ML20062F976
| ML20062F976 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 09/30/1989 |
| From: | Bamford W WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19310C929 | List: |
| References | |
| WCAP-12447, NUDOCS 9011280295 | |
| Download: ML20062F976 (112) | |
Text
..
WESTINGHOUSE CLASS:3?
WCAP-12447 BACKGROUND AND'. TECHNICAL BASIS:
HANDBOOK ON FLAW-EVALUATION.FOR THE FARLEY UNITS 1 & 2 i
MAIN C00LANT' SYSTEM AND COMPONENTS i
September 1989 W. H. Bamford K. R. Balkey Y. S. Lee H. L. Phillippi Reviewed by: [
f, T. A. Kozlosp
- (
[-
Approved by:
/
e 1
5 7. Palusamy,. Man'ager Structural Materials Engineering-i
?
WESTINGHOUSE ELECTRIC CORPORATION Nuclear. Energy Systems P.O. Box 355 Pittsburgh, Pennsylvania 15230 y
l l
9011280295 901120 t
PDR ADOCK 05000364 l
P PDC l.
. 3418t/41144910 l: i m-s
i TABLE OF CONTE 2' l
Section Title Page 1
1 INTRODUCTION 1 1.1 SCOPE OF WORK 1-1 1.2 CODE ACCEPTANCE CRITERIA 1-3 j
1.2.1 Criteria Based on flaw Size 4 i-1.2.2 Criteria Based on Stress Intensity factor 1-5
[
l 1.2.3 Primary Stress Limits 1-6
)
l 1.2.4 Criteria for Stainless Steel Components 1-6; q
1.3 GE0 METRY AND SOURCES OF DATA 1-7 2
LOAD CONDITIONS, FRACTURE ANALYSIS METHODS,.AND.
2-1 MATERIAL PROPERTIES.
2.1 TRANSIENTS FOR THE-REACTOR VESSELS 2-1 2.2 TRANSIENTS FOR THE STEAM GENERATORS AND PRESSURIZERS 2-2 2.3 TRANSIENTS FOR THE MAIN COOLANT PIPING 2-2 2.4 STRESS INTENSITY FACTOR CALCULATIONS-2 ]
i 2.5 FRACTURE TOUGHNESS' 2-6 '-
2.5.1 Ferritic Steel 2-6 2.5.2, Stainless Steel
'2 "
2.6 THERMAL AGING 2-9 l
2.7 IRRADIATION EFFECTS 2 2.8-ALLOWABLE FLAW SIZE DETERMINATION 2-12 2.8.1 'Ferritic' Steel 2-12 2.8.2 Stainless Steel 2-12 3
FATIGUE CRACK GROWTH 3-l' 3.1 ANALYSIS METHODOLOGY 3-1 3.2 STRESS INTENSITY FACTOR EXPRESSIONS 3 3.3 CRACK GR' WTH RATE REFERENCE CURVES - FERRITIC STEELS 3-3 0
3.4 CRACK GROWTH RATE REFERENCE CURVES---STAINLESS STEELS
.3 3.5~ RESIDUAL STRESSES 3-5 3.6 STRESS CORROS10N CRACKING SUSCEPTABILITY 3-6 l
3818s/111489 10 jj
-- 2m
2 TABLEOFCONTENTS(Cont'd.)
i Section Title Page 4
DETERMINATION OF LIMITING TRANSIENTS ~
4-1 1
4.1 INTRODUCTION
'4-1.
4.2 SELECTION OF GOVERNING EMERSENCY AND FAULTED 4-1 TRANSIENTS - REACTOR VESSELS 4.2.1 Background and History 4-1 4.2.2 PTS Risk for a Typical Westinghouse PWR 4-3 4.2.3 Treatment of Transient Severity 4 4.2.4 Emergency and f aulted -Conditions - Beltline ~
4-6
'I Region 4.2.5 Faulted Conditions Evaluation for.other 4 l Regions 4.2.6 Treatment of Low ~ Temperature Overpressurization 4-8 Transients-(LTOP) 4.3. DETERMINATION OF LIMITING TRANSIENTS -- STEAM 4-23 GENERATORS, PRESSURIZERS, AND MAIN COOLANT. PIPING i
5 SURFACE FLAW EVALVATION 5-1 5.1 CODE CRITERIA 5-1 5.2 LONGITUDINAL FLAWS VS. CIRCUMFERENTIAL FLAWS 5-2 5.3 BASIC DATA 5-2 5.3.1 Fatigue Crack Growth 5-3 5.3.2 Minimum Critical Flaw. Size for 5-3 Ferritic Steels 5.4-TYPICAL SURFACE FLAW EVALUATION CHART 5-4' 5.5 PROCEDURE FOR THE CONSTRUCTION OF A SURFACE FLAW 5-5 EVALUATION CHART -- FERRITIC STEELS 5.6 SURFACE FLAW EVALUATION -- STAINLESS STEEL COMPONENTS 5-8 6
EMBEDDED FLAW EVALUATION 6 6.1-EMBEDDED VS. SURFACE FLAWS 6-1
.mwmm io
$$3 l
-1 TABLEOFCONTENTS(Cont'd.)
Section
. Title Page 1
6.2 CODE CRITERIA -- FERRITIC STEELS-6-2 6.3-BASIC DATA -- FERRITIC STEELS 3 l
6.4 FATIGUE CRACK GROWTH FOR EMBEDDED FLAWS 6-4 I
6.5 TYPICAL EMBEDDED FLAW EVALUATION CHART 6-6 l
6.6 PROCEDURES FOR THE CONSTRUCTION OF. EMBEDDED FLAW 6-9 EVALVATION CHARTS 6.7 COMPARISON OF EMBEDDED FLAW CHARTS WITH' ACCEPTANCE 6-11 STANDARDS OF IWB-3510 l
7' REFERENCES 7-1 i'
I r
l l
-l l
f I
e
'l 3814s/111489 10 g
SECTION 1 INTRODUCTION The flaw
- evaluation handbooks, which have been published unter separate cover (1,2,3), have been designed for the evaluation of indications which may be discovered during inservice inspection of the farley main coolant system components.
The tables and charts provided therein allow the evaluation of any indication discovered in-the regions listed below without further fracture-mechanics calculations.
The fracture analysis work has been done in advance, and is documented in this report.
Use of these handbooks will allow the acceptability of much larger indications than would be alloweble by only using the standards tables of the ASME Code,Section XI [4).
This report provides the background and technical basis for these handbooks.
The flaw evaluation handbooks have been developed as a direct implementation of the analytical requirements of Section XI.
In all cases the latest material properties and analytical criteria have been used and the appropriate code editions appear in (4). These handbooks have been designed for use throughout the operating lifetime of.the plants, during which time the applicable edition of section.XI will change a. number of times.
1.1 SCOPE OF WORK l
Handbooks have been developed for the following locations in the Farley main coolant system:
Reactor Vessel:
o Beltline (core region) { Note that-this region includes both longitudinal and circumferential welds) o Inlet nozzle:
nozzle-to-shell weld, safe end-to-nozzle weld and inner radius (corner)
- The use of the term " flaw" in this document should be taken to be synonymous with the term " indication" as used in section XI of the ASME Code, wwume,o 11
l i
o Outlet nozzle:
nozzle-to-shell weld, safe ~end-to-nozzle weld and:
innerradius(corner).
o Lower head:
ring-to-lower head weld. and ring-to-lower shell weld o
Top head:
dome-to ring weld, and ring-to-flangeLweld o
W ssel lower flange-to-shell weld, and upper'_shell-to middle shell:weldL E
w e
E 5tesin Generatcr:
i o
Tubesheet-to channel head welo region 1
F o
Tubesheot.to-stub barraT weld' region g
o Stub barrel welds f
o lower shell-to-cone weld region o
Upper shell-to-cone' weld region i
I o
Uppor shell welds.
i E
o Upper shell-to-dome weld region I
o feedwater nozzle-to-shell weld' region y
o Steam outlet nozzle-to-shell weld region o
Primary nozzle safe-end weld; 3
Pressurizer:
o Upper shell-to-head weld o
Upper shell circumferential welds o
Upper shell longitudinal welds o
Lower shell circumferential welds o
Lower shell longitudinal welds o
Surge nozzle-to-head weld o
Spray nozzle-to-head weld o
Safety and relief nozzle-to-head w' eld-Main Coolant Piping:
Hot Leg Five locations, including junctions with reactor vessel and steam generator l
wwmeio 1-2 l
1
Cold Leg Four locations,. including junctions with reactor vessel and pump Crossover Leg Six locations, including junctions with pump and steam generator The geometry of these regions is shown in figures 1-1 through 1-6.
1 The highlight of the handbooks is a-series of flew eva'luation charts for both
'b surface flaws and the embedded flaws.
Since the characteristics of the two?
l types of flaws are different, the evaluation charts designed for each are i
distinctively different in style.
Section 5 of this technical basis documen.t deals with surface flaws at various locations, and section 6 concentrates on' the evaluation.of embedded flaws.
- h
'l The flaw evaluation charts were designed based on the Section XI Code criteria of acceptance for continued service without repair.
Through use of the charts, a flaw can be evaluated instantaneously, and no follow-up hand calcu-lation is required. ~ Host important-of all, no. fracture mechanics knowledge is needed by the user of the handbook charts..
It is important to note that indications which exceed the standards limits, and must be evaluated by fracture mechanics and will also require additional inservice inspection in the future, as discussed in Section XI,' paragraph IWB 2420.
The analytical evaluation results must be submitted.to.the regulatory 3
authority having jurisdiction at the plant site, 4
/=
1.2 CODE ACCEPTANCE D ITERIA There are two alternative sets of flaw acceptance criteria for continued service wit'hout repair in paragraph IWB-3600 of ASME' Code Section XI [4).
Either of the criteria below may be used, at the convenience of the user.
1.
Acceptance Criteria Based on Flaw Size (IWB-3611) 2.
Acceptance Criteria Based on Stress Intensity Factor (IWB-3612).
mwmm io 33 l
l
Both criteria are comparable in accuracy for thick sections, and the acceptance criteria based on stress intensity factor have been assessed by-past experience to be less restrictive for thin sections, and for outside surface flaws in many caras..In all cases, the most beneficial criteria have been used in developent of the flaw evaluation charts.
Since.the frau ure mechanics results for surface flaws have been presented in terms of critical flaw size, it is more straight forward to construct the-surface flaw evaluation charts by using criteria (1) in this handbook.
This has been done for inside surface flaws-in all cases except-the reactor vessel safe end region, and the thinner sections of the steam generator secondary side, where criteria (2) are more beneficial because of the small section thickness.
All of the embedded flaw charts and most outside surface flaw evaluation charts in this handbook were constructed using acceptance criteria (2), for ease of use, as well as to obtain the maximum benefit, sin g these criteria will generally be less restrictive for embedded flaws.
1.2.1 CRITERIA BASED ON FLAW SIZE-The code acceptance criteria stated'in IWB-3611 of section XI are:
.1 a For Normal Conditions af c
(Upset & Test Conditions Inclusive) f
.5 a$
For Faulted Conditions and a
l (Emergency Condition inclusive) l l
where The maximum size to which the detected flaw is calculated a
=
f to grow at the end of a specified period, or until the next inspection time..
1 L
The minimum critical flaw size under normal operating a
=
c conditions (upset and test conditions inclusive) 3816s/112789 10
. },. 4,
ag The minimum critical flaw size for initiation of
=
nonarresting growth under postulated faulted conditions.
(emergency conditions inclusive)
To determine whether a surface flaw is acceptable for continued service without repair, both criteria must be met simultaneously. However, both criteria have been considered in advanco before the charts were constructed, y
Only the most restrictive results were included in these charts.
1.2.2 CRITERIA BASED ON STRESS INTENSITI FACTOR As mentioned in the preceeding paragraphs, the criteria used for the evaluation of embedde.d flaws, including most outside surface flaws and those in the nozzle safe-end regions are from.lWB-3612 of.*2ction XI.
The term stress intensity factor (K ) is defined as_the driving force on a j
crack. -It is a function of the size of the crack and the applied stresses, as well as the overall geometry of the structure, in contrast, the fracture toughness (K,, Klc) is a measure of the resistance of the material to j
propagation of a crack, it is a material property, and a function of temperature.
1 The criteria are:
K la K; <
for normal conditions (upset & test conditions inclusive)
)
/ 10 K Ic f -
3 < /2 for faulted conditions (emergency conditions inclusive)
K where K).
The maximum applied stress intensity factu for the flaw
=
size a to which a detected flaw will grow, dur;ing the f
conditions under consideration, for a specified period, or to the next inspection.
wwwwo 1s
WESTINIH!USE PEOPRIETARY CLASS 2.
K, Fracture toughness based on crack' arrest for'the.
7
=
j corresponding crack-tip temperature.
fracture toughness based on. fracture initiation for the K
=
,1
!c corresponding crack tip temperature.
To determine whether a flaw is~ acceptable ;for continued service without repair, both criteria must be met simultaneously.
However, both criteria have j
been considered in advance before:the' charts were-constructed.
Only_the most-l restrictive results were included in.the ' harts, t
c 1.2.3 PRIMARY STRESS' LIMITS a
in addition'to satisfying the fracture criteria, it is. required-that the primary stress limits of-the ASME Code Seition Ill', paragraph NB 3000 be -
satisfied.
A local area' reduction of the pressure retaining membrane must be.
used, equal to the area of_ the indication, and _the stresses increased to-reflect the smaller cross section.
All the-flaw'a'cceptance tables provided in this handbook have included this consideration,-asidemonstrated herein.. The allowable flew depths determined using this criterion-have been summarized in.
tables 1-1 through 1-3 for each of the locations in the: reactor ve'ssel, veam '
generators, and pressurizer for which handbook charts' have been constructed.
Such calculations are not necessary for the main coolant piping, because -they.'
are directly included in the evaluation.
1.2.4 CRITERIA FOR STAINLESS STEEL COMPONENTS y
The evaluation procedures and acceptance criteria for indications in austeni-tic stainless piping are contained in paragraph IWB 3640 of;ASME Section XI.[4] The evaluation procedure is applicable to all the materials within a specified distance from the weld centerline, / rt, where r = the pipe nominal outside radius and t is the wall thickness.
For example, at :the cold leg elbow to valve region, this distance is calculated to be 6.2 inches, which encompasses regions of the elbow and weld.
L i
{.
wwswee so 16
st 7
l t
The evaluation process begins with a flaw growth analysis, with the require-ment to consider growth due to both fatigue and stress corrosion cracking.
I for pressurized water reactors only fatigue crack growth need be considered, i
as discussed in section 3.
The methodology for the fatigue crack growth l
analysis is described in detail in section 3.
i The calculated maximum flew dimensions at the end of the evaluation period are i
then compared with the maximum allowable flaw dimensions for both normal.
l operating conditions and emergency and faulted conditions, to determine j
acceptability for continued service.
