof Design Method Using
Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through Wall Cracks,

NUREG/CR-3464, September 1 983.
Scott, P., et al., IPIRG-2 Task I - Pipe System Experiments with Circumferential
Cracks in Straight-Pipe Locations, NUREG/CR-6389, February 1 997.
Wilkowski, G., et al., Analysis of Experiments on Stainless Steel Flux Welds,

NUREG/CR-4878, April 1987.
Wilkowski, G. M., et al., Summary of Technical Results and Their Significance to

Leak-Before-Break and In-Service Flaw Acceptance Criteria, NUREG/CR-4082, 8, March
1989.

Zahoor, A., Kanninen, M. F., 1981. A Plastic Fracture Mechanics Prediction of
Fracture Instability in a Circumferentially Cracked Pipe in Bending - Part I: J-Integral Analysis,

Trans. ASME J. Pressure Vessel Technol., 103, 352-358.

- 222 -

T

I~

~-

t

26
Crack
Crack section
Fig. 1

Circumferentially through-wall-cracked pipe subjected to combind bending and tension load.

(L)H

S

~,
cs

>
hc:$

(J

l

Neytral axis

Centerline,~~-T~~~t2et
1~l~~~~~~~~n Strai di tribution tress distribution
Fig. 2 Deflection of beam and related strain and stress distribution.
2.0
l .8

l.6

s~* 1.4~

Q,:S 1.2
c::o 1.0~ 0.8:' 0.6

F~

0.4
O. 2

0.0 o I O 20 30 40 50 60 70

80

90

80

90

Half Crack Ang]e e, degree

Fig. 3

Pipe geometry with equivalent reduced thickness. Fig. 4 Transition of np with O.
l .O

0.3

0.9

0.2

0.8
O. l

* 0.7

>(~ 0.0:tct 0.6~.~ 0.5•~ O ltSL)

c:::s 0.4:::~:' -0.2

p~ 0,3

-0.3

0.2

-0.4
-0.5 OO. 1

l O 20 30 40 50 60 70 80 90 0.0 O I O 20 30 40 50 60 70
Half Crack Angle e, degree Half Crack Angle O, degree
Fig. 5 Transition of w with O. Fig. 6 Transition of //P with O.
- 223 -

Aa = Aa'

l

Ji
I

C:$
hbJb

Q)

~a = Aa2
Aa = Aal

J,
:::
I-'

b,,l

Aa=0
l

Jin

I

l

I

Mi

Min

MomentFig. 7 Schematic of the approximate evaluation

method for ductile fracture analysis.

Fig. 8Flow of approximate evaluation for ductile fracture analysis.

Table 3 Matrix of cracked pipe fracture tests and their tensile properties.

* : Stress-strain relation described as

880= a+a(a)nolo ao80 = Co

E- 224 -

J-R Obtained from Pipe I .8

~ 1.5 J-R Obtained from Pipe 1.9 ~ 1.5 J = 0.305 + 0.088 (Aa)1.000~~:
Z J = 0.424 + 0.049 (Aa)1.000 ~:~' 0.5 ~,J-RfromObtainedCT
J-R 0.5
Obtained
CT J = 0.069from+ 0.173(Aa)0.779
J = 0.037 + 0.068 (Aa)0.46 l15

5

20

Crack Extension, mm Crack Extension, mm(a) AI06 Gr. B C/S for Experiment I .9 (b) A 106 Gr. B C/S for Experiment I .8

2.0 J-R
4.0Obtained
J-R from
Obtainedfrom Pipe 4131-5
Pipe 4131-7 J = 0.750 + 0.529 (Aa)0.645~

J = 0.264 + 0.207 (lia)0.649~ 1.0 ~ 2.0 J-R Obtained from Pipe 4131-l
o~' o
0.745
+ 0.255(Aa)0.860
0.5JJ-R~'=IObtained.O
J-RfromObtainedfromCT(Aa)0.834
CT J = 0.646 + 0.233-R Obtained from Pipe 41 3 1-3

J = 0.292 + 0.109 (Aa)0.753

J = 0.158 + 0.146 (Aa)0.658

15

20

O

20

5

Crack Extension, mm Crack Extension, mm(c) SA333 Gr. 6 C/S for Experiments 41 3 1-7 and 4131-3 (d) Type 304 S/S for Experiments 4131 -5 and 4131-lFig. 9 J-R curves used for ductile fracture analysis.

~~:
Uc~$

~
Extensi on around crack
Fi g .

lO Example of finite element mesh.

- 225 -

~

Z

~'

='

~
o

~
ed)
F:

F::

~S

1:'

~
eQ
~~

OCl

~
~

~

H
*o
~

~O
~'
~c~l

:1:

~::

IOO

50

l 50

300

Load-Point Displacement, mm Load-Point Displacement, mm(a) Experiment I .9 (b) Experiment I .8

(AI06 Gr. B C/S, pressure and dynamic bending) (AI06 Gr. B C/S, pressure and bending)

~

Z
~~

~:'::'~

~
o

~

~

~

eJ)

~'s::

;ej

~

o
eQ

s:

I~_

~
o

~

H
*o

~O

~
:~
~cE$

~~
:::

l OO

200

l 50

Load-Point Displacement, mm Load-Point Displacement, mm(c) Experiment 41 3 1 -7 (d) Experiment 41 3 1-3(SA333 Gr. 6 C/S, pure bcnding) (SA333 Gr 6 C/S, pressure and bending)

~

Z;
~:

1::'C:'~

~

~

o
.1

bO

bJ)

1:'

;5

S

F:

'::Q)

(L)

CQ

a:l

~
~

~
~o 20

tHO

*o

H

~~
~c~S

~~

::

:::

O I 50OO 1 50 200 250 300 O 50 1 OO 1 50 200 250
Load-Point Displacement, mm Load-Point Displacement, mm(e) Exporiment 41 3 1 -5 (O Experiment 41 3 l- l

(Type 304 S/S, pure bending) (Type 304 S/S, pressure and bending)Fig. I I Comparison between experimental and predicted load versus load-point displacement relation.

~ 226 -

300

Table 2 Matrix of sensitivity analysis.

I Predicted by FEM~ predicted by Proposed Method Using CTJ-R[] Predicted by Proposed Method Using PipeJ-R

l .2

~:'

CQO 1.0

1'c~'~lO~:'S 0.8
~I~~
~ .~ca.~~: 0.6

~:

~
c~
~l

(L)~ 0.4
o ~
~
u 'C:
;5~
~ o
c~
~:'

,~

~ O.2
0.0

l.9 1.8

4131-7 4131-3 4131-5 4131-l(B+P) (B+P)(B) (B+P) (B) (B+P)

AI06 Gr. B C/S

SA333 Gr. 6 C/S Type 304 S/S(B+P) : Bending and Pressure(B) : Pure Bending

Fig. 12 Comparison between experimental

and predicted maximum load.

2.0

50

40

~ 1.5

~

~ 30

.\
Z;
~:

~

~bo~ 1.0

F::

c:$H

o
~ 20
o
~:

Q)

~i
~

I~
~, 0.5

lO

O

0.00

0.02

0.04 0.06

0.0

0.08

O. I OoFig. 1 3 Effect of internal pressure on moment-rotation relation

lO

20

30

40

Moment, kN-m

Rotation, radians

Fig. 14 Effect of internal pressure on J-integral-moment relation(SA333 Gr. 6 C/S, 288'C, 4BXSch. 80, 2e = 30').

(SA333 Gr. 6 C/S, 288'C, 4BXSch. 80, 2e= 30').

- 227 -

50

lOOO

50

~

S,

2e~30'

~ 40

~ 800
~i

s:
~

~ 30
o
~:
~ 20
~
~

2e~60-

~ 400

e~90'
2 e~ 1 20-

'~ l
:~

O

~
o "

~:

2e~150-2e~1800.0 O. 1:'~'~ 200~:O

0.2 0.3

O. l

Normalized Tensile Stress, a,loo

Normalized Tensile Stress, atloo

Fig. 16 Dependency of crack angle on maximum moment

0.2

Fig. 17 Dependency of pipe diameter on maximum moment(SA333 Gr. 6 C/S, 288'C, 4-24BXSch. 80, 20= 30').

(SA333 Gr. 6 C/S, 288'C, 4BXSch. 80, 20 = 30-1 80').

l ll lf (e) = sgn (Sin e + sin fi) Isin 6 + sin PI~

f (e) = sgn (sin 0+ sin p) sin O + sin p 71+

Fig. A.1 Schematic of parameters necessary

Fig. A.2 Schematic of parameters necessary
to calculate equilibrium of bending momentto calculate equilibrium of tensile load

- 228 ~

IV-6

Low Alloy Steel Piping Test for Fracture Criteria of LBB

Koji Koyama, Nuclear Plant Component Designing

Section, Kobe Shipyard & Machinery Works, Mitsubishi
Heavy Industries, Ltd., Japan

LOW ALLOY STEEL PIPlNG TEST FOR FRACTURE CRrrERIA OF LBB
Koji Koyama and Itaru Muroya
Mitsubishi Heavy lirdustries, Ltd.

Toshihiko Tanaka
The Japan Atomic Power Company
Takao Nakamura
The Kansai Electric Power Co., Inc.

Abstract
Seven pipe fracture tests were performed to provide the data for establishing fracture
criteria of Leak Before Break for the low alloy steel pipe.

Test pipes were 6-inch and 8-inch diameter pipes made of SFVQIA or STPA24 Iow alloy steel. A
circunferential through-wall crack was introduced at the center of a pipe, and four-point bending
load was applied without internal pressure. Stable crack extensions were observed in all of theexperiments.
To estimate the maximum applied load, the net-section criteria (NSC), R6 method Option 2 andOption 3 were used. The predicted values by three kinds of evaluation methods were comparedwith the experimental loads. Most of all the predicted maximum loads agreed well with theexperimental maximum loads within 20% difference. The NSC gave the most accurate prediction
but also gave unconservative results in some test cases. The predicted maximum loads by R6 Op.2
were conservative in all of the test cases. From the view point of conservativeness R6 method
can be used for evaluation of the low alloy steel pipe fracture. Therefore, the LBB concept would
be applied to the protective design standard against pipe break for the material.

1. INTRODUCTION
The concept of leak before break (LBB) has been applied to the protective design standard
against pipe break since 1992 in Japan. The LBB applicability evaluation flow is shown in Fig.1-1.
The current application of LBB concept is limited to the piping systems in the reactor coolingpressure boundary(RCPB) and of austenitic stainless steels. Nevertheless, this concept would also
be applicable to other piping systems and other materials, provided that plant shutdown can beachieved before fatal pipe break, taking account of fracture behavior the pipe material and thereliability of the performance of leak detection systems.

The recent study of fracture mechanics confirmed the adequacy of application of the LBB
concept to the piping system of carbon steel piping/1/, consequently, the LBB concept would beapplicable to the piping systems, as well, of low alloy steel, which is the same ferritic material as
carbon steel.

Moreover, the existing leak detection system will be also effective to the application of LBB
concept for the PWR secondary piping systems inside the reactor containment vessel, such as main

feedwater system.
This paper covers the experimental study performed on the fracture behavior of low alloy steel
pipes, which is expected to be applied to the reactor coolant piping and feedwater piping in

advanced PWRS in Japan.

- 229 -

2. TEST DESCRIPTION
2.1 IeSIMatdx
Seven pipe experiments were conducted to investigate the effects of the crack angle, material
and pipe diameter on the fracture behavior. Table 2.1-1 shows the test matrix. In these tests, the 6inch and 8- inch diameter pipes were used.

2.2 I~sLEip~s
The pipes were made of srvQIA and STPA24 Iow alloy steel (Japanese Industrial Standard
Specification equivalent to ASME SA508 Cl.3 and SA335 Gr. P22 respectively). The test sections
were of base metal pipes or shop fabricated metal active gas (MAG) welds joining two pieces ofthe pipe. The pipe geometry is shown in Fig. 2.2-1.
A initial circunferential through-wall flaw was introduced into the center of the test section byelectric-discharge-machining(EDM) techniques, followed by cyclic loading until fatigue cracks
extended to approximately 5 mm to sharpen the crack tip. Fig 2.2-2 shows the through-wall crack

geometry.

2.3 I~sLJ~rQ~~ure
The test system is shown in Fig.2.3-1. This test system has a capacity of 100 tonf (980 kN) per

ram and the maximum stroke is 500mm. The pipes was subjected four-point bending with theinner span of 800 mm and the moment arm length of 750mm. The experiments were conductedunder displacement control with the displacement rate of 0.05 mm/sec.

