ML20106A763
| ML20106A763 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 08/31/1992 |
| From: | Caine T, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20106A760 | List: |
| References | |
| DRF-137-0010-5, DRF-137-10-5, GE-NE-523-70-06, GE-NE-523-70-0692, GE-NE-523-70-6, GE-NE-523-70-692, NUDOCS 9209290406 | |
| Download: ML20106A763 (88) | |
Text
7 GE-NE-523-70-0692 DRF # 137-0010-5 August 1992 OYSTER CREEK VESSEL FRACTURF. MECHANICS ANALYSIS FOR UPPER SHELF ENERGY REQUIREMENT Prepared for GPb Nuclear Corporation One Upper Pond Road Parsippany, NJ 07054 Prepared by GE Nuclear Energy 175 Curtner Avenue San Jose, CA 95125 9209290406 920922--
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0YSTER CREEK VESSEL FRACTURE MECHANICS ANALYSIS FOR UPPER SHELF ENERGY REQUIREMENT Prepared by: M
^
s H.S. Meht-Principal Engineer Materials Monitoring &
Structural Analysis Services Verified by:
2.s T.A. Caine. Senior Engineer Materials Monitoring &
Structural Analysis Services Approved by:
P"
- 5. Ranganath Manager Materials Mcr.itoring &
Structural Analysis Services
)
TABLE OF CONTENTS Pace 1.
INTRODUCTION AND BACKGROUND 1-1 1.1 General 1-1 1.2 Geometry and Material Property Data on Oyster Creek RPV 1-2 1.3 Report Outline 1-3 1.4 References 1-4 2.
EVALUATION METHODOLOGY 2-1 2.1 Acceptance Criteria 2-1 2.1.1 Level A and B Loadings 2-1 2.1.2 Level C Service Loadings 2-2 2.1.3 Level D Service Loadings 2-3 2.2 Calculation of the Applied J-Integral 2-3 2.3 Evaluation using Critcrion for Flaw Growth of 0.1 Inch 2-5 E.4 Evaluation Piocedures for Flaw Stability 2-5 2.5 Summary of Evaluation Methods 2-5 3.
IRRADIATED FRACTURE TOUGHNESS CHARACTERIZATION 3-1 3.1 Charpy Energy Versus J-R Curve Correlations 3-1 3.1.1 SA 302 Grade B Data 3-2 3.1.2 SA 533 Grade B (302B modified) Data 3-3 3.2 Charpy USE Values Selected for J-R Curve Determination 3-5 3.3 Determination of J-R CJrves 3-5 3.4 Summary of J-R Curve Determination 3-6 3.5 References 3-6 i-
TABLE OF CONTENTS (CONT'D)
Pace 4.
EVALUATION nF LEVEL A AND B CONDITIONS 4-1 4.1 Level A and B Service loadings 4-1 4.2 Evaluation per Jo,1 i
Cr ~,rion 4-1 4.3 Stability Evaluation 4-2 4.4 Summary of Service Level A and B Evaluation 4-2 4.5 References 4-3 5.
EVALUATION OF LEVEL C AND D CONDITIONS 5-1 5.1 Evaluation of Level C Conditions b-1 5.1.1 Transient Definition 5-1 5.1.2 Finite Element Stress Analysis 5-2 5.1.3 Fracture Mechanics Evaluation 5-2 5.1.4 Acceptance Criteria Evaluation 5-3 5.2 Evaluation of Level D Conditions 5-4 5.2.1 Description of LOCA Event 5-4 5.2.2 Fracture Mechanics Analysis Results 5-5 5.2.3 Acceptance Criteria Evaluation 5-5 5.3 SummLry of Level C and D Loriings Evaluation-5-6 5.4 References 5-6 6.
SUMMARY
-ANJ CONCLUSIONS 6-1 APPENDIX: DRAFT OF SECTION XI, APPENDIX XX
.II
- -- - A
)
1.0 INTRODUCTION AND BACKGROUND
)
1.1 General The Oyster Creek Nuclear Generating Station has a 620 MWe rated BWR i vpe nuc' ear steam supply system.
The reactor pressure vessel (RPV) was constructed to the requirements of Section I of the ASME Code (1959) witi.
Addenda through Winter 1963. ASME Code Cases 1270N and 1273N [1-1).
The pressure temperature (P-T) curv6s in the technical specifications meet the requirements of 10CFR50, Appendix 6 [1-2) to assure that the RPV temperatures and pressures during various operating conditions are such that the brittle fracture of the RPV is prevented.
The P-T curves account for irradiation embrittlement effects in the core region, or beltline.
The p
method described in Regulatory Guide 1.99, Revision 2 (1-3) was used in Reference 1-4 to account for irradiation embrittlement effects for the Oyster Creek RPV.
Reference 1-2 states that the RPV must maintain upper-shelf energy (USE) throughout its life of no less than 50 ft-lb, unless it is demonstrated in a manner approved by the Director, Office of Nuclear Reactor RegLlation. that lower values of upper shelf energy will provide p
margins of safety against fracture equivalent to those required by Appendix G of Section 111, ASME Code [1-5).
The analysis presented in the cover letter for Reference 1-4 indicated that the projected 32 effective full power year (EFPY) upper-shelf energy for son' of the RPV plates in the p
beltline region will fall below the 50 ft-lb value.
Based on this information, the NRC staff has requested a Oyster Creek specific analysis showing the basis for present nd continued vessel structural integrity when the 50 ft-lb requirement is not satisfied.
in recognition of the need for such an analysis for light water reactorsSection XI of the.ASME Lade has developed procedures and acceptance cr%ria for the evaluation of low upper-shelf energy sessels, p
These procedure and acceptance criteria, currently in the form of a draft 1-1
i 3
called Appendix XX, are expected to be approved by the Section XI Subcommittee in the near future.
The text of the most recent draft of l
Appendix XX, approved at the Working Group level in May 1992, is included j
l here as Attachment A of this report.
l The objective of this report is to present a fracture mechanics evaluation of Oyster Creek RPV using the precedures and acceptance criteria outlined in Appendix XX.
The Code of construction for the Oyster Creek RPV did not have the normal (Level A), upset (Level B), emergency (Level C) and faulted (Level D) condition categorization for the loadings and thermal transients defined in the loading diagram.
Therefore, more recent BWR thermal cycle diagrams were used in categorizing the loadings and to define the appropriate loadings to consider for the evaluation of level C and D conditions.
1.2 Geometry and Material Property Data on Oyster Creek RPV l
The geometry of the Oyster Creek RPV and the ASME Code stress analysis are documented in Reference 1-1.
The part of the RPV evaluated in this report is the cylindrical beltline region.
In that region, the inside radius and the thickness of the vessel is 106.7 inches and 7.125 inches, respectively.
The chemical compositions of the beltline plates and welds, obtained from Reference 1-4, are shown in Table 1-1.
A review of Table 1-1 shows that the two lower shell pla' s (G-307-1 and G-3081) have reported nickel content of 0.117..
This makes these two plates, which come from the same heat, confirm to the composition requirements of SA 3028.
All the other plates have nickel contents between 0.4Y. and 0.7Y., the requirement for SA 302B modified which was later designated as SA 533 Grade B.
Due to nickel addition, the SA 302B modified has, on the average, better fracture toughness properties than the unmodified SA 302B.
This distinction is important, as discussed later, in determining the appropriate material fracture toughness for the Appendix XX i
e"aluations.
1-2
4 i,
Table 1-2 from Reference 1-4 shows the projected Charpy USE values for i
j 32 effective full power years (EFPY) of operation.
It is seen that the projected USE values for the plates G-308-1, G-307-5 G-8-7 and G 8 6 are t
j lesc than 50 ft-lbs.
(Some of the projected USE values for the welds in j
{
Table 1-2 are also below 50 ft-lbs, but they were calculated based on the j
test temperature of 10' F.
If this overconservatism is removed, all of the i
calculated USE values fur the wolds will exceed 50 ft-lbs (1-6),)
Therefore, these plates were evaluated per Appendix XX procedures.
I J
j The projected USE values shown in Table 1-2 are for the transverse i
direction (normal to the rolling direction).
For the rolled law alloy steel plates, such as those used in the fabrication of RPVs the transverse I
direction gives lower USE as compared to the longitudinal direction (parallel to the rolling direction).
For the plants constructed to the ASME Code effective prior to Summer 1972 (which is the case for Oyster a
j Creek), only the longitudinal direction Charpy energy testing was. required.
i Therefore, the initial transverse USE values shown in Table 1-2 were r
l obtained by multiplying the initial longitudinal USE by 0.65, based on the l
guidelines given in Reference 1-7.
l 1
Since it is the projacted transverse direction USE that-falls below 50 j
ft-lbs, the postulated flaws are evaluated only in the orientation for which this direction toughness is relevant.
This means, in the case of biritline plates, only the circumferential flaws need to be evaluated.
j Therefore, the evaluations in the following Sections considered only the l
circur.ferential cracks.
i
+
1.3 Report Outline l
l Section 2 describes the methodology of Appendix XX and how it was implemented in this evaluation.
Irradiated fracture toughness properties i
of SA 302B and SA 3028 modified published in the-technical literature are l
reviewed in Section 3 to determine the initiation toughness (J e) and the i
l J-Resistance (J-R) curves based on the lower bound and tho'mean of the l~
data.
Section 4 describes the evaluation of 1.evel A and B conditions based l
1-3 i
e u..- - - - -.-.,-, - -. -. -,. - ~. x, -.,,,,--
on the lower bound toughness values.
Evaluation of Level C and D conditions is covered in Section 5.
Section 6 presents the summary of results and conclusions.
1.4 References 1-1 Pierson, T.M., Larkin, T. A. and Kinyon, B.W., " Analytical Report for Jersey Central Reactor Vessel," General Electric VPF # 1238-148-1 (1970).
1-2 " Fracture Toughness Requirements," Appendix G to Part 50 of Title 10, the Code of Federal regulations, July 1983.
1-3 " Radiation Embrittlement of Reactor Vessel Materials," USNRC Regulatory Guide 1.99, Revision 2. May-1988, 1-4 " Pressure Temperature Curves per Regulatory Guide 1.99, Revision 2, for the Oyster Creek Nuclear Generating Station " GE Report No. SASR 90-89, DRF # 137-0010 (1990).
1-5 " Fracture Toughness Criteria for Protection Against Failure," Appendix G to Section XI of the ASME Boiler & Pressure Vessel Code, 1989 Edition.
1-6 "Unirradiated Upper Shelf Energy Values for Reactor Vessel Belt Line Welds," GPUN Letter to NRC, Letter # 5000-91-2056, dated July 17, 1991.
1-7
" Fracture Toughness Requirements," USNRC Branch Technical Position HTEB 5-2, Revision 1. July 1981, 1-4 J
TABLE 1-1 Chem: Cal Composition of RPV Beltline Materials Composition by Weirbt Percent Identification IIcat/ Lot No.
C
.... Mn P
S SL Si Mo Cu l
Lower Shell Places:
C-307-1 T1937-2 0.2 1.4 0.011 0.022 0.24 0.11
- 0.51 0.17
- C-308-1 T1937-1 0.2 1.4 0.011 0.022 0.24 0.11
- 0.51 0.11
- C-307-5 P2076-2 0.2 1.28 0.019 0.030 0.21 0.53 0.52 0.27
- Lower-Intermediate Shell Plates:
C-P 7 P2161-1 0.19 1.35 0.019 0.021 0.24 0.48 0.46 0.21
- C-8-8 P2136-2 0.19 1.36 0.006 0.024 0.26 0.46 0.48 0.18
- G-8-6 P2150-1 0.2 1.25 0.013 0.026 0.23 0.51 0.46 0.20
- T w
Lower Shell Longitudinal Welds:
2-564 RACOf3, 860545 0.12 1.64 0.015 0.02 0.34 0.2
- 0.51 0.35
- A.B.C ARCOS B-5 Lot 4E5F Lower-Intermediate Longitudinal Veld:
2-564 RACo#3, 860548 0.12 1.67 0.013 0.02 0.41 0.2
- 0.50 0.35
- D.E.F ARCOS B-5 Lot 4D4F Lower to Lower-Intermediate Cirth Weld:
3-564 RACo#3, 1248 0.097 1.26 0.015 0.02 0.22 0.11
- 0.57 0.22
- ARCOS B-5 Lot 4M2F
- Valtres reported in GPUN TDR 725
I 2
i i
l l
TABLE 1-2 F
3 Upper Shelf Energy Analysis for j
Oyster Creek Beltline Materials i
a INITIAL INITIAL 32 EFPY 32 ErPY
- j TEST LONGIT.
TRANS.
FLUENCE % DECR.
TRANS.
l LOCA* ION IDENT TEMP USE USE teu (x10'18)
USE Ust PLA*ES:
i Leser G-307-1 160 9v 64.4-0.17 2.36 18.5 52.4 C-308-1 212 92 59.8 0.17 2.36 18.5 48.7 g
i G-307-5 212 95 61.8 0.27 2.36 23 45.7 1
i i
Low-Int.
