ML20154A881
| ML20154A881 | |
| Person / Time | |
|---|---|
| Site: | Davis Besse  |
| Issue date: | 05/04/1988 |
| From: | Shelton D TOLEDO EDISON CO. |
| To: | NRC OFFICE OF ADMINISTRATION & RESOURCES MANAGEMENT (ARM) |
| Shared Package | |
| ML20154A885 | List: |
| References | |
| 1510, TAC-66699, NUDOCS 8805160223 | |
| Download: ML20154A881 (2) | |
Text
/
f g -4 1
TOLEDO EDISON A Caterer tg Corrsey DONALD C. SHELION vu % sa.,
[419)243 2300 Docket No.
50-346 License No. NPF-3 Serial No. 1510 May 4, 1988 United States Nuclear Regulatory Commission Document Control Desk Vashington, D.C.
20555 Subj ect:
Supplemental Information Regarding the License Amendment Request to Revise the Reactor Coolant System Pressure-Temperature Operating Limits and Reactor Vessel Material Serveillance Program (TAC No. 66699)
Gentlemen:
In response to a request made during an April 6, 1988 telephone conversation between Mr. A. V. DeAgazio, the Nuclear Regulatory Commission (NRC)/ Nuclear Reactor Regulation (NRR) Davis-Besse Project Manager, and Toledo Edison, additional information is being provided to assist in the review of the License Amendment Request which vas submitted to the NRC on March 31, 1988 (Serial No. 1490). This License Amendment Request proposes revising the Reactor Coolant System Pressure-Temperature (F-T) Curves and other related changes necessary to allow operation to ten Effective Full Pover Years (EFPY). Additionally, changes to the Reactor Vessel Material Surveillance Program Schedule vere requested.
Each NRC question, followed by Toledo Edison's response, is provided below:
Question:
Provide the Reference Temperature (RTNDT) at 10 EFPY and at End of Reactor Vessel Life (E0L).
Response: The 10 EFPY reactor vessel material properties used in the preparation of the P-T curves are included in Babcock and Vilcox (B&V) Topical Report BAV 2011. "Pressure-Temperature Limits for 10EFPY", November 1987, as referenced by the License Amendment i
Request submitted in Serial No. 1490. The controlling beltline veld (VF-182-1) RT values are summarized below.
The NDT predicted E0L Reference Temperatures from Topical Report BAV 1882, "Analysis of Capsule TEl-A", September 1985, are also listed:
i f0f 1HE TOLEDO EDSON COVPANY EDSON PLAZA 300MADSON AVENUE TOLEDO. OHf 0 43652 i
{
8805160223 880504 PDR ADOCK 05000346 L
n rem
.s.
>e.
Docket No.
50-346 License No. NPF-3 L
Serial No. 1510 10 EFPY 10 EFPY EOL Region Predicted Assuvd Predicted RTNDT( F)
RTygg.('F)
RTNDT('F)
Beltline 1/4T 183 147 222 Beltline 3/4T 139 107 182 Closure Head 60 60.
60 Outlet Nozzle 60-60 60 Question:
Provide the calculations for the Pressure-Temperature (P-T)
Curves.
3 Response:, B&V Topical Report BAV 10046A, Revision 2, "Methods l
of Compliance with Fracture Toughness and Operational requirements of Appendix G to 10CFR50", describes the methodology used to calculate the P-T curves., B&V Topical Report BAV 2011 "Pressure-Temperature Limits for 10EFPY", describes the results of the calculation.
Note that Appendix A of BAV 2011 has not been included since it only presented B&V's recommendation for the wording of Technical Specification 3/4.4.9 and its Bases. Toledo Edison made additional changes to this and related Technical Specifications and the B&V recommendations do not accurately reflect those changes submitted by Serial No. 1490. Therefore, Toledo Edison has elected to omit Appendix A of the enclosed BAV 2011.
Each proposed change to the Technical Specifications has been described and justified in the License Amendment Request of i
Serial No. 1490.
Toledo Edison believes the above addresses the NRC concerns regarding this License Aaendment Request. Should there be any additional questions, please contact R. V. Schrauder, Nuclear Licensing Manager at (419) j 249-2366.
Very truly ours, 4
DRB/ tit At.tachments cc: DB-1 NRC Resident Inspector A. B. Davis, Region III Regional Administrator A. V. DeAgazio, NRC/NRR Davis-Besse Project Manager l
State of Ohio l
l
4 H%q d
Io, UNITED STATES f ", v.
~h NUCLEAR REGULATORY COMMISSION 8
WASHINGTON, D. C. 20555
/
APR 3 0 ja
' 8 e....
Babcock and Wilcox Company l
ATTN: James H. Taylor
>3 Manager, Licensing i3 Nuclear Power Generation P.O. Box 10935 Lynchburg, Virginia 24506-0935
Dear Mr. Taylor:
SUBJECT:
ACCEPTANCE FOR REFERENCING OF LICENSING TOPICAL REPORT BAW-10046, REY. 2 B&W OWNERS GROUP MATERIALS COMMITTEE "METHODS OF COMPLIANCE WITH FRACTURE TOUGHNESS AND OPERATIONAL REQUIREMENTS OF 10 CFR 50, APPENDIX G" We have completed our review of the subject topical report submitted by
'l Babcock and Wilcox (B&W) by letter dated December 21, 1984 We find the report to be acceptable for referencing in license applications to the extent specified and under the limitations delineated in the report and the associated NRC evaluation, which is enclosed. The evaluation defines I
the basis for acceptance of the report.
We do not intend to repeat our review of the matters described in the
- S-repcrt and found a
- ceptable when the report appears as a reference in license applications, except to assure that the material presented is applicable to the specific plant involved. Our acceptance applies only
,l to the matters described in the report.
In accordance with procedures established in NUREG-0390, it is requested that B&W publish accepted versions of this report, proprietary and non-I proprietary, within three months of receipt of this letter. The accepted ve sions shall incorporate this letter and the enclosed evaluation be-tween the title page and the abstract.
The accepted versions shall include an -A (designating accepted) following the report identification symbol.
Should our criteria or regulations chance such that our conclusions as to the acceptability of the report are invalidated, B&W and/or the applicants I
referencing the topical report will be expected to revise and resubmit their respective documentation, or submit justification for the continued ef fective applicability of the topical report without revision of their
.E respective documentation.
Sincerely, 1
/
8 Dennis M. Crutc6fie s.stan Director for Technical Support Division of PWR Licensing-B
Enclosure:
As stated 6f lYrC/ n 4 '-
9 mu o m.
ENCLOSURE SAFETY EVALUATION OF TOPICAL REPORT BAW-10046, Rev. 2. "Methods of Compliance With Fracture Toughness and Operational Requirements of 10 CFR 50, Appendix G"
SUMMARY
OF REPORT This Topical Report was submitted by letter of December 21, 1984 to Mr. Cecil 0. Thomas, Chief Standardi;:ation and Special Projects Branch.
It is an update of BAW 10046A, Rev. 1 issued in July 1977, which was eval-uated and accepted for referencing in licensing applications by(letter of June 22, 1s77. There are two principal changes in Revision 2:
- 1) additions and changes made to reflect the revised criteria in the amendments to Appendix G, 10 CFR 50, effective July 26, 1983, and (2) a new Chapter 6 that was added to describe the method of analysis and the material properties data to be used when a more advanced ductile fracture analysis is called for by evidence that the Charpy upper shelf energy of the reactor vessel beltline material has fallen below 50 ft lb.
Specifically, Chapter 6 was intended to comply with the requirement for continued operation given in paragraph V.C of Appendix G, 10 CFR 50.
The initial review of BAW 10046, Rev. 2 prompted a number of requests for explanatier. of the methods described in the report. The explanations were given in BAW 1868, March 1985 entitled "BAW 10046A Rev. 2 Supplement."
However, only BAW 10046, Rev. 2, which contains the requirements and criteria, is being accepted for referencing in licensing applications.
Another round of questions and corrents in December 1985 was answered at a meeting on January 8,1986 follcwed by a submittal of proposed additions and corrections by letter of April 24, 1986 to H. Denton.
In this evaluation it is assumed that these changes will be made in the final report.
Chapters 1 through 5 of Revision 2 of BAW 10046 cover the same material as did Revision 1.
They give an introduction (Chapter 1), discuss operating moces such as bolt up, heat up, cool down, operation, and testing (Chapter 2),
describe the measurement of the material properties pertinent to fracture toughness and to radiation damage of the reactor vessel beltline (Chapter 3),
show how B&W applies ASME Code and NRC requirements in the generation of pressure temperature (P-T) limits (Chapter 4), and give examples of P-T limits for 5 and 32 EFPY (Chapter 5).
Chapter 6, as described above, is new.
REGULATORY EVALUATION OF REPORT We have reviewed the B&W method of calculation of pressure - temperature limits against the requirements of Appendix G, 10 CFR 50, Regulatory Guide 1.99, and the Standard Review Plan.
Special attention was paid to the updates made in compliance with changes in Appendix G, 10 CFR 50 that became effective July 26, 1983.
These concerned consideration of the closure flange regions as potential controlling locations for the P-T limits, the hydrotest temperature requirement when there is no fuel in the
s o
l,
reactor, and the more specific requirements for consideration of reactor vessel integrity when the beltline material exhibits low upper shelf behavior. Assuming the final report is edited as indicated in the sub-mittal received April 24, 1986, Chapters 1-5 are satisfactory with the following comment on paragraph 3.1.3.
Chapter 6 is discussed separately.
3.1.3 Radiation Effects The B&W report states that the methodology used to adjust the RTNDIhe values as used in developing pressure-temperature limits will be currently accepted procedures." A similar statement is made concern-ing the decrease in Charpy upper shelf energy. As further stated in the report, the decision on these issues is simply passed to the utility, to be made when a submittal is made to the NRC.
It should be noted that Revision 2 of Regulatory Guide 1.99, which addresses these topics, was issued for public coment in February 1986. When issued in final form, it will be the basis for review of subsequent submittals, but if a utility chooses to use Rev. 2 as the basis for a submittal before that date, it will be evaluated on the basis of Rev. 2.
The Guide covers the use of plant surveillance data as well as calculative procedures to be used when there are no credible plant surveillance data.
As a cautionary note, alternatives to the use of either plant surveil-lance data or procedures that derive from analysis of a broad data base are difficult to justify; because significant scatter in these data gives rise to the possibility that a small subset of the data base will not give representative values of the mean and the uncertainty pertinent to the vessel in question.
Chapter 6 EPFM Analytical Procedures This chapter describes a method of fracture analysis for low upper shelf (below 50 ft lb) material.
Paragraph V.C of Appendix G requires such an analysis to demonstrate that the margin of safety is equivalent to that required by Appendix G of the ASME Code.
Chapter 6 has been reviewed against the guidelines for resolution of this issue given in NUREG-0744*, supplemented and in some cases superseded by the ongoing effort of the Working Group on Flaw Evaluation (WGFE) of the Sebcomittee on Nuclear Inservice Inspection of the ASME Boiler and Pressure Vessel Comittee. The report of the WGFE, "Development of Criteria for Assessment of Reactor Vessels With Low Upper Shelf Fracture Toughness" is still in preparation. Actually, the technology involved in both the analytical effort and the materials testing is still developing, as the foregoing illustrates.
Thus, this evaluation must be regarded as an interim one, subject to change as further developments occur.
l
- R. Johnson, "Resolution of the Task A-11 Reactor Vessel Materials Toughness Safety Issue, Vols. 1 & 2, October 1982 l
l t
O 3
The method of analysis used in BAW 10046, Rev. 2, meets the recommendation of NUREG-0744 that the analysis should use a J-integral formulation and that the material toughness should be characterized by J-R curves.
In addition, the B&W analysis meets the WGFE crack stability criteria:
dJ
< dU applied material da da at Japplied = Jmaterial The BL4 analysts have their own way of solving these simultaneous criteria, but it appears in the WGFE criteria document and it gives the same result as the other methods given there and is therefore acceptable.
To resolve any question about the physical meaning of the mathematical procedures (at the request of NRC staff), the B&W authors included a figure (Fig. 6-5) that summarizes the results by giving the pressures to cause crack initiation and tearing instability and the predicted amounts of crack growth at which initiation and instability occur.
The same figure also shows the pressure to cause plastic instability as a function of crack size.
The acceptance criteria for normal and upset conditions given in BAW 10046, Rev. 2, meet the criteria given in the WGFE draft report that margin should be based on load (not J) and that separate requirements should be given for crack stability and crack initiation.
The instability criterion given in the WGFE draft report is as follows:
"The postulated crack
- shall be demonstrated to be stable under ductile l
crack growth with a factor of safety of two on pressure and a factor of one on thermal loading for all service level A and B conditions except for hydrostatic tests.
For hydrostatic tests, the factor of safety shall be 1.5 on pressure and a factor of one on thermal loading."
This meets the requirement of Appendix G, 10 CFR 50, that margins against fracture should be equivalent to those required by Appendix G of the ASME 4
f Code.
The latter requires for the determination of allowable pressure during Service Conditions for which Level A and Level B Service Limits are specified, use of the following formula:
2Kyg + KIT IR The quantity K (membrane) is proportional to pressure.
This is the origin ofthefactorE2onpressure.
Thus, it seems clear that the acceptance criteria in BAW 10046, Rev. 2 should refer to Level A and 8 Service Condi-tions, rather than to cperating pressure.
The effect of the B&W criterion is that crack stability under ductile tearing conditions must be shown for a pressure of 4500 psi (2 x operating pressure) whereas the WGFE criterion, which meets the requirements of Appendix G,10 CFR 50, could require
" An ASME Section I!!, Appendix G flaw (a semi-elliptical surface flaw with the depth equal to one quarter of the wall thickness and length six times the depth).
e
1 5
4
[.
)
stability at as much as 5500 psi pressure, if the Design Specification 4
-designated a Level B service loading that produced a pressure stress of 110 percent of the design stress intensity value, S,.
4 The B&W criteria omit thermal stresses on the grounds that their effect is small and the methods used for calculation of J applied due to thennal stress are controversial.
It is agreed that for the Service Level A and B j
conditions where ductile tearing stability is an issue the effects of thermal stress are small; hence, the omission of thermal stress is accept-able.
With regard to crack initiation, the criterion in BAW 10046, Rev. 2 is that the crack initiation pressure must exceed 3000 psi where the J value for j
crack initiation is defined as Jmaterial at 0.01 inch crack growth. This is conservative, compared to the WGFE criterion, which is based on 2750 psi l..
and one millimeter (0.04 inch) crack growth.
In sumary, Chapter 6 of the Topical Report gives a satisfactory method of analysis of the resi~ stance to ductile tearing instability and plastic in-stability of the reactor vessel beltline with the exception that the accept-ance criterion for instability should be for all Level A and B loadings not j
just operating pressure.
4 REGULAT0kY p0SITION Tepical Report BAW 10046, Rev. 2 describes acceptable methods for the develop-ment of allowable pressure - temperature limits for normal operation and for l
test conditions to assure the prevention of non-ductile fracture.
It may be referenced in future applications for setting these limits in Technical Specifications.
It is understood that the report dated December 1984 will be edited per the submittal of corrections and additions to the text, sub-mitted by letter to H. Denton, dated April 24, 1986.
l Topical Report BAW 10046. Rev. 2 also describes acceptable methods for the j
ana!ysis and materials properties data required to demonstrate resistance of the reactor vessel beltline to ductile tearing instability when the Charpy 1
upper shelf energy of the beltline materials falls below 50 ft lb.
It may be referenced in license submittals made in conformance to the requirements of Appendix G, 10 CFR 50, except the acceptance criteria should be for all Level A and B loadings.
It should also be noted that the technology for l
treatment of ductile tearing instability is less mature than for example that for non-ductile fracture; hence, future revision of this requirement may be expected.
i l
4 i
I l
Babcock & Wilcox no.i, p.
, oi,ign a uceermott company April 24, 1986 po. 9oplo;p$
3l50 o t Road JHT/86-075 Lynenburt vA 24506.. 45 (804) 385 2000 Mr. Harold Denton Executive Director of operations
' office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C.
20555
Subject:
B&W Topical Report BAW-10046, Rev.
2, dated December 1984, "Methods of Compliance with Fracture Toughness and operational Requirements of 10CFR50, Appendix G.
References:
Letters, J. H. Taylor to C.
- o. Thomas, dated December 21, 1984 and April 25, 1985,
Attachment:
Additions and Clarifications to IAW-10046, Rev. 2.
Dear Sir:
In order to expedite the NRC's review and approval of the subject report, the attached supplemental information requested by the Staff's Dr.
P. N. Randall is being forwarded to you.
It is intended that this information will be included in the approved version of BAW-10046, Rev.
2.
Your immediate approval of the subject report is requested by Babcock & Wilcox, the B&W owners Group and the following Utilities:
Arkansas Power & Light Co.
GPU Nuclear Corporation Duke Power Co=pany Sacramento Municipal Utility Dist.
Florida Power Corporation Toledo Edison Company If you should have any questions, please do not ',tesitate to call.
Very truly youts,
.w b
'L J. H. Taylor Manager Licensing Services JHT/leh Attachment b Ah n e t / C O )O Q)f W Me.LQ
/
c
s
.cc:
P. N. Randall W.
Paulson B. J. Elliot t
B&WOG Materials Ccamittee D. F. Spond
- AP&L M. A. Maghi
- DPCo D. N. Miskiewicz - FPC R.
L. Miller
- GPUN S.
W. Rutter
- SMUD D. R. Cox
- SMUD R. J. Gradomski
- TED l
ADDITIONS AND CLARIFICATIONS TO BAW-10046, REV. 2 The folicring additions and corrections are proposed for BAW-10046 Rev. 2 Para 4.2.1.1 (New Paragraph) 6.
Appendix G of the ASME Code Section III recommends the temperature of the closure area be RT at bolt-up.
The forgoing procedura yielcs similar results with tbTexception of low stressac closures.
This procedure is considered consistent with the philosophy of the ASNE Code and will be used for establishing temperature requirements.
Para. 4.2.1.2 (New paragraph) 3.C 10CFR50 Appendix G Paragraph IV.A.2 requires the highly stressed regfons of the closure region to be at a temperature of at least RT 120 F for pressures above 625 psig.
The forgoing procedure results $7 +a g
similar temperature requirement.
The required temperature is icwor than 0
120 F if slow heat-up rates are specified and higher than 120 F for the 0
operating pressure condition and maximum heat-up rates. The forgoing procedure is considered to be consistent with the requirements of 102FR50 Appendix G is used in lieu of tha stated requirement.
YAa4 Pa ra. 4.2. 3. 2 (Add at end of paragraph)
The requirement of 10CFR50 Appendix G specifying a temprature of RT 90 F for highly stressed regions of the closure for pressures above b. +
0 8
psig is essentially met by this procedure.
As for the normal heat-up caso higher or 1cuer temperatures may De requirec depending on heat up rate.
Table 4-1 Revise materials property RT location for closure head to 1/4t.
NOT Tatie 5-1 Revise line (K.) in enemistry column uncer L 0.61 should be.016.
Also fer material E, revise initial RT I #30' NOT Taele 5-2 Revise table fcr typographical errors first heacir.g RT should e ART j
Fo.- MAT IO E 260 shoule ce 250.
- g..
"JGL.5 i Change "preventation" to "prevention" and change J IR
- J -R.
I l
Page 6-5 Revise paragrapn 6.2.5 to 6.3.
