ML20211D226
| ML20211D226 | |
| Person / Time | |
|---|---|
| Site: | Seabrook  |
| Issue date: | 08/26/1985 |
| From: | Cummings G, Ariuska Garcia LAWRENCE LIVERMORE NATIONAL LABORATORY |
| To: | Davis S Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML20209C800 | List:
|
| References | |
| CON-FIN-A-0801, CON-FIN-A-801, FOIA-87-6 RARE-85-141, NUDOCS 8702200359 | |
| Download: ML20211D226 (4) | |
Text
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s Lawrence Livermore National Laborato: y l ,
NUCLEAR SYSTEMS SAFETY PROGRAM RARE 85-141 August 26, 1985 i
Ms. Sarah M. Davis i Reliability and Risk Assessment Branch Division of Safety Technology i U.S. Nuclear Regulatory Commission Washington, D.C. 20555
SUBJECT:
Monthly Management Letter #15 Month of July 1985 NRC FIN-A0801 Review of the Seabrook Probabilistic Risk Assessment
Dear Ms. Davis:
- 1. Project Description and Objective The Office of Nuclear Reactor Regulation is conducting a probabilistic risk assessment (PRA) review program in which PRAs performed and submitted to the NRC by license applicants and licensees receive comprehensive review and evaluation. The program is the responsibility of the Reliability and Risk Assessment Branch (RRAB).
A PRA of the Seabrook Nuclear Power Plant (Seabrook PRA) has been submitted to the NRC by Public Service Company of New Hampshire, an operating license (OL) applicant. The review of this document, whose title is "Seabrook Station Probabilistic Safety Assessment," is being performed as one project in the larger NRC program.
The objective of this project is to perform an expeditious and cost effective review of those aspects of the Seabrook PRA leading to the estimates of the frequencies of each plant damage state and the associated uncertainty spread to determine the accuracy of these estimates. The review will cover method-ology, assumptions, data, information sources, models, plant understanding, completeness of the analysis and other areas where inconsistencies may arise which could affect the quantitative or qualitative results.
- 2. Progress for Period None, foyM - b ]~ C 0 &
- 3. Work to be Accomplished Next Month 0702200359 870211
[h3 None. PDR FOIA SHOLLYB7-6 PDR An Eoua'Owortaryy Emoor e'
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O. August 26, 1985
- RARE 85-141
- 4. List of Subcontractors and Consultants
- a. Subcontractors (1) Applied Risk Technology Corporation (ARTECH); P. J. Amico (2) Jack R. Benjamin and Associates, Inc. (JBA); J. W. Reed, M. W.
McCann, Jr.
- b. Consultants (1) P. R. Davis
- 5. Concerns None.
- 6. Other The $28,900 lien shown on Attachment A appears on this report because of procedural requirements for contract accounting. These funds have not and will not be spent.
Sincerely,
. s-sce ,(
Abel A. Garcia Principal Investigator
[ 6*tt Garth E. Cummings Program Leader Nuclear Systems Safety Program sr l
Enclosures (2)
I l Distributions l
DOE LLNL f NRC R. Frahm, DST W. J. Gallagher G. E. Cummings i J. Halvorsen, DST J.M. Johnson (3C F. Rowsome, DST C. A. Meier
- 1 L. Solander NRR A. C. Thadani, DST
h.
RARE 85-141 August 26, 1985 o
l ATTACHMENT A '
l 1
ESTIMATED PROJECT FINANCIAL STATUS ** Jul-85 (Figures given in thousands.)
FIN A0801 Summary Cost Analysist* Jul-85 Year to Date I Direct Staf f Ef fort (FTE's ) 0.0 1.0 **
II Direct Labor Costs 0.0 43.0 Material & Services 0.0 0.4 ADP Support 0.0 0.0 Subcontracts 0.0 56.5 Travel Expenses 0.0 2.4 Indirect Labor Costs 0.0 42.8 Other ( Recharge ) 0.0 4.5 Other (FY84 Unbilled Costs ) 0.0 0.0 General & Administrative 0.0 22.5 Total Costs 0.0 172.1 Liens 28.9 28.9 Total Costs & Liens 28.9 201.0 Percentage of Available Funds 116.9%
III Funding Status Prior Year FY 1985 Projected FY 1985 Funds FY 1985 Funding Carryover Funding Level Rec'd to Data Bal. Needed 121.9 171.9 50.0 0.0
- Note: These figures are for cost analysis only, and may differ slightly from final billing figures.
- FTE average.
_ m-,--
T RARE 85-141 August 26, 1985 e
i ATTACHMENT B FEE RECOVERY COST STATUS FIN: A0801 TITLE: Review of the Probabilistic Risk Assessment for the Seabrook Nuclear Power Plant PERIOD: July 1985 Docket Costs Facility Name Number Period Cumulative Seabrook 50-443 $ 0.0K $201.0K Common Costs 50-900 $ 0.0 0.0 Total Expenses $ 0.0K
$201.0K T
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[*g UNITED STATES I y e g NUCLEAR REGULATORY COMMISSION r,
.j WASHINGTON, D. C. 20655
%, . . . . + /
ggg Docket Nos.: 50 443 and 50-444 Mr. Robert J. Harrison President & Chief Executive Officer Public Service Company of Nw Fampshire Post Office Box 330 Manchester, New Hampshire 03105
Dear Mr.. Harrison:
SUBJECT:
SEABROOK PROBABILISTIC SAFETY ASSESSMENT (PSA) REVIEW A determination of accident initiation and propagation into core damage and meltdown sequences was reviewed by our contractor, the t.awrence livermore National I.aboratory (l.l.NL) and documented in a draft report. In an April 4, 1985, letter to you, we transmitted this draft 1.1.NI. Seabrook PSA review report.
Anott.er one of our contractors, Brookhaven National 1.aboratory, reviewed and evaluated the Seabrook PSA related to severe accident phenomena, containment response and radiological source tems and provided the NRC their report.
This report is enclosed for your information and use.
Questions or additional information regarding this matter should be directed to the Seabrook Project Manager, Mr. V. Nerses (301-492-8535).
M Victor Nerses, Project Manager PWR Project Directorate #5
~., Division of PWR I.icensing-A
Enclosure:
As stated cc: See next page ,
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4 Mr. Robert J. Harrison Public Service Company of New Hampshire Seabrrok Nuclear Power Station i CC*
Thomas Dignan, Esq. E. Tupper Kinder, Esq.
John A. Ritscher, Esq. G. Dana Bisbee, Esq.
, Ropes and Gray Assistant Attorney General 225 Franklin Street Office of Attorney General Boston, Massachusetts 02110 208 State House Annex Concord New Hampshire Mr. Bruce Beckley, Project Manager Public Service Company of New Hampshire Resident Inspector Post Office Box 330 Seabrook Nuclear Power Station i Manchester; New Hamps, hire 03105 c/o U.S. Nuclear Regulatory Comm.
Post Office Box 700 Dr. Murray Tye, President Seabrook, New Hampshire 03874 Sun Valley Association 209 Sumer Street Mr. John DeVincentis, Director Hacerhill, Massachusetts 08139 Engineering and t.icensing Yankee Atomic Electric Company i Robsrt A. Backus, Esq. 1671 Worchester Road O'Neil, Backus and Spielman Framingham, Massachusetts 01701
- 116 1.owell Street <
Manchester, New Hampshire 03105 Mr. A.M. Ebner, Project Manager United Engineers & Constructors Mr. Phillip Ahrens, Esq. 30 South 17th Street i Assistant Attorney General Post Office Box 8223 l State House, Station #6 Philadelphia, Pennsylvania 19101 Augusta, Maine 04333 William S. Jordan, III Mr. Warren Hall Diane Curran Public Service Company of Harmon, Weiss & Jordan New Hampshire 20001 S. street, NW Post Office Box 300 Suite 430 Seabrook, New Hampshire 03874 Washington, D.C. 20009 Seacoast Anti-Pollution f.eague Jo Ann Shotwell, Esq.
- Ms. Jane Doughty Office of the Assistant Attorney
. !i Market Street General Portsmouth, New Hamoshire 03801 Environmental Protection Division One Ashburton Place Ms. Diana P. Randall Boston, Massachusetts 02108 70 Collins Street Seabrook, New Hampshire 03874 D. Pierre G. Cameron, Jr. , Esq.
General Counsel Richard Hampe, Esq. Public Service Company of New New Hampshire Civil Defense Agency Hampshire 107 Pleasant Street Post Office Box 330 Concord, New Hampshire 03301 Manchester, New Hampshire 03105 Regional Administrator, Region I U.S. Nuclear Regulatory Commission 631 Park Avenue King of Prussia, Pennsylvania 19406
a 9
Public Service Company of Seabrook Nuclear Power Station New Hampshire cc:
Mr. Calvin A. Canney, City Manager Mr. Alfred V. Sargent, City Hall Chaiman 126 Daniel Street Board of Selectmen Portsmouth, New Hampshire 03801 Town of Salisbury, MA 01950 Ms. I.etty Hett Senator Gordon J. Humphrey Town of Brentwood ATTN: Tom Burack RF0 Dalton Road U. S. Senate Brentwood, New Hampshire Washington, D.C. 20510 Ms. Roberta C. Pevear' Town of Hampton Falls, New Hampshire Drinkwater Road Hampton Falls, New Hampshire 03844 Ms. Sandra Gavutis Mr. Owen B. Durgin. Chaiman Town of Kensington, New Hampshire Durham Board of Selectmen RDF 1 Town of Durham .
East Kingston, New Hampshire 03827 Durham, New Hampshire 03824 Ms. Anne Verga Chaiman, Board of Selectmen Charles Cross, Esq.
Town Hall Shatnes, Mardrigan and South Hampton, New Hampshire 03827 McEaschern 1 25 Maplewood Avenue Mr. Angie Machiros, Chaiman Post Office Box 366 Board of Selectmen Portsmouth, New Hampshire 03801 for the Town of Newbury Newbury, Massachusetts 01950 Mr. Guy Chichester, Chaiman Rye Nuclear Intervention Comittee Ms. Rosemary Cashman, Chairman c/o Rye Town Hall Board of Selectmen 10 Central Road
, Town of Amesbury Rye, New Hampshire 03870 Tpwn Hall Anesbury, Massachusetts 01913 Jane Spector l Federal Energy Regulatory Honorable Richard E. Sullivan Conunission Mayor, City of Newburyport 825 North Capitol Street, N.E. '
Office of the Mayor Room 8105 City Hall Washington, D.C. 20426 Newburyport, Massachusetts 01950 Mr. R. Sweeney Mr. Donald E. Chick, Town Manager New Hamphire Yankee Division Town of Exeter Public Service Company of New 10 Front Street Hampshire Exeter, New Hampshire 03823 7910 Woodmont Avenue Bethesda, Maryland 20814 Mr. William B. Derrickson Senior Vice President Public Service Company of New Hampshire Post Office Box 700, Route 1 Seabrook, New Hampshire 03874
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- ENCLOSURE i
NUREG/CR-4540 BNL/NUREG-51961 A REVIEW 0F THE SEABROOK STATION PROBABILISTIC SAFETY ASSESSMENT: CONTAll0ENT FAILURE MODES AND RADIOLOGICAL SOURCE TERMS M. Khatib-Rahbar, A. K. Agrawal, H. Ludewig and W. T. Pratt
- February 1986 Department of Nuclear Energy Brookhaven National Laboratory Upton, New York 11973 h-f3t7~/N @ f
- I
. .iv.
ACKNOWLEDGMENT The authors are grateful to R. A. Bari (8NL), W. Lyon, J. Rosenthal (USNRC), J. Moody (YAC), and A. Torri (PLG) for their review and many helpful remarks on this manuscript. The authors also wish to thank T. Skelaney for the excellent job in typing the many versions of the manuscript. The work reported herein was conducted under the auspices of the United States Nuclear Regulatory Commission (USNRC), Office of Nuclear Reactor Regulation.
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-111-ABSTRACT A technical review and evaluation of the Seabrook Station Probabilistic Safety Assessment has been performed. It is determined that (1) containment response to severe core melt accidents is judged to be an important factor in mitigating the consequences, (2) failure during the first few hours after '
core melt is also unlikely and the timing of overpressure failure is very long compared to WASH-1400, (3) the point-estimate radiological releases are compa-rable in magnitude to those used in WASH-1400, and (4) the energy of release is somewhat higher than for the previously reviewed studies. '
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. CONTENTS Page ABSTRACT................................................................ iii ACKNOWLEDGMENT.......................................................... iv LIST OF TABLES.......................................................... vi .
LIST OF FIGURES......................................................... vii
- 1. INTR 000CT10N....................................................... I 1.1 Background.................................................... 1 1.2 Obj ect i ve and Sc o pe . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . I 1.3 Organization of the Report.................................... 1
- 2. PLANT DESIGN AND FEATURES IMPORTANT TO SEVERE ACCIDENT ANALYSIS.... 2 2.1 As ses sment of Pl ant Desi gn . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 2.2 Comparison with Other P1 ants.................................. 4
- 3. ASSESS' MENT OF CONTAI NMENT PERFORMANCE . . . . . . . . . . . . . . . . . . . . . . . . . . .7. . .
3.1 Contai nment Analysi s Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 -
3.2 Containment Failure........................................... 8 3.2.1 Background............................................. 8 3.2.2 Design Description..................................... 8 4 3.2.3 Leakage Rate Cal cul ati on. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 3.2.4 Containment Failure Medel.............................. 22 3.2.4.1 Leak-Before-Failure........................... 22 1.2.4.2 Cl assi fi cation of Fail ure. . . . . . . . . . . . . . . . . . . . . 23 3.2.5 Containment Pressure Capacity.......................... 24 3.2.5.1 Concrete Containment.......................... 24 3.2.5.2 Liner......................................... 27 3.2.5.3 Penetrations.................................. 27 3.2.5.4 Containment Failure Probability............... 31 3.2.5.5 Conta i nment Encl osu re . . . . . . . . . . . . . . . . . . . . . . . . . 31 3.3 Definitien of Plant Damage States and Containment
. Re s po n s e Cl a s s e s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31 3.4 Containment Event Tree and Accident Phenomenology............. 33 3.5 Con t a i nme n t Ma t ri x ( C-Ma t ri x ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38 3.6 Rel ea se Catego ry Frequenci es . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44
- 4. ACCIDENT SOURCE TERMS.............................................. 48 l
4.1 Assessment of Severe Accident Source Te rms. .. ... .. .. .. .. . .. . . . 48 4.2 Sou rce Te rm Uncertai nty Analysi s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52 4 4.3 Rec ommend ed So u r c e Te rm s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56
- 5.
SUMMARY
AND CONCLUSIONS............................................ 60 i
- 6. REFERENCES......................................................... 62 l
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.vi.
, LIST OF TABLES Table Title Page 2.1 Comparison of Selected Design Characteri stics. .. . . ..... . ...... ... 5 3.1 Containment Operating and Design Parameters...................... 10 3.2 Contai nment Li ner Penet rati ons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 3.3 Leak Area Estimates fo Mechanical Penetrations................... 29 3.4 Frequencies of Occurrence of the Plant Damage States............. 35 3.5 Contai nment Res ponse Cl as s De fi ni ti ons . . . . . . . . . . . . . . . . . . . . . . . . . . . 36 3.6 Containment Class Mean Frequencies............................... 37 3.7 Accident Phase and Top Events for the Seabrook Containment Event Tree....................................................... 39 3.8 Release Categories Employed in the Seabrook Station Risk Mode 1............................................................ 40 3.9 Simplified Containment Matrix for Seabrock....................... 41 3.10 Frequency of Dominant Release Categories (yr 1).................. 45 3.11 Contribution of Containment Response Classes to the Total Core Melt Frequency.............................................. 46 3.12 Release Category Frequency as a Fraction of Core Melt Frequency........................................................ 47 4.1 Seabrook Point-Estimate Rel ease Categori es . . . . . . . . . . . . . . . . . . . . . . . 49 4.2 Late Overpressurization Failure Comparison....................... 51 a 4.3 Comparison of Releases for Failure to Isolate Containment and t he By-Pa s s Sequence . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53 4.4 Comparison of AB-c and TMLB'-c (BMI-2104) to 3DRr and 37.......... 55 4.5 Comparison of 100I (sum) to V-sequence (Surry).................... 57 4.6 BNL-Sugge sted So u rce Te rms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58 4.7 BNL-Suggested Release Characteristics for Seabrook............... 59 5.1 Comparison of SSPSA and WASH-1400 Containment Failure Frequencies...................................................... 61 1
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.vii.
. LIST OF FIGURES Figure Title Pace 3.1 A schematic representation of source term calculation............ 9 3.2 Equi pment h atch wi th pe rsonnel ai rl ock. . . . . . . . . . . . . . . . . . . . . . . . . . . 12 3.3 Pe rs on n el a i rl o ck . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 3.4 Typi cal hi gh energy pi pi ng penetrati on. . . . . . . . . . . . . . . . . . . . . . . . . . . 15 3.5 Typical moderate energy pi ping penetration. . . . . . . . . . . . . . . . . . . . . . . 16 3.6 Typical. electrical penetration................................... 19 3.7 Typi cal v entil ation pen et rati on. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 3.8 A pictorial representation of leakage categories................. 25 3.9 Estimated radial displacement of containment wa11................ 26 3.10 Estimated contai nme nt f ailu re fractions . . . . . . . . . . . . . . . . . . . . . . . . . . 32 .
3.11 Definitions of the plant damage states used in SPSS............... 34 e
- 1. INTRODUCTION
1.1 Background
Probabilistic Risk Assessment (PRA) studies have been undertaken by a number of utilities (as exemplified by Refs.1-4) and submitted to the Nuclear Regulatory Commission (NRC) in support of various regulatory actions. Brook-haven National Laboratory (BNL) under contract to the NRC, has been involved in reviewing the accident sequence evaluations, core melt phenomenology, con-tainment response and site consequence aspects of the PRAs.
This report presents a review and evaluation of the containment failure modes and the radiological release characteristics of the Seabrook Station Probabilistic Safety Assessment (SSPSA), which was completed by Pickard, Lowe and Garrick, Inc. (PLG) for the Public Service Company of New Hampshire and Yankee Atomic Electric Company in December 1983.5 1.2 Objective and scope The objective of this report is to provide a perspective on severe acci-dent propagation, containment response and failure modes together with radiol-ogical source term characteristics for the Seabrook Station. The determina-tion of accident initiation and propagation into core damage and meltdown -
6 sequences was reviewed by the Lawrence Livermore National Laboratory (LLNL) under contract to the Reliability and Risk Assessment Branch of NRC.
In the present report, principal containment design features are discuss-ed and compared with those of Zion, Indian Point and Millstone-3 designs.
Those portions of the SSPSA related to severe accident phenomena, containment response and radiological source terms are described and evaluated. Numerical adjustments to the SSPSA estimates are documented and justified.
1.3 Organization of the Report At brief review of the Seabrook plant features important to severe acci-dent analysis is presented in Chapter 2 along with comparisons to Zion Indian Point and Millstone-3 plant designs. Chapter 3 contains the assessment of
. containment performance. Specifically, the definition of containment response aclasses and plant damsge states, analytical methods, containment failure mod-
' el, containment event tree and accident phenomenology and the containment ma-trix are reviewed. Chapter 4 addresses the accident source terms together with justifications for adjustment where necessary. The results of this re-view are summarized in Chapter 5.
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- 2. PLANT DESIGN AND FEATURES IMPORTANT TO SEVERE ACCIDENT ANALYSIS In this section, those plant design features that may be important to an assessment of degraded and core melt scenarios and containment analysis are reviewed. These important features are then compared with the Zion, Indian Point and Millstone-3 facilities to identify commonalities for benchmark comparisons.
2.1 Assessment of Plant Design The Seabrook Station is comprised of two nuclear units each having an identical Nuclear Steam Supply System (NSSS) and turbine generator. The units are arranged using a " slide.-along" concept which results in Unit 2 being ar-ranged similar to Unit 1 but moved some 500 feet west. Each unit is a 1150 MWe (3650 MWt), 4-lbop, Westinghouse PWR plant. The turbine-generators are supplied by the General Electric Company and the balance of the plant is de-signed by United Engineers and Constructors.
Each containment completely encloses an NSSS, and is a seismic Category I reinforced concrete structure in the form of a right vertical cylinder with a hemispherical top dome and flat foundation mat built on bedrock. The inside face is lined with a welded carbon steel plate, providing a high degree of
- leak tightness. A protective 4 ft. thick concrete mat, which forf.s the floor ,
of the containment, protects the liner over the foundation mat. The contain-ment structure provides biological shielding for normal and accident condi-tions. The approximate dimensions of the containment are:
Inside diameter 140 ft.
Inside height 219 ft. .
Vertical wall thickness 4 ft. 6 in, and 4 ft. 7 1/2 in.
Dome thickness 3 ft. 6 1/8 in.
Foundation mat thickness 10 ft.
Containment penetrations are provided in the lower portion of the structure, and consist of a personnel lock and an equipment hatch / personnel lock, a fuel transfer tube, electrical, instrumentation, and ventilation penetrations.
Each containment enclosure (also known as secondary containment) sur-l' rounds a containment and is designed in a similar configuration as a vertical right cylindrical seismic Category I, reinforced concrete structure with dome and ring base. The approximate dimensions of the structure are: inside diam-eter,158 ft; vertical wall thickness, varies from 1 ft, 3 in. to 3 ft; and dome thickness, 1 ft, 3 in.
The containment enclosure and enclosure ventilation system is designed to collect leakage from the containment structure and to discharge the leakage to the filtration system of containment.* To accomplish this, the space between the containment enclosure and the containment structure, as well as the pene-tration and safeguards pump areas, are maintained at a negative pressure fol-lowing a design basis accident by fans which take suction from the containment
' Leakage via connections which pass into the steam and feedline penetrations are expected to bypass this system.
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enclosure and exhaust to atmosphere through charcoal filters. To ensure air tightness for the negative pressure, leakage through all joints and penetra-tions has been minimized.
A containment spray system is utilized for post accident containment heat removal. The containment spray system is designed to spray water containing boron and sodium hydroxide into the containment atmosphere after a major acci-dent to cool it and remove iodine. The pumps initially take suction from the refueling water storage tank and deliver water to the containment atmosphere through the spray headers located in the containment dome. After a prescribed amount of water is removed from the tank, the pump suction is transferred to the containment sump, and cooling is continued by recirculating sump water through the spray heat exchangars and back through the spray headers.
The spray is actuated by a containment spray ectuation signal which is generated at a designated containment pressure. The system is completely re-dundant and is designed to withstand any single failure.
The containment isolation system establishes and/or maintains isolation of the containment from the outside environment in order to prevent the re-lease of f tssion products. Automatic trip isolation signals actuate the ap-propriate valves to a closed position whenever automatic safety injection oc-
. curs or high containment pressure is experienced. Low capacity thermal elec-
- tric hydrogen recombiners are provided.
The emergency core cooling system (ECCS) injects borated water into the reactor coolant system following accidents to limit core damage, metal-water reactions and fission product release, and to assure adequate shutdown mar-
. gin. The ECCS also provides continuous long-term post-accident cooling of the core by recirculating borated water between the. containment sump and the reac-tor Core.
The ECCS consists of two centrifugal charging pumps, two high pressure i safety injection pumps, two residual heat removal pumps and heat exchangers, i
and four safety injection accumulators. The system is completely redundant, and will assure flow to the core in the event of any single failure.
The control building contains the building services necessary for contin-
. uous occupancy of the control room complex by operating personnel during all
.' operating conditions. These building services include: HVAC services, air purification and iodine removal, fresh air intakes, fire protection, emergency breathing apparatus, communications and meteorological equipment, lighting, and housekeeping facilities.
Engineered Safety Feature (ESF) filter systems required to perform a safety-related function following a design basis accident are discussed below:
- a. The containment enclosure exhaust filter system for each unit col-lects, filters and discharges any containment leakage. The system is not normally in operation, but in the event of an accident, it is placed in operation and keeps the containment enclosure and the building volumes associated with the penetration tunnel and the ESF equipment cubicles under negative pressure to ensure all leakage fron
I the containment structure is collected and filtered before discharge I to the plant vent.
- b. One of two redundant charcoal filter exhaust trains is placed in op- l eration in the fuel storage building whenever irradiated fuel not in '
a cask is being handled. These filter units together with dampers and controls will maintain the building at a negative pressure.
The emergency feedwater system supplies domineralized water from the con-densate water storage tank to the four steam generators upon loss of normal ;
feedwater flow to remove heat from the reactor coolant system. Operation of the system will continue until the reactor coolant system pressure is reduced to a value at which the residual heat .emoval system can be operated. The combination of one turbine-driven and one motor-driven emergency feedwater pump provides a diversity of power sources to assure delivery of condensate l i under emergency conditions.
The two units of the facility are interconnected to off-site power via three 345 kilovolt lines of the transmission system for the New England The normal preferred source of power for each unit is its own main states.
turbine generator. The redundant safety feature buses of each unit are power-ed by two unit auxiliary transformers. A highly reliable generator breaker is '
provided to isolate the generator from the unit auxiliary transformers in the event of a generator trip, thereby obviating the need for a bus transfer upon loss of turbine generator power. In the event that the unit auxiliary trans-formers are not available, the redundant safety feature buses of each unit are powered by two . reserve auxiliary transformers. Upon loss of off-site pcwer, each unit is supplied with adequate power by either of two fast-starting, diesel-engine generators. Either diesel-engine generator and its associated safety feature bus is capable of providing adequate power for a safe shutdown under accident conditions with a concurrent loss of off-site power. A con-stant supply of power to vital instruments and controls of each unit is assur-ed through the redundant 125 volt direct current buses and their associated battery banks, battery chargers and inverters.
2.2 Comparison with Other Plants Table 2.1 sets forth- the design characteristics of the Zion, Indian
. Point-2, and Millstone-3 facilities as they compare to the Seabrook station.
It 'is seen that the containment characteristics are quite similar with the exception of the containment operating pressure for Millstone-3 (subatmos-pheric design), the use of fan coolers in Zion and Indian Point for post-acci-dent containment cooling, lower reactor cavity configurations, and the chemi-cal composition of the concrete mix. The primary system designs are nearly I identical between the four units.
The Seabrook containment building basemat and the internal concrete structures are composed of basaltic-based concrete. As concrete is heated, water vapor and other gases are released. The initial gas consists largely of carbon dioxide, the quantity of which depends on the amount of calcium carbon-ate in the concrete mix. Limestone concrete can contain up to 80% calcium carbonate by weight, which could yield up to 53 lb of carbon dioxide per cubic l foot of concrete. However, basaltic-based concrete contains very little cal-cium carbonate (3.43 w% for Seabrook) and would not release a substantial
.s.
Table 2.1 Comparison of Selected Design Characteristics Zlon IndienPgint Seebroom Design Permeters MIllstony Unit 1 Unit 2 Unit 3 e Unit 1,2 neoctor Power leerttil 3,250 3,030 3,411 3,650 containment b ilding:
3 0 Free volume (ft ) 2.73 x 10 2.61 x 10 2.3 x 10 2.7 x 100 Design Pressure (psla) 62 62 59.7 67.7 Initial Pressure (psla) 15 14.7 12.7/9.1 15.2 teltlal Temperature (*F) 120 120 120/80 120 Primary syste r I
water Volume (ft l 12,710 11,347 11,671 13,I40 3
3+eam volume (ft ) 720 720 ? 2,012 Mass of 'J02 In Core (Ib) 216,600 216,600 222,739 222,739 Itss of Steel In C;re tib) 21,000 20,407 7 19,000 Moss of Zr In Core (lb) 44,500 44,600 45,296 45,234 Mass of enttom Heed (Ib) 87,000 70,130 87,000 87,000 Botton Hood Dimeter (ft) 14.4 14.7 14.4 14.4 Bottom Hood 7hickness (ft) 0.45 0.44 0.45 0.45 stese Generator Syste n Inventory per Generator (Ibe) 77,000 82,000 113,600 112,000 Conte.Innent Buildino coolers:
Spreys yes yes yes yes Fans (with safety function) yes yes no no Accoulstor Tanks:
, Total Moss of water (Ib) 200,000 173,000 348,000 213,000 s initial Pressure (psla) 665 665 600 615 l ,' Ta perature t'F) 150 150 80 100-150 Refueline water Storeas Tank:
0 Total Mass of water (Ib) 2.89 x 10 2.89 x 10 10 2.89 x 100
- Temperature (*F) 100 120 50 86 Reetter Cavity:
Configuration met Wet Dry Dry / wet l Concrete Meteelsi Limestone Basaltic Besattic 8asettle l
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- Minimum (Moulmue Qpecity = 3.9 4 10 l b) l l
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amount of carbon dioxide.5 Thus, pressurization of the containment as a 1 result of corium/ concrete interactions would be expected to take a very long '
time. In Table 2.1 the reactor cavity configuration is described as either )
wet or dry. Wet means that the reactor cavity configuration is such that for '
a wide range of accident sequences the cavity would be flooded. This means that if the reactor core melts down and falls into the cavity it would contact significant quantities of water. This has important implications in terms of ultimately cooling the core debris and perhaps preventing core / concrete inter-actions. It is also important in terms of containment pressurization because boiling water by the core debris is a much more efficient way of building up pressure in containment than by core / concrete interactions. Dry in Table 2.1 implies that the reactor cavity configuration is such that for a wide range of i accident sequences the cavity would not be flooded thus extensive core / con-l crete interactions would be expected.
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- 3. ASSESSMENT OF CONTAINMENT PERFORMANCE In this chapter, the review of containment response to severe accider.ts is described. Analytical techniques used to analyze core meltdown phenomena and containment response are reviewed, the containment failure model is assessed, plant damage states and containment failure modes are evaluated.
Parallels between this study and other PRAs are made. Finally, the relevance ,
and validity of the conclusions is addressed.
3.1 Containment Analysis Methods A brief description of the computer codes used to perform the transient degraded, core meltdown and containment response analyses is provided in this section.
The MARCH 8 computer code is used to model the core and primary system transient behavior and to obtain mass and energy releases from the primary system until reactor vessel failure. These mass and energy releases are then used :as input to the other computer codes for analysis of containment re-sponse.
For sequences in which the reactor coolant system remains at an elevatej pressure until the vessel failure (" time-phased dispersal"), the MODMESH ~
computer code is used. This code calculates the steam and hydrogen blowdown from the reactor vessel using an isothermal ideal gas model. The water level boil-off from the reactor cavity floor is modeled using a saturated critical ,
heat flux correlation. Additionally, the accumulator discharge following de-pressurization caused by the vessel failure is also considered.
A modified version of the CORCON8 code is used to replace the INTERe sub-routine of the MARCH code. CORCON models the core-concrete interaction after the occurrence of dryout in the reactor cavity. The mass and energy releases from the core-concrete interaction are transferred to the MODMESH code for proper seqgencing and integration into the overall mass and energy input to COC0 CLASS 9 code.
C0C0 CLASS 9, a modified version of the Westinghouse C0C0 computer code utilizes the mass and energy inputs to the containment as computed by MARCH to
- model the containment building pressurization and hydrogen combustion phenom-l
! 3ena. This code replaces the MACE subroutine of the MARCH code. The code also
,' models heat transfer to the containment structures and capability for contain-ment heat removal through containment sprays and sump recirculation.
Fission product transport and consequence calculations are performed using the CORRAL-II and the PLG proprietary CRACIT5 computer codes, respec-tively.
The analytical methods used to cerry out the core and containment thermal hydraulics, and fission product transport calculations are identical to those used for MPSS-3.7
3.2 Containment Failure 3.2.1 Background In order to assess the risk of the Seabrook-1 plant, radiological source terms have to be calculated. Many steps are involved in such calculations.
These are schematically shown in Fig. 3.1. The mode and time of containment failure directly impact on the radioactivity release categories. These, when I coupled with the status of reactor cavity and the spray system, determine the source terns. This section deals with the mode and time of containment fail-ure.
3.2.2 Design Description The primary containment of the Seabrook plant is a seismic Category I re-inforced concrete dry structure. It consists of an upright cylinder topped with a hemispherical ~ dome. The inside diameter of the cylinder is 140 feet and the inside height from the top of the basemat to the apex of the dome is approximately 219 feet. The cylindrical wall is 4'6" thick above elevation 5' and 4'7-1/2 " thick below that evaluation. The dome is 3'6-1/8" thick and 69'11-7/8" in radius. The cylinder is thickened to provide room for addition-al reinforcing steel around the openings for the equipment hatch and the per-sognelairlock. The net free volume of the containment is approximately 2.7 x 3
10 ft .
The inside of the containment is lined with a welded steel liner. The liner plate in the cylinder is 3/8" thick in all areas except penetration and the junction of the basemat and cylinder where it is 3/4" thick. This liner serves as a leak-tight membrane. Welds that are embedded in the concrete and 'i not readily accessible are covered by a leak chase system which permits leak testing of these welds throughout the life of the plant. The dome liner is 1/2" thick and flush with the outside face of the cylindrical liner. The op-erating and the design parameters of containment are noted in Table 3.1.
The containment building is surrounded by an enclosure. The containment enclosure is a reinforced concrete cylindrical structure with a hemispherical dome. The inside diameter of the cylinder is 158 feet. The vertical wall
- varies in thickness from 36 inches to 15 inches; the dome is 15 inches thick.
. The inside of the dome is 5'6" above the top of the containment dome. Located f.at the outside of the enclosure building is the plant vent stack, consisting
. ,' of a light steel frame with steel plates varying in cross-section. The stack carries exhaust air from various buildings.
The containment enclosure is designed to control any leakage from the containment structure. To accomplish this, the space between the containment and the enclosure building (approximately 4'6" wide) is maintained at a slight negative pressure (-0.25" water gauge) during accident conditions by fans whice take suction from the containment enclosure and exhaust to atmosphere through charcoal filters.
There are a number of containment penetrations which are steel components that resist pressure. These penetrations are not backed by structural con-crete and include the following:
9
- , - - - - - - , ,,.,-,,,-m
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CONTAINMENT TIME OF
, FAILURE FAILURE MODE WET OR DRY RELEASE SPRAY ,
REACTOR CAVITY CATEGORY SYSTEM SOURCE TERM l
Figure 3.1 A schematic representation of source term calculation.
9
10- ,
Table 3.1 Containment Operating and Design Parameters Parameter Value Normal Operation ,
Pressure , psig 0.5 Inside Temperature F 120 Outside Temperature . F 90 Relative Humidity , % 45 Service Water Temperature F 80 Refueling Water Temperature F 86 Spray Water Temperature , F 88 Containment Enclosure Pressure , inches w.g. -0.25 Design Conditions Pressure , psig 52.0.
Temperature , F ,
296
- Free Volume , ft 3 2.7x108 Leak Rate , 1 mass / day 0.2 Containment Enclosure Pressure , psig -3.5 i
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- 1. Equipment hatch, ,
- 2. Personnel air lock, '
- 3. Piping genetrations,
. 4 Electrical penetrations,
- 5. Fuel transfer tube assembly,
- 6. Instrumentation penetrations, and
- 7. Ventilation penetrations.
These components penetrate the containment and containment enclosure shells to provide access, anchor piping, or furnish some other operational requirement.
All penetrations are anchored to sleeves (or to barrels) which are embedded in the concrete containment wall.
Equipment Hatch The equipment hatch (Fig. 3.2) consists of the barrel, the spherical dished cover plate with flange, and the air lock mounting sleeve. The center-line of the hatch is located at elevation 37'1/2" and an azimuth of 150". The hatch opening has an inside diameter of 27'5". A sleeve for a personnel air lock, the inside diameter of which is 9'10", is provided at centerline eleva-tion 30'6". Thicknesses of the primary components are as follows:
Component Thickness (inches) ,
9 Barrel 3 1/2 Spherical 1 3/8 Flange 5 3/8 Air lock mounting 1 1/2 sleeve The equipment hatch cover is fitted with two seals that enclose a space which can be pressurized to 52.0 psig. The flange of the cover plate is at-tached to the hatch barrel with 32 swing bolts,1-3/8 inch in diameter. The barrel, which is also the sleeve for the equipment hatch, is embedded in the shell of the concrete containment. The equipment hatch cover can be lifted to
! clear the opening.
Inserted into the mounting sleeve through the equipment hatch cover is a .
. personnel air lock consisting of two air lock doors, two air lock bulkheads,
'and the air lock barrel. Significant dimensions of the air lock are as
'follows:
parameter Dimension Inside Diameter of Barrel 9'6" Barrel Thickness 1/2" Door Opening 6'8" x 3'6" Door Thickness 3/4" Bulkhead Thickness 1-1/8" Each door is locked by a set of six latch pin assemblies, and is designed to withstand the design pressure from inside the containment. To resist the test pressure, each door is fitted with a set of cast clamps. The doors are hinged and both swing into the containment. Each door is fitted with two seals that 1
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are located such that the area between doors can be pressurized to 52.0 psig.
The doors are mechanistically interlocked so that only one door can be opened at a time. The capability exists for bypassing this interlock to equalize the pressure by use of special tools. The doors may be operated mechanically.
Personnel Air Lock The personnel air lock (Fig. 3.3) consists of the air lock doors (2) and ,
the lock barrel. The barrel, which is also the sleeve for the personnel air lock, is imbedded in the shell of the concrete containment. The centerline of )
the barrel is located at elevation 29'6" and an azimuth of 315*. Significant dimensions are as follows:
Parameter Dimensions Clear Opening 7'0" 0.D. of F1ange on Door 7' 9 1/8" Barrel Thickness 5/8"
, Cover Thickness 5/8" The air lock barrel has a door on each end, each of which is designed to withstand the design pressure from inside the containment. The doors are hinged and swing away from the air lock barrel. Each door is fitted with two seals that are located such that the area between doors can be pressurized to
- 52.0 psi g. The locking device for the doors is a rotating, third ring, breach-type mechanism. These doors are also mechanically interlocked so that only one door can be opened at a time. The capability exists for bypassing this interlock and relieving the internal pressure by use of special tools.
The doors may be operated mechanically, i Piping Penetrations There are two types of piping penetrations: moderate energy and high energy. Moderate energy piping penetrations are used for process pipes in which both the pressure is less than or equal to 275 psi, and the temperature of the process fluid is less than or equal to 200*F. High energy piping pene-trations are used for that piping in which the pressure or temperature exceeds these values.
[
- l S High energy piping penetrations (Fig. 3.4) consist of a section of pro-cess pipe with an integrally-forged fluid head, a containment penetration sleeve and, where a pipe whip restraint is not provided, a penetration sliding support inside the containment. The sliding support provides shear restraint while permitting relative motion between the pipe and the support. The annu-lar space between the process pipe and the sleeve is completely filled with fiberglass thermal insulation. The pipe and the fluid head, are classified as ASME III Safety Class 2 (NC), whereas the sleeve is classified as part of the concrete containment, ASME III (CC). The sliding support inside the contain-j ment is classified as an ASME Safety Class 2 component support (NF).
Moderate energy piping penetrations (Fig. 3.5) consist of one or more process pipes, the containment penetration sleeve, and a flat circular end-plate. The pipe is classified as ASME III Safety' Class 2 (NC). The sleeve is classified as ASME III Div. 2 (CC). The end-plate is classified as Class MC.
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Table 3.2 gives a list of the containment piping penetrations. Included in this table is the penetration size. All of these piping penetrations are in the lower portion of the structure.
Electrical and Instrumentation Penetrations Electrical penetrations (Fig. 3.6) consist of a stainless steel header plate with an attached terminal box, electrical modules which are clamped to the header plate, and a carbon steel weld ring which is welded to the header plate and to the sleeve. The metallic pressure resisting parts, the sleeve, i
stainless steel header plate and carbon steel weld ring were designed as ASME III Safety Class MC components (NE); that portion of the sleeve which is backed by concrete was designed as part of the concrete containment. ASME III (CC).
Double silicone _ and Hypalon 0-rings provide a seal with a cavity for leakage monitoring between the header plate and the modules. The header plate is provided with a hole on the outside of the containment to allow for pressurization of the penetration assembly for leakage monitoring.
There are a total of 64 electrical penetrations out of which 14 are spare and 8 are unused. All of these electrical penetrations are below the grade.
Instrumentation penetrations are of two types -- electrical and fluid.
- The electrical type is similar in construction to the other electrical pene-trations. The fluid penetrations are similar in construction to the moderate j energy piping penetrations.
Fuel Transfer Tube Assembly The fuel transfer tube assembly consists of the fuel transfer tube, the penetration sleeve, the fixed saddle on the reactor side, and the sliding sad-die in the fuel storage building. The fuel transfer tube centerline is at elevation (-)9'4-1/4" and it has approximately 20" inner diameter. The fuel transfer tObe wall penetration sleeve, which is embedded in the concrete, has an inside diameter of about 25".
Ventilation Penetrations l
- There are two types of ventilation penetrations -- the containment air
' purge penetrations (HVAC-1 and HVAC-2) and the containment on-line penetra-tions (X-16 and X-18). The containment air purge penetrations (Fig. 3.7) each consist of a pipe sleeve (a rolled and welded pipe section, 36" outer diameter by 1/2" wall thickness) which is flanged at each end with 36" weld neck flanges and, attached to these flanges, the inner and outer isolation valves.
Together with the pipe, these valves form a part of the containment pressure bounda ry. The valves are 36" diameter butterfly valves with fail-safe pneu-matic operators. The weld between the pipe and the containment liner is equipped with a leak chase for pressure testing.
The containment on-line purge penetrations each consist of a pipe sleeve (a rolled and welded pioe section, 8" o.d. by 1/2" wall thickness). A short section of pipe with a nipple is welded to the sleeve on the outside of the containment, and a 3/4" valve and test connection is attached to it. The l
i Table 3.2 Containment Liner Penetrations Penetration Penetration Numbers Service Size X-1 to X-4 Main steam line 30" X-5 to X-8 Main feedwater 18" X-9, X-10 RHR pump suction 12" X-11 to X-13 RHR to safety injection 8" X-14 to X-15 Containment building spray 8" X-16, X-18 , Containment on-line purge 8" X-17 Hydrogenated vent header 2" X-20 to X-23 CCW supply and return 12" X-24 to X-27 Safety injection 4" X-28 to X-31 CVCS to pump seal injection 2" X-32, X-34 Drain line 3",2" X-33, X-37 CVCS 3" X-35, X-36, X-40 RCS test / sample control 1" or smaller X-52, X-71, X-72 X-38 Combustible gas control 10" X-39 Spent fuel pool cooling 2" X-43, X-47, X-50 Instrumentation lines <1" X-57 X-60, X-61 From containment recirculation sump 16" X-62 Fuel transfer tube 20" X-63 to X-66 Steam generator blowdown 3" X-67 Service air 2" HVAC-1,2 Containment purge supply / exhaust lines 36" l X-19, X-41, X-42 X-44 to X-46, X-48 Spare ?
X-49, X-51, X-58 X-59, X-68 to X-70
OUTBOARD CONTAINMENT WALL INBOARD STAINLESS STEEL HEADER PLATE I LINER PLATE THERMAL INSULATING Y I GASKET 1-BOX MOUNTING RING CANISTER n, o i il !
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ends of this resulting assembly are welded to 8" weld neck flanges which are through-bolted to the inner and outer isolation valves. These valves are 8" diameter butterfly valves having fail-safe pneumatic operators. The weld be-tween the pipe sleeve and the containment liner is equipped with a leak chase for pressure testing. These on-line purge penetrations are very similar to those for 36" lines shown earlier.
3.2.3 Leakage Rate Calculation Under severe accident conditions the pressure inside the containment builds up in the range of 75 to 200 psi. At these pressures, any leakage through the containment holes will essentially be choked. The leakage under choked flow condition is given as (Ref.10):
k+1 2 k-1 W= k(kT1) A dP (1) where W = discharge rate (kg/s),
A = leak area (m2),
P = absolute pressure (N/m2), .
p = mixture density (kg/m3), and k = ratio of specific heat at constant pressure to that at constant volume.
For air and water vapor mixture, k = 1.3. If the mixture density is expressed by perfect gas law p=h , (2) where R = gas constant, and T = the absolute temperature.
, Then Eq.(1) becomes
,' k+1 2 k-1 P W= k(UT) A y[llT (3)
The mass of mixture can be written as M = Vp or, M=h (4) where V is the free mixture volume in the containment. Equations (3) and (4) can be combined to get the leakage rate, in terms of mass fraction, as 1
22- .
k+1 h = k( h ) h AT A (5)
Note that the leakage rate, when expressed in terms of mass fraction, depends only on the leakage area. 1 For Seabrook-1, using V = 2.704x106 ft3 and T = 296 F. Eq.(5) gives f Leakage Rate = 0.721 Ain w/o per hour (6) where Ain is the leakage area in in 2 . Alternately, Leakage, Rate = 17.3 Ain w/o per day. (7) t The essentially intact containment leakage of 0.2 w/o per day, thus, corre-sponds to an equivalent leakage area of 0.012 in2 (or, an equivalent hole of 1/8-in diameter). A leakage area of 4 to 10 in2 would correspond to the leak-age rate of 2.9 to _7.2 w/o per hour. In other words, it will take about 14 i
hours to leak the entire content to the environment through a 10-in2 hole.
3.2.4 Containment Failure Model c
3.2.4.1 Leak-Before-Failure During accident sequences involving core damage, the containment struc-ture will be exposed to pressures and temperatures beyond those used in the design basis accident (DBA). Response of the containment building to these severe conditions is evaluated in SSPSA by employing, for the first time, a leak-before-failure model. In this model allowance is made for continuous leakage from the containment to the surroundings. This mode of containment failure is termed local failure. The containment leakage can occur at many locations and discontinuities such as mechanical and electrical penetrations, personnel lock, equipment hatch, fuel transfer tube, welds, and in between the liner and concrete. Depending upon the size of leakage area and the accident sequence, local failures may gradually relieve pressure, thereby gross con-
, tainment failure may be averted.
l' The leak-before-failure model is _a realistic one. The extent of leakage and the health consequences must, however, be carefully studied. In order to explain this issue it is observed that traditionally probabilistic risk as-sessment is made by using what is termed a threshold model. In the threshold model, the containment is considered intact until the internal loading equals l
or exceeds a pressure threshold (which may also be temperature dependent), at which it is deemed to have suffered a failure (gross). If the internal load-ing is below this threshold value, the containment is considered intact and hence the risk is quite low. In the leak-before-failure model, the release of activity, which is considerably small compared with that for the gross failure mode, must be considered in health consequences. However, such leakages can potentially prevent the internal pressure from approaching the threshold value and thus a catastrophic or gross failure may be avoided.
l_ __
23-l 3.2.4.2 Classification of Failure
~
The SSPSA report has classified containment failures in three categories:
. Containment Failure Category A. Includes containment' failures that develop a small leak that is substantially larger than the leak ac-ceptable from an intact containment, but not large enough to arrest the pressure rise in the containment. Category A failures thus cause an early increase in the rate of leakage of radionuclides over the de-sign basis leak rate but pressurization of the containment continues until either a category B or C containment failure occurs.