G-8-7 212 79
.4 0.21 2.36 21.5 40.3 1
G-6-8 212 100 65.0 0.18 2.36 19.5 52.3 C-8-6 212 81 52.7 0.2 2.36 21 41.6 WELDS:
Vertical 2-564 10 66 0.35 2.36 34 43.6*
A,B,C 2-564 10 32 0.35 2.36 34 21.1*
D,E,F s
Girth 3-564 10 65 0.22 2.36 2a 48.1 e
t f
Surveillance 350 83.5 34 55.1 i
- See discussion of paragraph 1.2 on page 1-3 i
I l-6 i
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, _. -. - _ _.. - _.. _ -... - - - -... - - -.. - -, ~, -...
-_.._._,-_....._.,.,--,-,_.u-.
(
2.0 EVALUA110N HETHODOLOGY The evaluation methodology followed in this report is that prescribed in Appendix XX, the latest draf t of which is included as Attachment A.
There are essentially four steps in this methodology: (1) postulate flaws in the reactor vessel, (2) determine the loading conditions at the location of the postulated flaws for tevel A, B, C and D Service loadings, (3) obtain ^he material properties, including E, ay, and the J-integral resistance curve (J-R curve), at the locations of the flaws, and (4) evaluate the postulated flaws according to the acceptance criteria.
Article A-3000 of the Appendix contains general description of procedures used to evaluate the applied fracture mechanics parameters, as well as requirements for selecting the J-R curve for the material.
Detailed calculation procedures for Level A and B Service loadings are given in A-4000.
The Appendix does not include a detailed calculation procedure for Level C and D Service loadings since it was concluded that the possible combinations of loadings and material properties which may be encountered during these Service conditions are too diverse.
The acceptance criteria and the calculations procedures from this Appendix as applicable to subject evaluation are described in this section, n
2.1 Acceptance Criteria 2.1.1 Level A and B Service Loadings I
j An interior semi-elliptical surface flaw with a depth one-quarter of the wall thickness and a length six times the depth is postulated.
Since it is the transverse direction USE that is projjeted to be below 50 ft-lb value, only a circumferential1y oriented flhv is postulated for this evaluation.
Thef1@isassumedtobe}ocaiedinthebasematerial.
(
g2-1
'j.1
l l
Toughness properties for the corresponding orientation (transverse orientation) are to be used in the evaluation.
Two criteria which shall be satisfied are (1)
The applied J-integral evaluated at a pressure which is 1.15 times the accumulation pressure as defined in the plant-specific Overpressure Protecticn Report, with a factor of safety of 1.0 on thermal loading for the plant specified heatup and cooldown conditions, shall be shown to be less than the J-integral characteristic of the material resistance to ductile tearing at a flaw growth of 0.10 in.
(2)
The flaw shall be shown to be stable, with the possibility of ductile flaw growth, at a pressure which is 1.25 times-the -
accumulation pressure defined in (1), with a factor of safety of 1.0 on thermal loading for the plant specifled heatup and cooldown conditions.
The J-integral resistance versus crack growth curve (J-R curve) shall be e conservative representation for the vessel material under evaluation.
The determination of the J-R curve for this evaluation is discussed in Section The mathematical expressions for the calculation of applied J-integral 3.
and for the evaluation of stability are discussed in Subsection 2.2.
[
2.1.2 Level C Service Loadings I
While the shape and the aspect ratio are the same as those for the j
Level A and B Service loadings, the depth of the-postulated flaw for this service condition is 1/10 of the base metal wall thickness, plus the cladding thickness, with total dep.th not to exceed 1.0 inch.
Smaller maximum flaw sizes may be used on an individual case basis when justified.-
Two criteria which shall be satisfied ares l
l 2-2 i
4 wy
..,w.y-e-+-es
--,r--y,w--r--+y+--,---e--,
.*-w,g-e w+W~1'
-te+w+-,s-rn-e--t
+-e-e s*-
+--e
,ww
+-
w e w w.w e-*
er e w a -,
e-r
= wt r rw,
.,y mei-+rw, rw---ee w v *=
I i
(1)
The applied J-integral shall be shown to be less than the i
J integral characteristic of the material resistance to ductile f
tearing at a flaw growth of 0.10 in., using a factor of safety j
of 1.0 on loading.
i l
)
(2)
The flaw shall be shown to be stable, with the possibility of f
ductile flaw growth, using a factor of safety of 1.0 in loading.
i The material J-R curve shall be the same as used in the evaluation of f
Level A and B conditions.
Thus, the key differences between the Level A/B
{
evaluation and the Level C evaluation are the postulated flaw size and the factor of safety.
i 2.1.3 Level D Service Loadings i
1 i
The postulated flaw geometry for this service condition is the same as l
that for the evaluation of Level C loadings.
The flaw shall be shown to be l
stable, with the possibility of ductile flaw growth, using a f actor of
{
safety of 1.0 on loading.
The J-R curve shall be a best estimate representation for the vessel material under evaluation.
The stable flaw cepth shall not exceed 75". of the vessel wall thickness, and the remaining ligament shall be safe from tensile instability.
2.2 Calculation of the Applied J-Integral i
j The calculation of applied J-integral consists of two steps: Step 1 is
{
to calculate the ef fective flaw depth which includes a plastic-zone j
correction; and Step 2 is to calculate the J-integral for <.nall.uale yielding based on this effective flaw depth.
l For the postulated circumferential flaw with a depth *a', the stress intensity factor due to internal pressure with a safety f actor (SF)- on 1
pressure, was calculated using the following:
Kip = (SF) p (1 + (Rj/(2t))) (na)0.5 F2 (2-1) 2-3 i
i
~
-._,--,-.--.m
....._.._.._......._.m,_.,_..-._._..
I j
F2 = 0.885 + 0.233(a/t) + 0.345(a/t)2 (2 2) i where, Rj, t and a are vessel inside radius, vessel thickness and crack j
depth, respectively.
This equation for K]p is valid for 0.2 s a/t s 0.50 l
and includes the ef act of pressure acting or, the flaw faces.
The units i
for K are ksi/in.
l l
For an axial or circumferetial flaw with a depth 'a',the stress intensity factor due to radial thermal gradient was calculated by using the 4
j following:
2 i
Kit = ((CR)/1000) t.5 F3 (2-3) i F3 = 0.584 + 2.647(a/t) --6.294;'It)2 + 2.990(a/t)3 (2 4) where CR is the cooling rate in 'F/ hour and the units of K are ksi/in.
This equation for kit is valid for 0.201 a/t 5 0.50, and 0 s (CR) <
~
100'F/ hour.
The effective flaw depth, a, is then calculated by using:
e i
e = a + (1/(6x))[(Kgp + kit)/oy]2 (2-5) a
{
where. 0y is the material yield stress.
The K'jp and K'It are calculated by substituting ae in place of a in l
equations 2-1 and 2-3.
The J-integral due to the applied loads for small j
scale yielding is then given by:
y J = 1000 (K']p + K'It)2/E' (2-6)
E' = E/(1-v2)
(2-7) 3' where, E is Young's Modulus und v is poisson's retio (=0.3).
The units of J are in-lb/in2, 2-4 1
4
,,--,p
,,y-----.--,-y,w,..--n.,r.,y,m..#%,,m_,,~,,,4y...
,,,.y.
,m y w r,,,.
,p.,m,,.+g,
,~w~
aw-,-~,.
.r ew s'
2.3 Evaluation Using Criterion for Flaw Growth of 0.1 Inch The J-inteoral due to the applied loads, J, for this case is 3
calculated using a factor of safety of 1.15 on the accumulation pressure.
The acceptance criterion for Level A and 8 Service loadings based on a ductile flaw growth of 0.1 inch (Criterion 1 in 2.1.1) is satisfied when J 1
< Jo,1, where J.1 is the value of J-integral in the material J R curve at 0
a Aa of 0.1 inch.
The thermal gradient contribution (K it) to the J-integral due to the applied loads for the Level C and 0 conditions, was calculated using the finite element stress analysis and available K solutions in the literature.
2.4 Evaluation Pror.edures for Flaw Stability Appendix XX provides three approaches that are equally acceptable for applying the flaw stability acceptance criteria.
The first is the J-R curve - crack driving force diagram approach.
In this approach, flaw stability is evaluated by a direct application of the flaw stability rules given in A-3400.
The other approaches are the failure assessmert diagram approach and the J-integral / tearing modulus approach.
The first approach was used in this report.
The J-R curve - crack driving force diagram approach is illustrated graphically in Figure 2-1.
The applied J-integral curve is evaluated at a constant load. -The 3-P. curve intersects the horizontal axis at the postulated initial flaw depth. ao.
Flaw stability at a given load is demonstrated when the slope of the J-R curve at the point on the-J-R curve where the two curves intersect.
The onset of flaw instability occurs-at an applied load corresponding to the point of tangency of the applied J-integral curve and the J-R curve.
2.5 Summary of Evaluation liethods The acceptance criteria and the evaluation methods of Appendix XX of Section XI, relevant to the USE requirement evaluation of Oyster Creek 2-5
,1 1
i 1
I i
plant RPV are summarized in this section. The key input in this evaluation 4
)
are the appropriate material J R curves and the applied J-integral values.
The selection of appropriate material J-R curves :onsidering Oyster Creek l
plant-specific irradiation data is described in Section 3.
Sections 4 and i
5 describe the rationale for the selection of appropriate pressure and thermal loadings and evaluation results for the Levels' A/B and C and D j
conditions, f
f l
L 4
i I
l i
t l
l i
l I
i
}
{
1 i
i d
2-6 l
1 1
i o
1
- ~ ----+ -.=
-,.--.n--,,--,----.-,
... -, - - =.
~-.,e+
s.e,.
m
..,,~-n,~-.,~..
,,-,n,w:,n--
n
-n--n--
,- ~.,,-- w
.-r---
m J-R Curve y
\\
J.
Evaluation Point i
i a,
o Figure 2-1 Comparison of the Slopes of the Applied J-integral Curve and the J-!! Curve 2-7
3.0- IRRADIATION FRACTURE TOUGHNESS CHARACTER!ZATION A key input in the evaluations based on the procedures of Appendix XX is the material J-R curve.
The beltline region material cf the RPV wall undergoes irradiation-induced toughness changes during the oporation of the plant.
Therefore, the material J-R curves used in this evaluation must factor-in the irradiation effects on the fracture toughness.
This section describes the selection of J-R curves for this evaluation considering the projected irradiated Charpy energies for the Oyster Creek RPV and the available correlations in the technical literature between the Charpy energy and the material J-R curves.
3.1 Charpy Energy Versus J-R Curve Correlations The available information to assess the state of vessel wall embrittlement generally consists of initial Charpy energy values, the caterial chemistry and the fluence level at the vessel wall.
Regulatory Guide 1.99, Revision 2 [ Reference 1-3] provides method to calculate the Charpy USE at any irradiation level given the preceding information This info *mation can be supplemented when the Charpy energy values from the testing of periodically removed surveillance specimens become rsailable.
Thus, most of the materi_al fracture test data available in the technical 5
literature include the Charpy energy values along with the J-R curve information.
Considerable data on the unitradiated anti irradiated low alloy steels have been reported in the technical literature (3-1 through 3-4).
Two broad categories of low alloy steels are generally covered in these studies: (1) SA 302 Grade B, and (2) SA-533 Grade B (equivalent to SA 302 Grade B, modified by nickel addition). The material for the beltline plate that has the lowest USE (G-8-7 in Table 1-2) is SA 302B modified. However, one of the SA 302B plate (G-308-1) in the beltline region has 32 EFPY transverse USE less than 50 ft-lbs. Therefore, to perform a bounding 3-1
i i
j evaluation, Charpy USE to J-R curve correlations for both the materials are j
discussed in this section and the appropriate J-R curves are developed.
3.1.1 SA 302 Grade B Data l
To establish a J]c versus Charpy USE correlation that can be used to l
assess the mean and mean minus two standard deviation J-R curve, test data j
on SA 302B base metal from References 3-2 and 3-4 were compiled and plotted 1
in Reference 3-5.
Figure 3-1 shows such a plot.
The data in Figure 3-1
]
j are for temperatures in the range of 400* F to 550' F for both the j
longitudinal and transverse orientations and irradiated as well as j
unirradiated material.
Fluences ranged from 0.6 to 3.5 x 1019 n/cm2, j
The data in Figure 3-1 show a trend for decreasing J c with decreasing I
f Charpy USE.
The mean and the mean minus two standard deviation lines shown l
in N ure 3-1 were obtained using linear regression analysis.
The mean l
line was used to determine the best estimate J-R curve for the evaluation of Level D loadings.
The mean minus two standard deviation line was used i
j to establish a conservative representation of the J-R curve, as required in l
l the evaluation of Levels A,B and C loading conditions.
j The J-R curves for most materials can be represented by an equation of I
l the following form:
i i
j J = C (Aa)D (3-1) l This equation gives a convex upwards form of J.R curve in which the J l
values keep increasing with increase in Aa.
Reference 3-3 suggests that SA 302 Grade B apparently exhibits a size effect where the J-R curve flattens j
significantly with increasing specimen thickness as illustrated in Figure 3-2.