In paragraph 6.3 second paragraph is revisec "as veil as the local" to "as well as the local plastic instability i
pressure calculated :y the ratio".
Ta:1e 6-12 shoule ce Ta:1e 6-1.
t Revise page 6-8 as fo11cvs:
In Figure 6-3 El-L9 shoulc be M1 = 6.9 anc 1cuer taele should be lacelee Table 6-1 Babcock & Wilcox a mcemon w uny l
s
." OWNERS GR@UP m e
,.ue u -,
me mume a.,., w ODC cw M 5
EODi,
7' '
Cl~. :",'"", -
C.-
M, C.
i Working Together to Economically Provide Reliable and Safe Electrical Power April 25,1985 Suite 220 7910 Woodmont Avenue Beth Mda, Maryland 20814 (301)951 3344 Mr. Cecil 0. Thomas, Chief Standardization & Special Projects Branch Division of Licensing U.S. Nuclear Regulatory Commission Washington, D.C.
20555
Subject:
B&W Report, BAW-10046, Rev. 2, dated December 1984, "Methods of Compliance with Fracture Toughness and Operational Requirements of 10CFR50, Appendix G."
References:
- 1) Letter, N. P. Kadambi to E. C. Simpson, dated February 20, 1985.
- 2) Letter, J. H. Taylor to C. O. Thomas, dated December 21, 1981..
Attachments:
- 1) B&W Report, BAW-1868, "BAW-10046A, Rev. 2, Supplement," dated March 1985,
- 2) B&W Report, BAW-1814, " Analysis of HSST Intermediate Yessel V-8A Test by the Deformation Plasticity Failure Assessment Diagram Method," dated November 1983.
Dear Mr. Thomas:
The attached reports are being submitted on behalf of the B&W Owners Group at the request of your Dr. P. N. Randall to facilitate the review and approval of the subject report.
BAW-10046, Rev. 2, was submitted by Reference 2 and it is understood that the review is still scheduled to be complete by June 1985 (Reference 1). contains additional information in support of SAW-10046, Rev. 2, including details on the deformation plasticity failure assessment diagram (JPFAD), calculation of the Ramberg-0sgood stress-strain relationship and the reactor vessel closure analysis. presents a OPFAQ analysis of the NRC sponsored c.ett of HSST vessel j
V-8A which serves to benchmark the analytical approach to experimental results.
h Should you have any questions or comments, please contact our Mr. C. J. Hudson (804-385-2550).
Very truly yours, J. H. Taylor Manager Licensing Services JHT/leh Attachments cc:
B. J. Elliot P. Kadambi D. Moran P. N. Randall G. Vissing B&WOG Materials Committee D. F. Spond
- AP&L M. A. Hachi
- OPCo R. A. Webb
- FPC J. A. Janiszewski - GPUN
- 0. R. Cox
- SMUD S. W. Rutter
- SMUD R. J. Gradomski
- TED E. C. Simpson
- FPC I
{
"' i m.
I
e,.[-
Babcock & Wilcox
= %om a Mcoermott company 3315 Old Forest Road December 21, 1984
$0$1$8 3m33 ESC-852
<eo4)3shooo Mr. Cecil 0. Thomas, Chief Standardization & Special Projects Branch Division of Licensing U.S. Nuclear Regulatory Commission Washingt,on. 0.C.
20555
Subject:
B&W Topical Report BAW-10046 Rev. 2, dated December 1984, "Methods of Compliance with Fracture Toughness and Operational Requirements of 10CFR50, App. G.
Dear Mr. Moran:
Enclosed are ten (10) copies of the subject report which is being submitted on behalf of the B&W Owners Group and the following Utilities:
Arkans'as Power & Light Company Duke Power Company Florida Power Corporation GPU Nuclear Corporation sacramento Municipal Utility District Toledo Edison Company BAW-10046, Rev. 1 has previously been approved by the NRC.
Revision 2 is being submitted as required by 10CFR50, Appendix G. Paragraph IV.A.1 which states that the Director, Office of Nuclear Reactor Regulation must approve the manner in which "equivalent margins of safety" are provided for reactor vessel beltline materials which do not maintain 50 ft-lbs of Charpy Upper Shelf Energy.
That manner is described in BAW-10046 Revision 2.
The method of Elastic-Plastic fracture mechanics known as the Failure Assessment Diagram has been under development for some time.
The submittal of BAW-10046, Revision 2 culminates years of work by the B&W Owners Group and the Babcock and Wilcox Company.
Approval of the methodology contained in the Topical Report is a key milestone in the B&W Owners Group Reactor Vessel Materials Program which has been underway since 1976.
In order to assure continued compliance with Federal Regulations, NRC approval of this document is requested by May 1, 1985.
()) I \\ 10 T I lalr) w&
33uw
O If you should have any questions please contact Mr. C. J. Hudson (804) 385-2550 or Mr. H. W. Behnke (804) 385-2417.
Very truly yours, h 50%
J. H. Taylor, Manager Licensing JHT/ met cc:
W/ Attachment D. Moran B. J. Elliot W. S. Hazelton P. N. Randall P. Kadambi R. Johnson B&W Owners Group Materials Committee D. F. Spond
- AP&L P. Guill
- OPCo R. A. Webb
- FPC J. A. Janiszewski - GPUN S. W. Rutter
- SMUD R. J. Gradomski
- TED E. C. Simpson
- FPC 4
I Dtacket No. 50-346 License No. NPF-3 BAW-10046A, Rev 2 serial No. 1510 Topical Report June 1986 lETH005 0F COMPLIANCE WITH FRACTURE TOUGHNESS AND OPERATIONAL REQUIREfENTS OF 10 CFR 50, APPENDIX G by H. W. Behnke A. L. Lowe, Jr.
J. M. Bloom W. A. Van der Sluys BABCOCK & WILC0X l
Nuclear Power Division / Alliance Research Center P. O. Box 10935 Lynchburg, Virginia 24506-0935 Babcock &Wilcox n w. Q, y j
P^
Babcock & Wilcox Nuclear Power Division / Alliance Research Center Lynchburg, Virginia Topical Report BAW-10046A, Rev 2 i
June 1986 Methods of Compliance With Fracture Toughness and Operational Requirements of 10 CFR 50. Appendix G H. W. Behnke, A. L. Lowe, Jr., J. M. Bloom, W. A. Van der Sluys Key Words: Ferritic Materials, Reactor Coolant Pressure Boundary, Reference Temperature, Charpy Upper Shelf Energy, Appendix G to 10 CFR 50, Appen.
dix G to AS4E Code, Fracture Prevention, Pressure Temperature Limitation, Technical Specifications, Ductile Tearing Instability, Elastic-Plastic Fracture Mechanics, Deforma-t tion Plasticity Failure Assessment Diagram ABSTRACT This report describes B&W's practices, methods, and criteria for compliance with the requirements of Appendix G to 10 CFR 50, "Fracture Toughness Re-quirements."
The ferritic materials and the operational parameters of the reactor coolant sys tem for nuclear power plants designed by B&W are de-scribed as are the methods for obtaining and estimating the reference tem-perature and the Charpy upper shelf energy.
The acceptance criteria for unirradiated Charpy upper shelf energy is given.
The adequacy of fracture toughness properties of bolting materials and type 403 materials are demon-I strated.
The methods employed to detemine the reactor coolant system pressure-temperature limit curves are given for each of the loading condi-r tions required by Appendix G to 10 CFR 50.
The pressure-temperature limit curves imposed by several regions of the reactor vessel are illustrated as is the development of the composite limit curves.
Furthermore this report describes the methods used to preclude ductile tearing instability.
This analysis applies to irradiated vessels with low upper shelf energies.
The l
Technical Specifications pressure-temperature limit curves and the Preser-vice System Hydrostatic Test limit curve of a typical 177 FA plant are also desc ribed.
i l
Babcock &WHcom
, 4x,
a MCDermott (Ompany
no
.g a
CONTENTS Page 1.
INTRODUCTION...........................
1-1 1.1.
Packground.........................
1-1 1.2.
Scope and Organization...................
1-2 2.
REACTOR COOLANT PRESSURE BOUNDARY 2-1 2.1.
Components.........................
2-1 2.2.
Ferritic Materials and RCPB Operational Parameters..... 2-3 2.3.
N o rma l O pe ra t i o n......................
2-3 2.3.1.
Bol t P rel oa d.................... 2-3 2.3.2.
Heatup.......................
2-4 2.3.3.
Cooldown...................... 2-4 2.4.
Preservice System Hydrostatic Test '.............
2-5 2.5.
Inservice System Leakage and Hydrostatic Tests.......
2-5 2.6.
Reactor Core Operation...................
2-6 3-1 3.
MATERIAL PROPERTIES.......................
3.1.
Impact Properties of Ferritic Materials 3-1 3.1.1.
Determination of RTNOT...............
3-1 3.1.2.
Determination of Charpy V-Notch Level 3-3 3-5 3.1.3.
Radiation Effects 3.2.
Impact Properties of Bolting Materials...........
3-7 3-7 3.2.1.
Code Requirements 3-7 3.2.2.
Estimating Method 3.3.
Impact Properties of Type 403 Modified Steel........ 3-9 3-9 3.3.1.
Code Requirements 3.3.2.
Demonstration of Adequate Toughness 3-10 3.4.
Supplemental Fracture Toughness Properties.........
3-11 3.4.1.
Terminology Related to Ductile Fracture Analysis.. 3-11
.l 3.4.2.
Toughness Properties of Ductile Materials 3-11 3.4.3.
Relationship Between Fracture Toughness Properties and the Fracture Mechanics Analysis 3-12 4
LEFM AN ALYTIC AL PROCEDURES.................... 4-1 4-1 4.1.
Basis 4-4 4.2.
Description 4.2.1.
No rmal Ope rati o n.................. 4-5 4.2.2.
Preservice System Hydrostatic Test......... 4-18 4.2.3.
Inservice System Leak and Hydrostatic Tests 4-20 4.2.4.
Reactor Core Operation............... 4-21
-x-Babcock & WHeos a hu oemen wmpany
O 7,.
CONTENTS (Cont'd)
- Pag, S.
TYPICAL PRESSURE-TEMPERATURE LIMITS 5-1 5.1.
Composi te Limi t Curves...................
5-1 5.2.
Technical Speci fication Limit Curves............
5-3 5.3.
Preservice System Hydrostatic Test Limit Curve.......
5-4 6.
EPFM MALYT IC AL PROC EDUR ES.................... 6-1 6.1.
Basis 6-1 6.2.
Elastic-Plastic Fracture Mechanics Analytical Model 6-2 6.2.1.
OPFAD Curve Generation............... 6-2 6.2.2.
Assessment Point Evaluation 6 -4 6.2.3.
Instability Pressure Prediction 6-4 6.3 Sample Calculation and Presentation of Data 6-5 6.4.
Th e rmal S tre s s....................... 6-5 6.5.
Acceptance Criteria 6-5 7.
SUMMARY
AND CONCLUSIONS 7-1 S.
R EF E R E NC E S............................ 8-1 List of Tables Table 2-1.
Ferritic Materials Used in Reactor Coolant Pressure Boundary.
2-7 3-1.
Summary of RTNDT Data and Estimated Temperatures.......
3-14 3-2.
Summary of Cy USE Data and Estimated Upper Shelf Energies 3-15 4-1.
Ou tl i ne of Me t ho ds......................
4 -2 2 5-1.
Unirradiated Impact Properties and Residual Element Content of Beltline Region Materials in a Typical 177 FA Plant....
5-5 5-2.
Typical Material Data for Preparing Beltline Region Pressure-Temperature Limit Curves 5-6 6-1.
Tabular Results of FAD Analysis Shown in Figure 6-3 6-8 List of Figures Figure 3-1.
Relationship Between Fracture Toughness Properties and the Fracture Mechanics Evaluation Methods........
3-16 4-1.
Reference Critical Stress Intensity Factor Vs Temperature Relative to RTN OT,.....................
4 -2 3 5-1.
Nomal Heatup Pressure-Temperature Limits Imposed by Several Reactor Yessel Regions and Composite Limit Curve for 5 EFPY 5-7 l
5-2.
Nomal Heatup Pressure-Temperature Limits imposed by Several Reactor Vessel Regions and Conposite Limit Curve f o r 3 2 EF P Y,,.......................
5-8 Babcock &WHess
- Xi a woermoit comceny
(
Figures (Cont'd)
Figure Page 5-3.
Normal Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve 4
for 5 EFPY 5-9 5-4.
Normal Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve for 32 EFPY..........................
5 10 5-5.
PSHT Heatup and Cooldown Pressure-Temperature Limits Imposed by Several Reactor Yessel Regions and Composite L i mi t C u rv e..........................
5-11 5-6.
PSHT Heatup and Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Canposite Limit Curve for 5 EFPY.............~.......
5-12 5-7.
Determination of Reactor Core Operation Pressure-Temperature Curve for 5 EFPY per Appendix G to 10 CFR 50 5-13 5-8.
Normal Operaiton Heatup Pressure-Temperature Limit Curves for Typical Plant Technical Specifications. Applicable 5-14 up to 5 EFPY 5-9.
Normal Operation Cooldown Pressure-Temperature Limit Curve for Typical Plant Technical Specification Applicable up to 5 EFPY...........................
5-15 5-10.
Inservice Leak and Hydrostatic Test Heatup and Cooldown Pressure-Temperature Limit Curve for Typical Plant Technical Specifications. Applicable up to 5 EFPY,......
5-16 5-11.
PSHT Pressure-Temperature Limit Curve for Typical Plant....
5-17 i
6-1.
Ty p i c al D PF AD Cu rv e...................... 6-6 6-2.
Assessment Point Illustration................. 6-6 6-3.
Failure Assessment Diagram Preocedure Applied to Typical 6-9 Beltline Region Weld 6 -4.
Typical Beltline Weld JR Curve 6-10 6-5.
Typical Resultant Tearing Pressure Prediction.........
6-11 t
a 4
- xi1 -
Babcock & Wilcox A McDernett tompany
O e
1.
INTRODUCTION i
1.1.
Background
On July 17, 1973, a new appendix to 10 CFR 50, entitled "Appendix G - Frac-ture Toughness Requirements" was published in the Federal Register.
10 CFR 50, Appendix G has been revised in subsequent years.
This report reflects the revised criteria including effective issue July 26, 1983.
This appen-dix specifies minimum fracture toughness requirements for the ferri tic materials of the pressure-retaining components of the reactor coolant pres-sure boundary (RCPB) of water-cooled power reactors and provides specific guidelines for deterraining pressure-temperature operational limitations on the RCPB.
The toughness and operational requirements are specified to provide adeqt. ate margins of safety during any condition of normal opera-tion, including anticipated operational occurrences and system hydrostatic tests to which the RCPB may be subjected over its service lifetime.
Al-though the requirements of Appendix G became effective August 13, 1973, they are applicable to all boiling water and pressurized wate r-cooled nuclear power reactors, including those under construction or in operation on the ef fective date.
At the time 10 CFR 50, Appendix G, became effective, immediate compliance with some of its provisions was not possible for plants whose pressure boundary components were ordered in accordance with an edition or addenda of Section III of the ASME Boiler and Pressure Vessel Code (hereafter ASME Code) published before the Summer 1972 Addenda.
For these plants, neither the fracture toughness data required by Appendix G nor the material for per-forming toughness tests is available.
Also, the stress calculations re-quired to quantitatively define the allowable pressure at any given tempera-ture were not readily available.
Appropriate, conservative methods of compliance for these plants have been developed and are described in this report.
1-1 Babcock &WHcom a McDermott company
~
1.2.
Scope and Organization This report presents B&W's ' practices, methods, and criteria for compliance with the requirements of 10 CFR 50, Appendix G.
It is applicable to all current B&W nuclear steam systems (NSSS).
The definitions and terminology of 10 CFR 50, Appendix G, and the ASME Code are used whenever appropriate.
The report is divided into seven parts and is summarized in Part 7.
Part 2 describes the reactor coolant pressure boundary (RCPB) and includes a list of the components and ferritic materials used in their construction.
Part 2 also describes the operational modes of the RCPB related to nonductile failure for each of the loading conditions for which pressure-temperature limit curves are required.
Part 3 presents the fracture toughness properties of the ferritic materials of the RCPB.
These materials are grouped as follows:
1.
Ferritic mstorials other than (6) bolting and (b) type 403 stainless steels 2.
Bolting materials
~
3.
Type 403 stainless steel For the first group, Part 3 describes methods for (1) detertaining the unir-g radiated reference temperature (RTNOT) for the ferritic materials and the l
unieradiated Charpy upper shelf energy (Cy SE) level of the beltline region U
material s.
The justification for use and acceptance criteria for unirrad-iated beltline region materials with CyVSE lower than 75 f t-lbs are pre-sented.
For the second group, bolting materials, Part 3 presents justification for allowing the lowest service temperature, and the minimum preload tempera-ture to be 40F.
The impact properties of these materials are also pre-sented.
For the third group of materials, Part 3 includes a demonstration of ade.
quate fracture toughness properties.
j Part 4 presents the basis for a step-by-step description of the calcula-
]
tional procedure to determine the pressure-temperature limitations of the reactor coolant system; this is done to ensure adequate fracture toughness i
under the loading conditions of interest.
1-2 Bath:ock & Wilcox 4 MCOtrmott (Cepay
r Part 5 gives an exarple of beginning-and end-of-life pressure-temperature l
limit curves that were developed using the material properties in Part 3 and the calculational procedure of Part 4.
Similar curves were developed i
for each plant and conservatively adjusted for use in the Technical Speci-i I
fications issued by the Nuclear Regulatory Commission (NRC) as a part of l
the plant operating license.
lypical limit curves, as they appear in the j
lj Technical Specifications, and the limit curve for the preservice system l
hydrostatic test are shown in Part 5.
1 Part 6 presents the supplemental analysis perfonned in the event a reactor r
t i
vessel beltline is predicted to be below 50 f t-lbs upper shel f.
This l
l analysis is an elastic plastic fracture mechanics assessment confinning
[
that the vessel has sufficient toughness to preclude ductile tearing insta-7 j
bili ty.
1 i
i i
l i
i 1
t l
1 i
f i
I f
{
L l
l L
l i
1 l
l 1-3 Babcock & WHeen i
4MCDU Ot t (Ompany
i s
2.
REACTOR C001. ANT PRESSURE EOUNDARY 2.1.
Components The RCPB is defined by NRC Regulation 10 CFR 50.2, (v) as follows:
"' Reactor coolant pressure bounda ry' means all those pres-sure-containing components of boiling and pressurized water-cooled nuclear power reactors, such as pressure vessel, piping, pumps, and valves, which are:
(1) Part of the reactor coolant system, or (2) Connected to the reactor coolant sy stem, up to and including any and all of the following:
(1) The outermost containment isolation valve in system piping which penetrates primary reactor containment, (ii) The second of two valves nonnally closed during normal reactor operation in system piping which does not penetrate primary reactor containment, (iii) The reactor coolant system I afety and relief valves.
For nuclear power reactors of the direct cycle boiling water type, the reactor coolant system extends to and includes the outennost containment isolation valve in the main steam and feedwater piping."
The reactor coolant system (RC system) for B&W nuclear power plants is made up of the following components:
reactor vessel, steam generators, pressur-izer, reactor coolant pumps, valves and interconnecting piping.
The RC system contains and circulates reactor coolant at the pressure and velocity necessary to transfer the heat generated in the reactor core to the sec-ondary fluid in the steam generators.