The experiments of srvQIA pipes and the STPA24 pipes were conducted at 325~C and
310~C, respectively. To achieve these temperature, the hot air was ventilated into the test section

by the fan, and the heaters were used for the fine temperature adjustment of the center of the test
section. The temperature was monitored by five thermocouples to be maintained at the designatedtest temperature. After reaching the temperature, the test pipes were allowed to stabilize at the test

temperature at least 2 hours before applying bending moment load. No internal pressure wasapplied.

2.4

The following data were collected using the personal computers. The grid lines with 2 mm
intervals were drawn near the crack tip of the pipe surface to observe crack extension behavior.
Crack extension behavior was also recorded by video camera.
' Applied load (at each ram)

' Load-1ine displacement (at each ram)
' Pipe ovalization' Crack opening displacements at three locations• Strain' Temperature near the crack tip• Crack extension in both crack tips

- 230 -

2.5

The chemical compositions of the srvQIA and STPA24 steels are shown in Table 2.5-1.
The stress-strain curves and J-resistance curves were obtained using the test specimens takenfrom the pipes with the same sizes of the bending test pipes. For the fracture toughness test, the

1/2T CT specimens and IT planform size CT specimens which had the thickness of 1/2-inch weretaken from 6-inch diameter pipes and 8-inch diameter pipes, respectively. The true stress-strain
curves of the base metals are shown m Frg 2 5 1-Frg 2.5-3. The J-resistance curves of the basemetals and the weld metals are shown in Fig. 2.5-4~Fig. 2.5-9. Also shown in those figures thefitting curves by Ramberg-Osgood's law for the stress-strain curves and the power law for the Jresistance curves are drawn in the same figures.
Table 2.5-2 shows the tensile properties and the Jlc Values of the test materials. The flow stressis defined by the average of the 0.2% offset yield strength and the ultimate strength.

2.6

The experimental results are summarized in Table 2.6-1. The total applied load versus load line displacement curves are shown in Fig. 2.6-1 and Fig. 2.6-2. At first from a fatigue crack tip,
two ductile cracks initiated and afterwards either of them propagated much longer than the other.
All of the cracks propagated stably with the angle of about 45' to the pre-cracked cross sections.

According to the measured crack mouth opening displacement, the symmetrical opening
behavior observed. The ovalization near the cracked section to the original diameter was within

2%.

3 . EVALUATION OF PIPE EXPERIMENTIn this study the net-section collapse criterion (NSC) and R6 method (Option 2 and Option 3)

were applied in order to estimate the maximum load. R6 Option 3 uses J-integral, which is
calculated by the finite element analysis or the simplified equation as proposed by Zahoor /2/. In
this study, the simplified J-integral calculation was chosen for R6 method, because J-integral

analysis by the finite element method would be time consuming to evaluate the pipe fracture
behavior for each of the seven pipe tests. Therefore by the following three steps the validity of thesimplified J-integral is confirmed and the applicability of three evaluation method is investigated.

i. Validity of the modeling of the finite element analysis
ii. Evaluation of the accuracy and conservativeness of the simplified J-integral comparing
with J-integral by the finite element analysisiii. Evaluation of the accuracy of the predicted maximum applied load from each evaluation

method

3.1

a(1) Description of the finite element(FE) models

The analyzed case was Experiment SFB1. The analyses were made with the FE code MARC
(version K6-1). The FE model is shown in Fig.3.1-1. The model is one quarter of the test pipe

was modeled due to symmetry. The element type was 20 node solid element. The total numberof elements and nodes was 1056 and 5727, respectively. The load-line displacement was appliedat the outer support points.

- 231 -

For the crack extension analysis, two models were prepared. One was Model I which had thecrack length corresponding to the initial crack length of half angle 45" . The other was Model 2

which had the extended crack length at the maximum applied load in the experiment. As
mentioned previously the cracks extended diagonally to the pre-cracked cross section, Model 2
had the projected crack extension to the pre-cracked cross section at the maximum load. The
experimental load-line displacement was applied to Model I until just before the ductile crack
initiation and the test load-line displacement at the maximum load was applied to Model 2.
(2) Global behavior of the test pipe

Fig. 3.1-2 shows the comparison of the calculated total applied load versus load-line
displacement relation with the experimental result. Fig. 3.1-3 shows the similar comparison in
the case of the crack mouth opening. In both comparisons, analytical results agree well with theexperimental results. These results gives the validity of the FE modeling including considerationof the diagonal crack propagation.

d3.2(1) Accuracy and conservativeness of the simplified J-integral
To calculate the simplified J-integral under the bending load, the stress-strain curve was fit
by Ramberg-Osgood's law expressed by the following equation.

E /E 0=(a/a )+a (a/a ) (3.2-1)

The fitting curve is already shown in Fig.2.5-1. The parameters of Ramberg-Osgood's law
are as follows;
a =1.47, a o = 413 MPa, E o =2.19 x 10~3, n = 7.91

Fig.3.2-1 shows the comparison of the simplified J-integral under the bending load with Jintegral by the FE analysis. In these calculations the crack extension was ignored. At thebending stress of 392 MPa which is close to the maximum value in the experiment, thedifference between both values is 10% and the simplified J gives conservative result.

(2) Comparison of Failure Assessment Curves (FAC)
Fig. 3.2-2 shows the comparison of FAC of R6 method Option 3 using the simplified J withthat using J-integral by the FE analysis. Also from this figure the simplified J analysis method
gives the accurate and conservative FAC.
3.3

a

t

da

a at'

t

(1) Analysis condition
Material properties, the stress-strain curves and J-resistance curves shown in Table 2.5-2 andFrg 2 5 1 -Frg 2 5 9 has been used for the analysis. Because the accuracy was confirmed, thesimplified J was used to obtain FAC in the case of R6 Option 3. To estimate the maximum
load in the case of the pipes with the weld metal, the stress-strain curves of the base metal and J-

resistance curves of the weld metal were used for the conservative assumption.

(2) Results
The maximum test loads normalized by the predicted maximum loads obtained from threedifferent kinds of evaluation methods are shown in Fig. 3.3-1-Fig. 3.3-3.

The average values of the normalized maximum loads of seven tests are I .05 from the NSC,

1.15 from R6 Op.2 and 1.09 from R6 Op.3. Most of all the predicted maximum loads agreed
well with the experimental maximum loads within 20% difference. There are some cases thatthe normalized maximum load is below 1.0, two cases from the NSC, one case from R6 Op.3

- 232 -

and no case from R6 Op.2.
Therefore, from the viewpoint of the accuracy, the NSC gives the most accurate prediction,

while from the viewpoint of the conservatism, the NSC may not always give conservative
prediction and R6 Op.2 gives the most conservative results.

4. CONCLUSION
In order to provide the experimental data for establishing the leak before break criteria of the
low alloy steel pipe, seven pipe fracture experiments were performed. Test pipes were 6-inch and8-inch diameter pipes of SFVQ IA or STPA24 Iow alloy steel. A circumferential through-wallcrack was introduced at the center of a pipe, and four-point bending load was applied without
internal pressure. Stable crack extensions were observed in all ofthe experiments.

To estimate the maximum applied load, the net-section criteria, R6 method Option 2 andOption 3 were used. The predicted values by three kinds of evaluation methods were comparedwith the experimental loads. Most of all the predicted maximum loads agreed well with theexperimental maximum loads within 200/0 difference. In detail, the NSC gave the most accurate
prediction but also gave two unconservative results. The predicted maximum loads by R6 Op.2
were conservative in all of the experiment cases. From the view point of conservativeness R6
method can be used for evaluation ofthe low alloy steel pipe fracture.
The further analyses on the pipe with the actual geometry is continuously followed to establish
the leak before break criteria of the low alloy steel pipe.

5. ACKNOWLEDGMENT
This test was funded by five Japanese PWR utilities, and the results on this study werediscussed and reviewed in the committee under Thermal and Nuclear Power Engineering Society.
The author acknowledges to the chairman Prof. Y. Asada, the committee members Prof. K. Miya,

Dr. K. Shibata, Dr. K. Kashima and the other committee members including those from theutilities.

6. REFERENCES/1/ Y. Asada, 'Development of Criteria for Protection Against Pipe Breaks in LWR Plants', pp.95l I I , International Journal of Pressure Vessel & Piping, vol. 43, 1 990.

/2/ A. Zahoor, Ductile Fracture Handbook Volumel, NP-6301 -D N 1 4-1 Research Project 1 757-69,

June 1989.
/3/ T. Tanaka, T. Nakamura, K. Koyama, I. Muroya, 'Four-Point Bending Tests of Low Alloy
Steel Pipes in Japan', presented at the 7th German Japanese Joint Seminar, Sept. 1 997.

- 233 -

~~~~T~~T~T~l~~~I r~~rlL-J

ot Psl(uT) PiPingNo

eak detectable(>5gpm)Inltlal defectEvaluate(shape, slze)

position

Load forevaluatlon

5 gpm crack
size

Yes
Crack size based on
propagation analysls

Crack propagation

Assumedanalysiscrack
Throuth wall
crack

Design load

Crack stability

8nalysls

InstableNof racture

Yes
Break

Leak
(LB8 appllcable)

Jrl l9~e a

L~tor~~~~~
desi9nFigure 1-1

LBB applicability evaluation flow

- 234 -

Test matrix ofthe

Table 2. I -1

Experimental

Material

numberNominalCrackdiameter(inch)

half angle(deg)

6

45
45

TWC
TWC

30

T\VC

TWC
TWC
TWC
TWC

SFVQ I A

SFB lpipe tests

SFWl

SFVQIA WELD

6

SFW2

SFVQ I A WELD

6

SFB2

SFVQ I A

45

SFW3

SFVQ I A WELD

45

STB l

STPA24

6

45

STW l

STPA24 WELD

6

45

Crack
geometry

TWC: Through wall crack

Table 2.5-1

Chemical compositions of the test materials(weight '/.)

NominalMateriai

diameter(inch)

C

Si

Mn

P

S

Ni

Cr

Mo

SFVQ I A

6

O. 1 8

O. 1 8

l .44

0.003

0.00 l

0.75

0.12

0.5

SFVQ I A

8

0.2

O. 1 9

l .47

0.003

0.001

0.76

0.12

0.5 l

STPA24

6

O. l

0.26

0.4 l

0.007

0.005

2.15

0.94

Table 2.5-2Tensile properties and Jlc Values ofthe test materials
0.2010 offset

Material(Base/~Veid)

DiameterTemperature(inch)('C)

yield
strength(MPa)
SPVQ I A(Base metal)

SFVQ I A(Weld metal)

STPA24(Base metal)

STPA 24(Weld metal)

strength(M:Pa)(MPa)

Jtc(k J/m2)

Flow stress

325

413

577

495

696

6

3~~5

477

628

553

l 23

3,-5

410

579

495

652

8

3~~5

457

630

544

270

6

310

245

482

3 64

623

6

310

513

598

556

l 52

SFVQ I A
SFVQ I A

Ultimate

6(Base metal)(Weld metal)V- 235 -

Experimental results

Tablc 2.6- l[3\pcrilnclltI) i pc

tlulnbcr

lllllterial

OutsideWa I l

C t-ack'l'otal lo',Id at the

Maximurn total

pipe

thicknessha I f' angle

crack initi'ation

experiment loacl

dlnnletcr

Crack extensionai (mm)(*)

Ct'ack extcnsiot]

a2 (n]m) (**]

(d cg)

(kN)

(kN)

l 5.7

45

i 96

282

I 4.8

9.0

I 5.7

45

216

254

6.4

3.l

l 5.7

30

226

339

I 3.3

().O

l 9.5

45

490

602

l 7.7

12.5

l 9.5

45

353

588

l 5.6

l 1.0

14.l

45

I 63

l 76

12. I

7.5

14.l

45

l 23

J 99

l 8. l13.0

mln(inch)

SI*VQ I A

SI'O i(6)

SI*VQ I A

l 66.5

lvtAG

(6)

SFW2

SFVQ I AMAG

l 66.5SFB2Sl'VQ I AlSI'W l~~cb

l 66.5I

sr'\v3
S'I'l3 l(6)
2 1 7.6(8)

SFVQ I A

2 1 7.6

lvi A G

(8)

S'I' r) A 2 4

l 65.4(6)
S'I'W l(*)(* *)S'I'PA24

MAG

l 65.4(6)

il lAbsolLttc crack extcnsion length

El 2

Projcctcd crack cxtcnsion lcngtll to tl]c pr-crirckcd cross section

unit: mm

SFVQ I A( )for 8-inch diameter pipe

Cracked cross section

,,\

\J

STPA24Cracked cross section

Fig.2.2- 1

Geometries of the test pipes

l

l

v

B

l

v

~i
l

0.2-0.3

A
JL

~5

5

3

unit:

Fatig uecrack

EDM:Electnc-Dlscharge-Machirunge : Halfcrack angle
Fi~.2.2-2

EDM

Throngh-\\'all crack geometry- 237 -mm

lrtltial crack!

ll!_ _1 I
l

l

L~

Ram

1

t

l

I

Ram

lt

A

't~

l

~~~__

~
t

~)
I 'I

I

'ft 'lll
il 111'l' 111'''

,

f

I

I

ll

l

Duct for heating test pipes

Fig.2.3- I Test system- 238 -

III

lI:

h~

600

tI

oo

400
~:

a o 413, E Q=2. 19E-3, a=1.47,n=7.