This result was also unusual in that the J R curve of e thicker (larger) specimen was lower than that of a thinner (smaller) specimen.
A l
subseque evaluation [3-6] examined several possibilities for this unusual j
size ef fect but could not cite any clear cause.
Consequently, an approximate method was used in Reference 3-5 to reflect this size effect.
l 3-2 i
l
I 9
i j
Data in Reference 3-3 show that the effect of thickness on J e is t
I relatively small.
Therefore, only the size effect on the J-R curve at larger Aa values was considered.
Results plotted in Figure 3 2 show that j
the J-R curve for the 6TCT specimen rises above J c by a factor of about I
l 1.3 and reaches a plateau, While it is not clear that this size effect l
seen at 180* F would be also present at 400' F to 550' f temperature range, i
it was assum&' that this was the case.
l Therefore, it was assumed in Reference 3-5 that the J-R curves for l
both longitudinal and transverse direction will flatten out at 30% above J c.
The same conservative assumption was also used in this report to I
j develop the J-R curve for SA 3028 material.
I 3.1.2 SA 533 Grade B (3028 modified) Data i
l A comprehensive multivariabic modeling of RPV and piping J-R data is f
reported in References 3-7 and 3-8.
Separate models were fitted for i
different materials groups, including RPV welds, Linde 80 welds, RPV base metals, piping welds, piping base metals, and a combined materials group.
The material data base did not include SA 3028 steel, but included SA 533 f
Grade B steel, which is relevant for SA 302B modified plate G-8-7.
The material J-R curve in Reference 3-8 is represented in the following form:
Jd = C1 (Aa)C2 (exp(C3(Aa)C4])
(3-2)
Jd i s the deformation J-integral.
The use of deformation J-integral, rather than the modified J-integral, is currently favored in the fracture-mechanics evaluations. The notation J is also used in this report' to i
indicate deformation J-integral. The expressions for C2 and C3 terms are:
i i
j C2 = di+d2 (in C1)
(3-3)-
l l
'C3 = d4^d5 (in C1)
(3-4) i.
I 3-3 I
i i
-e
-.._,.,_.,.,-_.m..-..._.--..._m..__,..
.,-..,.=m
._...-m.,-,,_,.,m..-__..._.,,,-.,-.....-_...-_--
The parameter C1 can be calculated from the following expression when the pre-irradiation Charpy US7 (CVN ) and the fluence (pt) values are p
available:
In (C1) = a1 + a2 in (CVN ) + a3T + a5ct (3-5) p The variable T is the test temperature.
When only the irradiated Charpy USE (CVN) is used. C1 is determined as follows:
in (C1) = a1 + a2 in (CVN) + a3T (3 6)
The parameters al a2+ a3, a5, d, d, d, d, and exponent C4 are l
2 4
5 constants, the values for which are given in Tabic 3-1.
When CVN, ct and p
CVN are available, as is the case for present evaluation, the parameter C1 can be calculated by either equation (3-5) or_ equation (3-6).
The J-R curves obtained by these two different approaches are, in general, different.
For conservatism, the lowest of the twc J R curves was used in the evaluation.
Although not necessary in the Appendix XX evaluations, the J c values may be obtained from the J-R curves using the definition given in ASTM Standard E 813, or the value of Jd at the intersection of the power law fit curve and the blunting line with 0.2 mm offset (see Figure 3-3).
The values of various parameters shown in Table 3-1 would give a best estimate or mean J-R curve.
Such a J-R curve can be used for the evaluation of Level D loadings.
A best estimate (or mean) minus two standard deviation J-R curve was considered a reasonably conservative representation of material toughness suitable for use in the evaluation of Level A through C condition loadings.
Such a J-R curve was obtained by multiplying the Jd values obtained from equation (3-2) by the value of the ratio (at the bottom of Table 3-1) corresponding to 2Se.
For example, the Jd values obtained using the CVNp model were multiplied by 0.741 to obtain the mean minus two standard deviation value of J -
d 3-4
_________-____----------------J
7_.__________
j i
3.2 Charpy USE Values Selected for J R Curve Determination i
d The lowest value of transverse USE in Table 1-2 is 40.3 ft-lbs.
This lj value tc for plate G-8 7 at 32 EFPY.
This is based on the fluence value at 1/4 thickress.
T'n pr.;tulated flaw for the Level C and D loadings is 1/10 thickness.
The transverse USE for thic case was calculated as 39.6 ft-lbs.
in view of the small difference between the two energy levels, a single 2
value of 40 ft-lbs was used in determining the J R curve.
If the plant life extension (PLEX) is implemented at Oyster Creek plant, the fluence level would be higher than that calculated for 32 EFPY.
)
To evaluate the PLEX case, a very conservative evaluation was conducted j
assuming a Charpy USE of 35 ft-Ibs for plate G 8-7.
This-is quite conservative since the fluence level associated with J5 f t-1bs USE is i
estimated as 1.3 x 1019 n/cm2, which is projected to be more than 170 EFPY l
of operation.
3.3 Determination of J-R Curves t
l Figures 3-4a and b show the J-R ;urves for SA 302B material at 40 and 35 ft-lbs USE levels. The mean minus two standard deviation J e values fer t
the preceding 'JSE levels are 170 and 150 in-1b/in2, respectively.
The corresponding mean values are 340 and 315 in-lb/in2, respectively.
Figure 3-4a shows the mean and 3-4b shows the mean minus two standard deviation i
j J-R curves.
The J-R curves flatten out at a J-integral valua equal to 1.3 l
times J]c, as described in bbsection 3.1.1.
]
l As discussed in Subsection 3.1.2, the J-il curves for SA 302B modified i
material can be calculated in two ways (1) using pre-irradiation Charpy f
USE and fluence level, and (2) irradiated Charpy USE.
Figure 3-5 show; the-
}
results of mean or best estimate J-R curve calculations for plate G-3 7.
The irradiated Charpy USE J-R curve was calculsted using 35 ft-It's energy.
The other J-R curve was calculated using a CVNp of 51.4 f t-lbs and a fluence level of 1.3x1019 n/cm2, it is seen that the J R curve based on j
irradiated Charpy USE is lower. Therefore, all of the J-R curvo 3-5 4
m
1 l
l j
calculations for SA 302B modified material were conducted using only the j
irradiated Charpy L'SE.
Figure 3-6a and b show the mean and mean minus two standard deviation J-R curves for the 40 and 35 f t 1bs irradiatet' Charpy USE levels, j
respectively.
A comparison of the J R curves in Figures 3 4 and 3-6 shews j
that the J R curves in Figure 3 4 (SA 3028 material) are lower than those j
in Figure 3 6 (0A 3028 modified or SA 533B).
For completeness, the J.R
)
curves for both the materials are shown in the evaluation of operating j
conditions presented in Sections 4 and 5.
I i
3.4 Summary of J-R Curve Determination J+R curves were developed for two Charpy USE levels (40 and 35 l
ft-lbs).
The first one corresponds to the lowest predicted USE level for
]
any beltline RPV plate at 32 EFPY and the latter energy level represents a i
boundirig case.
Although the material for plate G-8-7, which shows the t
{
lowest predicted USE, is SA 3020 modified, one other beltline plate of SA
}
3028 traterial is also predicted to have transverse USE below 50 f t-lbs.
Therefore, J-R curves are determined for both the SA 302B (Figures 3-4a and b) and the SA 3028 modified (Figures 3-6a and b).
The mean minus two f
stancard deviation J R curves are to be used in the evaluation of Levels A l
through C loadings, and the mean or the best estimate J-R curves are to be i
l used in evaluating the Level 0 loadings.
i l
3.5 References i
I l
3-1 J.R. Hawthorne, B.H. Henke, F.J. Loss, H.E. Watson, A.L. Hiser, and R. A. Gray
- Evaluation and Prediction of Neutron Embrittlement in l
Rcactor Preisure Vessel Materials," EPRI NP-2782, December 19C2.
32 A.L. Hiser and D.B. Fishman, "J-R Curve Data Base Analysis of Irradiated Reactor Pressure Vessel Steels," EPRI Research Project 1757-24 HEA-2024,-December 1983.
i 3-6
1 i
b l
3-3 A.L. Hiser and J.B. Terrel. " Site Efftc.t on J R Curves for A J02 8 j
Plate." NUREG/CR 5256. HEA-2320. January 1989, t
~
34 J.R. Sawthorne and A.L. Hiser. " Influence of Fluence Rate on l
Radiation Induced Mechanical Property Changes in Reactor Pressure 3
Vessel Steels." NUREG/CR 5493. HEA-2376. March 1990.
3-5 " Reactor Pressure Vessel Evaluation Report for Yankee Nuc' ear Power i
Station." Report # YAEC-1735. July 1990. Chapter # 3. Yankee Atomic 4
1 i
Clectric Co., Bo^.*6n, Mass.
J l
3-0 landes. J.D., "txtiapolation of the J-R Curve for Predicting Reactor
]
Ve% el Integrity." NOREG/CR-5650 January 1992.
j i
j 3-7 E.D. Eason niid E.E. Nelson
" Improved Model for Predicting J R Carves i
from Charpy Data." NUREU/CR-5356 April 1980, i
j 3-8 E.D. Eason. J.E Kright and C.E. Nelson "Hultivariable Modeliry of l
Pressure Vessel and Piping J R Data." NUREG/CR-5729. Hay 1991.
l 1
r s
i i
j 3-7 i
i i
, _ - _.. _. ~
. - - -. ~. - -
e TABl.E 3-1 Values of Variables for J R Curve Estimation - Base Metal RPV Base hietals Combined Database Parameter Variable CVN, hiodel Charpy hiodel InC/
ai (constant) 2.89 4.13 a3 InCVN or InCVN, 1.22 1.48 a3 T
0.00270
.00239 c,
$t 0.0104 G
d (constant) 0.0770 0.0770 i
d InC1 0.116 0.116 2
2 d.
(constant) 0.0812 0.0812 d,
inC1
-0.00920
-0.00920 Q4 (exponent) 0.417 0.455
- Points 2295 8463 S,
in units 0.150 0.229 Ratios 1.645 S, 0.781 0.686
-1 S.
0.861 0.795 2 S, 0.741 0.632 3 S, 0.637 0.503 3-8
I l
' ' ~ ~ "
l too soo v
l i
C too 5
han
/
i i
i so, n
o I
j%c o
3 h
I l
D Mean 2r 400 s
1
\\
l f 4
/
j l
~
l 11 pd soo
/
/
200
't
/
/1 10.
I o
o 10 20 30 40 50 so 70 80 so 100 Cv, ft lbs Figure 3-1 Least Squares Fit Mean and Mean Minus Two Sigma Lines, SA302B Plate, LT and TL Directions. 400-550' F (From Reference 3-5)
J 3-9
1400 1/27CT - 1/2 inch specimes 6 inch specimen p
1200 etet
\\
/
1000 k
Y 600 l
e l
400 4
-f 400
't
'l too I
o.co e.es e.10 0.1 s e.2e o.25 Crack tutension ( A -a), inch -
Figure 3-2 J-R Curves Illustrating Size Effect for SA 302B Plate, CVN Energy = 52 ft-1bs. 180' F 5
3-10
d i
1 I
et i
3 P~~~~-~1-____7
,l 4 *-
j
$ PO %f S 4880 H,is agoogggiom abattlig I
l l
\\
u 1
J B +6 me iaC6viso% t,%g I
i we I
g I
Lr l
a E en e
i I
l i
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Po*ta Lavv neoneSaios Lawa I
.Dj I
g I
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O3 m m c#ptgt gigg 10 mm gaCLUSiO4 Litt
/
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t t
i i
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!", u,...
t i
I !
I 1
I I
t i
- a em n
i.
in i.
CAAta tatthlion imm Figure 3 3 Schematic of ASTM Frocedure for J e Determination I
I 3-11 i
.___--_..____.__.m f
f J
i 1
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+
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(3 NI/En-NI) PF-r 3-12
l!Ij a
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o S
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+
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0 0
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3 2
1 0(EN31ZO D7 wL*
i ll llll li l:(il ll l
l
.l
J-R CURVES FOR 302B MOD.
@ 35 FT-LB USE MEAN VALUES -NUF{G/CR 5729 1.3 U
1.2 -
g_p 1.1 -
Y
/
1 -
O.3 -
0 0.8 -
(
- "i -.
Z, pr,
l 41 NF l
o7_
38 i
I w
l 3 c'
Zo 0.6 -
(
a
-K b
0.5 -
y 7
[
0.4 -
0.3 -
l 0.2 -
01 -
O -U i
i i
i i
i i
O O2 04 06 08 1
DELTA o IIN )
O CFRRPY,p
+
CHiF P (.i Figure 3-5 Mean J-R Curves for SA 302B Mcdified Material
[ __
~
t J-R CURVES FOR 302B MOD.
@ 40 FT-LB USE MEAN AND MEAN -2*SIG,tA VALUES 1
?