The intact containment is defined as the one in which leakage is lim-ited by the Technical Specification value. For Seabrook-1, this value is 0.2 w/o per day at the calculated peak accident pressure of approx-imately 47 psig. Note that the SSPSA study has used 0.1 volume per-cent per day for this leakage, although prior to the most recent amendment dated August 1984, the FSAR has cited both 0.1 volume per-cent and 0.1 w/o per day. The 10CFR50, Appendix J mandates the allow-able leakage to be quoted as w/o per day. The higher value noted here is based on Amendment 53, August 1984.*
. Containment Failure Category B. Includes failure modes that develop a "
large enough leak area so that the pressure in the containment no longer increases. The time during which a substantial fraction of the radionuclide source term is released is longer than approximately 1 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. Category B failures include self-regulating failure modes where the leak area is initially small but increases with pressure so
. that it becomes sufficient to terminate the pressure rise before a category C containment failure occurs.
The definition of " substantial" fraction is unclear.
- . Containment Failure Category C. Includes those containment failure modes that develop a large leak area. A large fraction of the total radionuclide source term is released over a period of less than 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. All gross failure modes are included in category C.
- Mathematically, these three failure categories can be expressed in terms
! l*of leakage areas as follows:
ADBA < AA i ANP Type A ANP < AB 1 AP Type B (8)
AC > AP Type C ADBA = leakage area corresponding to the technical specification limit for containment leakage,
- There appears to be substantial update / changes in the Engineered Safety Features flow diagram, including arrangements of motor operated valves and bypass lines, which may substantially change the frequency of events. BNL, however, is not reviewing this part of SSPSA.
'l
ANP = leakage area not large enough to arrest pressurization, and Ap = leakage area sufficient to release 100 w/o in one hour.
The leakage area required to release a substantial fraction of the radio-nuclide source term in approximately an hour can be computed using Eq. (6).
Assuming one-hundred percent turnover as " substantial" fraction in one hour, Eq. (6) gives the required leakage area to be equal to 138 in 2 or about 1 l ft 2 . If one were to use 50% turnover as " substantial" in an hour, the corre-sponding leakage area would be 60 in 2 or about 0.5 ft2 . Forthepgrposeof defining an upper bound for type B, it seems justified to use 1 ft , hence, leakage area in excess of this value should be considered type C failure Al-though this estimate of the leak area is a factor of two too high from the value stattd in SSPSA, it is not significant in the risk evaluation since failure categorization is somewhat arbitrary.
The leakage area required to arrest containment pressurization is in the range of 4 to 10 square inches. A leak area of about 6 square inches will re-sult in the release of about 100 w/o of activity in a dag (see Eq. 7). The upper bound leak area for Type A failure is taken as 4 in . This corresponds to release of the radioactive source term (100% turnover) 2 in about 2
36 hours4.166667e-4 days <br />0.01 hours <br />5.952381e-5 weeks <br />1.3698e-5 months <br />.
The Category B leak area is, thus, in the range of 4 in to 1 ft . Figure 3.8 is a pictorial representation of these leakage categories.
3.2.5 Containment Pressure Capacity 3.2.5.1 Concrete Containment The Seabrook PSA has examined failure modes for the containment structure itself, the steel liner, all penetrations, equipment and personnel lock hatch-es, and the secondary containment. The containment structure includes the cylindrical wall, the hemispherical dome, the base slab and the base slab and containment wall junction. The most critical membrane tension was found to occur in the cylinder in the hoop direction. The median pressure which causes yield of both the liner steel and the reinforcing bars was. found to be approx-
, imately 157 psi, with a coefficient of variation of 0.084 The ultimate hoop load in cylinder is 216 psig. The containment wall is, thus, assumed to fail at this pressure. At pressures beyond this, very large irreversthie defoma-tions occur which will cause cracks in the reinforced concrete but the loss of integrity of the pressure boundary may not occur until the liner tears. The compiled radial deformations of the containment wall are shown in Figure 3.9.
Note that the radial strain at the expected failure pressure of 216 psi is 4.71 (Ar/r).
The hemispherical dome was calculated to yield at a slightly higher pres-sure (163 psig). The failure pressure is predicted at 223 psig.
The median pressure for flexural failure of the base slab is 400 psig, with a logarithmic standard deviation of 0.25. However, the shear mode of failure is more restrictive. For this mode, the median failure pressure is estimated in SSPSA as 323 psig, with a logarithmic standard deviation of 0.23. Although the uncertainty for failure of the base slab is large, the
+
25-3
'1 s
'N
(_
g s, .
5 5
a E
I
- A :: B ; : C
\ l l l l l O.01 0.1 1 10 100 1000 2
EQUIVALENT LEAKAGE AREA, in
(
Figure 3.8 A pictorial representation of leakage categories, i
A,
~
i i
., c 1 I I I I 157 (yield) 180 200 Fatiure 216 Psi 9 Psig psig Pressure + psig _
20 l
i i 1
I : -
10 -
7 w l j j
- g Top of the Reactor Vessel -
p / +
/ k g 0 -
/ -
/ -
i sa d / -
/
/ ./
l -10 - / ,/ -
/ /
l /
W
./-
/
-20 -
/
/
/.#. _
- / . .
l -30 I I I i
! 10 20 30 40 r
! RADIAL DISPLACEMENT OF THE CONTAINMENT WALL AWAY FROM THE BASE (INCHES) -
l l Figure 3.9 Estimated radial displacement of containment wall.
1
r, probability of failure is small because th'e median capacities are high. Thus, failure of the base slab is not considered to be a critical failure mode and an estination of leak areas was, therefore, not considered for this mode of failure.
Secondary strasses in the cylindrical portion of the containment occur at discontinuity such as at 4the base slab contairnent wall junction, at the springline, and where the amount of reinforcing, changes. The flexural yield at the base of the cylinder occurs at 175, psi .- At higher pressures, 3 plastic hinge forms with considerable cracking of-the concrete. These cracks, how-ever, are $nall enough so as not to threaten the integrity of the liner. The loss of integrity of the liner is not expected until a median pressure of 408 psi is reached. Thus, the failure of the base slab and containment well junc-j tion is not limiting.
In summary, the . containment wall is expected to undergo significant de-formation (=4.7% ar/r) prior to its failure at 216 psig. At this pressure, Type C (i.e., gross) failure occurs.
3.2.5.2 Liner The elongation capacity of the steel liner is computed by neglecting the friction forces between the liner and the concrete. The possibility that the ,
liner stresses and strains could be different between two different pairs of tees was, however, considered. The SSPSA computed an elongation of 8.1 per-cent under unaxial conditions, or an elongation of 4.7 percent under plane strain conditions can be achieved without fracture. This would ensure integ-rity of the liner until fracture of the reinforcing bars. Additionally, the leakage of the containment at penetrations is considered likely before hcop failure of the liner occurs.
3.2.5.3 Penetrations At all major penetrations, the containment wall is thickened and addi-tional reinforcement is provided to resist stress concentrations. None of the meridional or hoop reinforcing bars are terminated at penetrations. Instead, they are continued around the penetrations, thus ensuring that excess hoop and meridional capacity is available. Table 3.2 lists all piping penetrations.
,' As the containment pressure increases beyond its yield value (157 psi),
large radial deformations begin to occur. This induces stresses in the pipes by relative displacements between the containment wall and the pipe whip re-straints. Therefore, the most critical penetrations are the areas where the pipe is supported close to the penetration. Also, stronger and stiffer pipes develop higher forces at the penetrations for a given relative displacement.
The SSPSA study selected the following penetrations for investigation as being among the lines most likely to fail:
Penetration X-23 12" schedule 40 carbon steel (also X-20 to X-22 by similarity)
Penetration X-26 4" schedule 160, stainless steel (also X-24, X-25, X-27)
J
Penetration X-71 1" - muMiple pipe penetration .
(also X-72 and possibly others) penetration X-8 18" main feedwater schedule 100 (also X-5 to X-7) carbon steel Fuel Transfer Tube Convoluted Bellows The probability of failure at these penetrations was computed by (a) establishing a pressure-displacement relation, (b) estimating the failure probability as a function of radial displacement and then (c) combining the two. The radial displacements for the containment wall were shown earlier (Fig. 3.9). The vertical displacement due to meridional strains is small (less than 3 inches) and hence its impact on the penetrations was ignored.
Since most of these. penetrations are in the lower part of the containment, =
the radial displacements experienced by them due to plastic deformation of containment would also be small.
The multiple penetration (X-71 and X-72) would not fail even for the most unfavorable forces which these pipes could sustain. For penetrations X-23 and X-26, the most likely location for failure is at the partial penetration fil- ,
let welds which join the pipe to the end plate. When failure of this weld oc-curs, the pipe remains in the hole provided in the end plate. The gap between the pipe and the end plate is likely to remain small unless the pipe wall buckles. Exact gap size is hard to compute. The SSPSA appears to use a uni-fonn gap size of 0.04 in., and 0.10 in, as median and upper estimates, respec-tively. The corresponding leak areas for X-23 (as well as X-20 to X-22) a.1d X-26 (as well as X-24, X-25, and X-27) penetrations are shown in Table 3.3.
The median failure pressure for X-23 penetration, is higher than the hoop failure pressure (216 psig) of the containment wall. These leak areas, there-fore, are not expected to develop.
penetration X-26 is expected to fail at a median pressure of 166 psig.
The combined leak area for a0 safety injection penetrations is obtained by independently adding individual median leak area of 0.5 in2 ,
Penetrations X-71 and X-72 are not likely to contribute to the overall
, leak area, as stated earlier.
- The main feedwater lines (penetrations X-5 to X-8) are 18-in. diameter, Schedule 100 pipes. The failure mode of most concern is failure of the flued head due to axial loads in the pipe at the penetration. At a median pressure of 180 psig, each one of these penetrations is likely to result in a leak area of 50 in 2each. A leakage area of 50 in 2, either from the failure of one or more feedwater line(s), will result in a substantial pressure reduction and thereby further deterioration of leakage area may not materialize. Neverthe-less if all four lines were to fail simultaneously, total leakage area of 200 in2 may result. In this case it will be category C failure.
The fuel transfer tube is fixed to an elevated floor inside the contain-ment. As the pressure in the containment increases, the containment wall moves outwards and thereby exerts pressure on the bellows. The most pertinent
' ~
Table 3.3 Leak Area Estimates for Mechanical Penetrations Median Median Line Penetration Leak Area Failure Pressure Size in2 psig X-20 to X-23 6.0 >216 12" CCW Supply and Return X-24 to X-27 2.0 166 4" Safety Injection X-71 and X-72
- Negligible i 1" Sample / Control X-5 to X-8 50 to 200 180 18" Main Feedwater Fuel Transfer Tube 3 172 --
8" X-16, X-18 See Text ,
On-line Purge HVAC-1,2 See Text 36" Containment Purge -
e
bellows from the viewpoint of containment leakage is the one inside the con- -
tainment (EP-2). Three potential failure modes, in their order of decreasing probability of failure, considered are (a) failure due to overall buckling of the bellows, (b) failure due to local buckling within the convolute, and (c)
- failure due to meridional bending strains. The SSPSA has estimated median leak area of about 3 in2 at a pressure of about 172 psig. This is a Type A failure.
There are two sets of containment penetrations which are open to the containment atmosphere a the inside. . The on-line penetrations (X-16 and X-18) are the 8-inch purge suction and discharge lines and containment purge 7
suction and discharge lines (HVAC-1 and 2) are the 36-inch lines. Each one of these four lines has two containment isolation valves, one inside and one outside the containment. operated t' utter-fly valves. At elevated All eight valves temperatures, are the pneumatically seal material (usually ethylene propylene) on these valves may deteriorate and lose its sealing function.
Any deposition of radioactive aerosols could further deteriorate the sealant material. Considering sealant degradation due to temperature alone, ethylene propylene seal life (Ref.10) is 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br />, 40 mts, or 20 sts if exposed to 400, 500 or 600 F, respectively.
In the event of the failure of the sealant material, a narrow crack leak -
path may develop and containment atmosphere may begin to leak into the space -
between the two isolation valves. Since the isolation v;1ves are closed from the containment isolation signal system, the leakage of containment atmosphere to the environment can occur only if the sealant of the outer containment iso-lation butterfly valve also fails. The time duration elapsed before this happens can be significantly long (of the order of hours). The SSPSA has es-timated .it to be long compared to the containment failure by other causes.
The SSPSA study, therefore, has disregarded this release path.
The available leakage area due to sealant degradation has been estimated (Ref.10) by assuming an eq~uivalent clearance of 1/16 inch between valve disc and body for ' low' and 1/8 inch for 'high' estimates. This gives a total leakage area of 17 in2 as low value and 34 in2 as high value. As noted ear-lier, the outer butterfly valves must also experience high temperatures prior to a through release path. This leak area is of Category B. The SSPSA study has argued that such a leak path is not likely to result prior to a gross
. containment failure (Category C).
Electrical penetrations can fail primarily due to overheating of the pot-ting compound. The SSPSA study has concluded that the failure of electrical penetrations is not expected to make a significant contribution to containment failure for any accident sequence. This conclusion, appears justified for the wet case, but, for the dry case, it is based on their esti.1 ate of slow over-heating of the potting compound. A careful thermal conduction calculation should be made to check this assessment. Such a calculation, similar to the problem of vent / purge line butterfly isolation valve failure, is beyond the scope of this work and hence it was not done.
The equipment hatch and personnel lock penetrations can fail either due to pressure loading or degradation of the sealant material (generally sili-cone). The structural failure, prior to containment failure, appears unlike-ly. The sealant material can degrade at high temperatures typical of a
. 1 severe accident. According to the 0-Ring Handbook (see Ref. 10), silicone can survive for twenty hours when exposed to 500 F temperature. Furthermore, the personnel air lock is a double door system so even if the sealant around one door were to become ineffective, substantial time delay would be required to make the second sealant also ineffective. It, thus, appears that the equip-ment hatch and personnel lock penetrations do not contribute significantly to Type B failure.
3.2.5.4 Containment Failure Probability The calculation of the probability of containment failure as a function of the pressure is quite involved. The method used and results reported in the SSPSA study seem reasonable except for the impact of all four main feed-l water lines failure. The SSPSA has categorized the failure of X-8 (one of the four main feedwater lines) penetration as Type B since anticipated leak area is 50 in2 It appears to us that when one such penetration fails, the remain-ing three will also fail at nearly the same pressure of 180 psig (195 psia).
Any depressurization due to a 50-ina hole is not likely to be fast enough to reduce the containment pressure substantially prior to the failure of the three' remaining penetrations. Assuming that all four main feedwater lines fail at 180 psig, an equivalent leak area of 200 in2 will result. This fail-ure, therefore, should be classified as Type C. The impact of this change on the containment failure probability numbers will be to reduce the rate for Type B with a corresponding increase in Type C. The total failure rate is not
- likely to change. Estimated containment failure fractions are compared with the SSPSA results in Fig. 3.10. Use of the modified failure fractions do not seem to cause any appreciable change in the risk estimates.
3.2.5.5 Containment Enclosure The containment enclosure building is designed to withstand 3.5 psi pres-sure difference between the enclosure and the environment. During normal operation, the internal pressure is about -0.25 inches of water gauge. The SSPSA study has calculated its pressure capacity to range from more than 1 psid to 10 psid. In view of relatively strong primary containment, the role of the secondary containment is important primarily for Type B failures of the primary containment. In the event of Type C failure, the secondary enclosure building might not play any significant role as far as the source tem calcu-lation is concerned. The SSPSA study, however, has not taken any credit for -
, the enclosure building.
.' 3.3 Definition of Plant Damage States and Containment Response Classes The grouping of accident sequences into plant damage states proceeds from the premise that the broad spectrum of many plant failure scenarios can be discretized into a manageable number of representative categories for which a single assessment of core and containment response will represent the response i
of all the individual scenarios in that category, The plant damage states classify events in accordance to the following three parameters:
- 1. Initiating Events "A" -
Large Loss of Coolant Accident "S" -
Small Loss of Coolant Accident "T" -
'. llET SEQUENCES 1.0 _, __
, , , , y g - - - SSPSA BNL 0.8 -
N -
m Benign s (Type B) \
t Failures N 3
O.6 -
E S ~
3 -
=
0.4 -
p p Gross /
g (Type C) /
o Fa11ures N
/
0.2 -
7
./
- I I I i 1 -
120 140 160 180 200 220 240 PRESSURE, PSIA Figure 3.10 Estimated containment failure fractions. .
G
i.
- 2. Timing of Cors Melt and Conditions at Vessel Failure "E" - No RWST Injection to RCS "L" - With RWST Injection to RCS
- - No Emergency Feedwater "FW" - With Emergency Feedwater l
- 3. Availability of Containment Systems !
"C" - Long-Term Containment Spray Cooling l "4" - Long-Term Spray Recirculation, No Cooling l "I" - Isolation Failure or Bypass Figure 3.11 gives the definition of the plant damage states and their re-spective frequencies (listed in Table 3.4) as used in the SSPSA risk model.
These damage states are categorized in a matrix of eight physical conditions in the containment (numerals (1) to (8)) and six combinations of containment safety function availability (letters A to F) for a total of 48 potential plant damage states. A ninth damage state type has been defined for accident sequences involving steam generator tube ruptures. Figure 3.11 indicates that only 39 plant damage states can be identified as credible sequences.
From the viewpoint of containment response, many of the plant damage states can be grouped into containment classes. The classes defined in Table 3.5 are differentiated primarily according to spray availability. The fre- ,
quency of each containment class is the sum of the frequencies of the plant states included therein.
Annual plant state frequencies calculated by the applicant 5 for both in-ternal and external events were reviewed by the Lawrence Livermore National Laboratory5 and were found acceptable. Table 3.6 presents the calculated con-tainment class frequency estimates for internal events, fires, floods and truck crashes; moderate and severe seismic events.
In order to comprehensively assess the risk from seismic events, it is necessary to make separate consequence calculations for those accidents which a are initiated by earthquences severe enough to impair evacuation. For this purpose, the seismic frequency estimates are divided into two categories in Table 3.6. The seismic events with instrument peak ground acceleration below 0.5g can be binned with internal events, fires, floods and truck crashes.
Seismic events with acceleration greater than 0.50g are judged to impair evac-l% tion, and must be treated separately in the consequence analysis.
These containment response classes (or plant damage states) are the starting point for the containment event tree analysis and they define the link or interfaces with the plant analysis.
(
' 3.4 Containment Event Tree and Accident Phenomenology An important step towards the development of the containment matrix in-volves the quantification of branch point probabilities in the containment event tree. These probabilities depend heavily on the analyses of degraded and core melt phenomenology and the containment building response described in Appendix H of the SSPSA.5 l
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Plant State Represents 2A AEC 4A TEC.SEC 2C/6C AE4 4C/8C TE4 10 AE 20/60 AL 30/70 SE TE/TEFW 4D/80 SL,TL 2E/6E AECI 4E/8E TECI 1F V 2F/6F AE1 3F/7F SEI 4F/8F SLI .
a Figure 3.11 Definitions of the plant damage states used in SPSS.
~
Table 3.4 Frequencies of Occurrence of the Plant Damage States Frequency Frequency Plant Damage (events per Plant Damage (events per State reactoryear) State reactor year) 10 3.03(-7) 6A 3.41(-7)
IF 1.89(-6) 6C 3.57(-10)
IFA 6.10(-11) 60 2.49(-7)
IFP 8.52(-7) 6E 5.30(-14) 2A 1.85(-6) 6F 2.08(-16) 2C -
1.91(-9) 6FA 1.11(-11) 2D 2.53(-7) 6FP 1.34(-12) 2E 1.40(-13) 7D 7.06(-5) 2F 1.06(-13) 7F 3.55(-8) 2FA 3.10(-11) 7FP 1.09(-5) 2FP 1.58(-10) 8A 4.50(-5) 30 1.94(-5) 8C 4.29(-8) 3F 5.00(-7) 80 5.51(-5) 3FP 6.21(-6) 8E 5.02(-11) 4A 1.28(-5) 8F 1.02(-10) 4C 1.65(-7) 8FP 1.95(-7) 4D 2.79(-6) 9A 7.51(-10) 4E 2.24(-11) 9C 3.62(-13) 4F 2.25(-13) 90 9.09(-9) i 4FP 1.18(-7)
TOTAL 2.30(-4)
NOTE: Exponential notation is indicated in abbreviated torm; i.e., 3.03(-7) = 3.03 x 10-7 e
Table 3.5 Containment Response Class Definitions
~
Class Plant State Represents 1 10 AE 2 2A/6A, 4A/8A AEC, TEC, SEC 3 2C/6C, 4C/8C AE4, TE4, SE4 4 3D/70 SE. TE, TEFW 5 . 20/60,4D/80 AL, SL, TL 6 IF, 2F, 3F, 4F, 6F, V 7F, 8F 7 2E/6E, 4E/8E AECI, TECI 8 1FP, 3FP/7FP Small leaks w/o RWST 9 2FP/6FP, 4FP/8FP Small leaks w/ RWST 10 1FA, 2FA/6FA Aircraft crashes 11 9A V2 (SGTR) 12 9C V2 (SGTR) 13 90 V2 (SGTR)~
i l
I i
9
. - --,-.-,,.,,,,,.-.,-,.,.,-,._,,,.,,..._.,,,,c,. -
. , - - - - - . _ , , , , - , . _._,,-,,n, , . - - . , - - . . , , . , . . . , - . , , . a,-- - ., ., ,
Table 3.6 Containment Class Mean Frequenciest
.l Frequency (per reactor year)
! Containment Internal, Fires. Internal l Response Class Floods and Truck Seismic <0.5g Seismic >0.59 Total Seismic and Crashes External 1
1 1.08E-7 -
1.%E-7 1.95E-7 3.03E-7 4 2 5.70E-5 1.54E-6 1.24E-6 2.78E-6 6.0E-5 i 3 1.80E-7 1.91E-8
- 1.91E-8 1.99E-7 i 4 8.60E-5 1.85E-6 2.27E-6 4.12E-6 9.0E-5 5 5.50E-5 1.10E-6 1.76E-6 2.86E-6 5.8E-5 6 1.80E-6 1.66E-7 3.93E-7 5.59E-7 2.4E-6 7 * * * *
- 8
- 5.29E-6 1.25E-5 1.79E-5 1.79E-5 L,
- 1.12E-7 2.40E-7
. 9 3.52E-7 3.52E-7 -d
! 10 * - - -
- 11 * - - - *
! 12 * - - -
- 13 * - - -
- i l tReference [5] Tables 5.1-3 and 9.2-9.
l
~
l
- Indicates frequencies less than 10-8 yr l.
l
[
l l
i l
The SSPSA containment event tree uses the twelve top events identified in -
Table 3.7 as major phenomenological phases which could occur with respect to the formation and location of core debris. These processes are grouped into four phases following an accident initiation (1) phenomena occurring while the core is still in place; (2) phenomena occurring while the core is located be-low the lower grid plate but is still in the reactor vessel; (3) phenomena oc-1 curring with the core debris located in the reactor cavity and on the contain- .
ment floor; and (4) the phenomena involving long-term cooling of the contain-ment and/or basemat penetration.
3.5 Containment Matrix (C-Matrix)
The twelve top events in the Seabrook containment event tree are summar-ized in Table 3.7. A negative response at any of the five nodes (a, 8,10, 11, and 12) in the containment event tree results in the failure of the con-tainment building by, a variety of failure modes. Each of these failure modes results in a particular radiological release category. For those paths that do not have a negative response at any of the five nodes, the path will even-
- tually result in no failure of the containment. The containment event tree thus links the plant damage states to a range of possible containment failure modes via the various paths through the tree. For a given tree, each path ends in a conditional probability (CP) of occurrence, and these cps should sum to unity. The quantification of an event tree is the process by which all the
, paths are combined to give the conditional probabilities of the various re-lease categories. In SSPSA, fourteen release categories are used for the quantification as summarized in Table 3.8. Note that two of these release categories (namely, SS and SS) correspond to intact / isolated containment.
Fission product release for this category would, therefore, be via normal leakage paths in the containment (and enclosure) building, which can be dif-ferent depending on availability of the enclosure building ventilation / fil-tration system.
Table 3.9 sets forth a simplified containment matrix (C-matrix) for the Seabrook plant using the containment response class definitions discussed in
( Section 3.3, and the release category definitions given in Table 3.8. In arriving at the C-matrix of Table 3.9 all of the very low probability values were disregarded. This was shown7 to be insignificant to the risk estimate.
. The present assessment of containment response for Seabrook plant is not
., based upon independent confimatory calculations of accident progression and containment response. Instead the knorledge gained from review of similar risk studies for otherbb" pr,essurized water reactors with large dry con-tainments is used to guide this assessment.
The mode and timing of containment failure cannot be calculated with a great degree of accuracy. Judgements must be made about the nature of the dominant phenomena and about the magnitude of several important parameters.
Furthermore, the codes and methods used for these calculations are approximate and do not model all of the detailed phenomena. Fortunately, risk measures are not sensitive to minor variations in failure mode and timing. It is im-portant, however, to properly characterize the major attributes of failure mechanisms; (1) whether the failure is early or late, (2) whether it is by overpressurization, bypass, or basemat melt-through and, (3) whether or not radionuclide removal systems are effective.
--,--------n--,,----, - , ,. - - , - - ,-ary>r-s.n.---.-,--, --,--.-~,,--,,,v-,.rm---, -r---*e--.---vo-w
- 1 l
Table 3.7 Accident Phase and Top Events for the Seabrook Containment Event Tree 1
l Accident Phase Top Event Initiator 1 Plant State Debris in Vessel 2 Debris Cooled in Place 3 No H2 Burn 4 Containment Intact Debris in Reactor Cavity 5 Debris Dispersed from Cavity 6 Debris Cooled 7 No H2 Burn 8 Containment Intact Long-Term Behavior 9 No Late Burn 10 Containment Shell Intact 11 Basemat Intact Failure Mode 12 Containment Failure * *
- Both event tree paths are containment failure; that is, success is small leakage and failure is gross leakage.
O e
ime--- ---w--re-r---,-- y o,-+,-,- , , , - - -
4 -,y - - - , - , ,-----p m ,w,,,-,,p_,,,--- ,, , , - - g- e,a w-----,mw-- , ,,,- ,--- _ -_.,ne--- - -_- - _ -- - - - - , - , - - - - _a -
i Table 3.8 Release Categories Employed in the Seabrook Station Risk Model Release Category Release
- Group Category Definition SS Containment intact / isolated with enclosure Containment air handling filtration working.
Intact / Isolated .
SS Same as S5 but with enclosure air handling filtration not working.
52 Early containment leakage with late over-pressurization failure and contair.sent building sprays working.
Tf Same as 52, but with containment building spray not working.
TN Same as Tf, but with an additional vaporiza-tion component of the source ters.
$3 Late overpressurization failure of the con- (
Long-Term tainment with no early leakage and contain-Containment ment building sprays working.
Failure TI Same as $3, but with containment building sprays not working.
TN Same as II, but with an additional vaporiza-tion component of the source term.
54 Basemat penetration failure, sprays operating TW Containment basemat penetration failure with containment building sprays not working and additional vaporization component of the source term.
56 Containment bypass or isolation failure with
. containment building sprays working.
TW Same as 56, but with containment building sprays not working and an additional vapori-Early zation component of the source term.
Containment Failure / Bypass 51 Early containment failure due to steam explo-sion or hydrogen burn with containment building sprays working.
IT Same as $1, but with containment building sprays not working.
"5 denotes applicability to Seabrook Station; number corresponds with contain.
ment failure mode; bar denotes nonfunctioning of containment building sprays; and V denotes achievement of sustained elevated core debris temperatures and associated vaporization release.
41 Table 3.9 Simplified Containment Matrix for Seabrook l
l Release Category Class 51 52 $3 SS S6 Tl TJ E!Y M T4V TW TW-d 1 0.60 0.40 2 0.01 0.99 3 1.0 4 O.89 0.11 5 1.0 1
! 6 1.0 7 1.0 8 1.0 9 1.0 10 1.0 11 1.0 12 1.0 13 1.0 l
4 6
-ce- -
The assessment of the containment response and failure mechanisms is '
based on the general understanding of that accident phenomenology and the con-tainment design characteristics discussed earlier. The phenomena of interest
- may be summarized as follows
Early Failure (51. 3T) which can result from a steam explosion or an early hy-drogen burn is believed to be unlikely. Although explosions in the reactor -
vessel lower plenum are probable, the resulting mechanical energy would be limited by the fraction of the core which could participate in a single explo-sion and by the efficiency of the process. In recent PRA reviews,6,7 we have assigned a conditional probability of 10 6 to steam explosion induced i containment failure. This probability leads to the conclusion that steam ex-
! plosions would have a negligible effect on risk, and consequently, the appli-
! cants 5x10 6 value is not included in the simplified C-matrix The conditionat . probability for an early containment failure due to ex-ternal events (i.e., aircraft crashes) is assigned 1 in the SSPSA as shown in Table 3.9. This simply indicates that an aircraft crash into the containment is assumed to fail the containment structure with certainty.
Early containment failure could also conceivably result from a rapid dis-persal of the core debris throughout containment in the form of aerosols which directly heat the atmosphere producing a rapid pressure / temperature pulse.
The dispersal could only be caused by the high primary system pressures that i
may exist at vessel failure for a number of transient sequences (recent calcu-4 lationsil indicate that there exists a potential for the establishment of a natural convection pattern inside the reactor vessel and the hot leg; which can cause rapid heatup of the RCS boundaries possibly leading to' failure and depressurization prior to bottom head melt through, thus eliminating, high pressure ejection sequences). The aerosols could rapidly pressurize contain-ment by direct heat exchange and concomitant chemical reactions. Scoping cal-i culations performed by the Containment Loads Working Group (CLWG) showed that l a very severe challenge to the containment integrity could result provided 25 l percent of the core mass were converted to aerosols.12 However, no consensus could be reached among the CLWG analysts as to the credibility of this param-eter value, and this failure mode is still speculative. Furthermore, the con-
- figuration of the Seabrook lower cavity would tend to reduce the dispersal of I
\ .
core debris beyond the reactor cavity boundaries.
S For the reasons outlined above (as well as the high containment failure
.' pressure for Seabrook), it is concluded that early overpressure failure has a very low likelihood.
Early Containment Leakage (S2, E, M) without gross failure of containment building is expected by the applicant to occur for large break LOCA sequences with RWST injection in the absence of sprays (T2), and for dry cavity sequences with a vaporization release ( M ).
The plant damage states for steam generator tube rupture sequences are assigned to the 52 release category without any attempt to compute a SGTR specific release category.
Late Overpressurization Failure (S3, U, N) can occur due to steam yoduc-j tion in a wet cavity or noncondensable gas production as a result of core-l l
1
concrete interaction for a dry cavity situation. For sequences in which early and intermediate failure is not expected to occur, and for which containment sprays are inoperable, failure is expected to be a certainty.
The conditional probability for a late overpressurization failure with a vaporization release (dry cavity) is shown to be 0.60 for large break LOCAs and 0.89 for small breaks and transients. This results from the relative com-petition between the late overpressure failure and the basemat penetration (TW) for accident sequences without the containment sprays for both low pressure (AE) and high pressure (SE, TE) scenarios.
The failure time for the late overpressurization failure mode is much longer than previously calculated for other large dry containment.1,3,'
l This is as a result of the very high failure pressure for the Seabrook con tainment. As a consequence of this high containment failure pressure (median pressure of 211 fo'r. wet and 187 psia for dry
- sequences) it is difficult to
, challenge the containment integrity by any conceivable event.
I 1 i Hydrogen deflagration early in the accident sequence or later after <
vessel failure when steam condensation occurring as a result of reactivation of sprays (due to regaining of ac power), or other natural heat sink mecha-nisms.7 which can produce a deinerted atmosphere is not expected to challenge the containment integrity. ,
The impact of changes in the containment failure distribution discussed 4
in 3.2.5.4 is not significant for late failures, j Basemat Penetration Failure (S4, TW) can result in the absence of containment heat removal system (sprays) for a dry cavity. A 26-inch high curb surrounds i
the reactor cavity that prevents the entry of water into the cavity unless all of the water from the RWST has been injected. The conditional probability of i
the basemat melt through is usually less than the late overpressurization failure, this is particularly true for Seabrook where there is a natural bed rock formation directly under the basemat foundation. Therefore, the basemat penetration failure probabilities are conservatively assigned.
No Failure (55. T!i) would result for all sequences with full spray operation.
The radiological releases are thus limited to the design basis leakage with essentially negligible off-site consequences.
Containment Isolation Failure (56. EY) is represented by an 8-inch diameter purge line. The accident sequences where the containment is either not iso-lated or bypassed (Event V) are assigned a conditional probability of unity to these release categories.
In the SSPSA, the conditional probability for failure to isolate contain-
, ment (8-failure mode) is assumed to be negligibly small. This is believed to
- be an optimistic assumption on the part of the applicant, because even for .
subatmospheric containment the e-failure mode is expected to have a i *For dry sequences, only primary system water inventory is available in the containment. In this case, the containment atmosphere becomes superheated and, at failure, the temperature can exceed 700*F.
I i
, _ - - - - - -m.,,___ _ , - - , . _ , . - - - _ _ . - ,--,----_c__.___.-,---,.--....--
i 40- .
i conditional failure probability ranging from 4x10 6 to about 2x10 3; there-fore, one expects the conditional probability s-failure mode probability for a large dry containment to be somewhat higher, and perhaps approaching -10 2 An interfacing systems LOCA (V sequence) results from valve disc rupture or disc failing open for series check valves that normally separate the high pressure system. This event results in a LOCA in which the reactor coolant bypasses the containment and results in a loss-of-coolant outside the contain-ment. Furthermore, the concurrent assumed loss of RHR and coolant make-up capability leads to severe core damage. In the SSPSA, only three possible interfacing systems LOCA sequences have been found and discussed. These are
- 1. Disc rupture of the check valve in the cold-leg injection lines of the RHR.
- 2. Disc ruptur'e of the two' series motor-operated valves in the normal RHR hot-leg suction.
- 3. Disc rupture of the motor-operated valve equipped with a stem mount-ed limit switch and " disc failing open while indicated closed" in the other motor-operated valve in the normal RHR hot-leg suction.
For the V-sequence, the core melts early with a low RCS pressure and a
, dry reactor cavity at vessel melt-through. The centainment sump remains dry and recirculation is not possible.
The core and containment phenomenology used to arrive at the split frac-tions for the containment event tree and thus the C-matrix are in general agreement with the other previous studiesl 3 4 for PWRs with large dry containments. Furthermore, the claimed unusually high strength of the Seabrook containment reduces the impact of sensitivity caused by uncertainties
! in the severe accident progression. However, should the claimed strength of 1
the containment be reduced to levels comparable to some of the other large dry containments, the impact of uncertainties may become significantly more pro-nounced, as discussed in our review of the MPSS-3.7 3.6 Release Category Frequencies Based on the containment class frequencies in Table 3.6 and the contain- '
ment failure matrix of Table 3.9, the release frequencies were computed and are sumarized in Table 3.10. Table 3.10 indicates that only eight of the release categories dominate the total release frequency.
Tables 3.11 and 3.12 set forth the contribution to core melt frequency from the various containment response classes and release categories, respec-tively. It is seen that containment classes 2, 4, and 5 dominate the core melt frequency while the release categories SS (containment intact), !IT and Tfi dominate the source term frequency.
9 h
.-..,,----,---..~--__,,,,,.-.e -
-._.-,.,,,,,,.,.,-.,_,._.,..,,_,..y,
--w - -,_,,_._,,._._-._,,--.___-w . - - . - - , - - _ - - - - - - - - . . - . - - - , - - -
-4 5-Table 3.10 Frequency of Dominant Release Categories (yr-1) l 1
4 Internal, Fires, .
Floods and Truck Internal and Category Crashes Seismic <0.5g Seismic >0.59 External
$3 7.50E-7 3.45E-8 2.69E-7 1.05E-6 SS 5.64E-5 1.52E-6 1.23E-6 5.92E-5 E
- 1.12E-7 2.40E-7 3.52E-7 S3 5.50E-5 1.10E-6 1.76E-6 5.79E-5 S2V
- 5.29E-6 1.25E-5 1.78E-5 S3V 7.66E-5 1.65E-6 2.14E-6 8.04E-5 1
S4V 9.50E-6 2.04E-7 3.27E-7 1.0E-5 56V 1.80E-6 1.66E-7 3.93E-7 2.36E-6
\
I t
Table 3.11 Contribution of Containment Response Classes to the Total Core Melt Frequency l
Internal, Fires, Internal Containment Floods and Truck and Class Crashes Seismic <0.5g Seismic >0.59 Total Seismic External 1 - - - - <0.01 2 0.25 <0.01 <0.01 O.01 0.2d 3 - - - -
<0.01 4 0.37 0.01 0.01 0.02 0.39 5 0.24 -
0.01 0.01 0.25 L T
6 0.01 - - - 0.01 y e s e s e 8
- 0.025 0.055 0.08 0.08 9-13 O
e 4
e
l .
Table 3.12 Release Category Frequency as a Fraction of Core Melt Frequency Release Internal, Firet. Internal Category Floods and Truck and Crashes Seismic <0.5g Seismic >0.5g Total Seismic External
)
l S3 <0.01 (0.01 <0.01 <0.01 <0.01 55 0.25 <0.01 <0.01 0.01 0.26 Si * <0.01 <0.01 <0.01 <0.01 i ST 0.24 40.01 <0.01 0.01 0.25 l 52V
- 0.03 0.05 0.08 0.08 i
! 7'
! S3V 0.33 0.01 0.01 0.02 0.35 S4V 0.04 <0.01 (0.01 (0.01 0.04 3Tf 0.01 <0.01 <0.01 <0.01 0.01 l
l i .
o
, , -4 8-4 ACCIDENT SOURCE TERMS In this chapter the approach utilized in the SSPSA to determine the frac-tion of fission products (originally in the core) that might be leaked to the outside environment will be outlined. The fission product source to the en-vironment, as calculated by this approach, will also be discussed for each release category.
4.1 Assessment of Severe Accident Source Terms The CORRAL-II code was used in the SSPSA for determining fission product leakage to the environment. This code takes input from the thermal-hydraulic analysis carried out for the containment atmosphere. In addition, it needs the time-dependent emission of fission products from the damaged fuel. The fission products were assumed to be released in distinct phases as suggested in the RSS 13 namely, the Gap Melt, and Vaporization phases. The time depen-dence of these phases is determined by the timing of core heatup, primary sys-tem failure, and start of core / concrete interactions. The methods used in the SSPSA differ from the RSS methods in the following ways:
- 1) The treatment of iodine was changed. Iodine was assumed to be in the form of cesium iodide. This was accomplished by merely using the same fraction of core inventory released for both the cesium group and the iodine group. *
] 2) Leakage releases are represented by a multi-puff model,
- 3) An uncertainty analysis was carried out in which it was attempted to account for shortcomings in the RSS methods.
In general, the net result of the SSPSA analysis was to reduce the fractional release of particulate fission products. This will be discussed in more de-tail later. In all, fourteen releases were determined (as shown in Table 3.8) ranging from containment bypass sequence to the no-fail sequence .
These release categories were evaluated by the applicant considering the containment failure mode, the availability of the spray system, and whether or not the cavity was wet or dry. Table 4.1 shows the point-estimate releases as determined by the methods outlined above. Containment failure mode $1 corre-
. ' sponds to a gross failure of the containment, resulting from a steam explo-sion, early pressure spike, or early hydrogen burn. Failure mode $2 repre-sents a loss of containment function early in the accident sequence. This loss of function takes the form of an increase in the leak rate to 40% per day where it stays until the containment fails due to overpressurization. Failure mode 53 represents a late overpressurization failure of the containment driven by decay heat or late hydrogen burn. Failure mode $4 represents a basemat melt-through, SS represents no containment failure and the leak rate is limit-ed to the containment design basis leak rate. Finally, failure mode S6 repre-sents sequences where the containment is failed or bypassed as part of the initiating event.
The second parameter considered in defining the source term is the avail-ability of sprays. This is determined by the plant damage states. Those i
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release categories with operating spray systeins are designated Si to S6, while '
those with spray systems not operating are designated U to W.
The third and final parameter considered in differentiating between source terms distinguishes between wet and dry cavities. In the case of a dry cavity a vaporization release due to core / concrete interactions was assumed to
' occur. For a wet cavity the core debris was assumed to be either quenched or '
it was assumed that the water in the cavity would scrub the vaporization re-lease thus effectively reducing the vaporization release to zero. The release
.! categories, which include a vaporization release, are shown in Table 3.8 and include the V sequence.
From the point of view of risk it was found5 that E E ST , and 3R were dominant either for acute or latent health effects. In view of this re-sult these four categories will be considered in more detail.
! ' Release categories U and 3W have late overpressurization failure modes, with no spray systems operating and differ only in the treatment of the vapor-ization release phase. The containment at Seabrook is calculated to fail at a median pressure of 211 psia for wet sequences and 187 psia for dry sequences.
At this pressure a gross failure is expected resulting in a puff release of approximately 0.5 hr release duration. From Table 4.1 it is seen that the U and T3Y sequences fail at 27.2 hrs and 91.5 hrs, respectively. These .
! failure times are several hours later than was calculated for Indian Point, Zion, and Millstone-3. The primary reason for the later failure in this case is due to the high calculate capacity of the Seabrook containment structure.
Table 4.2 compares the TT TJTrelease parameters with similar parameters for the other three reactors analyzed in References [1], [33, and [4]. Note that j a fair comparison should set (0l+1) equal to (Cs-Rb), since iodine was treated as Csl in the SSPSA but not in the other studies. It is seen that I, Cs, and Ba groups for Tf are approximately half the other releases, while the Te, Ru, and La groups are low by approximately an order of magnitude. This difference is due to the later failure time in the SSPSA (allowing more time for aerosol
- settling) and the absence of a vaporization release, (which dominates the 1
release of Te, Ru, and La). A similar comparison for the TR release indi-cates a uniform reduction of approximately an order of magnitude for all spe-cies. The reduction is entirely due to the late failure time for this
, sequence.
Another important consideration is the increased rate of fission product i release due to an increase in the leak area prior to attaining gross failure conditions. This can also impact the radionuclide transport mechanisms inside the containment due to changes in the containment thermal hydraulic conditions.
Release category TR is associated with early containment failure in which the containment function is compromised by increasing the leakage area in such a way that the leak rate increases from 0.1% per day to 40% per day, i This release rate is not enough to prevent an ultimate overpressurization failure. This release is modeled as a multi-puff release. The first puff corre'sponds to the release of fission products prior to the core debris melt-ing through the reactor pressure vessel (melt + gap). The second puff includes
- , , , - - - - , , - - - - - - - - - - - - , - - - - - - - - - ~ . , - - - - - - , - , - - - , . , - - - . _ . , _ - - , - - - - - . - - , - - , , - _ - - ,. ,,--- .-,------,.e
I Table 4.2 Late Overpressurization Failure Comparison 7
s M111 stone-3 Zion / Indiani " Indian 3 Seabrook Point Study Point l
1][ 1[FT M-7 TMLB' 2RW l
Xe 9.0(-1) 1.0 9 (-1) 9.6(-1) 1.0 10+! 1,.2(-1) 2.4(-2) 1.5(-1) 1.05(-1) 9.3(-2)
Cs-Rb 1.2(-1) 2.4(-2) 3.0(-1) 3.4(-1) 2.6(-1)
Te-Sb 2.2(-2) 3.0(-2) 3.0(-1) 3.8(-1) 4.4(-1)
Ba-Sr 1.5(-2) 2.6(-3) 3.0(-2) 3.7(-2) 2.5(-2)
Ru 4.4(-3) 2.3(-3) 2.0(-2) 2.9(-2) 2.9(-2) ,
La 4.4(-4) 3.9(-4) 4.0(-3) 4.9(-3) 1.0(-2)
T (release) 27.2 81.5 20 (hrs)
T (duration) 0.50 0.50 0.50 0.50 (hrs)
Energy 300E7 300E7 540E6 150E6 (Btu /hr) 4 4
l .
as-
. A ,.
\ .
the period of vaporization release and the third puff is equivalent to an
- overpressurization failure at the time of catastrophic containment failure.
In this model the duration of the melt release is seen to be 3.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br />, vapor-ization release 7.2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and the remaining release 78.0 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />. The long release time for this release is not tied to the release timing, i.e., melt
, release and vaporization release time, but determined by the blowdown charac-l teristics of the containment building. In order to determine the consequences from this release using the CRAC code the duration of release will have to be limited to the maximum allowed by the code model. Currently, this mode limits duration times of approximately 10 hrs. The suggested values for use in CRAC l are shown in Table 4.7. Also shown in Table 4.7 is the energy of release for the various releases. It should be noted that the high value of this param-
- eter relative to the corresponding RSS value is due to a different failure
- acchanism in the current study. The Seabrook containment building is a rein-forced concrete structure where as the RSS considered a pre-stressed concrete structure. Furthermore, the failure pressure of the Seabrook contr.inment is approximately 50 psi higher than the RSS containment failure pressure. This j will also tend to increase the energy of the released plume.
The total release of fission products from the sequences can be compared to the M-4 release determined for the Millstone-3 study. This comparison is made in Table 4.3. It is seen that, once adjustments are made for the dif-ferent ways in which iodine is treated, the SYTrelease is approximately half '
the M-4 release. Without the benefit of a calculation, it is difficult to judge whether the differences are reasonable.
Release category "5U has binned into it an isolation failure correspond-ing to an 8" diameter breach in containment and the interfacing LOCA (V-se-l~ quence). This sequence is also represented by a multi-puff release. In this case, as in the previous case, the total release time is long compared to the limited release times of the RSS13 consequence model. -
f The release fraction is compared (Table 4.3) to the M-4 release from the
! Millstone-3 study, PWR-2 from the RSS and the V-sequence from the RSSMAPIs study for Surry. Except for the iodine group, it is seen that the release fractions are comparable. If the iodine group were set equal to the cesium group value, it is seen that the value for 5Ti would be the lowest release fraction.
. 4.2 Source Term Uncertainty Analysis In this section we will briefly describe the uncertainty analysis carried out for the four dominant accident sequences and, where possible compare the fission product leakage to the environment to more mechanistic determina-tions. There are two contributors to the uncertainty in release characteriza-tion. First, uncertainty in timing of major events which are influenced by:
- 1) Prediction of key event times, and
- 2) The mix of accident sequences binned into a release category.
I Second, uncertainties in release fractions, which are influenced by:
i l 1) Uncertainties in timing of key events, and
- 2) Analysis methods and data.
I l.
i_ - - __ _ _ _
,. 'p ,'
Table 4.3 Comparison of Releases for Failure to Isolate Containment and the By-Pass Sequence Seabrook 5 Millstone-3 7 RSSA3* RSSMAP Is TN TW M-4 PWR-2 V-Sequence Xe 1.0 9.7(-1) 9.0(-1) 1.0 1.0 .
s -
OI+I 3.1(-1) 4.3(-1) 2.0(-1) 7.0(-1) 4.8(-1)
Cs-Rb 3.1(-1) 4.3(-1) 6.0(-1) 5.0(-1) 7.9(-1)
Te-Sb 3.2(-1) 4.0(-1) 5.0(-1) 3.0(-1) 4.4(-1)
Ba-Sr 3.4(-2) 4.8(-2) 7.0(-2) 6.0(-2) 9.0(-2)
Ru 2.5(-2) 3.3(-2) 5.0(-2) 2.0(-2) 4.0(-2) b La 4.2(-3) 5.3(-3) 7.0(-3) 4.0(-3) 6.0(-3)
T (release) 2.2 2.2 2.0 2.5 2.5 (hrs)
T (duration) 88.7 14 2.0 1.0 1.0 (hrs)
Energy (Btu /hr) 140E6 4E6 70E6 20E6 0.5E6 ,
a y
- The same as M1A release category in Millstone-3.7
, o 1
4 1
_ _ _ _ _ _ ______.___.------A----- - _ - - - - - - - - - - - - - - - - -
, .. . . r-The above principles were used in the $$PSA to determine source term mul-tipliers which would give a range of fission product leakage to the environ-ment. A probability was associated with each source term, and for later overpressurization failure modes ($T, M, and 577) the following discrete probability distribution was used in the SSPSA:
y Subcategory Probability U-a .02
(
U-b .08 U-c .30
. U-d .60 This indicates, for example that there is an 8% confidence level that U-b correctly defines the source term for the U release category.