:,,

91

U,

O
~
;1,

200

o

Measured data
Firting cu~'$ by Rarnberg-Osgood's law

0.01 0.02osrrain a

Fig.2.5- I True stress-strain curve of the test material(6-inch diameter SFVQ IA pipe(Base metal))

600

I ooo o
o

400
o

~

a 0=410. E 0=2.lSE 3 a 1 26 n 8 98

:1,
:1,

:-

o

~
:/:

200
O Measured dalaoFirtin~_ curve by Ramberg-Osgood's law

0.01 0.02

ostrain e

Flg.2.5-2 True stress-strain curve ofthe test material(8-inch diameter SFVQ I A pipe(Base metal))

600
a Q=24S, E 0=1. 30E-3, a =j. 64,n=5. 03

"+00

4
:1,

t,
V:.

(90

200

o

o

Mc,asurcd duta
Finins cur\c by R~mbcrs'Osgood's I;!Lv

0.0 1 0.02

O

strain e

Flg.2.5-3 True stress-strain cur\'e of the test material(6-inch diameter STPA24 pipe(Base metal))- 239 -

2000

i

o

Mcasured data'~)o,~~, Ia"v' fir,ing cul~lc .'J=646 (A a)Q'54

'~

,N

oo

'\~ 1000-,,,oo

ooo oo

~J

-~

o

o

o
l

3

,)~

l

o

A a (mm)Fig.2.5-4 J-resistance curve ofthe test material(6-inch diameter SFVQ IA pipe(Base metal))

•_ooo
O

Measured dataF•owcr law fining cl~~rc
J= /~24 (A a)o.51

o oo o'~t]~ 1000

'~,

'l

~
\J
-,

o
080
~-

1

o

3

:~ a (mm)Frg ~-5-5 J-resistance curve ofthe test material(8-inch diameter SFVQ I A pipe(Base metal))

2000

O

M:casurcd data

powcr isw fining curvc

l=56? (A a)a.4a~---x 1000
~
~
cPo

o

o
l

o

,-

1

~~a (mm)Fig.2.5-6 J-resistance cur~'e of the test material(6-inch diameter STPA24 pipe(Base meta]))- 240 -

3

2000

O

Measured data
powet law fit:ing cu~!* e

J=250 (A a)a.19

l~N~: 1000
.~
,-,
J'a

..,

-

o o oo o o o

o

oo o

I

3

2

l

o

~ a (mm)Fig.2.5-7 J-resistance curve ofthe test material(6-inch diameter SFVQIA pipe(Weld metal))

2000

O

Measured data
Po\iver I~w fitting F~r\'e-

J=3SS (~ a)Q'38

'-~

,N

~~ 1000
4
-;*~
,~

oo

3

,~

1

o

o

Aa (mm)Fig.2.5-8 J-resistance cun'e ofthe test material(8-inch diameter SFVQ I A pipe(\Veid metal))

2000

O

\L (c;IsuTcd data

Pol*'cr law nuing curve

i=2?9 (~ a)Q'Ior''\ 1000 [

-,J
JJ:

[

'~

[

[
I

o

1,

l

~a (mm)Flg ~.5-9 J-reslstance cur\'e of the test material(6-inch diameter STPA24 pipe(\Veld metai))- 241 -

3

800

I

,l

,~ 600

~

~::

SFB2

\

~l
IS
c~;

O'

~ 400
O
~n_
~
~ ~_oo
~:;

c:;

hlO

o

SFW3
sFw2
/ SFB1~~\\ SFW1

50
Load-line displacement (mm)Fig.2.6-1

Load \'ersus load-line displac~ment curve of SFVQ IA test pipes

300

'f~

STW l

~:
~:
\J'

~_oo

i:D

/

c::;

O

~
~O
~~ 1 OO
rl
fl

~O

rrl

STB 1

O

l OO

50

Load-line displacement (mm)Fi~.2.6-~

Load \'ersus load-line displacement cur\"e of STPA24 test pipes- 242 -

150

'•1 Ilg I C

Cr'ack ll'.llf

O = 115*(1\'~e~ ~\s~\~L:c;esc\:e~\~~~y\\~~ I~~

I

t~
~~
c~
I

Crack
\~~~I~~

Outcr diamctcr 16C.5mm'l' I I i ck n cs s

l 1"
.l ,t~~' I --' l

1 ro .7 mm

lllC'Sll Ol' tl LCSL
l,'il~llLc• C'ICll]CI~lLp i pCMaximum apphed load 28•_kN at load-line dis
300
Model 2f\
Z;

Model 1

~:

'_/

1'c:;

O 2_OO~~~n~~:j 100
O
H

O

vO

O
Experiment dataO

Analytical data50oLoad-line displacement (mm)Fig.3. I -2 Comparison ofthe calculated total applied load versus

load-line displacement curve with the experiment curve(6-inch diameter SFVQI A pipe, Experiment SFB I )

300
r~
Z:

Model 2

~:

~l
1:)

Model 1c:;

O ~_oO ~~1

~
iO
~C~
h~
~ 100
O
H
(L)

O
Experimental data

c:;

O

c::;

~

Analytical data

6

8

10

Crack mouth openlng displacement( mm)
Fig.3. I -3 Comparison of the calculated total applied load versus
crack mouth openlng displacement cur\'e \\"ith the experiment cur\'e(6-inch diameter SFVQI A pipe, E\periment SFB I )- 244 -

12

o

1 03

J calculated by FEIVI
Simplified J

'~,
c~'

~

'\~~

101

,~'

~)

10-1

o

300

200

l OO

Bending stress400

ab (MPa)

Comparison ofthe simplified J-integral with J-integral by FEM analysis

Fi~.3 .2- 1(6-inch diameter SFVQ I A pipe, Experiment SFB 1)

1.5

OJ calculated by FE~'l
Simplified J

l

~~{

O
0.5

o

O
O

1

0.5

1.5

Sr
Fl~.3.2-2

Comparison of FAC of P\6 OP.3 using the simplifled J-jntegral\vith that usin_".. j-jnte_~._ral by FEM: analysis(6-inch diameter SFVQ I A plpe \vith crack half angle 45 ' )- 245 -

~~

l

t\D

~~
(1)
l

l:jg.~)~',j- I Nol'Inilll7.cd c'xpcrlnlcntili inalximum lotlcl by tl]c prcdictcd nlalxlmunl loLld ll'oln NSC:

l

~e
d~

~
l

l:jg•3•],-2 Nol 111,llIZCCI C\pCI I111Cll[ Il i]1,1\lllILllll lO•1(1 by tllC pICCilctcd lll•lXlllILlill l0•lCI flOlll

I

t\)

~

co
I

I:jg•3•3 3 Nol'llltllil_CCI C'XpCl'illlC'I1[ill l]ItIXllllLllll lOil(1 by tl]C pl'C(lictccl lllalXll]Itllll lOilCl il•Olll R() Op

Session V
Material Characteristics

V- lEffect of Reverse Cyclic Loading on Fracture Resistance

Curves in CT Specimens
C. S. Seok, Sung Kyun Kwan University, Dept. of MechanicalEngineering, Korea

Y. J. Kim, Sung Kyun Kwan University, Dept. of Mechanical

Design, Korea
J. I. Won, Sung Kyun Kwan University, Graduate School Dept. of

Mechanical Design, Koera

Effect of Reverse Cyclic Loading on the Fracture Resistance

Curve in C(T) Specimen
C.S.Seoka, Y.J.Kima, J.1.Weona'School of Mechanical Engineering, Sung Kyun Kwan Univ., Suwon, Korea

Fracture resistance (J-R) curves, which are used for elastic-plastic fracture mechanics
analyses, are known to be dependent on the cyclic loading history. The objective of this paper
is to investigate the effect of reverse cyclic loading on the J-R curves in C(T) specimens. Twoparameters were observed to be effective on the J-R curves during the reverse cyclic loading.

One was the minimum-to-maximum load ratio (R) and the other was the incremental plastic
displacement ((5;.y.1/~:i), which is related to the amount of crack growth that occurs in a cycle.

Fracture resistance test on C(T) specimens with varying the load ratio and the incremental
plastic displacement were performed, and the test results showed that the J-R curves weredecreased with decreasing the load ratio and decreasing the incremental plastic displacement.

Direct Current Potential Drop (DCPD) method was used for the detection of crack
initiation and crack growih in typical laboratory J-R tests. The values of J, and (5;, were also

obtained by using the DCPD method.

1. INTRODUCTION
This paper presents experimental results from cyclic displacement-controlled C(T) fracture
tests. Seismic loading consists of dynamic loading and cyclic loading. According to a study
by the International Piping Integrity Research Group (IPIRG) program[1], which pioneered
the work on cyclic loading with C(T) and pipe specimens, the reverse cyclic loading was

infiuenced by the parameters of load velocity, Ioad ratio, and incremental plastic
displacement. The reverse cyclic loading could contribute to the reduction of material's
fracture toughness, i.e., J-R curve. This work is applicable to flawed-structure evaluations

under seismic loading conditions, i.e. Leak-Before-Break (LBB) and in-service flaw
evaluation criteria where seismic loading is addressed. In LBB and flaw evaluation criteriasuch as the ASME Section XI, the cyclic nature of the seismic loading is not implicitly
accounted for on the potential effects on fracture toughness.
The objective of this paper is to study the effects of reverse cyclic loading on the nuclear

piping integrity. In order to assess the effect of cyclic loading on fracture behavior, theexperiments considering two different parameters were performed. One parameter is load
ratio (R) and the other is. incremental plastic displacement ((5i,y.11~;i), which can be just as
important as the load ratio parameter on the load-carrying capacity.

2. EXPERIMENTS
The test was performed at quasi-static loading rates. The detailed test procedure will befollowed.
- 249 -

2. I Material and specimen.
Material used in the test was SA 5 1 6 Gr. 70 carbon steel, which is used for nuclear piping.
The chemical composition of the steel in weight percent is given by Table I .
The specimens for tensile and fractwe tests were cut from a rectangular plats. As specified

in the ASTM E8[2] and E21L3], the tensile specimens were designed with 50.8-mm gauge
length and 1 2.7-mm diameter. The specimens for the fracture toughness test were set to I -in.-thick compact tension specimens in accordance with the ASTM E813. The crack plane of
the compact tension specimens was set to the T-L orientation.

Table I Chemical composition oftest material (Wt. o/o)

2.2 Tensile and fracture test

A servo-hydraulic computer controlled material testing system was used. The test
temperature was set to 288 ~ .

Table 2 summarizes tensile properties for SA 5 1 6 Gr. 70 carbon steel. Fig. I shows theschematic diagram of the fracture toughness test system. In order to measure load-line

displacement, a high temperature capacitance type COD gauge was used. The DCPD
method was used for measuring crack length and crack propagation. This system consistsof a DC-power supply (30Amp.), pre-amplifier ( x l04), and filter (lHZ Iow pass filter).

The fracture toughness test procedure on cyclic loading is not clearly defined yet.
Therefore, the fracture toughness test on cyclic loading was performed on the basis of thesingle specimen method specified in the ASTM E8 1 3 [4] and E I 1 52[5] for static fracture
testing. The fatigue pre-cracking was preceded for a compact tension specimen in the samemauner. According to the ASTM E813, the maximum load for pre-cracking was limited toprevent the infiuence of plastic deformation in case of the static fracture test. Under thisload limit, the crack was extended to a/w=0.55. To produce the plane-strain condition, thepre-cracked C(T) specimens were I Oo/o side grooved on each side with an edge angle of 45 '.
The fracture resistance tests were performed by changing the load ratio of R from I to - I .