O.9 -
_a_-& PU~O l
JS-t 4
t s
b O.7 -
f I
d mN
(
O.6 -
n Zo l
u A,
a r
1 C
r r
8 0.5 -
l 3 i
w ZO v.c g
O.4 -
t i
O.3 -
I O.2 -
i i
O.1 -
i O O y
0 02 04 06 O8 1
1 DELTA o Iif1)
O MEAN
+
MEAN -2'51GMA Figure 3-6a Mear,and Mean Minus Two Sigma J-R Curves for SA 302B r
Modified Material at 40 ft-lbs USE l
IlIll 1l I
l v
v U
~
1 E
G
+
S
~
U 4
i
+
u B
}r p
L 8
T i 8 2
t 0
0 3
mF 9
A
+
S ro 5
4_
i f
3 se
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A u C
U M
E L
6 G R S A
i I
- U I
W i
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S J
.V s
DA 2 a b ml A
OG g -
i t
)l S f S
. l M'
NA I
o 5 2
M T o
t Bl1 9
s a X
A u
A T
n l
2E L
i a 4
E+
M i M
+
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0 r
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3l a t J
e a A
p M M 1
d d n e RA l
a i E
f OM n i N a d e o F
A
/
, O M M E
M b 6
S 2
3 E
e O r V
a u
R a
g i
F U
C f
i e
R I
t j
i O
l J
0 0
0 0
0 0
0 0
O 0
0 0
0 0
0 0
0 8
7 6
5 4
3 2
1 Rikbb V*
l 1lt ll l
4.0 EVALVATION OF LEVEL A AND 8 CONDITIONS 4
i The methodology for the evaluation of Level A and B Service loadings l
was described in section 2.
Key steps in that evaluation are the calculation of applied J-integral and the flaw stability evaluation.
This Section describes the selection and the evaluation of appropriate loadings j
for Service Levels A and B.
The J-R curves determined in Section 3 are used to determine if the acceptance criteria of Appendix XX are satisfied.
4 l
4.1 Level A and B Service L~ dings 4
The loadings for the Oyster Creek RPV are specified in Reference 4-1 and were used in the ASME Code analysis documented in Reference 1-1.
A review 6f Reference 4-1 indicates that the Service Level A and B loadings are essentially internal pressure and a heatup/cooldown rate of 100* F per hour.
The RFV design pressure is 1250 psig.
The numerical values of j
pressures, cooling rate, material strength properties and the beltline geometry used in this evaluation are summarized in Table 4-1.
The postulated flaw for this case is a 1/4 t (i.e., 7.125x.25 or 1.78 inch) and its orientation is circumferential.
The internal pressures and the cooldow.i rate shown' in Table 4-1 along with the flaw geometry information were input in equations (2-1) through (2-7) of Section 2 to j
determine the applied J-integral values.
4.2 Evaluation per Jo,1 Criterion l
Table 4-2 shows the calculated values of applied J-integral for internal pressure equal to 1581 psi (1.15x Accumulation pressure) and a cooldewn rate of 100* F per hour.
The applied J-integral value at 1.88 inch crack depth (1.78 + 0.1) is 63.8 in-lb/in2, i
4-1
w
. - - =
Figure 4-1 shows the plot of this J-integral value along with the mean minus two sigma J-R curves (40 ft-lbs USE)_for SA 302B and SA 302B f
modi fied.
It is seen that the applied J value of 63.8 is considerably less l
than the Jo,1 value for both J-R curves.
Figure 4-2 shows the same comparison with J-R curves corresponding to 35 ft-lbs USE.
In all cases the Jo,1 criterion is satisfied.
l_
4.3 Stability Evaluation i
The applied J-integral values for the stability evaluation are calculated using an internal pressure of 1719 psi (1.25x Accumulation l
pressure).
Figure 4-3 shows the applied J curve and the appropriate J-R curves for 35 ft-lbs USE.
Flaw rtability at a given applied load is j
assured when the slope of the applied J-integral curve is less than the j
slope of the material J-R curve at the point on the J-R curve where the two curves intersect (see Figure 2-1).
This condition is clearly satisfied i
even for the 35 ft-lbs USE case.
i l
l To determine the pressure at which the instability occurs, applied
}
J-integral curve was recalculated for a range of internal pressures.
With l
SA 320B modified J-R curve for 35 ft-lbs USE, the instability was predicted for an internal pressure of 3250 psi, well in excess of any credible j
pressures in BWR operations.
The stability assessment curve corresponding to 3250 psi pressure is shown in Figure 4-4.
4.4 Summary of Service Level A and B Evaluation 4
i The evaluation for Service Level A and B loadings was conducted with the conservative J-R curves for SA 302B and SA 302B modified materials.
The results of the evaluation show that even for the case with assuned USE of f
15 ft-lbs and the SA 302B J-R curve, both the Jo,1 and stability criteria are satisfied.
a 4-2 4
d J
y m--+c r.--
m-,,
9--
..e~..+.,--a e a v9*'--=te
-at--ww-w
-t-=mw'e w
---+^m--~v" w
=-weFM'm---*r'*-P e
Fd-'e m
l 4
4.5 References 4-1
" Reactor Vessel Loadings," GE Drawing No. 237E438. Revision 1, 4
December 1964.
4 i
4 4-2
" Reactor Cycles - BWR-6 Standard," GE Drawing No. 795E949. Revision 0, July 1981.
i 4
i i
s 4
4 k
i
)
1 i
i e
i J
J k
I i
i e
5 i
i i
4 4
6 e
I a
4-3
}
3 e
1
~
r
=m,e.
.e--
,,w.m
- v. ww-+
,<ww++
r-e v
,v,,e-,
.,-uw w < w-e-v y.sc i w -
w--ee-m ew -- r
_. - -. =..
i l
TABLE 4-1 RPV Geometry and Loading Information l'
Vessel Geometry for Beltline Region
{-
Vessel Radius 105.7 in.
Vessel Thickness 7.125 in.
l Plate Analyzed G-8-7 Flaw Orientation Circumferential I
j Conditions for level A & B Evaluatiort l
Pressure (design) 1250 psi l
Accumulation Pressure 1250x1.1 or 1375 psi 1
1.15xAccum. Pressure 1375x1.15 or 1581 psi 1.25xAccum. Pressure 1375x1.25 or 1719 psi j
Temperature 550' F l
Young's Modulus 27.7x106 psi l
Poisson's Ratio 0.3 Cooldown Rate 100* F per hour l
Unitradiated Properties of Plate-G-8-7 i
Yield Strength 67.2 ksi 0 room temperature ll U.ltimate Strength 88.6 ksi Transv. USE 51.4 ft-lbs i
l Irradiated Material Properties 4
i 0 32 EFPY a.
e.
Fluence (1/4t) 2.36x1018 n/cm2-S 0 550' F 69 ksi (estimated) y CVN Energy 40.3 ft-lbs 4-4 i
a
-e, er s,
,,n-
= - _ -. -...= -_.-
6 1
I i
TABLE 4-2 i
i Calculated Values of Applied J-Integral for 1.15xAccumulation Pressure Case I
1 i
- 55LE!;;5
- .
IE31 viii!. 4-
- s;.
.6'
.!53E. - :N;.
? Zi
- .:53 34*:': -4
!!. N:=
't;25
.Mi.
?
g
, :(;.
3 l
4 1
f i
i 0140VMFi1ENTIAL Flaw 4
a F2 F3 Ko Kt ae F2*
F3' Ktetal Js:p e
{
I 78 0.96 0,90 30.E3 12,1B 1.80 0.97 0.90 43.02
-60.81 1.83 0.97 0.90 31.15 12.19 1.85 0.97 0.90 43.55 E2.30 1
1 53 0.97 0.90 31.56 12.18 1.90 0.97 0.90 44.06 E3.79 j
4.93 0.97 0.9D 32.18 12.18 1.95 0.97 0.90 44.57 65.27 1.99 0.99 0.90 32.E9 12.16 2.00 0.98 0.90 45.08 -
E6.75 2.03 0.98 0.90 33.20 12.15 2.05 0.98 0,9D 45.57 69.23 2.08 0.93 0.89 33.71 12.12 2.10 0.9S 0.89 46,as -
69.70 2.13 0.99 0.89 34.22 12.10 2.16 -
0.99 0.89 46.54 71.17 1
2.18 0.90 0.S9 34.73 12.06 2.21 0.99 0.89 47.02 72.64 2.23 0.99 0.39 35.24 12.03 2.26 0.99 0,89 47.49. 74 10 l
2 28 0.99 0,88 35.74 11,9B 2.31 1.00 0.$8 47.96 75.57 i
2 33 1.00 0.8B 36.25 11.94 2.36 1,00 0.68 48.42 77.03 l
2.38 1.00 0.69 36.75 11.89 2.41 1.00 0.SB 45.88 78.49 i
2.43 1.00 0.87 37.25 11.83 2.46 1.01 0.87 49.33 79.95 2.4B 1.01 0.87 37.76 11.77 2.51 1.01 0.87
-49.73 81.40 1
2.53 1.01 0.86
-36.27 11.71 2.56 1.01
-0.36 50'22 82.66 2.53 1.01 0.85 38.77 11.64-2.61 1.02 0.86 50.66 8031 j
2.53 1.02 0.85 39.28 11.57 2.66-1.02 0.85 51.10 85.77 2.S~.
1.02 0.B5 39.79 11,49 2.71 1.02 0,84 51 53 87,22 2.73 1.03 0.84 40,29 11.41 2.76 1.03 0.84 51.96 88,E8 j
2.73 1.03 0.84 40.80 11.33 2.81 1.03 0.83 52.38 90.14 l
i 1
4 i
d k
4-5
LEVEL' A &
B EVAL.,1.15xACCU PRESSURE 40 FT-LB USE, MEAIl - 2'SlGMA J-R CUfivES 600 U'
e-e- -e4 e-fM-~~~
(
O' 500 -
[
,V i
400 -
1 m
N T
i
(
Z i
t*
-1 i
\\
'3 300 -
l g
6 v
]
,-:lll l l l l :
200 -
r
,a a
,o
- O i
i 100 -
O []
i i
i i
i i
i l
0 0.2 0.4 0.6 O8 1
i DELTA o (111)
O 3028 M OD.
t 3028 O
J-APPtlED
.i I
Figure'4-1 Evaluation Based on Jo,1 Criterion Based 40 ft-lbs USE
LEVEL A &
B EVAL.,1.15xACCU. PRESSURE 35 FT-LB USE, MEAN - 2'51GMA J-R CURVES 500
+
"~
W a-e-B 400 -
~
E
{
300 -
i 5
3 I
Z
.4 O f
200 -
' 4 444 : : : : : :
II n
I 100 -
0 O dj i
i i
i i
i i
i O
O2 0.4 06 0.8 1
i DELTA o l'IN )
O 2028 MOD.
1 3028 O
J-APPLIED Figure 4-2 Evaluation Based on Jo,1 Criterion Based 35 ft-lbs USE 1
LEVEL A &
B EVAL.,1.25xACCU. PRESSURE 35 FT-LB USE, MEAN - 2'S!GMA J-R CUfNES i,
500 g_ g 4 --n--B~0~ U ^
W A
x 400 -
t
,s 04 i,
300 -
6:
N 3
Z t
a O
[)
200 - ',,;3;;; ;;:. ;;
4 y
~)
100 -
r g
_4__-g c
r,
~
v g __ __ H O
G v
L I
I I
i l
i i
i l
x 0
0.2 0.4 0.6 0.8 1
i DELTA o (IN.)
~
O 3028 MOD.
+
3028 O
J-APPLIED Figure 4-3 Stability Assessment for Level A and B Loadings Based on 35 ft-lbs USE
LEVEL A &
B EVAL., 3250 PSI PRESSURE 35 FT-LB USE MEAL 1 - 2' SIGMA J-R CUfNES 500 O~G~
MV 400 -
U fr',g:#
i, 300 -
=
N 7#
1
,bb
[
200 -
- . rT.9 7.
y
/
(
l l
100 -
0 Il i
i i
i i
i i
e i
O O.2 0.4 0.6 08 1
DELTA o (111) 0 2028 MOD.
t 3028 O
J-AFPLIED Figure 4-4 Stability Assessment at 3250 Psi Pressure
i 1
4 5.0 EVALUATION OF LEVEL C AND D CONDITIONS i
i The Appendix XX procedures call for the evaluation of Service Level C f
and D loadings with a safety factor of 1.0 and a postulated flaw equal to i
1/10 of the vessel wall thickness.
The ASME Code of construction of the l
Oyster Creek RPV did noc have the Service Level classification for various leadings.
The later editions of the Code first introduced the Normal, i
Upset, Emergency and Faulted classification for the various plant transients and component loadings.
To avoid corfusion between the plant or j
system opt.ating condition: and the component operating conditions, this classificat1on was then changed for the components to Service Levels A l
through D.
]
Since the RPV loadings drawing [4-1] does not define Level C and D l
condition loads, guidance was taken from more recent RPV thermal cycle diagrams, such as Reference 4-2 for a BWR/S standard plant, to selet t the appropriate transients. Once the transient is selected, the first step in j
the evaluation is to determine the throughwall stress distribution in the i
RPV wall when the stresses reach their peak.