The other pressure-containing portions of the RCpB are the auxiliary system components.
These include the makeup and purification system piping and valves (including RC pump seal injection lines); the emergency core cooling 2-1 Babcocer &WHees
& MCOttmott (Omp4ny
0 system high-and low-pressure and core flooding injection piping and core flooding injection piping and valves; the vent, drain, and other piping and valves used for maintaining the RC system; and the incore instrumentation on piping.
-Portions of the RCPB are exempted from the requirements for Class 1 compo-nents of ASME Code Section III by foc ?"te 2 to NRC Regulation 10 CFR 50.55a, which reads as follows:
Components which are connected to the reactor coolant system and are part of the reactor coolant pressure boundary defined in 50.2(v) need not meet these requirements, provided:
(a)
In the event of pos tulate failure of the component during normal reactur operation, the reactor can be shut down and cooled down in an orderly manner, assuming makeup is provided by the reactor coolant makeup system only, or (b) the component is or can be isolated fr:xn the reactor coolant system by two valves (both closed, both open, or one closeo and the other open).
Each open valve must be capable of automatic actuation and, assuming the other valve is open, its closure tine must be such that, in the event of postulated failure of l
the component during normal reactor operation, each valve l
remains operable and the reactor can be shut down and cooled down in an orderly manner, assuming makeup is provided by the reactor coolant makeup rystem only.
Components of the RCP8 included under this exemption provision are gener-ally designed and fabricated in accordance with the requiremer.ts for Class 2 components in ASME Code Section III (see Regulatory Guide 1.29, "Quality I
Group Cl asM fications and Standards fo r Water,
Steam.
and Radioac-l tive-Waste containing Components of Nuclear Power Plants").
None of these components are constructed of ferritic material except in some instances the core flood tanks, which are carbon steel in some B&W plants.
Although the core flood tanks are isolated from the RC system by two valves during nonnal operation, connecting piping to the tanks (1-inch lines for nitrogen addition fill and drain) does penetrate reactor containment.
Therefore, the system is part of the RCPB to the outermost contai nment i sola tion 2-2 Babcock &Wilcox A McDermott Company
valvd Since these tanks are isolated fran the RC system during all condi-tiuns of normal operation, includi ng ant hi p., ted operational occurrences, they need not be considered in developing the RC system pressure-tempera-ture limita' ions and are not discussed in this report.
2.2.
Ferritic M3terials and RCPB Operational Parameters The ferritic materials used in construction of the RCPB for B&W nuclear power plants are li sted for each component in Table 2-1.
The pressure boundary of the RC system is fabricated primarily from ferritic materials, while that of the auxiliary systems is fabricated primarily from austenitic material.
Consequently, the RC sys tem components are the only ones that require special protection against nonductile failure and that must comply with the fracture toughness requirements of 10 CFR 50, Appendix G.
This protection against nonductile failure is ensured by imposing pressure-temperature lim-itations on operation of the RC system.
The margin of safety is controlled by the maximum calculated allowable pressure at any given temperature.
The following loading conditions require pressure-temperature limits:
1.
Norinal operations including bolt preloading, heatup and cooldown.
2.
Preservice system hydrostatic test.
3.
Inservice system leak and hydrostatic tests.
4.
Reactor core operation.
l To impart a better understanding of the required protection against nonduc-tile failure, typical operational parameters of the RC system are described l
in the following paragraphs for each of the loading conditions.
2.3.
Normal Operation 2.3.1.
Bolt Preload l
During bolt preload, the reactor vessel closure studs are tensioned to the specified load.
Bolt preloading is not allowed until the reactor coolant tempe rature and the volume tric average temperature of the closure head region (including the studs) is higher than the specified minimum preload j
tempe ra ture.
After the studs are tensioned, system pressure can be increased by the pressurizer until it is above the net positive suction head (NPSH) requi red for RC pump operation.
The heatup transient begins when the RC pumps are started.
Babcock & Wilcox l
2-3
,uco,,morecomp,ny l
s I
2.3.3.
Heatup During heatup the RC system is brought from a cold shutdown condition to a hot shutdown condition.
The heat sources used to increase the temperature of the system are the RC pumps and any residual (decay) heat from the core.
Normally, when the pumps are started, the temperature of the water in the pressurizer is about 400F; this corresponds to the pressure in the RC sys-tem of about 300 psig.
The coolant temperature is at or above the minimum specified bolt preload temperature.
Initially, the reactor coolant tempera-ture may be as low as room temperature for initial core loading or as high as 130F for subsequent refueling.
At any given time throughout the heatup transient, the temperature of the reactor coolant is essentially the same throughout the system except, of course, in the pressurizer.
The system pressure, as controlled by the pressurizer heaters is maintained between the minimum required for RC pump NPSH and the maximum to meet the fracture toughness requirements.
The heat-up rate is maintained below the maximum rate used to establish the maximum allowable pressure-temperature limit curve.
2.3.3.
Cooldown RC system cooldown brings the system from a hot to a cold shutdown condi-tion.
The cooldown is normally accomplished in two phases.
The first phase reduces the fluid temperature from approximately 550F to below the design temperature of the decay heat removal system (approximately 300F).
This temperature reduction is accomplished using the steam generators but bypassing the turbine and dumping the steam directly to the condenser.
Once below its design temperature (and pressure), the decay heat removal system (DHRS) is activated ia the second phase to further reduce the reactor coolant temperature to that desired.
Before cooldown, the RC system temperature is maintained constant by bal-ancing the heat removal rate from the steam dump with the heat contributed by the RC pumps and core decay heat.
The system pressure is maintained by the pressurizer.
The cooldown is nonnally ini tiated by stopping one RC pump in each loop.
The two remaining pumps provide coolant circulation through both steam generators, and the turbine steam bypass flow controls the cooldown rate.
The primary pressure during cooldown is controlled with 2-4 Babcock & Wilcox a McDermott company
the pressurizer heaters and spray.
After cooling down below the DHRS design temperature and pressure, the cooling mode is changed from the steam generators to the DHRS.
Before the switch, the RC system pressure is below 625 psig (207, of the preoperational system hydrostatic test pressure) and below the DHRS pressure but above the pressure required for the RC pumps to operate.
To minimize the thermal shock on the RCPB, the two RC pumps remain in opera-tion as the water flow of the DHRS is initiated.
The DHRS flow rapidly mixes with the reactor coolant, but during this period, the indicated RC temperature may fluctuate until mixing is complete.
After the switch is completed, the RC pumps are stopped.
During this phase, the cooldown rate is controlled by the temperature and flow of the DHRS.
2.4.
Preservice System Hydrostatic Test Prior to initial operation, the RC system is hydrostatically tested in ac-cordance with ASME Code requirements.
During this test, the system is brought up to an internal pressure not less than 1.25 times the system design pressure.
This minimum test pressure is in accordance with Article NB-6000 of ASME Section III.
Since the system design pressure is 2500 psig, the preservice system hydrostatic test pressure is 3125 psig.
Ini-tially, the RC system is heated to a temperature above the calculated mini-mum test temperature required for adequate fracture toughness.
This heatup is accomplished by runnir1 the RC pumps.
The pressurizer hedters are used to heat the pressurizer u the required temperature.
Before the test temp-erature is reached, the pressure is maintained above the NPSH required for the RC pumps but below the maximum allowable pressure for adequate fracture toughness.
When the test temperature is reached, the RC pumps are stopped and RC makeup water is added to fill the pressurizer.
The test pressure is then reached using either the pressurizer heaters or the hydrostatic pumps connected to the RC system.
The test pressure is held for the minimum spec-ified time, and the examination for leakage follows in accordance with the ASME Code.
2.5.
Inservice System Leakage and Hydrostatic Tests When inservice system leakage tests are required, the system is brought from a cold to a hot shutdown condition.
The means of heating the system 2-5 Babcock & Wilcox 4 MCDetmott Compar1y
-_ =_.
and increasing the pressure are the same as those used during normal heat-up.
If it is necessary to cool the system down af ter either test, normal cooldown procedures are used.
These tests are conducted in accordance with the requirements of ASME Section XI, Article IWA-5000.
The test pressure fo r the inservice leakage tests is the pressure that, for the component j
located at the highest elevation in the system, is no less than the system l
nominal operating pressure at 100% rated reactor power.
For the inservice hydrostatic test, ASME Section XI gives a table (Table IWB-5222-1) of the minimum test pressure versus the test temperature at which the system must be tested.
The test temperature for both the inservice leakage and hydro-static tests is determined by the requirements for fracture toughness.
2.6.
Reactor Core Operation The reactor core is not allowed to become critical until the RC system fluid temperature is above 525F except for brief periods of low-power physics testing.
This tempe rature is much higher than the minimum per-missible temperature for the inservice system hydrostatic pressure test, and it is al so at least 40F above the cal cula ted minimum temperature required at nomal pressure for operation throughout the service life of the plant.
l l
2-6 Babcock & Wilcox a WDermott corrpany
Table 2-1.
Ferritic Materials Used in Reactor Coolant Pressure Boundary Component Material Reactor Vessel Pla tes SA 533, Grade B, Class 1 Forgings SA 508, Class 2; SA 182, Grade F6 Bolting SA 540, Grade B-23 or -24 Wel ds SFA 5.5, SFA 5.17 Bars A276 Type 403 (Code Case 1337 or N-4)
Steam Generator Plates SA 533, Grade B, Class 1; SA 516, Grade 70 Forgings SA 508, Class 1 Bolting SA 540, Grade B-23 or -24 Wel ds SFA 5.5, SFA 5.17 Pressurizer Plates SA 533, Grade B, Class 1 Forgings SA 508, Class 2 Bolting SA 540, Grade B-23; SA 320, Grade L43 Wel ds SFA 5.5, SFA 5.17 Reactor Coolant Piping Plates SA 516, Grade 70 Forgings SA 105, Grade 2 Seamless Pipe & Tubing SA 106, Grade C Wel ds SFA 5.5, SFA 5.17 Reactor Coolant Pump Forgi ngs SA 508, Class 2; SA ~50, Grade LF2 Bolti ng SA 540, Grades B-21, -23, -24 Valves Forgings SA 105 Grade 2 2-7 Bat > cock & Wilcox a McDermott company
O
- 3. MATERIAL PROPERTIES 3.1.
Impact Properties of Ferritic Materials To determine the pressure-temperature operating limitations for the RCPB the reference nil-ductility temperature (RTNDT) of the ferritic materials must be established.
The RTNDT is needed to calculate the critical stress intensity factor (KIR).
In ASME Appendix G, XIR is related to temperature, T, and to RTNDT by the following equation:
XIR = 26.77 + 1.223 exp[0.0145(T - RTNDT + 160)]ksi /in.
This relationship is applicable only to ferritic materials that have a spec-ified minimum yield strength of 50,000 psi or less at room temperature.
Since the impact properties of the beltline region materials of a reactor vessel will change throughout its li fetime, periodic adjustments are re-quired on the pressure-temperature limit curves of the RCPB.
The magnitude of these adjustments is proportional to the shift in RTNDT caused by neu-tron fluence.
Therefore, it is essential to determine the radiation-in-l duced ARTNDT of the beltline region materials.
Since the ARTNDT is based on the temperature shift of the Charpy curves mea-sured at the 30 ft-lb level, it is necessary to know, by analysis or from the results of the material surveillance program, the magnitude of the Charpy 30 ft-lb shift.
l 3.1.1.
Detennination of RTNDT 3.1.1.1.
ASME Code Method The RTNDTs of the ferritic materials, which were specified and tested in accordance with the fracture toughness requirements of the ASME Section III Sumer 1972 Addenda (to 1971 Edition) or later Editions and Addenda, are de-termined as required by that Code.
When sufficient material is available, the RTNDTs of the beltline region materials (which were specified and 3-1 Babcock & Wilcox A McDermott company
tested in accordance with an Edition or Addenda of ASME Section III earlier than the Summer 1972 Addenda) are obtained by testing specimens oriented normal to the principal working direction.
The test procedure is in accor-dancc-with ASME Section III, paragraph NB 2300 (Summer 1972 or later Edi-tion and Addenda).
3.1.1.2.
Estimating Method The RCPBs of several plants were designed and constructed in accordance with the requirements of an edition or addenda of ASME Section III issued before the Summer 1972 Addenda.
Except for the beltline region materials for which sufficient test material is available, the RTNDTs of the ferritic materials must be estimated.
This is necessary because obtaining the test data required for the exact determination of RTNDT was not required by the applicable ASME Code.
Generally, drop weight tests were not performed, and the Charpy Y-notch tests were limited to "fixed" energy level requirements for specimens oriented in the longitudinal (principal working) direction at a temperature of 40F or lower.
To obtain an RTNDT estimate that is appropriately conservative, B&W has col-lected and evaluated the data from tests conducted on pressure-retaining ferritic materials to which the new fracture toughness requirements were ap-plied.
3.1.1.3.
Estimated RTNOT In the preceding section pertinent impact data for each type of ferritic ma-terial are discussed as a basis for estimating conservative RTNDTs.
E s ti-mated RTNDTs are needed for all materials that were specified to meet the requirements of an Edition or Addenda of ASME Section III earlier than the Summer 1972 Addenda.
This section summarizes the data and the estimated RTNDT of the ferritic materials used in construction of the RCPB.
The data are summarized in Table 3-1.
For each type of material, the table lists the number of cases considered; the highest measured RTNDT; the aver-age of tne measured RTNDTs; the estimated RTNDT; and the difference t,etween the average measured and the estimated temperatures.
3-2 Babcock & Wilcox a McDermott company
~
3.1.2.
Determination of Charpy V-Notch Level 3.1.2.1.
Specified Method Appendix G to 10 CFR 50 requires complete characterization of the unirradi-ated impact properties of all the beltline region materials of the reactor vessel.
This includes determination of RTNDT and Charpy (Cy) test curves for the directions normal to and parallel to the principal working direc-tion (other than the thickness direction).
Appendix G also requires a min-imum Charpy upper shelf energy (Cy SE) of 75 ft-lb for all beltline region V
materials unless it is demonstrated that lower values of upper shelf frac-ture energy provide an adequate margin against irradiation induced degra-dation.
To comply with Appendix G, the beltline region materials (not including HAZ) of reactor vessels for later plants meet the following test require-ments:
In addition to the Charpy V-notch impact tests needed to de-tennine RTNDT, 15 Charpy V-notch impact tests shall be con-ducted in each required di rection (for base metals the re-quired di rections are nonnal and parallel to the principal direction in which the materi al was worked, other than the thickness direction).
The tests shall be conducted at appro-priate temperatures over a temperature range sufficient to de-fine the C test curves (including upper shel f levels) in terms of both fracture energy and lateral expansion.
Three specimens shall be tested at each test temperature for the de-tennination of RTNDT.
The Charpy upper shelf energy shall be determined as follows:
l (1) Two sets of three Charpy specimens each shall be tested at two temperatures at which the percent of shear frac-ture is approximately 95%.
The Charpy upper shelf energy shall be the higher average energy value of the two sets of Charpy specimens.
l (2) If either of the two average upper shelf energy values of step (1) is below 75 f t-lb, another set of three Charpy specimens shall be tested at a temperature at least 50F higher than the highest temperature of step (1).
The l
Charpy upper shelf energy shall be the highest average value of the three sets of Charpy specimens.
The location and orientation of the impact test specimens t
l shall comply with the requirements of paragraph NB-2322 of Section III of the ASME Code.
l l
3-3 Babcock & Wiscos a MtOttmott company
6 The requirenents for the minimum Cy SE are described in section 3.1.3.3.
V The requirements above are also met for the HAZ of the beltline region base me tal( s) that are selected to be monitored by the reactor vessel surveil-lance program.
The requirements are not specified for the HAZ of the other beltline region materials because the ASME Code (Paragraph NB-4335 of the Winter 1974 Addenda) deleted the requirements for toughness testing of HAZs in the weld procedure qualification tests.
B&W has elected to follow the new ASME requirements.
For the beltline region materials of reactor vessels that were specified in accordance with the requirements of an Edition or Addenaa of ASME Section III issued before the Summer 1972 Addenda, the complete C test curves, in-y cluding C VSE, is detemined when the material is included in the reactor y
vessel materi al surveillance program.
For the beltline region materials that are not included in the surveillance program, and when sufficient ma-terial is available, the C test curve and USE are detennined only in the y
direction nomal to the principal working direction.
No minimum C VSE is y
requi red, other than the 50 f t-lbs/35 mils of lateral expansion for the beltline region materials of these reactor vesrals, one of the conditions required to establish RTNDT.
When the unirradiited Cy SE of these materi-V als is below 75 f t-lb, the currently accepted pt ocedure is applied to pre-dict the end-of-service CyVSE.
3.1.2.2.
Estimating Method The Cy USE must be estimated for reactor vessel beltline region materials that were specified in accordance with the requirements of an Edition or Addenda of ASME Section III issued before the Summer 1972 Addenda and for which insufficient material is available for testing.
All available data from tests conducted on reactor vessel beltline region materials were col-lected and evaluated in order to obtain an appropriately conservative esti-
~
mate.
Not all the data were obtained in accordance with the methods speci-fi ed in section 3.1.2.1 since in some cases the absorbed ene rgy wa s obtained only at one temperature.
3.1.2.3.
Estimated Cv SE u
The data used for estimating conservative Cy SE is discussed in the preced-V Cy SE is needed for all of those beltline i ng section.
The estimated V
3-4 Babcock & Wilcox a MCDermott company
region materials for which test material is not available, i.e., for which the actual C VSE data and the estimated energy for each type of beltline re-y gion material are summarized in Table 3-2.
For each type of material, the table lists the number of tested heats, the lowest measured, average mea-sured, and estimated C VSEs and the overage difference between the esti-y mated and measured C USE.
y 3.1.3.
Radiation Effects 3.1.3.1.
Adju3tment of RTNDT Adj ustment of the RTNDT to acconmodate the radiation-induced changes in fracture toughness of beltline region materials is an important factor in devel opi ng pressure-temperature limits.
Correlations have been developed for predict 1r.g the radiation-induced RTNDT to be used in adjusting the ini-tial RTNDT for pressure-temperature analyses.
These correlations are not perfected and, therefore, subject to continuous updating as additional data and infonnation is developed.
The methodology used to adjust the RTNDT values as used in developing pres-sure-temperature limits 'will be in accordance with the currently accepted licensing procedures.
The method used wil l be referenced in all pressure-temperature analyses and will be reported in the Owners licensing documents.
3.1.3.2.
Decrease in Cv0SE Neutron irradiation of the beltline region materials cause a decrease in CyVSE.
Correlations have been devel oped fo r predicting this decrease in Charpy USE.
These correlations are not perfect and, therefore, are subject to updating as additional data and information is obtained.
The methodology used to predict the decrease in CyVSE (used in the evalua-tion of beltline region materials) will be in accordance with the currently accepted licensing procedures.
The method used will be referenced in all analyses and will be reported in the Owners licensing documents.
3.1.3.3.
Acceptance Criterion for Unteradiated Cv SE u
l Appendix G to 10 CFR 50 requires that the CyV3E of the unieradiated belt-line region materials be equal to or greater than 75 f t-lb except if it is demonstrated by appropriate data and analyses that lower values still pro-l vide adequate margin for degradation resulting from neutron irradiation.
l 3-5 Babcock & Wilcom a McDermott company
This section demonstrates that for some beltline region materials, a CyVSE lower than 75 f t-lb still provides an adequate margin for degradation from i rradi ation.