The results of this analysis for the overpressurization failure modes is:
Particulate Release Factor (multiplier) *
. Probability U M M i
.02 .22 .63 .17
.08 .071 .22 .07
.30 .024 .065 .02
.60 .0071 .021 .007 V.
, From this table it is seen that for the most likely release, i.e., "d", the reduction factors of the source term are substantial.
2104I {he first two Volume releases V (Surry) forc.in the be compared TMLB'-c to releases and AB-c published sequences. in BMI-These two
. sequences correspond to late containment failures and are both binned into U
'and "!!T sequences . A comparison of these sequences is shown on Table 4.4.
.From this table it is evident that for the volatile species, Xe, Cs, and I, the release cEggories "Il and 337" bracket or exceed the mechanistic estimates carriedoutin'8MI-2104forboththeTMLB'andABsequengs. However, for the less volatile species Te, Ba, Ru, and La, the BMI-2104 calculated releases
. for the TMLB' sequence are higher than all the U and M releases. This dis-crepancy is primarily ~due to the comparatively early failure time. It is felt that agglomeration and settling would reduce the source for the TMLB' sequence
~
to valdts close to those reported for U and M. No comparative sequence for M was enalyzed in BMI-2104.
A s
\
n-,, ------,---,e-m--,-.,~,--- , - -
Table 4'.4, Comparison of A8-c and TML8'-c (BMI-2104) to 3Ti and 3T n.l s. rrecticas n.i Pron.6:alty ni Category Tle. thrs) X. Cs I T. Se she to 5-e 02 28 I.0 f.5(-2) 1.M-2) I .M-2) 8.6(-3) 1.M-3) 2.5(-43 1
s3V-b 08 36 9.0(-1) 5.3(-3) 5.3(-3) 6.6(-33 5.7(-4) 5.I(-4) 8.6(-5) b-c 30 54 8.0(-15 f.6(-3) 1.6(-3) 2.0(-3) 8.7(-4) 1.M-4) 2.M -5) 5-d 60 89 7.0(-13 5.0(-4) 5.0(-4) 6.3(-4) 5.M-Si 4.8(-5) 8.2(-6) 3-a 02 22 1.0 2.6(-2) 2.6(-2) 4.M -3) 3.M-3) 9.7(-4) 9.7(-5)
! 73-b 08 28 9.0(-1) 8.5(-3) 8.M -3) 1.6(-3) 1.l(-3) 3.II-4) 3.II-5) i .
73-c 30 34 8.0(-1) 2.M -3) 2.M -3) 5.3(-43 3.6(-4) 3.It-41 1.II-Si
[3-d 60 53 7.0(-il 8.M-41 8.M-4) 3.6(-4) 1.I(-4) 3.l(-5) 3. It-6) 7R 88-c -
12 1.0 2.8(-3) 6.0(-4) 8.M-21 f.7(-25 2.4(-53 4.M-4)
AS-c - 24 1.0 4.8(-5) 4 .7(- 51 4.0(-53 4.M-5) 2.4(-7) 3.6(-5) e O e
l
, . . ab-In the case of the 3ri release category a different probability distribu-tion was used. This change reflects the break location, which initiates the
. V-sequence. This break could be either in the hot-leg (b release subcategory) or the cold-leg (c release subcategory). This sequence is modeled as multi-puff release and each puff is treated separately. In this comparison only the sum of the release will be considered, since no adequate method of analyzing a multi-puff release is readily available. Table 4.5 shows a comparison be-tween the totals of the various TITreleases and two V-sequence releases com-puted for Surry and published in BMI-2104. One of the V-sequences is " dry,"
implying no water in the path of the release and the other is " wet," implying that the release passes through 3 feet of water before entering the atmo-sphere. From this comparison it can be seen that all the releases, except Cs for the " dry" V-sequence, are bracketed by the T6Y releases.
4.3 Recomended Source Tems The severe accident source terms used in the Seabrook Probabilistic Safe-ty Study reviewed in the previous sections, are aimed at the multi-puff cen-sequence model present in the CRACIT computer code. In order to make these source terms useful to the NRC staff for evaluation with the CRAC code, total releases must be used as summarized in Table 4.6. Furthermore, the suggested source terms of Table 4.6 together with their release category characteristics .
given in Table 4.7.
! It must also be noted that the suggested source term for the Steam Gener-l ator Tube Rupture (SGTR) sequence is assumed to be one-tenth of the source term for the event V (537). This is believed to be a conservative estimate and can be used in the absence of a more specific mechanistic calculation.
4 The suggested source tems in Tables 4.6 and 4.7 can be used to estimate the health and economic effects (consequences) due to radioactive atmospheric releases as a result of core melt accidents in the Seabrook Station.
The resulting consequences together with the frequency of radiological releases will enable the establishment of the severe accident risk at the Seabrook site.
4 4
I l
Table 4.5 Comparison of 56V (sum) to V-sequence (Surry)
Release Fractions Release Probability Category Xe Cs I Te Ba Ru La 56V-a .02 .97 4.3(-1) 4.3(-1) 4.06(-1) 4.2('-2) 3.32(-2) 5.3(-3) 56V-b .45 .97 2.95(-1) 2.95(-1) 1.36(-1) 3.53(-2) 1.52(-2) 2.0(-3) 56V-c .45 .97 1.295(-1) 1.295(-1) 3.2(-2) 1.593(-2) 5.2(-3) 5.3(-4) ,
O.
56V-d .08 .97 5.2(,-2) 5.2(-2) 1.3(-2) 6.6(-3) 2.0(-3) 2.2(-4)
V (dry) -
1.0 5.52(-1) 1.99(-1) 1.2(-1) * *
- V -
1.0 1.04(-1) 3.84(-2) 2.5(-2) * * * -
(submerged)
- Individually not reported.
o W
- e
- -bB-a ***
Table 4.6 BNL-Suggested Source Terms Release Category Xe 01 I-2* Cs Te Ba Ru La 51 0.94 - 0.023 0.023 0.24 0.0033 0.41 9.8E-5 52 0.89 -
2.1E-5 2.1E-5 4.4E-6 2.9E-6 8.8E-7 8.8E-8 53 0.90 7E-3 1.E-7 1.E-7 1.9E-8 1.3E-8 3.8E-9 3.8E-10 SS 0.0091 ,- 3.5E-8 3.5E-8 6.1E-9 4.0E-9 1.2E-9 1.2E-10 56 0 .90 -
3.6E-3 3.6E-3 6.7E-4 4.4E-4 1.3E-4 1.3E-5 l
lff 0.94 - 0.75 0.75 0.39 0.093 0.46 2.8E-3 lCT 0.90 - 0.31 0.31 0.057 0.038 0.011 1.1E-3 ,
1CET 1.0 -
0.31 0.31 0.32 0.034 0.025 4.2E-3 1CI 0.90 - 0.12 0.12 0.022 0.015 4.4E-3 4.4E-4 30GI 1.0 - 0.024 0.024 0.030 2.6E-3 2.3E-3 3.9E-4 317 1.0 -
0.058 0.058 0.072 6.2E-3 5.4E-3 9.1E-4 SS 0.014 7E-4 5.2E-7 5.2E-7 9.5E-8 6.3E-8 1.9E-8 1.9E-9 I ll? 0.97 -
0.43 0.43 0.40 0.048 0.033 5.3E-3 337-d 0.90 - 0.043 0 .043 0 .040 4.8E-3 3.3E-3 5.3E-4 l
l
- 337-d release is 1/10th of the 337 values.
i 8 Table 4.7 BNL-Suggested Release Characteristics for Seabrook Release Release Release Release Warning * ~
Duration Release Time Energy Height Time Category (hr) (hr) (Btu /hr) (m) (hr) 51 1.9 0.5 <140E6 10 0.35 52 2.6 1.0 <140E6 10 1.05 53 66.0- 0.5 300E6 10 63 55 1.9 10 <140E6 10 0.35 56 4.5 4 <140E6 10 0.50 TT 1.4 0.5 300E7 10 0.30
}
TF 27 10 <140E6 10 26 Tff 35 10 <140E6 10 35 TT 27 0.5 <140E6 10 26 TTi 81 0.5 300E7 10 76 TTi 50 0.5 300E7 0 49 SS 4.3 10 <140E6 10 0.30 Tif 2.5 1.0 3.7E6 10 1.0
. T07-d 2.5 1.0 3.7E6 10 1.0
- Warning time is defined as the time after core melt starts to the time of radiological release.
-u '
- 5.
SUMMARY
AND CONCLUSIONS The purpose of this report is to describe the technical review of the Seabrook Station Probabilistic Safety Assessment and to present an assessment of containment performance, and radiological source term estimates for severe core melt accidents.
The containment response to severe accidents is judged to be an important factor in mitigating the severe accident risk. There is negligible probabil-ity of prompt containment failure or failure to isolate. Failure during the first few hours after core melt is also unlikely and the timing of overpres-sure failure is very long compared to the Reactor Safety Study (WASH-1400).
Most core melt accidents would be effectively mitigated by containment spray operation. A comparison of SSPSA and RSS containment failure frequencies is given in Table 5.1.,
Our assessment of the containment failure characteristics indicate that, there is indeed a tendency to fail containment through a realistic benign mode compared with the traditional gross failures.
The point-estimate release fractions used in the S$pSA are comparable in magnitude to those used in the RSS. In those cases where comparisons can be made to the more mechanistic source term study carried out by the Accident Source Term Program Office (ASTPO) of the NRC and reported in BMI-2104 it was -
found that the SSPSA releases were either higher than or for the most part similar to the recent release fractions. It was also found that the energy of release was somewhat higher in the SSPSA than for other existing studies.
W b .
l' 1
I l
, lj o' '
l Table 5.1 Comparison of SSPSA and WASH.1400 Containment Failure Frequencies 1 of Core Melt Frequency WASH 1400 SSPSA Gross, Early Failure 34 1 Gradual Overpressure or Basemat Melt-through Failure . 66 73 Intact Containment -- 26 O
h e
4 l
. - ~ < , _ - _ . _ _ - _ _ - , , . . _ . . - . . _ _ - . . _ _ _ _ - , _ . - . _ . - _ _ . _ .
I I
t
- 6. REFERENCES
- 1. " Zion Probabilistic Safety Study," Comonwealth Edison Company (September 1981).
- 2. " Limerick Probabilistic Safety Study," Philadelphia Electric Co.
(September 1982).
- 3. " Indian Point Probabilistic Safety Study," Power Authority of the State of New York and Consolidated Edison Company (March 19R2).
- 4. " Millstone Unit 3 Probabilistic Safety Study," Northeast Utilities (August 19R3). .
- 5. B. J. Garrick, et al., "Seabrook Station Probabilistic Safety Assessment," pickard, Lowe and Garrick, Inc., PLG-0300 (December 1983).
6 A. A. Garcia, et al., "A Review of the Seabrook Station Probabilistic Safety Assessment," Lawrence Livermore National Laboratory Report (Dec.
12,1984).
- 7. M. Khatib-Rahbar, et al., " Review and Evaluation of the Millstone linit 3 Probabilistic Safety Study: Containment Failure Modes, Radiological Source Terms and Off-Site Consequences," NUREG/CR-4143, BNL-NUREG-51907 (September 19R5).
- 8. R. O. Wooten and H. Avci, " MARCH: Meltdown Accident Response Character-1stics - Code Description and liser's Manual," BMI-2064, NUREG/CR-1711 (1980).
- 9. J. F. Muis, et al., "CORCON-Mod 1: An Improved Model for Molten Core / Concrete Interactions," SAND 80-2415 (1981).
- 10. R. E. Miller, A. K. Agrawal, and R. E. Hall, "An Estimation of Pre-Exist-ing Containment Leakage Areas and Purge and Vent Valve Leakage Areas Re-sulting from Severe Accident Conditions," A-3741,11/15/84 (Oraft report
. dated August 1984) transmitted via letter to V. Noonon, June 29, 1984
- . See, also, A. K. Agrawal and R. E. Hall.
- 11. W. Lyon (organizer), "RCS Pressure Boundary Heating During Severe Acci-dents," USNRC Meeting, Bethesda, Maryland (May 14,1984).
- 12. " Estimates of Early Containment Loads From Core Melt Accidents," Con-tainment loads Working Group, NUREG-1079 (Draft 1985).
- 13. " Reactor Safety Study," U.S. Nuclear Regulatory Commission, WASH-1400, NUREG-75/014 (October 1975).
14 " Preliminary Assessment of Core Melt Accidents at the Zion and Indian '
Point Nuclear Power Plants and Strategies for Mitigating Their Effects,"
NUREG-0850, Vol. 1 (November 1981).
. . / .
- 15. G. S. Kol b, et al . , " Reactor Safety Study Methodology Application Program: Oconee #3 PWR Plant," NUREG/CR-1659/2 of 4
- 16. J. A. Gieseke, et al., "Radionuclide Release Under Specific LWR Accident Conditions," Battelle Columbus Laboratory Reports BMI-2104 (July 1984, Draft).
t O
4
SMA 12911.01 Rev. 1/NTS 1589.01 1
! Technical Report No. 1589.01 b f C- ,
- [ SEISMIC FRAGILITIES OF STRUCTURES AND
- COMPONENTS AT THE SEABROOK GENERATING STATION, UNITS 1 AND 2 by
~
D. A. Wesley
' R. D. Campbell R. B. Narver G.S. Hardy M. W. Salmon Prepared by NTS Engineering
/ 6695 East Pacific Coast Highway Long Beach, CA 90803
- Prepared for
{
" PICKARD, LOWE, AND GARRICK, INC.
l Irvine, CA l , l and New Hampshire Yankee Division Public Service Company of New Hampshire Seabrook, New Hampshire June 1986 h1M -87'6 0h bi M Wd/FA[A2 %\
o '" Y ' -
v p/ B/s t i
. - - - - - - , - , - . - - . . __ w,--- ---rr--,7 - - -
TABLE OF CONTENTS Section Title Pace 1 INTRODUCTION . . . . . . . . . . . . . . . . . . . . . . . . . 1-1 2 GENERAL CRITERIA FOR DEVELOPMENT OF MEDIAN SEISMIC SAFETY FACTORS ........................... 2-1 2.1 De fi ni tion of Fail ure . . . . . . . . . . . . . . . . . . 2-2 2.1.1 Seismic Category I Structures .......... 2-2 g, 2.1.2 Seismic Category I Equipment and Piping ..... 2-2 i 2.1.3 Non-Category I Structures ............ 2-3 d
2.1.4 Non-Seismic Category I Equipment and Piping ... 2-3 2.2 Basis for Safety Factors Derived in Study . . . . . . . . 2-4 2.2.1 Structural Response and Capacity . . . . . . . . . 2-4 2.2.2 Seismic Category I Piping and Equipment Response and Capacity .............. 2-4 2.3 Formulation Used for Fragility Curves . . . . . . . . . . 2-6 2.4 Design and Construction Errors ............. 2-10 3 DIFFERENCES BETWEEN CRITERIA USED FOR DESIGN OF SEABROOK AND PARAMETERS USED IN THE EVALUATION OF THE SEISMIC CAPACITY 3-1 3.1 Strength ........................ 3-2 3.2 Ductility . . . . . . . . . . . . . . . . . . . . . . . . 3-2 3.3 Earthquake Duration . . . . . . . . . . . . . . . . . . . 3-3 3.4 System Response . . . . . . . . . . . . . . . . . . . . . 3-4 3.4.1 Earthquake Characteristics . . . . . . . . . . . . 3-5 3.4.2 System Damping . . . . . . . . . . . . . . . . . . 3-5 3.4.3 Load Combinations ................ 3-6 3.4.4 Modal Combination ................ 3-8 3.4.5 Combination of Responses for Earthquake Di rectional Components . . . . . . . . . . . . . . 3-9 3.4.6 Structure Modeling Considerations ........ 3-9 4 STRUCTURES . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1
. 4.1 Median Safety Factors and Logarithmic Standard Deviations ....................... 4-1 4.1.1 S t ruc t u re Ca pa ci ty . . . . . . . . . . . . . . . . 4-5 4.1.1.1 Concrete Compressive Strength . . . . . . 4-5 _
4.1.1.2 Reinforcing Steel Yield Strength .... 4-7 4.1.1.3 Shear Strength of Concrete Walls .... 4-7 i
4
TABLE OF CONTENTS (Continued)
Section Title Page
, 4.1.1.4 Example of Shear Wall Failure in Shear . . . . . . . . . . . . . . . . . . 4-10 4.1.1.5 Strength of Shear Walls in Flexure i Under In-Plane Forces . . . . . . . . . . 4-12 4.1.1.6 Example of Shear Wall Failure in i Flexure . . . . . . . . . . . . . . . . . 4-13 g 4.1.2 Structure Ductility ............... 4-14 e :
4.1.2.1 Example of Inelastic Energy Absorption Factor ................. 4-15 4.1.3 Earthquake Duration ............... 4-16 g 4.1.4 Spectral Shape, Damping, and Modeling Factors .. 4-17 4.1.4.1 Example of Spectral Shape, Damping, and Modeling Factors .......... 4-21
, 4.1.5 Modal Combination ................ 4-22 4.1.6 Contination of Earthquake Components . . . . . . . 4-23 4.1.7 Soil-Structure Interaction Effects . . . . . . . . 4-24 4.2 Containment Building ................... 4-25 4.2.1 Containment Failure Modes ............ 4-26 4.3 Containment Enclosure Building ............. 4-27 1 4.3.1 Containment Enclosure Failure Modes ....... 4-28 4.4 Primary Auxiliary Building ............... 4-29 4.4.1 Primary Auxiliary Building Failure Modes . . . . . 4-29 4.5 Service Water Pumphouse and Circulating Water Pumphouse . 4-30 4.5.1 Service Water Pumphouse Failure Modes ...... 4-30 4.6 Service Water Cooling Tower . . . . . . . . . . . . . . . 4-31 4.6.1 Cooling Tower Failure Modes ........... 4-32
, 4.7 Condensate Storage Tank and Enclosure . . . . . . . . . . 4-32 4.7.1 Tank and Enclosure Failure Modes . . . . . . . . . 4-33 g 4.8 Control and Diesel Generator Building . . . . . . . . . . 4-34 4.8.1 Control and Diesel Generator Building Failure Modes ...................... 4-34 4.9 Fuel Storage Buil di ng . . . . . . . . . . . . . . . . . . 4-35 4.9.1 Fuel Storage Building Failure Modes ....... 4-35 ii l
l
TABLE OF CONTENTS (Continued)
Ti tle Pace Section 5 EQUIPMENT FRAGILITY ..................... 5-1 5.1 Equipment Fragility Methodology . . . . . . . . . . . . . 5-1 5.1.1 Fragility Derivation . . . . . . . . . . . . . . . 5-1 5.1.1.1 Equipment Capacity Factor . . . . . . . . 5-3 5.1.1.1.1 Strength Factor ....... 5-4 b 5-7 5.1.1.1.2 Ductility Factor . . . . . . .
4 5.1.1.2 Equipment Response Factor . . . . . . . . 5-8
- 5.1.1.2.1 Qualification Method Factor . 5-10 5.1.1.2.1.1 Static Analysis . 5-10 5.1.1.2.1.2 Dynamic Analysis 5-10 5.1.1.2.1.3 Testing . . . . . 5-11 5.1.1.2.2 Equipment Spectral Shape Factor . . . . . . . . . . . . 5-11 5.1.1.2.2.1 Peak Broadening l
and Smoothing . . 5-12 5.1.1.2.2.2 Artificial Time-History Generation 5-13 5.1.1.2.3 Modeling Factor ....... 5-14 5.1.1.2.4 Damping Factor . . . . . . . . 5-15 5.1.1.2.5 Mode Conbination Factor ... 5-17 5.1.1.2.6 Earthquake Component Combination Factor . . . . . . 5-17 5.1.1.2.7 Boundary Conditions Factor (Testing) .......... 5-19 5.1.1.2.8 Spectral Test Method . . . . . 5-19 5.1.1.2.9 Multi-Directional Effects .. 5-20 1
5.1.1.2.9.1 Biaxial Testing . 5-20 5.1.1.2.9.2 Uniaxial Testing 5-21
- i 5.1.1.3 Structural Response Factors . . . . . . . 5-22 5.1.1.4 Earthquake Duration Factor ....... 5-23 5.1.2 Infonnation Sources ............... 5-24 ii l
5.1.2.1 Seismic Qualification Analysis Reports . 5-25 5.1.2.2 Seismic Qualification Test Reports ... 5-25 5.1.2.3 Final Safety Analysis Report & SORT Summaries . . . . . . . . . . . . . . . . 5-25 l
?
iii I
TABLE OF CONTENTS (Continued)
Title Page Section 5.1.2.4 Vendor Drawings or Design Reports from which New Analyses are Conducted . . . . . . . . . . . . . . . 5-26 5.1.2.5 Past Earthquake Experience ...... 5-26 5.1.2.6 Specification for the Design of Equipment . . . . . . . . . . . . . . . 5-26
- s. .
5.1.3 Equipment Categories . . . . . . . . . . . . . . 5-27
, 5-28 5.2 Equipment Fragility Examples .............
"' 5.2.1 Example of a Plant-Specific Design Report Fragility Derivation . . . . . . . . . . . . . . 5-29
' 5.2.1.1 Spray Additive Tank Capacity Factor . . 5-30 5.2.1.2 Spray Additive Tank Equipment Response Factor ................ 5-33 -
5.2.1.3 Spray Additive Structural Response Factors ............... 5-35 5.2.1.4 Spray Additive Tank Earthquake Duration Factor . . . . . . . . . . . . 5-35 5.2.1.5 Spray Additive Tank Ground Acceleration Capacity ............... 5-35 i
5.2.2 Example of Qualification Test Report Fragility 5-36 Derivation . . . . . . . . . . . . . . . . . . .
5-36
- 5.2.2.1 Battery Charger Capacity Factors ...
5.2.2.2 Battery Charger Equipment Response Factors . . . . . . . . . . . . . . . . 5-39 5.2.2.3 Battery Charger Structural Response Factors . . . . . . . . . . . . . . . . 5-41 5.2.2.4 Battery Chargers Earthquake Duration Factor ................. 5-42 i
l =
5.2.2.5 Battery Chargers Ground Acceleration .
Capacity ............... 5-42 5.2.3 Example of Generic Fragility Derivation Based on Design Specifications ............. 5-43 5-43 5.2.3.1 Failure Modes of a Piping System ...
5.2.3.1.1 Piping Failure Modes . . . . 5-44 5.2.3.1.2 Support Failure Modes ... 5-45 5.2.3.2 Piping Capacity Factor ........ 5-45 5.2.3.2.1 Piping Strength Factor . . . 5-45 iv l
w
%- , ,m-- - + - . -
TABLE OF CONTENTS (Continued)
Title Pace Section 5.2.3.2.2 Piping Ductility Factor . . 5-50 5.2.3.2.3 Piping "Three-Hinge" Factor .......... 5-50 ,
5.2.3.3 Piping Equipment Response Factors .. 5-51 5.2.3.4 Piping Structural Response Factors . . 5-52
5.2.3.5 Piping Earthquake Duration Factor .. 5-52
' 5.2.3.6 Piping Ground Acceleration Capacity . 5-53 5.2.3.7 Piping Supports Capacity Factor ... 5-53 5.2.3.8 Ground Acceleration Capacity of Piping Supports ........... 5-57 5.2.4 Example of Fragility Derivation Based Upon New Analysis ................. 5-57 5.2.4.1 RWST Capacity Factor . . . . . . . . . 5-58 5.2.4.2 RWST Equipment Response Factor . . . . 5-60 5.2.4.2.1 Qualification Method ... 5-60
~
5.2.4.2.2 Spectral Shape Factor . . . 5-60
' 5.2.4.2.3 Damping Factor ...... 5-61 5.2.4.2.4 Modeling Factor . . . . . . 5-61
' 5.2.4.2.5 Mode Combination Factor . . 5-61 5.2.4.2.6 Earthquake Component Combi-nation Factor . . . . . . . 5-62 5.2.4.2.7 Overall Equipment Response
' Factor .......... 5-62
- 5.2.4.3 RWST Structural Response Factor ... 5-62 5.2.4.3.1 Spectral Shape Factor . . . 5-63 5.2.4.3.2 Damping . . . . . . . . . . 5-63
. 5.2.4.3.3 Modeling Factor . . . . . . 5-64 i
5.2.4.3.4 Soil-Structure Interaction. 5-64
- ' 5.2.4.3.5 RWST Structural Response l Factor .......... 5-64 l
5-64 5.2.4.4 Earthquake Duration Factor . . . . . .
5.2.4.5 RWST Capacity ............ 5-65 l
V i
- - _ = . . . . , . . _ , . . _ . . - - . _ . . . . . , _ , - . , , . . - _ . - . , , _ ~ . . _ _ _ _ _ . _ , -
I TABl.E OF CONTENTS (Contents) 4 Title Page Section 5.2.5 Example of Fragility Based on Engineering Judgment and Earthquake Experience . . . . . . 5-65 5.3 Equipment Fragility Results ............ 5-66 5.3.1 General Results ............... 5-67 4
' D REFERENCES
' APPENDIX A I
l s
4
- i J
4 i i t
vi i
o 1
REVISIONS - June 1986 Updated fragility descriptions for:
Non-vital Electric Power 4160 Volt Switchgear
- Emergency Feed Pumps Motor Driven
' Turbine Driven 120 V Instrument Bus 480 V Motor Control Centers 480 V Transformers and Busses (480 V Unit Substation)
Refueling Water Storage Tank Primary Component Cooling Water Heat Exchanger Diesel Fuel Oil Day Jank RHR Pump Safety Injection Pump
. Charging Pump Added fragility descriptions for:
Solid State Protection System Deleted Figures 5-2 through 5-7, (no longer applicable)
Revised text to incorporate updates of fragility. Revisions in text are marked with a vertical bar on the right-hand margin.
Deleted Tables 5-9, 5-10 and 5-11 (no longer applicable). Old
. . Table 5-12 becomes 5-9).
J i
- 1. INTRODUCTION A probabilistic risk assessment (PRA) of the Seabrook Nuclear Generating Station was completed by Pickard, Lowe and Garrick, Inc., for
. Public Service Co. of New Hampshire in December, 1983. In that evalua-tion, system models, event trees, and fault trees were utilized to deter-
> mine the frequency of radioactive release from the site due to random 4 .
equipment failure and failures initiated by natural hazard events. Earth-
, quakes were one of the extreme natural hazards considered in this PRA.
Structural Mechanics Associates, Inc. (SMA), under subcontract to
! Pickard, Lowe and Garrick, provided the required information on the response of plant structures and equipment to earthquake (seismic) events.
The original version of this report was included as part of the Seabrook Station Probabilistic Safety Assessment (SSPSA) report, Ref. 48, to document the seismic fragilities that were used to help estimate the risk
- contribution due to seismic events. While seismic events were found to make a small contribution to fisk, this contribution became more visible
- as a result of subsequent risk management efforts that reduced the risk contributions of non-seismic events. In addition, NRC conducted a review of the SSPSA including the portion dealing with seismic risk analysis, (Ref. 47). To account for new information about the Seabrook design, the l NRC review conrnents and the greater visibility of the seismic risk contribution in light of risk management activities, it was decided to reevaluate the seismic fragilities of selected key components. The
! purpose of this revision of our riport is to document these updated fragility analyses for the Seabrook Station.
i.
The frequency of seismically-induced failure as a function of
- f peak ground acceleration for both safety-related structures and equipment has been developed by SMA for the Seabrook facility. Also included is the expected variability in the frequency of failure. The determination l
of the seismic hazard is being conducted by others. The information for l
1-1 l
l
_ _ _ . . _ - ~ . _ . _ . . _ _ . _ _ _ , _ _ . . , _ _ _ _ _ _ _ . _ _ . _ . _ _ ,
both the frequency of occurrence of different levels of peak ground acceleration and the frequency of failure of the safety-related systems and components will then be incorporated into the risk models by Pickard, ,
Lowe and Garrick, to determine the frequency of seismic-induced radio-active release from the site.
In order to correctly interpret the fragilities derived in this report, it is necessary to define the effective acceleration to which 6 tbse fragilities are anchored. It is recognized that the damage poten-tial of an earthquake depends on many f actors, among which are magnitude, peak acceleration, and duration. For the Seabrook site, it is estimated that the majority of seismic risk results from earthquakes that have magnitudes between 5.3 and 6.3. This is the range represented by the site-specific spectra used to evaluate the fragilities. Because the site-specific spectra used in this study are centered around this magnitude range, the fragilities given in this report are to be anchored to the mean peak acceleration. This acceleration is the average of the peak accelerations from two orthogonal horizontal components. Note that if the magnitude range were higher, say 6.0 to 7.0, either a different ,
set of median spectra and a different duration factor (which will be discussed in Sections 3.3 a'nd 4.1.3) would required, or else it would be
' necessary to anchor the fragilities to an effective acceleration that ,
would be greater than the mean peak acceleration. Conversely, if the magnitude range were lower and this set of fragilities were to be used,
' the effective anchoring acceleration would have to be less than the mean r peak acceleration in order to accurately predict the damage potential.
The Seabrook Station was designed in the late 1970's and early 1980's in accordance with criteria and codes in effect at that time '
(Reference 1). The Seabrook systems and components which are essential
- ' to the prevention or mitigation of consequences of accidents which could affect the public health and safety were designed to enable the facility to withstand the effects of natural forces including earthquakes. The _
design criteria included the effects of simultaneous earthquake and I
l 1-2
- . _ , . _ _ . _ . _.- _ _ . _ _ _ _ , _ - . - ._.,.._~--,.,,,_.,% , _ , . . . , . _ _ _ - , _ . _ _ , _ _ . . . _ _ _ . . _ , _ _ _ _ _ . - _ , ..-,..-,_,,,,,-7 ,
loss-of-coolant-accident (LOCA) conditions. The plant was designed to withstand both an Operating Basis Earthquake (OBE) and a Safe Shutdown Earthquake (SSE). The structural design criteria for the SSE was based on 0.25g and the OBE on 0.1259 peak horizontal ground accelerations for all Seismic Category I structures.
Almost all Seismic Category I structures are founded on competent rock or concrete fill to rock. Some safety-related electrical manholes o are founded on engineered backfill consisting of offsite borrow or tunnel )
. cuttings. The maximum depth of the backfill to rock for these manholes l
is 18 feet. The manholes are small structures and soil-structure interac-tion effects for these structures is expected to be insignificant. The shear wave velocity of the bedrock is from 8,000 to 10,000 fps, and the compression velocity is 16,500 to 18,500 fps (Reference 1). Consequently, the effects of soil-structure interaction are considered to be negligible for the Category I structures. The ground response spectra used in the design are those recomended in USNRC Regulatory Guide 1.60 scaled to 0.259 for the SSE and 0.1259 for the OBE. Horizontal response spectra used for the SSE design analyses of Category I structures are shown in Figure 1-1. Both modal response spectrum and modal time-history analyses were conducted for the Seabrook Category I structures. In general, the response spectrum analysis results were used for evaluation of the struc-ture seismic loads and stresses while time-history results were used to
! generate in-structure response spectra for the design and evaluation of piping and equipment. Three synthetic time-histories were generated based on the ground response spectra. Comparisons of the ground response spectra generated by these artificial time-history records compared to the design spectra are shown in Figures 1-2 through 1-7 for 7 and 10 percent damping (Reference 1).
- The plant structures and equipment were originally divided into two categories acenrding to their function and the degree of integrity required to protect the public. These categories are Category I and non-Category I. Seabrook Station structures, systems and components ,
l 1-3 l
l l - - - - - - - _ . - -. . _,_ __ ._
important to safety, as well as their foundations and supports, were designed to withstand the effects of an OBE and an SSE, and were thus designated as Seismic Category I. These plant features are those necessary to assure:
- a. The integrity of the reactor coolant pressure boundary,
- b. The capability to shut down the reactor and maintain it in a safe shutdown condition, or
. c. The capability to prevent or mitigate the consequences c' accidents which could result in
. potential offsite exposures, comparable to the gJideline exposures of 10 CFR Part 100.
f Seismic Category I structures include the following:
Containment Structure and Internal Structures Containment Enclosure Building Containment Equipment Hatch Missile Shield Containment Enclosure Ventilation Area Control and Diesel Generator Building Control Room Makeup Air Intake Structures Emergency Feedwater Pump Building, Including i Electrical Cable Tunnels and Penetration Areas (Control Building to Containment)
Enclosure for Condensate Storage Tank Fuel Storage Building l . Main Steam and Feedwater Pipe Chase (East),
l Including East Penetration Area Main Steam and Feedwater Pipe Chase (West),
Including Mechanical Penetration Area and Personnel Hatch Area Piping Tunnels l
Pre-Action Valve Building i Primary Auxiliary Building, Including Residual Heat Removal (RHR) Equipment Vault t
s 1-4 l
l t
l
Safety-Related Electrical Duct Banks and Manholes Service Water Cooling Tower, Including Switchgear Rooms Service Water Pumphouse Tank Farm (Tunnels), Including Dikes and Foundations for Refueling Water Storage Tank (RWST) and Reactor Makeup Water Storage Tank Waste Processing Building In addition, the foundations and supports for all Category I structures and supports are designed for Category I criteria. Some structures such as the waste processing building, although classed as
~
Seismic Category I, are not essential to the safe shutdown of the reactor and were, therefore, not evaluated in these analyses.
Structures, equipment, and components which are important to plant operation, but are not essential for preventing an accident which would endanger the public health and safety, and are not essential for the mitigation of the consequences of these accidents are classified as non-Category I structures. An example of a non-Category I structure is the turbine building. Several structures, such as the service and circulating water punphouse, are classed as partially Category I.
Seismic Category I structures are sufficiently isolated or protected from non-Category I structures to ensure that their integrity is maintained for the design SSE. Several non-Category I structures are designed
. against collapse onto Category I structures due to SSE loads.
For the most part, results of existing analyses and evaluations of structures and equipment for the Seabrook plant were utilized in this
- study. As part of this evaluation, some limited analysis based on original design analysis loads was conducted to determine the expected
. seismic capacities of the important structures. The approach adopted in this study was to determine the median factor of safety and its statisti-cal variability which exists for the SSE in order to estimate the expected response at failure and, hence, the median peak ground acceleration for i
1-5
l failure. It is known that earthquakes with only one or two high accelera-tion spikes are not as damaging to structures and equipment as longer duration earthquakes with multiple peaks at close to the maximum accelera-tion level. The reason for this is that the shorter duration earthquakes do not have sufficient energy content to develop resonances. For this reason, the fragility evaluations described in this report are keyed to a mean peak acceleration belonging to an earthquake of short duration that develops narrow band response spectra.
l An evaluation of the individual important structures and some of
, the equipment was conducted in this manner. However, much of the piping and equipment were evaluated on the basis of a number of generic catego-ries. Although inelastic energy dissipation is included in determining the factors of safety, no nonlinear analyses have been conducted for either the structures or equipment for Seabrook, and all evaluations were based on elastic analysis and load distributions.
These results can be used together with the estimated annual frequency of occurrence of various ground motion levels to determine the frequency of seismic-induced failure for each safety-related structure or component in the plant. In the total study, these conditional component f ailure frequencies are used with systems models to determine the proba-bility of core melt frequencies and radioactive release frequencies.
These results are then combined with the results of the consequence analysis to determine the risks induced by earthquakes.
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! l-13 l
1
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- 2. GENERAL CRITERIA FOR DEVELOPMENT OF MEDIAN SEISMIC SAFETY FACTORS The factor of safety of a structure or component is defined as the resistance capacity divided by the response associated with the Safe ;
. Shutdown Earthquake (SSE) of 0.25g peak acceleration. The development of seismic safety factors associated with the SSE is based on consideration of several variables. The variability of dynamic response to the specified acceleration and the strength capacity of the structure or equipment component are the tw' o basic considerations in determining the variability in the factor of safety. Several variables are involved in determining both the structural response and the structural capacity, and each such variable, in turn, has a median factor of safety and variability associated with it. The overall factor of safety is the product of the factors of safety for each variable. The median of the
~
overall f actor of safety is the product of the median safety factors of all the variables. The variabilities of the individual variables also combine to determine that of the overall safety factor.
Variables influencing the factor of safety on structural capac-ity to withstand seismic induced vibration include the strength of the structure compared to the design stress level and the inelastic energy absorption capacity (ductility) of a structure or its ability to carry load beyond yield. The variability in computed structural response for a l given peak free-field ground acceleration is made up of many factors.
The more significant factors include variability in (1) ground motion and the associated ground response spectra for a given peak free-field ground
' - acceleration, (2) energy dissipation (damping), (3) structural mede Wiq, (4) muhod of analysis, (5) combination of modes, (6) combination of earthquake components, and (7) soil-structure interaction. The ratio between the median value of each of these factors and the value used in design of the Seabrook plant and the variability of each factor are 2-1
quantitatively estimated in Chapters 4 and 5 for various structures and components. These estimates are based on available test data for Seabrook, limited analysis, and engineering judgment and experience in the analysis of nuclear power plants and components.
2.1 DEFINITION OF FAILURE In order to estimate the median factor of safety against the
- structure or component failure for the SSE peak acceleration (0.259), it is necessary to define what constitutes failure.
2.1.1 Seismic Category I Structures For purposes of this study Category I structures are considered to fail functionally when inelastic deformations of the structure under seismic load are estimated to be sufficient to potentially interfere with the operability of safety-related equipment attached to the structure.
These limits on inelastic energy absorption capability (ductility limits) l chosen for Category I structures are estimated to correspond to the onset of significant structural damage. For many potential. modes of failure, this is believed to represent a conservative bound on the level of inelas-tic structural deformation which might interfere with the operability of l components housed within the structure. It is important to note that considerably greater margins of safety against structural collapse are believed to exist for these structures than many cases reported within this study. Thus, the conditional probabilities of failure for a given free-field ground acceleration presented in this report for Catetory I structures are considered appropriate for equipment operability limits and should not necessarily be inferred as corresponding to structure collapse.
l 2.1.2 Seismic Category I Equipment and Piping Piping, electrical, mechanical and electro-mechanical equipment vital to a safe shutdown of the plant or mitigation of an accident are considered to fail when they will no longer perform their designated functions. Rupture of the pressure boundary on mechanical equipment is 2-2 l
also considered a failure. Therefore, for mechanical equipment, a dual failure definition exists: failure to function and pressure boundary rupture. Depending upon the equipment type, one or the other definition will govern. For active equipment, the functional failure definition will usually govern as equipment pressure boundaries are usually very conservatively designed for equipment such as pumps and valves. For piping, failure of the support system or plastic collapse of the presure
. boundary are considered to represent failure. The inelastic energy
- absorption limits (ductility limits) associated with these failure modes have been conservatively estimated in order to define the margins of safety.
2.1.3 Non-Category I Structures In the Seabrook plant, no components identified as important to safety are located within non-Category I buildings. The service and circulating water pump house and waste processing building are partially Category I. The entire pump house structure was evaluated as part of this PRA. However, no equipment essential for the safe shutdown of the reactor is located in the waste processing building, and its capacity was not determined. Non-Category I buldings are separated from the Category I structures by seismic gaps. The non-Category I structures were either designed to Category I criteria or designed so that their failure would not damage any Category I structures. Since it was judged that failure of non-Category I buildings would not affect the seismic capacities of the Seabrook Category I structures, fragility evaluations were not conducted for the non-Category I buildings as part of this evaluation.
2.1.4 Non-Seismic Category I Equipment and piping i Failure of Non-Seismic Category I piping, electrical, mechanical and electro-mechanical equipment is defined as for Category I equipment; i.e., failure to perform its intended function or failure of the pressure boundary.
f 2-3
2.2 BASIS FOR SAFETY FACTORS DERIVED IN STUDY There was a general lack of detailed information available for this study on seismic fragility of specific Seabrook structures and equipment. This occurs because existing codes and standards do not require determination of ultimate seismic capacities, either for structures or equipment qualified by analysis, or for equipment or components qualified by testing. Therefore, most median safety factors, estimates of variability, and conditional frequencies of failure estimated in this study are based on existing analyses and qualified engineering judgment and assumptions. Limited additional analyses were
, conducted to evaluate the expected failure capacities of the important structures. The additional analyses were based on the original design analyses which were available, however.
2.2.1 Structural Response and Capacity The results from existing dynamic analyses of the important
, structures (Reference 1), which were used in the design, were extensively l
used in this study. These were_ supplemented as required to provide estimates of load redistributions resulting from localized failures, etc.
Levels of conservatism associated with the method of analysis used in i design were estimated such that safety factors reflecting this analysis could be estimated for the building structures and for the seismic excitation of equipment mounted within the building.
Detailed structural design calculations were not reviewed, but the design criteria used in design as defined in the FSAR (Reference 1)
! were reviewed. Some ultimate load capacity analyses were conducted which served as a basis for estimating the median factor of safety on struc-tural resistance to the SSE.
2.2.2 Seismic Category I Piping and Equipment Response and Capacity For most of the safety-related equipment, information on j analysis methods was available in sumary form in the FSAR. Seismic l
response information was obtained from the seismic qualification l
l 2-4 4
- _- - . - - - - _ - . . _ . _ - . . _ _ , ~
reports for specific components. In some cases such as for piping, only the seismic analysis requirements and stress acceptance criteria were known. Safety factors for response and structural or functional capacity were estimated from existing information. No new analyses were conducted.
In-structure response spectra for all Category I structures were generated during the design process. From these typical floor response spectra and knowledge or estimates of equipment fundamental frequencies.
- an estimate is made of the peak equipment response. The peak equipment response estimate is then compared to the dynamic response or equivalent
, static coefficient used in design to determine a median safety factor on response.
Capacity factors are derived from several sources of informa-tion: plant-specific design reports, test reports, generic fragility test data from military test programs and generic analytical derivations of capacity based on governing codes and standards. Two failure modes are considered in developing capacity factors for piping and equipment:
structural and functional. Equipment and piping design reports delineate -
stress levels for the specified seismic loading plus normal operating conditions. Where the equipment fails in a structural mode (i.e.,
pressure boundary rupture or loss of support), the median. capacity factor and its variability are derived in the same manner as for structures considering strength and energy absorption (ductility). In cases where equipment must function, the capacity factor is derived by comparing the equipment functional failure (or fragility) level to the design level of seismic loading. There are some fragility test data on generic classes of equipment that have been utilized in hardened military installations.
The equipment was off-the-shelf without special shock resistant design ,
and is similar to nuclear power plant equipment. These data provide estimates of the fragility levels, and thus, safety factors can be developed for the specified design earthquake. Fragility levels are not normally determinable from equipment qualification reports, but the achieved test levels can be utilized to update generic fragilities '
derived from the military data.
2-5
. - -_. - _ _ - - .J
2.3 FORMULATION USED FOR FRAGILITY CURVES Seismic-induced fragility data are generally unavailable for specific plant components and are certainly unavailable for the specific Seabrook structures. Thus, fragility curves must be developed primarily
- from analysis combined heavily with engineering judgment supported by very limited test data. Such fragility curves will contain a great deal of uncertainty, and it is imperative that this uncertainty be recognized in all subsequent analyses. Because of this uncertainty, great precisior.
in attempting to define the shepe of these curves is unwarranted. Thus, a procedure which requires a minimum amount of information, incorporates uncertainty into the fragility curves, and easily enables the use of engineering judgment, was used in this study.
The entire fragility curve for any mode of failure and its uncertainty can be expressed in terms of the best estimate of the median ground acceleration capacity, , times the product of random variables.
Thus, the ground acceleration, A, corresponding to failure is given by:
A=dcR 'U (2-1) are random variables with unit median representing in which cR and cU the inherent randomness (failure fraction) about the median and the uncertainty (probability) in the median value, respectively. Equation 2-1 enables the fragility curve and its uncertainty to be represented as shown in Figure 2-1; i.e., as a set of shifted curves with attached I uncertainty levels. Thus, it is assumed that all uncertainty in the l fragility curves can be expressed through uncertainty in the median alone.
Next, it is assumed that both cR and EU are lognormally
, , distributed with logarithmic standard deviations of SR and 80 '
respectively. The advantages of this formulation are:
- 1. The entire fragility curve and its uncertainty can be expressed by three parameters - A, BR .
and S U . With the very limited available data on 2-6
fragility, it is much easier to only have to estimate three parameters rather than the entire shape of the fragility curve and its uncertainty.
- 2. The formulation in Equation 2-1 and the lognormal distribution are very tractable mathematically.
Another advantage of the lognormal distribution is that it is easy to convert Equation 2-1 to a deterministic composite "best estimate"
. fragility curve (i.e., one which does not separate out uncertainty from underlying randomness) defined by:
A=Ec c (2-2) where cC is a lognormal random variable with unity median and logarithmic standard deviation SC given by:
SC= ER+SU (2-3)
This composite fragility curve (shown in Figure 2-1) can be used in preliminary deterministic safety analyses if one only needs a "best estimate" on failure fraction and does not desire an estimate of uncer-tainty. In this study, the guidelines used to estimate th'e values of ,
SR and SU for each variable affecting A were based on considering the ,
inherent randomness, RB , to be associated with the earthquake character-istics themselves, and BU to be associated with other lack of knowledge.
I Thus, such variability as resulting from earthquake response spectra shapes and amplification, earthr'uake duration, numbers and phasing of peak excitation cycles, etc., together with their contributions to repeated structure ductility and response characteristics is attributed to
. , randomness. In general, randomness is not considered to be significantly reduced from additional analysis or test based on current state-of-the-art techniques. Uncertainty is considered to result primarily from analytical modeling assumptions and other lack of knowledge concerning variables such as material strength, damping, etc. which could in many cases be reduced by additional study or test.
2-7 4
The lognormal distribution can be justified as a reasonable distribution since the statistical variation of many material properties (Reference 3 and 4) and seismic response variables may reasonably be represented by this distribution (Reference 5). In additicn, the central limit theorem states that a distribution consisting of products and quotients of distributions of several variables tends to be lognormal even if the individual distributions are not lognormal. Use of this distribution for estimating failure fractionc on the order of one percent
- or greater is considered to be quite reasonable. Lower fraction esti-mates which are associated with the extrane tails of distribution must be considered less accurate.
l l
Use of the lognormal distribution for estimating very low failure fractions of components or structures associated with the tails of the distribution is considered to be conservative because the low-frequency tails of the lognonnal distribution generally extend farther from the median than actual structural resistance or response data might indicate since such data generally show <ut-off limits beyond which there is essentially zero failure fraction. The degree of conservatism introduced
, into the probability of release is dependent not only on the conservatism in the fragility description, but also on the seismic hazard description i
at low seismic levels. If the seismic hazard for low seismic input levels is large enough, it is apparent that very low level earthquakes can govern the seismic-induced release. This is considered unrealistic for engi-neered structures and equfpment found in nuclear power plants. Structures and equipment are subjected to low level dynamic loads from a number of sources including wind on a repetitive basis which have never been known 1
to produce nuclear power plant structural failures. Similarly, for low level earthquakes, it is expected that below some threshold, there is
. virtually no chance of failure due to seismic excitation. riaterial strength data, for instance, normally does not fall to very low values compared to the median value, but instead normally exhibits some lower bound (Reference 3 and 4). Other variables, such as damping, also ,
l indicate both lower and upper bounds which are not zero or infinite.