This was done by finite element analysis.
The stress intensity factor, K,
values and I
correspondingly the applied J integral values are then-calculated using available handbook solutions for circumferential cracks.
The handbook j
approach is used since Appendix XX does not provide procedures to calculate K for temperature transients where heatup/cooldown rate exceeds 100* F per hour.
i 4
j 5.1 Evaluation of Level C Conditions l
5.1.1 Transient Definition I-A review of Reference 4-2 indicates that-among the transients j
specified for the Emergency (Level C) condition, Automatic Blow Down with Loss of High Pressure Feed (Event 23) is relevant for the Oyster Creek RPV.
5-1 1
a.
t
=
-vr.w,-r,-
w-ve-
--e=a*
+ -
- 4 is a more rapid blowdown than that occurs with a single relief or safety 4
l valve blowdown, a transient already postulated in Reference 4-1.
The temperature change in the beltline region during this transient is shown in i
Figure 5-1.
The temperature is assumed to drop to 375'_F in 3.3 minutes or 200 seconds.
The temperature is then assumed to drop to 259' F at the rate of 100* F per hour.
The internal pressure in the beltline region j
theocyhout this event is the saturation pressure corresponding to fluid temperature.
Thus the pressure at the end of 3.3 minute ramp is 170 psi, i
5.1.2 Finite Element Stress Analysis i
j Figure 5-2 shows an axisymmetric finite element model of the RPV wall j
in the beltline region.
The stainless steel clad on the ID surface, with a f
nominal thickness of 7/32 inch, is also included in the model.
ANSYS computer program (5-1] was used in Loth the transient temperature and the l
stress analyses.
The value for the convective heat transfer coefficiant, 2
h, at the ID surf ace was assumed as 10,000 Btu /hr-f t
- F, based on a l
previous analysis [5-2] of a more severe transient.
t The transient temperature distributions were calculated at several time points along the transient which were then used in the subsequent stress analysis.
A review of the stress distributions a^. different time l
points showed that the stresses reach a maximum at time ecual to 200
{
seconds.
Figure 5-3 shows this stress distribution.
The increased stress level in the clad (over and above the extrapolated trend from the base i
i metal stress) is due to the difference between the thermal expansion coefficients of low alloy steel and stainless steel.
This additional thermal stress in the clad was approximated as a point forca for the calculation of stress intensity factor, K.
5.1.3 Fracture Mechanics Evaluation The geometry of the postulated flaw for the Level C service loadings 4
is essentially the same as that for the Level A and B loadings except that-the flaw depth is 1/10 of the base metal wall thickness. The flaw i
5-2
~
4
-r-
. _, ~, _.
,-.m-
_m.
..y.,-,.._,_..,.,,,,_,,_,-,,.a3...
I L
i 4
orientation is circumferential.
For the calculation of applied K (and 4
correspondingly, J) values, a fully circumferentici (i.e., 360* around) l flaw geornetry was conservatively assumed, thereby it was possible to use a standard approach given in Reference 5-3.
The stress distribution in this approach is characterized in the form of a third order polynomial across the thickness:
o=ao + a1 x + a2 X2 + a3 x3 (5-1) l The stress in the clad over and above that which would be present based on the extrapolation from base metal, was integrated over the clad l
thickness and the resulting force, P, was assumed to be located at the
{
middle of the clad.
The following equation based on a solution given in Reference 5-4 was used to calculate the K contributed by the force P:
1 Kelad = 2P x 1.3// (na)
(5-2) i j
where, a is crack depth.
i l
Table 5-1 shows a summary of the calculated values of K and l
J-integral.
Both the pressure and thermal loadings are based on a safety factor of 1.0.
The last but one column (K' total) in Table 5-1 shows the e
total value of K including the plastic zone size correction.
Figure 5-4 shows a plot of K' total as a function of crack depth / thickness.
The last column in Table 5-1 shows the applied J-integral values obtained from K' total values by using equation (2-6).
The J-integral values are then used in the acceptance criteria evaluation.
5.1.4 Acceptance Criteria Evaluation i
Figure 5-5 shows the applied J-integral curve, anc the-J-R curves for SA 302B and SA 302B modified at 35 ft-lbs USE.
The 'J-R curves for this evaluation are conservative representations (i.e., mean minus two sigma values).
The J-integral value at 0.1 inch crack growth, J was obtained 1
as 64.98 in-lb/in2 from Table 5-1 at 'a' = 1.031 inch (which is 0.1x7.125 +
5-3
1 1
clad thtekness of 7/32 + 0.1).
This value is clearly less than the Jo,1 values for either of the J-R curves in Figure 5-5.
Therefore, the first criterion in Subsection 3.1.3 is satisfied.
Figure 5-5 also demonstrates that the stability criteria is also satisfied since the applied J-i, %qral i
value is less than Jge predicted by both J-R curves.
l Based on the preceding, it is concluded that the both acceptance j
criteria for Level C loadings are satisfied.
i 5.2 Evaluation of Level D Conditions l
As in the case of Level C loadings, there are no Level D condition loadings defined in the Oyster Creek RPV loading drawing.
A review of more i
recent RPV thermal cycle diagram [4-2] shows that the Loss of Coolant Accident (LOCA) event is the most limiting among the Level D events.
Therefore, this event was considered in the evaluation for Level D acceptance criteria.
4 A fracture mechanics evaluation of a BWR RPV following a postulated LOCA event has been described in Reference 5-2.
The analysis in Reference j
5-2 considered a 240 inch ID RPV with a thickness of 6 inches.
These I
dimensions differ slightly from those of Oyster Creek RPV, but it was judged that these differences are insignificant and, therefore, the results l
of that analysis were used in this evaluation.
}
l 5.2.1 Description of LOCA Event i
Two types of pipe rupture events can be postulated to cause a LOCA:
(1) steam line break (2) recirculation line break.
Both events assume a j
guillotine rupture of the line when the reactor is operating at full power.
j Following pipe rupture, depressurization occurs rapidly in both cases.
During steam line break, because of its higher elevation,-the annulus between the shroud and the beltline region continues to. be filled witF l
two phase mixture of water and steac, and the boiling continues for' qui.e some time after the initial rapid depressurization.
On the other hand, i
5-4
}
d
,,..,...-,,,,,--.-,,,._,...--,-,,,-,,,w
,,y..e.,_--,-,.-,,wnne.--,,,e
,-,yrew-7,..,.,_,,-
l i
i with a recirculation line break, the water in the annulus drains out after the initial depressurization and the beltline region is exposed to steam l
under natural convection heat transfer conditions.
Therefore, higher heat transfer conditions, and consequently, higher temperature gradients are expected to develop in the case of steam line break.
It was, therefore, j
concluded that the steam line break is more severe than the recirculation l
line break from the view point of thermal stresses and fracture failure mode.
Therefore, steam line break was analyzed in Reference 5-2.
I i
Figure 5-6 shows the pressure and temperature variations assumed during the LOCA event.
Based on the consideration of thermodynamic and heat tr3nsfer conditions, the convective heat transfee coefficient, h, during the depressurization phase (time 0 to 300 seconds) was i
2 conservatively assumed as 10000 Btu /hr-ft
- F.
After 300 seconds, there is i
significant subcooling from the ECCS flow and, therefore, the value of
'h' 2
i is much lower.
A value of 500 Btu /hr-ft *F was assumed for that portion of the event.
5.2.2 Fracture Mechanics Analysis Results 4
Figures 5-7 and 5-8 show tne calculated temperature and stress l
distributions, respectively, in the vessel wall at different times after the initiation of LOCA event.
The calculated values of stress intensity l
factors based on the stress distributions of Figure 5-8, are shown in Figure 5-9'.
A 360* circumferential crack geometry was assumed-in these calculations.
It is seen that the stress intensity factor values are the l
highest at time equal to 300 seconds.
The stress intensity factor value at a/t of 0.1 is approximately 90 ksi/in.
This is equivalent to a applied J-integral value of 266 in-lb/in2, i
i 5.2.3 Acceptance Criteria Evaluation i
The material J-R curves to be used for the evaluation of Level D.
j loadings are those based on the best estimate or the mean values.
Figure 5-10 shows the mean or best estimate J-R curves for J ft-lb USE and the j
5-5 i
.,-n.
- -..,.--._,,,, ----.. --.--__,., m,,
.-,.,.~.....--,,,.-.e.,
calculated value of applied J-integral for Level D loading.
It is seen j
that the value of applied J-integral, 266 in-lb/in2, is even less than the Jc value for the lower of the two J-R curves (315 in-lb/in2).
Thus, both I
the Jo,1 and the stability criteria are satisfied.
5.3 Summary of Level C and D Loadings Evaluation The evaluation for Service Level C was conducted with the assumed loading as that resulting from an Automatic Blowdown Transient.
The applied J-integral values were based on a finite element stress analysis of Oyster Creek RPV wall. The mean minus two sigma J-R curves for SA 302B and l
SA 302B modified materials were used in the acceptance criteria evaluation.
i For the Service Level D evaluation, results from a previous fracture mechanics analysis of LOCA event were used to determine the _ applied J-integral values. The mean or best estimate J-R curves were used for this t
case.
The results of both the evaluations show that even for the case with assumed USE of 35 ft-lbs and the SA 302B J-R curve, both the Jo,3 and l
stability criteria are satisfied.
i 5.4 References 5-1 ANSYS Computer Program. Version 4.1, Swanson Analysis Systems, Inc.
5-2 Ranganath, S., " Fracture Mechanics Evaluation of a-Boiling Water Reactor Vessel Following a Postulated Loss of Coolant Accident," Paper No. G 1/5, Transaction of the 5th SMiRT Conference, 1979.
i i
l 5-3 C.B. Buchalet and W.H. Bamford, " Stress Intensity Factor Solutions for Continuous Surface Flaws in Reactor Pressure Vessels," ASTM STP 590, 1976, pp. 385-402.
l 5-4 H. Tada and P.C. Paris, "The Stress Analysis of Cracks Handbook," Del j
Research Corporation, 1985.
5-6 i
h 1
.. ~. _, ~. _ _ _,.. _ _.,.-
TABLE 5-1 Applied J-Integral Values for Level C Loading a
f Irternal Clad Coef ficients for Ceefftetents for Pressure Force Thermal 5 tress Pressu*e $ttess (psig)
(Ibs)
AO:
29800 AO: 24070.26 170 1543.9844 A1:
-21730 41t 0
A2:
4242 A2:
0 A3:
-267.3 Ai:
0 a
Kth Kelad KD Ktotal se Kth' Kelad'
.Kp' Ktotal' Japp 0.000 0.0000 0.0000 0.0000 0.0000 0,0000 0.0000 0.0000 0.0000 0.0000 0.00 0.500 34.2232 3.6379 1.5891 39.a503 0.5173 34.5660 3.5597 1.6177 -39.7434 51.89 0.734 37.7102 2.8760 1.9457 42.5319 0.7545 37.9149 2.?307 1.9'41 t2.7196 59.95
.4 0.931>39.2444 2.5080 2.2111 43.9635 0.9528 39.3573 2.4757 2.2388 44.0719 63.89 1.000 39.5734 2.4092 2.2988 44.2814 1.0218 39.6595 2.3802 2.3263 44.3650 64.66 1.031 39.6940 2.3680 2,3380 44.4001 1.0532 39.7688 2.3403 2.3653 44.4745 64 9e 1.500 39.9634 1.9281 2.8883 44.7797 1.5223 39.9243 1.9128 2.9132 44.7503 65.79 2.000 38.4939 1.6536 3.4304 43.5778 2.0212 38.4123 - 1.6444 3.4528 43.5095 62.10 2.500 36.3855 1.4705 3.9544 41.8104 2.5195 36.2987 1.4646 3.9747 41.7380 57.23 3.000 34.1314 1.3373 4.4762 39.9448 3.0178 34.0508 1.J332 4.4948 39.f.i88 54.25 3.500 31.8534 1.2348 5.0056 38.0937 3.5162 31,7787 1.2318 5.0229 138.0335 47.52 4.000 29.4625 1.1527 5.5497 36.1649 4.0146 29.3892 1.1505 5.5658 36.1055 42.83 4.500 26.7450 1.0851 6.1138 33.9438 4.5128 26.6685 1,0835 6,1205 33.b605 37.71 5.000 23.4115 1.0281 6.7020 31.1417 5.0108 23.3302 1.0210 6.7150 31.0/32 31.12 5.500 19.1231 0.9793 7.3179 27.4203 5.5084 '19.0410 0.9785 7.3265 21,1489 24,57 6.000 13.4991 0.9368 7.9645 22.4004 6.0056 13.4271 0.9364 7.9719 22.3353 16.39 5-7
W 4
3.3.nin u tes i
528 F t
\\
\\
1 0
D M
\\
375 F b
300 F/ hour
\\
259 F
\\
I
(
Time 4
Figure 5-1 Temperature-Time Variations During Automatic Blowdown Transient (Level C Condition) 5-8
l l
1 i
i inside Radiu s 106.7"
>- - - Low Allc* Jteel
-H
,Le-va2 ss ci.oaine tu i
f i
F..jc inside i :.;:-
Radius I
llyi
!!I 1 ;
t - 11 Iq.l.
l
.l l
l I
. r'f:
'y
-- - _ _ 7 gg43ge i
.y i
Figure 5-2 Axisymmetric Finite Element Model of RPV Wall 5-9
.~ ~ - - -. - ~..
i
)
1 j
j i
4 I
1 ANSYS S.44 JU.L 7
1992 1
- 43 4
kk N17 i
53 ITERs1 i
45661 PATH PLOT nod 1:1 gD2=31 I
STRESS CLOBAL 49439 l
1 20
=1 DjST:g!g666 Ur :. S asz2a i
2r
=0.5 29996 i
i l
l
{
24775
\\
19553
\\,
\\
\\
14332 i
~\\
l i
l 9118 f
3888 i
1 l
-1333 ss.