This section also presents an acceptance criterion for CyUSE lower than 75 f t-lb which is applied to later plants.
The beltline region of the reactor vessel includes all the ferritic materi-al in the reactor vessel that (1) directly surrounds the effective height of the active core and (2) adjacent regions of the reactor vessel that are predicted to experience sufficient neutron radiation damage to be consid-ered in selecting the limiting material with regard to radiation damage.
The beltline region material above and below the effective height of the fuel element assablies are irradiated to a neutron fluence received by the materi al directly surrounding the fuel element assemblies.
Since not all beltline region material is subjected to the same neutron fluence, it is not neces sary for all of this material to have a CyVSE greater than 75 f t-l b.
Also, the radiation-induced drop in CyUSE depends not only on the neutron fluence but on the material's enemical composition.
The required CyVSE of unirradiated beltline region materials is defined in terms of the material's chemical composition and the predicted end-of-service neutron fluence to which the material will be subjected.
Complete Charpy V-notch impact curves are required for all of the unirradi-ated beltline region materials used in later reactor vessels.
The test re-quirements are in accordance with Appendix G to 10 CFR 50 and are described i n section 3.1.2.1.
C VSE requirments are as follows:
y 1.
The C VSE of the beltline region materials directly surrounding the ef-y fective height of the fuel assemblies shall be equal to or greater than 7 5 f t-l b.
2.
The CyVSE of the beltline region materials above and below the effec-tive height of the fuel assemblies shall be equal to or greater than the sum of the following energies:
a.
The energy calculated using the material's chemical composition, end-of-service neutron fluence at the 1/4T vessel wall location, and an accepted prediction technique which will provide an end of service life CyVSE no less than 50 ft-lbs.
b.
The energy equivalent to Si, of the energy calculated in step a.
3-6 Sabcock & Wilcox a McDermott comparty
e The minimum C USE above provides adequate margin for degradation from irrad-y iation.
All the beltline region materials of later reactor vessels have been specified to have a low copper content (<0.10%), and the predicted drop in C VSE is very small for the neutron fluence of interest.
y 3.2.
Impact Properties of Bolting Materials 3.2.1.
Code Requirements Appendix G to 10 CFR 50 requires that materials for bolting and other fas-teners mest the ASME Code.
In the early editions of the ASME Code, up to and including the Winter 1971 Edition, it was required that the bolting ma-terials exhibit a "fixed" minimum average energy at a temperature of 10F.
One specimen in a set of three was allowed to be less than the fixed f t-lb value, but not less than the fixed value minus 5 f t-lb.
In the Summer 1972 Addenda to the 1971 Edition, the fracture toughness requirements for bolt-ing materials were changed to be consistent with the requirements of Appen-dix G except that no requirements were made in tenns of absorbed energy (ft-lb).
The requirements were changed again by the Summer 1973 Addenda to the 1971 Edition.
In this revision and subsequent editions of ASME Section III, 45 ft-lb absorbed energy was required only for bolting materials hav-ing a nominal diameter greater than 4 inches.
All bolting materials ordered after the effective data of Appendix G to 10 CFR 50 (August 16, 1973) meet the requirements of Appendix G.
Bolting mate-rials ordered before this date must meet the requirements of the applicable i
ASME Code.
3.2.2.
Estimating Method To establish the minimum preload temperature and the lowest service temper-ature of a pressure-retaining component, it is necessary to know the lowest temperature at which the bolting materials have adequate fracture tough-ness.
This lowest temperature is either the temperature at which the bolt-ing materials exhibit a 25-mil lateral expansion and 45 ft-lb absorbed energy or the temperature at which the bolting materials are at the Cy SE.
ll For bolting materials of pressure-retaining components ordered before August 16, 1973, it is necessary to estimate the lowest temperature at which these Charpy impact properties are met.
The preload temperature and 3-7 Babcock & Wilcom a McDermott company
the lowest service temperature are defined by the applicable equipment specification for components ordered after August 16, 1973.
Impact data from 13 heats of SA 540 Class 3 bolting were evaluated in order to estimate the lowest temperature at which bolting materials have adequate fracture toughness.
The principal criteria defining the fracture toughness requirements for the bolting materials used in the reactor coolant pressure boundary are described in WRC Bulletin 175.3 The fracture mechanics analy-sis performed and described in WRC Bulletin 175 shows that for the refer-ence flaw size of 0.3 inch (nominal diameter over 3 inches), the required fracture toughness (XIC) is about 125 ksi6. for bolting materials with a specified minimum yield strength of 130 ksi.
To protect against nonductile failure, fracture toughness values exceeding 125 ksi6. would be needed at the lower service temperature at which maximum Code-allowed stresses occur.
In WRC Bulletin 175 KIC versus Cy energy correlations were used to estimate the Cy energy that would correspond to 125 ksiMi.
The KIC versus Cy cor-relations were those of Barson and Rolfe.4 Their empirical correlations are between slow-bend KIC tests and the results of standard Charpy V-notch impact tests for the transition-temperature and upper shelf regions.
The transition-temperature KIC-CVN correlation is (KIC)2 = 2(CVN)3/2 (1) t and the upper shelf KIC-CVN correlation is 2
=h CVN -
(2)
L Y.
1 The relationship in equation 1 suggests that at the transition-temperature region of the Charpy curve, 41 f t-lb corresponds to 125 ksi riT.
For the upper shelf region of the Charpy curve, the relationship of equation 2 re-lates 28 and 30 f t-10 to 125 ksiMi. for bolting materials having yield strengths of 160 and 130 ksi, respectively.
3-8 Babcock & Wilcox a uconmore company
Even though two of the bolting material heats evaluated do not meet the re-quirements of Appendix G, the materials have adequate fracture toughness to provide a conservative margin of safety against nonductile failure.
At
+40F, the bolting materials evaluated are at the upper shelf region of their Cy test curves.
For the bolting ma terial s under consideration, CyVSEs of 28 ft-lb would have sufficient fracture energy to prevent failure because the upper shelf KIC-CVN correlation shows that 28 f t-lb corresponds to 125 ksi /in.
The lowest C VSE of the data collected, 42 f t-l b, corre-y spondt to a fracture toughness value of 165 ksi/in.
To ensure adequate mar-gin of safety, the lowest service temperature and the minimum preload tem-perature are defined to be higher than 40F.
3.3.
Impact Properties of Type 403 Modified Steel 3.3.1.
Code Requirements Appendix G to 10 CFR 50 requests that the adequacy of the fracture tough-ness properties of ferritic materials such as Type 403 modified stainless steel be demonstrated to the Commission on a case-by-case basis.
The Type 403 modified steel is used as a RCPB material in the motor tube of the con-trol rod drive mechanism.
This section demonstrates that, for this appli-cation, the material has adequate fracture toughness for protection against non-ductile failure.
The nominal wall thickness of the motor tube section of interest is more than 1/2 inch and less than 5/8 inch.
In the early editions of ASME Sec-tion III up to the Winter 1971 Addenda to the 1971 Edition, materials with a nominal section thickness of 1/2 inch or less did not require impact test-ing.
Starting with the Summer 1972 Addenda, the nominal section thickness increased to 5/8 inch or less.
Thus, in the early editions of ASME Section III, the Type 403 modified steel required impact testing, but in the new editions it does not.
However, since this material was selected for use, B&W has ordered it to meet the impact toughness requirements of ASME Sec-l tion III, as if its nominal wall thickness exceeded 5/8 inch.
For materi-als crder to ASME Section III, Summer 1972 and later Addenda, the imposed acceptance standard for nominal wall thicknesses from 5/8 to 3/4 inch, in-clusive, is presented in Paragraph NB-2332.
The material has also been specified to meet the requirements of SA 182 Grade F6 (forgings) or ASTM A276 (bars) as modified by ASME Code Case 1337.
3-9 Babcock & Wilcom A MCDermott company
When ordered according to the early revisions of Code Case 1337 (including Revision 6) and to the early editions of ASME Section III, tho Type 403 modi fied forgings or bars were required to be impact-tested at 20F.
The minimum average energy of a set of three Charpy V-notch specimels was 35 f t-lb, with one specimen allowed to be less than 35 but not less than 30 ft-lb.
For both forgings and bars, the Charpy specimens we:,: oriented in the axial (longitudinal) direction.
In the Summer 1972 Addenda to the 1971 Edition of ASME Section III, the fracture toughness require,nents of all pressure boundary ferritic sterials changed; however, no acceptance criterion was given fo r the martensitic high-alloy chromium steels, such as Type 403 modified steel.
A year later, the Summer 1973 Addenda re-ettablished the acceptance criteria for the type 4XX steels.
Beginning with this addenda, the fracture toughness require-ments and acceptance criteria for the type 4XX steels are described in Para-graph NB-2332 of ASME Section III.
This paragraph requires that three Charpy V-notch specimens be tested at temperatures lower than or equal to the lowest service temperature.
The lateral expansion of each specimen must be equal to or greater than 20 mils.
The test temperature has been specified as equal to or less than 40F.
The orientations of the specimens are transverse (noriaal to principal working direction) for the forgings and axial for the steel bars.
The fracture toughness requirements of Code Case Summer 1337, starting with Revision 7, are the same as those of ASME Sec-tion III, Summer 1973 Addenda to the 1971 Edition.
3.3.2.
Demonstration of Adequate Toughness It is B&W's position that the fracture toughness requirements of the new editions of ASME Section III provide adequate protection against nonductile failure.
The proof of adequate toughness is based on demonstrating that the Type 403 modified steels used in the construction of conponents de-signed to an Edition or Addenda of ASME Section III nrior to the Summer 1973 Addenda meet or exceed the toughness requirenents of that Addenda.
Data from 1S lots of SA 186 F6 forgings and 15 lots of ASTM A276 bars were ev al ua ted.
Based on these data, the lowest service temperature of the con-trol rod drive mechanism can be as low at 40F; however, for additional pro-l tection against non-ductile failure, B&W has defined the component's lowest service temperature at 100F.
This specified lowest service temperature is 3-10 Babcock & Wilcox A MCDermott Company
60F above the temperature at which the fracture toughness requirements are specified and met.
The additional 60F provides margins of safety beyond that required by the ASME Code and by Appendix G to 10 CFR 50.
3.4.
Supplemental Fracture Toughness Troperties In the event the beltline naterial reaches a radiation level which causes the predicted Charpy upper shelf energy value to decrease below 50 f t-lb at 1/4T, supplemental fracture toughness data will be obtained to assess reac-tor pressure vessel integrity.
The data are used to demonstrate equivalent margins of safety as established in N)pendix G of ASME Code.
3.4.1.
Terminology Related to Ductile Fracture Analysis The terminology used in the development of material properties for analysis of the reactor vessel resistance to ductile fracture will be in accordance with the following standards.
3.4.1.1 Mechanical Properties -- ASTM Specification E6, Standard Defini-tions of Terms Relating to Methods of Mechanical Testing 3.4.1.2 Fracture Toughness Properties -- ASTM Specification E616. Standard Terminology Relating to Fracture Testing i
3.4.2.
Fracture Toughness Properties of Ductile Materials When the beltline region materials of the reactor pressure vessel reach an l
irradiation level which causes the Charpy upper shelf energy value of the l
material to decrease to a value below 50 ft-lb, supplemental fracture tough-l l
ness data to assess re r vessel integri tt are required by 10 CFR 50.
These data are used to provide input to the elastic-plastic fracture mech-anics analysis as described in section 6.
The data base for this fracture nachanics analysis is being developed in the integrated reactor vessel ma terial surveillance progran described in BAW-1543 and the interpretation of the nuterials data obtained f ran thi s surveillance program will be presented in RVSP reports.
The data that are most important to the analysis are those which define the initiation of ductile tearing and the resistance of the material to ductile tearing as a function of crack growth.
The interpretation of the data is presented in the load-displacenent curves obtained fran the individual tests and the 3-11 Babcock & Wilcox a McDermott company
resulting J-R curves derived from the data.
Supporting data is obtained from the stress-strain curves of the tension tests.
These data are analyzed to obtain the true stress-true strain curves to provide the work hardening coefficients.
The actual ma terial properties used in the establishnent of the reactor pressure vessel operating limitations and supporting references will be re-ported in the appropriate licensing document.
3.4.3.
Relationship Between Fracture Toughness Properties and the Fracture Mechanics Analysis The technical approach used in Appendix G of 10 CFR 50 is to establish the reactor vessel operating limitations with adequate margins of safety using a fracture mechanics analysis assuming that the vessel sterial may behave in a non-ductile manner.
In the temperature region characterized by the Charpy lower shelf and transition region a LEFM analysis is required using the procedure described in Appendix G of 10 CFR 50.
In the Charpy upper shelf temperature region no additional fracture mechanics analysis is re-quired as long as it is demonstrated that the Chany upper shelf energy is greater than 50 f t-lbs.
If the Charpy energy is predicted to drop below the 50 f t-lb level, it is required to provide supplemental fracture tough-ness information and an analysis to demonstrate an equivalent margin of safety as required by Appendix G of 10 CFR 50.
This necessitates the use of elastic-plastic fracture mechanics analysis methods.
As part of the required suppl emental analysis, a criteria must be estab-lished such that a smooth transition will occur in the vessel operating lim-itations between the required LEFM analysis and the supplemental EPFM analy-sis.
A conservative approach to establishing this transition is to perform both LEFM and EPFM analyses and establish the vessel operating limits as the lower bound of the two results.
The temperature at which the transition is made fran the LEFM coalysis to the EPFM analysis is therefore defined as the temperature at which the al-lowable pressure versus temperature curvu cciculated by the two procedures intersect.
Since the allowable pressure versus temperature curve obtained 3-12 Babcock & Wilcox J MCDermott Company
from the EPFM analysis is based on a stnJctural instability analysis which is a function of both the structure's geometry and the material properties, the temperature at which this transition is made in general is not a function of material properties alone (see section 6).
For the specific case where the J-R curve is obtained from small RVSP frac-ture toughness specimens, the temperature can be determined from the mate-rial properties data alone. Because of the limited crack extension and lim-itations on the maximum J values allowed by ASTM, the J value at calculated instability will always be the maximum J value measured on the surveillance specimen.
This limitation imposed by the specimen size provides additional conservatism in the EPFM analysis since the applied J value to cause insta-bility of the structure will always be greater than the maximum J value ob-tained from the surveillance specimens.
The temperature at which the transition is made from LEFM to EPFM can be ob-tained by the procedure shown schematically in Figure 3-1.
The KJR in this figure is obtained using the procedures for converting from J to X values found in ASTM E813.
This procedure for detennining the temperature for the transition fran LEFM to EPFM will always be conservative because it is based on the KlR curve.
Since the KIR curve is based on dynamic fracture tests (both dynamic load-ing and crack arrest), it is impossible for cleavage fracture to occur at temperatures greater than those obtained using this procedure and the K g i
curve.
3-13 Babcock & Wilcox a McDermott Comparty
_ Table 3-1.
Summary of RTNDT Data and Estimated Temperatures RTwor, F Diff between No.
ave. measured of High Avg.
and estimated Materf al/ type cases nyjt s meas Est RTNDT,F j
SA 508, Class 2 low-24 60 4
60pg) 56 alloy forgings TNDT SA 533 8 low-alloy 13 40 0
40 40 plates SA 516 C carbon 20 10
-11 10 21 steel plates Submerged-arc 10 20 0
20 20 Linde 80 weld Submerged-a rc 10
-50
-66 20 86 Linde 0091 weld Manual metal arc 9
-10
-67 20 87 weld "A 508 Class 2 HAZ 6
30
-20
-26
-20 6
SA 106 C piping 11 50 5
50 45 (a)60F or the drop weight temperature, if known.
4 l
3-14 Babcock & Wilcom A MCDermott company
Table-3-2.
Summary of Cy SE Data and Estimated V
Upper Shelf Energies CvUSE, ft-lb Diff between No.
ave. measured of Low Avg.
and estimated Material / type cases meas meas Est Cy USE, ft-lb SA 508, Class 2 5
91 124 75 49 l ow-alloy forgings SA 5338 low-8 85 91 75 16 alloy plates Submerged-20 66 81 66 15 arc weld 1
f t
i i
e t
3-15 Babcock & Wilcox a McDermott company
Figure 3-1.
Relationship Between Fracture Toughness Properties and the Fracture Mechanics Evaluation Methods
!I
~~ ~
I b
Kgg 1
Li l
g l
"4 l
=
=
w giIR l
l l
Analysis and acceptance criteria per l Analysis and acceptance criteria based ASME Code,Section III, Appendix G, Ionelastic-plasticfracturemechanics as described in section 4.
as described in section 6.
I L
Temperature Legend KjR
-- K relationship as defined in ASME Code,Section III, Appendix G, fbh a specific material and adjusted for initial properties and ef-fects of neutron irradiation.
K
-- Materials elastic-plastic fracture toughness relationship as devel-dR oped from appropriate data base.
T
-- Temperature at which the linear-elastic fracture mechanics and g
elastic-plastic fracture mechanics analytical methods interface.
3-16 Babcock & Wilcox 4 MCDermott company
4 9
4.
LEFM ANALYTICAL PROCEDURES 4.1.
Basis The calculational procedures used to determine the pressure-temperature
{
limitations on the reactor coolant (RC) system are based on ASME Appendix G,
as incorporated in the Winter 1973 Addenda, and on WRC Bulletin 175.3 To determine the minimum bolt preload temperature, the calculational pro-cedure is partially based on Appendix A to ASME Section XI since it uses the static critical stress intensity factor K c rather than the reference I
critical stress intensity KIR of ASME Appendix G.
Procedures for quantitatively obtaining the maximum allowable pressure at a given temperature for Class 1 ferritic pressure-retaining components are given in ASME Appendix G and are described in more detail in WRC Bulletin 175.
The methods of calculating applied stress intensity factor are simpli-fied, and the postulated flaw is defined by a reference flaw of specified size and shape.
The procedures are not applicable to pressure boundary regions near geometric discontinuities, such as nozzles, and in such cases the technology of Bulletin 175 is applied directly.
The components of the RC system in a typical B&W power plant have been analyzed to detennine the minimum required reactor coolant temperature for pressures of 626, 2250, and 3125 psig.
The 626 psig pressure was selected because it is 1 psig above the pressure corresponding to 20% of the pre-operational system hydrostatic test pressure.
This is the maximum allow-able pressure (625 psig) for a component when the reactor coolant tempera-ture (or the volumetric average metal temperature) is bel ow the lowest service temperature of the component.
The components for which a lowest service temperature must be defined include the RC loop piping and the con-trol rod drive mechanism (the CRDM is an appurtenance to the reactor ves-sel).
The lowest service temperature of these components is 150F (based on 4-1 Bat > cock & Wilcox A MCDf fmott Company r
- - - - ~ - -
RTNOT + 100F) for the piping and 100F (as derived in section 3.3) for the CROM.
The 2250 psig pressure was selected because it is approximately the normal operating pressure; 3125 psig was selected because it is the preser-vice system hydrostatic test pressure.
The reactor vessel closure head region, the reactor vessel outlet nozzles, and the beltline region are the only portions of the RC system with a rela-tively high minimum required temperature at 626 and 2250 psig.
The reactor vessel outlet nozzle and the closure head region show the highest miriimum required temperature at 3125 psig.
These three regions are the only ones that, at different stages of the vessel 's design life, regulate the pressure-temperature limitations of the RC system for nonnal operation and inservice pressure tests.
The outlet nozzles and the closure head region regulate the minimum allowable preservice hydrostatic test temperature.