2-8
Extensive studies have been conducted to c%velop response spectra from available earthquake records and while dispersion exists about the median values, spectra with essentially zero or infinite response do not occur (Refere;ce 5). For these as well as other variables contributing to the seismic fragility of a given structure or component, it is apparent that some lower and upper bound cutoffs on the tails of the dispersion exist.
,. Since the overall fragility curves are based on a combination of these
. variables, it is expected that a threshold exists below which no failures l will occur. This is supported by experience. Although quantitative data is lacking, this threshold value is expected to be at approximately minus two lognomal standard deviations for the median curves using the "best estimate" or composite fragility variability. The composite lognormal standard deviation, CS , is used for the basis of the cut-off rather
- than randomness or uncertainty since the composite value combines the j effects of both dispersions. However, it is also apparent that some variability should be associated with the cut-off.
Essentially no data ar'e available to establish the distribution of this variability or its range. A lognormal distribution is, therefore, assumed consistent with the majority of the other variables encountered in the PRA. The following approximation is recomended for establishing the cut-offs for the various fragility curves:
i l i The cut-off on the lower tails of the median (50 percentile) i fragility curve should be:
Aco = A exp (-28C. )
4 whereIco is the cut-off on the median curve, is the median peak
- ' ground acceleration for failure, and BC is the composite lognormal
~
standard deviation.
i !
2-9 i
I
- - . - - - . . _, ,. _.,-r,.._y, ,-, - , .___, ..,. .,,,,, ,_.,,,, ._,_m.,.,,_,..,,, ,_,_,.,.-,_.,____._._._....__,..-,___.,_..--,,.-mr_,
l The cut-off for the lower tails of the other fragility curves should be:
Aco = Aco ,exp(-ysC/1.65) _
where x is the ratio of the deviation divided by the standard deviation.
For instance, for the median curve, x = 0; for the 25 percentile curve,
. x = -0.67; for the 5 percentile curve and below, x = -1.65; and for the 95 percentile curve and above, x = 1.55.
It is recommended that the cut-off on the upper tails be established as +3a C for all fragility curves. Similarly, for fragility curves involving only uncertainty, it is recomended that the cut-offs be set at -38 Ufor the lower bound and +3SU for the upper bound, respectively.
Some characteristics of the lognormal distribution as applied to seismic capacities are discussed in Appendix A of this report.
2.4 DESIGN AND CONSTRUCTION ERRORS An inadequate data base exists upon which to determine explicitly the contributions of design and construction errors to most Seabrook structures and equipment seismic capacities. In one exception to this, the possibility of a large throughwall flaw was considered as a lower bound for generic piping. In general, for a plant as new as Seabrook with current design and QA procedures, the possibility that design and construction errors which can affect the seismic capacity of a component may exist is considered remote. Although some discrepencies have been identified and others may be in the future, these items have been
, - - modified as necessary or shown to have no safety implications. Thus.
l these items are r.ot expected to significantly affect the seismic capacity of the equipment or structures af ter they have been identified. However, there is a possibility that unidentified design and construction errors may exist which can affect the seismic capacity.
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- 3. DIFFERENCES BETWEEN CRITERIA USED FOR DESIGN OF SEABROOK AND PARAMETERS USED IN THE EVALUATION OF THE SEISMIC CAPACITY The seismic design of the Seabrook structures and equipment was based for the most part on currently accepted methodology and criteria in conformance with NRC licensing requirements. Essentially, the same cri-teria exist for the design and analysis of the Seabrook Nuclear Steam
. Supply System (NSSS) and non-NSSS Category I components. These criteria and methods together with the design codes in use at the time of the
. design form a conservative design basis and ensure that substantial fac-I tors of safety are introduced at various stages in the design procedure.
The exact magnitude of many of these safety f actors is still a matter of considerable discussion. Nevertheless, in order to establish a realistic value of the actual seismic capacity of a structure or equipment compo-nent, the amount of conservatism along with its variability must be estab-lished as accurately as possible. In this chapter, the design basis of the most important parameters affecting seismic capacity are identified,
~
and the general methods used in obtaining more realistic salues associated i with very high seismic response levels are discussed. The detailed determination of these parameters is described in Chapters 4 and 5 for i structures and equipment, respectively. The estimated seismic capacities of the most probable failure modes are also developed in Chapters 4 and 5.
Tim genersi approach used in the evaluation of the Seabrook seismic capacity is to develop the overall factor of safety associated
- with each important potential failure mode. Based on the governing design parameters, a median seismic capacity is then obtained in termr, of
~ '
l some representative seismic input such as free-field acceleration. The overall factor of safety is typically composed of several important j contributions such as strength, allowance for inelastic energy dissipation 3-1
(ductility), and differences in median structure response compared to design values resulting from such parameters as earthquake character-istics, damping, and directional load components.
3.1 STRENGTH The design strength of a structure or component is typically determined from applicable codes and standards such as the ACI building
- codes for concrete or the ASME boiler and pressure vessel code. Inherent in these design codes is a factor of safety on material strength. Some-times this factor is known reasonably accurately, such as the design allowable being one-half the ministan yield strength or some similar rela-tionship. At other times, it is less well defined or may be a function of the geometry or other physical characteristics of the component such as for reinforced concrete shear walls. For metal structures and compo-nants, the safety f actor included in the codes is usually fairly accu-rately known as are the relationships between minimum and mean or median strengths. For concrete structures, the factor of safety is normally less accurately known. In this" case, the strength of the element is a function of the concrete strength, the amount and strength of the reinforcing steel, and the configuration of the element including the element geometry and reinforcing steel details. In establishing the strength and seismic capacity of concrete components, the results of
{ concrete compression tests and reinforcing steel strength and elongation tests provide a valuable basis for establishing the element strength capacity. However, the increase in concrete strength with age together with the specific details of the element must also be considered. These effects are discussed in more detail in Chapter 4 for structures and
- Chapter 5 for the piping and equipment.
I~ 3.2 DUCTILITY In order to establish realistic seismic capacity levels for most structures and components, an assessment of the inelastic energy absorp-tion must usually be considered. Exceptions to this are some modes involving brittle f ailure, functional f ailure or elastic buckling.
3-2 4 ,
- . - _ . _____ _ _ _ _ _ - _ _ _ . . . _ _ _ _ _ _ . . _ _ _ _ . . _ _ _ _ . _ _ __._..--_.a
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However, most failures due to seismic response involve at least some degree of yielding. This is true of reinforced and prestressed concrete as well as the somewhat more ductile metal structures and components.
Consideration of structure ductility typically results in the ability of the structure to withstand greater seismic excitation than would be predicted using linear elastic techniques. In the design analysis of the Seabrook structures, all design analyses were based on linear elastic analyses. No nonlinear analyses of the structures were conducted. Although inelastic analysis would be desirable in order to more accurately quantify the inelastic effects, the dissipation of inelastic energy may be adequately accounted for without the time and expense of performing nonlinaar analyses. This can be accomplished by the use of the ductility-modified response spectrum approach (References 6 and 7) together with a knowledge of the elastic model results and the expected ductility ratios of the critical elements of the structure or component. This approach is based on a series of nonlinear time-history analyses using single-degree-of.-freedom models with various nonlinear resistance functions and levels of damping. For different levels of ductility, the reduction in seismic response for the nonlinear system compared to the equivalent elastic system response is calculated. This reduction has been shown to be a function of the frequency and damping of the system as well as the ductility. However, a reasonably accurate i assessment of the reduction in response of a structure or component can be made provided the results of the elastic analysis are available and a realistic evaluation of the system ductility can be made.
i . 3.3 EARTHQUAKE DURATION The earthquake duration assumed for design was from 10 to 15 seconds. This was used primarily to detemine the number of response cycles for the OBE for ASME Code fatigue analysis. No fatigue analysis is required by the ASME Code for the SSE since it is a Service Level D 3
(faulted) condition. For the non-NSSS equipment, 20 stress cycles were assumed for the analysis. For the NSSS equipment, an evaluation was i
3-3
conducted which indicated no more than 8 peak acceleration (above 90 percent of the maximum) cycles occurred for flexible equipment, and no more than 3 peak cycles for rigid equipment. For design purposes, 10 cycles were used for flexible equipment, and 5 for rigid equipment. For the magnitude earthquakes expected at the Seabrook site,10 to 15 seconds of strong motion excitation is considered extremely conservative. Con-sequently, fewer cycles would be expected than calculated for the design j
, earthquake since the design earthquake accelograms are approximately 15 l* seconds for all three components of motion.
The seismic capacity of structures is affected by the duration and resulting number of strong motion cycles since the expected available ductility of the controlling structural elements increases as the number of cycles is reduced. Element ductilities for concrete elements were developed from limited test data available in the literature. These data I are considered applicable for reversed load cycles representative of earthquakes with magnitudes of approximately 6.2 or greater. Estimated effects in terms of equivalent ductility from lower magnitude earthquakes, such as are expected for the Seabrook site, have been developed from
! limited nonlinear analyses of representative nuclear power plant struc-tures and observed damage to conventional structures. These effects are i included in the seismic capacity evaluations conducted for Seabrook.
I 3.4 SYSTEM RESPONSE A number of parameters must be evaluated when considering the expected system response near failure compared to the design ccnditions.
Among these are the expected compared to the design earthquake character-istics, directional combinations, system damping, load combinations, and i system modeling approaches and assumptions. In addition, the duration of the earthquake must be considered since short duration earthquakes do not possess sufficient energy to fully excite the structural systems. Some ~
of these parameters may be essentially median centered and introduce little change in the expected seismic capacity while other design criteria l
may be quite conservative. Several of the more important parameters l
l t
3-4 l
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_ . . _ . _.,.... _ _ _ _ --_ ___ __..__._ .. _ _ -._. _._. ___ __ _ ._.~._. _ __ _ . _ _ . _ _ . - _ _ - . _ _ .
h f required in evaluating the system seismic response are discussed below.
The f actors of safety associated with these parameters are developed in the following chapters for the specific failure modes identified.
3.4.1 Earthquake Characteristics The Seabrook Seismic Category I structures are founded on rock or concrete poured against rock. Equipment within the structures were i designed for an SSE of 0.25g defined by the USNRG Regulatory Guide 1.60 free-field ground response spectra shown in Figure 1-1. These spectra
! were developed from a number of earthquakes that occurred on both soil ll
- and rock sites. They were developed for design purposes and are smoothed envelopes of the actual earthquake spectra from which they were developed. Site-dependent spectra are not available for Seabrook. The spectra chosen as site-specific spectra were derived from Reference 8. A
- comparison between these spectra and the design spectra for different i
damping values is presented in Figure 3-1. Below 28 Hz, the design spectra accelerations exceed those of the site-specific spectra. At frgquencies below 8 Hz, the difference is substantial. Most of the structures analyzed in this report have their fundamental frequency below 8 Hz. These effects, together with the higher expected damping associated with seismic response levels at or near failure, provide a significant factor of safety compared to the Seabrook design criteria.
3.4.2 System Damping Damping values used for the SSE design analysis of the Seabrook plant are shown in Tables 3-1 and 3-2 for the non-NSSS and NSSS design, respectively. The values are typically the same as those specified in
. USNRC Regulatory Guide 1.61 (Reference 9), with the exception of the 3SSS primary coolant loop system and large piping. The values of damping specified in Regulatory Guide 1.61 are normally considered to be somewhat l ,
conservative. Therefore, the design values are considered to be quite .
conservative, particularly at response levels of structures and equipment near failure levels. Very little actual test data exist at failure 3-5 r-v +
--wy- =e ,i-9 --m-e-+,.-m,p -py, p gm-
l levels, particularly for structures. However, the damping values recom-mended in References 6,10 and 11 are considered representative. These damping values for structures and equipment at or near yield are shown in Tables 3-1 and 3-2 in comparison with those used for design analysis for the SSE. In accordance with the recomendations in Reference 10, the lower levels of the pairs of values shown in Tables 3-1 and 3-2 are con-sidered to be lower bounds while the upper levels are considered to be
. essentially average values. The values of damping used for this evalua-tion were taken from Tables 3-1 and 3-2 assuming the upper level to be a median value except in the case of piping. Review of piping damping values derived from experiments support the use of higher values
, (Reference 11). Composite modal damping ratios were developed based on l strain energy weighting for structures constructed of different materials.
For subsystems constructed of different materials, the damping ratio of
< the lowest damped material was used for all modes. This introduces some -
l conservatism in the response results for these subsystems. This was
(
considered in the evaluation of their seismic capacities.
~
l 3.4.3 Load Combinations l Load combinations on which the design of the Seabrook station Category I structures were based are shown in Tables 3-3 through 3-6
! (Reference 1). These load combination criteria define a large number of I load combinations that must be considered in design. For the reactor i
building structure and much of the equipment contained within the reactor building, these load combinations include a combination of a loss of coolant accident (LOCA) and the SSE loads. Random LOCA events have an
- l extremely low frequency of occurrence as do seismic events such that the frequency of both events occurring simulaneously is so small that their inclusion is judged to be not important to the risk analysis results.
Seismically-induced LOCA loads are considered to have a higher
! probability of occurring coincident with earthquake loads. Thus, the effects of seismic-induced LOCA loading on the LOCA mitigating systems (high pressure injection and low pressure injection, including their 3-6 4
. - _ _ , _ _ _ ~ . _ _ . _ _ , _ _ , . . _ _ . . . , _ _ , . _ _ _ _ _ . . _ _ _ _ _ _ _ _ _ . . _ . . _ _ . . . _ _ _ . _ _ _ _ . . . _
electrical power, control systems and supporting structures) were investigated. The conclusions of this investigation regarding the Seabrook plant are:
- 1. For PWR's, the major portions of each of the
- mitigating systems is contained within the auxiliary building. Equipment within the auxiliary building is not affected by transient LOCA loads; thus, their corresponding fragilities I
are unaffected by a LOCA. The only portion of these mitigating systems which would be affected
. by a LOCA are piping runs from the containment penetrations to the NSSS connections. Generic piping fragility derivations for Seabrook have resulted in very high capacities; thus, transient LOCA lohds much higher than the SSE would be required in order to significantly alter these piping fragility levels. During a small LOCA, the high pressure injection line would be slightly affected by LOCA loading. The change in fragility is judged to be inconsequential to the overall safety injection system fragility which is greater than 2.0 g's.
- 2. The steam generator has the lowest median ground
, acceleration capacity (1.71g's) of any of the Seabrook components which could cause a large 1.0C A . The only active components of the required LOCA mitigation system which are located in containment are motor-operated valves, which are qualified for a LOCA environment. It is judged that the valves subjected to LOCA effects will have greater than 2.0g seismic capacity.
- 3. In addition to the equipment adjacent to the NSSS system. LOCA's affect both the containment building and the containment internal structure.
The LOCA increases the internal pressure stress on the containment building, while transient LOCA loads are distributed directly into the internal 9 structure. The ground acceleration capacity -
levels for both of these structures has been l
shown in Section 4.2.1 to be 7.6g's or greater.
3-7 l
l e
. _ - - , . . _ . , _ _ _ _ . - . _ - . _ _ _ - . _ _ _ . - _ . . . - - _ , . , _ _ . . . _ . . _ - _ _ . , _ . , . _ . . , _ _ . . _ - . . _ _ . _ _ _ - ~ .__
These capacity levels are so great that the inclusion of seismic-induced loads does not significantly alter the computation of the capacity.
Therefore, special consideration of LOCA loading combined with seismic events in the development of fragility descriptions is not required for the Seabrook risk models.
. r 3.4.4 Modal Combination The Seabrook seismic design analysis was conducted on the basis of loads determined by the square-root-of-the-sum-of-the-squares (SRSS) method for both the NSSS and non-NSSS structures and equipment. Closely spaced modes were considered in accordance with USNRC Regulatory Guide 1.92 (Reference 12). The grouping method of R,eference 12 was used for closely spaced modes for the non-NSSS structures while the double sum method was used for the NSSS analysis. Both of these methods are considered to give approximately median centered results. Although some frequency shifts are expected as structures approach failure, these shifts in frequency are normally not large unless very high ductility
~
ratios exist. Also, the relatioqship between loads developed from individual modes may be ecpected to change once nonlinear response levels are reached. In the absence of a nonlinear analysis, the changes in the
, modal ratios are unknown. For the seismic evaluation of Seabrook, it is assumed that the load response relationships between modes does not change significantly once the structure reaches the yield point. For systems where most of the response results from one mode, this assumption intro-duces negligible possibility for error. For systems with a large number i . ,'
of modes with significant response levels, some additional uncertainty is introduced. The resulting assumed dispersion is discussed in Chapter 4 i for structures.
1 3-8 4
l l
l 3.4.5 Combination of Responses for Earthquake Directional Components In the Seabrook plant design analyses of both the NSSS and non-NSSS systems, the responses for the earthquake directional components were combined for structures and equipment by the SRSS of the vertical and two horizontal components. This is in accordance with USNRC Regulatory Guide 1.92 (Reference 12). This approach requires that the effects of two horizontal directional responses be combined with the vertical response, but does not require that the maximum response in each
~
direction occur at the same instant as the maximisn response in the other two directions. The SRSS combination of the three orthogonal components of response is considered to be essentially median centered so that no
, significant factor of safety exists for this aspect of the design. Some variability exists, however, and this is included in the evaluation of the individual structures and components.
3.4.6 Structure Modeling Considerations l In the seismic design analysis of Seabrook, multi-degree-of-freedom lumped mass models were developed for r.ast of the seismic Category I structures. For structures where the centers of mass and centers of rigidity were not coincident, three-dimensional models were developed in order to compute the torsional response of the structure.
Due to the symetry of the containment building shell, a two-dimensional l
lumped mass was developed. Because of the high relative stiffness of the foundation rock, fixed-base models were used. Since the containment internal structures are not connected to the containment shell structure above the base slabs, the high rock foundation stiffness allowed uncoup-ling of the shell and internal structures into two separate models.
. Separate models were also developed for the vertical analytical models for many of the structures. The models developed for the Seabrook buildings are considered to be able to characterize the seismic response consistent with the current state-of-the-art of seismic analysis and do not appear to introduce significant degrees of either conservatism or non-conservatism into the results.
3-9 4
, - - . - - _ , _ . - - ., _ _ - _ . , _ . . _ _ _ - _ _ _ _ _ _ _ . _ _ . _ _ _ _ _ _ _ - . . _ . _ ~ , . _ . , _ . _ _ _ _ . . . . _ - - - . , , _ . - -
Some aspects of the analysis procedure yield variations which can be quantifiably assessed compared to the design results. For instance, the increase in the actual concrete strength compared to the design values may be used to evaluate the change in stiffness and hence the change in frequencies of the concrete structures compared to the design values. The modified frequencies may, in turn, be used to re-evaluate the modal responses. Another area where modified responses are
, considered is in the load distribution through diaphraps containing l "
relatively large cut-outs. Neglecting the cut-outs typically overesti-mates the stiffness of the diaphragm and may consequently overestimate the seismic load calculated. For a single stick horizontal model, typically no diaphragm loads are computed. However, an estimate of the I
stiffness of the diaphrap with cut-outs, and, if necessary, in the f ailed condition, may be used to redistribute the seismic loads if redundant load paths are available, and hence provide a more realistic ultimate seismic capacity. The details of these and similar evaluations necessary to account for changes between parameter design values and values more representative of seismic response levels near failure are discussed in the following chapters.
3-10
TABLE 3-1 COMPARISON OF CRITICAL DAMPING RATIOS FOR DIFFERENT MATERIALS (NON-NSSS)
Percent Critical Damping
. Structure or Component Seabrook SSE Fragility Evaulation*
Design (Ref.1) (Ref. 6, 10)
Equipment and large-diameter piping systems, pipe diameter
, greater than 12 in. 3 5 Small-diameter piping systems, diameter equal to or less than 12 in. 2 5
~
i Welded steel structures 4 5 to 7 Bolted steel structures 7 7 to 15 Prestressed concrete structures 5 7 to 10**
Reinforced concrete structures 7 7 to 10 i
I
- Lower values are considered to be approximately lower bounds; upper values are considered to be essentially median centered
- With no prestress left 3
0 3-11
TABLE 3-2 COMPARISON OF CRITICAL DAMPING RATIOS FOR DIFFERENT MATERIALS (NSSS)
Percent Critical Damping Seabrook SSE Fragility Structure or Component Design (Ref.1) Evaluation (Ref.6.10)
Primary and large coolant piping loop 12" (system components Diameter and 4 5 greater)
Small piping 2 5 Welded steel structures 4 5 to 7 Bolted and/or riveted steel structures 7 10 to 15
- Lower values are considered to be approximately lower bounds; upper values are considered to be essentially median centered.
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TABLE 3-3 DESIGN FACTORED LOAD COMBINATIONS FOR CONTAINMENT INVOLVING EARTHOUAKES (REF. 1) i (1) U =
1.0D + 1.3L + 1.0To + 1.5Eo + 1.0Ro + 1.0P y (2) U = 1.0D + 1.0L + 1.25P, + 1.0Ta + 1.25Eo + 1.0Ra + 1.0Rrr +
1.0Rrj + 1.OR m (3) U =
1.0D + 1.0L + 1.0To + 1.0Ess
- I'Wt + 1.ORo + 1.0P y (4) U = 1.00 + 1.0L + 1.0P, + 1.0T, + 1.0Ess + 1. ora + 1.0Rrr +
1.0Rr j + 1.0R m where D = Dead Load L = Live Load T = Normal Temperature o _
Eo
= Operating Basis Earthquake Ro
= Normal Pipe Reaction ,
Py = Pressure Variation P, = Accident Pressure T, = DBA Temperature Ra = DBA Thermal Pipe Reaction Rrr = Reaction of Ruptured High Energy Pipe Rrj = Jet Impingement Loads l
Rm= Impact of Ruptured High Energy Pipe l
3-13 I
e TABLE 3-4 ALLOWABLE STRESSES AND STRAINS IN THE CONTAINMENT STRUCTURE (REF. 1)
Concrete Factored Loads CompressiveStress(ff=3000 psi)
Membrane 1.8 ksi Membrane + Bending 2.25 ksi CompressiveStress(ff=4000 psi)
Membrane 2.4 ksi Nembrane + Bending 3.0 ksi Shear Stress
- Radial CC-3431.4.1 Tangential CC-3431.5.1 Punching CC-3421.6 &
Code Case N-219 Reinforcing (fy = 60 ksi)
Tensile S, tress 54 ksi
- Allowable shear stress does not exceed 40 psi and 60 psi for load combinations (2) and (4) respectively from Table 3-3.
I
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3-14 4
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' TABLE 3-5
- DESIGN FACTORED LOAD COMBINATIONS FOR CONCRETE INTERNAL STRUCTURES INVOLVING l EARTHOUAKES (Ref.1)
(1) U = 1.40 + 1.7L + 1.9E, (2) U = 1.05D + 1.28L + 1.28T, + 1.43E, + 1.28R g (3) U = 1.0D + 1.0L + 1.25P, + 1.0T, + 1.25E, + 1.0R, +
1.0R rr + 1.0Rrj + 1.0R rm (4) U = 1.0D + 1.0L + 1.0To + 1.0E ss + 1.0R, (5) U = 1.0D + 1.0L + 1.0P, + 1.0T, + 1.0E3 , + 1.0R, +
1.0R rr + 1.0Rrj + I*0R rm + 1.0M where: M = Internal Missile Loads All other loads defined as in Table 3-3 Stresses in accordance with ACI 318-71 Criteria 4
w 0
3-15
-[
TABLE 3-6 DESIGN FACTORED LOAD COMBINATIONS FOR CONCRETE STRUCTURES OTHER THAN CONTAINMENT & INTERNALS FOR LOADS INVOLVING EARTHQUAKES (Ref.1)
(1) U = 1.4D + 1.7L + 1.9E, + 1.7H + 1.9H, (2) U = 1.05D + 1.28L + 1.28T, + 1.43E, + 1.28H +1.43H, + 1.28R, (3) U = 1.2D + 1.9E, + 1.7H + 1.9H, (4) U = 1.00 + 1.0L _+ 1.25E, + 1.0H + 1.25H, + 1.25P, +
1.0R, + 1.0Rrj + 1.OR, + 1.0R rr + 1.0T, (5) U = 1.0D + 1.0L + 1.0T, + 1.0Ess + 1.0H + 1.0H, + 1.OR, (6) U = 1.00 + 1.0L + 1.0E ss + 1.0H + 1.0Hs + 1.0P, + 1.0R,
+ 1.0Rrj + 1.0R rm + 1.0R rr + 1.0M + 1.0T, where: H = Lateral Earth Pressure H, = Earth Pressure due to OBE Hs = Earth Pressure due to SSE All other loads defined in Tables 3-3 and 3-5 Stresses in accordance with ACI 318-71 Criteria l r e
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- 4. STRUCTURES In this chapter, the median factors of safety and logarithmic standard deviations for the important structures are developed. Based on these factors of safety, median acceleration levels associated with seismic failure are presented. For these structures, existing dynamic models and response analyses were used to determine the median factors of safety and logarithmic standard deviations for each of the variables involved in calculating building response. All original analyses were based on lineer response model results, but some seismic loads were modified to more closely approximate the expected inelastic response at the high acceleration levels expected for failure.
4.1 MEDIAN SAFETY FACTORS AND LOGARITHMIC STANDARD DEVIATIONS As discussed in Section 2.3, the seismic fragilities of struc-tures and components are described in terms of the median ground accelera-tion, d, and random and uncertainty logarithmic standard deviations, SR and Sg . In estimating these fragility parameters, it it computationally attractive to work in terms of an intermediate random variable called the factor of safety F. The factor of safety is defined as the ratio of the ground acceleration capacity, A, to the Safe Shutdown Earthquake (SSE) acceleration used in design. For equipment and structures qualified by analysis, it is easier to estimate the median factor of safety, F, and variability parameters, B R and eg , based upon the original SSE stress analysis than it is to directly estimate the fragility parameters. Thus, d=F*ASSE (4-1) e From the existing analyses of the important structures together with a knowledge of the deterministic design criteria utilized, median factors of safety associated with the SSE ground acceleration of 0.25g i can be estimated. These are most conveniently separated into those 4-1 L
f actors associated with the seismic strength capacity of the structure, the inelastic energy absorption factor, the factor associated with earthquake duration, and those f actors associated with the expected building response.
The factor of safety for the structure seismic capacity consists of the following parts:
- 1. The strength factor Fs, based on the ratio of actual member strength to the design forces.
l l *
- 2. The inelastic energy absorption factor, Fy ,
related to the ductility of the structure.
- 3. The earthquake duration factor FED, used to account for the expected duration compared to that implicitly assumed in the determination of F.y i
- Associated with the median strength factor Fs,and the median ductility I
factor, F y, are the corresponding logarithmic standard deviations, 6, and 89 . The structure strength factors of safety and logarithmic standard deviations vary from structure to structure and according to the different failure modes within a given structure. Factors of safety for the most important modes of failure are summarized in subsequent sections.
i The earthquake duration factor, FED, as stated in Section 3.3, reflects the additional capacity due to the shorter duration and fewer number of strong motion cycles present in the Seabrook median expected earthquake as compared to the earthquake which would generate the number
, of cycles used in the determination of Fy. Limited experimental data were available for the determination of Fy . For the most part, the available data for reversed-load tests of reinforced concrete shear walls consisted of three to six or more cycles. As discussed in Section 4.1.3, for a median magnitude earthquake expected at the Seabrook site, one to three strong motion cycles are expected. FED accounts for the expected increase in F ydue to the above-mentioned items.
4-2
The factor of safety, FR, related to building response is determined from a number of variables which include:
- 1. The response spectra used for design compared to the median centered spectra for rock sites from multiple seismic events.
- 2. Damping used in the analysis compared with damping expected at failure.
- 3. Modal combination methods.
- 4. Combination of earthquake components.
. 5. Modeling accuracy.
- 6. Soil-structure interaction effects.
Based on the characteristics of lognormal distribution, median factors of safety and logarithmic standard deviations for the various contributing effects can be combined to yield the overall estimates. For instance, the capacity factor of safety of a structure, F e ,p, is obtained from the product of the strength, ductility, and duration factors of safety which, in turn, may include effects of more than one variable.
F cap = Fs xFy xF ED (#~2)
The methods of determining these safety factors are discussed in the following sections. The logarithmic standard deviation on capacity, ecap, is found by:
S cap = (82+62 + 6'ED I4~3)
. As discussed in Section 2.3, the logarithmic standard deviations are ,
I composed of both an inherent randonmess and uncertainty in the median value.
4-3 f ..
i pv Median factor of safety F, and variability, SR and SU '
estimates are made for each of the parameters affecting capacity and t response. These median and variability estimates are then combined using
! the properties of lognormal distribution (Equations 2-10 and 2-11) and in accordance with Equations 4-2 and 4-3 to obtain the overall median factor of safety and variability estimates required to define the fragility curve for the structure.
- For each variable affecting the factor of safety, the random variability, sR, and the uncertainty, 8 ,0 must be separately estimated. The differentiation is based on the following guidelines.
! Essentially, SR represents those sources of dispersion in the factor of safety which cannot be reduced by more detailed evaluation or by gathering more data. 6R is due primarily to the variability of an earthquake time history and, therefore, of a structure's response when the earthquake is only defined in terms of the peak ground acceleration. SU represents
! those sourc6s of dispersion which could be reduced only through better I understanding or more knowledge of the behavior of the item. 8 0 is due to such items as our lack of ability to predict the exact strength of l materials (concrete and steel) and of structures (shear walls and i diaphragms); errors in calculated response due to inaccuracies in mass and stiffness representations as well as load distributions; and use of engineering judgment in the absence of plant specific data on fragility levels.
The variables which are related to the structure response can be f
grouped into several main categories for which factors of safety and the logarithmic standard deviations may be combined in a similar manner.
Each of the f actors presented in Chapter 3 will now be discussed in more detail. To help in understanding the methodology involved, failure modes l for the control and diesel generator building will be derived at the same time.
l l
I l 4-4 l
1 4.1.1 Structure Capacity ,
The primary lateral load carrying systems of the Category I structures that were analyzed are of reinforced concrete construction with the exception of the condensate storage tank which is constructed from stainless steel. The containment buildings and the containment enclosure buildings are cylinders with hemispherical heads. The condensate storage tank enclosure is also a cylinder, but is open at the top. The rest of the structures are composed of shear walls (typically two to three feet thick) with diaphragms supported by structural steel beams and girders. All structures are separated from adjacent buildings
- by standard three-inch gaps except for the control building and the diesel generator building which share a comon wall.
Since the lateral load carrying systems are composed of rein-forced concrete, the two most important material strengths to determine are those for the concrete and for the reinforcing steel. The determina-tions of these strengths are presented in the following two sections.
4.1.1.1 Concrete Compressive Strength The evaluation of the strength of most concrete elements, whether loaded in compression or shear, is based on the concrete compressive strength, f .e Concrete compressive strength used for design is normally specified as some value at a specific time from mixing (for example, 28 or 90 days). This value is verified by laboratory testing of mix samples.
The strength must meet specified values allowing a finite number of failures per number of trials. As previously stated, there are two major f actors which justify the selection of a median value of concrete strength above the design strength, l
- 1. To meet the design specifications, the contractor I attempts to create a mix that has an " average" I strength above the design strength.
- 2. As concrete ages, it increases in strength.
f 4-5
For the Seabrook structures, results of concrete compression tests are available (Reference 13). These tests consist of 28 day strength tests and are divided into tests for the various mix numbers. Table 4-1 shows the average values and numbers of samples taken from each mix.
As concrete ages, its strength increases. This must also be accounted for in determining the median strength compared to the design strength. Figure 4-1 from Reference 14 shows the increase of the concrete compressive strength with time assuming the concrete poured in the field is adequately represented by the curve designated as " air
. cured, dry at test." At 28 days, the concrete has a relative strength of 50 percent which approaches 60 percent asymptotically. The factor relating the strength of aged concrete to the 28 day strength is, therefore, 1.2. No information is available on the standard deviation expected for aging. The estimated logarithmic standard deviation for aging is 0.10. For a small standard deviation, the median may be taken as approximately equal to the mean. Thus, the factor relating median compressive strength including aging effects to design compressive strength for the Seabrook struc'tures varies depending on the design strength and mix number. The average factor for 4000 psi concrete is '
approximately 1.65 and 1.9 for the 3000 psi concrete.
Other effects which could conceivably be included in the concrete strength evaluation include some decrease in strength in the in-place condition as opposed to the test cylinder strength, and some
- increase in strength resulting from rate of loading at the seismic response frequencies of the structures. Although experimental data on these effects are extremely limited, that which is available would tend to indicate these effects are relatively small and of the same order ,
! magnitude. Since the two effects are opposite, they were neglected.
f I
4-6
4 4.1.1.2 Reinforcing Steel Yield Strength Grade 60 reinforcing steel was used throughout the construction of the structures. Information pertaining to the strength of the rein-forcing steel used in the structures was supplied by Reference 15. A sunnary of this information is presented in Table 4-2. All structures other than the containment building used No. 11 bars or smaller. There-fore, for ease in calculating member strengths, an overall median yield strength encompassing these sizes (No. 4 through 11) was calculated. It is also presented in Table 4-2.
- Two other effects must be considered when evaluating the yield strength of reinforcing steel. These are the variations in the cross-sectional areas of the bars and the effects of the rate of loading. A survey of information (Reference 16) determined that the ratio of actual to nominal bar area has a mean value of 0.99 and a coefficient of varia-tion of 0.024. The same reference notes that the standard test rate of loading is 34 psi /sec. Accounting for the rate of loading anticipated in seismic response of structures results in a slight decrease in yield strength of reinforcing steel in tension. This effect is neglected in concrete compression.
4.1.1.3 Shear Strength of Concrete Walls Recent studies have shown that the shear strength of low-rise concrete shear walls with boundary elements is not accurately predicted by the ACI 318-77 code provisions (Reference 17). This is particularly true for walls with height to length ratios in the order of 1 or less.
Barda (Reference 18) determined that the ultimate shear strength of
. Iow-rise walls they tested could be represented by the following i relationship:
l 4-7 i
Vu*Vc+Vs
= 8.3 -3.4 ( (h,/t,-0.5) + ouy f (4-4) where:
, Vu = Ultimate shear strength, psi Vc = Contribution from concrete, psi Vs = Contribution from steel reinforcement, psi fc = Concrete compressive strength, psi hw
= Wall height, in tw = Wall length, in ou = Vertical steel reinforcement ratio fy = Steel yield strength, psi The contribution of the concrete to the ultimate shear strength of the wall as a function of h,/1, is shown in Figure 4-2. Also shown in Figure 4-2 are the applicable test values (References 18 through 21) and the corresponding At 318-17 formulation. The tests included load reversals and varying reinforcement ratios and h,/t w ratios. Web crushing generally controlled the failure of the test specimens. lasting l
, was performed with no axial loads, but an increase in shear capacity of N/4twh was recomended, where N is the axial load in pounds, and h is the wall thickness in inches.
l l
! 4-8 l
i i
The contribution of'the steel to the ultimate shear strength according to ACI 318-77 is:
V,=pfhy (4-5) where ph = horizontal steel reinforcement ratio.
Furthermore, one of the conclusions reached by Oesterle (Reference 21) is that for low-rise shear walls (specifically wh w /t = 1),
vertical steel has no effect, and the entire contribution to shear strength is due to the horizontal steel.
In order to estimate the effects that the horizontal and vertical steel have, the steel contribution to wall shear strength was determined from test values for the range of 0.5 < hw /tw < 2. Test data from the above references were used. The effective steel shear strength was assumed to be in the form: ,
(4-6)
Yse = AVsu + BVsh where A, B are constants and V
su " Duy f = vertical steel contribution to shear strength '
Vsh " Chyf = horizontal steel contribution to shear strength The constants A and B were then calculated assuming the concrete contribution to the ultimate strength is given as shown in Equation 4-4.
l Based on the results of this evaluation, the constants A and B can be shown to be:
4-9
(
a
A=1 B=0 h,/ t, < 0.5
=-2.0(hjt,)+2.0 = 2.0 (h,/ t,) - 1.0 0.5 s h,/ t, 5 1.0
=0 =1 1.0 5 hg t, and the median ultimate shear strength is given by:
Vu*Vc+Yse (4-7)
=8.3(-3.4 h,/t,-0.5 + 4t,h
- A se y where c3e = Acu + 0# hwith A and B determined as shown above.
Based on an evaluation of the same experimental data, the logarithmic standard deviation was calculated to be 0.15.
4.1.1.4 Example of Shear Wall Failure in Shear A plan view of the control and diesel generator building is shown in Figure 4-3. The east wall governs the shear capacity of this building.
This wall is two feet thick. The vertical reinforcing consists of No. 11 bars every 12 inches on each face while the horizontal reinforcing con-sists of No.10 bars every 12 inches on each face. The design concrete
, compressive strength was 3000 psi. From Table 4-1 and using the aging factor of 1.2 mentioned in Section 4.1.1.1, a median concrete compressive strength of 5700 psi was calculated. From Table 4-2, the reinforcing steel yield strength is 69.8 ksi. The wall length is 90 feet. The effective wall height is the base moment for this wall divided by the wall's base shear. This gives an effective height of 42.5 feet. Since .
4-10
h,/tw < 0.5, the steel's contribution to shear strength is assumed to come only from the vertical reinforcing steel. Thus, the concrete shear strength is:
v c = 8.3 /5700 -3.4 /5700 - 0.5 = 634 psi = 91.3 ksf
. The steel shear strength is:
v = 108.9 ksf 3= l, h"f (69.8 ksi) 144 The contribution due to axial load is typically very small and is usually neglected since the load is rarely known. Therefore, the ultimate shear capacity of this wall is:
Vu=(91.3+108.9)(2')(90')=36,000k The shear due to the design SSE is 7210 k. This gives a shear strength factor of 5.00. The variability due to randomness is negligible because the strength of the wall is independent of seismic events. The variability due to uncertainty is estimated to be 0.16 based on a steel strength variability of 0.01, a concrete compressive strength variability of 0.17, and an equation variability of 0.15. Thus, 0.04*(108.9)+( (91.3) 2 t 6 + 0.15 = 0.16 U. (108.9 + 91.3)2 i Note that the concrete variability is divided by two. This is due to the concrete strength being a function of the square root of the concrete compressive strength.
4-11
4.1.1.5 Strength of Shear,)lalls in Flexura Under In-plane Forces Data on reinforced concrete shear walls failing in flexure under in-plane forces can be found in Reference 21. Equations found in Reference 20 may be used to calculate the moment capacity for walls without chord steel. However, chord steel can be accounted for by increasing the depth from the extreme compressive fiber to the neutral axis to account for the yield strength of the tensile chord steel. The
. compression chord steel is neglected since it is near the neutral axis, and its effect on the moment capacity is small. The total moment capacity of reinforced concrete shear walls in flexure under in-plane
~
forces is then:
w I M=
s 1+ 1- +A chfy d~
(4-8) where:
c = Depth to neutral axis from extreme compression fiber As = Area of distributed steel Ach
= Area of chord steel t
w = Wall length fy = Steel yield strength l
N = Axial load d = Distance from the extreme compressive fiber i to the centroid of tensile chord steel 8, = Ratio of depth of equivalent rectangular concrete stress block to depth to neutral axis (c) i i 4-12 1
l
4.1.1.6 Example of Shear Wall Failure in Flexure The west wall of the control and diesel generator building governs the flexural capacity of this structure. This wall is two feet thick. The vertical reinforcing consists of No. 9 bars every 12 inches on each f> The overturning moment will be resisted by not only the west wair. but also by some of the north and the south walls acting as flanges. Reference 22 presents a procedure by which the effective flange width may be estimated. For this problem, an effective flange width of 13 feet (exclusiveofthewestwallthickness)iscalculated. The neutral axis is contained within one of these flanges while the other flange provides the chord steel. The vertical steel reinforcement in the flanges is No. 9 bars every 12 inches on each face.
From balancing the tensile and compressive forces, c = 1.645'.
The values used for the rest of the variables are as follows:
As
= /
2(1.00in2ft)(90')=180.0in2 Ach
=
2(1.00in2/ft) (13') = 26.0 in2 t.
y = 90' fy = 69.8 ksi N = 0 (as with the shear strength) d =
90'-f(2')=89' 6
3
= 0.85 - 0.05 (5700 - 4000)/1000 = 0.765 Using these values in Equation 4-8, the ultimate moment capacity for this wall is 718,000 k-ft. The moment derived from the design SSE is 246,000 k-ft. The flexural strength factor is thus 2.92. The variability _
due to uncertainty is estimated to be 0.18 based on a steel strength l
l variability of 0.04, an equation variability of 0.10, and a flange-width uncertainty of 0.15. Thus, i
4 13
2 S
U= 0.04 * + 0.10 + 0.15 * = 0.18 Note that there is no variability due to concrete strength in this ey ,
If the concrete strength was only half the calculated median value of 5700, the effect would be to roughly double the depth to the neutral axis, c. This, in turn Would lower the ultimate moment capacity by about one percent. Therefore, the concrete strength has no significant impact on the ultimate moment capacity.
4.1.2 Structure Ductility A much more accurate assessment of the seismic capacity of a structure can be obtained if the inelastic energy absorption of the structure is considered in addition to the strength capacity. One tractable method involves the use of ductility modified response spectra to determine the deamplification effect resulting from the inelastic energy dissipation. Early studies indicated the deamplification factor was primarily a function of the ductility ratio, 9, defined as the ratio of maximum displacement to displacement at yield. More recent analytic studies (Reference 7) have shown that for single-degree-of-freedom systems with resistance functions characterized by elastic-perfectly plastic, bilinear, or stiffness-degrading models, the shape of the resistance function is, on the average, not particularly important.
However, as opposed to the earlier studies, more recent analyses have shown the deamplification factor is also a function of the system damping. For systems in the acceleration region of the spectrum (i.e.,
approximately 2 Hz and above), Figure 4-4 from Reference 7 shows the deamplification function for several damping values as a functon of the ductility ratto.
4-14
Actual tests on walls show that there is a wide spread in the amount of ductility present in any given wall. For example, Reference 23 performed tests with load reversals and obtained ductilities of 2 to 10.
Reference 24 allows a ductility of four for reinforced concrete shear walls in a bearing wall system. A ductility of 4 was used in these analyses for simple one- or two-story systems, and 3.5 was used for more complex systems. When a ductility of 4 was used, the variabilities were estimated to be (By)g = 0.15 and (8 9 )U = 0.45. When a ductility of 3.5 was used, (89 )R = 0.15 and (89 )U = 0.40 were estimated for the variabil-
- ities. The uncertainty was judged to be much higher than the randomness
- because the ductilities are derived for strong motions with at least 3 strong cycles, and it is felt that the ductility will not vary by much for any earthquake meeting these requirements.
4.1.2.1 Example of inelastic Enerqy Absorption Factor Figure 4-4 presents the ductility, 9, versus the response spec-trum reduction factor, 49 , for different damping values. The inelastic energy absorption factor is then 1/4 9 For the control and diesel generator building, a duett 11ty of 3.5 and a damping value, s, of 10 percent of critical were used. From Figure 4-4:
q = 3.003-0.30 = 3.00 (10)-0.30 = 1.504 p = q + 1 = 2.504 r=0.48e-0.08=0.48(10)-0.08=0.399 Fy = -1= (pu-q)r , '2.504(3.5) - 1.504 .
0.399 = 2.21 ,
y .
l I
4-15 9
l It can be shown that the effect of (89)R ""d (8W)U '" #W **"
represented by:
' = "
I8 \
pu-q \ 9/R (8 4, q'* ,), p % (c u)u Foru=3.5,(8)R9 =0.15and(8)U9 = 0.40. Therefore, (8, ), = 0.07 and (8,y) = 0.19. There are no other factors to take into account for randokness, so g o = (8, )R = . . w, e is an equadon uncertainty of an estimated 0.10 that must be combined with (8, )U produce 8 g Therefore, the total uncertainty on Fy is 8
0= 0.19* + 0.10* = 0.21 4.1.3 Earthauake Duration -
The basis for this factor, as previously stated, is the expected increase in capacity due to the median expected earthquake being of less ,
duration and lower energy content than that from which the ductility f actor was developed. The median expected earthquake has a Richter magnitude of 5.8. Earthquakes of this size typically have a duration of ,
7 to 9 seconds and 1 to 3 strong motion cycles with an average of 2 cycles. It is felt that the lower bound on the duration factor is 1.0 I and the upper bound is 3.0. This means that the expected median i
earthquake would have to be scaled by a f actor between 1.0 and 3.0 in order to have the same dmage potential as an earthquake with 3 to 6
- cycles with both earthquakes having the same unscaled peak ground I
acceleration. This factor has an estimated median of 1.4. Note that !
I, I 4 16
,I
with lower and upper bounds of 1.0 and 3.0, respectively, this distribu-tion is not lognormal. However, since the region between 1.4 and 3.0 is of no concern, the variabilities were determined as if it is a lognormal distribution, at least around the region 1.0 to 1.4. It was assumed that there is only a five percent chance of the actual f actor being less than 1.0 and that this prediction is made with 95 percent confidence.
Therefore, 1.0 = 1.4 exp(-1.65Pg)] [exp(-1.6589)} = 1.4 {exp(-1.65(ep+8 U ) f Thus, eg + s u= 0.20. This factor is mainly an earthquake-dependent
. factor, but with a good deal of uncertainty, so 8R was set to 0.12 with 80 set to 0.08.
This factor is only used for f ailures that are ductile in nature. Brittle failures rely only on the acceleration reaching a given level. This level can theoretically be obtained in both large and small magnitude events.
4.1.4 Spectral Shape. Campina. and Modeling Factors As previously discussed, the important Seabrook structures were designed using the ground response spectra shown in Figure 1-1. For the design SSE, seven percent of critical damping was used for the reinforced concrete structures. For the reinforced concrete comprising the lateral load carrying structures for Seabrook, ten percent of critical damping is considered to be the median value expected at response levels near failure (Reference 10). The frequencies of interest for the Seabrook structures are in excess of 4 Hz. As is evident from Figure 31, the response of these structures using the ten percent damped median centered response spectrum for rock sites is less than the seven percent damped design spectum at all frequencies below 27 Hr. For the Seabrook structures, only the frequencies were known to that the spectral shape f actor of safety for the individual structures was calculated based on 4 17 i
m the fundamental frequency. Basing this factor on the ratio of the spectral acceleration of the design spectrum to that of the site-specific spectrum at the fundamental frequency is usually sufficient inasmuch as almost all of most structure's response is due to vibrating at the fundamental frequency. The spectral shape factor of safety is represented by:
5 0c= 7%
p$$ " 5g (4-g) c = 101 where 50 c = 7% represents the 7 percent damped design spectral acceleration and SM 10% represents the estimated spectral acceleration associated with thhm=ediansite-specificresponsespectrumfor10percentdamping.