'~~s,
-6553 EI T 9
- 1. 460 2.93C.
4.49d 5.871:-
7.34 e 734 2. 2 Ea2 3.67 5.138
-6.696 OYSTER CREEN MPV UFPER SHELF EMERGY ANALYSIS -
i Figure 5-3 Stress Distribution in RPV Wall at 200 Seconds 4
3 l
n 4
i 5-10
y
.r Oyster Creek itPV I!pl er Shelf Energy Analysis 45
[
/
\\
40 1
N 35 3
30 9
<c 5
20 5
\\
i x
20 b
15 l
10 5=
l 1
0 0'
O.2 '
O.4 0.6 08 1
a/t Figure 5-4. Plot of K' total as a Function of Crack Depth / Thickness
-1 i
ui..
LEVEL C LOADING EVALUATION 35 FT-LB USE, MEAN - 2' SIGMA J-R CUFiVES 500 O -- - B ~ d l - G - P F U o-P 400 -
- c4i, 3CO -
e N
I 3
i i
Z
[
ui 200 ~
-H P. '. ' '. ' '.
i, e,
i, i
m y
7 1
i r
f' i
4 1 0 0 --
o 9
1 i
O U l
0 0.2 0.4 06 0 13 1
DELTA o (111.)
O 3028 MOD.
t 3028 O
J-APPLIED I
Figure 5-5 Jo,1 and Stability Assessment Diagram for Level C Loading s
't e
i 1200 -
PRESSURI VARIATION
<>1050 psig DURING LOCA 4
l 1 800-a w
4 n
y 400 - t 4
1 120 psis 35 psig
~
sec.
t = 450 sec 0 psi @ l hr.
200 400 600 800 1060 TIME Sec BULK TEMPERATURE o
a 552 F DURING LOCA o
500 w
s 350 F @
H 305 sec.
h300 18 J81 F @ 450 sec.
j 212 F @ 300 sec.
H o
100 F @ 2 hrs
- l 100 I
I I
I I
I I
i 200 400 600 800 1000~
TIME SEC.
i i
Figure 5-6 Va.iation of Pressure and Temperature Assumed in.
the LOCA Evaluation i
5-13 v,~
e
---m-w--,m---w,-+
4.-- - -..,.,, ~ -, +-
.. ~.... - -. -. - -..
l J
l t
o00 i
i
' T = 50 sec.
-T 500 -i 150 300 i
i 100 g
i 450 o
I k
p T = 600 sec j
y400 -
AFTER LOCA t
I L
- 300,
,8 i
8 STAINLESS STEEL CLAD 0.2 in.
l fIi
-+-
l 200 O
1 2
3 4
5 6
j.
DEPTH, IN.
t l
l Figure 5-7 Tem'ierature Distribution in Vessel Wall at Different Times After the LOCA t
4 4
1 5-14
s 100 ~
90' 30 t=
100 see
,g G$0h j$0$'
t=
300 see s :.0 x
5 M
c - 500 see 30 20 a
s,
\\,
y 10 -
's i ?.
i i
e
% CLAD N\\
+
THICKNESS 10 -
,, % ~.
20 -
I t
t 0
1.0 2.0 3.0 4.0 5.0 6.0 DEPTH Inches Figure 5-8 -Thermal Stress Distribution at Different Times Following the LOCA 5-15
--m_--
- - - - - ^ - - - - - - - " ' - - - ' - - ' ~
j i
l i
I I
i 200 U';I fd.\\DIATED 1
-l1 i
ISO MINIMLH AVAILABLE TOUCHNESS 3
IN BASE MATERIAL
'2 l
- 200 Ksi-Vin 160 3
STAINLESS STEEL 140 CLAD, 0.2 inches C
i 5
120 j
- T TAL APPLIED K AT l
300 sec. AFTER LOCA D
I 100 -1 E
I l
80 -l h
'.7 * "#, 4=* N'N==,,'**,
s 60 -
f,#
m 40, (i
.N 3
m l
t = 500 N
sec 20 l
-t=
100 see i
1 i
i 0.-
0.4 0.6 0.8 1.0 a/t Figure 5 Applied Stress Intensity Factors for Different Crack Depths at Different Times Following the LOCA 5-16
- - - - - - - - - ~ ~ ~ ~^~
t t
I'
..lt 5
,I 1
t 1
~
G
~
G 1
g 4
l 0
I n
i da 8
D o
E L
N F
L D
P l
O A
ev I
e T
l J
L A
ro f
US O
E m
LV ar AR 6
g U
l 0
a VC i
D ER
+
t ne J
)
m G4 48 s
f f
i s
A O2 e
NE 0
s o3 s
I M
A A
DE T
y L
t S
E AU I
4 G+
i l
OB 0
i b
LL S
t'
-T F
d D
l an 53 1,
L o
E J
V l
0 2
1 E
0 D
5 h
O L
M er u
B g
2 i
i 0
F 3
+
m p
O
<?
0 0
0 0
0 0
0 0
0 0
0 0
0 0
0 0
0 8
7 6
5 4
2 2
1 m0(Z \\3IZ y>
4 mLw e
!14 i
f iIii 4
t i
e 6.0
SUMMARY
AND CONCLUSIONS i
The P-T curve analysis for Oyster Creek RPV had indicated that the projected 32 EFPY upper-shelf energy for some of the RPV plates in the j
beltline region will fall below the 50 ft-lb value.
Based on this information, the NRC staff has requested an Oyster Creek specific analysis l
showing the basis for present and continued vessel structural integrity when the 50 ft-lb requirement of 10CFR 50, Appendix G is not satisfied.
A draft Appendix (Appendix XX) to ASME Section XI provides evaluation j
procedures and acceptance criteria for such cases.
This report presents a j
fracture mechanics evaluation of Oyster Creek RPV using these procedures j
and acceptance criteria.
Levels A through D loading conditions were analyzed.
For the Oyster Creek vintage vessels, the reactor loading diagrams did not define any Level C (emergency condition) and D (faulted condition) loadings.
1 Therefore, based on a review of later BWR RPV thermal cycle diagrams, Automatic Blowdown transient a.id the LOCA event were selected as the i
limiting loadings for Level C and D conditions, respectively.
1 The beltline plates that have projected USE below 50 ft-lbs are made l
of SA 302B and SA 302B modified (equivalent to SA533 Grade B).
Technical literature on fracture toughness testing of both the materials was reviewed to determine the appropriate J-Resistance or J-R curves.
Best estimate minus two standard deviation 'J-R curves were used in the evaluation of Levels A, B and C service loadings.
Level D loadings were evaluated using best estimate J-R curves. The J-R curves for SA 3020 material was found to-be always lower than those for the SA 302B modified material at the same i
USE-level.
The beltline plate (G-8-7) that has the lowest projected USE-(40.3 f t-lbs) is made of SA 3028 modified stock but there are other l
beltline plates, made of SA 302B, which are also projected to have USE 6-1 1
m,
..y-,.-sw-,y,,.-
.y
-w, r, v uy,,. -e,,,., -,
slightly below 50 f t-lbs.
Therefore. J-R curves of both materials were shown on the evaluation di? grams.
)
To cover the case of plant life extension, the evaluation was also conducted using a very conservative USE level of 35 ft-lbs (projected to be more than 160 EFPYs for plate G-C 7).
)
There are two acceptance criteria in Appendix XX.
The first one compares the applied J ntegral with J-R curve value at a flaw growth of 0.1 inch.
The second criterion cone: erns with flaw stability.
The
)
evaluation results showed that even for the conservative case of 35 ft-lbs USE both the acceptance criteria are satisfied fnr all four Service Level loadings.
'n Based on the results of this evaluation of Oyster Creek RPV per Appendix XX procedures, it is concluded that even though some beltline plates would have projected USE below 50 f t-lbs at the end of current design life of 32 EFPY, these lower values would still provide margins of p
safety against f racture equivalent to those required by Appendix G of Section Ill, ASME Code.
Thus, the Oyster Creek RPV will continue to meet the requirements of 10CFR50, Appendix G.
This conclusion would remain valid also for any realistic plant life extension that may be considered by p
the plant owner.
D D
y 6-2 D
APPE!4 DIX DRAFT OF APPEliDIX XX, SECTION XI. ASME CODE i
i DRAFT CODE CASE N-XXX ASSESSMENT OF REAC*OR VESSELS WIT!!
LOW UPPER SHELF CRARPY ENERGY LEVELS May 27, 1992 REVISION 11 DRAFT HISTORY REVISION O
AUGUST 25, 1987 REVISION 1
JANUARY 19, 1988 REVISION 2
APRIL 19, 1988 REVISION 3
AUGUST 30, 1988 REVISION 4
NOVEMBER 30, 1988 REVISION 5
FEBRUARY 27, 1989 REVISION 6
JANUARY 5, 1990 REVISION 7
APRIL 12, 1990 REVISION 8
JANUARY 10, 1991 REVISION 8 - MARKED COPY APRIL 15, 1991 REVISION 9
JANUARY 17, 1992 REVISION 10 APRIL 17, 1992 REVISION 11 CURRENT
- - - - - - - - - - - - - - ^ - ' - - - - ^ ^ ^
ASSESSMENT OF REACTOR VESSELS WITH LOW UPPER SHELF CHARPY ENERGY LEVELS TABLE OF CONTENTS CASE li-XXX ASS"SSMENT OF REACTOR VESSELS WITH LOW UPPER SHELF CHARPY ENERGY LEVELS APPENDIX A ASSESSMEllT OF HEACTOR VESSELS WITH LOW UPPER SHELF CHARPY ENERGY LEVELS l
A-iOOO INTRODUCTION 3
A-1100 Scope A-1200 Procedure Overview A-1300 General Nomenclature A-2000 ACCEPTANCE CRITERIA A-3000 ANALYSIS A-3100 Scope A-1200 Applied J-Integral A-3300 Selection of the J-Integral Resistance Curve A-3400 Flaw Stability A-3500 Evaluation Approach for Level A and B Service Loadings A-4000 EVALUATION PROCEDURES FOR LEVEL A AND B SERVICE LOADINGS A-4100 Scope A-4200 Evaluation Procedure for the Applied J-Integral A-4210 Calculation of the Applied J-Integral A-4220 Evaluation Using Criterion for Flaw Growth of 0.1 in.
A-4300 Evaluation Procedures for Flaa Stability A-4310 J-R Curve Crack Driving Force Diagram Procedure 2
A-4320 Failuro Assessment Diagram Procedure A-43?1 Failure Assesstaunt Diagram Curve A-4322 Failure Assessment Point Coordinates A-4322.1 Axial Flaws A-4322.2 Circumferential Flaws A-4323 Evaluation Using Criterion for Flaw Stability A-4330 J-Integral / Tearing Ho.ulus Procedure A-4331 J-Integral at Flaw Instability A-4332 Internal Pressure at Flaw Instability A-4333 Evaluation Using Criterion for Flaw Stability A-5000 LEVEL C AND D SERVICE LOADINGS f
a 3
)
Case N-XXX Assessment of Reactor Vessels With Low Upper shelf Charpy Energy
)
Levels section XI, Division 1 Inquiry:
Section XI, Division 1,
IWB-3730, requires that i
during reactor operation, load and temperature conditions shall be maintained to provide protection against failure due to the presence of postulated flaws in the ferritic portions of the reactor coolant pressure boundary.
Under Section XI, Division 1, what procedure may be used to evaluate a reactor vessel with a low
)
upper shelf Charpy impact energy level as defined in ASTM E 185-82 to demonstr&te integrity for continued service at upper shelf conditions?
Reply:
It is the opinion of the committee that a reactor vessel with a low upper shelf Charpy impact energy level may be
)
evaluated to demonstrate integrity for continued service for upper shelf conditions.in accordance with the following, 1.0 EVALUATION PROCEDURES AND ACCEPTANCE CRITERIA Section XI, _ Division 1, Appendix G,
" Fracture Toughness criteria for Protection Against Failure", provides analytical procedures based on the principles of linear-elastic fracture mechanics that may be used to define load and temperature conditions to_ provide protection against nonductile failure due to the presence of I-pestulated flaws in-the territic portions of the reactor coolant pressure boundary.