Provisions are made for considering flaws projected both circumferentially and axially, in IWB 3640 the allowable flaw sizes have been defined in the tables based on maintaining specified safety margins on the loads at failure.
These margins are 2.77 for normal and upset conditions and 1.39 for emergency and faulted I
conditions.
The calculated failure loads are different for tM base metal and the flux welds, which have different fracture toughness valuu, as discussed in section 2.
The failure loads, and consequently the allowable flaw sizes, are larger for the base metal than for the welds. Allowable flaw sizes for welds are contained in separate tables, in IWB 3640, 1.3 GEOMETRY AND SOURCES OF DATA The geometry of the main coolant system components and piping is shown in Figures 1-1 through 1-6.
Where present, tle cladding on the inside of the vessel has been neglected in the stress analysis.. It has been accounted for in the thermal analysis by adjusting the film coefficient for the conditions analyzed. The outside surfaces have been assumed to be. insulated.
The notation used for both surface and embedded flaws in this work is illustrated in figure 1-7.
The fracture and fatigue crack growth evaluations carried out to develop the handbook charts have employed the recommended procedures and material properties for low alloy steels, as contained in Section X1, Appendix A.
The mwmm io 17
flaw evaluation for stainless steel-components-(safe-end. and piping) has used the analysis methods and properties'of_ stainless steel as prescribed in paragraph IWB 3640 and Appendix C of Section XI, me.inun io 1-8
I-TABLE 1-1
SUMMARY
OF ALLOWABLE FLAW. DEPTHS.'
BASED ON PRIMARY STRESS LIMIT CRITERIA - REACTOR-VESSEL i
ALLOWABLE l DEPTH-ALLOWABLE DEPTH OF FLAW, a/t 0F FLAW, a/t?
REGION-(longitudinal)
(circumferential)
H a,c,e Beltline Nozzle Shell-to Intermediate Shell Weld Inlet Nozzle to Shell Weld i
Inlet Nozzle Inner Radius Inlet Nozzle Safe-End Outlet Nozzle to Shell Weld Outlet Nozzle Inner Radius J
Outlet-Nozzle Safe-End Lower Head Ring to Shell Weld-Lower Head to Dome Weld
- Top Head Dome to Ring Weld Top Head Ring to Flange Weld
-i
(
t s
mia.niim io 19 f
9 TABLE 1-2
SUMMARY
OF ALLOWABLE FLAW MPf tG BASED ON PRIMARY STRES!; LIMIT CRITERIA - STEAM GENERA 70R Allowable Depth Allowable Depth Region of Flaw, a/t of Flaw, a/t (longitudinal)
(circumferential) 1 Inner Outer Inner Outer Surface Surfacs Sur ace Surface a,c.e
~~
not applicable t
nie.num to 1-10
l
?
l TABLE 1-3
SUMMARY
OF ALLOWABLE FLAW DEPTHS BASED ON i
PRIMARY STRESS LIMIT CRITERIA - PRESSURIZER I
Allowable Depth Allowable Depth Region of Flaw, att of Flaw, a/t (longitudinal, (circumferential)
Inner
<0 uter Inner Outer Surface Surface Surface Surface p-
a,c.e Upper Shell to Head Wald Upper Shell Welds Lower Shell Welds Surge Nozzle to Shell Weld Spray Nozzle to Shell Weld Lower S'a'l to Head Weld F
8 9
m e.nii. ie 1 31
]
l i
o 3
~
T--~~"
k l
i l
ed h
96.07" n
I I
n MIDDLE-TO-UPPER ~
~~----
CIRCUMFERENTIAL l
WELD 100.53"
~ 7.88" i
LOWER-TO-MIDDLE ~
o
~ " " - - - -
CIRCUMFERENTIAL n
l WELD i
100.66" l
LOWER HEAD RING TO u
1 l
LOWER SHELL WELD n 29.72" l
LOWER HEAD RING TO n
LOWER HEAD WELD l
. o
--,,111 h
5.00" l
NOTE: THICKNESSES DO NOT INCLUDE INSIDE CLADDING l
Figure 1-1, Reactor Vessel Welds (Unit I has peel segments in the bottom head and upper head regions) mwnsm
- 1-12 s
l~
TOP HEAD RING TO DOME WELD TOP HEAD RING 79.09R TO FLANGE WELD f
6.19 i
28,75 l
l 14.81 DI A 1
g a
s, -
i l 4P 14.560I A 35.06 TO S 155.500IA 9.125 +
+
l B.03 (BASE METAL)
=
79.53R (BASE METAL)
I LOWER HEAD RING TO LOWER HEAD WELD LOWER HEAD RING
{lAl5R TO LOWER SHELL. WELD 5.00 l
NOTES:
I. DIMENSIONS DO NOT INCLUDE CLADDING
- 2. ALL DIMENSIONS ARI IN INCHES Figure 1-2. Upper Head and Lower Head Regions.(dimensions in inches)
Note:
Unit I has peel segments in the upper head and bottom head-me.mv.. io y,g c.
l
\\
SIDE VIEW TOP VIEW w.i nni.
ht
- 155.5 10.
IN 9.12 -
l l
- i t
l
(
NC ZLE TO l
- - g, 3 l
CLADDING 38.48 L
j l
i i
O
- --- 27. 47 +j 33.07 %
~
=
55.5
=
77.75 TO l
VESSEL Q, jNNER RADIUS O.156 MIN.
btlet bule CLADO!NG
- 3.25 9.12 -
I f U
9 F 44 53
- - 0. 25 CLAD (TYP) g 35.4I h
- ~"" 28+97
- 12.13 s
g F
F
'I h
5 i
I
=
35.50
=
51.00 1-NOTES:
l
- 1. DIENSIONS DO NOT INCLUDE CLAD
- 2. ALL DIMENSIONS ARE IN INCE S Figure 1-3.
Geometry of reactor Vessel Inlet and Outlet Nozzles me.mme to 3 34
~.. --
1 i
3 823h o
o N
63.26 l
l h
3.62 168.40 f.D.
\\\\,
q' Il 113.49 p
e 1i q
"~
mammmer.
b OI
[jf MLJM 113 49 A,,.....
)
3
.g i
t i
3.48 76.94 A
C 412.00 I30 #
2.82 179 38 1.D.-
.38 123.60 1.D.-
i E
4 138.00 3.26 --
.76 76.00 h
f ) d';75, 9 I II U 4 IMfssh%AWT 27.00
- b. 62.8)(
" h 16.00 86.00 8
f 31.20
-- =
v v
l Figure 1-4.
Geometry of the farley Units 1 and 2 Steam Generators (Model 51) nimitano in y,
_ ~
i i
1 y
i l 5 :\\i,/ 't,1r
/
ikly 1
i n
/
d 4
.,r gs i
f i
a x
I i
k Ik Dh5 I
i.
1 l
e.
t s
ei
- y
\\
-yf-l-km
.k, ji j- -!
'I r,
1 1
i
!i
~,
l!
l 8
E i
_= lg as Figure 1-5.
Geometry of the Pressurizer me.nime in 1 16
.. = - -.. - -. -. -.. -.
.~.
i i
l i
I 1
4 i
=
as.ie
=
i o
i i
3.88 MIN.
l 43.75 a l l o
/
g i
44*-13' s
v 2
28.54 i'
U
- s c:
9 F"
n i
o i
Figure 1-6.
Head Geometry for Pressurizers 1-17'
=ie.mine io
v O
4 v
v bM i
av bf
%J V
l j
v Yi W
=&
1/g Figure 1-7.
Schematic of Reactor Coolant loop, Farley Units 1 and 2.
4 m..n un, u 1-18
i Wall Thickness t Wali Thickness t _
l
?
r i
f I
i 6"
W' v )
tr r.
E
$.-)
7 s --
G
-+
- a 1**
l k,
A i
i TYPICAL EiSEDED FLAW INDICATION TYPICAL SURFACE FLAW INDICATION Figure 1-8.
Typical Notation for Surface and Embedded Flaw Indications r
3310+/t 12F9919
.n.--..-..
-. - ~.
4,
,e
_,.~ _ -
SECTION 2 LOAD CONDITIONS, FRACTURE ANALYSIS METH005 AND MATERIAL PROPERTIES The loacing conditions used in the analyses described herein were taken directly from the equipment specification for each of the components covered, and from the piping design specification. The fracture analysis methods are the most advanced which are now available, and the material properties are the latest available properties contained in the ASME Code.
l 2.1 TRANSIENTS FOR THE REACTOR VESSEL l
The design transients.for the Farley reactor vessels red in Table 2-1.
l Both the minimum critical flaw sizes (such as a unde ermal operating c
l conditions, or ag under f aulted conditions for critt- - (1) of IWB-3611) and the stress intensity factors (K. for criteria (2) of iWB-3612) are a g
l function of the stresses at the cross-section where the flaw of interest is i
located, along with the material pr'operties.
Therefore, the first step for the evaluation of a flaw indication is to determine the appropriate limiting icad conditions for the location of interest.
It should be noted that the Farley plants were designed to a slightly different set of transients than those used in the analysis here.
Although the design transients are not identical, as can be seen in table 2-1, they are very similar, in general, the transients used in these analyses are more numerous than those for Farley.
The transients which have been used as the design basis for later plants are generally more accurate, because they reflect greater operating experience.
For this reason the transients used in the evaluation charts for the reactor vessels are considered to be conservative.
The selection of the most limiting tr,ansient for normal / upset / test conditions was straightforward.
The transient with the highest surface stress in the area where the flaw was postulated was chosen as the worst case.
Note that this can result in a different limiting transient for an inside flaw as me,mene a 2-1
f~
opposed to an outside flaw, as may be seen in the detailed treatments of the individual locations. The selection of the most limiting emergency and faulted condition transient is discussed section 4.
2.2 TRANSIENTS FOR THE STEAM GENERATOR AND PRESSURIZER The design transients for the steam generators and pressurizers are listed in Tables 'i-2 and 2-3.
For the pressurizer, a number of heatup/cooldown events have been included.
These correspond to differcnt spray activations, which affect only the pressurizer, and are within the original design specification.
These transients are used for accurate fatigue crack growth analysis of the pressurizer.
Both the minimum critical flaw sizes, (a e
under normal operating conditions, or ag under faulted conditions) and K; are functions of the stresses at the cross-section where the flaw of interest is located, along with the material properties.
Therefore, the first step for the evaluation of a flaw indication is to determine the appropriate limiting load conditions for the location of interest.
l The selection of the most limiting transient for normal / upset / test conditions was straightforward here also.
The transient with the highest surface stress in the area where the flaw was postulated was chosen as the worst case.
Note that this can result in a different limiting transient for an inside flaw as opposed to an outside flaw, as may be seen in the detailed treatments of the individual locations.
2.3 TRANSIENTS FOR THE biAIN C001. ANT PIPING The design transients for the primary system are contained in table 2-4.
Both the critical flaw size and the fatigue crack growth used in flaw evaluation are functions of the stresses at the cross-section where the flaw of interest is located, and the material properties'.
Therefore, the first step for the evaluation of a flaw is.to determine,the appropriate limiting load conditions for the location of interest.
This has been done, and is discussed next.
The loading conditions which were evaluated include thermal expansion (normal and upset), pressure, deadweight and seismic (OBE and SSE) loadings.
The mwm u
.2-2
forces and moments for rach condition were obtained from the ASME Code Section
!!! calculations prev'ously performed by Westinghouse.
Residual stresses were not used in this portion of the evaluation, in compliance with the Code guideline.
Stresses as appropriate were used in the fatigue crack growth analysis, as discussed in section 3.0.
The stress intensity values were calculated using the following equations:
51 = P, + Pb
x where F = axial force component (membrane) x M,, M, H *momentcomponents(bending) y 2
A = cross-section area Z = section modulus The section properties A and Z at the weld location were determined based on the minimum pipe dimensions.
This is conservative since the measured wall thickness at the weld is generally larger.
The following load combinations we e used.
A.
Normal / Upset - Primary Stress Pressure + Deadweight + OBE B.
Emergency / Faulted - Primary Stress Pressure + Deadweight + SSE C.
Expansion Stress - Secondary Stress i) Normal Thermal ii) Upset Thormal wwnme so p.3 l
I
I D.
Normal / Upset Total Strese-I i) Pressure + Deadweight + OBE + Normal Thermal i
l ii) Pressure + Deadweight + Upset Thermal l
E.
Emergency / Faulted - Total Stress i) Pressure + Deadweight + SSE + Normal Thermal ii) Pressure + Deadweight + Faulted Thermal In D and E above, load co,abination (i) is the governing case.
2.4 STRESS INTENSITY FACTOR CALCULATIONS One of the key elements of the critical flaw size calculations is the determination of the driving force or stress intensity factor (K ).
This j
was done for each of the regions using expressions available from the literature, in all cases the stress intensity factor for the critical flaw size calculations utilized a representation of the actual stress profile rather than a linearization.
This was necessary to provide the most accurate determination possible of the critical flaw size, and is particularly important for consideration of emergency and faulted conditions, where the stress profile is generally nonlinear and often very steep.
The stress profile was represented by a cubic polynomial:
o(x) = A0+Ag{+A2({} +A3({}
(24) where x is the coordinate distance into the wall t = wall thickness o = stress perpendicular to the plane of the crack
)
Ag = coefficients of the cubic fit for the surface flaw with length six times its depth, the stress intensity factor expression of (McGowan and Raymund (5))8'C was used. The stress intensity factor K; (e) can be calculated anywhere along the crack front.
m.. mom to 2-4
i The point of maximum crack depth is represented by v = 0.
The following expression is used for calculating K; (e), where e is the angular location around the crack.
~
a,c.e (2-2)
The magnification f actors H ('I' H)(,), H ($) and H (e) are obtained by the 0
2 3
procedure outlined in reference
.]aC,e i
The stress intensity factor calculation for a semi-circular surface flaw.
(aspect ratio 2:1) was carried out using the expressions developed by [
lac.e.
Their expression utilizes the same cubic representa-tion of the stress profile and gives precisely the same result as the expression of (
.)a,c.e for the 6:1 aspect ratio flaw, and the form of the equation is similar to that of [
.)a,c.e above.
The stress intensity factor expression used for a continuous surface flaw was that developed by [
Ja c.e.
Again the stress profile is represented as a cubic polynomial, as shown above, and these coefficients as well as the magnification factors are combined in the expression for Kj
.a,c.e (2-3) a,c.e where F, F, F, F are magnification factors, available in 3
2 3
4 The stress intensity factor calculation for an embedded flaw was taken from work by (
3a,c.e which is applicable to an embedded flaw in an infinite medium, subjected to an arbitrary stress profile.
This expression has been shown to be applicable to embedded flaws in a thick-walled j
pressure vessel in a paper by {
Ja.c.e, j
me,mo. i.