Each region has the following characteristics:
The beltline region di rectly surrounds the effective height of the fuel assemblies and is exposed to continual neutron flux throughout the service life of the reactor vessel.
The neutron fluence (flux x time) will change the mechanical properties of the beltline region materials.
This continual change necessitates periodic adjustments to the pressure-temperature operating limitation throughout the service life of the reactor vessel.
This region is remote from geometric di scontinui ties, and the applied stresses are proportional to the internal pressure and to the heatup or cooldown rates.
The closure head region of the reactor vessel is subject to significant stresses due to mechanical loads resulting from bolt preload.
In this re-gion, the applied stresses are not proportional to the internal pressure.
This region is subjected to high stresses at relatively low temperatures.
The highest stress levels occur at the head-to-head flange juncture of the closure head region.
The outlet nozzle of the reactor vessel is the largest nozzle in the RC system.
The inside corner of the nozzle is subjected to high local stress-es produced by pressure.
The local stresses can be two to three times the membrane stress of the shell.
As the radius of the nozzle increases, the magnitude of the stress intensity factor increases for a constant assumed flaw.
4-2 Babcock & Wilcox a McDermott company
For loading conditions other than the preservice systen hydrostatic test (PSHT), the nozzles and most other regions near geometric discontinuities are analyzed using the same safety margins as those required by ASiE Appen-dix G for shells and heads remote from discontinuities.
For the analysis of the head-to-head flange juncture of the closure head region, the safety factors are the same; however, the size of the postulated flaw is smaller than the referenced flaw.
The assumed flaw on the head-to-head flange juncture is a sharp surface flaw with a depth of 1/6 t and a length of t (where t is the section thickness).
The thickness of the juncture varies from 6.5 to 8 inches depending on the size of the reactor vessel.
This juncture is inspected prior to service and at several intervals throughout the service life of the power plant.
The inspection techniques can detect very small surface defects (defects with areas greater than 1 in.2 are considered detectable).
For the wall thickness of 6.5 inches, the area of the postula ted flaw ( semi-el liptical )
is 8.5 in.2 The area of the postulated flaw is 8.5 times la rger than the minimum detectable defect area.
For the PSHT all geometries are analyzed using a margin of safety of 1.0 on the stress intensity factor and postulated flaws that are smaller than the reference flaw of ASiE Appendix G.
Smaller postulated flaws are justifi-able since this test is perfomed before initial operation.
The postulated flaws employed to de temine the pressure-temperature limit curve fo r the PSHT are oescribed in section 4.2.2.2.
Additionally a pressure exceeding 2/3 of the test pressure is not allowed until the component temperature ex-ceeds RTNDT + 60.
The reference flaw of ASME Appendix G is a sharp surface flaw perpendicular to the direction of maximum stress, having a depth of 1/4 t and length of 1-1/2 t (for section thicknesses of 4 to 12 inches).
ASiE Section III also requires that the test coupons be at least 1/4 t frcn any surface unless the material is a very thick forging and the test location is very near the surface (0.75 inch from a heat-treated surface).
Since for most geometries the depth of the postulated flaw and the test location is 1/4 t ( from either surface), the analytical calculations used on all geometries depend on the metal temperature and impact properties (including ef fects of irradi-l ation) at 1/4 t and 3/4 t.
The impact properties of thick and complex i
4-3 Babcock & Wilcon l
a M(Jermott Company l
forgings at 1/4 and 3/4 t are assumed to be equal to the properties deter-mined near the surface.
The metal temperature and impact properties for the head-to-head flange juncture are taken at 5/6 t.
For the analysis of the head-to-head flange juncture, the impact properties at 5/6 t are as-sumed to be equal to those detennined by the ASME Code.
At the beginning of service life, the closure head region and the outlet nozzles control the pressure-temperature limitations of the loading condi-tions of interest.
Af ter several years of neutron irradition exposure, the RTNDT of the beltline region materials will be high enough for the beltline region to regulate parts of the pressure-temperature limit curves.
The maximum allowable pressure as a function of fluid temperature for the ser-vice period of the limit curves is obtained through a point-by-point com-parison of the limits imposed by the closure head region, outlet nozzle, and beltline region.
The maximum allowable pressure is the lower of the three calculated pressures.
For additional years' operation, the adjusted RTNDT of the beltline region materials will continue to increase; there-fore, periodic adjustments on the pressurization limit curves are required throughout the service life of the RC system.
Since every surveillance cap-sule withdrawal will produce pertinent irradiated beltline region material impact data, adjustment of the pressurization limit curves may be required after each capsule withdrawal.
The initial and subsequent adjusted pres-sure-temperature limits include the predicted radiation-induced RTNDT (detennined as described in section 3.1.3.1) for the period until the next capsule withdrawal.
Af ter each capsule withdrawal, the RTNDTs of the beltline region materials are predicted by adding the unirradiated values to the predicted radi a-tion-induced 1RTNDTs and then confirmed by the material surveillance pro-gram test results.
Both the predicted 1RTNDT and tne data obtained from the surveillance program are used to define the adjusted RTNDT that will be used to recalculate the pressurization limit curves.
4.2.
Description The methods used to obtain the pressure-temperature limitations for each of the loading conditions of interest are described in this section.
Table f
4-1 summarizes the analytical assumptions.
)
4-4 Babcock & Wilcox
& MtDttmott Company
4.2.1.
Normial Operation 4.2.1.1.
Bolt Preloading To. define the minimum preload temperature, it is necessary to analyze the bolt preloading conditions.
The minimum preload temperature can be the low-est temperature at which the bolting materials meet the toughness require-ments of the ASME Code or the calculated minimum temnerature required for protection against nonductile failure of the closure head region, whichever temperature is higher.
Section 3.2 of this report shows that at 40F the bolting materials meet the requirements of ASME Code; now it is necessary to calculate the minimum allowable temperature of the closure head region to determine whether it is higher than 40F.
During bolt preloading, the maximum tensile stresses occur at the outside surface (5/6 t) of the head-to-head flange juncture of the closure head region.
The stresses are primarily bolt preload bending stresses.
The pressure stresses are very small since the maximum allowable pressure for this loading condition is relatively low h450 psig).
The minimum tempera-ture required for protection against nonductile failure is first calculated at 0 psig and then at 626 psig.
Both pressures are analyzed because higher temperatures may be required at 0 than at 626 psig.
For both cases, the thermal stresses are nil since the coolant temperature is essentially at steady state throughout this loading condition.
The method used to calculate the minimum preload temperature is as follows:
1.
The membrane and bending stresses at the 5/6 t vessel wall location that result from bolt preload and internal pressure are calculated by the stress analysis of the head-to-head flange juncture at both 0 and 625 psig.
2.
Using the membrane and bending stresses calculated in step 1,
the stress intensity factor for both cases is calculated by the following equation:
M KI = 1.1 o Mm K q + ob B q Babcock & Wilcox 4-5
& MtOttmott (Ompafly
where the assumed flaw of a = 1/6 t; G + 0.64 O KI = 0.82 m b
Q Q
where KI = stress intensi ty factor based on reference flaw at S/6 t vessel wall location, M,MB = correction factors for membrane and bending load conditions, K
respectively (values from WRC Bulletin 175, Figures A3-1 and A3-2); for a 1:6 crack depth:
thickness ratio the values are 1.03 and 0.88, respectively; Q = flaw shape factor modified for plastic zone size (reference 3 gives basic expression) a = assumed crack depth, t = section thickness, gn = calculated mentrane stress, cb = calculated bending stress.
3.
The relative temperature T-RTNOT at which the critical static stress i ntensi ty factor K e equals the highest calculated stress intensi ty I
factor KI (from step 2) is calculated using Figure 4-1, which is based on Figure A-4200-1 from ASME XI, Appendix A.
4.
Using the relative temperature calculated in step 3 and the highest RTNOT of the closure head region materials, we can calculate the mini-mum temperature required for protection against nonductile failure.
S.
The mini >num preload tempe rature is the one calculated in step 4 or 40F, whichever is higher.
6.
Appendix G of the ASME Code Section III recommends the temperature of the closure area be RTNOT at bolt-up.
The forgoing procedure yields similar results with the exception of lod stressed closures.
This pr ecedure is considered consistent with the phil osophy of the ASiE Code and will be used for establishing temperature requirements.
4.2.1.2.
Heatup The heatup transient starts at the minimum preload tempe rature.
For tem-peratures above minimum preload, the heatup pressure-temperature limi t 4-6 Babcock & Wilcos a MCOf tmott Company
curve is calculated by a point-by-point comparison of the limits incosed by the closure head region, the c""
.ozzles, and the beltline region.
The heatup limit curve is the ccnposi te or lower bound curve of the limits im-posed by the three controlling regiuns.
The limits imposed by the closure head region are established by assuming a 1/6 t x t surface flaw louted at the outside surface of the head-to-head flange juncture.
During heatup all the stresses, including the bolt pre-load and the thermal stresses, are in tension at the outside surface of the closure head region.
The S/6 t location corresponds to the depth of the assumed flaw on the outside surface of the head-to-head flange juncture.
The minimum required fluid temperatures are calculated at several coolant pressures above 625 psig.
This is done by first calculating the fluid temperature as a function of metal temperatures for each beatup rate of interest and then calculati ng the minimum required metal temperature at each pressure.
For fluid temperatures between the minimum preload tem-perature and the minimum required fluid temperature at 626 psig, the maximum allowable pressure is 625 psig.
The limits imposed by the outlet nozzles are calculated by assuming a flaw at the inside corner of the nozzle.
Tho depth of the assumed flaw is 3 inches, which is the depth of toe reference flaw of ASME Appendix G for a section thicker than 12 i nch e s.
During heatup, the inside corners of nozzles are subjected to high local stresses produced by pressure; howeve*,
the themal stresses are in compression.
The limit curve is calculated by determining the metal temperature at the inside corner 1/4 t of the outlet nozzle as a function of fluid tempe rature.
The critical stress intensity f actor is indexed to the fluid temperature using the highest RTNDT of the two outlet nozzles.
The maximum allowable pressure is then calculated as a function of fluid tempe ra ture.
The themal stress intensity factors for these calculations are assumed to be zero.
This assumption is conservative since during heatup, the contributing themal stress intensity factor at the inside corner of the nozzle is negative.
The pressure-temperature limits imposed by the beltline region are calcu-i lated using the postulated reference fl aw of 1/4 t depth.
The reference i
i Bat > cock & WHcom 4-7
,ucoymoi, comp,ny l
flaw is assumed to be located at both the inside and outside surfaces of the beltline region.
During heatup, the themal stresses are in compres-sion at 1/4 t of the section thickness of the beltline region and are in tension at 3/4 t.
The 1/4 t location corresponds to the depth of the reference flaw on the inside surface of the reactor vessel wall.
The 3/4 t location corresponds to the depth of the reference flaw on the outside sur-face of the reactor vessel wall.
The metal temperatures at the 1/4 and 3/4 t lag the fluid temperature during the nomal heatup conditions.
Since the neutron fluence a *.tenua tes through the thickness of the beltline region material, the RTNDT at the 1/4 t location will be higher than that at the 3/4 t location.
Br.cause of these variables, two sets of calculations aust be performed to obtain the pressure-temperature limitations imposed by the bel tline region.
First, the pressurization limit for the steady-state condition is calcu-lated as a function of fluid temperature.
For this calculation the metal and fluid temperatures are the same and the impact properties used are those of the 1/4 t location.
There are no themal stresses in this case, and the only contributing stress intensity factors are those produced by pressure.
Second, the curve of pressure versus fluid temperature limit is calculated for each heatup ramp of interest assuming that the reference flaw is 10-cated at the outside surface of the beltline region wall.
For this calcu-lation, it is necessary to determine the metal temperature at 3/4 t as a function of fluid temperature and the stress intensity factor produced by the themal stresses.
The themal stress intensity facter is added to the pressure stress intensity factor.
The impact properties used in this cal-culation are those of the 3/4 t vessel w311 location.
The methods employed to obtain the limits imposed by the closure head re-gion, outlet nozzle, and beltline region and the pressure-temperature limit curve of the RCPB for normal heatup are described below.
Closure Head P.egion Heatup Limits The heatup limits imposed by the closure head ragion are calculated as follows:
4-8 a McDermott company
.-m-
1.
For each of the heatup ramps of interest, the metal temperature at 5/6 t of the head-to-head flange juncture is calculated as a function of fluid temperature.
2.
The minimum allowable fluid temperatures of the closure head region for coolant pressures of 626, 1250, and 2250 psig are calculated as follows:
a.
The membrane and bending stresses at the 5/6 t location resulting from bolt preload, thennal gradient, and internal pressure are cal culated ty a de tail ed stress analysis of the head-to-head flange juncture.
b.
Using the membrane and bending stresses calculated in itep a, the stress intensity factor is calculated by the following equation:
O KI=2 1.1( mb + @ )Mg M
M
+
bb B
+
bT B Q
T.
/T where the assumed flaw of a = 1/6 t; K1 = 1. 64 (omb + 3#) T T
W
+ 1.28 bb
+ 0.64 b T ---
where o mb i SP calculated membrane stresses due to bolt preload
=
and pressure, ebb >CbT = calculated bending stresses due to bolt prel oad and the rmal gradient.
(See section 4.2.1.1 for definition of other factors.)
c.
For each pressure, the minimum relative temperature is that at which the cal cula ted stress intensi ty factor (K )
equals the I
reference stress intensity factor (KIR) of Figure G-2110.1 of ASiE Appedix G.
d.
The minimum required metal temperatures are calculated using the minimum relative temperatures (calculated in step c) and the high-est RTNOT of the closure head region materials.
L Babcock &Wilcos 4-9 a uconmoa company
e.
The minimum all owable fluid tempe rature fo r the three coolant pressures are calculated using the minimum required metal tem-peratures calculated in step d and the fl uid-me tal temperature relationship of step 1.
3.
The pressure-tempe rature limits imposed by the closure head region during nomal heatup are defined as follows:
a.
For fluid temperatures between the minimum prel oad temperature calculated in section 4.2.1.1 and the minimum allowable fluid tem-perature calculated in step 2 for 626 psig, the maximum allowable coolant pressure is 625 psig.
b.
For pressures of 1250 and 2250 psig, the minimum allowable fluid temperatures are those calculated in step 2.
For coolant pres-sures between 626,1250, and 2250 psig, the minimum fluid tempera-tures are defined by linear interpolation.
c.
10CFR50 Appendix G Paragraph !Y.A.2 requires the highly stressed regions of the closure region to be at a temperature of at least RTNDT
+
1200F fo r pressures above 625 psig.
The forgoing procedure results in a
similar temperature requirement.
The required temperature is lower than 1200F if slow heat-up rates are speci fied and higher than 1200F for the operating pressure condition and maximum heat-up rates.
The forgoing procedure is considered to be consistent wi th the requi rements of 10CFR50 Appendix G and is used in lieu of the stated requirement.
Outlet Nozzle Heatup Limits The heatup limits imposed by the outlet nozzles are calculated as follows:
1.
For each of tha heatup ramps of interest, the metal temperature at a depth of 3 inches (at the inside corner) location of the outlet nozzle is calculated as a function of fluid temperature.
The thermal analy-sis calcalations are perfo med using a one-dimensional transient distribution program.
2.
The KIR curve of ASiE Appendix G is indexed to the highest RTNDT Of the two outlet nozzles.
Using the fluid-me tal temperature relation-ship cal culated in step 1,
the critical stress intensi ty factor is calculated as a function of fluid temperature, KIR(T )3a, f
Babcock & Wilcox 4-10 a uconmore company
3.
The pressure-temperature limit curve imposed by the outlet nozzles during heatup is calculated using the following equation:
KIR(T )3" f
p(T ) =
f r2 + r2 i
o gg 2F(a/r l 2 - ry n
r where KIR(T )3a = critical stress intensity factor ar. function of f
fluid temperature calculated in step 2, F(a/r ) = obtained from WRC Bulletin 175, Figure A5-1, n
en = apparent nozzle radius, p(T ) = coolant pressure as function of fluid temperature, f
r,rj = outside and inside radius of reactor vessel nozzle o
- belt, a = flaw depth, assumed to be 3 inches.
Beltline Region Heatup Limits The limits imposed by the beltline region are calculated as follows:
1.
For each heatup ramp of interest, the metal temperatures at the 3/4 t vessel (beltline region) wall locat'on are calculated as a function of fluid temperature (T ).
The themal analysis calculations are per-f fomed by a one-dimensional transient distribution program.
2.
Also, as part of the themal analysis of step 1 in the preceding para-graph, the temperature distribution through the vessel (beltline re-gion) wall is calculated as a function of fluid temperature for each heatup ramp.
3.
The KIR curve of AS4E Appendix G (Figure G-2110.1) is indexed to the highest postulated RTNOT of the 1/4 t wall location and to the highest postulated RTNDT of 3/4 t.
For each heatup ramp of interest the criti-cal stress intensity factor for the 3/4 t vessel wall location is cale-ulated as a function of fluid temperature using the data of step 1 and the RTNOT at 3/4 t.
Al so, for the steady-state condition, KIR 15 plotted as a function of fluid temperature using the RTNDT at 1/4 t.
4.
The K1 produced by the themal gradient across the vessel wall is calc-ulated as follows:
Embcock & WHcom 4*11 A McDermott company
a.
Utilizing the tempe rature distribution obtained in step 2 the equivalent linear bending stress is calculated due to the radial gradient.
This is done by either integrating the thermal distri-bution or stress distribution across the wall, b.
The KIT 'Mb x Sth whera Mb equals 2/3 Mm as defined -in ASliE Appendix G and 5th is the equivalent linear thermal bending stress.
5.
The pressurization limit for a steady-state condition is calculated as a function of fluid temperature by the following equation:
P(T )ss
- Kt"(T )1/4 t f
r2 + r2 2Mm r2 - r2 o
i X g(T )1/4 t = critical stress intensity factor for steadystate where I
f condition as a function of fluid temperature, based on RTNDT at 1/4 t, calculated in step 4; Mm = obtained from ASME Appendix G, Figure G-2214.1; a stress ratio > actual is used (Checks are made to confim that ~fhe proper Mm value is used.);
inside and outside radii of reactor vessel belt-rj,ro = line region, P(T )ss = allowed steady-state pressure as a function of f
fluid temperature.
6.
The pressure versus fluid temperature data for each heatup ramp of interest are calculated as follows:
P(T ) = Kin (Tr)3/4 t - KTT(TS) f r2 + r2 2Mm r2 - r2 o
i K g(T )3/4 t = critical stress intensi ty factor based on 3/4 t where I
f i
RTNOT, a function of fluid tempe rature cal cu-l lated in step 4, IT(T ) = KI produced by themal gradient across the ves-l K
f sel wall as a function of fluid temperature (cal-culated in step 5),
Babcock &Wilcon 4-12 a McOttmott company
P(T ) = allowable pressure as a function of fluid tem-f
- perature, and the other factors are as defined above.
7.
The pressure-temperature limits imposed by the beltline region during nomal heatup are obtained by a point-by-point conparison of the data obtained in steps 5 and 6.
The maximum allowable pressure is taken to be the lower of the two values.
Reactor Coolant Pressure Boundary Heatup Limits The pressure-temperature limits during nonnal heatup of the RCpB are ob-tained through a point-by-point conparison of the limits impcsed by the closure head region, ou tlet nozzles, and bel tline region.
The maximum allowable pressure at any given fluid temperature is taken to be the lower of the th.ee calculated pressures.