To prevent ductile failure of a reactor vessel with a low upper shelf Charpy impact energy level the vessel shall be evaluated using the principles of = elastic-plastic fracture mechanics.
Flaws shall be postulated in the reactor vessel at i
locations of predicted low upper shelf charpy impact energy and the applied J-integral for these flaws shall be calculated and compared with the J-integral fracture resistance of the material to determine acceptability.
Factors of safety on applied load for limited ductile flaw growth, and'on flaw stability due to ductile tearing, shall be satisfied.
All specified design transients for i
the reactor vessel shall be considered._ Evaluation procedures and-acceptance criteria based on the principles of elastic-plastic fracture mechanics are given in Appendix A of this code case.
The evaluation shall be-the responsibility of the owner and shall be subject to review by the regulatory and' enforcement-authorities having jurisdiction at the plant site, p
APPENDIX A TO CODE CASE N-III ASSESSMENT OF REACTOR VESSELS WITH LOW UPPER SHELF CHARPY ENERGY LEVELS ARTICLE A-1000 INTRODUCTION A-1100 SCOPE This Appendix provides acceptance criteria and evaluation procedures for determining the acceptability for operation of a
reactor vessel when the vessel metal temperature is in the upper shelf range.
The methodology is based on the principles of elastic-plastic fracture mechanics.
Flaws are postula*.ed in the reactor vessel at locations of predicted low upper shelf Charpy impact energy and the applied J-integral for these flaws is calculated and compared with the J-integral fracture resistance of the material to decermine acceptability.
All specified design transients for the reactor vessel shall be considered.
A-12OO PROCEDURE OVERVIEW The following is a summary of the analytical procedure which may be used.
(a)
Postulate flaws in the reactor vessel according to the criteria in A-2000.
(b)
Determine the loading conditions at the location of the postulated flaws for Level A, B, C and D Service loadings.
(c)
Obtain the material properties, including E, o,,
and the J-integral resistance curve (J D curve), at the locations of the postulated flaws.
Requirements for determining the J-R curve are given in A-3300.
(d)
Evaluate the postulated flaws according to the acceptance criteria in A-2000.
Requirements for evaluating the applied J-integral are given in A-3200, and for determining flaw stability in A-3400.
Three permissible evaluation approaches are described in t
A-3500.
Detailed calculation procedures for Level A and D Service loadings are given in A-4000.
A-1
A-1300 GENERAL HOMENCLATURE flaw depth which includes ductile a
=
flaw growth (in.)
effective flaw depth which includes
=
a, ductile flaw growth and a plastic-zone correction (in.)
affective flav depth at flaw a'
=
instability, which includes ductile flaw growth and a plastic-zone correction (in.)
postulated initial flaw depth (in.)
a,
=
amount of ductile flaw growth (in.)
Aa
=
Aa*
amount of ductile flaw growth
=
at flaw instability (in.)
material constants used to describe C,,
C,
=
the power-law fit to the J-integral resistance curve for the material, J, = C, (h a )
cooldown rate
('F/ hour)
(CR)
=
Young's modulus (ksi)
E
=
E/(1-v')
(ksi)
E'
=
geometry factors used to calculate F,,
F,,
=
F, the stress intensity factor (dimensionless)
F*,
F*,
geometry factors used to calculate a
F' the stress intensity factor at flaw instability (dhmnsionless)
J-integral due to the applied J
=
loads
( in. -lb / in. ' )
J-integral fracture resistance for J,
=
the material
( in. -lb / in. 8 )
=
A-2
l J-integral fracture resistance for J,, 2
=
the material at a ductile flaw growth of 0.10 in.
( in. -lb/ in. ' )
applied J-integral at a flaw J
3 depth of a, + 0.10 in.
(in. -lb/in. 3 )
J' 4
J-integral at flaw instability
( in. -lb /in. 8 )
mode I stress intensity factor (ksi Vin.)
K,
=
K,,
mode I stress intensity factor
=
due to internal-pressure, calculated with no plastic-zone correction (ksi Vin.)
Kr, calculated with a plastic-zone Kj,
=
correction (ksi Vin.)
Kj, K,, a t flaw instability, calculated
=
with a plastic-zone correction (ksi Vin.)
mode I stress intensity fhctor Krc
=
due to a radial thermal gradient through the vessel wall, calculated with no plastic-zone correction (ksi Vin.)
K,, calculated with a plastic-zone Kj,
=
correction (ksi.'in.)
Kj Kg, at flaw instability, calculated
=
with a plastic-zone correction (ksi Vin.)
ordinate of the failure assessment n,
=
diagram curve (dimensionless) ratio of the stress intensity Kl
=
factor to the fracture toughness for the material (dimensionless) 4 internal pressure (ksi) p
=
accumulation pressure as defined p,
=
in the plant-specific overpressure Protection Report, but not exceeding 1.1 times the design pressure (ksi)
A-3
p, pressure used to calculate the
=
applied J-integral / tearing modulus line (ksi) p' internal pressure at flaw
=
instability (ksi) reference limit-load internal p,
=
pressure (ksi)
R, inner radius of the vessel (in.)
=
S, abscissa of the failura assessment
=
diagram curve (dinensionices)
Sl ratio of internal pressure to
=
reference limit-load internal pressure (dirensienless)
(SF) satety factor (dinensicaless)
=
t
=
vessel wall thickness (in.)
T
=
tearing modulus due to the applied loads (dinensionless)
T, tearing modulus resistance for the
=
material (dinensionless)
R parameter used to relate the applied
=
J-integral to the applied tearing modulus wbrensionless)
- son's ratio (dinensionless)
V
=
reference flow stress, specified of
=
as 85 kai (ksi) y yield strength for the material (ksi) o
=
A-4
A ARTICLE A-2000 ACCEPTANCE CRITERIA
{
The adequacy of the upper shelf toughness of the reactor vessel shall be determined by analysis.
The reactor vessel is acceptable for continued service when the criteria of Paragraphs (a), (b), and (c) are satisfied.
(a)
Level A and B Servics Loadings When evaluating the adequacy of the upper shelf toughness for the weld material for Level A and D Service loadings, postulate an interior semi-elliptical surface flaw with a depth one-quarter of the wall thickness and a length six times the depth, with the flaw's major axis oriented along the weld of concern and the flaw plane oriented in the radial direction.
When evaluating the adequacy of the upper shelf toughness for the base material,.
postulate both interior axial and circumferential flaws with depths one-quarter of the wall thickness and lengths six times the depth and use the toughness properties for the correspending orientation.
Smaller flaw sizes may be used on an individual case basis when justified.
Two criteria shall be satisfied:
(1)
The applied J-integral evaluated at a pressure which is 1.15 times the accumulation pressure as defined in the plant-specific Overpressure Protection Report, with a factor of safety of 1.0 on thermal loading for the plant specified heatup and cooldown conditions, shall be shown to be less than the J-integral characteristic of the material resistance to ductile tearing at a flaw growth of 0.10 in.
(2)
The flaw shall be shown to be stable, with the possibility of ductile flaw growth, at a pressure which is 1.25 times the accumulation pressure defined in Subparagraph (1), with a factor of safety at 1.0 on thermal loading for the plant specified heatup and cooldown conditions.
The J-integral resistance versus crack growth curve shall be a conservative representation for the vessel material under evaluation.
A-5
(b)
Level C Service Loadings When evaluating the adequacy of the upper shelf toughness f or the weld material f or Level C Service loadings, postulate interior semi-elliptical surface flaws with depths up to 1/10 of the base metal wall thickness, plus the cladding thickness, with total depths not to exceed 1.0 in.,
and a surface length six times the depth, with the flaw's major axis oriented along the weld of concern and the flaw plane oriented in the radial direction.
When evaluating the adequacy of the upper shelf toughness for the base material, postulate both interior axial and circumferential flaws, and use the toughness properties for the corresponding orientation.
Flaws of various depths, ranging up to the maximum postulated depth, shall be analyzed to determine the most limiting flaw depth.
Smaller maximum flaw sizes may be used on an individual case basis when justified.
Two criteria shall be satisfied (1)
The applied J-integral shall be shown to be less than the J-integral char. cteristic of the material resistance to ductile tearing at a flaw growth of 0.10 in., using a f actor of safety of 1.0 on loading.
(2)
The flaws shall be shown to be stable, with the possibilit o of ductile flaw growth, using a factor of safety of 1.0 on loading.
The J-integral resistance versus crack growtb curve shall be a conservative representation for the vessel material under evaluation.
(c)
Level D Service Loadings When evaluating the adequacy of the upper shelf toughness for Level D Service loadings, postulate flaws as specified for Level C Service loadings in Paragraph b), and use the toughness properties for the corresponding orientation.
Flaws of various depths, ranging up to the maximum postulated depth, shall be analyzed to determine the most limiting flaw depth.
Smaller maximum flaw sizes may be used on an individual case basis when justified.
The flaws shall be shown to be stable, with the possibility of ductile flaw growth, using a f actor of safety of 1.0 on loading.
The J-integral resistance versus crack growth curve shall be a best esti. mate representation for the vessel material under evaluation.
The stable flaw depth shall not exceed 75% of the vessel wall thickness, and the remaining ligament shall be safe from tensile instability.
A-6
ARTICLE A-3000 ANALYSIS A-3100 SCOPE This Article contains a general description of procedures which shall be used to evaluate the applied fracture mechanics parameters, as well as requirements for selecting the J-R curve for the material.
References are made to acceptabl:, approaches to apply the criteria.
A-3200 APPLIED J-INTEGRAL The calculation of the J-integral due to the applied loads shall account for the full elastic-plastic behavior of the stress-strain curve for the material.
When the conditions fall into the-category of elastic fracture mechanics with small-scale yielding, the J-integrsi nay alcernately be calculated by using crack-tip stress intensity factor formulae with a plastic-zone correction.
The method of calculation shall be validated and documented.
A-3300 SELECTION OF THE J-INTEGRAL RESISTANCE CURVE When evaluating the vessel for Level A,
B and C Service loadings, the J-integral resistance versus crack growth curve (J-R curve) shall be a conservative representation of the toughness of the controlling beltline material at_ upper shelf temperatures in the operating range.
When evaluating the vessel for Level D Service
- loadings, the J-R curve shall be a
best estimate representation of the toughness of the controlling beltline material at upper shelf temperatures in the operating range.
One of the following options shall be used to determine the J-R curva..
(a)
A J-R curve generated for the actual saaterial under consideration'by following accepted-test procedures ~ may be used.
The J-R curve shall be based on the proper combination of crack orientation, temperature and fluence level._
The crack growth. hall include ductile tearing with no occurrence
-of cleavage.
l A-7
(b)
A J-R curve generated f rom a J-integral database obtained f rom the same class of material under consideration with the same orientation using appropriate correlations for the ef fects of temperature, chemical composition and fluence level may be used.
The crack growth shall include ductile tearing with no occurrence of cleavage.
(c)
When the approaches of (a) or (b) are not possible, indirect methods of estimating the J-R curve may be used provided these methods are justified for the material under consideration.
A-3400 FLAW STABILITY The equilibrium equation for stable flaw growth is J = J, where J is the J-integral due to - the applied loads for the-postulated flaw in the vessel, and J, is the J-integral resistance to ductile tearing for the material.
The inequality for flaw stability due to ductile tearing is BJ dJa Ti da where aJ/da is the partial derivative of the applied J-integral with respect to the f3w depth a with load held constant, and dJ,/da is the slope o;; the J-R curve.
Under a condition of increasing load, stable flaw growth will ' continue as long as aJ/da remains less than dJ,/da.
A-3500 EVALUATION APPROACE FOR LEVEL A AND B SERVICE LOADINGS The procedure given in A-4200-shall be used to evaluate the applied-J-integral for a specified amount of ductile flaw growth..
4 There ai e three. approaches that are equally acceptable for applying the flaw stability acceptance criteria according to the governing flaw stability rules in A-3400.
The first is a J-R curve
- crack driving force diagram approach.
_In this approach flaw stability is evaluated by a-direct - application of the flaw stability rules given in A-2400.
Guidelines for using this s
approach are given in A-4310.
The second is a failure assessment-diagram approach.
A procedure based on this approach for the A-8 s
postulated initial one-quarter wall thickness flaw is given in A-4320.
The third is a J-integral / tearing modulus approach.
A procedure based on this approach for the postulated initial one-quarter wall thickness flaw is given in A-4330.
ARTICLE A-4000 EVALUATION PROCEDURES FOR LEVEL A AND B SERVICE LOADINGS A-4100 SCOPE This Article contains calculation procedures to be used to satisf y the acceptance criteria in A-2000 for Level A and B Service loadings.
A procedure to be used to satisfy the J-integral criterion for a specified amount of flaw growth of 0.10 in. is given in A-4200.
Procedures to satisfy the flaw stability criterion are given in A-4300.
These procedures include the axial and circumferential flaw orientations.
A-4200 EVALUATION PROCEDURE FOR THE APPLIED J-INTEGRAL A-4210 CALCULATION OF THE APPLIED J-INTEGRAL The calculation of the applied J-integral consists of two steps: Step 1 is to calculate the offective flaw depth which includes a plastic-zone correction; and Step 2 is to calculate the J-integral for small-scale yielding based on-this effective flaw depth.