2-5
.a j
2.5 FRACTURE 100GHNESS 2.5.1 Ferritic Steel The other key element in the determination of critical flaw sizes is the fracture toughness of the material.
The fracture toughness for ferritic steels has been taken directly from the reference curves of Appendix A, i
Section XI.
In the transition temperature region, these curves can be represented by the following equations:
K;c = 33.2 + 2.806 exp. (0.02 (T-RTNDT + 100*F))
(2-4)
Kla = 26.8 + 1.233 exp. (0.0145 (T-RTNDT + 160'F))
(2-5) where K and K;, are in ksi/in.
je The upper shelf temperature regime requires utilization of a shelf toughness which is not specified in the ASME Code.
A value of 200 ksi/in has been used here for all the regions.
This value is consistent with general practice in such evaluations, as shown for example in reference (10), which provides the background and technical basis of ppendix A of Section XI.
The other key element in the determination of the fracture toughness is the value of RTNOT, which is a parameter determined from Charpy V-notch and drop weight tests.
The material chemistry and initial RTNDT values for all the welds, plates and forgings in the Farley reactor vessels are provided in tables 2-5 and 2-6.
The core region materials are identified in figures 2-1 and 2-2.
This information was determined from the vendors material certification reports, surveillance capsule tects, and weld chemistry studies by Westinghouse, EPRI, Combustion Engineering, and others.
When no information on the chemistry or RTNDT was available, conservative assump-tions were made, and these cases are clearly marked in the tables.
The limiting material properties from all the vessels were used in the analyses here.
mwnne io 2-6
The fracture toughness of steam generator and pressurizer materials has been examined in recent years relative to the reference toughness curves of the ASME Code.
(
.)a,c,e, The RT values used for the steam generator and pressurizer materials were NDT based on estimation procedures, because precise values for this parameter were not always determined at the time of manufacture. The materials were required (by the ASME Code) to meet a specification of Charpy energy equal to 30 f t-lb,;
as the average of 3 tests at 10'F.
Using the (
3a,c.e, 2.6.2 Stainless Steel The primary loop piping is cast stainless steel SA351, type CF8A or CF8M.
The welds in the primary loop piping were made by either submerged arc or shielded metal are processcss.
The fracture toughness of the base metal has been found to be very high, even at operating temperatures (16), v.here the Jlc values have been found to be 2
well over 2000 in-lb/in. Fracture toughness values for weld materials have been found to display much more scatter, with the lowest reported values significantly lower than the base metal toughness.
Although the J;e values reported have been lower, the slope of the J-R-curve is still large for these J
cases.
Representative values for J were obtained from the results le le of Landes, et al. [17), where the following values were obtained, and used in the development of the fracture evaluation methods:
wwume so g.y l
l-
S
(
)a c.e 2.6 THERMAL AGING Thermal aging at, operating temperatures of reactor primary piping can reduce the fracture toughness of cast stainless steels and, to a lesser degree, stainless steel weldments, The cast stainless steel piping and elbows of the primary loop are very tough, usually exhibiting J;c values exceeding 2000 2
in-1b/in and a tearing modulus, Tmat, well over 200.
NRC procedures i
exist for addressing the impact of thermal aging on fracture toughness for full-service life.
The approved procedures were applied to all the components for which flaw evaluation charts were constructed.
(
3a,c.e,
(
3a c.e
(
)a,c.e it should be recognized that these value: are believed to be very conservative estimates of the end of life fracture properties.
2.7 IRRADIATION EFFECTS - REACTOR VESSEL Neutron irradiation has been shown to produce embrittlement which reduces the toughness properties of reactor vessel steels.
The decrease in the toughness properties can be assessed by determining the shift to higher temperatures of ws,is uae so y.g 1
the reference nil-ductility transition temperature, RTNDT.
Because the chemistry (especially copper and nickel content) of reactor vessel steel has been identified as a major contributor to irradiation enbrittlement, trend curves have been developed to relate the magnitude of the shift of RTNDT to the amount of neutron fluence.
The reference fracture toughness curve, indexed to RTNDT, will shif t along the temperature scale with a value equal to the increase in the RT for given levels of irradiation.
NDT Based on the initial RTNDT value and the material chemistry of the limiting values are determined from core region materials, the post irradiation RTNDT the trend curves.
These final RTNDT values are subsequently used to calculate K and K as a function of the fractional depth through the yg ia wall.
Irradiation effects were accounted for in all regions analyzed, but only had a significant impact on the properties in the beltline region.
The extent of the shift in RT is enhanced by certain chemical elements NDT (such as copper, nickel and phosphorus) present in reactor vessel steels.
i Westinghouse, other NSSS vendors, the U.S. Nuclear Regulatory Commission and others have developed trend curves for predicting adjustment of RTNDT as a function of fluence and copper, nickel and/or phosphorus content.
The Nuclear Regulatory Commission (NRC) trend curve is published in Regulatory Guide 1.99.
Regulatory Guide 1.99 was originally published in July 1975 with a Revision 1 issued in April 1977, and revision 2 (18) issued in July 1988.
The chemistry f actor, "CF" (*F), a function of copper and nickel content identified in Regulatory Guide 1.99, Revision 2 is given in table 2-7 for welds and table 2-8 for base metal (plates and forgings).
Interpolation is permitted. The value, "f", is the calculated value of the neutron fluence at the location of the postulated defect (n/cm3 (E > 1 MeV)) divided by 19 10 The fluence factor is determined from figure 2-3.
The Adjusted Reference Temperature (ART) based on the methods of Reg. Guide 1.99 Revision 2 can be compactly described by the sequence of equations listed below:
ART = Initial RTNDT + 6RTNDT + Margin (2-6) mwmano 2-9
I i
i X=Depthintovesselwallfrominner(wetted) surface (1/4Tand3/4T).
(2-7) e,RT SVRFACE = [CF)F (0.28 - 0.10 LOG F)
(2-8)
)
~
19 F = Neutron fluence divided by 10
.(2-9)
F=F EXP(-0.24X) surf CF=Chemistryfactorfromtables*(ifnodatause 0.35%Cuand1.0%Ni)
(2-10) 3 )0.5 (2-11) 2 MARGIN = 2 (ej2+o og = Maan value of' initial RTNDT; if initiai RTNDT measured, og = 0, otherw';.e og obtained from set of data,to get (2-12) initial RTNDT.
Standard deviation of initial:RT (2'13) o 3
NDT 28'F for welds-17'F for base metal (o need not exceed 1/2 times RTNDT surface) 3 i
l l
- See tables 2-7 and 2-8.
i uss.msess so
^
g.gn.
2.8 ALLOWABLE FLAW SIZE DETERMINATION 2.8.1 Ferritic Steels The applied stress intensity factor (K;) and the material fracture toughness values (K;, and K;c) can be used to determine the critical flaw size values used to construct the handbook charts.
For normal, upset and test conditions, the critical flew size a is determined from the depth at which c
the applied stress intensity factor Kg exceeds the arrest fracture toughness Ela' For emergency and faulted conditions the minimum flaw size for crack initia-tion is obtained from the first intersection of the applied stress intensity factor (K;) curve with the static fracture toughness (K;g) curve.
IntersectionoftheK;curvewiththecrackarresttoughness(K),) curve determines the crack arrest size.
The critical flaw depth for emergency and faulted conditions (ag) as defined earlier, is the minimum flaw depth for initiation of non-arresting growth.
Non-arresting growth is defined as growth which arrests at a depth greater than 75 percent of the wai' depth.
An example of this type of calculation is shown in figure 2-4.
The critical flew depth is determined at point A in this figure.
2.8.2 Stainless Steels 4
The critical flaw size is not directly calculated as part of the flaw evalua-l tion process for stainless steels.
Instead, the failure mode and critical flaw size are incorporated directly into the flaw evaluation technical basis, and therefore into the tables of
- Allowable End-of-Evaluation Period Flaw Depth to Thickness Ratio," which are contained in paragraph IWB 3640.
Rapid, nonductile failure is possible for ferritic materials at low tempera-tures, but is not applicable to stainless steels.
In stainless steel materials, the, higher auc M H ty leads to two possible modes of failure, plastic collapse er unstable ductile tearing.
The second mechanism can occur when the applied J integral exceeds the Jlc. fracture toughness, and some mwm*"
2-11
l stable tearing occurs prior to failure.
If this mode of failure is dominant, the load carrying capacity is less than that predicted by the plastic collapse mechanism.
The allowable flaw sizes of paragraph IWB 3640 for the high toughness base materials were determined based on the assumption that plastic collapse would
-i be achieved and would be the dominant mode of failure, l.
l 1
Ja.C e i
i l
l I
mie.n new ie 2-12 1
TABLE 2-1 dVMMARY OF REACTOR VESSEL TRANSIENTS 4
NUMBER OF OCCURRENCE 5 USE0 IN THE NUMBER TRANSIENT IDENTIFICATION SPECIFIED ANALYSIS Normal C,onditions 1
Heatup and Cooldown at 100*F/hr (pressurizarcooldown200'F/hr) 200 200 2
Load follow Cycles (Unit loading and unloading at 5% of full power / min) 18300 18300 3
Step load increase and decrease of 10% of full power 2000 2000 4
Large step load decrease, with steam dump 200 200 6'
5 Steady state fluctuations Infinite 10 80 6
Refueling 26400 7
Boron Concentration Equalization 80 8
Loop Out of Service Upset Conditions 9
Loss of load, without immediate turbine 80 200 or reactor trip 1
10 Loss of powe- (blacksut with natural circulation in the deactor Coolant System 40 40 11 Loss of flow (partial loss uf flow, one l
pumponly) 80 80 12 Reactor trip 400 400 no cold shutdown (230) e no safety injection (160)
E with cooldown and S.I. (10) 13 Inadvertent Auxiliary Spray 10 60 mwmm io 2-13
'i TABLE 2-1
SUMMARY
OF REACTOR VESSEL TRANSIENTS (cont.)
l NUMBER OF OCCURRENCE 5 l
USED IN THE NUMBER TRANSIENT IDENTIFICATION SPECIFIED ANALYSIS 30 14 Excessive feedwater Flow 4
15 Inadvertent Depressurization 20 10 l
16 Inadvertent Startup of an Inactive Loop 80l j
17 Control Rod Drop
.70 f
18 Loop Out of Serv. ice Faulted Conditions 19 Large Loss of Coolant Accident (LOCA) 1 1
f 20 Large Steam Line Break (LSB) (other transients described in section 4) 1 1-l 21 Safe Shutdown Earthquake 1
1
~;
i Test Coaditions 22 Turbine roll test.
-20
-80 23 Primary Side Hydrostatic leak test conditions.
50 280
~
24 Cold Hydrostatic test 9 3105.psig 5
10 v
i i
a m e.n m...io 2-14 i'
n:
-l
TABLE 2-2
SUMMARY
OF FARLEY STEAM GENERATOR TRANSIENTS (MODEL 51)
PRIMARY SIDE TRANSIENTS GROUP #
DESCRIPTION OCCURRENCES
- 1 Heatup/Cooldown 200 2
Unit Loading /Unicading 18300 3
Reactor Trip 490 Turbine Roll Loss of Flow 4
Loss of Load 120 Loss of Power 5
Large Step Decrease a200 Small Step Inc./Dec.
6 Hot Standby 183006 7
SS Fluctuation 3.187 x 10 Boron Cone. Equalization 8
OBE 400 9
Primary Hydro 5
10 Primary Leak Test 50 SECONDARYSIDETRANSIENTS GROUP #
TRANSIENT CYCLES FOR THIS TOTAL CYCLES TRANSIENT IN GROUP 1
Heatup and Cooldown 200 210 Turbine Roll Test 10 2
Plans ioading and Unioading 18300 18300 3
Small Step Load increase 2000 4000 Small Step Load Decrease 2000 4
Loss of Power 40 800 Large Step Load Decrease 200 Loss of Load 80 Loss of Flow 80 Rector Trip 400 5
Hot Standby Operation 18300 18300 6
Loss of Power - 32*F Cold Water 1
1 into Hot Dry Etiipty Steam Generator 7
OBE 50 50 6
6 8
Steady State Fluctuations 3.15 x 10 3.15x10 9
Secondary Hydrotest 5
5
- 0ccurrences indicates number for each transient group.
For example the reactor trip group includes cycles for both turbine roll and loss of flow, since the reactor trip umbrellas the other two.
. me.nn.e io 2-15
=
9
- h TABLE 2-3 l
SUMMARY
OF PRESSUR12ER TRANSIENTS FARLEY UNITS 1 AND 2 t
GROUP #
TRANSIENT NO. CYCLES 1
Heatup/Cooldown 200 l
2 Heatup 1/Cooldown 6 400
[
3 Heatup 2/Cooldown 5 400 j
' I 4
Heatup 3/Cooldown 4 400 5
Heatup 4/Cooldown 3 400 l
6 Heatup 5/Cooldown 2 400 7
Heatup 6/Cooldo*
400 8
Cooldown 7 1200 9
Unit Load /llnload 36600-t 10 Group #1 4280 Loss of Flow Large Step Load Decr.
Small Step Load incr.
Small Step Load Decr.
11 Loss of Load 120 Loss of Power 12 Reattor Trip 400 i
13 Boron Concentration 36600 14
. Inadvertent Auxiliary Spray 10 15 Primary Hydro 5
16 Primary Leak 60 Turbine Roll 17 OBE 50 l
wi.n o.n io 2-16 g
TABLE 2-4
SUMMARY
OF PRIMARY SYSTEM TRANSIENTS - PRIMARY COOLANT SYSTEM TRANSIENT NUMBER TRANSIENT IDENTIFICATION CYCLES 1
Partial loss of flow
'80 2
Inadvertant S.I. Actuation 60 3
Heatup 200 4
Cooldown 200 5
Unit loading at 5% per minute 18300 6
Unit unloading at 5% per minute 18300-7 10% step load increase 2000 8
10% step load decrease 2000 9
Large step load decrease with steam dump 200 10 feedwater cycling at hot shutdown 2000 11 Loop out-of-service, normal loop startup 70 12 Loop out-of-service, normal-loop shutdown 80 4
13 Reactor trip fro'n full power -- no cooldown 400 14 Reactor trip from full power -- cooldown, no S.I..
480 15 Reactor trip from full power -- cooldown and S.I.
10 16 Inadvertant startup of an inactive loop --
inactive loop 10 17 Inadvertant'startup of an inactive loop --
active loop 10 l
18 SmallLOCA(E/F) 5 19 Small steam line break-(E/F) 5 l
i 20 Complete. loss of flow (E/F) 5 21 Turbine roll test'-- heat up 20' mwm..no 2-17 i
i
TABLE 2-4(cont.)