4.2.1.3.
Cooldown The method used to obtain the cooldown pressurization limit curve for the RCPB is very similar to that used for the heatup curve.
From the nomal operating temperature to the minimum bolt preload temperature, the cooldown pressure-temperature limit curve is calculated through a point-by-point comparison of the limits imposed by the closure head regi on, the ou tl e t nozzles, and the beltline region.
The cooldown liinit curve is the lower bound curve of the limits imposed by the three controlling regions.
The cooldown limits of the closure head region are established, as for heatup, by assuming a 1/6 t x t surface flaw located at the outside surface of the head-to-head flange juncture.
Although the inside surface is sub-jected to positive thermal stresses during cooldown, the to tal stress is higher at the outside than at the inside surface.
This ir due to the high bolt prel oad bending stresses on the outside surface.
The cooldown and heatup limits of the closure head region are calculated very similarly.
The only differences are that (1) the fluid and metal temperatures are as-sumed to be equal (steady-state), and (2) the themal stresses at the out-side surfaces are assumed to be zero.
The steady-state assumption is con-servative since the metal temperature, especially at 5/6 t, is higher than the fluid temperature during cooldown.
The assumption that the themal stresses are zero is also conservative since the thermal stresses at the outside wall of the closure head region are negative during cooldown.
Babcock & Wilcom 4-13 a McDermott company
The cooldown limits imposed by the outlet nozzles are calculated, as for heatup, assuming a 3-inch-deep flaw at the inside corner of the nozzle.
During cooldown, the inside corners of the nozzles are subjected to high local stresses produced by the pressure and temperature gradient.
To calc-ulate the limit curve, the metal temperature 3 inches from the inside cor-ner locations is calculated as a function of fluid temperature.
When calc-ulating the maximum allowable pressure, the contributing themal stress intensity factor is assumed to be equal to that calculated for the nozzle bel t vessel wal l.
This assumption is conservative because the themal stress intensity factor for a nozzle corner flaw is also lower than that for a surface flaw on the nozzle vessel wall owing to the lower postulated crack penetration (crack depth over section thickness) on the nozzle corner.
The method used to calculate the cooldown pressure limit curve imposed by the beltline region is also similar to that used for the heatup limit curve; however some differences exist.
During cooldown, the themal stress-es are in tension at 1/4 t and in compression at 3/4 t.
Because the ther-mal stresses are in tension at 1/4 t, and the RTNDT at 1/4 t will be higher than that at 3/4 t af ter exposure to neutron irradiation, only the metal temperature and the impact properties of the 1/4 t location are used to ob-tain the cooldown limit curve.
However, three calculational steps are re-quired to obtain the cooldown limit curve of the beltline region:
1.
The pressure limit curve for a steady-state condition is calculated as a function of fluid tempe ratur e.
The assumed steady-state condition makes the fluid and metal temperatures equal.
The impact properties are those of the 1/4 t location.
The contributing thennal stress in-tensi ty factor is zero.
This step is required because the metal temperature may not be higher than that of the fluid during an upset cooldown condition as it is during nomal cooldown.
2.
The pressure limit curve is calculated for each cooldown ramp of in-terest assuming that the reference flaw is located at the inside sur-face of the beltline region wall.
For this calculation, the metal temperature at 1/4 t is determined as a function of fluid temperature, and the themal stress intensity factor is added to the stress inten-sity factor produced by pressure.
The impact properties at 1/4 t are used in this calculation.
Babcock & Wilcox 4-14 J M(Derrnott company
e 3.
A point-by-point comparison of the data obtained in the first two steps will obtain the lowest pressure at any temperature of the two data sets.
The calculated lowest pressure becones the maximum pres-sure at any temperature for the reactor vessel beltline region.
The methods used to obtain the limits imposed by the closure head region, outlet nozzle, and beltline region and the pressure-temperature limit curve for the RCPB for nonnal cooldown are described below.
Closure Head Region Cooldown Limits The cooldown limits imposed by the closure head region are calculated as follows:
1.
For each cooldown ramp of interest, the metal temperature at 5/6 t of the head-to-head flange juncture is assumed to be equal to the fluid tempe rature.
2.
The minimum allowable fluid temperatures of the closure head region for pressures of 626,1250, and 2250 psig are calculated as follows:
a.
The membrane and bendi ng stresses at 5/6 t resulting from bolt preload and internal pressure are calculated by a detailed stress analysis of the head-to-head flange juncture, b.
Using the membrane and bending stresses calculated in step a, the stress intensi ty factor is calculated using the foll owi ng equa-tion:
E+
"bb ad KI=2 1.1F mb + mp)MK M
M W-where the assumed flaw of a = 1/6 t; KI = 1.64( mb + mp)
+ 1.28 c b b
Q, a, and t are defined in section 4.2.1.1 and where Kg, Mg, Mg", Closure Head Region Heatup Limits," step 2b.
other factors in I
i Babcock & WHcom 4-15
,,o,,,,, nom,,n, t
c.
For each pressure the minimum relative temperature is that at which the calculated stress intensity factor KI equals the refer-l ence stress intensi ty factor (KIR of ASiE Appendix 3,
Figure G-2110.1).
d.
The minimum required fluid temperatures are calculated using the minimum relative temperatures (calculated in step c) and the high-est RTNOT of the closure head region materials.
3.
The pressure-temperature limits imposed by the closure head region dur-ing nomal cooldown are defined as follows:
a.
For fluid temperatures between the minimum preload temperature cal cula ted in section 4.2.1.1 and the minimum allowable fluid temperature calculated in step 2 for 626 psig, the maximum allow-able pressure is 625 psig.
b.
The minimum allowable fluid temperatures for pressures of 1250 ar.;d 2250 are those calculated in step 2.
For coolant prissures be-tween 625,1250, and 2250 psig, the minimum fluid temperatures aee defined by linear interpolation.
Outlet Nozzle Cooldown Limits The cooldown limits imposed by the outl et nozzles are cal cul atet as follows:
1.
For each cooldown ramp of interest, the metal tempe rature at the 3-inch depth (from sne inside corner) location of the outlet nozzle is cal culated as a function of fluid tempe rature.
The thermal analysis is perfomed using a one-dimensional transient distribution program.
2.
As part of the themal analysis in step 1, the temperature difference through the nozr.le belt vessel wall is calculated as a function of fluid temperature.
l l
3.
The KIR curve of AS4E Appendix G is indexed to the highest RTNDT Of i
the two ou tlet nozzles.
Using the data calculated in step 1,
the l
critical stress intensity factor at the inside corner of the nozzle is l
IR(T )3a.
calculated as a function of fluid temperature, K f
l l
Babcock & WHcom l
4-16 a ucoumore company l
I
1 4.
The K1 produced by the thennal gradient ac ross the outlet nozzle corner is calculated by same method as step 4 in heatup procedure.
5.
The pressure-temperature limit curve imposed by the outlet nozzles dur-ing cooldown is calculated using the following equation:
P(T ) = Kin (Tcha - KtT(TS) f d + r2
- m 2F(a/r )
~N n
- ere KIR(T )3a = critical stress intensity factor as a function of i
f l
fluid temperature calculated in step 3, KIT (T } = thennal stress intensity factor as a function of f
fluid temperature calculated in step 4, and all other factors are defined in "Outlet Nozzle Heatup Limits,"
step 3.
Beltline Region Cooldown Limits The limits imposed by the beltline region during cooldown are calculated as follows:
1.
For each cooldown ramp of interest, the temperature at 1/4 t (beltline region) is calculated as a function of fluid temperature (T ).
f 2.
As part of the thennal analysis of step 1, the temperature difference through the vessel (beltline region) wall is also calculated as a func-l tion of fluid temperature for each cooldown ramp.
3.
The nest limiting adjusted RTNDT at 1/4 t is also used in the cooldown 4
analysis.
1 i
4.
The KIR curve of AS4E Appendix G is indexed to the adjusted RTNDT of l
step 3.
For each cooldown ramp of interest, KIR is plotted as a func-l tion of fluid temperature using the data from step 1.
For the steady-state condition, KIR is also plotted as a function of fluid tempera-ture using the same adjusted RTNDT.
S.
The XI produced by the thennal oradient across the vessel wall during cooldown is calculated as describe. in step 4 of the heatup procedure.
However, the AT values are those calculated in step 2 for each cool-down ramp.
I Babcock &WHeos 4-17
. ucoumoncom,,ny
6.
The pressurization limit for steady-state condition is calculated as a function of fluid temperature, as described in "Beltline Region Heatup Limits," step 5.
7.
The pressure-versus-fluid temperature data for each cooldown ramp are calculated as follows:
P(T ) = K a(Tr)1/4 t - KtT(Tf) f r2 + r2 2Mm r2-rj K g(T )1/4 t = Kg based on RTNOT, also a function of fluid tem-where I
f perature (see step 4),
IT(T ) = themal stress intensity factor as a function of K
f fluid temperature (see step 5).
and the other factors are as defined in "Beltline Region Heatup Lim-its," steps 6 and 7.
8.
The pressut e-temperature limits imposed by the beltline region during nomal cooldown are obtained through a point-by-point comparison of the data obtained in steps 6 and 7; the maximum allowable pressure is taken to be the lower of the two values.
Reactor Coolant Pressure Boundary Cooldown Limits The pressure-tuperature limits during normal cooldown of the RCPB are ob-tained through a point-by-point conparison of the limits imposed by the closure head region ("Closure Head Region Cooldown Limits," step 3), the outlet nozzles ("Outlet Nozzle Cooldown Limits," step 5), and the beltline region ("Beltline Region Cooldown Limits," step 8).
The maximum allowable presure at any given fluid temperature is taken to be the lowest of the three calculated pressures.
4.2.2.
Preservice System Hydrostatic Test (PSHT) 4.2.2.1.
Bolt Preloading The minimum preload temperature for the PSHT is calculated by following the basic methods employed for nomal operation (section 4.2.1.1).
For the PSHT the minimum preload temperature is calculated using a postulated sur-face flaw 1/8 t deep and 3/4 t long (1/8 t x 3/4 t) located in the outside 4-18
- ** U E ***
4 MCDermott Comparty
i surface of the head-to-head flange juncture.
This assumed flaw is smaller than that assumed during nonnal operation (1/6 t x t).
The smaller flaw is 1
conservative since the PSHT is perfomed after the nondestructive testing required by ASME Section III, and the system has not been subjected to cyclic loading.
For the smaller postulated flaw, the equation used to calculate the stress intensity factor (step 2) takes the following form:
Kt = 0.70 c m
+ 0.57 ob 4
4 where KI is the stress intensity factor based on a 1/8 t x 3/4 t flaw, and all other factors are as Jefined in section 4.2.1.1.
The values of o and % are calculated as described for nomal operation m
(step 1) for the higher specified preload.
All other steps of the pro-cedure for calculating minimum preload teroperature for normal operation are followed when calculating the minimum preload temperature for PSHT.
4.2.2.2.
Heatup and Cooldown As described in section 2.4, the PSHT pressure is nomally reached when the metal temperature of the controlling pressure boundary is at steady state, and it is higher than the calculated minimum test temperature.
At tempera-tures lower than this ntnimum, the maximum allowable pressure is only 625 l
psig.
However, for some plants, it may be necessary to gradually increase the maximum allowable pressure as the metal temperature increases, just as for normal heatup and cooldown.
For these plants the themally induced stresses are considered when calculating the pressure-temperature limit curve.
The methods for calculating the PSHT limit curve are similar to those for nomal operation except for the following deviations:
1.
The analysis is only perfomed for the two regions of the RC5 that potentially control the PSHT pressure-temperature limits:
the closure head region and the outlet nozzle.
The beltline region does not con-trol these limits since the materials have not been affected by irrad-1ation.
l l
l 4-19 Bat > cock & WHeos a McDermott company l
t_
2.
When calculating the limits imposed by the closure head region, the postulated flaw is a 1/8 t x 3/4 t semi-elliptical surface flaw in the outside surface.
The applied factor of safety in the stress intensity f actor is 1.0, and the minimum allowable temperature is also calcu-lated for 3125 psig.
The postulated flaw is the same as that assumed I
when calculating the minimum preload temperature for the PSHT (section 4.2.2.1).
3.
When calculating the limits imposed by the outlet nozzles, the postu-lated flaw is a surface flaw 1.0 inch deep located at the inside corner, and the factor of safety applied on the stress intensity fac-tor due to pressure is 1.0.
The justification for the smaller postu-lated flaw (1.0 rather than 3.0 inches deep) is again the nondestruc-tive examination prior to PSHT and the impossibility of fatigue crack growth.
4.
The pressure-temperature limits are calculated for both heatup and cooldown; however, for simplicity, the most limiting curve is used to define these limits from initiation to completion of the PSHT.
The PSHT limit curve for the RCPB is the composite or lower bound of the limits imposed by the two controlling regions during both heatup and cooldown.
4.2.3.
Inservice System Leak and Hydrostatic Tests (ISLHT) 4.2.3.1.
Bolt Preloading The minimum prel oad temperature for the ISLHT is the same as that for nomal operation since the same load is specified.
4.2.3.2.
Heatup and Cooldown Since the ISLHT can be perfomed throughout the service life of the power plant, the effects of irradiation are considered when establishing the pres-l sure-temperature limit curve for each test.
As for nomal heatup and cool-down, the closure head region, the outlet nozzles, and the beltline region are the only regions of the reactor vessel that control the pressurization l
limits of the RC system during ISLHT.
The nomal means of heating or cool-ing the sys tem, before or af ter reaching the desired pressure for each test, are those used during nomal heatup and cooldown.
Consequently, the 4-20 Babcock & WHeos a McDermott company
methods used to obtain the pressure limit curves of these loading condi-tions are similar to those used for normal heatup and cooldown.
As for the PSitT, the ISLHT pressure-temperature limits are calculated for both heatup and cooldown; however, for simplicity, the most limiting curve is used to define the pressure-temperature limits from initiation to completion of the ISLHT.
Another deviation from the methods employed for normal heatup and cooldown is the magnitude of the applied factor of safety.
The-factor of safety applied to calculate the stress intensity factor and the allowable pressure in the preceding procedures is 1.5 rather than 2.0.
The ISLHT pressure-temperature limit curve is the composite or lower bound curve of the limits calculated for heatup and cooldown.
The requirement of 10CFR50 Appendix G speci fyi ng a temperature of RTNDT + 900F for highly stressed regions of the closure for pressures above 625 psig is essentially met by this procedure.
As fo r the nonnal heat-up case higher or lower temperatures may be requied depending on heat up rate.
4.2.4.
Reactor Core Operation Except fo r l ow-powe r physics tests, the pressure-temperature limits for reactor core operation are as follows:
1.
The fluid temperature must be equal to or higher than the minimun re-quired for the ISLHT as calculated by the method described in section 4.2.3.
2.
In addi tion, the fluid temperature must be at least 40F higher than the ninimum pressure-temperature limit curve for both nonnal heatup and cooldown as calculated by the methods described in section 4.2.1.
3.
The fluid temperature must be at least 525F.
These pressure-temperature limits for reactor core operation are in accor-dance with Appendix G to 10 CFR 50.
{
i 4-21 Babcock &WHeen I
a MCDermott Company 1
=.
I s
Table 4-1.
Outline of Methods APPI safety factoria) y,t.1 property Flaw Temperature Loading Re fon E*
K=
K relationship RT (b) g condition analyzed Loc *n
_ Depth NOT I
Steady-state 1/4 t KIC Nonnal bolt Closure head 00 1/6 t 1
1 preload T (T )
1/4 t KIR Nonnal Closure head 00 1/6 t 2
-1 1
f heatup Outlet nozzle ID 3 in.
2 Ty(Tm) 3/4 in.
KIR Steady-state 1/4 t KIR Beltline ID 1/4 t 2
T (T.)
3/4 t KIR OD 1/4 t 2
1 f
Steady-state 1/4 t XIR Nonnal Closure head 00 1/6 t 2
1 p
cooldown T (T )
3/4 in.
KIR g
Outlet nozzle ID 3 in.
2 1
f Beltline ID 1/4 t 2
Steady-state 1/4 t KIR ID 1/4 t 2
1 Ty(T.)
1/4 t KIR Preservice Same as nonnal bolt preload, heatup and cooldown; however the depths of the postulated Sit test flaws are 1/8 t and 1.0 inch for the closure head region and outlet nozzle, respec-bolt preload, tively, the applied safety factor is always 1.0, and the beltline region is not con-heatup and sidered. The limit curve is the composite of the limits imposed by the two controlling cooldown regions during both heatup and cooldown.
Inservice Same as onnal bolt preload, heatup and cooldown, however, the applied safety factor is SLH test 1.5 rathe:/ than 2.0.
The limit curve is the composite of the limits imposed by the y
bolt preload, three controlling regions during both heatup and cooldown.
p heatup and cooldown g
5O f5 (a)K5 = stress intensity factor resulting from primary stresses, K" = stress intensity factor resulting a{
from secondary stresses.
Is (b) Location of the RTNDT used in the calculation.
4 9
e e
e
Figure 4-1.
Reference (Static) Critical Stress Intensity Factor Vs Temperature Relative to RTNDT (T-RT )T)
M 200 2>
aC-3 ISO y
KIC 5
E 120 T
C C
~
u A
u" 00 r
G O
40 m
N?
E 9()
gn oW
~
5D 2
0 t
t i
e 3%
-100
-50 0
+50
+100
+150
?=
J$
Temperature Relative to RTNOT(T-RTNOT), F M
l 5.
TYPICAL PRESSURE-TEMPERATURE LIMITS 5.1.
Composite Limit Curves The methods described in sections 3 and 4 have been applied to a typical 177 FA type plant to illustrate the development of the composite limit curves.
The methods were applied for each of the loading conditions of in-terest.
The analysis for normal heatup and cooldown was performed for the service periods ending at 5 and 32 effective full-power years (EFPY).
The analysis for the inservice leak and hydrostatic tests (ISLHT) was performed for the service period ending at 5 EFPY.
For consistency, the analysis was perfonned using a 100F/ hour temperature ramp.
For some transients the as-sumed 100F/ hour ramp is not practical.
The actual pressure-temperature limit curves for B&W plants may be different from those presented in this section because of the different maximum allowable temperature ramp rates and variations in tRTNDT.
The figures included here are for illustration only.
The analysis used the unirradiated impact properties, residual el ements,
predicted neutron fluence, and predicted radiation-induced tRTNDT for the beltline region materials typical of an 177 FA-type plant.
The unirrad-lated RTNDTs of the closure head region materials and outlet nozzles are
[
al so those of a typical plant.
The unirradiated impact properties and residual elements of the beltline region materials are listed in Table 5-1.
The predicted neutron fluence values at the 1/4 t and 3/4 t beltline region locations for 5 and 32 EFPY and the corresponding aRTNDTs and adjusted RTNDT are listed in Table 5-2 for each of the beltline region materials.
Figures 5-1 and 5-2 illustrate the development of the conposite pressure-temperature limit curves for a 100F/h normal heatup.
The figures are applicable for the service periods ending at 5 and 32 EFPY, respectively.
In addition to the composite limit curve, both figures show the limit curves imposed by the outlet nozzles, closure head region, and beltline region based on steady state and by the beltline region based on a finite 5-1 Babcock 8WIfcom 4 MCOttmott (0mpany
n.
heatup rate.
As shown in Figure 5-1, the composite limit curve for 5 EFPY is the lower bound curve of the limits imposed by the outlet nozzles, clo-sure head region, and beltline region based on a finite heatup rate.