Fig. 2 shows the variation of load vs. Ioad-line displacement for R=-1. To observe thevariation of J-R eurves with varying the incremental plastic displacement, (5;.y.//6i, the
DCPD method was used. Fig. 3 shows the load vs. Ioad-Iine displacement for 6.y.//6,=1 and
R=0. Firstl)r; the crack initiation displacement ((5:i) was obtained from the load-line
displacement vs. potential drop (voltage) curve as shown in Fig. 4. This crack initiation
displacement was used for a criterion in determining the number of reverse cyclic loading
for the fracture resistance tests.

- 250 -

Fig. lSchematic diagram of the J-R test systemqo

R = -l5cvcle /~ = O' l

3o

2o

10

Z
:X

-O
1:(!'

S -10
•20
•eo

•40
~: .5

o.o 0.6 1 o 1 5 2 o 2 6
Dis placement.,

Fig.

2

3.0

3 .5 4.0 4 5 5 .o(m m )

Load vs. Ioad-line displacement curve for cyclic load test

(R=-1 (5; /~:..=01)' .yct , '2.3 Crack initiation[6 -8]

DCPD method is one of the most accurate and sensitive techniques for detecting crack
initiation. The Crack initiation during the tests was determined by plotting a linear line, so-called an E.P. blunting line, on the load-line displacement vs. potential drop curve, as
illustrated in Fig. 4. In the beginning of the test, the crack tip area was controlled under- 251 -

Monoteniceta

intt iaaon

1'11'

s
8c~8i = 0.1

R=0

8=~ ~

8i

Displacemen~ 8

Fig. 3 11lustration ofcyclic load test (R=0 (5; 1~:, = O I )' .yct i '

20
le
lo

14

E
E
4J

12

E.P.
10

co
E
aJ
u

e

a!
(1)
b~

8

blunting line
/

(l9

4
2
o

~~

.2

1,5

20

2.5 3.0

3 .5

4.0

Voitage, (V)Fig. 4 Load-line displacement vs.

EP data for SA 5 1 6 Gr.

70 steellinear-elastic mode, and thus no change in the potential drop was observed. As the crack tip
blunted with increasing load, a linear relation between the load-Iine displacement and thepotential drop was observed up to the point of crack initiation. Therefore, the E.P. blunting
line can be used to determine the crack initiation displacement ((5:i). At the same time thevalue of J-integral corresponding to (5:, was proposed as J,. In the present paper, the averagevalue of crack initiation displacement determined with the monotonic loading was 0.92mm.
- 252 -

3. RESULTS
A summary ofthe fracture toughness test results performed at 288 ~ was listed in Table
3 . The J-R curves were fitted to a power-law equation, which is given by
JR = C] (A a)c2
The C1 and C2 Values were obtained from the test and are also listed in Table 3 .

In order to evaluate the effect of cyclic loading, the experiments were also conducted at

quasi-static loading rates with a load ratio (R) ranging from - I to I . The effect ofincremental plastic displacement was investigated by changing 6,y,/1~;, from I /40 to I /10
with a fixed load ratio of - I considering the reverse cyclic loading condition.
Table 3 Summary of fracture toughness test results at 288 ~)f*•~ ~
Mates.~

s.c~~ '~~~~';~ ~; ' ~~ ' ';~l'~f~~~ ~;~~;~;~~;;-'~~"~1/1"~~'~'i~~~/'~~~;~"~~i~://~' :;~~;:(~:~~~;":~~;/;~; ~ '///:t:;:;i:i;~~...;;i=~;/~;~;/:~_=':~~':~~~!'~;t~ ~~'~':~~~~~"~~~~'~~~/~i~ ~ .. ,s;;'~~~~:""#~i!~+~:+'~;~':~'~~*;." ~ +t::ip;:::.~4~ip '' ~1;c~~~' ~.*~~'f'~~'/~t~i.:"i'""~:

Ratc.>' (s::•) ,~~'~i7~~;~(;; ;;,,/i~:""~ 'i':;~~~~t~"';.....~.j'~ ;:/,j.t;;~~ ~'"'?'~~~~ ~*/"'"+~~= ~+ ~~
;~~

**=*

Quasi -

Gr. 70Static

~l

1 .O

313

345

o.35

o.5

292

363

o.35

o.o

262

323

o.33

249

311

0.34

-O.6

l 82

239

O.3 l

-0.8

153

215

o.35

-1 .O

134

180

o.28

1 15

134

191

o.35

l /20

98

1 42

o.33

l/30

85

121

o.30

l /40

77

l 07

o.27

-o.3

SA 516

_~ ~

~~r

1/10

=1 .O

3. I Eff_ect of load ratio on C(T) specimens

The experimental results were presented in Fig. 5 in terms ofJ-integral with increasing
crack extension. In these tests, only the load ratio was varied for 7 different values
covering monotonic loading (R=1), R=0.5, R=0, R=-O.3, R=-O.6, R=-O.8, and R=1 with afixed incremental plastic displacement of (5i,y,/16:i = 0.1. The resulting J-R curves showthat the J values are significantly reduced with decreasing the load ratio in the range of

R=0.5 was not evident. This result proves' the validity of ASTM E1152 requirement~ 253 ~

which restricts the amount unloading up to 500/0 of the current load. Fig. 6 shows the J,
values observed from the tests with 7 different load ratios. The corresponding Ji Value,
the J-integral at crack initiation, was nprmalized by the Ji,value under monotonic lpading
condition. The normalized Ji Value was decreased with decreasing the load ratio. The J,
value was about 60 percent of one for monotonic loading condition when the load ratio
reached -0.6. When the load ratio reached - I , the Ji value was about 40 percent of one formonotonic loading condition.

10Cl]

~K

R= 1

1)

c R=0.5A

A R= -0.3
o R= -O.6
v R= -0.8

R=0.0

900

e

v

8CID

~~(

o R= -1:~~$
AAAAA AJLA~~AA AA A A
c~ AA +~ C'(F. A4+ ~+~ A A
(F.-

7Cx)

f~

IE:

le~

5Cx]

t~

~
~
'~ 500c (bLCiA$~AA+AA AA ')q) QPe~eoeoee()Oc~(~~~_t i +2~:A~a q,o ~

15L*

O'
O
~,

~~A/A ~,) IP

400

V Vy vvVv vvv:Vvvvy V~((~^~~: o. oe=
J:b-i~:q,~r:~~ 8dirlbVlv ~1: IQ;Vv; ID ID IE! IDI Dl

300~riV :

200
ICx3

O
O

567

2

Crack Extension, (mm)Fig. 5 Cyclic J-R curves for SA 5 16 Gr. 70 steel at 288 ~) (6.y.11~;, =0.1)

1 ,o

o.o

e

o .8

,

i)

~5 07
c:

o
c
oo
O
E '
:~
cO o.6
'~5

e
e

Z

t

e
e
e
e

0.4

0.3

o .2

o .2

0.0 JD:2.o.4 4.8 -0.8 . I .o - I .2
Load Ratio

Fig. 6 Effect of load ratio for J at crack initiation (6.yctl~:, =0. I )- 254 ~

3 .2 Effect of incremental plastic displacement on C(T) specimens

The resulting J-integral vs. crack extension behavior was presented in Fig. 7. The
experiments were also conducted at quasi-static loading rates, and the incremental plastic
displacement, (5:.y.//(5:i, was set to 4 different values of l/lO, l/20, l/30, and l/40 with afixed load ratio of - I . The resulting J-R curves showed a decreasing tendency of J-

the J-R curves were considerably decreased when the 6.y.1/6/ Value was changed froml/lO to 1/20. Fig. 8 shows the variation of the J, value with decreasing the incremental
plastic displacement. In this Fig., the corresponding J, value was normalized by the J,
value for (5;.y.1/~;,= I /lO. The normalized J, value was apparently decreased with a decreasein the incremental plastic displacement as shown in Fig. 8.
3 .3 Discussion

It was observed from the tests that the J-R curves and the Ji Values were decreased
with decreasing the load ratio and the incremental plastic displacement as shown in
Figures 5 to 8. In general, the value ofJ-integral is measured from the load vs. Ioad-line

displacement curve. For the monotonic loading, the area defined as A(BCD) m Fig. 9 isused for the J-integral calculation. However, the phenomenon of crack opening andcrack closure observed from the reverse cyclic loading test is quite different from themonotonic loading test as illustrated in Fig. 9. While the crack opening is occurred atpoint B for the monotonic loading, it occurred at point A for the reverse cyclic loading
condition. Therefore, the value ofJ-integral which is the actual energy consumed with

400

350

ee

e

: vV 7VVVV'V

300

eV ,V'.V
V:1~r~~E 250
.~

1
~
_s
~200
Ca

Lo,

~y

l I llolE]Ialal ala IDI o D D a

. cc
ccc
e~ Dolol i( 1~(CF~~cA AAAJdL_JF AAAJLAA

VV aol I~IL~A~:A:AALA_0= 150 ~;Ii!~~l~Z~f.~ ! A AA~~AA;AA':L AA AJ2~AA A A ~~ A
' ~I~:::~AAA: e 6cycJq= I / 5: v v 60yde/q=1/10

!:,

100

I D ~ ce,lq=1120
11' c 6cyde/q=1/30

50o

5 G(mm)7 8 9 Io
1 2 3 Crack
4 Exten8ion,

Fig. 7 Cyclic J-R curves for SA 5 16Gr. 70 steel at 288~C (R=-1)- 255 -

1 .1

e

lp

1 .O

1 o .e
CD

co

(,

(1)

=a) o.e

E
co

~5

Z

0.7

e
e

1)
o .e

e

Smaller Plastc
displacement incrementso .5

1'10 1120 Il30 1/40
seyd.
/ s. .

Fig. 8 Effect of incremental plastic displacement for J at crack initiation (R=- I )crack growih must be measured from the area specified as A(ABCDE) covering the point P
rather than A(BCD). In other words, calculating the J-integral for the reverse cyclic loading
based on the area A(BCD) will underestimate the J-integral which should consider the area
A(ABDE) to account for the effect of reverse cyclic loading. It will be one of reasons ofdecreasing the J-R curves under the reverse cyclic loading. Therefore, the way of J-integral
calculation for the reverse cyclic loading should be studied further regarding both tensile andcompressive loading condition.

Cl:,C:f~///r/f/ ' ll.

/

l"
/

/ //'

I

.* ,

ll

I

..ll

ll

//'

B

*./ t

' '/~ "' ll

/

J

$

)
.'* I

/
.l

'! "~*
;'~~
"

~;' '

./i

/ ' /*.
~~

F

D i sp I acemen t' '\\>~7'\' "A_~~/ 't/

Po penE

Fig. 9 Definition of area for the J calculation under reverse cyclic loading- 256 -"p'"

4. CONCLUSIONSThis paper presented the fracture resistance test results regarding the effect of reverse
cyclic loading in C(T) specimens. The conclusions are summarized as follows:(1) For the SA 5 1 6 Gr. 70 carbon steel, the J-R curves and the Ji Value were decreased withdecreasing the load ratio and the incremental plastic displacement, respectively.
(2) When the load ratio was set to - I , the resulting J-R curves and the J, value were about

40 - 50 percent of those for the monotonic loading condition.
(3) In calculating J-integral under the reverse cyclic loading, both tensile and compressive

loading should be considered. A further study on the test procedure regarding thereverse cyclic loading is recommended.

5. REFERENCE[ I J P. Scott, G. Kramer, P. Vieth, R. Francini, and Gery M. Wilkowski, The Effect ofDynamic and Cyclic Loading During Ductile Tearing on Circumferentially Cracked Pipe:

Experimental Results, ASME PVP Vol. 280 (1994) 207 - 220.
[2] Standard Test Methods of Tension Testing of Metallic Materials, ASTM E8.
[3] Standard Test Methods for Elevated Temperature Tension Tests of Metallic Materials,
ASTM E2 1 .
[4] Standard Test Method for JIC, A Measure of Fracture Toughness, ASTM E8 1 3 .
[5] Standard Test Method for Determining J-R Curves, ASTM El 1 52.

[6] G. M. Wilkoski, J. O. Wambaugh, and K. Prabhat, Single Specimen J Resistance Curve

Evaluation Using the Direct-Current Potential Drop Method and Computerized Data
Acquisition System, Fracture Mechanics : 1 5th Symposium, ASTM STP 833 (1984) 553 576.
[7] M. G. Vassilaros, and E. M. Hackett, J-Integral R-Curve Testing of High Strength Steels
Utilizing the Direct-Current Potential Drop Method", Fracture Mechanics : 1 5th Symposium,

ASTM STP 833 (1984) 535 - 552.
[8] Ad. Bakker, A DC Potential Drop Procedure for Crack Initiation and R-Curve
Measurements During Ductile Fracture Tests, Elastic-Plastic Fracture Test Method, ASTM

STP 856 (1985) 394 -410.