SUMMARY
OF PRIMARY SYSTEM TRANSIENTS - PRIMARY COOLANT SYSTEM 1
NUMBER TRANSIENT IDENTIFICATION CYCLES i
22 Turbine roll test -- cooldown 20 23 Loss of load 80 24 Control rod drop 80 f
4 25 Loss of power
-40 26
-Inadvertant RCS depressurization
- 10-6
-27 Steady stato fluctuation 10 28 Operating basis earthquake 50 I
'i j
i i
i
- a
-]
me.n m.. io 18
~
TABLE 2-5 CHEMISTRY AND PROPERTIES OF JOSEPH FARLEY UNIT 1 REACTOR VESSEL MATERIALS T
RT Upper Shelf Energy Material Cu P
Ni NDT NOT ICI Component Code No.
Type
(%)
(%)
(T)
(*F)
(*F)
NMWD(d)
MWO Closure head dome B6901 A533,B,C1.1 0.16 0.009 0.50
-30
-20[a]
140 Closure head segment B6902-1 A533,B,C1.1 0.17 0.007 0.52
-20
-20[a]
138 Closure head flange B6915-1 A508, C1.2 0.10 0.012 0.64 60[a]
60[a]
75[a]
Vessel flange B6913-1 A508, C1.2 0.17 0.011 0.69 60[a]
60[a]
106[a]
Inlet nozzle 86917-1 A508, C1.2 0.010 0.83 60[a]
60[a]
110 Inlet nozzle B6917-2 A508, C1.2 0.008 0.80 60[a]
60[a]
80 Inlet nozzle B6917-3 A508, C1.2 0.008 0.87 60[a]
60[a]
98 Outlet nozzle B6916-1 A508, C1.2 0.007 0.77 60[a]
60[a]
96.5 Outlet nozzle B6916-2 A508, C1.2 0.011 0.78 60[a]
60[a]
97.5 Outlet nozzle B6916-3 A508, C1.2 0.009- 0.78 60[a]
60[a]
100 Upper shell B6914-1 A508, C1.2 0.010 0.68 30 30[a]
148 Inter. shell B6903-2 A533,B,C1.1 0.13 0.011 0.60 0
0 97 151.5 Inter. shell B6903-3 A533,B,C1.1 0.12 0.014 0.56 10 10 100 134.5 "f
Lower shell B6919-1 A533,B,C1.1 0.14 0.015- 0.55 15 90.5 133 E$
Lower shell B6919-2 A533,B.C1.1 0.14 0.015 0.56
-10 5
97 134 Bottom head ring B6912-1
.A508 C1.2 0.010 0.72 10 10[a]
163.5 Bottom head segment 86906-1 A533,B,C1.1 0.15 0.011 0.52
-30
-30[a]
147 Bottom head dome B6907-1 A533,B,C1.1 0.17 0.014 0.60
-30
-30[a]
143.5-Inter. shell long.
M1.33 Sub Arc Weld 0.25 0.017 0.21 ia]
0[a]
weld seam Inter. to lower Gl.18 Sub Arc Weld 0.22 0.011 <0.20[b]
0[a]
0[a]
shell weld seams Lower shell long.
G1.18 Sub Arc Weld 0.17 10.022 <0.20[bl 0[a]
0[a]
weld seams
[a] Estimate per'NUREG-0800 "USNRC Standard Review Plan" Branch Technical Position MTEB 5-2. [11]
[b] Estimated (low' nickel _ weld wire used in fabricating vessel weld seams).
[c] Major. working direction.
[d] Normal to' major working direction.
3818s/l t l40910
TABLE 2-6 CHEMISTRY AND PROPERTIES OF JOSEPH FARLEY UNIT 2 R ACTOR VESSEL MATERIALS Average Upper Shelf Energy-Normal to-Principal Principal W rking Working T
RT Cu P
.Ni NDT NDT Oirection Direciion Component Code No.
Grade
(%)
. {%)
(%)
(*F)
(*F)
(ft-lb)
(ft-lb)
.CL. HD. Dome B7215-1 A533,B,CL.1
- 0. D 0.010 0.49
-30 16(a) 83(a) 128 CL. HD. Flange B7207-1 A508,CL.2 0.14 0.011 0.65 60(a) 60(a)
>56(a)
>86(c)-
Vessel Flange B7206-1 A508,CL.2 0.10 0.012 0.67 60(a) 60(a)
>71(a)
>109 Inlet Noz.
B7218-2 A508,CL.2 0.010 0.68 50(a) 50(a) 103(a) 158 Inlet Noz.
B7218-1 A508,CL.2 0.010 0.71 32(a):
32(a) 112(a) 172 Inlet Noz.
B7218-3 A508,CL.2 0.010 0.72-60(a) 60(a) 98(a).
150 Outlet Noz...
B7217-1 A508,CL.2 0.010 0.73 60(a)'
60(a) 100(a) 154 0.010 0.72 6(a) 6(a) 108(a) 167 Outlet Noz.
B7217-2 A508,CL.2 Outlet Noz.
B7217-3 A508,CL.2 0.010 0.72 48(a) 48(a) 103(a) 158 7
Upper Shell B7216-1 A508,CL.2 0.010 0.73 30 30(a) 97(a) 149 5
Inter Shell B7203-1 A533,B,CL.1 0.14 0.010 0.60
-40 15 99 140 Inter Shell B7212-1 A533,B,CL.1 0.20 0.018 0.60 -10 99 134 Lower Shell
~B7210-1 A533,B,CL.1 0.13 0.010 0.56
-40 18 103 128 Lower Shell B7210-2 A533,B,CL.1 0.14 0.015 0.57
-30 0
99 1,5 Boi. tom Head Ring-B7208-1 A508,CL.2 0.010 0.73 40 40(a) 89(a) 137 Bottom Head Dome
_B7214-1 A533,B,CL.1 0.11 0.007: 0.48
-30
-2(a) 37(a) 134 Inter. Shell A1.46 SMAW 0.02 ' O.009 0.96 0(a)
O(a)
>131 Long Seams A1.40 SMAW 0.02 'O.010 0.93
-60
-60
>106 Inter Shell to Lower Shell-
.G1.50 SAW 0.13 0.016 <.20(b) -40
-40
>102 Lower Shell Long Seams' G1.39.
SAW:
0,05 0.006 <.20(b) -70
-70
>126 (a) Estimate per NUREG 0800 "USNRC Standard Review' Plan"' Branch Technical Position MTEB 5-2. [11]1 (b) Estimated.
(c). Upper shelf not'available, value represents minimum energy at the highest test temperature.
3818s/111489-10
TABLE 2-7 CHEMISTRY FACTOR FOR WELDS, 'F l
0 0.20 0.40 0.60 0.80 1.00 1.20 20 20 20 20 20 20 --
20 0.01 20 20 20 20 20 20 20.
y 0.02 21 26 27 27 27 27 0.03.
22 35 41 41-41 41 41 1
0.04 24 43 54 55 54 54 54' 0.05 26 49 67 68 68 68 68 0.06 29 52 77 82 82 82 82 O.07 32 55 85 95 95-95 95 0.08
,36 58-90 106 108 108 108-0.09 40 61 94 115
'122-122 122 0.10 44 65 97 122 133 135 135 b
0.11 49 68 101 130 144 148 148 0.12 52
.72 103 135 153 161 161 j
0.13 58 76 106 139
'162 172 176 O.14 61 79 109 142 168 182 188 0.15 66 84 112 146 175 191 200 0.16 70 88 115-149 178 199 211 0.17 75 92 119 151 184 207 221 0.18 79 95 122 154 187 214 230 0.19 83 100 126 157 191 220
-238 0.20 88 104 129 160 194 223 245 0.21 92 108 133 164 197 229-252 0.22 97 112 137 167 200 232 257 0.23 101 117 140 169 203 236 263 0.24 105 121 144 173 206 236 268 0.25 110 126 148 176 209 243 272 0.26 113 130 151 180 212 246 276 0.27 119 134 155 184-216 249 280 0.28 122 138 160 187 218 251 284 O.29 128-142 164 191 222 254.
287.
0.30 131 145 167 194 225
.257 290 0.31 136 151 172
.198 228 260 293.
0.32 18.0 155 175 202-231 263 296 0.33 144 160 180 205 231 266 299 0.34 149 164 184 209 238 269 302 n
0.35 153 168 187 212 241 272 305 0.36 158 172 191 216 245 275 308 0.37 162 177 196 220 248 278 311 0.38 166 182 200 223 250-281 314 0.39 171 185 203 227 254 285 317 0.40 175 189 207 231 257 288 320 mwman io 2 _
1
TABLE 2-8 CHEMISTRY FACTOR FOR BASE METAL, 'F
Wt-%
0 0.20 0.40 0.60 0.80 1.00 1.20 0
20 20 20 20 20 20 20 0.01 20 20 20 20 20 20 20 0.02 20 20 20 20 20 20 20 0.03 20 20 20 20 20 20-20 0.04 22 26 26 26 26 26 26 0.05 25 31 31 31 31 31 31 0.06 28 37 37 37 37 37 37 0.07 31 43 44 44 44 44 44 0.08 34 48 51 51 51 51 51 0.09 37 53 58 58 58 58 58 0.10 41 SB 65 65 67 67 67 0.11 45 62 72 74 77 77 77 0.12 49 67 79 83-86-86 86 0.13 53 71 85 91 96 96 96 0.14 57 75 91 100 105 106 106 0.15 61 80 99 110 115 117 117 0.16 65 84 104 118 123 125 125 0.17 69 88 110 127 132 135 135 0.18 73 92 115 134 141 144 144 0.19 78 97 120 142 150 154 154 0.20 82 102 125 149 159 164 165 0.21 86 107 129 155 157 172 174 0.22 91 112 134 161 176 181 184 0.23 95 117 138 167 183 190 194 0.24 100 121 143 172 191
~199 204 0.25 104 126 148 176 199 208 214 l
0.26 109 130 151 180 205 216 221 0.27 114 134 155 184 211 22S 230 0.28 119 138 160 187 218 233 239 0.29 124 142 164 191 221 241-248 0.30 129 146 167 194 225 249 257 0.31 134 151 172-198 228 255 266 0.32 139 155-175 202 231
~260 274 0.33' 144 160 180 205 234 264 282 1
0.34 149 164 184 209 238 268-290 r
0.35 153 168 187 212 241 272 298 0.36 158 173 191 216 245-275 303 i
0.37 162 177 196 220 248 278 308 l
0.38 166 182 200 223 250 281 313 0.39 171 185 203 227 254 285 317 0.40 175 189 207 231 257 288 320 2
C!AD M IRtwitAL stAM5 VSTIMLlLAfB 19-8948 86903 3 l
t 10 494 j.
g 45' I
l 8.4"
- ?
CORE' g
F CORE 1
=
A 1,44.0*
f T.
R6903 2 19-894A E
1 u..
C t _20.1" q
gg,g,4 20-8948 86919 2
?
E
'45 I
CORE' p
l o
'48.75" C
20-894A 86919-1 Figure 2-1.
Identification and Location of Beltline Region Material for the Farley Unit 1 Reactor Vessel wswomao g_g3
l.
i tfRCUMPtatWTIAL SD M5 IRIIMLIDlB 19-9238 97212-1 l
e
^
"eig
- 10 923 9
8.4' 45 E
g CQRt
\\
=
2 I."'E A
B7203-l' 19-923A
]
- m 5
ll e at..
g 20.1" 4
,11-923 g
B7210-2 20-923B 1
user,-
- 45 3,
=
e i
a 48.79*
20-023A B7210-1 Figure 2-2.
Identification and Location-of Materials for the Farley Unit 2 Reactor Vessel wimmoimio -
2-24
j gi~'
,y.
l'y {j.1. l ygtg.g,] g g. t
.t' j e g] '-
y' lg.4 g
~
15-
!,1,;"U0:,;e !..n,,mik h EWWg!-.g j
.ER-) i-
"Ii!WHIMC junilUnm e
.1 c
l"h"s n 11 'j!""M"i v -l iim n,r.
n i l "N"ff"k'umnirmtn""2 m
o illnlir.itettutrenucc -l'iinimili p T 1 l
"w
,e estippliinstit+.
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Example of Critical Flaw Size Determination mie. moue,o 2-26 l
SECTION 3 FATIGUE CRACK GROWTH i
In applying Code acceptance criteria as introduced in section 1, the final i
flaw size a used in criteria (1) is defined as the minimum flaw size to f
which the detected flaw is calculated to grow at the end of a specified period, or until the next inspection time.
In this handbook, ten, twenty-i and thirty year inspection periods are assumed.
These crack growth calculations have been carried out for all the regions in the Farley crimar, system components for which evaluation charts have been constructed.
This section will examine each of the calculations, and provide' the methodology used as well as the assumptions.
l 3.1 ANALYSIS METHODOLOGY l
The methods used in the crack growt'h analysis reported here are the same es those suggested by Section XI of the ASME Code.
The analysis. procedure involves postulating an initial flaw at specific regions and predicting the growth of that flaw due to an imposed series of loading transients.
The input required for a fatigue crack growth analysis-is basically the information necessary to calculate the parameter AKj which depends on crack and structure geometry and the range of applied stresses in the area where the crack exists. Once AK; is calculated, the growth due to that particular stress cycle can be calculated by equations given in section 3.3 and figure 3-1.
This increment of growth is then added to the original crack size, and the analysis proceeds to the next transient.
The procedure is continued in this manner until all the transients known to occur in the period of evaluation have been analyzed.
The transients considered in the analysis are all the design transients contained in the vessel equipment' specification, as shown in section 2.
These transients are spread equally over the-design lifetime of the vessel, with the exception that the preoperational teste are considered first.
Emergency and faulted conditions are not considered because their frequency of occurrence is too low to affect fatigue crack growth, mwnme io 31
Crack growth calculations were carried out for a range of flaw' depths, and three basic types.
The.first two were surface flaws, one with length equal to six times the depth and another with length equal twice the depth.
Next was a continuous' surf ace flaw, which represents a worst case for surface flaws, and the last type was an embedded flaw, with a range of. shapes and locations.
For all cases the flaw was assumed to maintain a constant shape as it grew.
3.2 STRESS INTENSITY FACTOR EXPRESSIONS-Stress intensity factors were calculated.from methods available in the literature for each of the flaw types analyzed.
The surface flaw with aspect ratio 6:1 was analyzed using an expression developed by.(
)*'C where the stress intensity factor K is calculated from the actual l
stress profile through the wall at the location of interest.
The maximum and minimum stress profiles corresponding to each transient are represented by a third order polynomial, such that:
2 3
e (X) = A0+A3 [+ A2 3
+A
.(3-1)
The stress intensity factor K (e) can be ce Nulated anywhere.along the y
crack front.
The point of maximum crack depth is represented by e =.0.
The-following expression is used for calculating K; (o).
a,c,e (3-2)
The magnification factors H I*)' "1(*)' "2(t) and H (9) are obtained by the 0
3 procedure outlined in reference (5).
l nu,nune io 32
q l
The stress intensity-factor for a continuous surface flaw was calculated using an expression for [
-Ja,c.e The stress distribution is linearized through the wall thickness to determine rnembrane and bending stress and the applied K is calculated from:
j j.c.e (3 3) a The magnification f actors 'Y, 'and YB are taken from (19) and a is the crack f
depth.