In computing the spectral shape factor of safety, it is convenient to combine the damping and ground response spectrum effects.
In the development of logarithmic standard deviations on spectral shape, however, it is informative to consider the damping effects separately.
This implies a factor of safety of unity on damping alone since it has already been included in the factor of safety on spectral shape.
The logarithmic standard deviation on spectral acceleration, og may be estimated from References 8 and 10. Reference 8 provided the mean site specific spectrum and associated variabilities for five percent damping. It also presents a procedure by which mean spectra at other damping values could be calculated, but does not give the variabil-ities for these other spectra. These variabilities were estimated with the aid'of Reference 10.
J f
4 18 I
The deviation on spectral acceleration resulting from damping, 8;, can be estimated from:
6 as an
[IM C*") (4-10) c*10%)
where Sg , 7g is the spectral acceleration from the median site-specific spectrum at seven percent damping, and SM = 10% * 'E'" * ""#'
tionfromthetenpercentdampedmediansfte-specificspectrum. Seven percent damping is estimated by Reference 10 to be one standard deviation below the median damping value of ten percent.
The modeling factor of safety is usually taken to be unity.
Among the items that would change this are:
- 1. Story stiffnesses that did not compensate for large openings and are, therefore, too stiff: and
- 2. Diaphragm stiffnesses used to connect the various sticks in a multi stick and lumped mass model that also do not account for openings.
Variability in modeling predominantly influences the calculated mode shapes and modal frequencies. Since the concrete strength and, consequently, the stiffness of the structures is above the design values, calculated frequencies would be espected to be somewhat less than actual values, at least for low to moderate levels of response. At response levels approaching f ailure, sof tening of the structures due to concrete cracking occurs, and for structures analyzed using uncracked section properties, some decrease in the actus) frequencies compared to the calculated values is espected. As can be seen from Figure 3 1 for frequencies in the 5 to g Ha range and greater than 20 Hz, the response 4 19
accelerations are fairly constant so that a small shif t in frequency does not result in much change in amplitude. Between 8 and 20 Hz, an increase in frequency results in a decrease in amplitude. Calculated frequencies and mode shapes were assumed to be median centered unless material properties used in the original analyses differed from the material
,! properties calculated from test data enough to change the calculated frequencies by at least 15 percent. The new frequencies were calculated
, based on the new material properties. The mode shapes were assumed to
- stay the same.
Modeling uncertainties from both the mode shapes and modal frequencies enter into the uncertainty on calculated modal response as t
defined by og . Thus, f g = (8k + 8kF (4-11) where 833 and 8gg are estimated logarithmic standard deviations on structural response of a given point in the structure due to uncertainties in mode shape and due to uncertainties in modal frequency, respectively.
Based upon experience in performing similar analyses, ogs is estimated to be about 0.10. The modal frequency variability shifts the frequency at which spectral accelerations are to be determined, so thatt l'ef , f )
8 8gg = in (412)
(3Hff/ g where fg is the median frequency estimate, and f8 is the 84 percent exceedence probability frequency estimate. The logarithmic standard deviation on frequency is estimated to be approximately 0.20 for the structures.
4 20
l 4.1.4.1 Example of Spectral Shape. Damping, and Modeling, Factors The fundamental frequency of the control and diesel generator l buildingfortheN-Sdirectionwascalculatedtobe5.69Hz(Reference 1).
The spectral amplification factor from the seven percent damped design spectrum at 5.69 Hz is 2.42. The spectral amplification factor from the ten percent damped site-specific spectrum at 5.69 Hz is 1.77. These amplification f actors are obtained from Figure 3-1. Their ratio gives F$$ . 1.37. The variability due to randomness is estimated to be 0.24
- at this frequency. The variability due to uncertainty is estimated to be
% of 8 g, or 0.08.
For damping, the spectral acceleration from the seven percent
- e damped site-specific spectrum is 1.96. Therefore, 8; = in = 0.10 8; needs to be broken down into 8g and BU which are judged to be equal to each other. Since f isg an SRSS combination of eg and BU '
en.8u' I 8c " 0 07 The model of the control and diesel generator building is judged ,
to provide a median centered representation of the structure response, so Fg = 1.0. To calculate gge , f6(Equation 412)is6.95Hz. The
. spectral amplification factor from the ten percent damped site specific spectrum at 6.95 Hz is 1.81. Thus, sg7 e in = 0.02 i
4 21 l
l s
i When combined with E gg, Sg=V0.022 + 0.10* = 0.10 4.1.5 Modal Combination The seismic design analysis of Seabrook structures was performed by response spectrum analysis; therefore, phasing of the individual modal responses was unknown. Most current design analyses are normally con-
. ducted using response spectra techniques. The current recomended prac-tice of the USNRC as given in Regulatory Guide 1.92 (Reference 12) is to combinemodesbythesquare-root-of-the-sum-of-the-squares ($RSS). This was the methodology used in the Seabrook analyses. Many studies have been conducted to determine the degree of conservatism or unconservatism obtained by use of SRSS combination of modes. Except for very low damping ratios, these studies have shown that SRSS combination of modal responses tends to be median centered. The coefficient of variation (approximate logarithmic standard deviation) tends to increase with increasing damping ratios. Figure 4-5(takenfromReference25)shows the actual time history calculated peak response versus SRSS combined modal responses for structural models with four predominant modes. Based upon these and other similar results, it is estimated that for ten percent structural damping, the SRSS response is median centered.
Therefore. FMC = 1.0 for Seabrook structures.
Where individual modal responses are known, the absolute sum of these responses can be used to estimate the coefficient of variation. .
The absolute sum is an upper bound considered to be three standard deviations above the median SRSS response. Where the individual modal responses are not known, past experience with structures where one mode predominates indicates that the coefficient of variation is on the order of 0.05.
4 22
4.1.6 Combination of Earthquake Components The design of the Seabrook structures was based on the current recomended practice of combining the responses for the three principal directions by the SRSS method. Alternatively, it is recomended (Reference 10) that directional effects be combined by taking 100 percent of the effects due to motion in one direction and 40 percent of the effects from the two remaining principal directions of motion.
The effect of SRSS combination of three components compared to
- the direct addition of two depends on the relative magnitudes of the two
. horizontal load components together with the vertical component and the geometry of structure. For instance, if the two horizontal load components are approximately equal, and the vertical component is small, the SRSS method results in an increase in stress of from approximately 40 percent for a square structure to O percent for a circular structure.
Combining the effects by the 100, 40, 40 percent method for the same case results in the same 40 percent increase in stress for a square structure as for the SRSS method and an 1.ncrease of approximately 8 percent for a circular structure, such as the containment structure.
Depending on the geometry of the particular structure under consideration together with the relative magnitude of the individual load .
or stress components, the expected variation in stresses due to the 100, 40, 40 percent method of load combinations is from -5 to +10 percent when compared with the original design method. For shear wall structures
' where the shear walls in the two principal directions act essentially independently and are the controlling elements, the two horizontal loads l
i do not combine to a significant degree except for the torsional coupling.
I Thus, only the vertical component affects the individual shear wall stress. A moderate amount of vertical load slightly increases the I ultimate shear load carrying capacity of reinforced concrete walls.
However, there is an equal probability that the vertical seismic
- component will add to or subtract from the deadweight loads at the time of maximum horizontal loads. Thus, while the dead load is usually l
I 4 23 I
.I l
included in the analyses, the vertical seismic component is ignored.
Consequently, the factor of safety is not strongly influenced by the directional component assumptions.
The coefficient of variation is calculated in the same manner as it was for the modal combination factor. The absolute sum of the three components is an upper bound, assumed to be three standard deviations above the median. Lacking the individual components, SR is assumed to be 0.05.
I o 4.1.7 Soil-Structure Interaction Effects Two types of soil-structure effects are considered in the analysis of nuclear power stations. The first involves the variation in frequency and response of the structure due to the flexibility of the soil and the dissipation of energy into the soil by radiation (geometric) damping. For structures founded on competent bedrock such as the Seabrook Category I structures, these effects are usually small and are typically neglected in current design analyses. A second effect is the amplification of the bedrock motion through the soil. Again, for structures founded directly on the bedrock, essentially nu amplification occurs, and the motion is normally specified at the foundation level as was done in the design of the Seabrook structures. Thus, the design of the structures at Seabrook was conducted using current state-of-the-art assumptions and methods of analysis in regard to the soil-structure interaction effects.
One other possible area of concern is the slab uplif t of the structures at high input acceleration levels. For structures founded on competent rock, there is insufficient energy in the low frequency earth-quake waves to sustain overturning motion of the structure at the very I long response periods required to overturn an auxiliary building or containment structure. At the frequencies of maximum input energy l
content, although a very small amount of uplift may occur, the direction of input motion is reversed before any significant rocking motion can l
4-24 l i 1
occur. So long as significant rock or concrete crushing does not occur, relative motion sufficient to cause piping or electrical conduit failure is not considered a possible failure mode. The bedrock at the Seabrook site is considered to be of adequate strength to preclude failures resulting from base slab uplift.
4.2 CONTAINMENT BUILDING The containment building is a reinforced concrete upright cylinder with a hemispherical dome roof. It is supported on a reinforced concrete foundation mat which is keyed into the bedrock by a depression for the reactor pit and by continuous bearing around the periphery of the mat . The inside diameter of the cylinder is 140'-0". The cylinder wall thickness is typically 4'-6" including the Ye" thick liner. The dome thickness is 3'-65/e " including the /"2 thick liner. The foundation mat is 1
153'-0" in diameter with a thickness of 10'-0".
The containment wall is reinforced with No. 18, Grade 60 reinforcing bars at the inner an_d outer faces, each consisting of one meridional and two hoop layers. Below Elevation (-)15'-0", there are two
~
additional meridional layers on the inner face to resist discontinuity moments and radial shears caused by the restraint on the cylinder at the cylinder-base mat connection. There is also an orthogonal set of bars inclined at 45 degrees to the horizontal on the outer face. This set is provided to resist in-plane seismic shear forces and membrane tension from other loads. Where there are large openings in the cylinder, the bars are continued around the openings without interruption. No main l reinforcement is terminated at any opening.
The liner plate is carbon steel conforming to ASME SA 516, Grade
- 60. It is provided with an anchorage system that will maintain leak tightness in the event of accidents. The anchorage system consists of vertical tees spaced every 20 inches around the circumference of the cylinder wall. The webs of the tees are welded to the liner plate with continuous fillet welds.
4-25
A number of structures important to safety are located within the containment. Among the most critical to seismic response are the reactor support system and the primary and secondary shield walls. The internal structures are supported on and anchored to a 4-foot thick fill mat which is not anchored to the containment base mat. The internal structure is separated from the containment wall by a minimum 0.5" gap.
Potential damage due to impact between the internals and the containment wall was considered in this analysis.
4.2.1 Containment Failure Modes The lowest level significant failure mode for the containment structure is failure of the wall due to flexure at Elevation (-)9'-0".
The median peak ground acceleration capacity is estimated to be 7.69 Median f actors of safety and their variabilities are reported in Table 4-3. Four other failure modes were considered: failure due to shear, failure due to impact, and f ailure of the internals due to flexure and shear.
Failure of the containment due to shear is also critical at Elevation (-)9'-0"; however, the structure is stronger in shear than in flexure. The median ground acceleration capacity for shear is estimated to be 8.89 The factors of safety and variabilities for this failure mode are presented in Table 4-4.
Containment impact was evaluated at three locations: one involving the internal structure and two involving the enclosure building. These latter two will be discussed in the section pertaining
, to the enclosure building since any potential damage is to it and not to the containment building. The location of greatest concern for impact between the containment and the internal structure is at Elevation 25'-0" on the south side. There is a 0.5" seismic isolation gap between the two structures at this level. An estimated median peak ground acceleration of 1.0g is required to close this gap. It is expected that higher accelerations will lead to some penetration of the containment wall, 4-26
but that the penetration will be insufficient to cause the liner to break. Spalling off the outer face of the containment wall is not considered possible since the collision velocity and penetration appear to be less than those required to cause spalling.
Failure of the concrete internal structure is expected at higher accelerations than the containment. The internal structure is a massive l . structure with a dead load equal to a third of the dead load of the l containment. The primary shield wall was checked for both flexure and shear. For flexure, the factor of safety on strength is estimated to be 13.6; for shear the factor is 7.78. Note that both of these factors are greater than the strength factor of 7.28 reported in Table 4-3. The other factors for the primary shield wall would be the same as those listed in Table 4-3.
4.3 CONTAINMENT ENCLOSURE BUILDING The purpose of the containment enclosure building is to provide leak protection for the containment and to protect it from certain loads such as wind loads and tornado loads. Due to its presence, there are also no dynamic soil effects on the containment.
The enclosure building has the same shape as the containment building. The inside diameter is 158'-0", and the wall varies in thickness from 15 inches above Elevation 45'-6" to 36 inches below Elevation (-)11'-0". The hemispherical dome roof is 15 inches thick.
The space between the containment and the enclosure building varies from 4'-6" between the cylindrical sides to 5'-6" between the domed roofs.
The enclosure building wall is supported on a spread footing that is 10'-3" wide by 10'-0" deep. This footing abuts the containment base mat,
< but the two are not connected except by water stops. Everywhere else, l
the two structures are separated by at least 3 inch gaps.
4-27 r
l
~ , . - - .-. , - . - - . , _ . - . . - . _ . - _ _ _ - - _ - . _ _ _ . - _ _ _ , , .
I 4.3.1 Containment Enclosure Failure Modes Three failure modes were considered for this structure: failure of the wall due to shear and flexure and f ailure due to impact between this structure and other buildings. The lowest level failure mode is due to impact between the containment building and a concrete shield located i at Elevation 49'-6" on the northwest face of the enclosure building. It is estimated that a median ground acceleration of 4.lg is necessary to initiate spalling of the outside of the containment enclosure. This failure mode is only a local f ailure and will not damage any safety-related equipment. The lowest level significant failure mode is failure of the wall at the base due to shear. The median peak ground accelera-tion capacity for this mode is estimated to be 8.2 9. Table 4-5 presents the f actors of safety and related variabilities for this failure.
Failure due to flexure has a median capacity of 10.4g. The median f actors of safety and their variabilities are listed in Table 4-6. In comparing the flexural capacity of the enclosure building to that of the containment building (Table 4-3), the main difference is in the strength factor. The fact that the strength factor is larger for the enclosure building does not mean that the enclosure building can withstand a greater overturning moment than can the containment building. At the critical sections in both buildings, the containment building can withstand a moment that is 78 percent greater than the maximum moment for the enclosure building. However, the containment building is required to withstand a moment that is 2.54 times the moment required to be taken by the enclosure building. Hence, the strength factor, which is the j ultimate moment capacity divided by the SSE moment, is greater for the e enclosure building by a factor of 2.54/1.78 or 1.43.
Although an impact failure capacity of 4.lg has been discussed, it is possible that due to phasing and directional components, impact will not occur at the shield at Elevation 49'-6". Thus, impact was also checked at Elevation 22'-0" where there are concrete shields on the east 4-28
and west f aces of the enclosure building. A peak ground acceleration of 9.5g is necessary just to close the three inch gap at this level. Impact of the fuel storage building roof at Elevation 84'-0" against the enclo-sure building was also considered. To close the gap at this level, a peak ground acceleration of about 7.2g is necessary. This acceleration, however, is greater than the acceleration needed to cause significant failure in the fuel storage building. Therefore, this impact mode cannot
, occur.
4.4 PRIMARY AUXILIARY BUILDING The primary auxiliary building is a four-story structure which contains such equipment as heat exchangers, pumps, demineralizers, filters, tanks, and ventilation equipment. The residual heat removal (RHR) equipment vault is located no*th of and shares a wall with the primary auxiliary building. The vault contains containment spray pumps, residual heat removal pumps, and heat exchangers. The main portion of the auxiliary building is 79 feet wide,145 feet long, and extends from a maximum of 46 feet below grade to 88 feet above grade.
4.4.1 Primary Auxiliary Buildina Failure Modes l
Two failure modes were considered for this structure: f ailure of the shear walls due to shear and due to flexure. Failure due to flexure is the lowest level significant f ailure mode with a median peak ground acceleration of 2.6g. The median factors of safety and their variabil-ities are given in Table 4-7. This failure mode pertains to the north wall. Since this is the common wall with the RHR vault, this 2.69 acceleration can also be considered as the median capacity for the vault.
Failure due to shear, the other failure mode that was analyzed, is not expected until the peak ground acceleration reaches 4.09 This f ailure mode pertains to the south wall. Table 4-8 presents the safety f actors and variabilities for this failure mode.
4-29 1
t .
4.5 SERVICE WATER PUMPHOUSE AND CIRCULATING WATER PUMPHOUSE
' The service water pumphouse (SWPH) is a two-story structure that contains the service water pumps and screen wash pumps. The reinforced concrete basin under the SWPH is approximately 91 feet wide by 74 feet long. It extends from the operating floor 1 foot above grade to 63 feet below grade. The SWPH itself is approximately 118 feet wide by 78 feet
- - long and extends 28 feet above the operating floor. An electrical equipment room is attached to the west end of the building.
The circulating water pwnphouse (CWPH) is attached to and shares i a foundation with the SWPH. The concrete basin under the CWPH is approximately 110 feet wide by 123 feet long and extends to the same depth as the SWPH basin. The CWPH itself is non-Category. I and consists of a steel frame covered with metal siding. It is approximately 119 feet wide by 123 feet long and extends 28 feet above the operating floor. It contains the circulating water pumps and screen wash pumps.
- Water is supplied to the CWPH by means of an intake tunnel bored through the bedrock at a depth of 150 to 250 feet below sea level. The tunnel is 19 feet in diameter and extends for over three miles. At a minimum, the tunnel is supported by patterned rockbolts. Where more support is needed, due to the presence of faults, diabase dikes, or closely spaced joints, 6- to 8-inch steel ribs at 2.5 to 5 foot intervals i are used. The tunnel connects to a transition structure by means of a i 200-foot vertical shaf t. At this structure, the flow is separated into four sections and channeled into the CWPH basin via a flume.
4.5.1 Service Water Pwnphouse Failure Modes The lowest level significant f ailure mode is flexural failure of the shear walls due to N-S motion. The median peak ground acceleration is estimated to be 2.1g. The median factors of safety and their varia-i.
1 I bilities for this mode are listed in Table 4-9.
C 4-30 1
Due to the number of openings in the SWPH roof, its capacity was also analyzed. It was found to be slightly stronger than the shear walls, requiring a median peak ground acceleration of 2.4g to cause failure.
This type of f ailure is expected to be a local failure. It would not greatly affect the overall building response and would only cause damage to that equipment located innediately beneath falling debris. The median safety factors and variabilities associated with this failure are i . presented in Table 4-10.
The lowest significant failure for the CWPH-flume-transition structure-intake tunnel system occurs in the intake tunnel. The intake tunnel capacity is based on the past behavior of tunnels during the San Fernando earthquake of February, 1971, and during weapon tests at the Nevada Test Site. It is estimated to have a median peak ground accelera-tion capacity of 4.6g with a variability due to uncertainty of 0.39 and a variability due to randomness of 0.49. The median was derived by comparing past ground velocities due to weapon tests at different stress regions in the rock to the damage in nearby tunnels. In a stress region of about 8700 psi, tunnel collapse has been fairly universal; at about 1800 psi, there has been almost no damage. The ground velocities at these points are 50.2 fps arid 10.7 fps. The maximisn ground velocity of l an earthquake was assumed to be 2.5 fps /1.09 The weapon test velocities were divided by two to try to correlate the pulsive blast accelerations to cyclic earthquake accelerations. These scaled velocities were then assumed to be the +2B and -28 bounds on the capacity. The variability due to uncertainty reflects the unknown rock strength while the variabil-ity due to randomness contains the variabilities associated with the 2.5
' fps /1.0g earthquake ground velocity and the factor of two used to scale I the weapon test ground velocities. This capacity is based on the tunnel not having experienced rockfalls or showing signs of creep. If any of these assumptions is not correct, the tunnel capacity could be greatly
! diminished.
i 4-31 l
r v
k
4.6 SERVICE WAT3 C001.ING TOWER
, The cooling tower is a rectangular building approximately 300 by 54 feet in plan, extending 28 feet below grade and rising 57'-6" above grade. It serves as the ultimate heat sink in the event that the cooling water tunnels are rendered inoperative. The cooling tower contains pumps, f ans, and a water distribution system. The water level in the tower is
. 16 feet above grade. The switchgear rooms that are attached to the east and west ends of the cooling tower are approximately 54 by 26 feet in plan and are located below the mechanical equipment rooms. They contain the switchgear, substation, and motor control center for the cooling tower.
4.6.1 Cooling Tower Failure Mode Two f ailure modes were investigated for this structure: flexural failure of the shear walls due to in-plane dynamic loads and due to out-of-plane hydrodynamic loads. Shear failure of the walls does not appear to be likely. Failure of the roof is considered most likely in the area between column lines 0 to K. Failure of the roof in this region will not damage any equipment nor affect the overall capacity of the strJcture. Flexural failure of the outside switchgear walls due to N-S j motion is predicted at a median peak ground acceleration of 2.49 The f actors of safety and their variabilities for this failure mode are presented in Table 4-11.
Failure of the north and south walls above grade (E1. 22'-0")
due to out-of-plane hydrodynamic forces was also analyzed. Depending on how failure is defined, different capacities can be calculated. For instance, if it is vital that very little, if any, water be allowed to leak out, then the capacity must be based on the pressure that causes the wall to crack near grade. This capacity is estimated to have a median peak ground acceleration of 0.74g with su = 0.43 and BR = 0.10. On the other hand, if larger amounts of water can be lost and failure is i defined as the wall reaching its ultimate capacity, then the median peak ground acceleration is expected to be 3.75g with BU = 0.24 and l
l 4-32 1
1 BR = 0.20. Note that if it is possible for the water level to drop to l Elevation 22'-0" without impairing any other safety-related function, then this failure mode becomes imaterial.
4.7 CONDENSATE STORAGE TAhX AND ENCLOSURE The condensate storage tank is a cylindrical stainless steel tank with a diameter of 42'-0" and a height of 42'-7". It has a
~
spherical dome roof that has a 42'-0" radius and rises almost 5'-8" above the tank. The tank walls vary in thickness from 0.500 inch at the base
, to 0.188 inch at the top. The roof is 0.455 inch thick. The bottom plate is 0.25 inch thick. It connects to the tank wall through a 0.800-inch footer plate that extends 4.75 inches inside the tank and 7.25 inches outside the tank. There are 46 anchor bolts spaced typically at 8 degree intervals around the tank. These bolts have a diameter of 1.75 inches and are made from SA 193, Grade B7 steel. The tank's capacity is 400,000 gallons.
The tank is enclosed in a reinforced concrete cylinder, 42'-11" in diameter, 2' thick and 42'-10" high. It surrounds the tank from the base up to the spriigline and provides protection from horizontal tornado-generated missiles. It is capable of retaining the tank contents should the tank be ruptured.
4.7.1 Tank and Enclosure Failure Modes The lowest level f ailure mode is for the welds connecting the l
anchor bolt chairs to the tank to yield in tension; thus, permitting that i ,
side of the tank to lift up and causing the bottom plate to bend. The weld connecting the footer plate to the bottom plate lies between the edge of the tank and the point at which the plastic hinge will form; thus, causing bending across the weld. Since this weld is only a single lap fillet weld, this weld is expected to fail with loss of the tank contents through the base. The median peak ground acceleration for this failure is estimated to be 1.09 Failure of the tank, however, does not mean f ailure of the system since the contents will simply flow into the enclosure. At most, the water level in the tank would drop about 1.5',
4-33 l
e.- a--,--
leaving outlet nozzles still below the water level. Failure of the enclosure is not expected below a peak ground acceleration of 4.29 . The f actors of safety and the variabilities for this failure mode are listed in 'able 4-12. Note that the two factors dealing with ductility, the inelastic erergy absorption f actor and the earthquake duration and cycles factor, are absent from this table. The reason for this is that failure of the enclosure is defined to be that point at which the concrete begins to crack. This point is well below the yield point of the reinforcing.
4.8 CONTROL AND DIESEL GENERATOR BUILDING
. The control and diesel generator bulding is a rectangular struc-tura approximately 233 feet long by 90 feet wide. A 4-foot thick wall separates the 138-foot long control building on the east side from the 95-foot long diesel generator building on the west. The control building has three floors and extends from grade to about 79 feet above grade. It contains the electrical switchgear, motor generator sets, battery rooms, cable spreading room, and the main control room. The diesel generator building has two floors and extends from 36 feet below grade to approxi-mately 59 feet above grade. This building houses diesel fuel storage tanks, the diesel generators, air intakes for the generators and building j ventilation equipment.
4.8.1 Control and Diesel Generator Building Failure Modes Two f ailure modes were checked for this building: flexural failure and shear failure of the shear walls. Flexural failure gives the lower median peak ground acceleration: 3.09 At this acceleration, f ailure is expected in the west wall, which is the outer well of the
, diesel generator building, due to N-S motion. The same motion results in the lowest level shear capacity, 5.29, this time at the east wall, which
- . is the outer wall of the control building. The median factors of safety and their variabilities for the flexural failure are given in Table 4-13; those for the shear failure are presented in Table 4-14.
4-34
Reference 26 provides an analysis of the control room ceiling.
Because the ceiling is wire-suspended, a nonlinear analysis had to be perfomed to account for the fact that the wire can carry only tensile loads. Thus, it is difficult to say, for instance, whether an earthquake with a peak ground acceleration of 10 times the one used in the analysis will generate stresses and displacements that are greater than or less than 10 times the reported values. However assuming the connections are as detailed in Reference 26, it does appear that the ceiling capacity should exceed the lowest calculated structure acceleration of 3.0g.
4.9 FUEL STORAGE BUILDING The fuel storage building is a two-story, square building approximately 98 feet on each side that extends approximately 44 feet below grade and rises 66 feet above grade. Ne building contains the new fuel storage area and the spent fuel pool. The spent fuel pool e),ists entirely below grade and has walls of a minimum thickness of 6 feet. It is lined with stainless steel plates 0.188 inch thick on the walls and 0.25 inch thick on the floor.
4.9.1 Fuel Storage Building Failure Modes l The lowest level significant failure mode for this structure is flexural f ailure of the west wall due to N S excitation. It is expected to f ail at a median peak ground acceleration of 4.99 The median safety f actors and their variabilities are presented in Table 4-15 for this mode. The only other significant f ailure mode that was considered was shear failure of the walls. The lowest level shear failure has a median peak ground acceleration of 6.5g and also occurs to the west wall due to
. N-S motion. The factors of safety and variabilities for this mode are listed in Table 4-16.
4-35
TABLE 4-1 SEABROOK CONCRETE COMPRESSIVE STRENGTHS -
Average Number Design Tested Standard of Mix Class Strength Strength Deviation Samples (psi) (psi) (psi) 4AWR67 (M.53) 3000 5260 487 414 4AWR67 (M.60) 3000 4327 484 405 4AWR67 (M.45) 4000 5859 582 57 4MAWR67(M.45) 4000 6057 576 456 4AWR67 (M.48) 4000 5557 490 54 4AWR67 (M.48)special 4000 5592 466 135 4AWR67 (M.50) 4000 5442 546 1599 4MN67 (M.45) 4000 6369 354 30 4AR67 (M.48)special 4000 5858 376 50 4AS67 (M.429) 4000 5747 406 207 4 MAS 67 (M.44) 4000 5829 452 60
- 4000 5400 ** 78 4AWR67 (M.48) 4AWR67 (M.48)special* 4000 5360 ** 54
- 4000 5560 ** 75 4WR67 (M.50) 4AS67
- 4000 5640 ** 30 (M.429) 1
- These four mixes were specifically for containment. The rest of the mixes were assumed to apply to all other structures.
- Values not reported. A COV (see Appendix A) of 0.08 was assumed.
4-36 I
TABLE 4-2 SEABROOK REINFORCING STEEL STRENGTHS Number Median Yield Standard '
Bar Size of Tests Strength Deviation
. (ksi) (ksi)
- 4 20 70.5 2.18
- 5 12 67.9 2.11
- 6 96 67.9 2.66
- 7 32 70.0 2.45
- 8 128 69.6 2.37
- 9 80 70.3 2.81
- 10 84 70.9 2.65 fil 136 70.6 2.82
- 14 16 73.6 2.32
\
- 18 92 72.3 2.04 Median yield strength of all bars except #14 and #18 = 69.8 ksi Standard deviation of all bars except #14 and #18 = 2.79 ksi l
l 4 37 l
TABLE 4-3 Structure: Containment Building Failure Mode: Flexural Failure of Wall Median Item F.S. B R
8 8 C
U Strength 7.28 0 0.16 0.16 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.46 0.24 0.08 0.25 Damping 1.0 0.08 0.08 0.11 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.93 0.09 0 0.09 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 30.6 0.31 0.32 0.44 Median Acceleration Capacity = 30.6(0.25) = 7.69 TABLE 4-4 Structure: Containment Building Failure Mode: Shear Failure of Wall Median Item F.S. B R
8 U
8 C
Strength 8.36 0 0.16 0.16 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.46 0.24 0.08 0.25 Damping 1.0 0.08 0.08 0.11 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.93 0.09 0 0.09 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 35.1 0.31 0.32 0.44 i
Median Acceleration capacity = 35.1(0.25) = 8.89 4-38 i
l l
TABLE 4-5 Structure: Containment Enclosure Building Failure Mode: Shear Failure of Wall Median Item F.S. B R
8 U
6 C
Strength 8.25 0 0.25 0.25 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14
. Spectral Shape 1.39 0.24 0.08 0.25 Damping 1.0 0.07 0.07 0.10
- Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.93 0.09 0 0.09 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 33.0 0.30 0.37 0.48 Median Acceleration Capacity = 33.0(0.25) = 8.29 TABLE 4-6 Structure: Containment Enclosure Building Failure Mode: Flexural Failure of Wall Median Item F.S. 8 R
8 U
6 C
Strength 10.4 0 0.25 0.25 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.39 0.24 0.08 0.25
(;amping 1.0 0.07 0.07 0.10 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.93 0.09 0 0.09 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 41.6 0.30 0.37 0.48
- Median Acceleration Capacity = 41.6(0.25) = 10.4g 4-39
Table 4-7 Structure: Primary Auxiliary Building Failure Mode: Flexural Failure of Shear Walls Median l
Item F.S. s R
8 U
8 C
Strength 2.32 0 0.18 0.18 Inelastic Energy Absorption 2.21 0.07 0.21 0.22
~
Earthquake Duration and Cycles 1.4 0.12 0.08 0.14
- Spectral Shape 1.44 0.25 0.08 0.26 Damping 1.0 0.08 0.08 0.11
- Modeling 1.0 0 0.10 0.10 1.0 0.05 0 0.05 Modal Combination Combination of Earthquake Components 0.99 0.01 0 0.01 1.0 0 0.05 0.05 Soil-Structure Interaction 10.2 0.30 0.33 0.44 Total ,
Median Acceleration Capacity = 10.2(0.25g)=2.69 Table 4-8 Structure: Primary Auxiliary 8uilding Failure Mode: Shear Failure of Shear Walls Median Item F.S. s R
8 0
8 C
3.66 0 0.16 0.16 Strength 2.21 0.07 ' O.21 0.22 Inelastic Energy Absorption 1.4 0.12 0.08 0.14 Earthquake Duration and Cycles 1.44 0.25 0.08 0.26 Spectral Shape 1.0 0.08 0.08 0.11 Damping 1.0 0 0.10 0.10 -
Modeling 1.0 0.05 0 0.05 Modal Combination 0.99 0.01 0 0.01 Combination of Earthquake Components 1.0 0 0.05 0.05 Soil-Structure Interaction 16.1 0.30 0.32 0.43 Total Median Acceleration Capacity = 16.1 (0.25g) = 4.0g 4 40
.-. - - - -. , , , . . _ ~ - _ _ - - _ _ _ _ _ , _ , ,- . . - - . . _ -,,. ,,. . ..-- ._
Table 4-9 Structure: Service Water Pumphouse Failure Mode: Flexural Failure of Shear Walls Median Item F.S. B 8 8 R 0 C Strength 3.05 0 0.18 0.16 Inelastic Energy Absorption 1.62 0.04 0.14 0.15 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.26 0.10 0.03 0.11
, Damping 1.0 0.01 0.01 0.01 Modeling 1.0 0 0.14 0.14
. Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.97 0.02 0 0.02 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 8.45 0.17 0.29 0.33 Median Acceleration Capacity = 8.45 (0.259 ) = 2.1g Table 4-10
! Structure: Service Water Pumphouse i Failure Mode: Diaphragm Failure Median Item F.S. e p ag s C
1 Strength 4.14 0 0.18 0.16
- Inelastic Energy Absorption 1.32 0.02 0.11 0.11
., Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.26 0.10 0.03 0.11 Damping 1.0 0.01 0.01 0.01 Modeling 1.0 0 0.14 0.14 Modal Combination 1.0 0.05 0 0.05
[
Combinatior, of Earthquake Components 0.98 0.01 0 0.01 l Soil-Structure Interaction 1.0 0 0.05 0.05 l- Total 9.45 0.17 0.27 0.32 l
Median Acceleration Capacity = 9.45 (0.25g) = 2.49 4-41 l
Table 4-11 Structure: Service Water Cooling Tower Failure Mode: Flexural Failure of Shear Walls Median Item F.S. s R
8 U
8 C
Strength 2.23 0 0.19 0.19 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles l.4 0.12 0.08 0.14 Spectral Shape 1.43 0.24 0.08 0.25 Damping 1.0 0.08 0.08 0.11 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.98 0.05 0 0.05 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 9.67 0.30 0.33 0.44 Median Acceleration Capacity = 9.67 (0.259) = 2.4g Table 4-12 Structure: Condensate Storage Tank Enclosure Failure Mode: Flexural failure of Wall Median Item F.S. s R
S U
8 C
Strength 15 0 0.22 0.22 Inelastic Energy Absorption - - - -
Earthquake Duration and Cycles - --
Spectral Shape 1.20 0.06 0.02 0.06 Damping 1.0 0 0 0 Modeling - 1.0 0 0.12 0.12 Modal Combination 1.0 0.03 0 0.03 Combination of Earthquake Components 0.93 0.09 0 0.09 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 16.7 0.11 0.26 0.28 i
Median Acceleration Capacity = 16.7(0.25g)=4.2g i
4-42 l
-w -- , - , - . --- ,,-,, ,-n,-- , - . - - - , , , - - ,.,e, ,,. - - -
Table 4-13 Structure: Control and Diesel Generator Building Failure Mode: Flexural Failure of Shear Walls Median Item F.S. sp s u C Strength 2.92 0 0.18 0.18 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14
. Spectral Shape 1.37 0.24 0.08 0.25 Damping 1.0 0.07 0.07 0.10 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.98 0.05 0 0.05 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 12.1 0.29 0.33 0.44 Median Acceleration Capacity = 12.1 (0.259) = 3.0g Table 4-14 Structure: Control and Diesel Generator Building Failure Mode: Shear Failure of Shear Walls Median Item F.S. s R
8 U
3 C
Strength 5.00 0 0.16 0.16 Inelastic Energy Absorption 2.21 0.07 0.21 0.22 1
Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.37 0.24 0.08 0.25 Damping 1.0 0.07 0.07 0.10 Modeling 1.0 0 0.10 0.10 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.98 0.05 0 0.05 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 20.8 0.29 0.32 0.43 l
Median Acceleration Capacity = 20.8 (0.25g) = 5.29 f
4-43 l
i k
Table 4-15 Structure: Fuel Storage Building Failure Mode: Flexural Failure of Shear Walls Median Item F.S. sg s u
8 C
Strength 5.53 0 0.18 0.18 Inelastic. Energy Absorption 2,08 0.06 0.20 0.21 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.25 0.18 0.06 0.19 Damping 1.0 0.05 0.05 0.07 Modeling 1.0 0 0.13 0.13 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.97 0.02 0 0.02 Soil-Structure Interaction 1.0 0 0.05 0.05 Total 19.5 0.24 0.32 0.40 Median Acceleration Capacity = 19.5(0.25g)=4.99 Table 4-16 Structure: Fuel Storage Building Failure Mode: Shear Failure of Shear Walls Median Item F.S. s R
8 U
8 C
Strength 7.33 0 0.22 0.22 Inelastic Energy Absorption 2.08 0.06 0.20 0.21 Earthquake Duration and Cycles 1.4 0.12 0.08 0.14 Spectral Shape 1.25 0.18 0.06 0.19 Damping 1.0 0.05 0.05 0.07 Modeling 1.0 0 0.13 0.13 Modal Combination 1.0 0.05 0 0.05 Combination of Earthquake Components 0.97 0.02 0 0.02 i Soil-Structure Interaction 1.0 0 0.05 0.05 Total 25.9 0.24 0.35 0.42 Median Acceleration Capacity = 25.9(0.259)=6.5g 4-44 ,
80 70 Air cured, dry at test
, a g ,
e y M
f40 -
- 30 to '
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4 47
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l - 4 Ill- 4 gg i1 1 1 1111 111111111 ll1111111 111111111 1111 litI lill IIII IIll i i LS 1 -
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1 FIGURE 4-4 DEAMPLIFICATION FACTORS FOR ELASTIC-PERFECTLY PLASTIC SYSTEMS IN THE ACCELERATION AMPLIFIED RANGE (FROM REFERENCE 7) 4 48 l
l i
i
i I
' i, Soo ,
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,, t IW SQ
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FIGURE 4-5. HISTOGRAMS OF RATIO OF PEAK RESPONSE TO SRSS COMPUTED RESPONSE FOR FOUR-DEGREE- -
0F-FREEDOM DYNAMIC MODELS (FROM REFERENCE
- 25) ,
4 49
1
- 5. EQUIPMENT FRAGILITY This chapter describes the fragility development for the seismically critical equipment within the Seabrook Nuclear Generating
- Station. Pickard, Lowe and Garrick, Inc. have identified those equipment items which are essential to plant safety during and af ter a seismic event, and a fragility level and associated variabilities are determined for each of these components. Section 5.1 contains a general description of the equipment fragility methodology with a mare in-depth treatment than was provided in Chapter 3. Section 5.2 pres (nts a set of represen-tative example fragility derivations which provide the reader with further insight into the equipment fragility determination process. Section 5.3 presents the resulting equipment fragilities for the Seabrook plant.
5.1 EQVIPMENT FRAGILITY METHODOLOGY Fragility as used in probabilistic seismic safety studies is defined as a conditional probability of failure for a given hazard input. In this case, the fragility of a component or system is defined as the frequency of failure as a function of peak ground acceleration.
The development of these fragility levels combined with a discussion of the available information sources and the selection of equipment categories are the subject of this section.
5.1.1 Fragility Derivation The procedure used in deriving fragility descriptions is similar to that used for structural fragility descriptions, wherein, factors of ,
safety and their variability are first developed for equipment capacity, earthquake duration and equipment response. These three factors, along with the factor of safety on structural response, are then multiplied together to obtain an overall f actor of safety for the equipment item.
FE*FEC FER
- F E0 *F $g ($,g) ,
51 1
F EC is the capacity f actor of safety for the equipment relative to the i floor acceleration used for the design, FER is the factor of safety inherent in the computatien of equipment response, FED is the earthquake duration factor of safety associated with the predicted number of strong motion cicles within a seismic event, and F $g is the factor of safety in the structural response analysis that resulted in floor
~
spectra for equipment design. Sections 5.1.1.1, 5.1.1.2, 5.1.1.3, and 5.'1.1.4 of this report containIa Pure thorough explanation of these four factors'(iEC, FER, FE and Fjg; rispectively). The overall factor of safety, E E,D, is then maltiplied by the reference earthquake 4
peak ground acceleration to obtain fragility in terms of peak ground i acceleration, a C A=FE'ASSE (5-2) where:
d = Median ground acceleration capacity
, i ASSE
= Peak ground acceleration of the safe shutdown earthouake I
4 In most instances, the SSE was used as the reference aarthquake; however, the OBE was used as a reference for those cases where the OBE acceptance l
criteria governed the equipment design,- -
! The logarithmic standard deviation, S, for cach of these factors is obtained using the logarithmic standard deviations for each of the above f actors and based upon the lognormal model (Appendix A).
8 E"(8hC+SER + 8ED + 85R l
n.
5-2 l
l l
F where sEC' 8FR' OED, and 6SR are the logarithmic standard deviations of the equipment capacity, equipment response, earthquake duration, and structural response, respectively. The logarithmic standard deviations are further divided into random variability, BR '
and uncertainty, SU , as described in Chapter 3.
5.1.1.1 Equipment Capacity Factor The equipment capacity factor is defined as the failure threshold divided by seismic design level. For the purposes of this study, the ultimate failure threshold is the acceleration level at which the component ceases to perform its intended function. This failure threshold could consist of a breaker tripping on a motor control center, excessive deflection of the control rod guide tubes or a support failure of the steam generator. Where several failure modes pertaining to the same component are found to have roughly the same capacity level, all significant failure modes are analyzed and reported.
The f actor of safety for the equipment seismic capacity consists
- of two parts:
- 1. The strength factor, Fs, based on the components static strength and
- 2. The ductility factor, F ,yrelated to the equipment's inelastic energy absorption capability.
FEC = F3Fu (5-4)
~
The logarithmic standard deviation on the capacity can be derived by taking the SRSS of the logarithmic standard deviations on the strength factor and the ductility factor. The randomness and the uncertainty por-tion of the variability can each be derived individually from Equation 5-5, by substituting the random or the uncertainty 83 for the strength f actor and the ductility f actor (i.e., SS for 83 and B y for B y, etc.).
S EC*(85+8)2 9 (5-5) 5-3
5.1.1.1.1 Strength Factor - The strength factor, FS , is derived from the equation:
P P C N pS, P b b P
(5-6)
T N O
b b where CP is the median limit state load or stress, NP IS O' "O**I operating load or stress, PT is the total normal plus seismic load or stress and PD is the code design allowable load or stress.
Alternatively, this equation can be written:
PC-PN F
3= p (5-7)
- SSE where P SSE is the seismic load or stress corresponding to the safe shutdown earthquake. The normal and the seismic loads (PN and PSSE) i are typically derived from the seismic qualification reports and the other information sources described in Section 5.1.2. The calculation of l
the collapse load, PC , is a function of the failure mode for the specific equipment item. Equipment failures can be classified into three
! categories: ,
. 1. Elastic functional failures
- 2. Brittle failures
- 3. Ductile Failures.
t 4
5-4
Elastic functional failures involve the loss of intended function while the component is stressed below its yield point. Examples of this type of failure include:
- 1. Elastic buckling in tank walls and component supports.
- 2. Chatter and trip in electrical compon.ents.
- 3. Excessive blade deflection in fans.
- 4. Shaft seizure in pumps.
The limit state load for this type of a f ailure is defined as the median load or stress level where functional failure occurs.
i Brittle failures are defined in this study as those f ailure modes which have little or no system inelastic energy absorption capability. Examples of brittle type failures include: ,
- 1. Anchor bolt failures.
- 2. Component support weld failures.
- 3. Shear pin failures.
Each of these failure modes have the ability to absorb some inelastic energy on the component level, but the plastic zone is very localized and the system ductility for an anchor bolt or a support weld is very small.
Thus, the collapse load for a brittle failure mode is defined as the median ultimate strength of the material. For example, consider a trans-l fomer structure whose anchor bolts have been determined to be the criti-cal failure mode. Under seismic loading, the massive transformer will typically be stressed well below its yield level while the bolts are being stressed well above the bolt yield level. The amount of system inelastic energy absorption provided by the bolts' plasticity is negligible when compared to the seismically induced kinetic energy of the transmission structure, and thus, these bolts will fail in a brittle mode once the ultimate bolt strength is reached.
f 5-5
Ductile failures coincide much more closely with the structure failures which were described in Chapter 4. Ductile failure modes are those in which the structural system can absorb a significant amount of energy through inelastic deformation. Examples of ductile failure modes include:
- 1. Pressure boundary failure of piping
. 2. Structural failure of cable trays
- 3. Structural failure of ducting
- 4. Polar crane failure.
The collapse load for ductile failure modes consists of the median yield strength of the material for tensile type loading conditions. For bending type failure modes, the yield point is defined as the limit load or stress to develop a plastic hinge. The ductility factor will then quantify the inherent safety factor above the yield strength to the failure threshold.
Each variable within Equation 5-6 has an associated lognormal probability distribution to express its combined randomness and uncer-tainty. To find the overall variance on the strength factor, a technique comonly referred to as the "Second Moment Method" was utilized. The
,i mean and variance of a function comprised of lognormally distributed l
variables can be derived utilizing the moments (i.e., the mean and variances) of the logarithms of the distribution of each variable (Reference 27). The resulting equation for the logarithmic standard
- deviation on the strength factor is given below:
"2 "2r l
C ..,2
.s .
4 . (PT-PN (PC-P) N
~
(5-8)
(PC-P)T P
N 2 Y2 (PT-P) N .(PC-PN .
5-6 l
l
where:
SC
= Logarithmic standard deviation on the collapse load (stress).
= Logarithmic standard deviation on the total load (stress).
8N
= Logarithmic standard deviation on the normal
- load (stress).
5.1.1.1.2 Ductility Factor - The inelastic energy absorption capability of a piece of equipment is quantified by the ductility factor. Brittle f ailure modes and functional failure modes typically have a ductility f actor of 1.0, while ductile type f ailure modes have ductility factors which are a function of a deamplification factor (Reference 7). At moderate damping levels, the ductility factor (F,) for equipment that responds in the amplified acceleration region of the design spectrum is approximately:
F,=c(2u-1)2 (5-9)
The symbol u represents the ductility ratio and c is a random variable with a median value equal to 1.0 and a logarithmic standard deviation ranging from about 0.02 to 0.10 (depending on the ductility), which represents the uncertainty in the use of Equation 5-9. For rigid equipment, Reference 7 recomends that the ductilty f actor be represented by:
0.13 '
F, = c.9 (5-10)
. Again, c is a variable of median equal to 1.0 with a logarithmic standard deviation ranging from about 0.02 to 0.10, as a function of the ductility.
I In both Equations 5-9 and 5-10, the logarithmic standard deviation on e increases with the ductility, u. Equation 5-9 applies to I equipment with a fundamental frequency in the arhplified acceleration 5-7
range (about 2 to 8 Hz) and Equation 5-10 applies to equipment with a fundamental frequency greater than about 30 Hz. Between about 8 and 30 Hz the ductility factor and the logarithmic standard deviation of are interpolated.
The ductility ratio, p, itself is based upon the recommendations given in Reference 6. This reference gives a range of ductility values to be used for design. The upper end of this range is considered to be a median value while the lower end of this range is considered to be about a -2 logarithmic standard deviation value. Engineering judgment was utilized to match the applicable category from Reference 6 to a particular f ailure mode for the equipment component.
5.1.1.2 Equipment Response Factor The response factors are an estimate of the conservatism or unconservatism that may have existed in the computation of seismic response during the design process. In this section, individual response f actors are described for both plant specific and generic equipment.
These factors differ according to the seismic qualification procedure which was used in the equipment design.
! There are three types of seismic qualifications which were perfomed for Seabrook plant equipment:
- 1. Dynamic Analysis l ;
- 2. Static Analysis
- 3. Testing.
For equipment qualified by dynamic analysis, the important variables that affect the computed response and its dispersion are:
'l
- 1. Qualification Method (FgM)
- 2. Spectral Shape (FS$)
5-8
3.