- 257 -

V-2

The Frequency Effect on the Fatigue Crack Growth Rate of 304
Stainless SteelJien-Jong Chen & Yan-Shin Shih, Institute of Nuclear

Energy Research Atomic Energy Council, Taiwan, Republic of China

The frequency effect on the fatigue crack growth rateof 304 stainless steel

Jien-Jong Chen"'b and Yan-Shin Shihb"Nuclear Division of Institute of Nuclear Energy Research, Box 3-3, Lung-Tan,

Taoyuan Taiwan, Republic of Chinab Mechanical Engineering , Chung-Yuan Christian University ,Chung-Li

32023,Taiwan,Republic of China

ABSTRACT
Under cyclic loading condition, the fatigue crack growth (FCG) rate governed bystress intensity factor and stress ratio is well known. Walker's equation, Forman's
equation and Elber's equation are the typical formulae of the FCG rate. However, theloading frequency effect on the FCG rate is still need to explore. The behavior ofloading frequency effect has been studied on some of the visco-elastic material recently,

and gives a more clear aspect of the frequency effect on the FCG rate. In physicalsense, The knowledge is still lacking on for the loading frequency effect on the FCG
rate for the 304 stainless steel. James had provided a lot of experiments, through thedata analysis, he suggested an evaluation equation which is merely based upon theexperimental illustration. In this study, the alternation of physical properties of thematerial under various loading frequency is used to illustrate the modification of FCG

rate, and a new formula based upon modified Forman's equation has been provided.

1. Introduction

The modified Forman's equation(Forman et. al. ., 1992) which accounts the stress- 259 -

intensity factor range, the mean stress level and other important parameters is one of thebest formula for all the fatigue crack growth analysis. This equation is shown as
following:da C~" [AK - ~~,h JP

dN ~ [(1 -R)Kc ~~C]q (1)where C, n, p, q are matenal constants R rs stress ratro ~~ rs the cntical stressintensity factor, and AKth is threshold of stress intensity factor. Therefore, the modern

codes of FCG analysis, such as NASA/FLAGRO(EMT,1995), NASCRACK(F~~,1994), etc., have already incorporated this model to evaluate FCG rate of structuralcomponents. However, the FCG is a very complicated subject which can be affectedby material properties, Ioading history, global structure with local crack geometry andenvironmental condition, etc. Hence, in the fatrgue llfe analysrs of machme
components, there are still somewhat difference between predictions and experimental

data. For experimental result, a great mount of these deviations may be caused bydifferent testing conditions. Hence. considering as many as the important parameters
into evaluation formula is sti]] as a major topic for fatigue life assessment. Many ofthe modern structures may be operated for a wide range of loading frequency, andfatigue failure occurs during their service life. Therefore, the effect of cyclic loading

frequency on the FCG rate is as an important research subject.

In the review by James(James, 1985), beside of the stress intensity factor, the FCG

rates are precisely affected by some other parameters, and the cyclic loading frequency

is a significant one of all the important parameters. However, due to the complicatedinteraction between material and loading, the behavior of the loading frequency effects
on the FCG rate is still not clear.

- 260 -

In the analysis of FCG rate with loading frequency effect, there are two approaches

have been published in the literature. One constructs the model of FCG rate with thefrequency effect by analyzing experimental data, the other explains the underlying

reason of cyclic loading frequency. From experimental data, the increasing theloading frequency generally decreases the FCG rate for various materials. In early

70's, a power law of loading frequency was added to the model by Muknerjee(Mukherjee, 1971)

dadN = Cof (~~) (2)wheref is loading frequency, Co ' ;L and n are constants for a given combination of thematerial, temperature and environments. This model has also been adopted by
Solomon and Coffm ( Solomon et. al., 1973), James (James 1979), Plumtree andSchafer (Plumtree et. al., 1984), and Yokobori and Sato (Yokobori et. al., 1976).
Alternatively, James (James 1985) performs a curve fitting of experimental data, andsuggested a formula as follow,dadN = Co F(~C) (3)log(F) = Bo + BIX + B2X2 + B3X3 (4)where F is frequency correction factor, Bo ' Bl 'B2 and B* are fitting constants, and X islogarithm of the frequency in cycle/min. This approach gives a closely frtting withexperimental data. Recently, some experimental datum show that a portion of FCG

rate curve is independent of cyclic loading frequency. Hence, a new concept on

critical frequency was presented by Makhlouf and Jones (Makhlouf, 1993) They
accomplished a series of experiments on 1 8010-Cr. ferritic stainless steel with various- 261 -

frequency at elevated temperature, a step function of relationship was presented asfollows[da f
dadN = C(AK) for f > f (6)where f. (T) is a critical loading frequency on each temperature, such as f. (655 "C)is closed to 60 cpm and f.(500 "C) is near by 3 cpm.

For the underlying reason of loading frequency effect can be speculated by
Yokobori and Kamei (Yokobori et. al., 1975) who use dislocation dynamic theory toanalyze the behavior of strain rate sensitivity of material and derive the relationship asfollow,~+1f~+1
da = Co~'2(~~)~'2 (7)dN

where m is the exponent in the dislocation velocity equationv=vo(T~ / T;)",T. isapplied shear stress, T; is a constant which represents the stress required to give v = l

cm/sec. This equation reaches the same form as eq.(2), and can be regarded asphysical basis of the power law equation, but the exponents of frequency and stressintensity factor in Yokobori's formula have the same value with negative sign is hard to

confirm with experiments data.
The object of the present investigation is to study the effect of frequency on FCG

rate of 304 stainless steel. A new FCG rate model is developed based on Modified
Forman equation for better prescribing the loading frequency effect on FCG rate, it can
not only offer closer prediction on FCG rate but also with physical in sense.

- 262 -

2. Data acquisition and analysis
The experimental FCG rate analysis of 304 stainless steel has been carried out by

several authors (James et al 1971, Ohta et. al. 1978, Hudson et. al.
1978,1982,1989,1991) for investigating the effect ofthe various parameters. All ofthe
experimental data used in this analysis is obtained from the open literature. To avoid

the loading hold time effect, most of the adopted data was generated by sawioothloading waveform with a loading/unloading ratio of 7.78. The total FCG rate withdifferent loading frequency versus stress intensity factor range is shown in Figure I .

The datum appeared on various articles were shown with several typical forms andunits. These basic raw datum are treated by using Forman's equation, firstly. For
minimizing the effect of stress ratio, critical stress intensity factor and threshold ofstress intensity factor, the crack growth rate can be expressed into the similar form with

Paris equation and shown as following,da [(1 - R)Kc ~ ~~]* CAK (8)dN [AK - ~~,, J' ~3. Loading Frequency Analysis

From the constructed raw data, the coefficients C and n in eq. (8) for different

loading frequency can be calculated and shown in Table I . The power coefficient, n,

for different loading frequencies in Table I is very similar. Therefore, one may

assume the power coefficient be not affected by loading frequency. A mean pwer
coefficient has been determined using the following relation:n^,..~ - ~n ' M (9)~M= 3.13546

where n and M are the power coefficient and number of data with this power coefficient.

~* 263 -By the curve fitting of eq. (8) for all experimental datum, C and n are obtained, and theresult shown in Fig. 2, One can fmd that the n value is the same as eq(9). To reduce

the error of simulation, one considers that C depends on loading frequency, The new

frtting material constant C ' based on the power exponent of value 3.13546 can beobtained and shown in Table 2.
In the analysis of FCG rate with frequency effect, increasing the loading frequency

increases dislocation density and decreases FCG rate. From mention above, the FCG
rate can be expressed as a form of power exponent. Hence, a modified power law has
been proposed as,r"

da C~C"[AK- AK,h]P L f )dN = [(1- R)Kc ~ ~K]q f..f (lO)
where fref is a reference state loading frequency, and 2 cpm is used in this study.

4. Result and Discussion
The material constant ratio ( C' / C~f ) with respect to loadmg frequency ratro rs shownas Figure 3, where C:.f is the material constant of the reference state. One may finsthat increasing the loading frequency ratio decreases the material constant ratio, and thematerial constant approaches to a low bound value 0.5 while f >3000 cpm. Thisfeature reach the same conclusion with Makhlouf (Makhlouf, 1995). The curve frtting
equation.ofthe material constant ratio is shown as,

C'

f )~f
f .,f

Substituting into equation (lO), one obtains- 264 -

o.0802

)da C:,f~~"[AK- AK,h]P Lrf~o.oeo2

dN = [(1-R)Kc ~ AK]q f..f (12)The fatigue crack growth rates for different loading frequency by James, eq. (12) andexperimental data are shown in Figure 4-6. A closer fitting has achieved by the present

model.

5. Conclusion
The wide-range loading frequency effect on 304 stainless steel at room temperature

has been studied, and a new modification based on modified Forman's equation is also

provided . Some conclusion are made as follows.

l . A new equation is proposed to replace the modified Forman's equation for

counting the loading frequency effect on FCG rate. The deviation between
James and new equation are very small.

2. Below the critical loading frequency, the FCG rate increased slightly withdecreasing loading frequency at room temperature, this is caused by the higher
strength feature of 304 stainless steel at room temperature.

3. The lower bound ofFCG rate exists while f > f. .

Acknowledgment
This work was supported by National Science Council, Taiwan, Republic of China
under grant number NSC-86-22 1 2-E-033 -004.

Reference

Engineering Mechanics Technology, Inc., 1995. NASA/FLAGROFatigue Crack Growth Computer Program Version 2.0, published byEngineering Mechanics Technology, Inc.

Failure Analysis Associates, Inc., 1994. NASCRAC-NASA Crack- 2,65 -Analysis Code. Version 3.0, published by Engineering Mechanics
Technology. Inc.

Forman, R. G., Mettu.S. R., 1992. Behavior oi' Surface and corner
cracks subjected to tensile and bending loads in a Ti-6Al-4V alloy,

Fracture Mechanics 22'th Symposium(Volume l).ASTP STP I 131,
Ernst. H. A., Saxena, A.. McDowell, D.L., Eds.. American Society
for Testing and Materials, Philadalphia, pp. 5 19-646.

Hudson C. M., 1978. "Fracture Toughness and Fatigue Crack Growthfor Metallic Alloys-Part I" International Journal of Fracture, Vol.
14, pp. R15 l-R184.

Hudson C. M., 1982. "Fracture Toughness and Fatigue Crack Growthfor Metallic Alloys-Part ll" International Journal of Fracture, Vol.

20, pp. R59-RI 17.

Hudson C. M., 1989. "Fracture Toughness and Fatigue Crack Growthfor Metallic Alloys-Part lll" International Journal of Fracture, Vol.

39, pp. R43 R63

Hudson C. M., 1991. "Fracture Toughness and Fatigue Crack Growthfor Metallic Alloys-Part IV" International Journal of Fracture, Vol.

48, pp. R19-R43.
James L.A., 1979. Frequency effects in the elevated temperature crack

growth behavior of austenitic stainless steel -A design approach,

Journal of Pressure Vessel Technology-Transactions of the ASME,
Vol. 101, pp. 171-176.

James L A., Jones D. P., 1985. "Fatigue Crack Growth Correlations
for Austenitic Stainless Steels in Air", in Conference on "Predictive

Capabilities in Environmentally Assisted Cracking. (Presented at

the Winter Annual Meeting of the American Society of MechanicalEngineers.), ASME PVP Vol. 99, pp. 353-414.
James L.A., Schwenk E. B.,1971. Fatigue-crack propagation behavior- 266 -

of Type_ 304 stainless steel at- elevated temperature"~ MetallurgicalTransactions of ASME, Vol. 2, pp. 491 496

Makhlouf K. & Jones J.W., 1993. "Effect of Temperature andFrequency on Fatigue Crack Growth in 180/0 Cr. Ferritic Stain]ess

Steel" Internatlonal Journal of Fatlgue Vol 15 No 3 pp. 163171.

Mukherjee B., & Burns D. J., 1971. Fatigue-crack growth inpolymehylmethacrylate, Experimental Mechanics, Vol I l, pp. 433-

439
Ohta A., Kosuge M. and Sasaki E., 1978. "Fatigue Crack Closure over

The Range of Stress Ratios from -1 to 0.8 Down to The Stress
Intensity Threshold Level in HT80 steel and SUS 304 StainlessSteel" Internatlonal Journal of Fracture Vol. 14, pp. 251-264

Plumtree A. & Schafer S., Waveform and frequency effects on thehigh temperature fatigue crack propagation rate of stainless steel"in Advances in Fracture Research (Fracture 84), VOL. 3. Pub]ishedby Pergamon Press, Oxford, England New York, USA, pp. 22492256.