For embedded flaws, the stress intensity._ factor expression of (
Ja,c.e was used, as discussed earlier in section 2.
The flaw I
shape and the eccentricity were both varied.
The calculated crack growth was very small for these cases.
3.3 CRACK GROWTH RATE REFERENCE CURVES -- FERRITIC STEEL The crack growth rate curves used in the analyses were taken directly from Appendix A of Section XI of the ASME Code. Water environment curves were used for all inside surface flaws, and the air environment curve was used for embedded flaws and outside surface flaws.
The materials used for the pressure boundary of steam generators for Farley-units are basically higher strength versions of the reactor vessel steels, SA508 Class 2 and 3 and SA533 Gr B Class 1, and early designs had exactly the same materials as the reactor vessel.
A large number of specimens of the higher strength steam generator materials, SA508 Cl. 2a, SA553 Gr. A Cl 2, SA508 C1. 3a, and SA533 Gr. B Cl. 2 materials and two associated submerged are weldments were tested at Westinghouse (20).
The environments used were low and high temperature air, PWR primary water, and secondary side steam. These environments cover all the possible environments for both the primary and secondary side of the steam generator, but do not include any contaminants which could be present in the secondary side environment.
Load ratios of 0.2 and 0.7 were employed for the air environment, and values of 0.2 and 0.5 were used in the PWR and steam environments.
wwnm io 3-3
Results showed that the reference crack growth rate curves for ferritic steel contained in Section XI were also applicable to these steels.
The PWR l
environment was found to produce the highest growth rates, but the data were well below the ASME reference curves.
The data obtained in the steam environment showed crack growth rates equal to or below the rates obtained in-the PWR environment under the same conditions.
These results are discussed in.
reference 20. Therefore the ASME Code reference curves are applicable.
For water environments the reference crack growth curves are shown in Fig.
3-1, and growth rate is a function of both the applied stress intensity factor range, and the R ratio (Kmin/Kmax) f r the transient.
For R<0.25 (aKi <19 ksi / En)df = (1.02 x 16 ) 3g{5.95 (34) 6 (aK; >19 ksi / in)h = (1,01 x 10-3) 3g}1.95 whereh=CrackGrowthrate, micro-inches / cycle.
For R>0.65
( AK; <12 ksi / in)dj = (1.20 x 10-5) 3g 5.95 (3-5)
(aK; >12 ksi / in)d
=(2.52x10'1)'aK;1.95 For R ratio between these two extremes, interpolation is_ recommended.
The crack growth rate reference curve for air environments is a-single curve, with growth rate being only a function of applied aK.
This reference curve is also shown in figure 3-1.
h=(0.0267x10~3) aK 3.726 (3-6) 3 mwman io 3-4 i
where,h=Crackgrowthrate, micro-inches / cycle 1
f AK; = stress intensity facter range, ksiilin 1
1
= (K
~Kimin) imax 3.4 CRACK GROWiH RATE REFERENCE CVRVES - STAINLESS STEEL The reference crack growth law used for_the stainless steel portions of the system was taken from that developed by the Metal Propertie Council -
Pressure Vessel Research Committee Task. Force in Crack Propagation Technology.
The reference curve has the equation:
k=CFSAK~
(3-7)
D whereh=crackgrowth_ rate,'in:hespercycle l
materialcoefficient(C='2.0x10-19)
C
=
frequency coefficient for loadings (F = 2.0)
~
l F
=
R ratio correction coefficient (S = 1.0 = 0.502 R ) 4.0 2-l S
=
material property slope (=3.0321) n
=
AK stress intensity factor range, psi / in This equation appears in Section XI, Appendix C-(1989 Addendum) for air-environments and its basis is provided in reference (21-), and shown in figure 3-2.
For water environments, an environmental factor of 2 was-used, based on the crack growth tests in PWR environments reported by Bamford (22).
3.5 RESIDUAL STRESSES Since the piping welds were not stress relieved, residual stresses are clearly
~
present.
For fatigue crack growth analyses, these stresses were included directly, using as a guide the residual stress values in the technical basis document for the ASME Code flaw evaluation procedures (23).
me,n nwo 3-5
Although there is significant residuai stress variation from one weldment-to another, there have been a large number of measurements made, and these were collected in reference (23).
The resulting guidelines are summarized in figure 3-3.
These stresses were used directly in the analysis, except in a few instances where the elastically calculated stresses were-significantly above yield stress, in these cases it would be unrealistic to add the residual stresses, so they were not used, it should be noted here that the residual stresses were added to both the maximum and minimum stresses, and therefore do not affect the stress range.
Their effect is seen only through the R ratio, as illustrated in figure 3-2, 3.6 STRESS CORROSION CRACKING SUSCEPTIBILITY In evaluating flaws, all mechanisms of suberitical crack growth must be evaluated to ensure that proper safety margins are maintained during service.
Stress corrosion cracking has been observed to occur in stainless steel in operating BWR piping systems.
The discussion presented here is the technical basis for not considering this mechanism in the present analysis.
For all Westinghouse plants, there is no history of cracking failure in the reactor coolant system loop piping.
For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously:
high tensile stresses, a susceptible material, and a corrosive environment.
Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized I
by proper selection of a material immune to SCC as well as preventing the occurrence of a corrosive environment.
The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other materials.in the system, applicable ASME Code rules, fracture toughness, welding, fabrication, and processing.-
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Reference Fatigue Crack Growth Curves for Carbon and Low Alloy Ferritic Steels l
mi..mna io 37 l
^
a
i l
l l
l TIwough WallMesidualStress '
Well Thickness t
Antal
. Circumferentfal
'S S
<1 inch 0.5 ~
0 0
I7 I
I u
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'8 Considerable variation with weld heat input.
3 g eg [1.0 - 6.91 (a/g + 8.69 (a/Q'- 0.48 (a/4 - 2.03 (a/tf]
8 L
ei = stress at inner surface (a = 0) 1 Figure 3-2.
Reference Crack Growth Rate Curves for Stainless Steel-in Air Environments [23) aie.iorme io 3-8
9 l
3.0x10~4 g gy f,
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J 10~4 SOLID LINES FOR 70 F.
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Recommended Axial and Circumferential Residual Stress Distributions for Austenitic Stainless Steel Pipe Welds (from reference 25) nie.i-io 3-9
=
l l
SECTION 4 DETERMINATION OF LIMITING TRANSIENTS u
4.1 INTRODUCTION
The two key parameters used.in the evaluation of'any. indications discovered during inservice inspection are:
first,ithe critical flaw depth for the governing normal, upset, and test conditions,'and second, that for-the governing emergency.and faulted conditions.
The selection =of the governing transient for normal, upset, and test conditions was done based on the highest surface stress for each location for which a chart was to be constructed.
For emergency and faulted conditions, this choice was.not as straightforward, as a result of developments on the pressurized thermal shock issue.
This. issue has resulted in a great deal of study of various transients which could occur in operating. plants, including consideration of the overall frequency of.each transient.in addition to.its severity.
The following section will provide a summary of the generic work performed for:
PTS, along with a detailed comparison of the various emergency and faulted transients that are possible in the beltline region'of the Joseph Farley Unit 1 and 2 reactor vessels.
4.2 SELECTION OF GOVERNING EMERGENCY AND FAULTED TRANSIENTS 1
4.
2.1 BACKGROUND
AND HISTORY The issue of reactor vessel pressurized thermal shock (PTS) has focused significant attention to the evaluation of the vessel beltline location.
Until early 1982 reactor, vessel integrity was evaluated for PTS events, which.
generally fall into the category of emergency and. faulted conditions, usually using only design basis transient scenarios.
For instance, a summary report on reactor vessel integrity for Westinghouse plants, WCAP-10019 [24), was w e. m ussso 4-1
submitted to the NRC staff in December 1981 and addressed the large LOCA and large steamline break' transients along with a conservative evaluation of the small break LOCA and small steamline break events. The Joseph Farley Units 1 and 2 reactor vessels were evaluated as part of this generic evaluation supported by the Westinghouse Owners Group.
Following the submittal of this information, the NRC was concerned, as a result of recent plant operating events, that other more likely events with dominating transient characteristics were not being addressed.
I To respond to the above concern, an innovative methodology was developed that l
coupled probabilistic event sequence analysis results with thermal hydraulic and fracture mechanics analysis results to identify all potential transient scenarios of comern for reactor vessel PTS.
This methodology efficiently evaluated over 8,000 possible transient scenarios on a generic basis and the results demonstrated adequate safety margin for the Westinghouse domestic operating plants.
This work, which was submitted to the NRC via the Westinghouse Owners Group (WOG) in References [25, 26, 27) was extensively used by the NRC Staff in the development and improvement of their own position on PTS.
The NRC used the Westinghouse probabilistic results to better quantify total plant risk from PTS and to support their licensing position as described in NRC Policy Issue SECY-82-465, November 1982 (28).
(This document provides the technical basis for the PTS Rule [29) that was issued in 1985.)
A key aspect of this work is that the principal contributors (dominating transients) to the total frequency of significant flaw extension in the vessel from PTS can be identified.
However, this work was done in'an approximate generic manner and both the Westinghouse Owners Group and the NRC agreed that more work should be done to investigate additional. candidate transient sequences and characterizations and to validate some of the approximations made in the supporting analyses.- For instance, the 2"-6" small LOCA results used detailed calculations of system response (including fluid mixing effects in the cold leg and vessel downcomer as predicted from cxperimental results, heat input from hot piping walls, and assumed benefits from the effect of warm prestressing) whereas the extended high pressure injection category (i.e.,
events that could lead to extended high pressure safety injection operation l
asiwimas ic 4-2
with stagnated loop (s)) used very conservative transient characterizations.
This approach lead to a conservative assessment of the total frequency of significant flaw extension.
4.2.2 PTS Risk for a Typical Westinghouse PWR In order to address'all candidate transient scenarios in a thorough manner, the Westinghouse Owners Group (WOG) undertook a Stagnant loop Code Evaluation Program in late 1982.
One key purpose of this program was to~ demonstrate that the overall risk from PTS on a typical Westinghouse plant is dominated by l
small steamline breaks, small LOCA's, and steam generator tube ruptures, as suggested-in previous WOG work during 1982, and not by other transient scenarios, including those involving loop stagnation. WCAP-10319(30]
presents the results of this exhaustive study.
The important results and the relationship to previous fracture analyses performed for'the Joseph Farley Units 1 and 2 reactor vessels are discussed below.
The event sequence analysis performed in the WOG Stagnant Loop Code. Evaluation resulted in the following broad categories of events that could potentially result in a pressurized thermal shock of the reactor vessel:
1.
Secondary Depressurization (SD) 2.
Loss of Coolant Accident (LOCA) 3.
Steam Generator Tube Rupture (SGTR) 4.
Loss of Secondary Heat Sink (LOHS) 5.
Excessive feedwater (EXFW) 6.
Anticipated Transients Without SCRAM (ATWS) 7.
FeedlineBreak(FB)
Combinations of these categories were also considered if they met certain criteria defined in WCAP-10319 (30). Some of these PTS-categories were further subdivided into a number of small-bins to offer greater resolution and accuracy in the risk assessment and.in the identification of the dominating transient scenarios, m.,n n asio 4-3
m The summary results of the above WOG risk assessment for PTS (see Figure 4-1) showed that the key contributors to the total risk occur from the LOCA and SGTR categrries because of the combination of severe transient characteristics with relatively high frequencies of transient occurrence.
The LOHS transient, while much lower than LOCA or SGTR, was the third most dominating transient in terms of contributing to the total PTS risk.
This is primarily because LOCA transient cheracteristics were conservatively used for the LOHS analysis.
If the true LOHS transient results had been used, it is believed that the resulting transient characteristics would be less severe than those that were used.
The other PTS transient scenarios,. including those involving loop stagnation (i e., 50, EXFW, ATWS, and FB), do not contribute significantly to the overall risk.
The ASME Code in its present f *.rm, however, does not take transient frequencies into consideration and requires an evaluation of flaw indications using the most limiting emergency / faulted condition transient.
Therefore, the above PTS risk analysis results could not be used directly, but they were used to guide the determination of the key transients to be considered further, as will be seen in the next section.
4.2.3 Treatment of Transient Severity Probabilistic fracture mechanics (PFM) results, used in the above WOG risk assessment for PTS, were utilized to evaluate the severity of the transients used in the generic study that were major contributors to the risk of vessel failure.
Figure 4-2 shows an example of PFM results that quantify the conditional probability of reactor vessel failure (i.
e., significant flaw extension) given that a PTS event occurs.
The results shown in figure 4-2 were based upon the evaluation of stylized exponential cooldown transients characterized by three quantities: a final temperature (T ) reflecting the depth of the f
cooldown; a time constant (B) reflecting the rate of the cooldown; and a characteristic pressure (P) as described in figure 4-3.
The curves in figure 4-2 were generated from PFM analyses using the Monte Carlo technique. A mwmmo 44
matrix of cases for given T, B, and inner surface RTNDT values were f
evaluated to obtain results for generation of the curves. The RT values NDT are calculated as a function of initial RTNDT, material residual elements and fluence using the methodology discussed in Section 2.
For each case, a 6
large number of deterministic fracture mechanics analysis trials (~10 )
were simulated using random values selected by a random generator from distributions defined for the pertinent input properties. The input properties that have been treated as random variables include:
initial crack depth, initial RTNDT, copper content, fluence, and the critical stress intensity values for flaw initiation and. arrest.
The probability of vessel failure for each case was determined by dividing uie number of failures by the number of trials. The curves in Figure 4-2 were plotted from the matrix of f r assumed longitudinally against RTNDT results by normalizing Tf criented flaws.
The pertinent aspect of the PFM results for determining the governing transient (s) is that, for a given inner surface RTNDT value, the higher the conditional probability of vessel failure, the more limiting the transient.
Using the stylized transient characteristics for the WOG generic transients within all of the various transient categories (30), the most limiting 3
transients were determined from the WOG PFM results as shown in Table 4-1.
The transients are shewn in order of decreasing severity. The associated transient frequencies of occurrence are also given for the purpose of information.
-2 The conditional probability of failure values ranged from 1 x 10 to
-2 5 x 10 for the above transients at an inner. surface RT value which is NDT value for the Joseph Farley 1 near the projected end-of-life (32 EFPY) RTNDT and 2 reactor vessels (see Section 2).
For all other transient events, the
-2 conditional probability of failure values were much less than 1 x 10 From the standpoint of statistics, however, the conditional probability of failure values were essentially the same for the above limiting transients, and any one of them could be the " governing" event.