Modelino)(effectsmodeshapeandfrequency results (FM )
- 4. Damping (FD )
- 5. Combination of Modal Responses (for response spectrum method) (FMC)
- 6. Combination of Earthquake Components (FECC)
. For equipment qualified by static analysis, two subdivisions must be considered. For rigid equipment, spectral shape, combination of modal responses, damping,'and for the most part, modeling errors are eliminated. If the equipment is flexible and was designed via the static coefficient method, the dynamic characteristic variables and their i variability must be considered. This involves estimating the range of frequency of the equipment and introduces a much larger uncertainty in quantifying the response factor.
1, Where testing is conducted for seismic qualification, the response factor must take into account:
- 1. Qualifiction Method (FQM)
- 2. Spectral Shape (Fss)
- 3. Boundary Conditions in the Test vs Installation (FBC)
- 4. Damping (FD )
- 5. Spectral Test Method (sine beat, sine sweep, compexwaveform,etc.)(FSTM)
- 6. Multi-directional Effects (FMDE).
. The overall equipment response factor is the product of each of these variables. The overall variabilities (uncertainty and randomness) are calculated by taking the SRSS of the individual logarithmic standard deviations for each of the variables. A brief description of each of the variables used to develop the equipment response factor is provided below. A more detailed discussion is contained within Reference (28). ,
i 5-9 l
4 5.1.1.2.1qu,alificationMethodFactor-Thequalificationmethodfactor is a measure of the conservatism /unconservatism involved in the seismic qualification method used to seismically qualify the component.
Analytical qualifications can be separated into static analysis and dynamic analysis techniques. The inherent safety factor in using these 4
qualification techniques is discussed below, while the variability on this factor is accounted for within the damping, modeling and mode combination factors (i.e., 8 =8 =0.0).
1
- 5.1.1.2.1.1 Static Analysis - The static coefficient method is intended to be a conservative upper bound method by which simple components may be qualified. Typically, the peak spectral acceleration is multiplied by a coefficient and this product is multiplied by the. weight of the component to determine an equivalent static load to be applied at the subsystem center of gravity. If the component is comprised of more than one lumped mass, the same procedure may be applied at each lumped mass point in the static model or may be applied as a uniformly distributed load on the static model. If the component..is rigid (i.e., its fundamental frequency
- is above the frequency where the response spectrum returns to the zero period acceleration), the degree of conservatism in the response level
. used for design is the ratio of the specified static coefficient divided
- by the zero period acceleration of the floor level where the equipment is
' mounted. If the equipment is flexible and responds predominantly in one i
mode, the degree of conservatism is the ratio of the static coefficient to the spectral acceleration at the equipment fundamental frequency.
5.1.1.2.1.2 Dynamic Analysis - Response spectrum, mode superposition
- time-history and direct integration time-history dynamic analysis methods may be applied in subsystem response analyses. If response for a single l
degree-of-freedom model with best estimate material properties and damping l ,
are computed by the response spectrum method, the mode superposition ,
time-history method or the direct integration time-history method, we would expect to obtain equal median centered results assuming that the l
! response spectrum and time-history inputs are compatible.
5-10 f .
i
The response spectrum method was extensively used for dynamic analysis of components and systems within the Seabrook plant. If the applicable Seabrook floor response spectra were utilized in the design analysis, the qualification method factor, Fy , is equal to unity and the variability is zero. If a conservative generic spectra was used to seismically qualify a component, Fy is the ratio of the spectral acceleration from the generic spectra divided by the spectral acceleration from the Seabrook site specific spectra evaluated at the components' fundamental frequency.
5.1.1.2.1.3 Testina - In vibration testing, the test response spectrum
> generally envelopes the required response spectrim. by approximately ten
- percent or more depending on the frequency range. If the test response spectra are available within the test report, the overtest safety factor will be accounted for in the strength factor, the qualification method 4
factor (Fy) and variability (By) will be unity and zero, respec-i tively. If the component fragility is bf.ing based on testing where the
, test response spectra are not av'ailable, Fp and 8p account for the j overtest safety f actor and variability on a generic case-by-case basis.
t
! 5.1.1.2.2 Equipment Spectral Shape Factor - The Seabrook floor response l spectra (Ref erence 29) were computed by means of a time-history (T/H) seismic analysis. The overall dynamic response of each of the critical buildings was modeled by lumping the mass of the structure and rigidly attached components generally at each of the floor levels. Two artificial T/H's (each with a peak acceleration of 0.25g) were used in the T/H analyses. One T/H was applied horizontally in each of the N-S and E-W directions while the second T/H was applied in the vertical direction. The two time-histories were developed to envelope, as closely as possible, the ground response spectra for the Seabrook site. The conservatism /unconservatism involved in developing the floor response spectra from the ground response spectra is quantified with the equipment spectral shape factor. The conservatism /unconservatism involved in using the specified Seabrook design response spectra in lieu of a median Safe Shutdown Earthquake spectra is quantified in development of the spectral shape f actor associated with the structural response f actor.
5-11
~
f n -- .-
l The response spectrum method is often referred to as being conservative, however, the conservatism compared to a time-history analysis is primarily due to the method of developing the spectrum.
Spectra used for design purposes are generally smoothed and the peaks are widened such that the resulting design spectrum is conservative. In addition, conservatism is generally introduced in the development of the artificial time-history. The combined effect of the two conservatisms make up the equipment spectral shape factor.
5.1.1.2.2.1 Peak B,roadening and Smoothing - The effect of smoothing and peak broadening varies with structure, elevation, frequency and damping.
Comparisons between broadened and unbroadened floor spectra for various
' elevations within the Containment Building, Primary Auxiliary Building, Fuel Building and Control / Diesel Generator Butiding art contained within Section 3.7 of the Seabrook FSAR (Reference 1). Figure 5-1 shows a typical example of these spectra comparisons.
Factors of conservatism due to peak broadening and smoothing were generated for these four buildings based on these FSAR spectra. For I any particular frequency, this safety factor can be computed from Equation 5-11.
, 5,(broadenedandsmoothed)
. F 33
. (5-11) 5, (unbroadened and unsmoothed)
I where:
FSSg = Spectral shape factor due to peak broadening and smoothing Sa = Spectral acceleration value ,
5-12
Due to the inexact nature of predicting fundamental frequencies, these safety factors were computed for frequency bands of roughly 5 Hz incre-ments. The median ratio within these 5 Hz bands was used as a factor, F
SS and a minus two logarithmic standard deviations were estimated toEx,istbetweenthemedianandtheminimumratioswithina5Hzband.
Therefore,
= 1/
SS I(median)\ (5-12)
- 835 1 2sn(FS$g(minimum))
where:
6 55
= Logarithmic standard deviation on the peak broadening and smoothing factor l
Since the variability in 833 is due to the shift in the frequency, it is considered to be all uncertainty.
For equipment located in buildings other than the four which were analyzed, average median spectral shape factors from the four butidings were used and the variability 835 was computed by taking a f actor of unity as a -28 lower bound. Table 5-1 shows the peak broadening and smoothing factors for each of these buildings together with the loga-f rithmic standard deviation due to uncertainty. These values are appli-
! cable to equipment on any floor level except the basemat. Equipment on
- the basemat were qualified to the ground spectra, and have a factor of
{
1.0 and a variability of 0.0 since floor spectra were not generated at
. this elevation.
5.1.1.2.2.2 Artificial Time-History Generation - Studies have been con.
ducted which show that conservatism is involved in the current practice of generating floor spectra in structures using artificial time-histories.
l These artificial time histories result in response spectra that conserva-tively envelope the applicable ground spectra. For instance, Reference l 30 indicates that the average industry-generated artificial time-history tends to introduce about 10 percent conservatism except at high frequen-l 5-13
cies for which the conservatism is about 20 p'ercent at 33 Hz. Our own experience is compatible with these numbers. Since Stabrook structures have fundamental frequencies in the amplified region of the spectra, an artificial time-history f actor of 1.1 is applied. This relatively low f actor of conservatism is a reflection of the substantial industry effort to reduce this arbitrary source of conservatism.
- It has also been observed that two different artificial time-histories, both of which result in response spectra which adequately envelope the Regulatory Guide 1.60 response spectra, can lead to floor spectra which may differ by a factor of 2 or more (for instance, see
. Reference 31). Use of the artificial time-histories method for testing and analysis can result in a small arbitrary amount of conservatism on the average and considerable dispersion in the resultant response.
Reference 31 reports that a coefficient of variation of 0.2 is reasonable based on a comparison study of 44 synthetic time-histories. This vari-ability on the artificial time-history generation factor is classified as being all randomness since it r'epresents the variety of possible earth-quakes.
The overall spectral shape factor was generated by taking the product of the peak broadening and smoothing factor times the artificial time-history f actor. Table 5-1 shows the spectral shape factors and their variabilities for each of the seismically critical buildings as a function of the frequency interval. Note that a 5 to 20 Hz interval has been included to reflect the most possible frequency ranges for complex mechanical systems such as piping and cable trays.
5.1.1.2.3 Modeling Factor - In any dynamic analysis there is uncertainty l l' in resonse due to assunptions made in modeling the structure, modeling boundary conditions and representing material behavior. Modeling of complex systems is usually conducted using nominal dimensions, weights, and material properties and is done in such a manner that further refinement of mesh size in a finite element representation will not significantly alter the calculated response. Representation of boundary 5-14
conditions in a model may have a significant influence on the response.
The misrepresentation of boundary conditions in the dynamic model by assuming greater or lesser stiffness or treating nonlinear gap effects linearly cannot be quantified generically and each model must be treated specifically to determine a response factor for modeling. Assuming that the analyst does his best job of modeling, modeling accuracy could be considered to be median centered (i.e., Fg=1.0)withthevariability in each of the modeling parameters amounting to variability in calculated mode shapes and frequencies. The error in calculation of mode shapes and frequencies then has an effect on the computed response.
For complex state-of-the-art dynamic analysis, the coefficient
+
of variation on response (approxinate logarithmic standard deviation) is about 0.20. For simple single-frequency systems the coefficient of variation is about 0.10, and for systems of medium complexity the coefficient of variation is about 0.15. These variabilities are considered to be all uncertainty and are based on past experience and engineering judgment.
5.1.1.2.4 Damping Factor _ - The basis for the damping factor has been addressed in Section 3.4.2 of this report. Tables 3-1 and 3-2 show the
. danping values used for the SSE design analysis of Seabrook equipment.
Median damping values and their variabilities are a function of the material, construction details, size and stress level. Reference (28) suggests that median damping for equipment at the $$E level is about five percent. Thus, for single-degree-of-freedom systems the damping factor for Seabrook equipment is: .
s S (qual)
(5-13)
FD" 5 (median) 4 5-15
where:
S(qual) a
= Spectral acceleration using the qualification design analysis damping and evaluated at the equipment fundamental frequency Sa(median) = Spectral acceleration using the expected median damping and evaluated at the equipment fundamental frequency.
for multi-degree-of-freedom systems Equation 5-13 can be altered to
- reflect the sismation of the spectral accelerations at each of the frequencies multiplied by their associated mass participation f actors.
There is variability in damping and associated response that must be considered. It is indicated within Reference 28 that for a median damping value of 5 percent, the minus one logarithmic standard deviation value is about 3.5 percent. The variability in damping results in a logarithmic standard deviation in response equal to:
[I * '
h 8 " I" (5-14)
D 5 U
N*s=5.0%)
where Sa is the 5 percent damped spectral acceleration and Sa
, $g c = 3.5%
is the 3.5 percent damped spectral acceleration taken at the equipment
- fundamental frequency using the applicable floor response spectra. The resulting logarithmic standard deviation on the damping response f actor,
. from Equation 5-14 above, is considered to be all uncertainty. An additional randomness variability estimated at approximately 20 percent of the uncertainty variability reflects the earthquake time-histories' effect on the median damping value.
1 l
t 5-16
l 2
k' 5.1.1.2.5 Mode Combination Factor - The modal combination technique utilized within the Seabrook seismic design analysis was described in
]! general in Section 3.4.4 of this report. A square-root-of-the-sum-of-the-squares ($R$$) methodology was used for all Seabrook equipment. This
! SRSS method (allowing for absolute sum for closely spaced modes) is in l accordance with current Regulatory Guide 1.g2 (Reference 12) recomended practice and is considered median centered.
1 The response factor for combination of modes is then considered to be 1.0. The variability associated with mode combination depends upon r
the complexity of the model. For multi-degree-of-freedom systems.
- Reference (28) recomends that the coefficient of variation due to mode
! combination is approximately 0.15. For single-degree-of-freedom flexible l systems, the coefficient of variation due to mode combination is estimated I ,
within Reference 28 to be approximately 0.10. For a single-degree of.
j freedom rigid system, the C0V is by definition zero. The variability due to mode combination is considered to be all random due to the random
{
j phasing of modes. .
5.1.1.2.6 Earthquake Component Combination Factor - The Seabrook plant l
) - design analyses earthquake components were required to be combined by the
! SR$$ of the vertical and the two horizontal components. This approach ,
requires that the effects of two horizontal directional responses be combined with the vertical response, but does not require the maximum response in each direction occur at the same instant as the maximum response in the other two directions. Reference 10 recomends that the median response can be represented by combining the worst case horizontal l
response with 40 percent of the orthogonal horizontal response and 40 i percent of the vertical response. Comparing this suggestion to the Seabrook design criterion results in a response factor for combination of l
! earthquake components. The magnitude of the factor depends, however, on the orientation and response characteristics of the component under
- consideration.
4
)
j 5-17 i
A generic study was conducted to develop earthquake component combination response factors and their variabilities for comon two- and three-dimensional equipment idealizations. The amount of conservatism /
unconservatism and the associated variability on this factor are a function of the following:
- 1. The number and direction of earthquake components which affect the failure mode under consideration (e.g., piping failures are influenced by all three directional responses, but a particular relay can fail due to a particular horizontal
. seismic excitation while remaining unaffected by) the vertical and the other horizontal directions
- 2. The amount of coupling that exists between directional response (i.e., does an x direction excitation cause a response in the y and z directions)
- 3. The attachment configuration for anchor bolt failures (rectangularanchorboltpatternswill behave differently than circular patterns when combiningdirect1.onalresponses)
- 4. The end item of response which was combined by SRSS(i.e.,differentsafetyfactorsexistfor the desiga analysis where the directional loads were combined by SRSS and then aplied to the model to find the resulting stress, than for the design analysis where stresses were calculated for each of the three earthquake excitations separately and then combined by SRSS).
8 Table 5-2 contains the earthquake component combination response f actors for those cases which were applicable to Seabrook equipment. The
. variability involved in the phasing of the three earthquake directional components was considered to be all random, while the variability due to the degree of coupling involved between directions was considered to be all uncertainty. ,
5-18
5.1.1.2.7 Boundary Conditions Factor (Testing) - The boundary conditions utilized in equipment seismic testing can be a significant source of variability that depends almost solely upon the diligence of the test laboratory and the qualification review organization. In general, a component that is bolted to the floor in a nuclear power plant and which is similarly bolted to a shake table for qualification testing, will experience little variability in response factor due to boundary conditions. Carelessness on the part of the various organizations involved in design, fabrication, testing and installation can result in a significant variability. For instance, the lack of a specified bolt
. torque at the mounting interface can result in a difference between the testing and installation condition which could have a pronounced impact on the response factor.
The variability of the subsystem response due to test boundary conditions would come primarily from mode shape and frequency shift. The variability of mode shape and frequency and resulting response due to boundary conditions varies considerably for different generic types of equipment. For a large majority of tests conducted by reputable testing laboratories, the boundary condition factor is 1.0. Engineering judgment must be utilized in calculating boundary condition factors for those cases where the component to test table attachment mechanism is not representative of the actual in-plane condition. The variability is all uncertainty and can be calculated based on spectral accelerations obtained from estimating a 90 percent confidence interval on the equipment frequency.
5.1.1.2.8 Spectral Test Method . Synthesized time-histories are currently developed directly from the Required Response Spectrum at most testing laboratories. A much better approach, as recomended in Reference 32, is to synthesize a time-history that corresponds to a power spectral density which closely envelopes the RR$ rather than make the direct step from the RR$ to the synthesized time-history. This approach tends to smooth out the input time-history, resulting in less chance for an equipment i
l 5-19
mode to coincide with a significant peak or valley. Reference 28 recomends a spectral test method factor of unity and a total variability of 0.11. This variability is entirely uncertainty since the use of better equipment and techniques could eliminate most of the uncertainty.
5.1.1.?.9 Multi-Directional Effects - The multi-directional effects factor is a measure of the conservative /unconservatism and corresponding variability involved in testing the three different earthquake directional components. Seabrook equipment fragilities were developed from plant specific and generic test data and are based on two types of testing:
- biaxial and uniaxial. Biaxial qualification tests are conducted by exciting the equipment in one horizontal direction at a time along with the vertical direction, using randomly phased input time-histories.
Uniaxial qualification tests, on the other hand, are conducted in each of the three directions independently. Biaxial testing was required for all plant specific equipment qualified for the Seabrook plant. The shock tests conducted during the SAFEGUARD program were, in many cases, single axis tests with complex waveforms consisting of superimposed sine beats.
Some blaxial testing data were included when deriving the generic SAFE-GUARD fragilities, but were scaled to an equivalent uniaxial input.
. Thus, multi-directional effect factors were developed for biaxial testing (usedforfragtlitiesdevelopedforplantspecificSeabrooktesting)and uniaxial testing (used for fragilities based on generic SAFEGUARDS test data).
5.1.1.2.9.1 Biaxial Testino - There is a slight unconservatism involved j in biaxial testing in that the actual input during a seismic event is
. three-dimensional. This unconservatism along with its associated variability is a function of both the phasing and the coupling between earthquake directional corponents. Assuming that the median acceleration vector can be defined as recommended in Reference 10 as 100 percent of the acceleration in one direction plus 40 percent of the a:celeration in the other two orthogonal directions, the degree of unconservatism associated with biaxial testing can be defined as the median response i
I 5-20 l
1
l 1
vector for biaxial testing divided by the median three-axis response.
The resulting response factor based on both phasing and coupling is calculated to be 0.86. The variability due to phasing is a function of the earthquake, and thus, is all random. The phasing variability is identical to that which has been calculated for the general three-2 dimensional condition (case No. 1 of Table 5-2) for the earthquake
- component combination factor and is equal to 0.12. The variability due
' - to coupling is very small since median coupling in testing axists by definition for the two input directional components. Using the uncoupled case and the 100 percent coupling case as 13e extremes on coupling, an uncertainty logarithmic standard deviation of 0.01 is calculated.
i 1
- The multi-directional effects factor and its associated S's for random vibration biaxial testing is:
i IMDE = 0.86 ENDER
= 0.12 e = 0.01 MDEU 5.1.1.2.9.2 Uniaxial Testing - A uniaxial test is, in general, unconserva-tive in that coupling and phasing between the three-directional earthquake
- components is not accounted for. Again, assuming the median acceleration vector can be defined as recomended in Reference 10 as 100 percent of the acceleration in one direction plus 40 percent of the acceleration in the other two orthogonal directions, the degree of unconservatism associated I
with uniaxial testing can be defined as the median response vector for
- untaxial testing divided by the median three-axis response. The resulting response factor based on both phasing and coupling is calculated to be ,
0.735. The phasing variability is random and is identical to that for the blaxial case, i.e., 0.12. The uncertainty variability due to coupling, based on the uncoupled case and the 100 percent coupling case 4
being 138 extremes, is calculated to be 0.07.
l I
a 5-21
( <
l l
Thus, the multi-directional effects factor and its associated s's for untaxial testing is:
IMDE = 0.735 l 8MDEg
= 0.12 8 = 0.07 MDEg 5.1.1.3 Structural Resoonse Factors Structural response factors as they relate to structural capacity
- for the safety-related structures within Seabrook are derived in Chapter
- 4. These structural response factors were computed assuming the structure was at or above its yield point. For equipment whose seismic capacity level has been reached while the structure is still within the elastic range, these structural response factors are optimistic. Reference 10 recomends seven percent damping for reinforced concrete at the yield condition and five percent damping for reinforced concrete at the one-half yield condition. Thus, a second set of structural response factors which reflect the one-half yield condition of the structure (using five percent damping) were calculated. The variables pertinent to the struc-tural response analyses used to generate floor spectra for equipment design are the only variables of interest relative to equipment fragility.
Time-history analyses, using the same structural models used to conduct structural response analyses for structural design, were used to generate floor spectra., The applicable variables from those analyses are:
- 1. Spectral Shape 4
- 2. Damping
- 3. Modeling I 4. Soil-Structure Interaction.
ll The explanation of these variable effects are contained in Chapter 4 and will not be repeated here. Note, the combination of earthquake components is not included in structural response since that
- 5-22
l l
variable is addressed for specific equipment orientation in the treatment !
of equipment response. The resultant structural response factors derived from each of the above variables that pertain to equipment fragilities are included in Table 5-3.
Equipment capacities were compared to capacities of the struc-tures that housed the equipment in order to determine the appropriate damping f actor of safety to use in developing the structural response factor. The approximate yield level for each of the buildings was
. estimated by taking the ground acceleration capacity for the lowest structural f ailure mode (see Chapter 4) and dividing it by the inelastic energy absorption factor. For those cases where tne equipment was determined to f ail at a level where the structure remained elastic, the spectral shape f actor appropriate for five percent damping was substituted for the seven percent damped spectral shape f actor.
It should be noted that equipment that is located on the base mat have had their structural response factors calculated on a case by case basis. Since the ground spectrum was utilized in their qualifica-tion, the spectral shape f actor reflects the conservatism /unconservatism involved in using the ground spectra as opposed to the site specific median spectra. In this instance, only the srectral shape and soil structure interaction factors apply. The structure modeling factor and the damping f actor do not apply since the equipment response is independent of the building response.
5.1.1.4 Earthquake Duration Factor The earthquake duration factor has previously been addressed in Section 3.3 and 4.1.3. The basis for this factor is the increase in capacity due to the median expected earthquake being less severe than the earthquake on which the equipment ductility factors were based. Equipment ductility f actors were taken from the recomendations in Reference 6 and -
are based on three to five cycles of strong seismic motion. The median -
expected earthquake for the Seabrook site has a duration of seven to nine seconds and one to three strong motion cycles. The seismic capacity of I 5-23 i e
, _ , - . _ _ _ , _ _ _ . . . _ - . . . , , , . . . . _ - - .__.-__.-__-7___,._., -- w,_ -._.,.3 _., ,, _ -,_-________.y _
, _ _ - ~ - _ ~ - _ , ,
4 j equipment that fails in a structural mode is affected by the duration and
} resulting number of strong motion cycles since the expected available ductility of the controlling structural elements increases as the number 3
of cycles is reduced.
As derived in Section 4.1.3, the duration factor and its
! Iogarithmic standard deviations for equipment that fails in a ductile manner is:
! En = 1.40 i
8 = 0.12 l* 0g s
g . 0.08 For equipment that fails in a brittle manner or for functional I failure modes within the elastic range, the duration factor and its l variability are unity and zero, respectively. ,
5.1.2 Information Sources -
lj Several sources of information are utilized in a PRA from which 1~ to develop plant specific and generic fragilities for equipment. These sources include:
l 1. Seismic Qualification Design Reports 4
6 2. Seismic Qualification Test Reports
! 3. FinalSafetyAnalysisReport(FSAR) l 4. SeismicQualificationReviewTeam($QRT) +
Submittals
- 5. Seismic Qualification Report Sumaries f
- 6. Vendor drawings from which new analyses are
- conducted l
l 7. Past Earthquake Emperience 8.- Specifications for the Seismic Design of Equipment i
5 24 l
4 i
The first 6 of these information sources are termed " plant specific" since they pertain to specific equipment within the Seabrook plant. The remaining information sources are termed " generic" since they constitute data generated for similar types of equipment or are definitions of design requirements, in lieu of actual design results.
Plant specific sources are preferred since they have been generated for the specific items in question and their uncertainty level is reduced from that of the generic sources.
5.1.2.1 Seismic Qualification Analysis Reports
, The majority of the fragility levels for critical Seabrook equipment were developed from the review of seismic qualification analysis reports. Westinghouse provided qualification mport sumaries for most of the equipment items which they had supplied to Seabrook, and fragility levels were calculated based on these summaries. In some caseh, the Westinghouse supplied data was based on a generic analysis where generic spectra, which enveloped the response spectra for several plant sites, had been used for the loading. In these cases, the stresses have been scaled down to reflect the response for the Seabrook site, and thus, these cases essentially constitute a plant specific analysis.
5.1.2.2 Seismic Qualification Test Reports Several examples of test reports for equipment qualified by testing were reviewed. Qualification test reports, by themselves, cannot be utilized to directly develop full fragility relationships unless the equipment has been tested to increased vibration levels up to failure.
Fragility at levels greater than design criteria can, however, be inferred from test data and if the test levels were sufficiently high, the inferred fragility levels may be such that there is high confidence failure will not occur below the maximum vibration level postulated for the site.
5.1.2.3 Final Safety Analysis Report & SQRT Sumaries The Seabrook FSAR contained very few SSE stress sumaries on critical equipment; thus, was not significantly utilized within the 5 25
seismic portion of the PRA. Seabrook SQRT summaries were used to a limited degree to demonstrate that selected items of equiment had median ground acceleration capacities exceeding 2.0g.
5.1.2.4 Vendor Drawinesor Desien Reports from Which New Analyses are '
Conducted Occasionally, it is fruitful to conduct new analyses for critical components in order to remove hidden conservatisms inherent in the
- original design analysis and to reduce the uncertainty in the fragility derivation. New analyses were conducted for the diesel fuel oil day tank i and the refueling water storage tank.
5.1.2.5 Past Earthquake Experience Past earthquake experience is valuable for establishing fragilities for equipment which have historically been vulnerable. Most equipment survives without any apparent damage and the historic experience must be treated the same as a qualification test. Earthquake experience was used to estimate fragility levels for non vital AC power.
i 5.1.2.6 Specification for the Desion of Equipment I Specifications for seismic qualification of Seabrook equipment were provided by United Engineers and Constructors, Inc. In cases where l plant specific qualification reports were not reviewed, knowledge of the l l vendor requirements plus generic fragility and qualification test data l were combined to develop fragility descriptions.
L i
I i
5 26 i
r
l
! i 5.1.3 Eauipment Catenories ,
Depending upon the uniqueness of the equipment, the failure mode.
inelastic energy absorption capability and the dynamic characteristics of the equipment, a plant specific or a generic derivation of the fragility j description may be appropriate. The factors of safety relative to the i Safe Shutdown Earthquake are widely variable. In general, fienible I j equipment such as piping, which possesses the ability to undergo large ,
inelastic deformation, will have a factor of safety against f ailure of j j . many times the Safe Shutdown Earthquake even if stressed to the maximum
) code allowable stress. Such equipment is a prime candidate for generic i j* derivation of fragility descriptions. The increased uncertainty inherent !
j in a generic derivation does not have much influence on the outcome of l l the seismic risk analysis if large safety facto:s can be demonstrated. ;
I On the other hand, if rigid equipment with relatively brittle failure !
modes are Stressed to code allowable for the Safe Shutdown Earthquake, l l j the factor of safety against failure my be considerably smaller and a ;
l generic treatment may result in unsatisf actory risk predictions. ;
Fortunately, most rigid equipment stress response is very much less than ;
- the allowable and large safety factors are present. l I ,
The extensive number of safety related equipment items within a [
ll nuclear power plant, coupled with practical limits on resources available with which to perform a PRA analysis, necessitates limiting the amount of j l i detailed analysis devoted to any single piece of equipment. Importance f l gained as to relative component fragility levels must be utilized in j deciding which components to treat generically and which components to l treat specifically. Those components which have been shown to possess a .
high degree of resistance to seismic loading can be treated in more of a !
generic f ashion. The higher degree of uncertainty d ich results from the ;
j use of generic analysis can be accepted due to the inherently high }
l fragility level for components of this type. Likewise, cogonents which !
can be shown through esperience to possess low fragility levels should j have a more thorough analysis undertaken to provide as much accuracy as !
l possible. Note that of ten af ter a preliminary risk analysis has been ,
l l
I l 5 27 :
i i 1 :
I 1 0
conducted, components which contribute heavily to the overall risk of the plant are reevaluated in more detail to reduce sources of uncertainty that may have arisen from an approximate or generic derivation of fragility. Table 5-4 contains a listing of the relative fragility level for general equipment categorles based on the data provided in Reference
- 33. The information provided within Table 5 4 is utilized within the PRA study to discern which equipment can be treated generically and which equipment should be treated on a plant specific basis.
. 5.2 EQUlPMENT FRAGILITY EXAMPLES
- Because of the amount of equipment to be included within the risk model, it is impractical to describe the specific fragility deriva.
tion for each piece of equipment. This section contains selected exangles of fraglitty derivations which are judged to be representative of the different types of analyses which had to be undertaken for Seabrook equipment. Tne equipment fragility derivation categories applicable to the Seabrook PRA are:
- 1. Equipment whose fraglitty descriptions are based on plant specific design reports.
- 2. Equipment whose frag 111ty desciiptions are based on qualification test reports
- 3. Equipment whose fragility descriptions are based on knowledge of the design specifications and the f actors of safety inherent in the governing codes and standards
- 4. Equipment whose fragility descriptions are based on new analysis a 5. Equipment whose fragility descriptions are based on englineering judgment and past earthquake exper<ence(nonseismicallyqualifiedcomponents).
An example of Seabrook equipment whose fragility derivation stems from each of the above categories is included in this section.
5 28
4 5.2.1 Example of a Plant Specific Design Report Fragility Derivation The majority c,f the fragility derivations for the Seabrook'PftA ,
fall within the plant specific design mport category. United Engineers w and Constructors, Inc. (UE&C) supplied the seismic qualification reports s for all of the critical equipment for which they were responsible in the 5eabrook plant development. These UE&C supplied reports were used to derive fragilities for components except for those cases where a generic ,
fragility development was deemed appropriate (empected high capacity) or
- where the qualification report was not descriptive enough to accurately derive a capacity.
I The example chosen for the plant specific design report i fragility derivation is the spray additive tank. The spray additive tank has a 6.67 foot diameter, is 44.25 feet high and holds 10,700 gallons.
3 It is located at Elevation 25' within the Tank Farm. The tank is i
constructed of SA 240 Type 304 stainless steel and was designed to ASME Code Section !!! Subsection ND Class 3 criteria. The qualification analysis (Reference 34) was performed by the Pittsburg, Des Moines Stee)
- Company using a response spectrum approach.
1 l
A review of Reference 34 reveals the critical areas during
' seismic loading are
! 1. Buckling of the tank walls
- 2. Anchor bolts ,
- 3. Anchor Bracket.
l Strefigth factors were calculated for each of these failure modes and the resulting governing case was the anchor bolts. The other two fail'are modes had strength factors significantly above that of the ancnar bolts; thus, only the anchor bolts were utilized in the fragility derivation.
i Table 5 5 contains all of the factors which make up the final fragility values for the spray additive tanks. Each of these f actors is discussed .
t briefly. . .
i 6
5 29 S
5.2.1.1 Spray Additive Tank Capacity Factor _
The anchor bolt pattern consists of eight, two-inch diameter SA193 B7 bolts on a 7'-25 /8" diameter bolt circle. Reference 34 states the favited condition maximum load to be a tensile load of 214.9 kips and a shear load of 18.53 kips. The state of stress in the most highly stressed bolt can be calculated as:
o= = 85.96 ksi (tensilestress) 2.50 in
- 18. 3 s t= = 8.65 ksi (shearstress) 3.142 in These stresses are essentially all due to the safe shutdown earthquake 4
. (SSE) since the normal stress in the anchor bolts is negligible.
~ Anchor bolt failures are considered to be a brittle type of f ailure mode. The system inelastic energy absorption capability for anchor bolts is negligible in comparison to the amount of kinetic energy which is input into the system by the earthquake. There exists only a small region in the bolt which goes inelastic and when the median ultimate strength of the bolt is reached, failure occurs. Therefore, in
. using Equation 5-7 to derive the strength factor, the collapse load is defined as the median ultimate strength.
For SA 193 87 bolts, the ASME code specified ultimate strength is 125 ksi. For high strength steel, Reference 35 reports that the average strength is about a factor of 1.1 above the ASME code specified ultimate tensile strength. Considering the ASME code specified tensile
. strengthtobea95percentprobabilityofexceedancevalue(i.e..-1.65e) we'can calculate the logarithmic standard deviation for uncertainty in bol't strength ast 6g,3
- h
- in h = 0.06 (5-15) 5-30
Since the state of stress within the anchor bolt includes both shear and tensile stresses, an interaction equation was utilized in order to calculate the strength factor. The AISC code (Reference 36) recommends that the shear and tension stresses be combined utilizing Equation 5-16 below:
2 2 f f (5-16) d+is1.0 F F y
t where:
ft = Computed tensile stress fy = Computed shear stress Ft = Ultimate tensile stress Fv
= Ultimate shear stress Using a shear ultimate of 60 percent of the tensile ultimate and inserting values into Equation 5-16:
85.96 , 5.9 = 0.396 (1.1 x 125)2 (1.1 x 125 x 0.6)2 Since the normal stress in zero, the strength factor becomes:
f 3= 1/(0.396) 2 = 1.59 This states that the tensile and shear stresses could be scaled up by a factor of 1.59 before bolt failure will occur. The variability on this l
strength factor was calculated using Equation 5-8. Since the normal stress and its variability have been stated to be negligible, the BN tenn drops out of the equation. In addition, the SSE stress has been 5-31
l given in the qualification report and its variability is accounted for in the equipment response factor. Therefore, the variability on the applied stress is zero and drops out of Equation 5-8 since the BT term is a function of SN and SSSE.
What remains from Equation 5-8 is:
1 l
83=BC The variability on the capacity factor is made up of the 0.06 material strength variability from Equation 5-15 together with an uncertainty variability associated with the actual failure mechanism of the bolts,
. Sg. This latter uncertainty, Sm , accounts for miscellaneous conditions such as prying action, load paths to the bolt pattern, corrosion, etc. which are difficult to include within the analysis. The 95 percent confidence bound on failure is judged to be the yield point, and the resulting logarithmic standard deviation is 0.11. The SRSS 4 combination of these two variabilities gives the strength factor logarithmic standard deviation of:
~
S S = (0.06' + 0.11 ) 2 = 0.13 This variability is all uncertainty and the resulting strength factor and variability for the spray additive tank are:
is = 1.59 8
sg
= 0.0 BS U
As stated previously, the ductilty factor will be unity and the variability will be zero since the anchor bolts are considered to fail in a brittle mode. Therefort., the capacity factor and its variability are are equivalent to the strength f actor and its associated variability.
t 5-32 i
I
(
IC *I S = 1.59 eCR=SSR = 0.0 BCU=esg = 0.13 5.2.1.2 Spra.y Additive Tank Equipment Response Factor The response spectrum method was used to qualify the spray additive tank using Seabrook site specific spectra. As stated in Section 5.1.1.2.1.2, the qualification method response factor is unity and its variability is zero for this case.
The spray additive tank is located at Elevation 25' in the tank farm building. The tank has a sloshing frequency of 0.67 Hz and a constrained frequency of 3.3 Hz. Since the tank has a high slenderness ratio, the sloshing effect is negligible and the fundamental frequency was treated as being 3.3 Hz. The applicable equipment spectral shape 1 factor is contained within Table 5-1 in the "All Other Buildings" category for the 3 to 10 Hz range.
Iss = 1.43 -
SSSR
= 0.20
= 0.13 635U The modeling of the spray additive tank is considered to be median centered; thus, a factor cf 1.0 is applicable. The tank model is judged to be of median complexity, and as stated in Section 5.1.1.2.3, the applicable variability is:
BR = 0.0 and SU = 0.15.
The damping factor was computed as the ratio of the spectral accelerations for the design damping value of three percent and the median damping value of five percent taken at the tank fundamental l
5-33
frequency of 3.3 Hz (Equation 5-13). The three percent damped response used in the analysis was 1.14g's and the average between the two horizontal five percent damped spectra was 0.96g's. Therefore, I"
D
" 1 19 The variability on this damping factor was calculated from Equation 5-14 considering three percent damping to be about a minus one logarithmic
. standard deviation value.
SD g " A" (0 6)=0.16 Section 5.1.1.2.5 recomends a mode combination factor of unity and a random logarithmic standard deviation of 0.1E for a multi-degree-of-freedom system such as the spray additive tank. Table 5-5 reflects these values.
The spray additive tank is anchored via a circular bolt pattern.
For a tank of these proportions, the vertical earthquake component does not significantly affect the anchor bolt failure mode. Thus, earthquake component combination case number 5 from Table 5-2 applies and:
NECC = 0.926
~ = 0.06 SECCR
= 0.0 i SECCu The combined equipment response factor is then:
N ER
= 1.43 x 1.19 x 0.926 = 1.58 5-34 er
- w. y*-- - ,.,,wy y------ -,,-mw,
I The random and uncertainty variabilities are.
2 BERR
=
(0.202 + 0.152 + 0.06 )V2 = 0.26 2
S erg =
(0.132 + 0.152 + 0.16 )Y2 = 0.25 5.2.1.3 Spray Additive Tank Structural Response Factors The structural response factors for equipment are based on the data presented in Chapter 4 and are tabulated in Table 5-3. The tank
. farm building was assumed to remain elastic at the spray additive tank l fragility level. The structural response factors must be updated to the inelastic level on a trial and error basis if the fragility level of the tank exceeds the approximate structural yield level of 1.2g's. In this case, it did not and the appropriate structural rt.sponse factor and its variability from Table 5-3 are:
fSR = 1.20 B
SRg
= 0.31
= 0.16 ESRU 5.2.1.4 Spray Additive Tank Earthquake Duration Factor The anchor bolt failure on the spray additive tank is a brittle f - type of failure with very little system ductility. Therefore, a short duration earthquake is nearly as damaging as a longer duration earthquake and the appropriate duration factor is unity with a variability of zero.
5.2.1.5 Syray Additive Tank Ground Acceleration Capacity The ground acceleration capacity for the spray additive tank was calculated using Equations 5-1 and 5-2.
~
5 = 1.59 x 1.58 x 1.20 x 1.0 x 0.25g's = 0.75g's The variability was calculated by taking the SRSS of the variabilities for each of the four factors contributing to overall capacity (Equation -
5-3).
1 2
BR= (0.02 + 0.262 + 0.31 2+0.0)4=0.40 (randomness) 2 BU = (0.132 + 0.252 + 0.162 + 0.0 )V2 (uncertainty)
= 0.32 5-35 9
, , - . , , -%~,m__ _- . _ - - _, _ . - - - - _.,,--.-,,..,y -
The combined variability, S C
, is a measure of the overall variability contributed by earthquake randomness and uncertainty and can be obtained by taking the SRSS of SR and SU
- 2 8C = (0.40 2 + 0.32 )h = 0.51 This value of SC, along with the four factors making up the overall I I
. fragility (FEC, FER, FSR, FED) are tabulated in Table 5-12 along with the rest of the equipment which were addressed in the Seabrook PRA study.
5.2.2 Exawle of Qualification Test Report Fragility Derivation The battery chargers will be used as an example of a fragility derivation based on a plant specific test report. The battery chargers were qualified by subjecting them to 30-second duration simultaneous horizontal and vertical phase-incoherent inputs of random motion consisting of frequency band widths spaced one-third octave apart over the frequency range of 1 Hz to 40 Hz. The chargers were subjected to generic required response spectra which were designed to envelop the requirements for a number of nuclear power generating stations.
5.2.2.1 Battery Charger Capacity Factors During the seismic qualification testing of the battery charger, the indicator lights went out and the alam relay contact opened. The explanation given within the qualification report (Reference 37) was that the indicator light filament broke and made intemittent contact causing both a fuse and a relay contact to open. In spite of this anomaly, the
~
Wyle Laboratories test report states that the battery charger demonstrated sufficient integrity to withstand, without compromise of structural or electrical functions, the prescribed simulated environment. Due to the indicator light failure and the relay tripping, it is judged that the seismic SSE test level provided in Reference 37 is a fragility level for the battery chargers. It is estimated that any further increase in the loading would cause additional anomalies which could cause the functional failure of the chargers.
5-30
The battery chargers are located at Elevation 21'-6" within the control building. Their fundamental frequencies are not specified within the qualification report, but past experience has shown these electrical components to be in the 5 to 15 Hz range. The capacity factor was, there-fore, based on average spectral acceleration values in the 5 to 15 Hz range.
The test response spectra, from the test report, have the following average spectral accelerations in the 5 to 15 Hz range:
. Horizontal = 6.0g's Vertical = 7.5g's These loads were utilized to define the failure load, Pc, in EqJation 5-7. The normal loads on the battery charger are very small and were neglected. The seismic loading, PSSE, was taken from the ground response spectra which are the appropriate floor spectra for the battery chargers. Two percent damped ground spectra were utilized for comparison to the test spectra since the test response spectra were given for two percent damping. The average spectral accelerations from the ground spectra between 5 and 15 Hz are:
Horizontal = 0.86g's Vertical = 0.87g's Assuming that each earthquake directional component contributes equally to the functional failure mode of the battery chargers, the test level and the required level were ratioed by their resultant peak vectors.
2 p , (6 +2 7.5 )Y2 = 6.41 (0.862 + 0.862 + 0.872 )y Note that the test level has only two components since biaxial testing was conducted.
5-37
I l
The variability on the computed strength factor has two separate i components. The first variability component pertains to the uncertainty in the fundamental frequency of the battery charger. The strength factor was calculated based on average values in the 5 to 15 Hz range. From inspection of the spectra, a 5 Hz frequency will produce the lowest corresponding strength factor (FS = 5.38) within this frequency range.
Treating this as a 95 percent probability of exceedance (-1.658) value, the variability is calculated to be:
6 3 =h in ( h ) = 0.11 The second component of the strength factor variability, 83, stems from uncertainty in the exact failure threshold of the battery charger. No anomalies were reported during the OBE testing which was at one-half the level of the SSE tests; thus, the OBE test level was treated as a lower bound for fragility. Estimating this OBE test level to represent a -26 lower bound on strength:
e 3 = fin (f)=0.35 The overall strength factor variability due to uncertainty is
. then:
8 , 2 3 ,e3 p = 0.37 3
The ductility factor is unity and the variability on this factor ,
is zero since the ba'.tery charger failure mode is an acceleration sensitive functional failure. Thus, the capacity factor and its variability are equal to the strength factor and its variability.
5-38
IC "YS = 6.41 SCR "SR = 0.0 BCU "05U = 0.37 5.2.2.2 Battery Charger Equipment Response Factors The battery chargers were qualified to generic spectra whose safety factor and variability were accounted for in the strength factor.
Thus as described in Section 5.1.1.2.1.3, the qualification method factor and variabilties are:
fqg = 1.0 ,
SQMR"8QMU = 0.0 The battery chargers are located on the basemat in the control building and were qualified to the ground response spectra. The spectral shape factor and its variablity are unity and zero, respectively, for this condition since the ground _ spectra conservatism and variability are included in the structural response factor. Therefore, iss = 1.0 eSsR* SSU = 0.0 The qualification test report states that the battery chargers were installed onto the test table exactly as they are anchored in the field. Thus, as is discussed in Section 5.1.1.2.7:
l EBC = 1.0 8BCR"8800 = 0.0 5-39
The test response spectra (TRS) at two percent damping was compared to the required response spectra (RRS) at two percent damping to develop the strength f actor. Even though actual damping is expected to be much higher, approximately the same ratio exists between the TRS and RRS for median damping. Therefore, the damping factor is considered to be unity and the variability equal to zero.
I D = 1.0
. BDR"8Du = 0.0 The spectral test mcthods factor quantifies the variability involved in synthesizing a time-history to represent the RRS. Section 5.1.1.2.8 discussed this subjected and the recommended values are:
t I STM = 1.0 SSTMg = 0.0 B
STMg = 0.11 Multi-directional effects of using biaxial testing was accounted for in the development of the strength factor. In the strength factor I
development, the vector resulting from two components of biaxial excita-tion were compared directly to the vector resulting from three components of the median earthquake response. Thus, the difference between biaxial and three directional effects was accounted for and the factor, fMDE' equals unity. The uncertainty associated with the degree of coupling between earthquake directional components has also been accounted for within the strength factor derivation. There still exists the variability due to random phasing of the earthquake which was stated in Section
. 5.1.1.2.9.1 to be about 0.12.
MDE = 1.0 BMDER = 0.12 l '
c BMDE U
l l
5-40
The overall equipment response factor,ERf , and its logarithmic standard deviations, FER, and BERn, were calculated by taking the product of the factors aHd the SR35 of the logarithmic standard deviations of each of the contributing variables (Section 5.1.1.2).
I ER = 1.0 8
erg = 0.12 8
Eg = 0.11 5.2.2.3 Battery Charger Structural Response Factors The structural response factors listed in Table 5-3 are not applicable for the battery chargers since the chargers are located on the base mat and were qualified to ground spectra. Building response does not affect the response of equipment mounted on the basemat. Only the effects of soil-structure interaction and the spectral shape of the ground spectrum affect the response. From Chapter 4 the soil-structure interaction modeling for buildings at the Seabrook site resulted in a SSI factor of 1.0 and variabilities"of SR = 0.0 and eu of 0.05.
The spectral shape factor, in this case, is a measure of the conservatism involved in qualifying the battery chargers to the ground spectra instead of median site-specific spectra (Figure 3-1). The spectral shape factor was evaluated at the equipment fundamental frequency and for an estimated median damping level of five percent. The average spectral acceleration in the 5 to 10 Hz range is 0.70g's. The average spectral acceleration in the 5 to 10 Hz range is 0.50g's.
Therefore, the spectral shape factor is:
e f 33 O J0g 's = 1.40 0.50g's
! 5-41 l
t l
s Methodology utilized in deriving the random and the uncertainty portion of the variability on this spectral shape factor was developed in Chapter 4 for seismically critical buildings and are applicable to equipment mounted on the base mat. The resulting variabilities for the spectral shape factor on the battery chargers are SR = 0.30 and 80 = 0.10.
The overall structural response factor and its variability for
. the battery chargers was calculated by combining the soil-structure interaction f actor with the spectral shape factor, ISR = 1.40 85RR = 0.30 2
85Rg = (0.102 + 0.05 ) 2 = 0.11
! 5.2.2.4 Battery Chargers Earthquake Duration Factor Testing on battery chargers and similar electrical equipment has shown that failures due to seismic loads consists of chatter, breaker trip and similar electrical malfunctions. These functional type failures have no ductility associated with them; thus, the earthquake duration factor is unity and the variability is zero.
5.2.2.5 Battery Chargers Ground Acceleration Capacity ,
! The ground acceleration capacity for the ba'.tery chargers was calculated using Equations 5-1 and 5-2:
= 6.41 x 1.0 x 1.40 x 1.0 x 0.25g's = 2.24g's l The variability was calculated using Equation 5-3:
BR = (0.12 2 + 0.302 )Y2 = 0.32 2
l Su=(0.112 + 0.11 2 + 0.37 )Y2 = 0.40 l
I l 5-42 i
l l
l .