The composite limit curve for 32 EFPY is controlled by the limits set by the beltline region based on steady-state and finite heatup and the closure head region.
At 5 EFPY, the limits set by the closure head region largely control the composite limit curve, and at 32 EFPY the same region only con-trols a small portion.
This is because the limit curves set by the closure head region and outlet nozzles do not change throughout the service life of the power plant.
Also, note that the limits set by the beltline region based on steady state do not control the composite limit curve for 5 EFPY, but they largely control the composite limit curve for 32 EFPY.
This is due to the large difference in RTNDT between 1/4 t and 3/4 t.
Both figures illustrate the crossover of the limit curves imposed by the several regions and the need for composite limit curves.
Figures 5-3 and 5-4 are very similar to 5-1 and 5-2; however, they are for nomal cooldown.
Note that the beltline region steady-state limit curves for normal cooldown control the composite limit curves for 5 and 32 EFPY.
At the high fluid temperatures (Tf > 205F) the normal cooldown composite limit curve for 5 EFPY (Figure 5-3) is less restrictive than the curve for normal heatup (Figure 5-1).
This is primarily due to the large difference between the fluid temperature and the closure head region wall metal tem-perature at 5/6 t that occurs during heatup.
However, at the lower fluid temperature (Tf< 124F), Figure 5-3 is more restrictive than Figure 5-1 because of the contributing thermal stresses at the inside corner of the outlet nozzles.
The presence of the themal stresses reduces the maximum allowable pressure.
Again, Figures 5-3 and 5-4 illustrate the need for the composite limit curves.
Figures 5-5 and 5-6 present the limits imposed by the several regions of the reactor vessel and the composite limit curves for the PSHT and ISLHT.
The allowable pressure-temperature combinations of these figures differ be-cause of the different sizes of the postulated flaws, applied margins of sa f e ty, and 3ssumed RTNDTs.
For both tests the limit imposed by the clo-sure head region during heatup control the composite limit curves.
For the 5-2 Babcock & Wilcos 4 MCDermott Company
ISLHT the limits imposed by the outlet nozzles during cooldown also control the composite limit curve at the low temperatures.
However, for the ISLHT, the beltline region would eventually control at higher EFPY.
At 5 EFPY the temperature difference between the RTNDis of the beltline and the closure head region materials is not large enough to compensate for the higher stress intensities that the closure head region is subjected to (at the same internal pressure).
Figure 5-7 shows the development of the minimum pressure-temperature limit curve for reactor core operation up to 5 EFPY based on Appendix G to 10 CFR 50.
The references used here are the limit curve for nonnal heatup and the minimum permissible temperature for the ISLHT pressure.
The data used for Figure 5-7 are the composite limit curve of Figure 5-1 and the minimum per-missible temperature for 2500 psi obtained from Figure 5-6.
The critical-ity limits imposed by the Technical Specifications are based on other con-siderations since these limits are not controlling.
5.2.
Technical Specification Limit Curves The Technical Specificiations for each plant give allowable pressure and temperature combinations and require that the RC system be maintained with-in these limits during normal heatup and cooldown, criticality, and inser-vice leak and hydrostatic tests.
The objective of these pressure-tempera-ture limits is to prevent stresses from exceeding the ASME Code maximum allowable design stresses and the stresses allowed by ASME Appendix G for protection against nonductile failure.
Since the stresses allowed by ASME Appendix G are generally more restrictive than the Code maximum allowable design stresses, the Technical Specifications pressure-temperature limits are the nonductile fracture prevention limits presented in section 5.1.
However, there is one exception:
Ouring cooldown, the stresses in the steam generator tubing may exceed the ASME Code maximum allowable stresses if cooldown rates are high, and the allowable pressure-temperature combination during cooldown is calculated according to ASME Appendix G.
When high cooldown rates are desired, the pressure-temperature limit curve is modified by reducing the allowable pres-sure, which reduces stresses in the steam generator tubing.
J 5-3 Bat > cock & Wilcox a McDermott company
Figures 5-8 through 5-10 are pressure-temperature limit that illustrate the limit curves that appear in the Technical Specifications of a typical 177 FA type plant.
Figure 5-8 represents the nomal heatup limits applicable for the first 5 EFPY.
Figure 5-9 represents the normal cooldown limita-tions, and Figure 5-10 represents the ISLHT limits.
These figures were ob-tained from Figures 5-1, 5-3, and 5-6, respectively.
Figure 5-7 was also used to develop Figure 5-8.
Figures 5-1, 5-3, and 5-8 were adjusted as follows:
1.
The maximum allowable pressure had been reduced by the pressure differ-ential between the point of system pressure measurement and the limit-ing region of the reactor vessel for all operating pump combinations.
The applied pressure differential is 100 psig when either the beltline region or the outlet nozzles control the pressure-temperature curves and 75 psig when the closure head region controls them.
These pressure differentials have been conservatively calculated.
2.
Figures 5-3 and 5-6 were adjusted to include the pressure-temperature limits imposed by the steam generator tubing.
For some 177 FA plants and other B&W plants, the Technical Specification limit curves will be different from those presented in Figures 5-8 through 5-10.
The differences are caused by the lower maximum allowable remp rates and the material's RTNDT, wall thickness, neutron fluence, etc.
5.3.
Preservice System Hydrostatic Test Limit Curve The Technical Specifications do not include the RC system pressure-tempera-ture limits for the PSHT since this test is conducted before the plant operating license is issued.
The limits for the PSHT are imposed by the test procedure.
Figure 5-11 is the PSHT limit curves as developed by adjusting the com-posite limit curve of Figure 5-5.
The adjustments are the same as those used to develop Figures 5-8 through 5-10.
Figure 5-11 is labeled as the PSHT limit curve for the typical 177-FA type plant.
The actual curves may differ since this curve was calculated using 100*/ hour heatup and cooldown ramp rates and during the PSHT, the ramp rates are much lower than 100F/
nour.
f 5-4 Babcock & Wilcox A MCDermott company
Table 5-1.
Unirradiated Impact Properties and Residual Element Content of Beltline Region Materials in a Typical 177 FA Plant Core MP Transverse Material identification to weld Cy USE, RTNOT.
type, location CL, in.
ft-lb F
Cu P
S A.
SA 508 Class 2, nozzle 183
+10 0.0 54 0.008 0.006 belt B.
SA 533 8 upper shell 88
+30 0.20 0.008 0.016 C.
SA 533 B upper shell 90
+20 0.20 0.006 0.016 D.
SA 533 B lower shell 119
-20 0.12 0.013 0.015 E.
SA 533 8 lower shell 99
+40 0.12 0.013 0.015 F.
Weld, upper long.
(66)(a)
(+20) 0.20 0.009 0.009 G.
Weld, upper circ 123 (66)
(+20) 0.19 0.021 0.016 En H.
Weld, mid circ (100%)
-63 (66)
(+20) 0.27 0.014 0.011' I.
Weld, lower long. (100%)
(66)
(+20) 0.22 0.015 0.013 J.
Weld, lower circ (100%)
-249 (66)
(+20) 0.20 0.015 0.021 K.
Weld, outlet nozzle 244.8 (66)
(+20) 0.19 0.021 0.016 (a) Numbers in parentheses indicate predicted values.
sit
?R 2sr 5 C) i 11 s ir
- O M
l
.=..
T &le 5-2. Typical Material Data for Priparing Beltline Region Pressure-Tagerature Limit Curves End of 5th EIPY End of 32nd EFPY Fluente, E > 1 MeV Fluence, E > 1 MeV 1
n/ m2 ARTg, F RTg, F n/an2 tRTg, F RTg, F Mat 10 F
1/4 t 3/4 t 1/4 t 3/4 t 1/4 t 3/4 t 1/4 t 3/4 t 1/4 t 3/4 t 1/4 t 3/4 t A
+10 2.63E18 5 5 17 30 16 40 26 1.68E19 3.8E18 105 3
115 45 B
+30 2.63E18 5.9E17 75 3
105 68 1.6E19 3.E18 220 90 250 120 C
+20 2.63E18 5 5 17 76 3
95 58 1.68E19 3 K 18 220 90 240 110 D
-20 2.63E18 5.9E17 45 21 25 1
1.5E19 3.E18 140 59 130 39 E
440 2.63E18 5.9E17 45 21 f6 61 1.68E19 3.E18 150 69 180 99 F
+20 2.2SE18 5.1E17 70 34 90 54 1.44E19 3.2SE18 212 85 232 110
[
G
+20 2.63E18 5.9E17 75 3
95 58 1.68E19 3K18 220 90 240 110 H
+20 2.63E18 5.9E17 75 38 95 58 1.68E19 3.E18 220 90 240 110 1
+2D 2Kl8 4.53E17 64 32 84 52 1.2BE19 2.9E18 200 80 220 10 0 J
+20 4.75E16 0
0 20 20
<5.6E18
<50
<?O K
+20
<8.75E16 0
0 20 20
<5K18
<SO
<70 1
f sr ER 35D SE i iF
- O M
e
,w-
1 Figure 5-1.
Nomal Heatup Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve for S EFPY 2400 2200 7
j LEGEND
/
/
/
2000
- - -- --- Beltline, SS, 1/4 t
/
Beltline, 100 F/h, 3/4 t
/
I 1800
/
Closure Head, 100 F/h 7
1600 Outlet Nozzles, 100 F/h
/
Composite Limit Curve-f
/
1400 Note: At temperatures > 140F, the composite f
T limit curve is controlled by the f
u.L elesure head.
/
,/
1200 f
f 3
/
a 1000
,/
/
/
/
800 f
~
/'O
~ W =.
600
,~
f A** " d NDT Temp. F e er gn 400 Beltline 1/4 t 105 eR j3r flinimum Bolt-up Temperature Beltline 3/4 t 68 5
200 Closure head 60
$k 60
}k 0
l l
l lOutletnoyles y
O 40 80 120 160 200 240 280 320 Fluid Temperature, F
Figure 5-2.
Nonnal Heatup Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Comp 3 site Limit Curve for 32 EFPY 2400
./
e 2200 LEGIDED 2000 seitline, SS, 1/4 t seitline, 100 F/h, 3/4 t g
I
~
Closure Head, 100 F/h gg Outlet Mozzles, 100 F/h Composite Limit Curve l
1 1400 3
=
o>
- 1200 j
g 5
./
E 1~
/
/
800
/
/
/
Assumed RT g
y W
\\ m_ #
g__%
Beltline 1/4 t 250 g
Beltline 3/4 e 120 g
400
!2 Closure head 60 D
gg 200 Outlet nozzles 60 2:
'<-Minimum Bol t-up Temperature J {l 0
I I
I I
I I
I 60 100 140 180 220 260 300 340 380 420 Fluid Temperature, F e
Figure 5-3.
Normal Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve for 5 EFPY 2400 l
2200 -
t I
LCCEND 2000 -
- - - -- Belt 11ce, SS, 1/4t Beltline, 100 F/h, 1/4t l
800 -
Closure licad, SS
~
~
I 1600 Composite Limit Curve
- 1400
/
Y' O
e a
/
/
- 1200
/
/
- 1000
/
sL*
800
' /
s,-
600
,- ~
Assumed RTM T% F
.R f
400 p
Beltline 1/4t 105
- k Closure head 60 3 3r M
200 Mininm Bolt-up Tecperature Outlet nozzles 60
$k i
i i
i l
1 28 0
40 80 120 160 200 240 280 320 M
Fluid Temperature, F i
i E
l
Figure S-4.
Normal Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve for 32 EFPY 2400 2200 -
/
I ucrenn I
2000 '-
- -- --- Beltline. SS. 1/4t Beltline. 100 F/h. 1/4t 1800 Closure Head, SS 1600
- Outlet Nozzles. 100 F/h a
Composite Limit Curve j
7 g 14~
l l
/
j 1200 I
1000 5
i
,/
/
800
/
/
p/
Assumed RT l
m Tm F s
~
~
f pp Beltline If4t 250 600
,[
" " ~ ~
Closure head 60 5{
400 Outlet nozzles 60 in J
oW
- p 200 B, g
+- Minimum Bolt-up Temperature g
g g
0 60 100 140 180 220 260 300 340 380 420 M
Fluid Temperature. F O
e e
r__
Figure 5-5.
PSHT Heatup and Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Composite Limit Curve 3200 3000
-.- Heatup Outlet Nossles I
--- Cooldown Outlet Nozzles,
I 2800 Heatup Closure Head 2600
- - Cooldown Closure Head I
Composite Limit Curve #
2400 Note:
The composite limit I
curve is the heatup closure head curve.
I 2000 1800
[
1600 p
1400 1200 f
1000 j
/
\\
800 J
J 600
^**""*
NDT Temp. F 400 fiininum outlet nozzle 60 Bolt-up -+
Temperature Closure head 60 200 l
1 l
l l
l 0 0 40 80 120 160 200 240 280 Fluid Temperature, F 5-11 Babcock & Wilcox
& MCOtti'v'tt COmpefty
~
ISLT Heatup and Cooldown Pressure-Temperature Limits Imposed by Several Reactor Vessel Regions and Figure 5-6.
Composite Limit Curve for S EFPY q
3400 LEGEND g
- Composite Limit Curve
-- --- Heatup and Cooldown Beltline, SS,1/4 e
~
3200 -"
- Bestup Beltline,100F/h, 3/4 t
~ A - Cooldown Beltline 100F/h, 1/4 t 3000 -
- Beatup Outlet Nozzles
- Cooldown Outlet Nozzles 2800 -
- Heatup Closure Head
- ~ ~ Cooldown Closure Head 2600 -
I
/
l f
5F j
/
t W eratures the composite limit i
/
2400 -
curve is controlled
/
by the closure head.
/
/
j 2200 f
N/
i 2000 -
[/
/
1800
/
/
E
/
~
/
1600
/
I if,/
5 1400
/
p
/
//
1700 1000-0y- ~
F 80 J
600 Assumed Rt 105 Beltline 1/4 68 Beltline 3/4 t 60
~/
Closure head 60 7
outlet nozzles 4~- flinimum Bolt-up Temperature 200 OJ i i e e i i i t
1 J f I I f f f I I t I i
I 8
.l 300 260 220 180 140 100 60 Fluid Temperature, F Babcock &Wilcox 5-12 J McDermott company
l o-1 Figure 5-7.
Determination of Reactor Core Operation Pressure-Temperature Curve for S EFPY per Appendix G to.10 CFR 50 2600 2400 Inservice Tests 2200 2000 1800 1600 O
Normal Heatup (100F/h) P-T Limits m.
=
1400 U
.7 1200 Core
=
Operation u
6 40F Limits 1000 800 W
600 g
2R g,
400 5m Sk 200 2=
E o l
M 060 100 140 180 220 260 300 340 380 l
Fluid Temperature, F
Figure 5-8.
Normal Operation Heatup Pressure-Temperature Limit Curves for Typical Plant Technical Specifications, Applicable up to 5 EFPY me i 2400 i
j POINT PRES 3uRE TEMPERATURE D
A 405 60 l
B 550 148
~
C 550 273 h 2000 j
e 2250 303 E
J c
1
._ 1800 f
=
E
~
l 3 1600 THE ACCEPTABLE PRESSURE AND TEMPERATURE COMBINATION ARE
(
[
BELOW AND TO THE RIQlT OF THE LIMIT CURVE. THE REACTOR r
3 MUST NOT BE MADE CRITICAL UNTIL Tile PRESSURE-TEMPERATURE l
T l*.
1400 IS 525F OR GREATER.
5 E
~
j
\\
1200 s
r B
N APPLICABLE FOR NEATUP 9ATES OF 1000 g
< 100*F/t ({50*F IN ANY I/2 2
NOUR PERIOD)
$ 800 C,g 3 600 8
a an 5E A
_<-- Bol t-u )
Beltline 3/4 t 6 8,_.---
Minimu a Beltline 1/4 t 105 5D j
jg
- 200 Temper iture Closure head 60 3 ;7 Outlet nozzles 60 x o i
i i
e i
i 60 100 140 180 220 260 300 340 380 Indicated Reactor Coolant System Temperature, Tc, F w
e i.
I e
I
,o cm wo M
l w
w N
>= %
as 4 -
m a
w w
ooe N w y\\
ceoo_.m e oo-w o
m =
C.
C@@@
S M
x o a
.J w
o H
w ow uO oa as o M@
N w
w oooo e.
w== a as y a g
C p
o
- e e as e -
n H
@ "' w m e e on w N
w o
449-a n
en E
.J z -
N c6 N w
f a% as
- it; N.4 M ar=
J a
w w w l-e c
@ N as c
uo a.
Cai
.c O a.
o
@ U C
~
G) a o as y
g g g 6,O.
- ~
- vs we W u c4
- # U U xo o a.
- N/l
- 9 W "u
_a W W O U m >=
x w J
'b ""
O
,s y aw o
us
- >= J
_m M **
.o a.
(4 3
o cm e Q) %
r
> n as a Q. ae=
uw o
a
- C CQ cQ U O a.
N= v EO o
as cs xw o a.
w z a z c) as
.a w
s ow as >- -- ma p c, a z
& & w g g m
e a
a.
>=
u o
z w. o o i.e.
>=
b "O
- >=
3 3 W
.a
>- o za-as 4 U
w M U sa 3
ao O,
u w
w >= o
.J z
M va
. w o as
.J as >- as e
- =
OC N m s
a as as L.c a.s
.a. >=
= w as cw Q,,
y e
aj
- = g 3
- a. o a as
==
e e as oo- - o cg a
as w 4
as
- a. w
~
as G
5 Q
o w L 4 xL L a E
99y "O C Q.
- e. x g w w o an as p
ow
>=
>=
& f0 6 4
- a. m z
- e. e
- > x p-
==
0 m La.J z
0 c.
om o E e tw as o O as Q
Lf')
- en as w
>= W e Q
N
- X x=
0 as e
C re C o
w >=
in a uw w ou4 u
- w
>= c as 4 J
to n as w as w
w tsJ W es w
M O O-as e x oa z a.
3 c=
O h3
'L H
.m w
- >- xo eo w
w
.J "o'"
OJ as as M o a M >=
.J Q. L m as z >
m >= =
Q o o O.c w
oe as -
x en CJ
% re a.
w z 3s o as n
-=
x a:
u w n -
g.U, b
r-w J X S E M o
q
>=.
e a w o -ww w
as
-w as M o >
co n >= W es Um O 3 C.
s c;J xw
>- e o as w as Z U EC
< m w >=
as as x as CE W
as as e
Es w M oO e
w uo w a.
w w as >- >-
-w
- a. o w a o C)
C L
o u m -
.a on x
w a a
m
. M x w E m W x o >= A as c
N **
= m g
n -
z -
>- o n x a p
u m
w
.J w
as 4 es u g c, rg w
> w e
- a. m z m as 3 3 g.
'e-g
- a. w w a w
u as e,
E f CJ
~
=c u >- a
>- n x
(g b
"C r M
.z.
w
.J e a.
w I zw 3
w a
Cn
.J o.=
x as a
p.
- e-O C
w=
mw
.o z xm na w E cc H w
e o w ae a a. > no a.
z -
x
- e. ow
<=
- e..=
a
-u. xo-
=
w =
. w
.a su w = - a 3
uo 3 as a
=
- a. a.
u =*
w wM xJ z w w
=s as a
- a. w
.J w u a z as o>
p-as >. >-
ty V
w nsJ z z e
e
~
<c l
1 1
I t
t I
f f
I t
i t
<=
c=
c=
n C=
m o
e es a
<=
=
<=
c Q
Cs Q
O O
O
<=
Cs Q
C=
c,
~
o
=
m n
co.