The fact that stylized transient characteristics were used in the
<aluation rather than the actual transient histories lends further support to the above statement, wwmne to 4.s a
Although the large LOCA and large steamline break (LSB) events are not significant contributors to the overall risk of failure because the frequency of occurrence for these events is negligible (~1_x 10 /r yr), the severity of these events still needs to be considered in the. selection of the most 1 iting event for the flaw handbook.
The plant specific results for these events from prior Joseph Farley analyses are considered as shown in the next section.
Therefore, we see that the large number of-thermal shock and pressurized thermal shock transients (>8000)-can be reduced to a list of a few key transients, as shown in Table 4-1.
Fracture analysis was then concentrated on these transients, as discussed in the following section.
4.2.4 Emergency and Faulted Conditions Evaluation -- Beltline. Region To determine the governing emergency and faulted conditions for the Joseph Farley reactor vessels, a series of' transients were studied.
These transients
~
included the large LOCA~and large steamline break (LSB) and the dominating transients from the Westinghouse Owners Group pressurized thermal shock studies.
This work, which took into account the differences in plant system characteristics between Joseph Farley and the typical plant in the generic WOG evaluation, led to the conclusion that the following transients should be coasidered in the deterministic assessments for the beltline regions to be used for this handbook.
l o
steam generator tube rupture (SGTR) o small LOCA o
large LOCA o
large steamline break (LSB)
The transient frequencies for these limiting events are also given in the table in Section 4.2.3.
i 3618:1112789 10 g
l Thermal stress and fracture analyses we-e performed for the beltline region, utilizing the characteristics of the above four transients these are repre-sented in the form of Figure 4-3.
The limiting circumferential weld and the limiting longitudinal weld for both units were used in performing the fracture analyses.
The resulting critical flaw depths for a range of shapes are shown in Table 4-2.
From this table it may be seen that the large steamline break transient evaluated previously is the governing transient for the beltline region.
The detailed assessments performed for the tube rupture and small LOCA transients-servn to verify this conclusion.
Also, from the standpoint of total risk it is worthy of note that these latter two transients are the dominant ones.
Section XI of the ASME Code presently requires that only the most severe transient be evaluated, regardless of its probability of occurrence, so the large steamline break is the governing transient for the handbook.
4.2.5 Faulted Conditions Evaluation For Other Regions A number of analyses were performed by means of linear elastic fracture mechanics methods to determine the postulated minimum critical flaw size at which unstable flaw growth could occur in the Joseph Farley Units 1 and 2 reactor vessel beltline regions.
The critical flaw size required for unstable flaw growth was determined from the intersection of the K; curve with the K
curve, as described in Section 2.
lc The conclusions reached as to the governing transients for the beltline region will not necessarily be applicable to the other regions, because the fracture toughness is not reduced from irradiation.
The conditions which could lead to fracture in these other regions will be governed primarily by pressure stresses, while the conditions for the beltline regions are governed by thermal stresses.
This conclusion is even more true for regions of stress discontinuity, where most of the welds,are found.
For this reason the severe thermal transient with the largest pressurization level was found to be generally the governing transient, i.e., the large steamline break (LSB).
Although not true in general for all plants, this is the same transient found to be governing for the beltline region.
3318:1112789 10 47
'4.2.6 Treatment of Low Temperature Overpressurization Transients (LTOP)
In this section, the frequency of occurrence of a significant low temperature overpressurization (LTOP) challenge to'the Joseph M. Farley Plant is calculated.
For this calculation, a (
ja.c,e,
l 4.2.6.1 Model Assumptions The PRA model is constructed on the basis of the following assumptions:
1.
The LTOP event is of concern if the primary coolant temperature is less than (
Ja,c,e degrees F.
i 2.
There are written operational procedures for overpressurization 5
mitigation.
The operators are trained in' mitigating the LTOP event..
(
3.
During plant cooldown prior to reducing reactor coolant = system.
temperature below the minimum temperature allowable for the inservice l
pressure test, valves 8702A/B and 8701A/B are opened to allow any
?
excess pressure to be vented through the Residual Heat Removal System Relief Valves (RHRSRV) 3 4.
During cooldown, the RHRSRV are placed in service between 350 and 310 degrees F. while a steam bubble remains in the pressurizer.
After RHR i
is placed. inservice at less than [.
']a,c.e degrees F.,'the-reactor i
l coolant system is cooled and taken solid after reaching 160 degrees F.
s i
5.
Normal operating procedures maximize the use of a pressurizer cushion (steam bubble) during periods of-low temperature-operation.
A steam i:
bubble is formed in the'pressuri'.er at a cold leg temperature of i
i sais,*imse ic 4.g
]
approximately 160*F when thi plant is being started up.
it is collapsed at a cold leg temperature of approximately.160'F when the plant is being cooled down.
This cushion substantially reduces the' severity of some potential transien,ts such as RCP-induced heat input and slows the rate of pressure rise for others.
This provides reasonable assurance that most potential transients can be terminatsd by operator action before an overpressure condition exists.
6.
Additional limitations placed upon plant operations inclu'de:
a) When the reactor coolant system is not open to_ the atmosphere 'and the temperature of one or-b;
- e. tor coolant system cold legs is less than or equal to (
Ja.c." degrees F., no'more than one' high pressure safety injection (SI) pump. shall be operable.
The second SI pump breaker shall be racked out, b) A reactor coolant pump (RCP) shall not be started,when the reactor coolant system temperature is less than-the minimum temperature for the inservice pressure test unless:
l 1.
There is a pressure absorbing volume in the pressurizer, or i
j 2.
The secondary water temperature of each steam generator is less than (
Ja,c.e degrees F. above the temperature of the reactor coolant system.
.1
-l 7.
Several control room alarms have been'provided. A seismic category 1 alarm designed to the requirements of IEEE-279-1971 alerts the operator if the RHR isolation valves are not fully open when the RC5 temperature is less than or equal to (.
)"'C.
Another alarm provides indication to the operator of any~ overpressure transient occurring when the-RCS pressure exceeds (
fla,c,e j
i I
wwn use to
'49
' 8.
The motor-operated valves ($0Vs). upstream of the safety' valves can close from either a spurious closure, or a falso high pressure input signal which can close both MOV trains.
For'a general surveillance time of these valves, a conservative period of-10 hours will be assumed.
9.
Although the PORVs are available to the operator to mitigate an overpressure event, no credit.will be given to the PORVs'and corresponding operator action (s) to reduce an LTOP event.
j
- 10. The failure rate for a sensor (
3a,c.e includes'all.
failure modes (i.e., incorrect /no signal and spurious signal).
Althoup> a spurious signal may make up aLvery small percent of this-nucec, ' ne conservative nunber of (
')a,c.e will be used i
for a sptcious signal i
4.2.6.2 Event Tree Analysis Event tree analysis is used to model end quantify the progression and frequency of significant LTOP challenges. The event.trc.e considers the following six items in the progression of a significant LTOP challenge:
i 1.
An LTOP precursor challenges the plant safety systems as an initiating event.
This challenge may ta due to (
ja,c.e,
2.
Primary coolant temperature during the challenge is below (
_)a,c.e degrees F.
If not, the LTOP event is not a concern.
I 3.
The RHRSRV system is availab"s.
This system is available-if one or two of the trains are available, i
4.
The pressurizer has bubble formation.
If the bubble exists, operator action is possible even if the RHRSRV system fails.
wwn ua.io 4-10
' 5.
'iigh pressure alarms and valve closure work.
If-they do, the operator can take manual-action. Operator action is only credible if.a bubble is present. Otherwise, the event progresses-too fast for operator response.
6.
Operator, mitigates the event. This is credible if the bubble is present and the alarms work.
I 4.2.6.3 Succese. Criteria i
The success criteria to avoid a significant LTOP event is:-
a) At least one mitigating system train is available.
4 b)
If both trains of the mitigation system fail, then there must be bubble in the primary and the alarm must work and the operator action must be successful.
4.2.6.4 DATA USED Two sets of data were used for the calculation-of the frequency of a significant LTOP event.
The first set was a best estimate calculation.,- the second one'was a conservative estimate calculatior 1
1.
Initiating event frequency. (Event tree node OPC) i Since 1983, there has Loen-(
Ja c.e Thus the best i
estimate challenge frequency is calculated.as
[
ja.c.e -
As a conservative estimate approximately twice the above value will be used:
[
Ja,c.e, wiwm4a io 4 11
'l Primary coolant ter@erature during the challenge is less than 350 degrees F. (Event tree node TMP)-
6 a
The probability of the temperature being in the range of concern is unknown.
It will be taken as (
ja.c.e in the temperature range of concern.
The same value will be used for the conservative estimate.
3.
Overpressure protection is availabie.(Ever,t tr:: r,ede OP$)
The unavailability of each train upon demand will be calculated as follows:
(
ja.c.e
(
ja.c.e
[
3a,c.e ja.c.e
(
ja c.e I
Probability of both trains being unavailable is calculated as follows:
(
3a,c.e Assuming a beta factor of (
.)a,c.e for common cause failure,
(
A*(
3.c.e a
q[
3a,c.e l
3416s/111449 10 gg l
4.
The pressurizer has a bubble formation.(Event tree node WS0)
It will be assumed that there is no bubble formed in (
Ja.c.e percent of the challenges.
This is due to operational practices in which the bubble is formed after heat-up, and primary being taken solid during cooldown.
The same value will also be used for conservative estimate.
5.
Highpressureandvalveclosurealarmswork.(EventtreenodeALR)
[
I i
It will be assumed that the failure of these alarms occurs once in hundred challenges.
This is considered to be conservative.
Operator mitigates the event.(Event tree node OPE)
The operator failure probability to respond to the alarm within a (
Ja.c.e, assuming that there are at least two operators in the control room and the alarm works; and there is a bubble formation in the primary, giving the operators a three to ten minute response time. However, as a conservative estimate and no credit is given for the PORVs, operator failure probability of (
Ja,c.e,g))
be used.
4.2.6.5 Data Summary The following data were used to quantify the event tree for best estimate and conservative estimate scenarios for the frequency of a significant LTOP challenge per calendar year, per plant:
BEST ESTIMATE CONSERVATIVE a,c.e emm >
me.m u.e io 4-13
BEST ESTIMATE CONSERVATIVE
- a,c.e i
4.2.6.6 Calculations and Conclusions The vent trees in tables 1 and 2 are used to calculate the frequency of a significant LTOP event per calendar year per plant on the farley site as follows:
The best estimate frequency is [
Ja.c.e The conservative estimate frequency is (
la.c.e Based on the above results, the frequency of a significant LTOP event is (
ja.c.e, In review of the preliminary results of a study presently in progress,
" Residual Heat Removal System Autoclotnre laterlock Removal. Report for the Joseph M. Farley Nuclear Plant Units 1 and 2." WCAP-11746, which uses
" Residual Heat Removal System Autoclosure Interlock Removal Report for the Westinghouse Owners Group," WCAP-11736 as a basis, it is noted that the results of this study, which is more detailed in showing the various plant states resulting from an LTOP initiatirg event, indicates a frequency of high/
significant overpressure (similar to the Turkey Point incident) of 3.0E-07/ shutdown year or 1.0E-07/ plant year.
Thus, the results of this analysis (2.3E-05 and 4.7E-05/ plant year) may be more conservative than assumed.
Bothanalysesindicate(
3a,c,e, m.,m... i.
49
4.3 DETERMINATION OF LIMITING TRANSIENTS - 3 TEAM GENERATORS, PRESSURIZERS, AND MAIN COOLANT PIPING The key parameters used in the evaluation of indications discovered during in service inspection are two critical flaw depths.
The first of these critical flaw depths is c.alculated using stresses from governing normal, upset, and test conditions. The second is calculated based on stresses for the governing emergency and faulted conditions.
Critical flaw depths are calculated based on inesi two sets of conditions to correspond to the two ASME Code criteria outlined in Section 1.2.
To allow for the evaluation of indicaU s of various shapes, critical flaw sizes are calculated for emradded flaw. as well as surface flaws of other shapes, with lengths up to continuous flaws.
Critical flaw sizes have been calculated for emergency and faulted conditions and the results for the single
{
l I
)a.c.e, The selection of the governing transient for normal, upset, and test conditions has been done based on the results of the stress analyses at the sections of interest.
Each transient or grouping was analyzed separately, and this made the selection of the governing transient based on stress very straightforward.
nis,n m esto 4-15
TABLE 4-1 KEY PRESSURIZED THERMAL SHOCK TRANSIE'NiS 4
WOG Frequency of Occurrence Per Reactor Year For Transient limiting Events
-- a.c.e o
3" Small Break LOCA in Hot Leg at Zero Power with Accumulator Injection Flow o
3" Small Break LOCA in Hot Leg at full Power o
Loss of Secondary Heat Sink o
Steam Generator Tube Rupture at Zero Power, 30 Minute Delay in S1 Termination o
Steam Generator Tube Rupture at Moderate Decay Heat, 30 Minute Delay in S1 Termination l
xiwm... s o 4-16 s,
w
U.
N I
I i
.E Dw W
I 2:
>=
J W
CC 8
N I
>=
T %
W J
CD
< V)
W W
N e-.
N 2
- C I
W
.j J<
.U N
o.=
UU 4-17
i i
l TABLE 4-3 LTOP EVENT TREE FOR BEST ESTIMATE CALCULATION
~
8,C,4 t
]
i l-CATEGORY DESCRIPTION OK OVERPRESSURE EVENT IS MITIGATED OVPRES OVERPRESSURE SPIKE OCCURS TOTAL FREQUENCY OF SIGNIFICANT OVERPRESSURE EVENT IS
(
3a,c.e l
l m e.n n...,.
4-18
)
TABLE 4-4 LTOP EVENT TREE FOR CONSERVATIVE CALCULATION ll l
l a,c.e J
{
l l
l l
~
CATEGORY DESCRIPTION OK OVERPRESSURE EVENT IS MITIGATED OVPRES OVERPRESSURE SPIKE OCCURS TOTAL FREQUENCY OF SIGNIFICANT OVERPRESSURE EVENT IS
[
jac.e wssm use,o 4 39
TABLE 4-5 EVENT CATEGORIZATION RELATIVE TO EVENT FREQUENCY B C,9 e
i e
f i
1
{
4 4
i 1 :
9 N
L I
i -
=
i i
30tes-9 3 9400 90 i
]
100 7
)
30-5 r
1 l
10-2 r
1 t
WCAP-10319 tO-s NRC TOTAL' WOGTOTL 1
5
~, -
W Loss of Coolant 8-4 rAccident 5
/,#
g i
/
9*
l t
g
,/
10-5 7
,r'
\\
s l lO-, /h ##
s
",j#
/
t 10-s [r,-
P,/
.0-. /.
/
3 t,ce. 4ve Feedwater,,.
/,
30.se200 210 220 230 240 250 250 270 280 290 300 WEAN SURFACC RTNOT l
Figure 4-1.