Solomon H.D. & Coffin L.F , 1973. Effect of frequency andenvironment on fatigue crack growth in A286 at I 100 F, in Fatigue

at Elevated Temperature, Edited by Carden A.E. et. al., ASTMSTP 520 , Philadelphia, pp. I 12-122.

Yokobori T. & Sato K., 1976. " The effect of frequency on fatigue
crack propagation rate and striation spacing in 2024-T3 Aluminumalloy and SM-50 stee]", Engineering Fracture Mechanics, Vol. 8, No.
l, pp. 81-88.

Yokobori T., Yokobori A. T., Jr. and Kamei A., 1975. "Dislocation
Dynamics Theory for Fatigue Crack Growth", International Journalof Fracture, Vol. I l, No. 5, pp. 781-788- 267 -

Table I :The material constant of FCG of various loading frequency

Table 2The material constant C '~ 268 -

1 E-3

~~~~

1 E-4

~; ~~~

'~
o

~>~~~
~~~

o
>
.\o

_= I E-5~~~~~; ~

Z
~
.\

CU

~o

l~~~~~ ~
1 E-6

~~ ~~

1 E-7

1 O0.00

1 0.00AK (ksi~ in),Figure I :The total FCG rate data for 304 stainless steelof various loading frequency and 75-- 269 -