=
~
=
m m
n 3!sd '(Jaz!JnssaJd ulig dool) aJnssaJd 2815AS tueloo3 Joloeau paleo!put Babcock & Wilcox 5-15 4 MCDerrnott Company
Figure 5-10.
Inservice Leak and Hydrostatic Test Heatup and Cooldown Pressure-Temperature Limit Curve for Typical Plant Technical Specifications, Applicable up to 5 EFPY 3000 THE ACCEPTAILE PRESSURE AND TEMPERATURE CGKllu T10NS ARE BELOW
~
AND TO TME Rl6MT OF THE LIMif CURVE.
~. NMEN COOLING 13 ACNIEVEB tnt 0USN THE STEAM GENERATots. TME P.T
~~
1 LIMITS FROM PolNT E TO POINT 0 0F THE MONNAL C00LOOWN LIMIT CutvE 2600 Att APPLICABLE (FIGURE s-9) 2.
FOR C00LDOWN. NOTES 1 AND 2 ON FIGURE 5-,
0 as I
{ 2400 ARE ACCEPTAsLE I
O 2200 l
.5 E
y 2000 E
0 1800 E
APPLICABLE FOR NEATUP AND C00LDOWN(f.2)
[IC0'F/M
(<~5 0' F IN ANY 1/2 Mou: PEtl00) a 1600 N
j j
1400 1
2 a
1200 I
h 8
1000 NDT Temp. F j
Beltline 1/4 t 105 Beltline 3/4 t 68 g
800 u
Closure head 60 g
Outlet nozzles 60 f
3 f
600 B
Cj
/
l
[=
400
/
Pol'T PatssutE TEMPERATutE A
265 60 g
A g
550 100 3
200 C
550 255 7
0 2500 272 Minimum Bolt-up Temperature
_j l
t i
I I
I I
I I
I i
i t
I i
i e
i t
i t
l 60 100 140 160 220 260 300 indicated Reactor CO0lant System Temperature, TC. F 5-16 Babcock & Wilcox a McDermott company L
Figure 5-11.
PSHT Pressure-Temperature Limit Curve for Typical Plant 3400 THE ACCEPTABLE PRES 5URE AND TEMPERATURE. COMBINATIONS ARE BELOW
~
AND TO THE RIGHT OF THE LIMIT CURVE.
3200 C
1.
WHEN COOLING 13 ACHIEVED THROUGN THE STEAM GENERATOR $,
f THE P.T LIMITS FROM PotNT E TO POINT 0 0F THE NORMAL f
C00LDOWN LIMIT CURVE Ar4E APPLIC ABLE (flGURE J -9),
2800 - 2.
FOR C00LOOWN. NOTES I AND 2 ON FIGURE 5-9 ARE APPLICA8LE.
I 2600 C
2400 1
5
)
l 2200 l
[
3 2000 i
G l
1800 U
APPLICA8LE F4R MEATUP AND f
C00LDOWN II'2IRATES OF h
< 100*F/H
(< 50'F IN ANY U
~
5/2MOUR PERIOD) 1400 e
i
=
3 1200 m
l E
1000 O
3 800 l
u l
2 1
600 A
1 u
~
S fiinimum Bolt-up 400 550 70 F
Tempera ture
=
~
8 550 215 C
3:25 252 200 l
l t
t 1
i i
t i
i I
0 i
i t
i f
I 40 80 1-20 160 200 240 280 Indicated Reactor CO0lant System Temperature, TC, F i
l l
5-17 Babcock & Wilcox a McDermott company
o
,. ~ '
6.
EPFM ANALYTICAL PROCEDURES 6.1.- Basis The analytical procedures given in Section 4 are applicable for the areas of the pressure boundary which comply with the material restrictions of ASiE Appendix G.
If the material does not comply with the restrictions then supplemental analysis is required to assure the reactor coolant pres-sure boundary i ntegri ty.
The only anticipated divergence from the mater-ials restrictions is the failure of irradiated materials, in particular weld metal, to exhibit a Chany upper shelf exceeding 50 f t-lbs.
If a material exhibits less than 50 ft-lbs absorbed energy but greater than is determined in accordance with 10 30 f t-lbs the adjusted shif t in RTNDT CFR 50 Appendix G.
The analysis of section 4 is carried out in the same manner previously discussed.
10 CFR 50 Appendix G further specifies that Charpy upper shelves below 50 f t-lbs are permitted if the component is verified to still have a margin of safety equivalent to that specified in ASME Appendix G.
The only area of the reactor coolant boundary which is predicted to potential ly fall bel ow 50 f t-lbs is the reacto r vessel bel tline.
This evaluation will be restricted to that area but similar evaluation could be perfomed on other areas.
Appendix G of the ASME Code is a design guideline for the prevention of non-ductile failure. The general philosophy is to index the fracture tough-ness to temperature and require that the component be operated at a suffi-ciently high temperature to preclude non-ductile failure.
ASME Appendix G is not adequate to control operating conditions in the higher temperature regime.
In the high temperature regime ductile tearing is the controlling mechanism for possible loss of vessel integrity.
Evaluation for ductile tearing can be accomplished utilizing the J-integral and the J -R curve for t
the material.
Babcock & Wilcom 6-1
,uco,, moi,com,,ny
6.2.
Elastic-Plastic Fracture Mechanics (EPFM) Analytical Model An elastic-plastic fracture mechanics (EPFM) procedure based on defomation plasticity J-integral solutions in the fomat of a failure assessment dia-gram will be used to set the pressure-temperature limits for upper shelf material behavior.
The procedure fo r setting these pressure-temperature limits consists of four steps:
1.
J-integral fomulation.
2.
Failure assessment diagram curve generation.
3.
Assessment point evaluation.
4.
Instability pressure prediction.
For reactor vessel materials which can be modeled by defomation plasticity and whose stress-strain behavior can be represented by a power law strain-hardening equation, the J-integral response (Japplied) can be evaluated for the reference flaw using the expression J = J0(a ff,P) + JP(a,P,n)
(1) e where de is the el a stic contribution based on Irwin's effective crack is the defomation plasticity contribution derived in d ep th, a ff, and Jp e
reference 6.
P is the applied pressure and n is the strain-hardening expo-nent.
A convenient way to use this equation is to construct a defomation plasticity failure assessment diagram (DPFAD).
The details of this proce'-
dure are found in references 10 and 11.
The process is summarized here only for the beltline flaw evaluation.
6.2.2.
DPFAD Curve Generation The DPFAD curve expression is obtained by nomalizing the sum of the elas-tic and plastic response by the "elastic" J-integral of the flawed reactor vessel in tems of "a", where (1. v2 )g{(a)
(2)
Je(a) =
g and K! is the linear-elastic fracture mechanics (LEFM) stress intensity fac-to r.
E and v are You'ng's modulus and Poisson's ratio, respectively.
The normalized J-response is then defined by Babcock & Wilcox 6-2
, ucoumon comp ny
Kr=
= f(Sr)
(3) where S = P/P (a).
r L
P is the applied pressure and PL is the reference plastic collapse pressure or limit pressure, a function of "a" and the material yield strength, g.
Equation 3 defines a OPFAD curve which is a function of the flaw geometry, structural configuration, and stress-strain behavior of the material of interest.
This curve, in tems of K *Sr is independent of the magnitude of r
the applied loading.
For the beltline area of the reactor vessel assuming a semi-elliptical axial flaw on the inside of the vessel.9 Kg =
F(a/t.a/t)
(4) where P = applied pressure, Q = 1 + 4.593(a/t)1.65 Rt = inside radius, t = length of flaw t = thickness, a = flaw depth 1 + H (a/t)2 + M3 (a/t)4]fc, F =.97[M 2
Mi = 1.13 -0.18 aft, M2 =
.54 +.445/ (.1 + a/t),
M3 =.5 - 1/(.65 + 2a/t) + 14 (1-2a/t)24, fc = 1.152
.05 /a/t.
then J8(a) = P R!za_ F2 (1.y2) 2
()
t2 Q
E The effective crack correction is given by 2
1 (n-1) 1 a f f = a + E (n+1) 4I
- 2 e
j 1 + Sr where n = strain hardening exponent,(Ramberg-Osgood) o = engineering yield stress, c
l S = P/P,
r L
(t - a*)
p l * /3 0(Rj + a*)
'-0 l
I 6-3 Babcock & Wilcox a McDermott company i
-~
The limit pressure expression Pt is based on a continuous axial fl aw.
A correction is applied in the form of a* to account for the partial length flaw.
a(1 - s) a* = 1 - la/ t)s (6) where:
/2t )-1/2 2
s = (1 + 12 The plastic portion of J is given by the following expression o
a(1 - a/t)h (P/P )n+1 (7) i L
where a is obtained from the Ramberg-Osgood stress-strain relation and h1 is a dimensionless tenn which is a function of a/t, a/E, n and t/Rj.
This latter constant is evaluated from finite element results.8 Combining all of the above tenns into equation (3) results in an equation which when plotted has the shape shown in Figure 6-1.
The DPFAD curve is unique for a given set of stress-strain parameters, flaw size and vessel geome try.
6.2.2.
Assessment Point Evaluation Having defined the DPFAD curve the beltline of the vessel can be evaluated for a given set of material properties.
Assessment points are denoted by Kf, Sh and are defined as follows:
(ao + aa)
(3)
Xf(ao + oa) =
JR (aa)
(9)
Sf(ao + aa) = p
(,
33) where terms are as defined in section 6.2.1 with J ( A3) being the material R
I I
is the initial assumed flaw.
J -R resistance property.
ao 6.2.3.
Instability Pressure Prediction To evaluate the structure, the applied pressure is held constant and succes-sive points are cal cula ted incrementing the crack size.
The assessment point which is the minimum distance from the origin represents the maximum crack growth which the structure can sustain before becoming unstable. The 64 Babcock & Wilcox a McDermott company
~
correspondi ng point on the DPFAD curve then represents the instabili ty
~
pressure designated by Pcrit. This proce y it illustrateo ir Figure 6-2.
If the initial point eval ua ted is Jg(ad = J IC then the pressure can be detemined which corresponds to the initiation of ductile tearing.
This pressure is designated F nit and is also illustrated on Figure b.2.
i l
6.3.
Sample Calculation and Presentation of Data For further clarification of the failure assessment diagram approach te predicting tearing pressure, a sample calculcation is presented of an AStE Section III, Appendix G flaw in a beltline weld in a reactor pressure vessel under a pressure of 2500 psi.
Figure 6-3 and Table 6-1 present the FAD fomat while Figure 6-4 presents a pl ot of the Jg (a a) curve for the weld material.
Figure 6-5 shows the resultant tearing pressure versus stable crack growth, i a, as well as the local plastic instability pressure calculated by the ratio 2500/Sr' (Sr' is given in Table 6-1 as a function of aa).
The critical pressure is the lower value of the two curves.
In al l the figures and the table, the numbers refer to the points plotted; initiation is the point numbered #1 (J =JIC) while the instability or the critical point is numbered #5.
6.4.
Thermal Stress Themal stresses are not considered when eval ua ting ductile tearing.
Themal stresses arising fra radial gradients through the wall are self-limiting and will decrease with crack propagation.
Furthemore, for the conditions being considered (i.e., normal and upset transients in the power operating range) the themal contribution to the J applied is small calcu-lated on an elastic basis.
6.5.
Acceptance Criteria The acceptance criteria for the evaluation are two fold.
Although the flaw used in the evaluation is hypothetical it is necessary to demonstrate that ductile initiation wil l not occur-to preclude assuming inc remental ref-erence flaws throughout the life of the plant.
Therefore, the first crite-ria is that the initiation pressure, Pinit, must be greater than 3000 psi.
This value is one-third above the nominal operating pressure and ten-per-cent above any nomal or upset anticipated transients.
6-5 Babcock & Wilcox a McDermott company 4
The second criteria is that the instability pressure, Pcrit, must exceed two times the highest level A or 8 operating pressure.
For B&W designed nuclear vessels this correspo nds to 5500 psi.
These criteria will be reflected in the Owner's licensir.g document by specifying a maximum allowed pressure in the Technical Specification of 2750 psi for temperatures in the operating range.
i 6-6 Babcock & Wilcon a MCDermott company I
4 7.
S'JMMARY AND CONCLUSIONS B&W's ' methods of compliance with the material properties and operational limit requirments of Appendix G to 10 CFR 50 have been described.
Since Appendix G specifies fracture toughness requirements for the ferritic mater-f als used in the RCPB and provides guidelines for determining its operating limitations, the RCPB is described first.
Section 2.3 describes the operational parameters of each loading condition for which pressure-temperature limit curves are required; these conditions are as follows:
1.
Nomal operation, including bolt preloading, heatup, and cooldown.
2.
Preservice system hydrostatic test.
3.
Inservice syste leak and hydrostatic test.
4.
Reactor core operations.
Section 3 describes the methods of cmpliance with these material require-ments.
Section 3.1 covers ferritic materials other than bolting and type 403 stainless steels.
As required by Appendix G to 10 CFR 50, the RTNDTS of these materials must be established in order to determine the pres-sure-tmperature limit curves for the RC systs.
For later plants (ordered according to the Summer 1972 Addenda to ASME Section III or subsequent editions or addenda), the RTNDTs were obtained as required by the appli-cable ASME Code.
The RTNDTs of the other ferritic materials in the older plants were conservatively estimated using the fracture toughness data obtained on low-alloy steel forgi ng s, plates, carbon steel plates, weld metals, HAZs and piping.
Appendix G (10 CFR 50) also raquires full Charpy test curves on the belt-line region materials to determine the USEs for the more recent reactor ves sel s.
B&W has specified emplete Charpy test curves (nomal and paral-1el to the principal worki ng direction) on the base metals; for wel d l
7-1 Babcock & Wilcox A MCDermott CompJny i
4
'.~.
metals, only one curve is needed.
These curves were also obtained for the HAZs of the beltline region base metals (s) selected to be nonitored by the reactor vessel surveillance program.
For older reactor vessels, Charpy test curves (both directions) were obtained on the materials from the sur-veillance progaam.
Where enough material was available for testing, Charpy test curves for spdmens oriented normal to the principal working direc-tion were obtained for materials not included in the program. For any belt-line region materials for which no test material was available, the Charpy USE was conservatively estimated from da ta obtained on beltline region low-alloy steel forgings, plates, and weld metals.
The fracture toughness properties of bolting materials are discussed in section 3.
The requirements and acceptance criteria of Appendix G to 10 CFR 50 are compared to those of AS4E Section III.
Since the bolting mater-ials ordered before August 16, 1973, meet only the requirements of the applicable AS4E Code, it is demonstrated that these materials have adequate toughness for protection against nonductile failure.
The fracture toughness of type 403 modified steel is covered in section 3.3.
The test results demonstrate that the material has adequate fracture toughness for protection against failure at 40F.
However, B&W specifies a minimum service temperature of 100F for the CRDM, which provides appropri-ate consenatism.
Section 3.4 describes the supplemental material properties generated to as-sess the reactor vessel for resistance to ductile tearing instability.
These properties are the stress-strain characteristics and the materials resistance to ductile tearing as a function of crack growth.
This section also discusses the method of merging the LEFM criteria of section 4 with the EPFM criteria of section 6.
Section 4 describes the basis for the methods used by B&W to obtain the pressure-temperature limit curves.
The calculational procedures are pri-marily based on ASME Appendix G and WRC Bulletin 175. Yhe method of deter-mining the pressure-temperature limit is described for each loading condi-tion of interest.
7-2 Babcock & Wilcox a McDermott company
In section 5, the methods presented and described in sections 3 and 4 are applied to a typical 177 FA plant.
Figures in section 5 illustrate (1) the development of the composite limit curves for each loading condition of interest, (2) the development of the reactor criticality limit curve, (3) the limit curves appearing in the Technical Specifications for a typical plant, and (4) the pressure-temperature limits for the preservice system hydrostatic test.
In section 6 the methods are described for qualifying the reactor vessel in the event that a Charpy upper shelf energy of 50 f t-lbs is not obtained.
This section determines the pressure limits for ductile tearing instability and statec the acceptance criteria.
The basis of this analysis is the J-integral and the supplemental fracture toughness data described in para-graph 3.4.
As described in this report, the fracture toughness requirements imposed on the ferritic materials of pressure-retaining components of the RCPB of B&W reactor coolant systems are in compliance with the fracture tcughness re-quirements of Appendix G to 10 CFR 50.
In addition, the report demon-strates that the ferritic materials ordered before the ef fective date of Appendix G to 10 CFR 50 have adequate toughness for protecting against non-ductile failure when the system is operated in compliance with the pres-sure-temperature limit curves developed by B&W.
The analytical method employed by B&W to calculate the maximum allowable pressure of the RC sys-tem as a function of fluid temperature includes all the margins of safety required by Appendix G.
7-3 Babcock & Wilcox a McDermott company
__._..a
ns.~-
c
,x r
8.
REFERENCES 1.
H.
S. Palme,- Radiation Embrittlement Sensitivity of Reactor Pressure Yessel Steels, BAW-10056A, Babcock & Wilcox, Lynchburg, Virginia, August 1973.
2.
H. S. Palme, et al., Reactor Vessel Material Surveillance Program --
Compliance With 10 CFR 50, Appendix H,
for Oconee Class Reactors, BAW-10100A, Babcock & Wilcox, Lynchburg, Virginia, February 1975.
3.
PVRC Recommendations on Toughness Requirements for Ferritic Materials, PVRC Ad Hoc Group on Toughness Requirements, WRC Bulletin 175, August 1972.
4.
J.
M.
Barsom and S.
T.
Rolfe, "Correlation Between Kic and Charpy V-Notch Test Results in the Transition Temperature Range," Impact Testing of Metals, ASTM STP-466 (1970), pp 281-302.
5.
ASME Code Case 1337, "Requirements for Special Type 403 Modified Forg-ings or Bars."
6.
Y. Kumar, et al., "An Engineering Approach for Elactic-Plastic Frac-ture Analysis", EPRI Topical Report NP-1931, Research Project 1237-1, Electric Power Research institute, Palo Alto, CA., July 1981.
7.
Intentionally lef t blank.
8.
H.
G.
delorenzi, "Elastic-Plastic Analysis of the Maximum Postulated Flaw in the Beltline Region of a Reactor Vessel," ASME Journal of Pressure Vessel Technology, Vol.104, No. 4, November,1982.
9.
J. C.
Newman and I. S.
Raj u, "Stress-Intensity Factors for Internal Surface Cracks in Cylindrical Pressure Vesseb," Transactions of ASME Vol.102, November,1980.
l l
8-1 Babcock & Wilcox a McDermott company
~.w.,
4 10.
J. Bloom and S. N. Malik, "A Procedure for the Assessment of the Integ-rity of Nuclear Pressure Vessels and Piping Containing Defects," EPRI Topical Report NP-2431, Research Project 1237-2, Electric Power Re-search Institute, Palo, Alto, CA June 1982.
11.
J. M. Bloom, "Extensions of the Failure Assessment Diagram Approach --
Semi-Elliptical Flaw in Pressurized Cylinder," presented at the 1983 ASME Winter Annual Meeting, Boston, MA, reprint 83-WPVP-3, American Society of Mechanical Engineers, November 1983.
1 8-2 Babcock & Wilcox 4 MCOttmott Company
-