Frequency of Significant Flaw Extension for Longitudinal Flaws in a Typical Westinghouse PWR wi.mna io 4-21
5
- 6 22
=
s JJ s
e 5
E L
T G
0 l
I
- 4. I a
p_
I S
o n
e i
d A
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t N
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- 2. F e
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g i
n I
s i
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o G 5 r:i -
r: : -
r_: : ~
r::
r.:.
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f A
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s G
=
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R 1
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i P D l
i L
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b R W wg f
weeg go C a._m4moao a.4zo ozoo b
. ww U
or L
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'8 r
I F L 5
u 1
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i 2..
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=
s a
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e 5
F SS D T
le g E L s i s s V A
- 4. I e p R
N 0
s V
0 0
O D 5
I r 0 o5 T
1 t 1 C U
- 3. in c =
i
=
a P A
P e -
l m.
E G R d l
R I
l l
e
- 2. p I
a W L O n
L s
o e A
i n l
t i l
i l
o 1
d t l
n l
T o e O
C B D
r: E r: : -
7:: ~
r::
rE :
r:-
N 1
2 3
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O 0
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1 1
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1 1
erug i
F w5.0,_wesU go U s_ _m4moaa a_<zo;.ozoo we s...e m
.,~~
Pcharacteristic i
Tinitial
'I Tg,4, P T
If 1
Tims LARGE STEAM STEAM GEN.
PARAMETER SMALL LOCA LARGE LOCA LINE BREAK TUBE RUPTURE l
a,c.e i
1 I
i Figure 4-3.
Schematic Representation of Emergency and Faulted Transients for Joseph Farley, along with actual values used for Transients Evaluated.
mie.im.= io 4-23 i
SECT;0N 5 SURFACE FLAW EVALUATION 5.1 CODE CRITERIA The acceptance criteria for surface flaws have been presented in paragraph 1.1.
For convenience they are repeated as follows:
af 51 a For Normal Conditions c
(Upset & Test Conditions inclusive) and af 5 5 a$
For faulted Conditions (Emergency Condition inclusive) where af The maximum size to which the detected flaw is calculated to
=
grow until the next inspection.
10, 20, and 30 year periods have been considered in this handbook.
l l
The ainimum critical flaw size under normal operating a
=
c conditions (upset and test conditions inclusive) ag The minimum critical flaw size for initiation of nonarresting
=
growth under postulated faulted conditions.
(emergency conditions inclusive)
Alternatively criteria based on applied stress intensity factors may be used:
K Kg5 For normal conditions (upset & test conditions inclusive) j$[K For faulted conditions (emergency conditions inclusive)
K u i..,iii... i.
5-1
where The maximum applied stress intensity factor for the final flaw K
g size after crack growth, given the conditions under consideration.
K, Fracture toughness based on crack arrest for the corresponding
=
g crack tip temperature.
Fracture toughness based on, fracture initiation for the K
=
gg corresponding crack tip temperature.
5.2 LONGITUDINAL FLAWS VS. CIRCUMFERENTIAL FLAWS Longitudinal flaws may be defined as flaws oriented in a radial plane, such that circumferential or hoop stresses wuld tend to open them.
On the other hand, circumferential flaws would be oriented in a radial plane such that longitudinal or axial stresses would open then.
These two types of flaws are portrayed graphically in the geometry figure of each section o' the flaw handbooks (1,2,3).
5.3 BASIC DATA in view of the criteria, it is noticed that three groups of basic data are required for the construction of charts for surface flaw evaluation.
- Namely, af, a, and ag, respectively.
c The preparation of these three groups of basic data will be discussed in the following paragraphs.
5.3.1 FATIGUE CRACK GROWTH l
l The first group of basic data required for surface flaw chart construction is the final flaw size a determined from fatigue crack growth. As defined in f
IWB-3611 of Code Section X1, af is the maximum size resulting from growth 1
wiwiinn to 5-2
during a specific time period, which is the next scheduled inspection of the component.
Therefore, the final depth, af, after a specific service period of time must be used as the basis for evaluation.
The charts have been constructed t allow the initial (measured) indication size to be used directly.
Charts have been constructed for operational periods of 10, 20, and 30 years from the time of detection.
The final flaw size can be calculated by fatigue crack growth analysis, which has been performed covering the range of postulated flaw sizes, and flaw shapes at various locations of the reacto.r vessel nneded for the construction of surface flaw evaluation charts in this handbook.
Note that all the finite surface flaws and embedded flaws analyzed are semi-elliptical in shape. All crack growth analyses for finite surface flaws utilize an aspect ratio (length to depth) of 6:1.
This is conservative in all cases.
In some of the regions, it is noted that only the crack growth analysis for l
longitudinal flaws was performed. The crack growth results for the longitu-l dinal flaws can be used for circumferential flaws at the same location with l
some slight conservatism.
In regions where differences are significant, separate analyses have been done, as may be seen in the various sections of Appendix A.
1 5.3.2 MINIMUM CRITICAL FLAW SIZE FOR FERRITIC STEELS By definition o is the minimum critical flaw size for normal operating c
conditions, it is calculated based on the load of the most limiting transient for normal operating conditions. By the same token, a$ is defined as the minimum critical flaw rize for faulted conditions, it is. calculated based on the most governing transient of faulted conditions.
The governing transients are often different for different regions.
The theory and methodology for the calculation of a$ and a has been provided in Section 2.
c w w w.seso s.3 1
.i..,,.
-in..-.n 5.4 TYPICAL SVRFACE FLAW EVALUATION CHART Two basic dimensionless parameters can fully address the characteristics of a surface flaw, and are used for the evaluation chart construction, Namely:
o Flaw Sha.pe Parameter a/t
=
o Flaw Depth Parameter a/t
=
- where, t
wall thickness, in.
=
flaw depth, in.
a flaw length, in.
t A typical chart was chosen for illustration purpose as follows:
(Refer to
~
Figure 5-1) 4 iM M
Ja C e i
h me.n n4.. io 5-4
[
)a c.e The surface flaw evaluation charts constructed for various locations of the reactor vessel are presented in the flaw handbooks (1,2,3).
5.5 PROCEDURE FOR 1HE CONSTRUCTION OF A SVRFACE FLAW EVALVATION CHART A numerical example is used here to show how a surface flaw evaluation chart can be constructed.
Example Required:
To construct a surface flaw evaluation chart for the longitudinal flaws at the beltline region, at the inside surface.
Step 1
[
?
ja c.e i
i N/U/T-normal, upset, and test conditions E/F emergency and faulted conditions w ss:nnse,o 55
13'J
( :
7 -
~
5 k
g fl f,
g{,
,.t,-
I-
?
5 A
6 g
1^g '-
5
~
5 g
'+f^
. t
(
- 9
+
~1 4
=
.m
. J m
4 1
]OsCet Steo' 2 5
L c
s)4,Ci tx r
step 3
(..
A b
4
)3
,C,t'
.q 4
,r e
o b
i^
)
1 ).,
pg?
'9
- 1. ;,,
.aq
? -
s<
i g-
',5 ' $,'
- i%.,g N M !\\\\ M l$.~
w...,
7 1
. - y c.
s f
I
/
k
^h.
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4
>j-i 4
y
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,'Y, -
_ 'I' f.,.
n,
.. u k
.... h { ' ;-
(-
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r
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r
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- :.3 's
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i m um mie imm m u ii 4
ja.c.e Step 4
(
ja,c.e Step 5
[
}8,C,0 ws,? mass so s.y
Note that the allowable flaw depths'here exceed 20 percent of the wall thickness, which has been set as an arbitrary limit, based on engineering judgement. The charts therefore reflect this value as an upper limit.
Step 6 L
l ja.c.e Step 7 Plot a/t vs. a/t data from the standards tables of Section XI as the lower curve of Figure 5-2.
For example, the values of Table IWB-3510-1 for Code editions up until the Winter 83 addendum are:
Aspect Surface
- Ratio, Indication, a/t a/t. %
0.00 1.8 0.05 2.0 0.10 2.2 0.15 2.4 0.20 2.7 0.25 3.1 0.30 3.5 0.35 3.5 0.40 3.5 0.45 3.5 0.50 3.5 I
The above.seven steps would complete the procedure for the construction of the surface flaw evaluation charts for 10 years, 20 years, or 30 years of operating life.
[
ja c.e m..n ii... i.
5-8 L
li.
_ia e
IL 85 lIli
.l k
55
=
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r S,
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(%)2/tHid30MO3 t
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5-9
~.,
LEERB A - The 10, 20, 30 year acceptable flaw limits.
8 - Within this asse. the surface flow is acceptable by A9E Code analytScal critwfa la I M.
10,20,30 Yrs
-= --
A C - A9E Lede a11susble slace 20 - '
4 1983 utster Addenden.
gg :.
=
~... -
D - ASE Code allemoble prior
.s
.z to 1983 Nfater Addendum.
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I l
Figure 5-2.
Evaluation Chart for Reactor Vessel Beltline X
Inside Surface X
Surface Flaw Longitudinal Flaw Outside Surface Embedded Flaw T Circumferential Flaw I
l l
SECTION 6 EMBEDDED FLAW EVALUATION 6.1 EMBEDDED VS. SURFACE FLAWS According to IWA-3300 of the ASME Code Section XI, a flaw is defined as embedded, as shown in Figure 6-1, whenever,
% 3 a (For Editions prior to 1980) or S 3 0.4 a (for Editions of 1980 and thereaf er) t where S - the mi'nimum distance from the flaw edge to the nearest vessel wall surface (clad-base metal interface for flaws near the inside of the vessel) a - the embedded flaw depth, (defined as the semi-minor axis of the elliptical flaw.)
Surface Proximity Rules The surface proximity rules were liberalized with the 1980 Code, allowing flaws as near the surface as four-tenths their width to be considered erbedded. This change resulted from the finding that the original proximity rules had been more restrictive for near-surface embedded flaws than for known surface flaws, which is clearly not technically correct. Specifically, the criterion fer a flaw to be considered embedded was changed to S g 0.4 a, so substituting into the definition for 6 we now find:
wie,mi.e io 6-1
(
6-S a
=
6
> 1.4 a Therefore, the limit for a flaw to be considered embedded is a, = 0.714 6 for Code editions of 1980 and thereafter.
This more accurate criterion has been used throughout this handbook, and is recommended for all inspections, regardless of the edition of the Code which is used for the inspection.
A flaw lying within the embedded flaw domain is to be evaluated by the embedded flaw evaluation charts generated in this section of the handbook.
On the other hand, a flaw lying beyond this domain should be evaluated as a surface flaw using the charts developed in Section 5 of the handbook instead.
The demarcation lines between the two domains are shown graphically in figure l
6-3, for both earlier and later Code editions, in other words, for any flaw indication detected by inservice inspection, the first step of evaluation is to define the category to which the flew actually belongs, then, choose the appropriate charts for evaluation.
6.2 CODE CRITERIA l
As mentioned in Section 1, the criteria used for all the embedded flaws are from IWB-3612 of ASME Code Section XI. Namely, K
K<
For normal conditions (upset & test conditions inclusive) 3 K
i<hForfaultedconditions(emergencyconditionsinclusive)
K 4
33 3
I where The maximum applied stress intensity factor for the flaw K
=
3 size a to which a detected flaw will grow, during the f
conditions under consideration.
K, Fracture toughness based on crack arrest for the
=
3 corresponding crack tip temperature.
I acture toughness based on fracture initiation for the L
K
=
ic corresponding crack tip temperature.
4 The above two criteria must be met simultaneously, in this handbook only the most limiting results have been used as the basis of the flaw evaluation charts.
6.3 BASIC DATA In view of the criter's based on stress intensity factor, three basic groups of data are needed for construction of embedded flaw evaluation charts.
They je, K,, and K, respectively.
The units used herein for all are:
K g
j these three parameters are ksi /in.
K and K are the initiation and arrest fracture toughness values ic la (respectively) of the vessel material at which the flaw is located.
They can be calculated by formulae.
K;c = 33.2 + 2.806 exp[.02(T4TNDT+100*F))
(6-1) and K, = 26.8 + 1.233 exp(.0145( M TNDT+160*F))
(6-2) 3 K is the maximum stress intensity factor for the embedded flaw of 3
interest. The methods used for determining the stress intensity factors for embedded flaws have been referenced in Section 2.
me.n noe io 6-3
and K, are a function of crack tie temperature T, Notice that both K g
lc and the material property of RTNDT at the tip of the flaw.
The upper shelf fracture toughness of the reactor vessel steel is assumed to be 200 ksi/in in all regions.
K used in the determinadon of the flaw evaluation charts is the maximum y
stress intensity factor of the embedded flaw under evaluation.
It is important to note that the flaw size used for the calculation of K; it not the flaw size detected by inservice inspection, instead, it is the calculated flaw size which will have grown from the flaw size detected by inservice inspection.
That means that the embedded flaw size used for the calculation of K; had to be determined by using fatigue crack growth resulti, similar to the approach used for surface flaw evaluation, as illustrated ii the previous section.
6.4 FATIGUE CRACK GROWTH FOR EMBEDDED FLAWS Unlike the surface flaw case, the fatigue crack growth for an embedded flaw l
(even after 40 years of service life) is very small in comparison with that of a surface flaw with the same initial depth.
Consequently, in the handbook l
evaluations, the detected flaw size has been used for evaluation by the charts without any appreciable error.* This simplifies the evaluation procedure without sacrificing the accuracy of the results.
A detailed justification of this conclusion is provided in this section.
The environment of an embedded flaw is considered to be inert, or air.
The crack growth rate for air environment is far smaller than that of the water environment, to which the surface flaw is conservatively considered to be exposed.
Consequently, the fatigue crack growth for an embedded flaw must be far smallar than that of an inside surface flaw (of the same size and under-This conclusion holds for the range of flaw sizes acceptable by the rules of section XI, IWB-3000.
It would not necessarily hold for very large flaws of the order of 50 percent of the vessel wall thickness.-
wis,wwe io 6-4
the same transient conditions). h'umerically, the fatigue crack growth of an embedded flaw is so low that the difference between the initial flaw depth and its final crack depth is negligible.
This engineering judgment has been demonstrated by an illustrative example, as follows:
Example The beltline region of the Joseph Farley reactor vessels was used as a demonstration.
The crack growth results for circumferential inside surface flaws (a/t = 0.167) are as follows.
These flaws were assumed exposed to the water environment.
Postulated Initial Crack Depth Crack Deptb (in.) After Year
~'
a,c.e A similar crack growth analysis was performed for an embedded flaw, using the same set of transients
- and the number of cycles
- as the surface flaw run, and the results follow.
The air crack growth reference law was used.