1 .OE-3

Y = pow(X,3.13546) * I .74928E-OIO ~~~~~~~~~>

~
~

~~

~

~~

'\

:L,

~:~x_

~e

~;~~~~~ ~d) 1.0E-5
~e
OC

~
-Z~

~>~
~~~;

~-

~ ~~

~
~.
~ 1.0E 6

~> ~

CU

~>

1 .OE-7

10,00

1 oO. OO

AK (ksi~ in),Figure 2 :The mean slope of the total FCG rate data for
304 stainless steel.

- 270 -

1 E-3/~/~~

Experimental data/

James [1985]//

eq (12)

1 E-4.A~~~~
/

Z
._~

CU

/

1 E-5

/

1:

/

/

/

/

1 E-6

1 E-7

AK
Figure 4 : The fatigue crack growth rate for SS 304 on 2 cpm
1 E-3

1 E-4

of loading frequency and 75 •F~

Experimental data,

James [1 985]/

eq (1 2)

/

Z
'\~
U

/

~

~

/

/

/

/~~>
• '~)

1 E-5

cu

1 E-6

~

,~

1 E-7

AK
Figure 5 : The fatigue crack growl:h rate for SS 304 on 180 cpm

of loading frequency and 75•F

- 271 -

1 E-3d'~ Experimental data~

James [1 985]

1 E-4

• Eq(12)
Z
~D
•\
I E-5
~
CD

~~

~~ ~~l
1 E-6

~

~
c
1 E-7

AK
Figure 6 : The fatigue crack growth rate for SS 304 on 400 cpm

of loading frequency and 75•F

- 272 -

V-3

Investigation on Fracture Toughness Evaluation Method for
Reactor Pressure Vessel Surveillance

Kunio Onizawa, Tohru Tobitz & Masahide Suzuki,
Reactor Component Reliability Laboratory, Japan Atomic Energy
Research Institute, Japan

Investigation on Fracture Toughness Evaluation Method
for Reactor Pressure Vessel Surveillance

Kunio ONIZAWA', Tohru TOBITA and Masahide SUZUKI
Reactor Component Reliabilty Laboratory, Japan A tomic Energy Research Institute,
Tokai-mura, Naka-gun, Ibaraki-ken, 319- I 195, Japan

Abstract
To evaluate cleavage fracture toughness in the ductile-brittle transition range of RPV

steels, the applicability of precracked Charpy-v (PCCv) specimens has been investigated. An
approach based on the weakest link theory was applied to analyze the specimen size effect andthe scatter of fracture toughness values. Four kinds of ASTM A533B class I steels made by
Japanese steel makers were used.

The specimen size effect on cleavage fracture toughness was clearly seen among PCCv,l T-CT and 4T-CT specimens. To obtain the equivalent data from PCCv specimens to I T-CT
specimens, the size correction scheme based on the weakest link theory was applied to the PCCv
data. The scatter of fracture toughness values was considered to have a Weibull type distribution.

lrradiation effect on the scatter was shown to be insignificant. The fracture toughness transition

curves were determined by PCCv data according to the drafi ASTM and the JEAC. The shifi ofthe master curve by irradiation was somewhat greater than the Charpy 4 1 J shifi. The scheme forthe estimation of lower bound fracture toughness was also studied.

1. INTRODUCTION
The brittle fracture of an RPV must be prevented, because it may lead to a catastrophic

fracture of the vessel followed by severe accident. Therefore, the RPV is operated at higher
temperature than the ductile-to-brittle transition temperature of the RPV steel . However, fastneutrons from the reactor core during operation can cause the increase of the transition
temperature, i.e., the irradiation embrittlement. To assure the structural integrity of an RPV
throughout the operational life, the fracture toughness after neutron irradiation are needed. To dothis, the surveillance program for the RPV is performed according to a certain code andregulation. The results obtained from the Charpy impact tests according to the codes are used forthe estimation of fracture toughness after irradiation(1, 2). Most of the codes assume that thedegree of irradiation embrittlement obtained from Charpy impact tests are equivalent to that offracture toughness for the material concerned. Recent research, however, has indicated that thecorrelation is not always conservative in predicting the degree of irradiation embrittlement interms of fracture toughness(3'4). Fracture mechanics methodology is, therefore, necessary for the
precise evaluation of irradiation embrittlement to assure the structural integrity of an RPV.

A methodology to evaluate fracture toughness of RPV steels has not been established' Corresponding author. Tel: +81 29 282 5147; Fax: +81 29 282 5406;
e-mail : konizawa@popsvr.tokai j aeri .go jp- 273 -

completely, particularly in the transition range. As the fracture toughness values from test
specimens of the RPV steel show a large scatter in the ductile-to-brittle transition range, thefracture toughness should be evaluated with the statistical nature. This means that a certain
statistical approach is necessary for the evaluation of the fracture toughness. Fracture toughness

values from small size specimens, such as surveillance specimens, also show large effects on
specimen size. Although intensive research("g' 5) has been conducted to establish the statisticaltreatment and the evaluation method of size effects in the past, the methods need a large numberof specimens and/or relatively large size specimens(6). The number and the size of surveillance
specimens are limited by the small volume of a surveillance capsule. Therefore, further research
is desired to resolve the limitation and to establish an improved surveillance test method that is
applicable for the small nulnber and the small sizes of specimens.
This paper describes the results of fracture toughness tests using four kinds of pressure

vessel steels including an IAEA reference material JRQ. Based on the results of the fracture
toughness tests by precracked Charpy specimens and standard size I T-CT and 4T-CT specimens,a statistical analysis was performed to evaluate the scatter and the lower bound of fracture
toughness in the transition temperature range . A part of this report was performed at JAERI in

the framework of the IAEA Coordinated Research Program Phase-3(7).

2. EXPERIMENTS
2.1 Materials
Materials used in this study were four kinds of pressure vessel steels of ASTM A533B

class I . The one is designated JRQ, which was used as a correlation monitor material in the

IAEA CRP-3(8). The another material w~s chosen to have a large number of database forcleavage fracture toughness from small size specimens to large ones. The material is designatedas JSPS A533B-1 which was used in the round robin study organized by The Japan Society for
Promoting of Science (JSPS)(9). The other materials were made as A533B-1 steels which werecalled as Steel A and Steel B that corresponded to the old age RPV steel and the modem steel,respectively. Chemical compositions of the materials are summarized in Table I . Mechanicalproperties at room temperature for the materials are listed in Table 2.

2.2 Fracture Toughness Tests
Four types of specimens were used in this study, namely precracked Charpy (PCCv),
0.5T-DCT, I T-CT and 4T-CT. In some cases the specimens were side-grooved by I OO/o on eachside of the specimen after precracking. For JSPS A533B-1 steel, one of the authors performed

fracture toughness tests using PCCv specimens machined from halves of broken 2T-CT
specimens at SCK/CEN(ro). The notch orientation of the specimens is T-L direction for all
specimens. All specimens were machined from the approximately quarter thickness position of
the material . Fracture toughness tests were performed at mainly lower transition temperature

range which a cleavage fracture was occurred before the limit load of the specimen. Some
specimens were also tested at an upper shelf temperature range to obtain J-R curves.

Testing and evaluation for fracture toughness were performed according to ASTM
standards(ll,12) and JSME method(13). A valid plane-strain fracture toughness, Klc, could not be- 274 -

obtained from the specimens except for 4T-CT specimens. The following method, therefore, wasapplied to obtain elastic-plastic fracture toughness. Fracture toughness values at cleavage
fracture, Jc, were calculated based on the area under load-displacement curve up to the sudden
load drop and the cleavage fracture load (see eq. I ). Fracture toughness KJC Values were thenconverted from Jc by the following equation (2). Elastic modulus E and Poisson's ratio v were
fixed to 206 GPa and 0.3 in this study, respectively.

J - Jcl + Jpl

C~

K2
-C
+ nA
P!

(1)

E' BN bowhere Kc: Stress intensity factor at the cleavage fracture,n =2+0.522*(a~/W) for compact specimens,n =2 for PCCv specimens,
Apl: Plastic part of the area under load-displacement curve,

BN: Net thickness, bo: Initial ligament (=W-ao),
ao: Initial crack length, W: Specimen width,

E'=E for ASTM, E'=E/(1- v 2) for JSME.

In the ASTM draft method(14), n factor has been changed from 2 O to 1 9 based on the

3-D detailed analysis for PCCv specimen(15). This makes the fracture toughness value lower by

maximum 50/0 in terms of J-integral. When some equations in the ASTM draft method areapplied to the data, the new 7? factor is used.

At upper shelf temperature range, Jlc Values and J-R curves for the materials except for

the JSPS A533B-1 were obtained according to the JSME method. Ductile crack growih in a
specimen during a test was measured by unloading elastic compliance method.

2.3 Irradiation

PCCv and 0.5T-DCT specimens of JRQ were irradiated in Japan Materials Testing
Reactor (JMTR, thermal output 50MW). Fast neutron fluence for each specimen was calculated
to be within the range from I .9 to 2.5 x I 019 n/cm2 (E> I MeV). The displacement per atom (dpa)
for iron by the irradiation to 2 x I 019 n/cm2 (E> I MeV) was about 3 x I 0-2. Fast neutron flux was

about I .2 x I 013 n/cm2s (E> I MeV). Temperature for the specimens during irradiation was
controlled in the range from 280 to 300 ~C . The shifts of Charpy transition temperatures and thehardening by neutron irradiation are summarized in Table 3(7)Many specimens of Steel A and Steel B were also irradiated at JMTR. Sixteen PCCv
specimens of the steels were encapsuled with ten Charpy-V specimens for the JMTR irradiation.
Post irradiation tests for those specimens are now underway.

3. RESULTS AND DISCUSSION
3.1 Fracture Toughness Database
Figures I to 5 show cleavage

fracture toughness data
- 275 -

obtained in thisstudyl 1 6).

Thevalidity limit for the PCCv data based on equation (3) is indicated. This criterion has been
stipulated in the JSPS method.
bo ' B > JcaY

(3)25
The equation (3) used in the JSPS method is the same as the validity criterion on thedetermination of a Jlc in the ASTM E813. For the PCCv specimen, the ligament size, bo, issmaller than the specimen thickness, B. Therefore, bo is considered as a key parameter for thevalidity of PCCv data.
Figure I shows KJC values of unirradiated JRQ specimens tested in the transition range.

At room temperature, cleavage fracture was not observed from any PCCv specimens, but fromall 0.5TDCT and ITCT specimens. When PCCv and IT-CT data are compared at -80 ~C, it isseen that I T-CT data are situated within the scatter range of PCCv data. The Jlc Value at I OO ~C

obtained from IT-CT and 0.5T-DCT specimens by means of the JSME method is also indicatedin the figure. Some data points of PCCv and compact specimens tested at upper transition
temperature range exceeded the Jlc Value.

Figure 2 shows all cleavage fracture toughness data of irradiated JRQ specimens. Test

temperatures for five PCCv and six 0.5TDCT specimens, performed in the frame work of IAEA
CRP-3, were sparse from O to I OO ~C . We tested additional irradiated PCCv specimens; twelve
specimens around 1 7 ~C and ten at 35 ~C so that a statistical analysis could be done.

Figure 3 shows fracture toughness values from PCCv specimens of JSPS A533B-1
steel with a large number of database by means of compact specimens from 0.5T to 4T (13). The

database includes -25, O and 25 ~C data while PCCv were tested at O ~C. Obviously PCCv data
show higher fracture toughness values than the other compact specimen data.

The cleavage fracture toughness data of Steel A are shown in Figure 4. The data
include some valid Klc data from 4T-CT. The Klc data locates nearly the lowest bound of theother data. The PCCv data show higher KJC Values than KJC Of I T-CT and Klc of 4T-CT data.
The upper shelf fracture toughness, Jlc, obtained by I T-CT at 20 ~C is indicated in the figure.

More than halfofPCCv data at -50 ~C exceeded the Jlc Value.
Figure 5 indicates the results of fracture toughness tests by specimens of Steel B. ValidKlc data points from 4T-CT are also shown in this figure. The Klc data lie nearly lowest bound of

the other data. Similarly to Steel A, PCCv data indicate higher toughness values than the otherCT data. The upper shelf fracture toughness, Jlc, obtained from I T-CT at 20 ~C is indicated in
the figure. Several PCCv data at -80 ~C exceeded the Jlc Value.

3.2 Specimen Size Effect on Cleavage Fracture Toughness
As is known well, a fracture toughness value from a small specimen can be higher thanthat from a large specimen. In addition to this, in the transition range, a statistical effect oncleavage fracture is apparent. This has been explained by the weakest link theoryll7,18) that isbased on the local inhomogeneity in the material. The theory treats the probability to find aparticle that is located randomly in the material and that leads to the fracture of the specimen.
Another reason is that the small specimen has insufficient length in thickness to keep small scale

yielding and maintain a certain level of constraint at a crack tip, as compared with a large size
specimen or a real heavy-section structure(19).
- 276 -

Figure 6 shows the effect of specimen thickness on fracture toughness value. Eachvalue was normalized by Klc Values for each testing condition. This figure clearly shows the size

dependence on the mean value. The highest PCCv data was larger than four times of Klc Value.
The mean values of PCCv data except for invalid data were I .5 to 2.2 compared to Klc Value. On
the other hand, the lowest value of PCCv data was approximately the same as the Klc Value.

To obtain a reliable fracture toughness value from a small size specimen, several
correction schemes have been proposed. Wallin(20) has proposed the following equation forcorrecting specimen size applicable to pressure vessel steels;=K y K~,~ )(B. / By)1/~ (4)KJC~i*+ (KJC _'where KJc_. ., y: KJC for specimens of thickness, B. ., yK~j. = 20 MPal/~

m=4
Figure 6 includes the size effect correction curve based on the equation (4). The upper
curve is corresponding to Steel B, the lower curve is for Steel A. The correction curve for SteelB overestimated 300/0 of Klc Value by 4T-CT specimen. Unfortunately Klc Value of JRQ was notavailable. However, the size correction was compared between PCCv and I T-CT data. As aresult, the correction provided the I T-CT equivalent values from PCCv data. To verify thecorrection equation, we need more database including sufficient number of PCCv data.

3.3 Scatter of Fracture Toughness
The equation (4) is based on the weakest link theory. The theory assumes that the
distribution of the fracture toughness values of a material can be described as a Weibull typestatistics. Using three parameters, K~i., m and Ko, the fracture probability based on Weibull
statistics can be expressed by the following equation(20).

K KJc_rr ~20 "Ko ~ K~j. Ko ~ 20 (5)~f =1-exp Jc_lT ~K~i~ ~ =1-exp

where Pf: Median ranking( = (i-0.3)/(N + 0.4) )m, Ko: Fitting parameters

All data obtained by the same-type specimens at a certain temperature can be analyzed
by these equations (4) and (5). This equation (5) has also been incorporated into the ASTM draft
method to describe the statistics of cleavage fracture toughness.

Figures 7 to 9 compare Weibull plots of PCCv data and I T-CT data. PCCv data werecorrected by equation (4) to I T-CT thickness for comparison. These figures also include thefitting curve by equation (5). Figure 7 indicates that the corrected PCCv data and I T-CT data of

unirradiated JRQ are in good agreement. The parameter m in equation (5) were 2.3 and 2.5 for

PCCv adjusted data and I T-CT data, respectively. These values were much less than thetheoretical value of four. Figure 8 shows the Weibull plots of PCCv data of irradiated JRQ. Them values were 2.8 and 3 . I for the data tested at 1 7 ~C and 35 ~C, respectively. The slopes wereagain smaller than the theoretical values. However, it is noted that the m value became a bit large- 277 -

after irradiation.' In other words, neutron irradiation could 'make the data scatter smaller. Thismeans that the scatter after irradiation may be estimated conservatively by applying the scatter
before irradiation.

Figure 9 shows Weibull plot of steels A. In this case, I T-CT data were well fitted bymeans of a theoretical m value. The mean value of corrected PCCv data was higher than that ofl T-CT data. For the data scatter, the corrected PCCv data showed the same tendency as PCCv of
JRQ, that is, the m value was smaller than the theoretical value of four. For the lower bound data,

however, approximately the same trend was seen between both data below 50/0 fracture
probability. If the theoretical m value was used for data fitting, the lower tail of PCCv data ismuch higher than that of I T-CT data.

Figure I O summarizes the Weibull parameter m obtained from this study. Although alarge scatter of m values was seen, the m value tends to decrease as specimen thickness decrease.

To estimate the lower bound from a limited number of PCCv data, it may be helpful to use theexperimental m value instead of the theoretical value .

3.4 Temperature Dependence of Fracture Toughness
The median value of the data can be obtained by Ko

from equation (5) and thefollowing equation as a fracture probability Pf=0.5.

[I

KJC IT(med) = (K0~20) In(2) /4 + 20

(6)Using this median value, the temperature dependence of cleavage

fracture toughness

within the transition range is expressed by the following equation(14.20).

Jc IT(med)30 + 70{ • exp 0.019 • (T - To)}

This equation has been established by statistical analyses using a huge databaseincluding unirradiated and irradiated RPV steels . This (20)equation, which is called a master curve,

uniquely defines the cleavage fracture toughness transition curve. A reference temperature , To'
can be calculated by

T = T I . In{(KJc IT(med) ~ 30)170}

Table 4 Iists the To Values obtained from materials used in this study. Figure I I shows

the values of To based on the method above as a function of specimen thickness. The values of To

from unirradiated JRQ and JSPS A533B-1 showed little effect on specimen thickness, while
those of steels A and B showed a size effect to some extent. The To Values from PCCv data wereabout 20 ~C Iower than those values from I T-CT data for both steels A and B. The deviation of

20 ~C is larger than the deviation of 5-lO ~C reported for some RPV steels in reference (21).

The reason of the deviation is not clear for the moment. Additional tests might help tounderstand the reason because the number of valid data from PCCv specimen of Steels A and B
Figure 1 2 shows all data as a function of the difference between test temperature andTo determined by PCCv data. The fracture toughness values were size-corrected to I T thickness
by equation (4). Below 30 ~C of T-To, the master curve agrees well with the data. At the higher

-278-

(7)

transition range the master curve tends to overestimate the data. Of course, the use of the mastercurve to higher temperature should be restricted because the PCCv data is no longer valid above-200 MPalrm.
Comparing the To before and after irradiation, we obtain the shifi of fracture toughness.

Figure 1 3 shows the master curve for JRQ steel before and after irradiation. Before irradiation,

the TO value was determined by PCCv data tested at -80 ~C and -50 ~C . The To Value after
irradiation was the average oftwo To Values from two PCCv data sets tested at 1 7~C and 35~C.
The shift of To by irradiation to JRQ was then calculated as I 08 ~C. This shift is larger than theCharpy shifi listed in Table 3 by about 20 ~C.

Based on Weibull statistics, tolerance bounds for the master curve can be drawn.

Assuming the Weibull slope m=4, the third parameter K~j.=20 and an mfinrte samplmg slze a
lower bound curve in the same form as equation (7) can be defined. Figures 1 2 and 1 3 indicate

the lower tolerance bound curves. As mentioned before, the experimental data indicated larger

scatter then the theory. The 50/0 Iower bound curve is not enough to bound the experimental
values. Instead of 50/0 bound curve, the I o/o bound curve is in good agreement.

The Japanese code on the method for the fracture toughness test and evaluation, JEAC

4206(22) stipulates the method to obtain the lower bound toughness curve for the integrity
assessment of RPV during a PTS transient. The curve can be defined as the lowest envelope ofall the fracture toughness data points from the surveillance tests. The form of the curve isexpressed by the following:

Figure 1 4 shows the JEAC Klc curve together with I o/o tolerance bound curve based on

the weakest link theory. To compare all data with those curves, the data points without sizecorrection were plotted as a function of T-To' The shape of JEAC Klc curve coincides well withthe lowest data points. The I o/o tolerance bound curve agrees with the JEAC Klc curve in thelower transition range but tends to become higher at higher temperature. Although thiscomparison is limited to only four material, the JEAC method could give the conservative
estimation of the lower bound curve.

Applying equation (8) to the data of JRQ steel, we obtain the lower bound curves
before and after irradiation, and hence the shift of the lower bound curves. Figure 1 5 indicates

the JEAC Klc curves before and after irradiation together with all test data. The curve after
irradiation was determined by shifting the Klc curve before irradiation by the amount of Charpy
shift. The lower bound curves agreed well with the data up to upper transition range, Ieading to aconservative evaluation. When we compare the test data before and after irradiation, we obtainthe shift of ToJ value by irradiation of -80 ~C. This values was slightly less than the Charpy shift.

Therefore, the Klc curve after irradiation became conservative. This may depend on the statisticaltreatment of the data. If we had more data from irradiated specimens or less data fromunirradiated specimens, we might have got less conservative Klc curve after irradiation.

The following equation (9) is the Klc curve in the ASME Sec. XI(1) to describe thelower bound fracture toughness of RPV steels.
RTN/)7' ) f

In this case, we need an RTNDT Of the steel to defme the Klc curve. The RTNDT is- 279 -

determined by the combination of the results from drop weight test and Charpy impact test. The
drop weight test can not be done afier irradiation because the test requires relatively large size
specimens. Therefore, Charpy impact results are used instead of the results of drop weight test to

determine the irradiation effects. The adjusted RTNDT Value after neutron irradiation is thencalculated with the Charpy shift added to the initial RTNDT. The Klc curve afier irradiation isdefined by an equation (9) using the adjusted value. This means that the irradiation effect on

Charpy property and fracture toughness should be equivalent. Figure 1 6 shows the comparisonof the experimental data and the ASME Klc curves. This figure clearly shows that all data arebounded by the ASME Klc Curve based on RTNDT. For the moment, the ASME method gives agood approximation for the lower bound curve. However, the shift of fracture toughness can belarger than Charpy shift as obtained from JRQ results. The lower bound curve for highly
irradiated material should be carefully checked by fracture toughness tests.

4. SUMMARY AND CONCLUSIONSTo establish the method to evaluate fracture toughness values in the transition region,

fracture toughness tests using PCCv to 4T-CT specimens were perfonned. Although post
irradiation tests have finished only for JRQ, the following conclusion were drawn from the testsof four types of Japanese RPV steels and their analysis;( I ) The size effect correction based on the weakest link theory was not always enough to obtain

an equivalent value from a PCCv specimen to a large specimen.
(2) For data scatter, experimantal value of the Weibull slope was not much affected and slightly
decreased by neutron irradiation.

(3) According to the master curve, the shift of the curve by irradiation was evaluated to be
somewhat greater than the Charpy 4 1 J shift.

(4) JEAC Klc curves for JRQ data before and after irradiation could be conservative. However,large number of data before irradiation would be necessary.

ACKNOWLEDGMENTS
A part of this work was carried out under the contract between JAERI and Science andTechnology Agency of Japan under the auspices of the special account law for electric power
development promotion. The authors would like to thank Dr. K. Shibata and Mr. Y. Nishiyama,

Japan Atomic Energy Research Institute for their encouragement and helpful discussion, andmany staffs at Hot Laboratory of JAERI for their careful testing of irradiated specimens. They

are also grateful to Dr. T. Iwadate and Dr. Y. Tanaka of the Japan Steel Works, Ltd. forproviding the JSPS A533B-1 material. The tests of PCCv specimens of JSPS A533B-1 weredone at SCK/CEN during the period that one ofthe authors had stayed there.

REFERENCES(1) ASME Boiler and Pressure Vessel Code, Section XI, "Inservice Inspection
Power Plant," American Society for Mechanical Engineers, 1 995.
- 280 -of Nuclear

(2)Japanese Electric Association, "Methods of surveillance tests for structural materials of
nuclear reactors," JEAC 4201 - 1 99 1 . (in Japanese)(3)

R. Nanstad et al., "Irradiation Effects on Fracture Toughness of Two High-Copper

Submerged-arc Welds, HSSI Series 5," NUREG/CR-5913 Vol.1, October 1992,
U.S.NRC .
(4)

K Wallm et al "Beyond the Charpy - Improved toughness predictions and advanced
surveillance," presented at Radiation Damage and Embrittlement in Pressure Vessel Steels- IGRDM Open Workshop. Santa Barbara. May 6, 1 994.
(5)

M.T. Miglin, et al., "Analysis of Results from the MPC/JSPS Round Robin Testing
Program in the Ductile-to-Brittle Transition Region," Fracture Mechanics: Twenty-fourth

Volume, ASTM STP 1207, J.D. Landes, D. E. McCabe and J. A. M. Boulet, Eds.,
Americal Society for Testing and Materials, Philadelphia, 1995, pp. 342-354.
(6)

T. Iwadate et al., "An Analysis of Elastic-Plastic Fracture Toughness Behavior for JIC

Measurement in the Transition Region," Elastic-Plastic Fracture: Second Symposium,

Volume II-Fracture Resistance Curves and Engineering Applications, ASTM STP 803,
American Society for Testing and Materials, 1 983, pp. II-53 1 -II-561 .
(7)

K. Onizawa et al., "JAERI contribution for the IAEA coordinated research program phase
III on optimizing of reactor pressure vessel surveillance programmes and their analysis,"

JAEFLl-M 93-201 , Japan Atomic Energy Research Institute (October 1 993).
(8)

CRP Sub-Committee, "Manufacturing History and Mechanical Properties of JapaneseMaterials Provided for The International Atomic Energy Agency " Japan Weldmg
Engineering Society, Oct. 1986.

(9)

JSPS 1 29 Committee, "Standard Test Method for Fracture Toughness within DuctileBrittle Transition Range," Japan Society for Promoting Science, 1 995.

( I O)

R. Chaouadi, "Fracture Toughness Measurements in the Transition Regime Using Small

Size Samples," Small Specimen Test Techniques, ASTM STP 1329, W. R. Corwin, S. T.

Rosinski, and E. van Walle, Eds., American Society for Testing and Materials,
Philadelphia, 1 998, to be published.

(ll)

ASTM E399-83, "Standard Test Method for Plane-Strain Fracture Toughness of Metallic
Materials," Annual Book of ASTM Standards, Vol.03.01, American Standard for Testingand Materials, 1983.

( 1 2)

ASTM E813 89 "Standard Test Method for Jlc, A Measure of Fracture Toughness "Annual Book of ASTM Standards, Vol.03.01, American Standard for Testing andMaterials, 1989.

(13)

JSME SOOl-1992, "Standard Method of Test for Elastic-Plastic Fracture Toughness Jlc(Supplement, I st Edition)," The Japan Society of Mechanical Engineers, Feb., 1992.

( 1 4)

ASTM Draft 1 5, "Test Method for The Determination of Reference Temperature, TO, forFerritic Steels in the Transition Range," Rev. 6-13-97, ASTM.

(15)

M. Nevalainen and R Dodds "Numencal Investrgatron of 3 D Constramt Effects on
Brittle Fracture in SE(B) and C(T) Specimens," NUREG/CR-63 17, UILU-ENG-95-2001,
July 1996.

( 1 6)

K. Onizawa, et al., "Investigation on the Evaluation of Cleavage Fracture Toughness Using

PCCv Specimens in the Ductile-Brittle Transition Range of Reactor Pressure Vessel
Steels," JAERI-Research 97-08 1 , Nov. 1 997.
( 1 7)

J.D. Landes and D.H. Shaffer, "Statistical Characterization of Fracture in the Transition- 281 -

region," Fracture Mechanics: Twelfth Conference, ASTM STP 700, American Society for
Testing and Materials, 1980, pp. 368-382.
(18)

K. Wallin, "The Scatter in Klc~Results," Engineering Fracture Mechanics, Vol. 19, N0.6,

pp.l085-l093, 1984.
( 1 9)

T. Anderson and R. Dodds, "Specimen Size requirements for Fracture Toughness Testingin the Ductile-Brittle Transition Region," Joumal of Testing and Evaluation, Vol. 1 9, 1 991 ,pp. 123-134.
(20)

K. Wallin, "Irradiation Damage Effects on the Fracture Toughness Transition Curve Shapefor Reactor Pressure Vessel Steels," Intemational Journal of Pressure Vessel and Piping,55(1993) pp.61-79.
(2 1 )K Wallm "Validity of Small Specimen Fracture Toughness Estimates Neglecting
Constraint Corrections," Constraint Effects in Fracture Theory and Applications: Second

Volume, ASTM STP 1 244, M. Kirk and Ad Bakker, Eds. American Society for Testingand Materials, Philadelphia, 1 995.

(22)

Japanese Electric Association, "Fracture Toughness Test Methods for Nuclear Power Plant

Components," JEAC 4206-1 991 . (in Japanese)

Table I Chemical compositions of materials used in this study.

Table 2 Material properties at room temperature ofmaterials used in this study.

Table 3 Neutron irradiation effect on material properties ofJRQ.

- 282 -

(Wt.olo)Table 4

Comparison of the values of To determined according to the ASTM drafi method.

B: Specimen thickness
N: Total number of data usedr Number of valid data
SG: Side grooving- 283 -

300
JRQ UnirradiatedD
D K(Jlc JSN'E)e

O PCCv
D 0.5TDCTO I TCT

E 200
L.

a

CL

=

~~ 100at 100~C

Ooe~~~~o~~-- -~~~
o
g
o
o

D~

KJc_validity limit(PCCv)oo

Test Temperature, 'CoFigure I --Fracture toughness test results of unirradiated JRQ steel.

300l

JRQ Irradiatede~rl:~~~J~~CT

l 0'5TDCT

I

E 200
L?

~
[~
:~

- - e_ _ _6:~, 100

eC

e

I KJc_validity limit(PCCv)

Jl i

eRI o

Fast Neutron Fluence:

nvi=-2xlO19n/cm2,E>1MeVO
1 oo

Test Temperature, 'C
Figure 2--Fracture toughness test results of irradiated JRQ steel.

300
JSPS A533B-1
O PCC+
D 0.5TCT

E
L* 200~~=e I TCTA 2TCT
X 4TCToClo- - - - - - - - -L- - KJcLvalidity limit

l~ 100~(PCCv)~eo

Test Temperature, 'C
Figure 3--Fracture toughness test results of JSPS A533B- I steel.

- 284 -

300

~ : ITCT

E

x : 4TCT o

~ 200
~
~
:~

l

8

Steel AKJc_validity nmit(Pccv) - - - O

6K(J,c JSME) at 20cc

~
O

~

~o~~~

~__
(O~

~e*~ 100

O-1 50 -1 OOTest Temperature, 'C-50

O

Figure 4--Fracture toughness test results of Steel A.

300
steel B
K(Jlc JsME) at 20'C^o". : !ic""ci= ITCTE
L? 200* = 4TCT

~

c5

~
~:

~-- --- _q~ __ _ ~_KJc_validity limit(PCCv)

~~ ~

o~_

~Z

~ 100

~4

O

- 1 50 - I OO
-50 'C
Test Temperature,O

Figure 5--Fracture toughness test results of Steel B.

54~ JSPS A533B-1 , OPC

A

I Steel A, -80PC
ASteel B, -1 1 O~C

~e~

3

~

Size Effect Correction based\o~41

on Weakest Link Theory

2

1

1

~c=55~80 (MPafm)
O

80
Specimen Gross Thickness, mm

Figure 6--Normalized fracture toughness values as a function of specimen gross thickness.

- 285 -

~~ 1.0~~

JRQ unirradiatedTest temp. -80PC>~' 0.8
.-

~
~ .l

~:
c9
JD

Je

OL 0.6O_

O : PCCv data

o
> 0.4
~'

~,

: Weibull ftting curve

,

cu

l

IS
E:: 0.2l~ : I TCT dataO

- - ,:Weibull type fitting curve

l

0.0

O

1 OO
1 50
50
200
Corrected Fracture Toughness (MParm)

Figure 7--Weibull plots of unirradiated JRQ data at -80 ~C.

1 .O

~
> 0.8

~~

JRQ irradiatedOl-2. I xl 019 n/cm2 (E>1 MeV)

:~:

e

JD
CD
JD 0.6: PCCv data
Test temp. 1 6'C

l

9

ll

o
> 0.4

l

~

•••.••• : Weibull type fittingocurve (m=2.8)

O : PCCv data
Test temp. 35~C

l

CD

~

0.2

~

al o

: Weibull type fitting

curve (m=3.1)

,

0.0
1 OO

1 50

Corrected Fracture Toughness (MParm)
Figure 8--Weibull plots ofirradiated JRQ data at 17 and 35 ~C.

1 .O

~
Q~

~:

Steel A

:E!_> 0.8

9

1 TCT at -80~C

~

,~

,D

,~

JD
9 06

O_

~0

q

~ 0.4

~,.

;5

,g

~>j

E:':' 0.2

O

lll

PCCv at -80~C

q
~:
l

0.0
1 OO

1 50

Corrected Fracture Toughness (MParm)
Frgure 9 Weibull plots of Steel A data at -80 ~C .

~ 286 -

200

10

8

E

~6
o

~

~4
~

~0

~
~

~)

1~;

A

i

: JSPS A533B-1
A: JRQ unirradiated
JL : JRQ irradiatedA

2

O: Stdel A
e: Steel B

O

40 50 6010

30Specimen thickness, mm

Figure I O-Comparison of Weibull slope m values as a function of specimen thickness.
50

O

JSPS A533B-1

~F

c

O
=
~o
E

O

~:

H
(1)

-50

JRQ unirradiated

o

~bo
c

Steel Aooo,U

Steel B

~L -1 OO

Ho
-1 50

O

30
20
40
Specimen thickness, mm

10

50

60

Figure I I --Comparison of reference temperature, To, according to the ASTM draft method
as a function of specimen gross thickness.

400

~E

~ : I TCT

L?co 300

A : 4TCT

Master curve

Au data co~ected to I T sizeO_=~~ 200
F:

- 100
,
~e
o'

(,~

~)e

o'
~~~