. - - . _. . . - . _ _ _ _ - _ = . - - - .-
Note that since k is greater than 2.0g's, Table 5-12 will contain "NA" (not applicable) for the individual response factors and their variabili-
, ties and ">2.0 g's" for the ground acceleration capacity. Equipment with ground acceleration capacities greater than 2.0 g's do not contribute significantly to the overall risk of the plant; thus, do not necessitate a detailed derivation and subsequent inclusion in Table 5-12.
5.2.3 Exangle of Generic Fragility Derivation Based on Design Specifications In the majority of cases in risk studies, all detailed informa-
' - tion regarding resulting stresses, deflections, bearing loads, etc., for safety-related equipment is not readily available to the risk analyst.
Classes of equipment must then be treated generically and the fragility descriptions derived from knowledge of design criteria, analytical methods, service experience, etc. In this section, an example of a i fragility description is developed which represents those items of equip-ment whose failure modes are structural and for which design reports or sumaries were not reviewed. Balance of plant piping and miscellaneous pressure vessels and heat exchangers are typically addressed in this manner. Seismic capacities of Class 2 and 3 piping and supports were derived in a generic manner and have been chosen as the example for this section. Tables 5-7 and 5-8 contain the fragility derivation parameters for the balance of plant piping and the piping supports, respectively.
5.2.3.1 Failure Modes of a Piping System In order to determine the most probable failure mode for piping, the design margins inherent for various pipe fittings, when designed to
- the governing code, must be compared to those for supports designed to
- the applicable codes. The fragility description for piping systems is based upon the single component type most likely to fail (i.e., pipe l
! fittings, straight pipe, pipe support, etc.). Failure of a pipe support does not necessarily mean failure of a piping system pressure boundary; however, the scope of this study does not permit side studies to deter-mine the increased probability of a piping system failure, given a support failure. Consequently, it is assumed that a support failure results in a failure of the piping system.
5-43
5.2.3.1.1 Piping Failure Modes - References 38 and 39 compare pipe fitting collapse loads to code allowable load for Class 1, 2 and 3 piping for Service Levels C and D. Both studies used almost identical data bases and both studies were based on current code criteria which are essentially identical to the Seabrook piping design criteria. For illustration purposes, development of generic capacities for Class 2 and 3 piping is portrayed. Capacities for the Class 1 piping systems were developed in a '
similar manner.
Equation 9 for Class 2 piping was the governing equation of Seabrook piping design for seismic induced loading combined with other loading and is:
0.751 M P*{n D +
Z
'"b h (5-17)
- The equation accounts for the axial stress due to pressure and the bending stress due to deadweight, hydrodynamic and earthquake induced moment, M . H a is a constant that depends upon the classification of the load combina-tion (i.e., normal, upset, eaergency or faulted) and Sh is a basic code allowable stress intensity. The criteria only take credit for 75 percent of the combined deadweight and earthquake moment; however, the combina-tion of 0.751 cannot be less than 1.0, where i is a stress intensification i
factor. This can have a slight effect on the most critical type of pipe.
f fitting selected. .
Reference 38 ranks pipe fittings in order of least to most conservative design as: _
- 1. Straight pipe
- 2. Elbows and bends l 3. Branch connections
- 4. Tees 5-44 b
- , . - _ . . . , _ . . . . . _ - , . - . - _ - _ _ . . ,..m,_m . . , . _ , ,c . . _ , . ~ -vr-+ + - - - - - - - - - - - - - - - '
l d
Review of the data base reveals, however, that at room tempera-ture the elbows have a slightly less conservative design basis than straight pipe. The same conclusion can be drawn from Reference 39. -
However, two factors must be considered in the ranking. First, for i
elevated temperatures, straight pipe has a slightly less conservative
! design basis than elbows. This is due to a change in the governing
, criterion for establishing the allowable Sh as temperature increases
. (i.e.,S ish based upon yield strength instead of ultimate strength).
Secondly, the largest moments usually occur at tenninal points in piping ,
l (anchors). Butt weld joints are then a logical candidate to define ;
~
fragility descriptions for piping since they occur at almost all terminal points, in most cases have less margin against failure if stressed to code allowables, and are most likely to contain flaws.
- 5.2.3.1.2 Support Failure Modes - Supports for restraint of seismic inertial loads can be in the form of snubbers or rigid rod type supports and can be both horizontal and vertical. Vertical rigid rod type supports i
must also carry deadweight; thus, they would carry proportionally less i
seismic load than theoretically allowed for lateral supports or vertical
) snubbers. If it is assumed that the resulting stresses in each support j type are at code allowable, a larger seismic margin would exist for vertical rigid supports than for lateral rigid supports or vertical j snubbers. Thus, the fragility description for supports is based on j supports that carry only seismic load. In the case of snubbers, the j snubbers themselves would be less likely to fail structurally under the l seismic loading than attachments to the pipe or the building.
I 5.2.3.2 Piping Capacit.y Factor
! The steps utilized in establishing a median factor of safety on j ; piping capacity are:
- 1. Establish a range of piping capacity
- 2. Estimate the range of loading on piping due to ,
weight, pressure and seismic events 1
i j
5-45 4
f l I i
[
- - -e- -- c w . w ww w ,e---------r-.~--ww.----+a - erm-----,------tw-v-,,,-w_ . wem + - m ww w- w w -. me-
- 3. Estimate the range of ductility
- 4. Estimate the threshold of piping system collapse vs individual pipe element collapse.
The range of piping system collapse is based on two extreme case models of piping failure. The upper bound on capacity can be represented by modeling the piping failure mechanism as occurring when the entire cross-section of the pipe is at the flow stress level. The flow stress is defined as midway between the yield strength and the tensile strength of the material. This value has been shown from tests to be an upper
- bound on capacity.
The lower bound on capacity can be represented by modeling the piping failure mechanism as occurring on a flawed piece of straight pipe with a circumferentially oriented flaw length equal to six times the pipe wall thickness. A flaw of this size is estimated to bound the possible flaws which could occur at butt welded joints. The theory for analyzing the moment capacites of through-wall flawed piping has been developed in Reference 40.
5.2.3.2.1 Piping Strength Factor - Strength factors were calculated for both of these extreme cases for a piping configuration that was estimated to be typical of the Seabrook plant critical piping. The assumptions made for this analysis were:
- 1. Flow stress around the cross-section of an unflawed pipe represents the +2 logarithmic standard deviation (+28) upper bound
- 2. A flawed pipe configuration with the net section at flow stress represents the -2 logarithmic standard deviation lower bound (-28)
- 3. A 10-inch pipe model was used
- 4. Two pipe thicknesses bound the possibilities, high energy lines (Schedule 160) and low energy lines (Schedule 40) t 5-46
~w--vv,,w------,---------,-w.- - - - - - - ~ - - - -
i
- 5. Pipe material was assumed to be 304 stainless !
steel type SA 312 J
A ten-inch pipe is considered representative of typical Seabrook piping, and the schedule 160 and the schedule 40 piping thicknesses are estimated
]
to be the thickest and the thinnest piping comonly used for safety-l related lines. In addition, past PRA studies on plants with piping .
! design criteria similar to the Seabrook Plant have shown that the collapse capacity of stainless steel fittings is more critical than for carbon steel, and that the upset (OBE) load case always governs the design in the absence of pipe break loading coupled with the SSE.
The median collapse load for piping, based on the values
, calculated from the +,28 bounding cases (i.e., the flow stress model and
' the through wall flaw model), was calculated to be 3.11 times the ASME code yield strength,oy(code)withalogarithmicstandarddeviationof -
0.16. These values are used later in Equations 5-7 and 5-8 to derive the
- strength factor and its variability.
For essential Class 2 and 3 subsystems, the resulting stresses from normal loading plus OBE are held to upset allowables (1.2 Sh )-
I The normal loading due to pressure and weight, combined with seismic
, loading, will vary considerably among piping systems. In a generic
) treatment of piping, fragility estimates for loading ranges must be I
made. In a piping system, the axial pressure stress typically will be
, less than 1/2 Sh and will never exceed this value. Weight supports are i
nominally spaced to result in deadweight bending stress of about 1500 psi. The combination of pressure plus weight stress is generally much less than the allowable value of Sh in order to accomodate seismic loading.
1 5-47
i Expressing the pressure and weight stress approximations in terms of the allowable stress and allowing for variations, the normal loading stress range for piping systems will typically be from about 0.35 to 0.7 times the allowable stress, Sh . This translates to a median normal stress of 0.5 x Sh and a logarithmic standard deviation of 0.17. The total stress (eN+ OBE) for Seabrook piping systems are assumed to typically vary from 0.65 to 1.0 times the allowable design stress of 1.2 Sh . This translates to a median total stress of approximately 0.80 x (1.2 x Sh ) with a logarithmic standard deviation
, of 0.13.
For SA 312 Type 304 stainless steel, the ratio of the code yield strength of oy (Code) to the allowable value of Sh varies.from 1.60 to 1.14 (depending on the operating temperature) with a median value of 1.35.
Thus, we can put the components of the strength factor all in like terms.
Collapse Load = c C = 3.11 x y (Code) = 3.11 x (Sh x 1.35) = 4.20 x Sh 8co11 apse = 0.16 Normal Stress =
N = 0.5 x Sh SN = 0.17 i
Total Stress =
T = 1.2 x 0.8 x Sh = 0.96 x Sh ET = 0.13 Using Equation 5-6, p , 4.20 - 0.5 = 8.04 S 0.96 - 0.5 Using Equation 5-8 83 = 0.36 5-48
i l
1 l
1 This strength factor of 8.04 was calculated based on the pipe being capable of reaching the flow stress. Piping capacity tests sumarized in Reference 38 have shown that thick-walled pipe are capable of reaching this flow stress, but that thin-walled pipe (schedule 40 or schedule 80) buckles before this flow stress is realized. Thus, the calculated strength factor of 8.04 is not applicable to thin-walled piping. For piping with D/t ratios (diameter to thickness) between 25
~
and 50, which is representative of the standard weights of pipe (schedule 40), the median shape factor for both stainless steel and carbon steel
. ranges from about 1.4 to 1.7. Since a broad range of pipe sizes, materials and schedules is being considered, a shape factor of 1.5 is selected as a median value for all piping within this D/t range. This median pipe capacity under static load is then 1.5 times the yield moment. Considering that the median yield strength is about 1.25 times the ASME code specified yield strength, the median moment capacity is about 1.87 times the yield moment determined from code yield properties.
The strength f actor for thin-walled piping is calculated using Equation 5-6:
F 2.52 x Sh - 0.5 x Sh = 4.39 3 = 0.96 x Sh - 0.5 x Sh where the Collapse Load = C = 1.87 x (1.35 x Sh) = 2.52 x Sh
- , The logarithmic standard deviation, SCollapse, remains the same at a 0.16 value; thus, the strength factor variability will be identical to that which was previously calculated (i.e. 85 = 0.36).
The overall peak ground acceleration capacity of the Seabrook plant critical piping is conservatively based on the buckling of the thin-walled piping.
i is = 4.39 -
8 s = 0.36 5-49
5.2.3.2.2 Piping Ductility Factor - Reference 6 recomends a ductility of 1.5 to 3 for design of critical piping systems. These are design recomendations; thus, the value of 3 is considered to be a median value with 1.5 representing an approximate lower bound or approximately a minus two logarithmic standard deviation value. Using these asstanptions and applying Equation 5-9, the r.edian factor of safety for ductility was ,
computed to be 2.24 with a lugarithmic standard deviation, Sy , of 0.24.
The random portion and the uncertainty portion of the ductility variability are considered to be about 0.16 each.
5.2.3.2.3 Piping "Three-Hinge" Factor - For complex piping systems, there is an additional source of design conservatism. ln order for a piping system to completely collapse, usually more than one collapse mechanism must form in the system; thus, basing fragility on the moment capacity of one fitting is conservative. A lower threshold of collapse could be likened to a simple beam where only one hinge is necessary for collapse.
An upper threshold of collapse could be likened to a fixed-fixed beam where three hinges must form. The elastica 11y calculated maximum moment i in the latter case would be 1.5 times the pipe element collapse moment.
These bounds were considered to be approximately a +28 range and the median system collapse factor was computed to be 1.22. The logarithmic l standard deviation, which is all uncertainty, is approximately 0.1.
Combining all the factors and variabilities results in a median capacity f actor of safety relative to the OBE and variability expressed in terms of logarithmic standard deviations of:
f C = 12.0 BCR = 0.16 8C0 = 0.41 5-50
5.2.3.3 Piping Equipment Response Factors The equipment response factors for piping are contained within Table 5-7 and are explained briefly below. The bulk of the piping was
- qualified by response spectrum analysis; thus, a factor of unity and a variability of zero are applicable for qualification method. The spectral shape factor was taken from Table 5-1 in the 5 to 20 Hz frequency range for the "All Other Buildings" category. This category we.s used from Table 5-7 for the example, while individual factors based on the location of the piping system were used on a piping system by piping system basis.
The modeling of the piping systems is felt to be median centered and of median complexity, thus, i IM = 1.0 6pR = 0.0 Sgu = 0.15 The damping factor was computed by comparing response for a two perce.nt damped spectrum to response for an expected median damping value of five percent at or near failure. Two percert damping was utilized in designing large diameter piping to the OBE event. For the example problem, spectra for Elevation 53' in the Primary Auxiliary Building was used. Applying Equations 5-13 and 5-14:
ID = 1.34
- 8 0R = 0.03 8
Du = 0.17 i
The mode combination factor is unity with random variability of 0.15 as suggested in Section 5.1.1.2.5 for multiple degree-of-freedom systems.
5-51
The earthquake component combination factor for piping is considered to be the general case (Case No.1 in Table 5-2) since all three directions of earthquake motion generally excite the piping system.
The overall equipment response factor, PER and the variability, Bg and 60 for piping were computed to be:
. P ER " 2 14
, SERR = 0.28
+
8ERu = 0.28 5.2.3.4 Piping Structural Response Factors The structural response f actors for piping are a function of the piping location (building and clevation). These structural response f actors are shown in Table 5-3. Piping systems are typically in the 5 to 20 Hz frequency range and values in Table 5-3 corresponding to this frequency band were used. The structural response factor used as an exangle in Table 5-7 for piping is the most conservative value within Table 5-3, and is for the containment internals.
5.2.3.5 Piping Earthquake Duration Factor _
Piping is a very ductile system and the duration factor and
- associated logarithmic standard deviations are specified in Section 5.1.1.4 to be:
FED "l4
- S EDR = 0.12 BEDU = 0.08 5-52
.,-----.e , , - - , - , , - ,- - - - - - - - - - - - - - - - - - - . , . , .---,---e--e
- - ..--- -- --m --- ---
5.2.3.6 Pipino Ground Acceleration Capacity The ground acceleration capacity for piping was obtained by' multiplying the four f actors within Equation 5-1 by the OBE level of 0.125g's, since the capacity factor was developed based on the OBE and not the SSE as is generally the case.
A = 4.99g's 8R = 0.43 8U = 0.53 ,
5.2.3.7 Piping Supports Capacity Factor
' It was stated previously that for supports stressed to their design limit under seismic conditions, the minimum margin would occur if the load were all seismic (i.e., the support carried no normal load).
The fragility description for supports was developed on this basis.
I In order to apply Equation 5-7 to compute a strength factor or -
t upper and lower bound factors, the range of material properties, the ;
- support failure mode, the allowable design load and a range of applied l
load for the seismic event must be established.
i Piping supports are generally constructed of carbon steel, and three of the most comon carbon steels were utilized for this generic study. The three carbon steels are SA 36, SA 675-G70 and SA 516-G70.
t The ratio of the allowable stress value, S h to the yield strength was evaluated for each of these materials with a resulting median ratio of 0.53 and logarithmic standard deviation of 0.11.
1 Almost all piping supports in a modern nuclear power plant such as Seabrook are welded carbon steel structures, and from analytical experience the most critical section of these supports due to seismic j loading is the welds. Since the pipe supports at Seabrook were designed .
- (
a ,
5-53 ,
to current ASME Code criteria contained in Appendix XVII of the Code for linear type supports. Appendix XVII specifies that for Service Level B (UpsetCondition):
OBE + Deadweight + Thermal s 0.4 Sy(shearinwelds)
, where Sy = ASME Code specified yield strength.
For Service Level D (Faulted Condition) the allowable of 0.4 Symay be
. increased up by the lessor of "1.2 x (cy /ea llowable)" or "0.7 x
' ult /8 allowable)"whereinthiscase allowable is the tension allowable of 0.6Sy . Using the properties of A36 carbon steel, the lesser of these factors is 1.88. Thus, the allowables for Level O conditions are:
SSE + Deadweight + Thermal + Hydrodynamic s 0.75
- Sy Since deadweight and themal are taken to be zero, the OBE loading condition gcverns, just as it did for the piping.
A classical strength of materials analysis of fillet welds which are subjected to a combined state of shear and tensile loads, as are typically produced in pipe supports, has shown that yield will occur in the weld throat when the principal stress is 73 percent of the material
. yield strength. The variability on this 73 percent factor can be quantified by estimating the pure shear condition (shear yield is
- approximately 60 percent of the yield strength) to be 95 percent
- probability of exceedance value (-1.658) and a logarithmic stanhrd deviation of 0.12 results. The yield level in the weld throat was utilized as the collapse load in determining the strength f actor (Equation 5-7), realizing that the ductility factor will quantify the supports' capacity past the yield level and into the plastic region, oc = collapse Stress = 0.73 x cy 5-54
It was assumed that the median seismic stress is 0.7 times the allowable design load for welds, 0.4 x Sy , with 1.0 times the design load being about a +1.658 (90 percentconfidenceupperbound). Thus, the median OBE stress and its logarithmic standard deviation are:
C OBE = 0.7 x 0.4 x Sy = 0.28 S y
. 8 0BE = in = 0.22 For the most limiting case, the normal stress is zero and Equation 5-7 becomes:
f5 . *C
'0BE References 35 and 41 have established that the median yield strength of both carbon and stainless steels is about 25 percent above the ASME code specified yield strength with a variability of 0.14. Therefore.
I S= 0.73 0.28 x 1.25 x5 x5 = 3.26 y
The overall logarithmic standard deviation on the strength factor was computed using Equation 5-8 to be 0.29.
The system ductility for piping systems is considered applicable to supports since piping may be well into the inelastic range prior to f ailure of a support. The median ductility factor. Fy. is then about 2.24 with logarithmic standard deviations due to both randomness and
. uncertainty of 8 =8 = 0.16.
Combining the strength and ductility factors and their varia-bilities, the resulting capacity factor relative to the OBE and its variabilities were computed to be:
5-55 1
l l
i l I EC = 7.30 8ECR = 0.16 ,
l 8ECg = 0.33 l
Pipe sgports have a much lower capacity factor than piping and would be the governing element in piping systems.
Concrete anchors used to attach the pipt supports to the
- surrounding walls are not qualified with the supports themselves; thus,
!, must be addressed separately. The NRC has required a factor of safety of i four for wedge anchors and a factor of safety of five for shell anchors l (Reference 42) to be used'in the anchorage design of piping and equipment.
} These safety factors are based on the median ultimate capacity of the particular anchor bolt in question in relation to the design allowable.
l ,
In the design of a piping system, the load on a pipe support will j generally be less than the allowable and it is estimated that the median l
load on a support is 70 percent of the design allowable. Thus, strength
[ ,
factors of 5.7 and 7.1 exists for the wedge anchors and the shell anchors, j i respectively.
1 i ,
i Effective ductility for the anchorage pullout depends upon the j degree of inelasticity in'the piping system. If the piping system i remains elastic up to the point of anchor pullout, the failure mode is j brittle, the elastica 11y calculated load is valid and the ductility 1
factor would be 1.0. On the other hand if the piping system were highly l inelastic, the calculated support loading would not develop and the i , ductility factor computed for piping systems would be appropriate. The ,
I effective ductility factor can then be assumed to range from 1.0 to '
j 2.24. The median ductility factor is about 1.5 with a logarithmic l standard deviation of 0.2. Combining the median strength and ductility l factors results in:
t
! ,[
1 i
- 5-56 b
._...._..___._m.. -
W I EC = 8.55 (for wedge anchors)
(for shell anchors)
PEC = 10.65 Thus, the concrete anchor bolts will not govern the fragility description of piping systems since piping supports have a lower capacity factor.
5.2.3.8 Ground Acceleration Capacity of Piping Supports
- The equipment response factor, the structural response factor and the duration f actor for piping supports are identical to those which were developed for piping. Table 5-8 contains all of the fragility f actors for pipe supports. The ground acceleration capacity of pipe supports and
, its variability are: .
I = 3.039's fR = 0.43 60 = 0.47 l
Pipe supports are then the governing failure mode for piping systems and were utilized in their fragility descriptions. However, since the ground i acceleration capacity of the pipe supports is greater than 2g's they will j' not contribute significantly to the overall plant risk.
5.2.4 Example of Fragility Derivation Based Upon New Analysis A new analysis was conducted for the refueling water storage tank in order to remove some of the conservatisms and uncertainties that are inherent in scaling from the vendor's design analysis.
The RWST is 44'-0" inside diameter and is composed of six rings of different thicknesses. The tank has a spherical dome roof and is anchored to a slab in the tank farm area at EL 20'. Amplified response spectra applicable to the tank location in its supporting structure were specified by UE&C and are contained in the vendor design report, Reference 46. The tank is anchored by 46, 2-inch diameter SA 193 Grade 8-7 high strength bolts.
5-57
w ,
' e Il y7 ,
'~'
+ Ttie Yank was analyzed using computer program TANK. TANK computes the fundamental frequency and base shear and overturning moment for a specified input acceleration. The program was developed around the methodology of Veletsos. Reference 46. The fundamental flexural frequency
'I of the tank was computed to be 7.0 Hz. The first 2 modes of sloshing frequency were computed to be 0.26 Hz and 0.44 Hz. The tank was assumed
^ to be rigid in the vertical direction.
The base shear for the specified design spectrum was 2100 kips and the overturning moment was computed to be 53,700 ft/ kips. These
- loads were used to assess the potential failure modes in the tank.
i l' Several potential failure modes were examined to determine the governing capacity. Failure Modes examined included: ,
Anchor bolt tension in threads Anchor bolt shear and tension in unthreaded shank
! ! Shell buckling
,C ,
, Anchor bolt chair gusset welds Anchor bolt chair gusset capacity i .
Anchor bolt chair top plate I
It was determined that the minimum capacity was controlled by the anchor .
s bolt chair gussets. They are subjected to a combination of compression and bending and when a fully plastic section (compression plus bending) was formed, the gussets were assumed to become unstable, allowing tank
.lif t-off' and subsequent buckling of the compression side.
5.2.4.1 RWST Capacity Factor The tensile bolt load at failure of the bolt chair gussets was computed to be 266 k. The computed tensile load from the specified SSE anplified response spectrum was 106 k. The resulting strength factor is d computed as:
1.'
- FS = 266/106 = 2.51 5-58 I
, , - - , - - - - - - - - , ~ , - _ . . ,__ _ . , . _ , _ _ , _ .-._n_ . _ _ - . . _ _ , _ _ , , - _ . , . _ , . . - - . , _ . , . . . - _ , . , , ,. ,n,.__. , _ _ .._ ,.,_, -_
Tnere are two sources of uncertainty in the bolt chain capacity, the uncertainty in the strength calculation and the uncertainty in the ,
material properties. Failure level of the bolt chair gussets is a complex problem to assess. The failure model assumed that when the free edge of the bolt chair gusset was fully plastic, i.e., the combination of axial load and bending produced a fully plastic section, the increased eccentricity would result in the gussets buckling. The next higher failure mode was computed to be tensile failure in the anchor bolt in the thread area. Median tensile capacity was calculated to be 344 kips. It was reasoned that plastic buckling of the gussets would possibly either not occur or that the deformed bolt chairs would not result in tank failure until the ultimate strength of the bolts was reached. The bolt ultimate strength was assumed to be a 95% confi:lence upper bound. The uncertainty,SU , on the failure strength was computed to be:
SU strength
=(1/1.65)in (344/266) = 0.16
, The uncertainty on gusset material strength is well defined. Median strength is about 25% above the code specified strength, where code strength is a 95% confidence value, Reference 41. The uncertainty on material strength is then:
eg = (1/1.65) in 1.25 = 0.14 The resulting uncertainty on strength is:
8 '
3 U
There is no random variability assigned to strength thus, 83 = 0.
R The failure mode, whether gusset buckling or bolt tension, will occur with very little displacement and is considered a brittle failure; thus, there is no credit taken for ductility.
5-59
5.2.4.2 RWST Equipment Response Factor 1
The variables that must be addressed in deriving an equipment response factor and its uncertainty include:
Qualification Method Spectral Shape Damping Modeling Mode Combination Earthquake Component Combination 5.2.4.2.1 Qualification Method - The dynamic analysis was conducted by the response spectrum method which is considered to be median-centered.
The uncertainty in the response analysis is defined by.other variables; thus, the qualification method factor and its variability are considered to be unity and zero, respectively.
5.2.4.2.2 Spectral Shape Factor The aniplified response spectra, ARS, 4
used in the analysis were peak broadened 10% and smoothed. The ARS peaked at 2.5 Hz wherein the tank fundamental frequency was 7.0 Hz; thus, there shoJld not be a large factor of conservatism well away from the spectral peak. Comparisons of raw and smoothed spectra were not readily available for the tank locations and the spectral shape factor was conservatively assumed to be 1.1. This is consistent with factors in Table 5-1 for frequencies well away from the peak of the ARS. Since the tank was mounted very low in its supporting structure, the ARS ZPA was only amplified about 20%, thus it was assumed that there was virtually no additional conservatism applied in generating the ARS. The uncertainty in the spectral shape factor was computed to be:
8 33 =(1/1.65)tn(1.1/1)=0.06 Random variability,8 33 , is considered zero for the equipment spectral R
j shape factor.
( 5-60 l
~
5.2.4.2.3 Damping Factor - The ARS used in the response analysis was for 3% damping. Median damping is considered to be 5%. Five percent damped ARS were not specified for the tank location. The ratio of 3% to 5%
amplified response was, therefore, estimated from ground motion amplifica-tion factors contained in Reference 10. The ground motion amplification factors are applied at the tank fundamental frequency. Use of ground motion aglification factors is reasonable, since the ARS spectral shape is similar to that for the ground motion input at the tank fundamental frequency. The resulting factor is conservatively estimated as:
Sa3% _
. F D
" 1*18
" Sa5%
. Three percent dag ing was assumed to be an approximate 95% lower bound;
~
thus:
8 0 (1/1 65) in 1.15 = 0.09 0
The actual equipment damping is considered to be all uncertainty; thus, SD = 0.
R 5.2.4.2.4 Modeling Factor - The dynamic model was considered to be median-centered, FM = 1.0 The geometry is simple and the error in calculating the fundamental frequency was estimated to be no greater than 15%. A 15%
error in frequency at the tank frequency results in very little difference in spectral acceleration due to the shallow slope of the ARS
, at 7 Hz. The resulting uncertainty, e g, was only about 0.01. Sg is considered zero for modeling.
5.2.4.2.5 Mode Combination Factor - Sloshing and tank flexural mode response were combined by the square-root-of-the-sum-of-the-squares, SRSS, rule which is considered to be median and the response factor is considered to be unity. The tank response is dominated by the flexural l
mode response; thus, the variability is very small. Mode combination variability is considered random anu:
I egg = 0.03 R
l l 5-61
5.2.4.2.6 Earthquake Component Combination Factor - The base shear and overturning moment were computed for one direction of response. The SRSS response combination for vertical, circular components would result in no increase in bolt load. There is, however, a possibility of having two orthogonal acceleration vectors in-phase, in which case, the resulting vector would be d times the specified horizontal acceleration for each direction. An estimate of the most probable phasing is taken from Reference 10 where it is suggested that 100% of the load in one direction be combined with 40% of the load in the other two orthogonal directions.
Vertical acceleration contributes very little to the tank failure mode; thus, the resulting response factor for the horizontal resonse is:
FECC = 1/(1 + 0.4 )b= 0.93 The phasing is all random variability. Considering the ratio of completely in-phase vectors to out-of-phase vectors to be a three log standard deviation spread, s
g = (1/3) In 1.4 = 0.14 R
5.2.4.2.7 Overall dquipment Response Factor - The overall equipment response factor and its variability reflects the degree of conservatism and uncertainty in the tank response analysis that was conducted using the specified SSE ARS.
R
= (1)(1.1)(1.15)(1.0)(1.0)(0.93) = 1.18
. S
= . . = 0.12 ER R
S '( + + + + + '
ER U
5.2.4.3 RWST Structural Response Factor The RWST is mounted in a Tank Farm structure for which ARS were generated. The ARS peaks at 2.5 Hz reflecting the fundamental frequency 5-62
of the structure. Concrete structures were analyzed for a damping level of 7%. The variables to be considered in quantifying the conservatism and uncertainty in the structural response are:
Spectral Shape Damping Modeling Soil-Structure Interaction 5.2.4.3.1 Spectral Shape Factor - At the 2.5 Hz fundamental frequency of the structure, there is a pronounced difference in the amplification between the Reg. Guide 1.60 spectrum specified for design and the site-specific spectrum considered to be median for f. agility analysis. Figure 3-1 compares these two spectra. The spectral shape factor is defined as:
Sa R.G. 1.60 F
33 = 3U
= 1.81 site-specific The-variability of the amplification in the site-specific spectrum varies with frequency. From data in Reference 29, at 2.5 Hz, the random variability is:
8 = 0.38 33 p The uncertainty in the spectral shape,833 , is estimated to be U
approximately 1/3 633 R
6 b3 0
- 5.2.4.3.2 Damping - Structure response was calculated using 7% damping.
For structures that are below yield, 7% is considered to be median. The supporting structure is considered to be below yield and the design danping is considered to be median with the resulting damping factor being unity (FD = 1.0). Ten percent (10%) damping is considered to be an approximate plus one log standard deviation value, thus:
Sa ,
E = 0.10 D
V " Sa7%
5-63 f
l
_m... _,.
. . . , _ _ _ _ . _ . _ _ _ . , _ _ _ ~ . , . _ , _ . . . ~ . _ _ _ , _ _ . _ ~ . . . . ._
There is some randum variability assigned to damping that is a function of the earthquake time history input. The BD is estimated to l R
be equal to BD*
U 5.2.4.3.3 Modeling Factor - The structural model was considered to be median-centered.
Fg = 1.0 The structural model is simple and the uncertainty, 6 is estimated to be 0.1. S g is taken as zero.
R 5.2.4.3.4 Soil-Structure Interaction - Seabrook is a rock site and all structures are either founded on rock or on concrete fill. Therefore, uncertainty in using fixed-base structural models is very low and estimated to be:
8 337 U
5.2.4.3.5 RWST Structural Response Factor - The overall structural response factor and its variability are:
F SR
= 1.81 (1.0)(1.0)(1.0) = 1.81 .
E SR (0.382 + 0.102 + 0 + 0) = 0.39 R
e 3p = (0.132 + 0.102 + 0. d + 0.05D = 0.20 g
5.2.4.4 Earthquake Duration Factor The tank failure mode is considered to be essentially brittle.
The duration factor is accordingly considered to be unity with zero variability, i
5-64
-P - .~..
~ _
5.2.4.5 RWST Capacity The peak ground acceleration capacity of the RWST is the product of the four factors time the SSE peak ground acceleration.
A=(2.51)(1.18)(1.81)(1.0)(0.25)=1.34g The variability is the SRSS of the variabilities of the four factors:
sg = (0 + 0.122 + 0.392 + 0)b = 0.41
- sg = (0.212 + 0.112 + 0.202 + 0)b = 0.51 5.2.5 Example of Fragility Based on Engineering Judgment and Earthquake Experience There are several equipment items within the list of components for the Seabrook PRA for which no seismic qualification was required.
These components were not designed for seismic loading; thus, they will generally have a lower capacity and a higher uncertainty than seismically qualified components. The methodology which has been utilized on the previous examples of developing capacity factors, response factors, duration factors, etc., is ~ generally not applicable for unqualified components. The fragility levels for components must be derived based on earthquake experience and engineering judgment. The example which has been chosen in this category is the reserve auxiliary transformer.
This transformer is a large steel structure which sits unanchored on a concrete slab. It is located on the base mat, so building amplifica-4 -
tion will not be a consideration. Inspection of the transformer during the Seabrook site visit led to the judgment that sliding of the trans- ,
former relative to the slab would occur before overturning of the trans-former. A static coefficient of friction between steel and concrete of 0.4 is recomended in Reference 43. Due to uplift effects from the vertical portion of the earthquake combined with the horizontal ,
components, an acceleration level of about 0.39's could be tolerated i
5-65 l
l I
before transformer sliding will occur. The derivation of this 0.39 ground acceleration fragility level was based on engineering judgment combined with seismic analysis using simplified models.
As a check of the transformer fragility level which was derived in the previous paragraph, some earthquake experience data exists for transformers which are unanchored. Reference 44 reports that an unanchored two station auxiliary transformer sustained significant damage during the 1971 San Fernando Earthquake. This transformer experienced as much as 18 inches of movement relative to its concrete pad and the
. resulting damage included broken 4.16 KV bushings, damaged bushing terminal plates and terminals, control box malfunction and surface conduit failures. Reference 44 reports that the San Fernando Valley earthquake which caused this transformer damage had an. estimated peak ground acceleration in the range of 0.3g to 0.5g. Thus, actual experience confirms the 0.39 capacity calculated by approximate methods. There is considercble variability on this capacity due to both the uncertainty involved in the failure mechanism and the randomness of the earthquake.
Earthquake experience has shown that for numerous small earthquakes in the 0.1g peak ground acceleration range very little damage has occurred to unanchored equipment. Thus, the 0.1g level was used as a 98 percent probability of exceedance value (-2.08) which results in an overall ! =
0.55. Based on engineering judgment this can be broken down to ER" 0.25 and SU = 0.50. Table 5-9 shows the ground acceleration capacity and its associated logarithmic standard deviations for loss of offsite power which would include the reserve auxiliary transformer as well as other high voltage electrical equipment in the switchyard as well as on
! , the grid.
5.3 EQUIPMENT FRAGILITY RESULTS Table 5-9 contains fragility descriptions for all of the equip-ment which were included in the PRA. Fragility derivations were conducted j for each of the components and are reported for those items which have a ground acceleration capacity less than 2g's. Pickard, Lowe and Garrick, Inc. has stated that equipment which possess ground acceleration capaci-ties greater than 29 's will not contribute to the overall plant risk because of the extremely low frequency of events causing this level of i
5-66
acceleration; thus, detailed fragility descriptions need not be included.
Equipment which fall into th' category have been labeled with a
" > 2.0g's" in the ground act mtion capacity column, and have "NA" (not applicable) in the respcnse ic. and variability columns. In addition, for those components ich have capacities less than 2.09's, only the capacity factor, equipment response factor, structural response factor and the earthquake duration factor are given in Table 5-9. Intermediate factors which make up these four main factors were determined but not included in the table. .
. 5.3.1 General Results The following are some general results on the equipment portion of the PRA:
- 1. The majority of the equipment within the Seabrook plant which were seismically qualified have rela-acities. This tively is duehigh ground to the acceleration relatively high SSEcap (0.25g ZPA) j
, coupled with the more sophisticated qualifi-cation techniques and applicable codes which are associated with equipment in a modern plant.
- 2. Non-seismically qualified components listed in Table 5-9 generally have relatively low capacities and relatively high uncertainties.
These results are understandable due to the lack of a requirement for seismic design.
- 3. The category of " instrument sensors" within Table 5-9 represents a group of information gathering
(
, devices (thermocouple probes, pressure trans-ducers,flowtransmitters,etc.)whichhavebeen shown during vibration testing to be inherently rugged. The weak link in these instrument sensor systems is estimated to be the small tap lines which connect the piping or component to the transmitters. These lines are judged to have a capacity similar to that of small piping which is greater than 2g's.
- 4. Where failure modes are listed in Table 5-9 as
" chatter," the consequences of chatter should be determined by the systems analyst by examining the electrical circuits and capability of the i
l operators to recover from any spurious signals or trips resulting from chatter. '
I 5-67
[
TABLE 5-1: EQUIPMENT SPECTRAL SHAPE FACTORS Peak Broadening Overall Spectral Shape and Smoothing Factor Equipment Building Frequency (Hertz) F 33 8
33 F
33 s 33 s 5S 1 1 U R Containment 3-10 1.23 0.10 1.35 0.10 0.2
. Building 10-15 1.24 0.11 1.36 0.11 0.2 15-20 1.16 0.07 1.28 0.07 0.2 20-33 1.05 0.02 1.16 D.D2 0.2 5-20* 1.21 0.11 1.33 0.11 0.2
. Rigid 1.0 0.0 1.1 0.0 0.2 Primary 3-10 1.47 0.19 1.62 0.19 0.2 ('
Auxiliary 10-15 1.24 0.11 1.36 0.11 D.2 Building 15-20 1.35 0.15 1.48 0.15 D.2 20-33 1.30 0.13 1.43 0.13 0.2
- 5-20* 1.35 0.19 1.48 0.19 0.2 Rigid 1.0 0.0 1.1 D.0 0.2 Control / Diesel 3-10 1.25 0.11 1.37 0.11 0.2 Generator 10-15 1.13 0.06 1.24 D.05 0.2 Building 15-20 1.09 0.04 1.20 0.04 0.2
, 20-33 1.17 0.08 1.29 D.08 0.2 ,
5-20* 1.16 0.11 1.28 0.11 0.2
, Rigid 1.0 0.0 1.1 ,
D.0 0.2 Fuel Buildings 3-10 1.25 0.11 1.37 0.11 0.2 10-15 1.60 0.23 1.76 ' O.23 0.2 15-20 1.11 0.05 1.22 0.05 0.2 20-33 1.10 0.05 1.21 0.05 0.2 5-20* 1.32 0.23 1.45 0.23 0.2 Rigid 1.0 0.0 1.1 0.0 0.2 ,
All Other 3-10 1.30 0.13 1.43 0.13 0.2 Buildings 10-15 1.30 0.13 1.43 0.13 0.2 15-20 1.18 0.08 1.30 0.08 0.2 20-33 1.15 0.07 1.26 0.07 0.2
, 5-20* 1.26 0.13 1.39 0.13 0.2 Rigid 1.0 0.0 1.1 0.0 0.2 ,
- Piping, Cable Tray and Conduit are typically in the 5-20 Hz range depending on the support configuration.
?
5-68 L
. . o l
j TABLE 5-2: EARTHQUAKE COMPONENT COMBINATION FACTORS i
i 1
D h tion l '
Case Number Responses Description F 8 8 ECC R 0 1 2 Horizontal + General Case (Piping) 1.15 0.12 0.10 Vertical 2 2 Horizontal General Case, Coupled 1.11 0.10 0.08 3 2 Horizontal Rectangular Anchorage Failure Mode Uncoupled, 1.01 0.12 0.0 Designed by SRSS of the resultant stresses 4 2 Horizontal Circular Anchorage Failure Mode, Uncoupled, 1.31 0.06 0.0 4
Designed by SRSS of the applied moments 1
, 5 2 Horizontal Circular Anchorage Failure Mode, Uncoupled, 0.926 0.06 0.0 g Designed by SRSS of the resultant stresses 6 1 Horizontal + Coupled 1.04 0.07 0.05 Vertical 7 1 Horizontal + Uncoupled 1.15 0.03 0.0 Vertical 8 1 Horizontal General Case 1.0 0.0 0.0 I Refers to the earthquake directional components which affect the failure mode under consideration
TABLE 5-3: STRUCTURAL RESPONSE FACTORS FOR EQUIPMENT Approximate Elastic Structure Inelastic Structure Building Yield Level) F SR 8
R 8
U F
SR 8
R 8
U (g's)
Waste Processing /
1.2 1.20 0.31 0.16 1.45 0.24 0.16 Tank Farm Containment 2 2 2 Internals high 2 1.11 0.26 0.16 NA NA NA
. Fuel Storage 2.15 1.14 0.22 0.16 1.25 0.19 0.16 Containment 2.87 1.22 0.30 0.16 1.46 0.25 0.16 Control /0iesel 1.15 1.16 0.31 0.14 1.37 0.23 0.16 Primary Auxiliary 1.2 1.20 0.31 0.16 1.45 0.24 0.16 5ervice Water Pump Structure 1.28 1.24 0.14 0.16 1.26 0.10 0.15 RHR/ Containment
! Spray Vault 1.2 1.20 0.31 0.16 1.45 0.24 0.16 Cooling Tower 0.90 1.19 0.30 0.16 1.43 0.25 0.15 2
0.30 0.16 2 2 gg 2 Pipe Chase high 1.17 NA NA Emergency Feed-water Pumphouse 0.90 1.24 0.33 0.16 1.50 0.28 0.17 -
Containment Enclosure Ventilation Area 1.2 1.15 0.22 0.13 1.26 0.18 0.13 l IThe approximate yield level estimates the ground acceleration level at which yielding is reached in the building. Equipment with capacities less than this value should utilize the elastic structural response factors.
The containment internals and the pipe chase have very high yield levels, and thus the 1/2 yield structural response factors will always be appropriate.
5-70
_ . _ _ _ _ _ . . - . , _ _ _ . -~ _ . , . _ _ _ . _ . , .
TABLE 5-4: NUCLEAR POWER PLANT EQUIPMENT CATEGORIES Relative Capacity Equipment Category Level High 1. Piping, Ducting, Cable Trays and Electrical Conduit
- 2. All Valves (Except small motor operated valves)
Medium-High 3. Small Vessels and Heat Exchangers
- 4. Horizontal Pumps. Compressors and Turbines
- 5. Fans and Air Conditioning Units
- 6. Diesel Generator
- 7. Reactor Coolant Loop Components Medium 8. Small Motor Operated Valves
- 9. Large Vessels and Heat Exchangers
- 10. Batteries and Racks
- 11. Vertical Pumps ~
- 12. Reactor Internals and Control Rod Drive Mechanism Low-Medium 13. Motor Control Centers, Switchgear, Control Panels, Instrument Racks
- 14. Non-seismically Qualified Components (e.g. Offsite PowerSystem) i f
5-71
TABLE 5-5: FRAGILITY DERIVATION OF THE SPRAY ADDITIVE TANKS Median Random Uncertainty Factors Safety Variability Variability Factor og s u
. Capacity Factor (FEC) i 1. Strength Factor 1.59 0.0 0.13
- 2. Ductility Factor 1.0 0.0 0.0 Combined- F 1.59 0.0 0.13
, EC 4
Equipment Response Factor (FER}
- 1. Qualification Method 1.0 0.0 0.0
- 2. Spectral Shape 1.43 0.20 0.13
- 3. Modeling 1.0 0.0 0.15
- 4. Damping 1.19 0.0 0.16
- 5. Combination of Modes 1.0 0.15 0.0
- 6. Earthquake Component Com-bination 0.926 0.06 0.0 Combined - F EC 1.58 0.26 0.25 StructuralResponseFactor(FSR) 1.20 0.31 0.16 Earthquake Duration factor (FED) 1.0 0.0 0.0 l Ground Acceleration Capacity (A) 0.75 g's 0.40 0.32 e
6 5-72 d
. . , _ _ _ _ . - _ _ _ , _ , . , _ , _ _ - _ . _ _ _ . _ _ _ _ _ _ - . _ , _ _ , _ _ _ . , , _ , , _ _ _ _ , . . _ _ _ . . _ , _ _ . _ _ _ , _ , , , , _ . , . . , . _ , . _ _ _ , _ , . _ , , . _ - _ _ _ - _ _ . . - . _ . , _,y_, ,
TABLE 5-6: FRAGILITY DERIVATION OF BATTERY CHARGERS Median Random Uncertainty Factors Safety Variability Yariability Factor ag 8 U
Capacity Factor (FEC)
- 1. Strength Factor 6.41 0.0 0.37
- 2. Ductility Factor 1.0 0.0 0.0 Combined--- F 6.41 0.0 0.37 EC Equipment Response Factor (F I ER
- 1. Qualification Method Factor 1.0 0.0 0.0
- 2. Spectral Shape Factor 1.0 0.0 0.0
- 3. Boundary Conditions Factor 1.0 0.0 0.0
- 4. Damping Factor 1.0 0.0 0.0
- 5. Spectral Test Method Factor 1.0 0.0 0.11
- 6. Multi-Directional Effects Factor 1.0 0.12 0.0 Combined --- F ER 1.0 0.12 0.11 StructuralResponseFactor(FSR) 1.40 0.30 0.11 EarthquakeDurationFactor(FEO) 1.0 0.0 0.0 GroundAccelerationCapacity(A) 2.24 g's 0.32 0.40 l
5-73
TABLE 5-7: FRAGILITY DERIVATION OF BALANCE OF PLANT P! PING Median Random Uncertainty Factors Safety Variability Variability Factor og og
. Capacity Factor (FEC)
- 1. Strength Factor
- 4.39 0.0 0.36
- 2. Ductility Factor 2.24 0.16 0.16
- 3. 3 Minge Factor 1.22 0.0 0.10 Combined - F EC
- 12.0 0.16 0.41 EquipmentResponseFactor(FER) .
- 1. Qualification Method 1.0 0.0 0.0
- 2. Spectral Shape 1.39 0.20 0.13
- 3. Modeling 1.0 0.0 0.15
- 4. Damping 1.34 0.03 0.17
- 5. Combination of Modes 1.0 0.15 0.0
- 6. Earthquake Component Combination 1.15 0.12 0.10 .
Combined - F ER 2.14 0.28 0.28 StructuralResponseFactor(F SR P* 1.11 0.26 0.16 EarthquakeDurationFactor(FED) 1.40 0.12 0.08 Ground Acceleration Capacity (A) 4.99 g's 0.43 0.53
- Based on OBE (0.125 g's)
- Based on Containment Internals Structural Response Factor and E 's which are the Most Conservative 5 74
)
i TABLE 5-8: FRAGILITY DERIVATION OF PIPING SUPPORTS Median Random Uncertainty Factors Safety Variability Variability Factor S R
8 U
. Capacity Factor (FEC)
- 1. Strength Factor
- 3.26 0.0 0.29
- 2. Ductility Factor 2.24 0.16 0.16 Combined - EC F* 7.30 0.16 0.33 EquipmentResponseFactor(FER)
- 1. Qualification Method 1.0 0.0 0.0
- 2. Spectral Shape 1.39 0.20 0.13
- 3. Modeling 1.0 0.0 0.15
- 4. Damping 1.34 0.03 0.17 .
- 5. Combination of Modes 1.0 0.15 0.0
- 6. Earthquake Component Combination 1.15 0.12 0.10 Combined- F ER 2.14 0.28 0.28 StructuralResponseFactor(F 7* 1.11 0.26 0.16 SR 1.40 0.12 0.08 Earthquake Duration Factor (FEO)
Ground Acceleration Capacity (A) 3.03 g's 0.43 0.47
- Based on OBE (0125 g's) .
- 8ased on Containment Internals Structural Response Factors which are the Most Conservative 1
1 i
5-75
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.,n,r,-m.-,.,.,, a.,--...,-,.-,----.-,-n,,,.,,.---- ~
.,---n ., n- . , - , - - - - - - - - - - , . -
TABLE 5-9
SUMMARY
OF SEABROOK COMPONENT FRAGILITIES The following pages (5-77 through 5-78) comprise Table 5-9 O
e 5-76 4
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u :.. ::. tu n: .:p . !* :! u
/ ...:. . . .u .