Step 1 For an axial flaw with a depth a,
calculate the stress intensity f actor due to internal pressure with a safety f actor (SF) on pressure by using K
=
ip (SF) p (1 + (Rs/t)) (na)**' Fs (1)
Ps 0.982 + 1.006(a/t)'
=
This equation for K, is valid for 0.20 s a/t s 0.50, and includes i
the effect of pressure acting on the flaw faces.
A-9
For a circumferential flaw with a depth a,
calculate the stress intensity factor due to internal pressure with a safety factor (SF) on pressure by using (SF) p (1 + (Rs/(2t))) (:t a ) **
- Fo (2)
Kz,
=
0.8C5 + 0.23)(a/t) + 0.34S(a/t)'
F,
=
This equation for K,,is valid for 0.20 s a/t s 0.50, and includes the etfeet of pressure acting on the flaw faces.
For an axial or circumferential flaw with a depth a, calculate the stress intensity f actor due to radial thermal gradients by using
((CR)/1000) t
Fs (3)
Kz,
=
F
=
3 0.584 + 2.647(a/t) - G.294(a/t)' + 2.990(a/t)'
This equation for X,,is valid for 0.20 s a/t s 0.50, and 0 s (CR) s 100* F / hou r.
Calculate the effective flaw depth for small-scale yielding, a, by using a,
a + (1/(6n))[(Kg, + Kss)/o )'
=
y Sten 2 For an axial flaw, calculate the stress intensity factor due to internal pressure for small-scale yielding, Kj,, by substituting a, in place of a in equation (1), including the equation for F,.
For a circumfarential flaw, calculate Kj, by subntituting a,
in place of a in equation (2), including the equation for F,.
For an axial or circumferential flaw, calculate the stress intensity factor due to radie1 thermal gradients for small-scale yielding, K;t, by substituting a, in place of a in equation (3), including the equation for F,.
Equations (1), (2) and (3) are valid for 0.20 s a,/t s 0. 50.
The J-integral due to the applied loads for small-scale yielding is given by 1000(Kjp + K j,) */E '
J
=
A-10
_1
)
A-4220 EVALUATION USINO CRITERION FOR FLAH GROWTH OF 0.1 IN.
)
Calculate the J-integral due to the applied loads, J,,
by following A-4210.
Use a flaw depth a equal to 0.25t + 0.10 in.; a pressure p equal to the accumulation pressure for Level A and B Service loadings, p,; and a safety f actor (SF) on pressure equal to 1.15.
The acceptance criterion for Level A and B Service loadings
)
based on a ductile flaw growth of 0.10 in. in A-2000(a)(1) is satisfied when the following inequality is satisfied.
J, < Jo a
)
where J =
the applied J-integral for a safety factor on 2
pressure of 1.15, and a safety factor of 1.0 on thermal loading, s
1 J, 2 = the J-integral resistance at a ductile flaw growth of 0.10 in.
i
)
A-4300 EVALUATION PROCEDUKES FOR FLAM STABILITY A-4310 J-R CURVE - CRACK DRIVING FORCE DIAGRAM PROCEDURE
)
In this procedure flaw stability is evaluated by a direct application of the flaw stability rules given in A-3400.
The applied J-integral is calculated for a series of flaw deptha corresponding to increasing amounts of ductile flaw growth.
The applied J-integral for Level A and B Service loadings shall be f
calculated by using the procedures given in A-4210.
The applied pressure p is set equal to the accumulation pressure for Level A and B Service loadings, p,; and the safety f actor (SF) on pressure-is equal to 1.25.
The applied J-integral is plotted against crack depth on the crack driving force diagram to produce the applied J-l integral curve, as illustrated in Figure A-4310-1.
The J-R curve is also plotted on the crack driving force diagram, and intersects the horizontal axis at the initial flaw depth, a.
Flaw stability at a given applied load is demonstrated when the slope of the applied J-integral curve is less than the slope of the J-R curve at j
the point on the J-R curve where the two curves intersect.
k A-11
'\\Interial J, J
P lied 1 EvaIU3 tion Poim ao a
f FIGURE A-4310-1 COMPARISON OF THE SLOPE 8 OF THE APPLIED J-INTEGRAL CURVE AND TRE J-R CUP.VR.
A-12
4 A-4320 FAILURE ASSESSMENT DIAGRAM PROCEDURE This procedure is restricted to a postulated initial flaw depth equal to ene-quarter of the wall thickness.
A-4321 FAILURE ASSESSMENT DIAGRAM CURVE The same fallure assessment diagram curve shs.11 be used for axial and circumferential flaws, and is given in rigure A-4320-1.
The coordinates (S,, K,) of the f ailure assessment olagram curve are given in Table A-4320-1.
This curve is based on material properties which are characteristic of reactor pressure vessel steels.
A-4322 FAILURL ASSESSMENT POINT COORDINATES The flaw depth a for a-ductile-flaw growth of Aa is given by 0.25t + Aa a
=
The failure assessment point coordinates (Sl, _Kl)-for a ductile flaw growth of Aa shall be calculated by using the following expressions:
K; Kg ( 2 000/(E 'J,) )** *
=
where the stress intensity factor shall be calculated using the flaw depth a without the plastic-zone werection, and:is given by Ks Kg, + Kge
=
and Sl (SF) p/p,
=
where (SF) is the required safety f actor on pressure.
The procedure for calculating K,,, Kg, and p, for axial flaws is _ given in A-4322.1, and for ejreumferential. flaws in-A-4322.2.
A-13
A-4322.1 Axial Plaws The stress intensity f actor due to internal pressure f or axial flaws with a safety factor (SP) on pressure is given by equation (1).
The stress intensity factor due to radial thermal gradients is given by equation (3).
The reference limit-load pressure is given by (2/6 )o y (0.905 - 0.379 ( Aa/ c) )
E' ~
[0.379 + (R / c) + 0.379 ( Aa/ c)T 3
For materials with a yield strength 0 greater than 85 kat, set o 7
f equal to 85 kai in this equation.
This equation tor p, is valid for 0 s aa/t s 0.10.
A-4322.2 circumferential riaws s
The stress intensity factor due to internal pressure for circumferential flaws with a safety factor (SP) on pressure is given by eonation (2).
The stress intensity factor due to radial thermal gradients is given by equation (3).
The refsrence limit-load pressure is given by (1 - 0.91 ( 0.25 + ( Aa/ t) ): ( g/p,) )
o y
~
(1 + (R /(2 c) ))
3 For materials with a yield strength or greater than 85 kai, set o f equal to 85 kai in this equation.
This equation for p, is valid for 0 s aa/t s 0.25.
A-4323 EVALUATION USING CRITERION FOR FLAW STABILITY Assessment points shall be calculated for each loading condition according to A-4322, and plotted on Figure A-4320-1 as follows.
Plot a series of assessment points for various amounts of ductile flaw growth Aa up to the validity limit of the J-R curve.
Use a pressure p equal to the accumulation pressure for Level A and B Service loadings, p,; and a safety f actor (SF) on pressure equal to 1.25.
When one or more assessment points lie inside the failure assessment curve, the acceptance criterion based on flaw stability in A-2000(a)(2) is satisfied.
A-14
TABLE A-4320-1 COORDINATES OF THE FAILURE ASSESSMENT DIAGRAH CURVE OF FIGURE A-4320-1 S,
K, 0.000 1.000 0.050 1.000 i
0.100 0.999 0.150 0.999 0.200 0.996 0.250 0.993 0.300 0.990 0.350 0.987 0.400 0.981 0.450 0.973 3
0.500 0.960 l
O.550 0.939 0.600 0.908 0.650 0.864 0.700 0.807 0.750 0.737 0.800 0.660 0.850 0.581 0.900 0.505 0.950 0.435 1.000 0.374 1.050 0.321 1.100 0.276 1.150 0.238 D
A-15
1.2 i
1.0 I
i
'N 0.8 l
N 0.6 i
\\
K, 0.4 0.2 0.0 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 S,
FIGURE A-4320-1 FAILURE ASSESSMENT DIAGRAM FOR THE ONE-QUARTER WALL THICKNESS FLAW.
c A-16
A-4330 J-INTEGRAL / TEARING MODULUS PROCEDURE This procedure is restricted to a postulated initial flaw depth equal to one-quarter of the wall thickness.
A-4331 J-INTEGRAL AT FLAW INSTABILITY Referring to Figure A-4330-1, the onset of flaw instability is the pH nt of intersection of the applied anc material curves plotteu on a graph of the J-integral versus tearing modulus (J versus T).
The expression for the applied J/T curve is given by (2 000 W t og'/2) T (4)
J
=
where o is a reference flow stress which is set to 85 kai in f
equetion (4).
For axial flaws 0.235[1 + (0. 083 x 10*')(CR) t'/((SF)p,))
(5)
W
=
where p, is the pressure under evaluation.
Equation (5) is valid for 6 s t s 12 in.,
2.25 s ((SF)p,) s 5.00 kei, and 0 s (CR) s 100* F/ hou r.
For circumferential flaws
- 0. 21(1 + (0. 25 7 x 10) (CR) t'/((SF)p,))
(6)
W
=
Equation (6) is valid for 6 s ts 12 in.,
2.25 s ((SF)p,) s 9.00 kai, and 0 s (CR) s 100'F/ hour.
Equations (4), (5)'and (6) are based on material properties which are characteristic of reactor pressure vessel steels.
The tearing modulus for the material is determined' by differentiation of the J-R curve with respect to flaw-depth a.
(E/(1000 og')) AJ,/da (1)
T,
=
The same values for E and o shall be used in equations-(4) and r
(7).
The J-integral versus tearing modulus Ja/Ta curve for'the material is given - by plotting Ja. against T for a series of increments in ductile. flaw growth.
Each coordinate for. Ja is evaluated at the - same amount of- - ductile - flaw growth - as -the coordinate for Ta.
A-17
Tno value of the J-integral at the onset of flaw instability, J', corresponds to the intersection of the applied J/T curve given by equation (4) with the material J./T, curve, as illustrated in rigure A-4330-1.
The J-integral at the onset of flaw instability may be determined analytically when a power-law curve fit to the J-R curve of the form J, - C ( A a ) #8 g
is available.
The J-integral at the onset of flaw instability, J',
in this case is given by J' - C ( W C Cg ) '*
1 A-4332 INTERNAL PRESSURE AT FLAW INSTABILITY 1
The calculation of the internal pressure at the onset of f3 w instability is based on the value of the J-integral at the c: 4L of flaw instability, J'.
The ductile flaw growth at the onset of flaw instability, Aa', is taken f rom the J-R curve.
The effective flaw depth at the onset of flaw instability includes the ductile flaw growth Aa', and is given by a'
0.25t + ba' + (1/(6n))[J'E'/(1000 ol))
=
The stress intensity factor due to radial thermal gradients at the onset of flaw instability, Kje, for axial or circumferential flaws i
is given by X},
((CR)/1000) t * *
- F]
=
F}
0.584 + 2.64 7(a /t) - 6.294(a;/C)* + 2.990(a /t)'
=
This equation for KI, is valid for 0.20 s a;/t s 0.50, and 0 s (CR) s 100*F/ hour.
The stress intensity f actor for small-scale yielding due to internal pressure at the-onset of flaw-instability, Kj,, is-given bv Kj, (J'E '/1000) ** * - K;,
=
A-18
For a given value of X;,, the internal pressure at the onset of flaw instability for axial flaws is given by K;, / ((1 + (R,/t))(na')**' F;)
p'
=
- 0. 982 + 1. 006(a */t j' F;
=
and for circumferential flaws by K;, / [(1 + (Rs/(2t))) (na;)'*
- F;)
p'
=
- 0. 885 + 0. 23)(a'/t) + 0. 345(al/t)'
F;
=
These equations for p' are valid for 0.20 s a;/t s 0.50, and include the effect of pressure acting on the flaw faces.
A-4333 EVALUATION USING CRITERION FOR FLAW STABILITY Calculate the value of the J-integral at the onset of flaw i
instability, J',
by following A-4331 using a pressure p,
in equations (5) and (6) equal to the accumulation pressure for Level A and B Service loadings, p ; and a safety f actor (SP) on pressure equal to 1.25.
Calculate the internal prosesre at the onset of flaw instability, p*,
by following A-4332.
The acceptance crit' 'on based on flaw stability in A-2000(a)(2) is satisfied when the following inequality is satisfied.
p'
- 1. 25 p, ARTICLE A-5000 LEVEL C AND D SERVICE LOADINGS The possible combinations of loadings and material properties which may be encountered during Level C and D Service loadings are too diverse to allow the application of pre-specified procedures and it is recommended that each situation be evaluated on an individual case basis.
l A-19
Instability
- l' Material 3n vs Ta lc i
U Appiled 3 vs T
.c 4
T FIGUgg g.4 0-1 ILLUSTRATION OF THE g. INTEGRAL /TEARiyo MODULUS PROCEDURE A-20
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