e

o

e

~p
Se5 ~efan# hound1 ~efo~ra~e taurtdO

T-To (oO
Figure 1 2--Size corrected data to I T, the master curve and the I o/o and 50/0 tolerance boundcurves detemrined by PCCv data.
- 287 -

300

JRQ

Master curve

O: PCCv

Master curveby PCCv

by PCCv~e~ : CT&DCT oa orre edtolT '

e

E 200

1~

L?
cl'

1 % tolerance bound

o

CL
:~

~

~ ATO

F:

lolo tolerance bound

o'

' 100
~e

o

e

o
o

,

o

~e

e

Open: UnirradiatedClose: Irradiatednvt=-2xl 019nlcm2 , E>1 Mev

O

50
Test Temperature, 'C

1 oO

Figure 1 3 Master curves according to the ASTM draft method before and after irradiation.

400~ : i;'•.~~E 300

,

L?

~
-=200

e,

~F
~~

e

:~~

o 100

o

~

JEAC Ktc curve
A.LS?t, {~~~ Icierance bcund

o
O

-50

50

1 OO

T-To ("OFigure 1 4--Comparison of the JEAC Klc curve and the ASTM I o/o tolerance bound curve.
300

JRQ

~

No size-adiustment

O: PCCv

E 200

L?

(~ : CT&DCT O

o

e

o
9
o
o

O~

a

~
~
~Z 100

~
O

o

~
(,)

~
)1~1'

o

~
e Margin

A T41
Open: Unirradiated

JEAC Lower Bound CurvO

Close: Irradiatedl9
nvt=-2 Io n/cm2
E>1Mev- I oo -50 o 50 1 oo
Test Temperature, 'C

Figure 1 5--Lower bound curves according to the JEAC method steel before and after irradiation.

- 288 -

400^~OA : i ~:";~f~=1TcT

~

E 300

o
O

A = 4TCT

L?

~
-=200

o

oe

o

~e~

~e*~ 100

o

i
o

~
.$,

$
ASME Xte cuwe

O
-1 OO

T-RTNDT ("O50

Figure 1 6-Comparison of all data and ASME Klc curve as a function of T-RTNDT.

~ 289 -