As specified in Table 2-1.
ain,mn. 3o 6-5
Initial Crack Depth Crack Depth (in.) After Year
- ' a,c.e l
l l
l
(
l 3a,c.e 1
Postulated Final Crack Dersth (in)
Crack Growth for Initial Crack After 40 Years Embedded Flaws, in (%)
Depth, (in)
Embedded Flaws _ _
,c.e a
(
l
)a c.e 6.5 TYPICAL EMBEDDED FLAW EVALVATION CHART The details of the procedures for the construction of an embedded flaw evaluation chart are provided in the next section.
l l
me,n nue io 6-6 1
y
In this section, instructions for reading a chart are provided by going through construction of a typical chart, Figure 6-3, step by step. This will' help the users.to become familiar with the characteristics of each part of the 1
chart, and make it easier to apply.. This example utilizes the surface /.
embedded flaw demarkation criteria of the.1980 Code, and later editions.
following are the highlights of a typical embedded flaw evalustion chart.-
(Refer to Figures 6-2 and 6-3)'.
(-
k i
i i
I ya,c,e l.
- Note that aspect ratio AR = t/a r
m i..n n4 io 6-7.-
I
[
O e
1 I
I l
l
]&,C,9 I
l-i
! 3414s/111440 10 S-8 l-
\\.
l
s i
(
i
\\
e o
b i
.ja c.e These embedded flaw evaluation charts, constructedifor various locations-of l
the reactor vessel, are presented in Appendix A.
6.6 PROCEDURES FOR THE CONSTRUCTION OF EMBEDDEO FLAW EVALUATION CHARTS A numerical example was used in this section to show how'an embedded flaw evaluation chart was constructed step by step as follows:
Example To construct an embedded ' law evaluation chart for circumferential flaws at the beltline. The excess feedwater flow transient was determined to be the governing condition for this example.
9 Step 1
[
3a c.e miwuu e io 6-9 M
[
l I.
Ja.c.e Step 2 i
[
i ja.c.e i
The 141 analyzed cases are tabulated in Table 6-1.
Step 3
[
Step 4
[
3a,c.e m i," ' * "
6-10
(
-- )
-- i I
l The basic concept of the evaluation is that the part of the curves under t~ne-K hlineareacceptablebytheCode' criteria. _Therefore, the intersection K
of-acurvehwiththedrivingforceKI curve indicates the maximum flaw depth acceptable by the Code criteria.
1
{
)
=)
d 1 3 ja,c.e The above four steps.have completely described the procedr es of the~ construc-tion of an embedded flaw evaluation chart for circumferential flaws at the:
inlet nozzle to shell weld.
j I
I
'l 5
_ja.c e 6.7 COMPARISON OF EMBEDDED FLAW CHARTS WITH ACCEPTANCE >. STANDARDS OF IWB-3510
{
)a,c,e
- mwume io 6-11 4
i l
[
3a,c.e me.ii n=uo 6-12 Y
1(
O
+
v
TABLE 6-1 EMBEDDED FLAW CASES' ANALYZED FOR THE INLET N0ZZLE.
TO SHELL WELD-a,c.e i
(
l l
mi..iniam to 6-13 l
SURFACE e
g
\\ \\
\\
- \\
4 h
4 4
//
/
/
= the maximum embedded flaw size
/
a (in depth direction) allowable 1
/
0
_/
per ASME XI* -
S,= the corresponding, minimum depth-L DDED
/
of an embedded flaw (less than D_',AIN which -it must be considered a-L surfaceflaw) a = a, a
0 0
FOR ALL EMBEDDED FLAWS:
If a > a, the flaw must be
~
r
- NOTE:
asa I
charactefized as a surface 0
flaw, with depth = a + 6.
Figure 6-1.
Embedded vs. Surface Flaw wie. mow io 6-14
ed.
e J
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6-15
'1 SURFACE / EMBEDDED PLAW DEMARCATION UNE. SEGINNING WITH 1980 CODE -
0.13 ud5ii 5 M ih si seu8sooao PLaw5i/
i: -! " P W ".. - MW8 WITH A
=
=
.u e.g
., 'gcompieuRAviow 3 p j.j t
0.12 M*..-
D".i.e M 7g ly "t J.6.k. 3 ABOVETHIS UNE ARE c
l im MW
-ii, NOT ALLOWABLE
- .5 n-5
=;
- =
=y
"' :~~.l r '7c y'
t ri
- 1 1 m-
- is
~
- u".s.
=
0'11 8"~~.8iilE
- SURFACE / EMBEDDED -
~
~'iMi f.
Si2 ld i jp.f E Mil,[p=p' J
1 -
jCJ.'"./
UNE UP TIL 1980 CODE
= 'lii s! i j- ~_ id((
F FLAW DEMARCATION F i!!
.p:;: ::..j 0.10 m
m~r
- l l7.-
~w 9
I - ;p4:=r - ;
- =
r.
- -.u ::::
tu.. rt.a l...
+
- g _
_.:: lr :cc.yu.
=a,
- 4. 0.09
.....=: i
!!b. +...
m
..-..y< g,.jj <4 fu.
-w w u.:t
,a
.2 m
w-.
a l:-
- n-ur z%
iiE 15 i, e.se de ?! 17. AW5i L i
~
r es
,= in
] {1 gfg g gg g,g
@ ;;j q mt:
n-0.07 suarace g
.==. g..t
' F
~ M ia
~e
- =
,n....II=! E3""
4 ;;ll N)
M PLAWS lN TMi$
I' 0.06
- RaoioN MUST St *
- =
n.: n::
y v
g;..l : rail :==:
n:
i CONSIDERED 1
d a
-~
=
f.-=.
.. ".3
"~
~'
~.. * *. *.7c.0 -
=
l 2
2
- SURPACE J
w 0.05
" PLaws Eij B
- =...::e~#.
7/..
- =
n-. -=1 !!E !!:1 _-
__irh.liiiiI.
I O 04 E SW ~.M"4 ir M
+ms&
- AR EMBEDDED MWS
. (ON THIS SIDF CF r
n
=
r- =.
,g373="....
- .==-
DEMARKATION UNE)
,;; u..
.. nd r,..
.=.... =
.an,7:
...if :.:. f.......
ra. :
=...:::.,
,.-..a..:...l :=:.
e.....'t....,
ARE ACCEPTABLE PER
. =... u.. a..
0.03 m.: = :=Jr= f-
- u- =
- - ' u
- ' -
- r:
e
- =
!':=
L
.: n7:
.:t-- n CRITERIA OF IWB 3600 m
=
.:: :==
=
=:.
,m
.=3y.: y..
.=:n
.:= :
. - -:un
- m:
- - - -m::.:a,u., =.
A8 LONG A8 g,60.25 0.02 =.::=n..
=
- r. ; av :- - en J /
- e
- .u _w +-
-4
-.::.:.g.:.:
- :=
- .=d:.
g m =
ra 0.01 J
" un = q
=::.
, : ;. a::.u. :::: ::a :m :.-,:.
.-~.,
r
..a
'~'
~
- ~~
- ~
':n n.v.~
~ ' ~
I:A[.
a.-.u: n.. :.1
- --3. r.
r.a.=a:cnu:=
- r.: - n. ::::
P i j=. :::.it.;.:- := au=
m-!:!s h:: !!"
- ni-s;r.-- ::n==]=n =!
o 0.06 0.10 0.15 0.20-0.3 DISTANCE PhtOM SURFACE k)
Figure 6-3.
Embeddad Flaw Evaluation Chart for Inlet Nozzle to Shell Wald (for Longitudinal and Circumferential Flaws) 4 f
m.. ion io 6-16
1 i
p 7*
.e NU 8.
NOTE:-Y={
w
,14 l
1 a.
.i
- r
+
o
.I3 l
l i
,jg
- a..
o 7,
)
.: m
.11 e
i i
a:.
.1
- a:
.10 a
s I
.a
.09
'C' s
.I
.M j-
= Y = 1.0
/
- 07 f
e '
- i..
- f.. y. 0.8-'
E_
- g f ;- _.s f f
ff+
.06
.j y gg, l
j7 f
- ,,,e Y = 0.6 sgr c f"
jr.
sk am i
,a
.03
- -Y = 0.4 s.
=s!" e
,.m r 4
7 p i, as==s d'"" ;-
A
--y
==.ad2-
.02 Py
== ""
[s s'y="
'"'"Y
,i;;;;iille*"""
iusr1983 Ed' tion WD Md.
. 01 liiiiiiiiu'"" "" """"' ""
" Prior to Sanner
~
-~
79 Add.
.I.
0.1 0.2 0.3 0.4 0.5 FLAWSHAPE(a/a)
Figure 6-4.
Illustration of Advantages gained by Analysis for Embedded Flaws at the Inlet wozzle to Vessel Hsid min.m-o 6-17
)
T(s.fd L
- I'
^
r
SECTION 7 REFERENCES' 1.
WCAP-12212, " Handbook on Flaw Evaluation for Joseph Farley Units 1 and 2 Reactor Vessels", by W. H. Bamford, et. al., July 1989.,
2.
WCAP-12213, " Handbook on flaw Evaluation for Joseph Farle'y Units 1 and 2 Steam Generators and Pressurizers,." W. H. Bamford et. al., March 1989.
3.
WCAP-12214, " Handbook on Flaw Evaluation for Joseph Farley Units 1 and 2 Primary Coolant System Piping," W. H. Damford et. al., June 1989.
4.
ASME Code Section XI, " Rules for Inservice Inspection of Nuclear Power Plant Components," 1983 edition (used for updated code 'llowable limits);
1983 edition, Winter 1985 Addendum (used for flaw evaluation of austenitic stainless steel piping); 1989 edition (used for-reference crack growth curve, stainless steel).
l S.
McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for-l Internal Longitudinal Semi-elliptic Surface Flaw in a Cylinder Under Arbitrary Loading", ASTM STP 677, 1979, pp. 365-380, 6.
Newman, J. C. Jr. and Raju, I. S., " Stress -Intensity Factors for Internal' Surface Cracks in Cylindrical Pressure Vessels", ASME Trans., Journal of-Pressure Vessel Technology, Vol. 102, 1980, pp. 342-346.
7.
Buchalet, C. B. and Bamford, W. H., " Stress Intensity Factor Solutions for Continuous Surface Flaws in Reactor Pressure Vessels", in Mechanics of Crack Growth, ASTM, STP 590, 1976, pp. 385-402.
8.
Shah, R. C. and Kobayashi, A. S., " Stress Intensity Factor for an Elliptical Crack Under Arbitrary Loading", Enaineering Fracture Mechanics, Vol. 3, 1931, pp. 71-96.
h,s,mme to.
71
9.
Lee, Y. S. and Bamford, W. H., " Stress Intensity Factor Solutions for a Longitudinal Buried Elliptical ~ Flaw in a Cylinder Under Arbitrary Loads",
presented at ASME Pressure Vessel and Piping Confcrence,-Portland Oregon, June 1983. Paper 83-PVP-92.
- 10. Marston, 1, U. et.'a1. " Flaw Evaluation Procedures:
ASME Section XI" Electric Power Research Institute Report EPRI-NP-719-SR, 1ugust-1978.
- 11. Logsdon, W. A.,
" Dynamic Fracture Toughness cf ASME SA508 C2a Base and' Heat-Affected Zone Material,-in Elastic-Plastic' Fracture, ASTM STP 668, 1979.
- 12. Logsdon, W. A., " Dynamic Fracture Toughness of Heavy Section, Narrow Gap Gas Tungsten Arc Weldments" Engineering Fracture Mechanics, Vol. 16, No.
6, 1982.
- 13. Logsdon, W. A., " Dynamic Fracture Toughness and Fatigue Crack Growth Rate Properties of ASME SA508 C1.3 and SA508 C1.3a base and Heat Affected Zone l
Materials" in ASTM Journal.of Testing and Evaluation, Vol. 10, July 1981.
l
- 14. Logsdon, W. A., and Begley, -J. A., " Dynamic Fracture Toughness-of SA533 Grade-A Class 2' Base Plate and Weldments" in Flaw Growth and Fracture, ASTM STP 631, 1977.
- 15. USNRC Standard Review ilan, NUREG 0800, July 1981.
- 16. Bamford, W. H. and Bush, A.
J., " Fracture of Stainless Steel," in Elastic Plastic Fracture, ASTM STP 668, 1979.
- 17. 'Landes, J. D., and Norris, D. M., " Fracture Toughness of Stainless Steel Piping Weldments," presented a' ASME Pressure Vessel Conference,1984,
- 18. USNRC Regulatory Guide 1.99, Effects of Residual Elements on Predicting Rauiation Damage to Reactor Vessel Materials, July 1975; Rev.1:
April i
1977; Rev. 2 1986.
3 l'
m.,n m.."
7-2 l.
I
i
- 19. Plane Strain Crack Toughness Testing of High Strength Metallic Materials, ASTM STP 410, March 1969.
- 20. Logsdon, W. A., Liaw, P. K., and Begley, J. A., " Fatigue Crack Growth Rate Properties of SA508 and SA533 Pressure Vessel Steels'and Submerged Arc Weldments in Room and Elevated Temperature Air Environments" Engr.
Fracture Mechanics, Vol. 2, No. 3, 1985.
l l
- 21. James, L. A., and Jones, D. P., " Fatigue Crack growth Correlations for l
Austenitic Stainless Steel in Air," in Predictive Capabilities in Environmentally Assisted Cracking," ASME publication PVP-99, Dec. 1985.
- 22. Bamford, W. H.,
' Fatigue Crack Growth ~of Stainless Steel Piping in a Pressurized Water Reactor l Environment," Trans ASME, Journal of' Pressure Vessel technoi:,9y,- Feb.1979,
- 23. " Evaluation of Flaws in Austenitic Steel Piping," Trans ASMC, Journal of-Pressure Vessel Technology, Vol. 108, Aug. 1986, pp. 352-366.
- 24. WCAP-10019, " Summary Report on Reactor Vessel Integrity for Westinghouse Operating Plants," December, 1981.
- 25. " Summary of Evaluations Related to Reactor Vessel Integrity," report performed for the Westinghouse Owner's Group, Westinghouse Electric Corporation, Hay, 1982,
- 26. Letter 0. D. Kingsley, WOG, to H. Denton, NRC, " Westinghouse Owner's Group Activities Related to Pressurized Thermal Shock," 0G-73, July 15,1982.
- 27. Letter from 0. D. Kingsley, WOG, to H. Dentuc, NRC, " Westinghouse Owner's Group Activities Related to Pressurizcd Therma, Shock, ' OG-79, September 2, 1982.
uiwmue io 7-3
- m. w
-~
f
- 28. SECY-82-465, United States Nuclear Regulatory Commission Policy lssue,
" Pressurized Thermal Shock (PTS)," November 23, 1982.-
- 29. U. S. Nuclear Regulatory Commission, 10CFR50, " Analysis of Potential l
Pressurized Thermal Shock Events," Federal Register Vol. 50, No.141, July l
23, 1985.
i
- 30. WCAP-10319, "A Generic Assessment of Risk from Pressurized Thermal Shock of Reactor Vessels on Westinghouse Nuciear Power Piants," July, 1983.'
i.
-1 4
k wwmasuo y.4