. lf
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!. . . .a . 53 ! .. n! .!!. . 33 ::
. u.!. .u!.n
/ ...:. u.: o ..n . ::. g in ..: .::. . :::: u a
I .up su:n. ...o . :p : p. n: .:p . :::: ..
/
I l .= u.!. so:n. ...n . g:. : :n u! .::. . :::: u g
S
- / u .i. u
- .o ...n . g:. 3 n:. .p . :::: n E
I f .u!. u.:.o ..u. . ::. : su n: .:p . :::: ..
s '1 ... 3.. .. 33 . :::: ..
. ...:. ...... . . ... . i 1
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'I 1 i
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1 III lI, ;
+ il lilli -l l l !
! !!!Il ill: 1Itil 1:1 I ii m fiff i ii II i ji jil fill I ii li lI l -
jjjij tilli 1111111 11111111 1 111 I g;gji i 111 Ill I 11 til Illi i Ii 11 h L,,s lli l'I l Il i i, l l3 ::: !... It 1 I .N l
- 11 se;;I!, t, 1***. .u 11 11 , j; l l ,,uI .m. ,,
. .i ...
I. I.. i;'i ffffff! t
,If I fI til !!!! . I I r
g fitr5 1111 1111111 a6lll ii... } }I ! lI 111 1111 1 1:1: l l l I lhg!
! 1 ishl ll"'lh>ihIioai l l i "' d " 'l '
i @lLI.m.I I,s...i.>. .!.i.a 1..i l P
.a i in i n "i I. F ill i
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8 HORIZONTAL (E.'W) RESPONSE.EL.75' k o AMPLIFIED RESPONSE SPECTRA - SSE
- CONTROL / DIESEL GENERATOR BUILDING
& 7% STRUCTURAL DAMPlNG; 2% EQUIPMENT DAMPING I FINAL SAFETY ANALYSIS REPORT FIGURE 3.7(8) 29. SH. 2 O
~
a zo co-H G E wo Smoothed and Broadened u G f
#8 M k l u
w
- 0.
Woo k 8 h c E l 1.0 ; i 3 i 6 6 i66'10" i h i $ $ i $ $ 'l O' NATURAL PERTOD (SECONDS) FIGURE 5-1: COMPARISON OF A SMOOTHED AND BROADENED SPECTRA TO AN UNSM00THED AND UNBROADENED SPECTRA 5-79 .
.,.--n. , . , _ - . . - - . . , , . . , , , , , , . _ . . _ . . _ , _ _ , . . _ _ . _ . . . _ . _ . , _ _ _ . _ _ _ , . - - _ _ _ _ - , _ _ _ _ _ , _ . . . _ , _ . , _ . - - , . , , . . , _ , _ _ _ _ _ , , - - , , _ . _ , , , _
REFERENCES
- 1. Seabrook Station Final Safety Analysis Report, Public Service Company of New Hampshire, Revision 45, June, 1982.
- 2. USMtC, " Design Response Spectra for Seismic Design of Nuclear Power Plants" USNRC Regulatory Guide 1.60, Revision 1. December, 1973.
- 3. Freudenthal, A. M., J. M. Garrelts, and M. Shinozuka, "The Analysis of Structural Safety", Journal of the Structural Division ASCE.
- ST 1, pp. 267-325, February, 1966.
- 4. Kennedy, R. P., A Statistical Analysis of the Shear Strength of Reinforced Concrete Beams, Technical Report No. 78 Department of Civil Engineering, Stanford University, Stanford, California, April, 1967.
- 5. Newmark, N. M., "A Study of Vertical and Horizontal Earthquake Spectra", WASH 1255, Nathan M. Newmark Consulting Engineering Services, prepared for USAEC, April,1973.
- 6. Newmark, N. M., " Inelastic Design of Nuclear Reactor Structures and Its Implications on Design of Critical Equipment" SMiRT Paper K 4/1, ,
1977 SMiRT Conference, San Francisco, California.
- 7. Riddell, R., and N. M. Newmark, " Statistical Analysis of the Response of Nonlinear Systems Subjected to Earthquakes", Department of Civil Engineering, Report UILU 79-2016 Urbana, Illinois, August, 1979.
- 8. Bernreuter, D. L., " Seismic Hazard Analysis, Application of Methodology, Results, and Sensitivity Studies", NUREG/CR-1582, Vol. 4. Lawrence Livermore National Laboratory, October,1981.
- 9. USNRC, " Damping Values for Seismic Design of Nuclear Power Plants".
USNRC Regulatory Guide 1.61, October, 1973.
. 10. Newmark, N. M., and W. J. Hall, " Development of Criteria for Seismic
. Review of Selected Nuclear Power Plants", NUREG/CR-0098, May,1978.
- 11. Kennedy, R. P., et al., "Probabilistic Seismic Safety Study of an Existing Nuclear Power Plant", Nuclear Engineering and Design, Vol. 59, No. 2, pp. 315-338.
- 12. USNRC, " Combining Modal Responses and Spatial Components in Seismic Response Analysis". USNRC Regulatory Guide 1.92, Rev.1, February, 1976.
R-1
N REFERENCES (Continued)
- 13. Letter from H. L. Ruffner, Pittsburgh Testing Laboratory Seabrook Site Manager, to R. A. Rebel, Resident Construction Manager for United Engineers and Constructors, Inc., Septeder 7,1982.
- 14. Troxell, G.E., H.E. Davis and J.W.' Kelly, Composition and Properties of Concrete McGraw-Hill, 1968.
- 15. United Engineers and Constructors, Inc., Calculation Set No. MA-29
. (final), sheets 62 through 137 of 162 Revision 0, May 27, 1981.
- 16. Mirza, S. A., M. Hatzinikolas, and J. G. MacGregor, " Variability of Mechanical Properties of Reinforcing Bars", Journal of Structural Division. ASCE, May, 1979.
- 17. ACI 318-77, " Building Code Requirements for Reinforced Concrete",
American Concrete Institute, 1977.
- 18. Barda, F., J. M. Hanson and W. G. Corley, " Shear Strength of Low-Rise Walls with Boundary Elements". ACI Symposium, " Reinforced 4 Concrete Structures in Seismic Zones", ACI, Detroit, Michigan, 1976.
- 19. Shiga. T., A. Shibata and J. Tabahashi, " Experimental Study on Dynamic Properties of Reinforced Concrete Shear Walls" Conference on Earthquake Engineering, Rome, Italy,197b.Sth World
! 20. Cardenas. A. E., et al., " Design Provisions for Shear Walls", ACI Journal, Vol. 70, No. 3. March, 1973.
- 21. Oesterle, R. G., et al., " Earthquake Resistant Structural Walls -
Tests of Isolated Walls - Phase !!" Construction Technology Laboratories (Division of PCA), Skokie. Illinois, October,1979.
- 22. Hadjian, A. H., and T. S. Atalik " Discrete Modeling of Symmetric Box-Type Structures", International Symposium on Earthquake Structural Engineering, St. Louis, Missouri, August, 1976.
- 23. " Earthquake Resistant Structural Walls - Tests of Isolated Walls".
PB-271 467, Portland Cement Association, prepared for National . science roundation, November, 1976.
- 24. " Tentative Provisions for the Development of Seismic Regulations for Buildings", ATC 3-06, prepared by Applied Technology Council for i National Science Foundation, June, 1978. .
l
- 25. Merchant. H. C., and T. C. Golden, " Investigations of Bounds for the Maximum Response of Earthquake Excited Systems", Bulletin of the Seismological Society of America Vol. 64 No. 4, pp. 1239-1244, August, 1974. .
R-2
I REFERENCES (Continued)
- 26. " Final Report. Testing and Seismic Qualification of the Seabrook Control Room Ceiling", prepared by CYGNA Energy Services, San Francisco, California, Revised April, 1982.
- 27. Ang, Alfredo H. and Wilson H. Tang, Probability Conceots in Engineerina Plannine and Desien, John vuey ana sons, Inc.,1975.
- 28. NUREG/CR-1706, UCRL-15216. " Subsystem Response Review, Seismic Safety Margin Research Program". October,1980.
- 29. " Amplified Response Spectra for Seismic Category I Structures" J.0.
9763.006 Public Service Cogany of New Hagshire, February 15, 1980.
- 30. Smith, P. D. and O. R. Maslenikov, "LLNL/ DOR Seismic Conservatism Program Part 111: Synthetic Time Histories Generated to Satisfy NRC Regulatory Guide 1.60" UCID-17964 (draft report), Lawrence Livermore Laboratory, Livermore, California, April, 1979.
- 31. Smith, P. D., 5. Bumpus and 0. R. Maslenikov. "LLNL/ DOR Seismic Conservatism Program, Part VI: Response to Three Input Components".
UCID-17959 (draft report), Lawrence Livermore Laboratory Livermore, California, April, 1979.
- 32. " Seminar on Understanding Digital Control and Analysis in Vibration Test Systems", sponsored by Goddard Space Flight Center, Jet Propulsion Laboratory and The Shock and Vibration Information Center held at Goddard Space Flight Center on 17-18 June 1975 and at the JPL on 22-23 July 1975. P
- 33. Hardy, G.S. and R. D. Campbell, " Development of Fragility
. Descriptions of Equipment for Seismic Risk Assessment of Nuclear Power Plants", paper to be presented at the ASME Pressure Vessel and
; Piping Conference in Portland, June, 1983.
- 34. " Seismic Qualification Report for the Seabrook Spray Additive
. Tanks", by Pittsburgh-Des Moines Steel Company, July,1981. Foreign 3
Print No. 52802-08.
- 35. NUREG/CR-2137, " Realistic Seismic Design Margins of Pumps, Valves
and Piping", by E.C. Rodabaugh and K. D. Desai. June,1981.
! 36. Manual of Steel Construction, Seventh Edition. American Institute of 5 teel Construction, Inc. i 37. " Qualification Program for the Class 1E Battery Chargers for the Seabrook Nuclear Generating Station", by Power Conversion Products, j Inc., Document No. QR-15629UE&C Foreign Print No. 31601. !I i R-3 i i l
- REFERENCES (Continued)
- 38. NUREG/CR0261, ORNL/ SUS-2913/8, " Evaluation of the Plastic Characteristics of Piping Products in Relation to ASME Code Criteria", by E. C. Rodabaugh and S. E. More, Battelle Columbus Laboratories, July, 1978.
- 39. Sargent and Lundy Report, " Evaluation of the Functional Capability
- of ASME Section Ill, Class 1, 2 and 3 Piping Components MARK I Containment Program, Task 3.1.5.4", September 21, 1978.
- 40. Witt, F. J., W. H. Samford and T. C. Esselman, " Integrity of the
- Primary Piping Systems of Westinghouse Nuclear Power Plants During Seismic Events", WCAP No. 9283 Westinghouse Electric Corporation, March, 1978.
- 41. ASTM DS-552, "An Evaluation of the Yleid. Tensile, Creep and Rupture Strengths of Wrought 304, 316, 321 and 347 Stainless Steels at Elevated Temperatures", American Society of Testing Materials.
- 42. IE Bulletin No. 79-02, " Pipe Support Base Plate Designs Using
. Concrete Expansion Anchor Bolts."
- 43. PCI Design Handbook, Published by Prestressed Concrete Institute, First Edition, Second Printing 1971.
- 44. " San Fernando Earthquake of February 9, 1971: Effects on Power System Operation and Electrical Equipment", Prepared by the Design and Construction Division of the Department of Water and Power of the City of Los Angeles. October, 1971.
- 45. Pittsburg-Des Moines Steel Company Design Report, Refueling Water Storage Tank for Seabrook, New Hampshire. June 1981 UE&C Foreign Print Number FP 52801.
- 46. Veletsos. A. S. and J. Y. Yang, Dynamics of Fixed-Base Liquid Storage Tanks, presented at U.S.-Japan Seminar for Earthquake
, Engineering Research with Emphasis on Lifeline Systems, Tokyo, Japan, November 8-17, 1976. .
- 47. USNRC letter dated April 4, 1985.
Subject:
Seabrook PRA Review. G. W. Knight to R. J. Harrison.
- 48. Pickard, Lowe and Garrick, Inc., Seabrook Station Probabilistic Safety Assessment, Report PLG 0300, December 1983, prepared for Public Service of New Hampshire.
R-4
E APPENDIX A CHARACTERISTICS OF YHE LOGNORMAL 015TRIBUT!0N
. Some of the characteristics of the lognormal distribution which - are useful to keep in mind when generating estimates of A, ag, and 8 U are sumarized in References Al and A2. A random variable X is said to be lognormally distributed if its natural logarithm Y given by:
Y=in(X) (A-1) is normally distributed with the mean of Y equal to in i where i is the median of X, and with the standard deviation of Y equal to 8, which will be defined herein as the logarithmic standard deviation of X. Then, the coefficient of variation. COV, is given by the relationship: 2 COV=Yexp(8)-1 (A-2) For 8 values less than about 0.5, this equation becomes approximately: C0Vis e (A-3) and COV and 8 are often used interchangeably. For a lognormal distribution, the median value is used as the
- characteristic parameter of central tendency (50 percent of the values areabovethemedianvalueand50percentarebelowthemedianvalue).
The logarithmic standard deviation, 8. or the coefficient of variation, COV, is used as a measure of the dispersion of the distribution. A1
59 Therelationshipbetweenthemedianvalue,i, logarithmic standard deviation, 8, and any value x of the random variable can be expressed as: , x=i exp (n 8) (A-4) wherenisthestandardizedGaussianrandomvariable,(meanzero, standard deviationone). Therefore, the frequency that X is less than any value x' equals the frequency that n is less than n' where: ni. in(x'/i) (A-5) Because n is a standardized Gaussian random variable, one can simply enter standardized Gaussian tables to find the frequency that n is less than n' which equals the probability that X is less than x'. Using c mulative distribution tables for the standardized Gaussian random variable, it can beshownthati*exp(+8)ofalegnormaldistributioncorrespondstothe 84 liercentile value (i.e., 84 percent of the data fall below the + 6 value). Thei*exp(.8)valuecorrespondstothevalueforwhich16 percent of the data fall below. One implication of the usage of the legnormal distribution is that if A, B, and C are independent lognormally distributed random vari. ables, and if e Ar , gs , D= t 4 (A-6) A-2
s
/
where q, r, s and t are given constants, then D is also a lognormally distributed random variable. Further, the median value of D denoted by 6, and the logarithmic variance 8h, which is the square of the logarith- , mic standard deviation, SD , of D, are given by* 6= ,' q (A-7) C and 2 22 22 22 g (A-8) 0=rSA+s68+t8C where 4, 5 and C are the median values, and 6A , 6g, and 8C are the loga-rithmic standard deviations of A, 8, and C, respectively. i .
' The formulation for fragility curves given by Equation 2-1 and shown in Figure 2-1 and the use of the lognormal distribution enables easy development and expression of these curves and their uncertainty.
However, expression of uncertainty as shown in Figure 2-1 in which a range of peak accelerations are presented for a given failure fraction is not very usable in the systems analyses for frequency of radioactive release. For the systems analyses, it is preferable to express uncertainty in terms
, of a range of failure fractions (frequencies of failure) for a given ground acceleration. Conversion from the one description of uncertainty
.to the other is easily accomplished as illustrated in Figure A-1 and sunnarized below.
e With perfect knowledge (i.e., only accounting fcr the random -
- variablity 8 A), the f ailure fraction, f(a), for a given acceleration a can be obtained from: .
f f(a) = ti '"I'/ I \ 8 (A-9) ( R/ A-3
*9 e
, ._ _ _ _ _ _ . , _ _ . . _ _ _ _ . _ _ _ _ _ - - . _ . - _ _ _ _ _ _ , . _ _ . _ . _ _ _ _ . _ _ _ _ _ _ , _ _ _ _ . _ _ _ _ . _ _ _ _ _ . _ ._.-,_.4,_ _,._ .__ _ ,.
in which 4(-) is the standard Gaussian cumulative distribution function, 1 and SR is the logarithmic standard deviation associated with the underlying randomness of the capacity. . For simplicity, denote f = f(a). Similarly, f' is the failure fraction associated with acceleration a', etc. Then, with perfect knowledge (no uncertainty in the failure fractions), the ground acceler-ation a' corresponding to a given frequency of failure f' is given by: a' = exp Sg c'*(f') (A-10) The uncertainty in ground acceleration capacity corresponding to a given frequency of failure as a result of uncertainty of the median capacity can then be expressed by the following probability statement:
" "^
P A > a"l f' =1C (A-11)
. U .
in which P[A > a"lf'] represents the probability that the ground accelera-i tion A exceeds a" for a given failure fraction f'. This probability is shown shaded in Figure A-1. However, it is desirable to transform this probability statement into a statement on the probability that the failure fraction f is less than f' for a given ground acceleration a", or in symbols P[f s f'la"). This probability is also shown shaded in Figure A-1. It follows that: P[f s f'la"] = P[A > a"lf'] (A-12) l Thus, from Equations A-10 and A-11: l ina"/Iexp 8p e~l(f') P[f s f'la"] = 1- C g U _ (A-13) A-4
- - , - - . - -,,--.n - . - - - - , , - , , - , , . ,---,-,y,, -, - -- - -- -
from which:
' ~ *
(ina"/Iexp8 p 4
-1 (f')g}
P[f > f' la"] = c g (A-14) which is the basic statement expressing the probability that the failure fraction exceeds f' for a ground acceleration a" given the median ground acceleration capacity d, and the logarithmic standard deviations BR and , , Sg associated with randomness and uncertainty, respectively. As an example, if: 4 E=0.77, BR = 0.36, BU = 0.39 then from Equation A-14 for typical values of f and a", P[f > 0.5 la" =0.40g]=0.05 which says that there is a 5 percent probability that the failure frequency exceeds 0.5 for a ground acceleration of 0.40g. i e G A-5
i . . . i -
, I i
,/
/
/
/
Failure Frequency / 1
/
in(a/M ~ /j f(a) = e 8 ~ f - R -
/
' /
/
] / 1
. T /
U /
- -.... - -- p g 0.5 . . .
9 f a 1 8 k > E l l 4 a e f,, - . --- .--- . g/ Probability
= 1-0 in(a"/A) l P A > a"lf'
/
#1 l
- g / -
9 i '. - . - . I f' 4 I / f I # ) I i 1/ l I . . . ; I P
.fsf'la".-= P A>a"If' l l .
{
#p l l -
8 8 l #p 1 1 8 I
# 1 I 3
! p# e e l a J s 1 0 v , a' a" A i Ground Acceleration, a l FIGURE A-1. RELATIONSHIP BETWEEN UNCERTAINTY IN GROUND ACCELERATION FOR l AND UNCERTAINTY IN FAILURE FRACTION FOR A GIVEN GROUND ACCELERATION i
REFERENCES A.1 Genjamin, J. R., and Cornell, C. A., Probability. Statistics and Decision for Civil Engineers, McGraw Hill Book Company, New York,
, 1970.
A.2 Kennedy, R. P., and Chelapati, C. V., " Conditional Probability of a Local Flexural Wall Failure of a Reactor Building as a Result of
, Aircraft Impact" Holmes and Narver Inc., prepared for General a Electric Company, San Jose, California, June, 1970.
6 4 l e I r
. A-R-1 l
l l l i
3P 4- . ; ff gDf
/ g) f , f fu o1 pi .e v G o
,4 o Y f p (
/f, e o o a ad f ,t; The cormlent numbers correspond to report sections identified with the same number:
(1) Page 5: Other key parameters to include in this table are:
- a. Containinent design - reinforced.
- b. Steam generator secondary inventory - 112,000 lbm/ steam generator. -
(2) Page 21: After the vessel blowdown transient to about 70-80 psia, the RCS pressure builds up rather slowly. (3) Page 22 and 23: The SSPSA used 0.1%/ day as the design basis leak rate per Amendment 47. (4) Page 23 and 24: The SSPSA used 100% in 2 hours as the boundary between Type A and Type B containment leaks. Anything treater would be treated as a single puff release. This is conservative compared to the BNL definition of 100% in 1 hour. (5) Page 28: The Type A leak area of 0.5 square inches per pipe is based on failure of the weld at the worst and probably weakest location in the annular seal plate. (6) Pages 28, 29 and 31: Simultaneous failure of all 4 feedwater lines is unlikely given the uncertainties in the failure pressure. After failure of the first feedwater pipe, the leak area would be large enough to result in a steady or decreasing containment pressure. This would prevent failure of the other feedwater pipes. Even if all four feedwater pipes were to fail, the combined leak area would not be much larger than i that for a single pipe failure. The failure mode is self-regulating and the total leak area is determined by the pressurization source, i.e., the steaming rate or the gas generation rate. (7) Page 30: Failureofthecaterpurgeisolation'valveandtheelectrica[ penetrations is very unlikely for the same reason. The heat losses into the concrete are too large to raise the temperature in the penetration before mass transfer into the interspace stops. Steam condensation leads to a noncondensable atmosphere in the interspace which is in pressure equilibrium with the containment. From this point on, only conduction through the sleeve can transfer heat into the penetration. , fe3p C H 1 8/E7
i., -, l (8) .Page 30: Electrical penetration integrity is based on: ! a. Location of the plugs on the outside of the penetration.
- b. Calculated heat losses to the concrete. (See comment 7).
(9) Page 31: Use of Figure 3.10 would not have changed any of the SSPSA results. l (10) Page 31: The SSPSA took no credit for the secondary enclosure building except for intact containment leakage. Even type A and B leaks develop at a pressure where interference at the equipment hatch is expected to fail the enclosure locally. (11) Page 33: Impaired evacuation was modeled for the earthquake scenarios, assuming larger delay times and slower evacuation speeds. l (12) Page 38: For top event 12 on the containment event tree, both paths are containment failure. Success is small leakage and failure is gross leakage. (13) Page 42: Release Category S2 only represents sequences where the leak rate increases substantially (to 401/ day) at vessel melt-through. It does not represent failure of the Type A penetrations at 181 psia during slow pressurization sequences. Release Category 52 was introduced because a PWR1 or PWR2 type release at Seabrook is extremely unlikely. The probability of accident sequences with an isolated containment resulting in an 51 or S2 release category was based on an uncertainty analysis for the pressure spike at vessel melt-through compared to the containment pressure capacity; i.e., Pr(52) = Pr (P at VMT)181 psia) Pr(51) = Pr (P at VMT>211 psia) Plant damage states for steam generator tube rupture sequences were assigned to the most appropriate release category without an attempt to compute SGTR specific release categories. (14) Page 51: The energy release of 3 x 10 9 Btu /hr only applies to gross containment failures and resulted from two considerations, namely:
- 1. Mechanistic failure propagation for containment membrane failure modes.
- 2. An energy release which limits the plume rise to the inversion laybr.
The latter limitation (2) was found more constraining. ,e (15) Page 52: The start of the second puff release coincides with the beginning of the vaporization release. The duration of the vaporization release is only 2 hours but the duration of the second puff is 7.2 hours to better balance the release fractions associated with each puff. In any event, the entire release for the second puff occurs at the beginning of the second puff. Therefore, the duration is only of secondary importance. 2 1
, , , - , - - , . - , - - - - , - , - - - , , , , . . , . , . _ _ , . - - - - - _ - -----v-- - , - - , - - _ - - - - - - , , y - ,-----~,-e,: , ,r, ,- . , -
i ~ (16) Page 52: We feel it would be appropriate to point out that your recommendation for reducing release times and durations, as well as the conversion to single puff releases is only for the purpose of bringing the release category parameters within the CRAC code limitations and not l to'make the release more realistic. We do not believe that the RSS times are more reasonable. l Page 56: To us the single puff release concept is not a realistic one (17) because we now know that all the risk significant releases at Seabrook extend over several hours. The purpose of an " equivalent" single puff release can only be to either assess the conservatism of a single puff release versus a multipuff release or to bring the analysis within the constraints of the CRAC code. In either case, this should be clearly stated. (18) Pages 52 and 56: Table 4.7 lists BNL recomended release characteristics which are based on the MARCH / CORRAL calculated point estimates, without taking into consideration any of the advances since WASH-1400. In the SSPSA source term uncertainty analysis, it is shown that these release characteristics represent something like a 99: confidence level for nonexceedance. We think it would be appropriate to list in a separate table the single puff equivalents of the SSPSA best estimate release categories. These would be U-d, DV-d, (T6VI-d + IFV2-d + 55V3-d) and ($2V1-d + 52V2-d + DV3-d). These could be listed simply for comparison without a recommendation, to give the future user a choice of using whatever he believes to be appropriate for his purpose. (19) Pages lii and 60: We believe that the following two conclusions are equally significant: o Local containment failures with self-regulating leak areas are shown to occur before gross containment failure. These result in extended releases with reduced consequences. _. D o An analysis of the source term uncertainties has shown that the best estimate source terms can be expected to be between one and two orders of magnitude lower than the point estimate releases ' calculated for conservative accident sequences using the WASH-1400 methodology. (20) Page 52: The point estimate release categories were determined on the basis of a containment capacity corresponding to a wet containment condition. For the dominant release categories which are all dry I containment conditions, the release times and release fractions were corrected to the dry conditions and are_shown in the uncertainty analysis. Therefore, release category $2V-a on Tab _le 11.6-16 should be used for this comparison. This release category (S2V-a) reflects the*
)
release timing calculated for this type of accident sequence using the l ( WASH-1400 methodology. l JK O f/s % n p./rwfuwu g//f t f/ - /-- ). g f y 3
1 6 . 47 , em r , BNL/NUREG- I NUREG/CR- i A REVIEW OF THE SEABROOK STATION PROBABILISTIC SAFETY ASSESSMENT: CONTAINMENT FAILURE MODES AND RADIOLOGICAL SOURCE TERMS M. Khatib-Rahbar, A. K. Agrawal, H. Ludewig and W. T. Pratt September 1985 Department of Nuclear Energy Brookhaven National Laboratory Upton, New York 11973 e S l l l 1
9 s . 111 ABSTRACT e A technical review and evaluation of the Seabrook Station Probabilistic Safety Assessment has been performed. it is determined that (1) containment response to severe core melt accidents is judged to be an important factor in mitigating the consequences (2) there is negligible probability of prompt containment failure or failure to isolate, (3) failbre during the first few hours after core melt is also unlikelg (4) the point-estimate radiological releases are comparable in magnitude tf those used in WASH-1400, and (5) the energy of release is somewhat higher 'an for the previously reviewed studies.
/9 n1-s % ..in y u .
jaLa K Aaky k w/w WA:y/-/400 w ..a' x .h-
/
F-0 I e e 0 ew
=
0
,-~+.,e, n--,-, - - - - , , , - -
-r,---, -- r-- - - - - - - - - -- - - - - - - ,- , - - - , , -- -
? 6 .
O
. iv
. ACKNOWLEDGMENT
( i s e om e l -
t >
.v.
CONTENTS
- Page l l
iii J ABSTRACT................................................................ iv ACKNOWLEDGMENT.......................................................... v1 LIST OF TABLES......c................................................... vii LIST 0F FIGURES......................................................... I
- 1. INTR 000CTION.......................................................
I 1.1 Background.................................................... I 1.2 Obj ecti v e s a nd Sc o pe . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.3 Org a ni za ti on o f the Re po rt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2
- 2. PLANTDESIGNANDFEATURESIMPORTANTTOSEVEREACCIDENTANALYSIS....
2 2.1 As s e s sment of Pl ant De sign . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 2.2 Compa ri son wi th Othe r Pl ants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
- 3. ASSESSMENT OF CONTAINMENT PERFORMANCE..............................
7 3.1 Background.................................................... 8 3.2 Conta i nme nt Fa il u re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 3.2.1 Background............................................. 8 3.2.2 De s i g n De s c r i pt i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . l 21 3.2.3 Leakage Rate Calculation............................... 22 3.2.4 Co n t ai nmen t Fa il u re Mod e1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 l 3.2.4.1 Le a k-Be f o re-Fa il u r e . . . . . . . . . . . . . . . . . . . . . . . . . . . 23 3.2.4.2 Cl assi fication of Fail ure. . . . . . . . . . . . . . . . . . . . . 24 3.2.5 Cont ainment Pre s sure Ca pa ci ty . . . . . . . . . . . . . . . . . . . . . . . . . . 24 3.2.5.1 Concrete Containment.......................... 27 3.2.5.2 Liner......................................... 27 3.2.5.3 Penetrations.................................. 3.2.5.4 Containment Fail ure Probabil ity. . . . . . . . . . . . . . . 31 + 31 3.2.5.5 Co nt ai nment Encl o s u re . . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Definition of Plant Damage States and Containment l 31 Re s po n se Cl a s s e s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33 3.4 Containment Event Tree and Accident Phenomenology............. 38 3.5 Contai nment Matri x (C-Matrix) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44 3.6 Rel ease Category Frequencies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48
- 4. ACCIDENT SOURCE TERMS..............................................
48 4.1 Asse ssment of Severe Accident Source Te rms . . . . .. . . . . . . .. .. .. .. 52 4.2 So u rc e Te rm Unce r t a i n ty An al y si s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56 4.3 Re c omme nd e d So u rc e Te rm s . . . . . . . . . . . . . . . . .
- 5.
SUMMARY
AND CONCLU S 10 N S . . . . . . . . . . . . . . . . . . . 61
- 6. REFERENCES.........................................................
o G e
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LIST OF TABLES Titl e Pace Tabl e Comparison of Selected Design Characteristics... .... . .... . . ..... . 5 2.1 3.1 Con tai nment Ope rating and De sign Pa ramete rs . . . . . . . . . . . . . . . . . . . . . . 10 18 3.2 Containment Linir Penetrations................................... 29 3.3 Leak Area Estimates fc Mechanical Penetrations................... Frequencies of Occurrence of the Plant Damage States............. 35 3.4 3.5 Co ntai nment Re sponse Cl ass De finitions . . . . . . . . . . . . . . . . . . . . . . . . . . . 36 3.6 Co ntai nment Cl as s Mean Fre quenci e s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37 3.7 Accident inase and Top Events for the Seabrook Containment 39 Event Tree....................................................... 3.8 Release Categories Employed in the Seabrook Station Risk 40 Model............................................................ 41 3.9 Simpl i fi ed Cont ainment Matrix for Se ab rook. . . . . . . . . . . . . . . . . . . . . . . 3.10 Frequency of Dominant Release Categories (yr-1)...... ... .. . .. . . . . 45 3.11 Contribution of Containment Response Classes to the Total Co r e Mel t F re qu e n cy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 3.12 Release Category Frequency as a Fraction uf Core Melt 47 Frequency........................................................ 4.1 Se ab rook Poi nthEstimate Rel ease Categori es.. . . . . . . . . . . . . . . . . . . . . . 49 Late Ov erpressuri zation Fail ure Compari son. . . . . . . . . . . . . . . . . .. . . . . 51 4.2 4.3 Comparison of Releases for Failure to Isolate Containment 53 a nd t he By- Pa s s Se que nc e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
-y -
4.4 Comparison o AB- c and TMLB'- c (BMI-2104) to llY and 37. . . . . . . . . . 55 57 4.5 Comparison of 1907 (sum) to V-sequence (Surry) . . . . . . . . . . . . . . . . . . . . SS 4.6 BNL-Suggested Source Term........................................ f ' BNL-Suggested Release Characteristics for Seabrook................ 59 4.7 ' f G l D e.--,_. , _ , , , - ,,r .__.m, - , _ . %-._77--, m. ----, - - - - - _ - _ - , ---...-.-.,y
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-vii- !
LIST OF FIGURES Title Page Figure 3.1 A schematic represent; tion of source term calculation............ 9
~
Equipment hatch with personnel air 1cck........................... 12 3.2 3.3 Pe r s o n n el a i rl o c k . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 Typi cal hi gh energy pi pi ng penetrati on . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 3.4 Typi cal moderate energy piping penetration. . . . . . . .. . . . . . . . . . . . . . . 16 3.5 19 3.6 Typi cal el ectri ;al penetration. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 3.7 Typi c al ventil ati on pene tration. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25 3.8 A pictorial representation of laakage categories................. Estimated radial displacement of containment wa11................ 26 3.9 . fractions.......................... 32 3.10 Estimated containment failure 34 3.11 Definitions of the plant damace states used in 5PSS............... l l ~ 9 e t--- ,,,.,---r r---, y _ r , , - . . - , , --- - - y----.- . -, - - - - . - - - ~ - - , - - - - - . , ,
- 1. INTRODUCTION 1.1 Back cround Probabilistic Risk Asessment (PRA) studies have been undertaken by a num-ber of utilities (as exemplified by Refs.1-4) and submitted to the Nuclear Regulatory Commission (NRC) for review. Brookhaven National Laboratory (BNL) under contract to the NRC, has been involved in reviewing core malt phenome-nology, containment response and site consequence aspects of the PRAs.
This report presents a review and evaluation of the containment failure modes and the radiological release characteristics of the Seabrook Station Probabilistic Safety Assessment (SSPSA), which was completed by Pickard, Lowe and Garrick, Inc. (PLG) for the Public Service Company of New Hampshire and Yankee Atomic Electric Company in December 1983.5 1.2 Objective and Scope The objective of this report is to provide a perspective on severe acci-dent propagation, containment response and failure modes together with radiol-ogical source term characteristics for the Seabrook Station. Accident initia-tion and propagation into core damage and meltdown sequen:es were reviewed by - the Lawrence Livermore National Laboratory (LLNL) as reported in an incomplete report [6] prepared for the Reliability and Risk Assessment Branch of NRC.
"In the pres
- report , principal contai nment design features are dis-cussed and compared w'th hose of Zion, Indian Point and Millstone-3 designs.
Those portions of th ' PSS Telated to severe accident phenomena, containment source terms are described and evaluated. Numerical 1 , respcnse ar.d radio adjustments to the SPS ' estimates are documented and justified. )( 1.3 Organization of tke Report i At brief review of the Seabrook plant features important to severe acci-dent analysis is pre $ented in Chapter 2 along with comparisons to Zion, Indian Point and Hillstone '3 plant designs. Chapter 3 contains the assessment of contai vent perfo rmance. Specifically, definition of contairrnent response classes and plant damage states, analytical eho containment failure model, contai nment event $ree and accident phenomeno o y and the containment matrix are reviewed. Cha source terms together with justifications for/pter adjustment 4 addresses the accideThe where necessary. t results of this review are summarized in Chapter S. , f pWA l - m s6 . em l B i
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- 2. PLANT DESIGN AND FEATURES IMPORTANT TO SIVERE ACCIDENT ANALYSIS In this section, those piant design features that may be important to an assessment of degraded and core melt scenarios and contairnent analysis are reviewed. These important features are then ccmpared with the Zion, Indian Point and Millstone-3 facilities to identify comonalities for benchmark comparisons.
2.1 Assessment of Plant Design The Seabrook Station is comprised of two nuclear units each having an identical Nuclear Steam Supply System (NSSS) and turbine generator. The units are arranged using a " sling-along" concept which results in Unit 2 being arranged similar to Unit 1 but moved some 500 feet west. Each unit is a 1150 MWe (3650 MWt), 4-loop, Westir; house PWR plant. The turbine-generators are supplied by the General Electric Company and the balance of the plant is designed by United Engineers and Constructors. Each containment completely encloses an NSSS, and is a seismic Category I reinforced concrete structure in the fonn cf a right vertical cylinder with a hemispherical top dome and flat foundation mat built on bedrock. The inside face is lined with a welded carbon steel plate, providing a high degree of leak tightness. A protective 4 ft. thick concrete mat, which forms the floor of the contai nment , protects the liner over the foundation mat. The containment structure provides biological shielding for normal and accident conditions. The approximate dimensions of the containment are: Inside diameter 140 ft. Inside height 219 ft. - Vertical wall thickness 4 ft. 6 in, and 4 ft. 71/2 in. Dome thickness 3 ft. 6 1/8 in. Foundation mat thickness 10 ft. Contairnent penetrations are provided in the lower portion of the structure, ' and consist of a personnel lock and an equipment hatch / personnel lock, a fuel transfer tube, electrical, instrumentation, and ventilation penetrations. Each containment enclosure (also known as secondary contairnent) sur-rounds a containment and is designed in a similar configuration as a vertical right cylindrical seismic Category 1, reinforced concrete structure with dome and ring base. The approximate dimensions of the structure are: inside diam-eter,158 ft; vertical wall thickness, varies f rom i ft, 3 lr.. to 3 ft; and dome thickness,1 f t, 3 in. The containment enclosure is designed to collec leakage from the O contai nment structure other than leakage associated with piping, electrical ( and access passage penetration and discharge to the filtration system of e i contai nment. To accomplish this, the space betweerr the contairnent enclosure and the containment structure, as well as the penetration and safeguards " pump. areas, are maintained at a negative pressure following a design basis accident by fans which take suction from the containment enclosure and exhaust to ( atmosphere through charcoal filters. To ensure air tightness for the negative. pressure, leakage through all joints and penetrations has been minimized.
.a.
A containment spray system is utilized for post accident containnent hea* renoval. The containment spray system is designed to spray water containing boron and sodium hydroxide into the containment atmosphere after,a major acci-dent to cool it and remove iodine. The pumps initially take suction from the refueling water storage tank and deliver water to the containment atmosphere through the spray headers located in the containment dome. Af ter a prescribed amou,t of water is removed from the tank, the pump suction is transferred to the containment sump, and cooling is continued by recirculating sump water through the spray heat exchangers and back through the spray headers. The spray is actuated by a containment spray actuation signal which is generated at a designated containment pressure. The system is completely re-dundant and is designed to withstand any single failure. The con:ainment isolation system establishes and/or maintains isolation of the containment .from the outside environment in order to prevent the re-lease of fission products. Automatic trip isolation signals actuate the ap-propriate valves to a closed position whenever automatic safety injection oc-curs or high containment pressure is experienced. Low capacity thermal elet-tric hydrogen recombiners are provided. The emergency core cooling system (ECCS) injects borated water into the reactor coolant system following accidents to limit core damage, metal-water reactions and fission product release, and to assure adaquate shutdown mar-gin. The ECCS also provides continuous long-term post-ac ident cooling of the core by recirculating borated water between the containment sump and the reac-tor core. The ECCS consists of two centrifugal charging pumps, two high pressure safety injection pumps, two residual heat removal pumps and heat exchangers, and four safety injection accumulators. The system is completely redundant, and will assure flow to the core in the event of any single failure. . The control building contains the building services necessary for contin-uous occupancy of the control room complex by operating personnel during all operating conditions. These building services include: HVAC services, air purification and iodine removal, fresh air intakes, fire protection, emergen:y breathing apparatus, communications and meteorological equipment, lighting, and housekeeping facilities. Engineered Safety Feature (ESF) filter systems required to perform a safety-related function following a design basis accident are discussed below:
- a. The containnent enclosure exhaust filter system for each unit col-lects, filters and discharges any containment leakage. The system is not normally in operation, but in the event of an accident, it' is placed in operation and keeps the containment enclosure and the building volumes associated with the penetration tunnel and the ,ESF equipment cubicles under negative pressure to ensure all leakage f rom, the containment structure is collected and filtered before discharge to the plant vent.
9 4 Y
-4
- b. One of two redundant charcoal filter exhaust trains is placed in operation in the fuel
- storage building whenever irradiated fuel not in a cask is being handled. These filter units together with dampers and controls will maintain the building at a negative pressure.
The emergency feedwater system supplies demineralized water from the con-densate' water storage tank to the four steam generators upon loss of nomal feedwater flow to remove heat from the reactor coolant system. Operation of the system will continue until the reactor coolant system pressure is reduced The removal system can be operated. tc a value at which the residual heat combination of one turbine-driven and one motor-driven emergency feedwater pump provides a diversity of power soJrces to assure delivery of condensate under emergency conditions. The two units of the facflity are interconnected to foroff-sitethe Newpower England via three 345 kilovolt' lines of the transmission system states. The nomel preferred source of p>wer for each unit is its own main turbine generator. T.1e redundant safety feature buses of each unit are power- - A highly reliable generator breaker is ed by two unit auxiliary transfomers. provided to isolate the generator from the unit auxiliary transfomers in the everet of a generator trip, thereby obviating the need for a bus transfer upon loss of turbine generator power. In the event that the unit auxiliary trans-formers are not available, the redundant safety feature buses of each unit are Upon loss of off-site power, powered each unit by two reserve auxiliary transfomers.is supplied with adequate power diesel-engine generators. Either diesel-engina generator ar.d its associated safety feature bus is capable of providing acequate power for a safe shutdown A con-under accident conditions with a concurrent loss of of f-site. power. i stant supply of power to vital instruments and controls of each unit is assur-ed through the redundant 125 volt direct current buses and their associated battery banks, battery chargers and inverters. 2.2 Comoarison with Other Plants Table 2.1 sets forth the design characteristics of the Zion, Indian Point-2, and Millstone-3 f acilities as they compare to the Seabrook station. l It is seen that the containment characteristics are quite similar with the exception of containment operating pressure for Millstone-3 (subatmospher-i ic design), and the use of fan coolers in Zion and Indian Point for post-acci-dent containment cooling, the lower reactor cavity configuration, and chemical composition of the concrete mix. The primary system designs are nearly iden-l tical between the four units. l The Seabrook containment building basemat andAsthe internalisconcrete concrete heated, structures are composed of basaltic-based concrete.The initjal gas consists largely of i I water vapor and other gases are released. carbon dioxide, the quantity of which depends on tr.e amount of calcium caibon-Limestone concrete can contain up to 80% cal _ciua. ate in the concrete mix. carbonate by weight, which could yield up to 53 lb of carton dioxide per cubic foot of concrete. However, basaltic-based concrete contains very little cal-cium carbonate (3.43 ws for Seabrook) Thus, pressurization and would of the not release a containment assubstantial a amount of carbon dioxide.5
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Table 2.1 Comparison of Selected Design Characteristics Zion Indian Point Hillstone Seabrook Unit 2 3 Unit 3 .7 Unit 1,2 5 Design Parameters Unit 11 3,250 3,030 3,411 3,650 Reactor Powe'r (HW(t)] Containment Building: /g 2.73 x 10 6 2.61 x 10 8 2.3 x 10 8 2,.7/x 10 6 Free Volume (ft 3) 59.7 67.7 Desigr Pressure (psia) 62 62 (psia) 15 14.7 12.7/9.1 15.2 Initial Pressure 120 120/80 120 Initial Temperature (*F) 120 Pr d ,ary System: 11,671 13,140 Water Volume (ft 33) 12,710 11.347 1 017 , 7 - Steam Volume (ft ) 720 216,600 720 216,600
?
222,739 222,739 Mass of UO2 in Core (it.) 20,407 ? 19,000 Mass of Steel in Core (ib) 21,000 44,500 44,600 45,296 45,234 Mass of Zr in Core (1b) 87,000 87,000 (ib) 87,000 78,130 , Mass of Bottom Head 14.4 14.4 (f ) 14.7 . Bottom Heid Diameter 14.4 0.45 0.44 0.45 0.45 Bottom Head Thickness (ft) Containment Building Coolers: yes yes yes Sprays yes yes no no Fans (with safety function) yes Accumulator Tanks: 173,000 348,000 213,000 Total Mass of Water (1b) - 200,000 600 615 Initial Pressure (psia) 665 665
-) , ' - "
150 80 Temperature __ (*F) , 150 Refuelina Water Storace Tank: 2.89 x 10 8 2.89 x 10 6 10 7 2.89 x 105
- Total Mass of Water (1 b.) .
50 86 Temperature ('F) 100 120 Reactor Cavity: Wet Dry Dry / Wet Configuration Het Limestone Basaltic Basaltic Basaltic Concrete Material .
- Minimum (Maximum Capacity = 3.9 x 106 lb)
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.6-result of corium/ concrete inter 3Cti ns would be expected to take a very long tirae.
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- 3. ASSESSMENT OF CONTAINMENT PERFORMANCE In this chapter, the review of containment responce to severe accidents is described. Analytical techniques used to analyze core meltdown phenomena and containment response are reviewed, containment failure model is assessed and plant damage states and containment failure modes are evaluated.
Parallels between this study and other PRAs are cet forth. Finally, the rel-evance and validity of the conclusions is addressed. 3.1 Containnent Analysis Methods A brief description of the computer codes used to perform the transient degraded, core meltdown and containment response analyses is provided in this section. The MARCHE computer code is used to model the core and primary system ' transient behavior and to obtain mass and energy releases from the primary system until reactor vessel failure. These mass and energy releases are then used as input to tne other computer codes for analysis of containment re-sponse. For sequences in which the reactor coolant system remains at an elevated5 pressure until the vessel failure (" time-phased dispersal"), the MODMESH computer code is used. This code calculates the steam and hydrogen blowdown from the reactor vessel using an isothermal ideal gas model. The water level boil-off from the reactor cavity 'loor is modeled using a saturated critical heat flux correlation. AdditionUly, the accumulator discharge following de- . pressurization caused by the vessel failure is also considered. ' A modified version of the CORCON 8 code is used to replace the INTERe sub-routine of the MARCH code. CORCON models the core-concrete interaction after the occurrence of dryout in the reactor cavity. The mass and energy releases from the core-concrete interaction are transferred to the M00 MESH code for proper sequencing and integration into the overall mass and energy input to 5 COC0 CLASS 9 code. COC0 CLASS 9, a modified version of the Westinghouse COC0 computer code utilizes the mass and energy inputs to the containment as computed by MARCH to model the containment building pressurization and hydrogen combustion phenom- - ena. This code replaces the MACE subroutine of the MARCH code. The code also models heat transfer to the containment structures and capability for contain-ment heat removal through containment sprays and sump recirculation. Fission product transport and consequence calculations 5 are perfo rmed using the CORRAL-Il and the PLG proprietary CRACIT computer codes, respec-tively. The analytical methods used to carry out the core and containnent thernal hydraulics, and fission product transport calculations are identical to those used for MPSS-3.7 4 1
'.1 1
3.2 Containment Failure 3.2.1 Eackground In order to assess the risk of the Seabrook-1 plant, radiological source terms have to be calculated. Many steps are involved in such calculations.
'These are schematically shown in Fig. 3.1. The mode and time of containment failure directly impact on the radioactivity release categories. These, when coupled with the status of reactor cavity and the spray system, determine the source terms. This section deals with the mode and time of containment fall-ure.
l 3.2.2 Design Description The primary containment of the Seabrook plant is a seismic Category I re-inforccd concrete dry structure. It consists of an upright cylinder topped with a hemispherical dome. The inside diameter of the cylinder is 140 feet and the it. side height from the top of the basemat to the apex of the dome is ' approximately 219 feet. The cylindrical vall is 4'6" thick above elevation 5' and 4'7-1/2 " thick below that evaluation. The dome is 3'6-1/8" thick and 69'11-7/8" in radius. The cylinder is thickened to provide room for addition-al reinforcing steel around the openings for' the equipment hatch and the per . sonnel airlock. The net free volume of the containment is approximately 2.7 x 106 ft 3. ya The liner The[inside of the containment isjwelded y1%kSer steel liner. plate it the cylinder is 3/8" thick in all areas except penetrationThis andliner the ' iunction of the basemat and cylinder where it is 3/4" thick. i serves as a leak-tight membrane. Welds that are embedded in the concrete and not readily accessible are covered by a leak chase system which permits leak testing of these welds throughout the life of the plant. The dome liner is 1/2" thick and flush with the outside face of the cylindrical liner. The operating and the design parameters of containment are notec in Table 3.1. The containment building is surrounded by an enclosure. The contair. ment enclosure is a reinforced concrete cylindrical structure with a hemispherical dome. The inside diameter of the cylinder is 158 feet. The vertical wall varies in thickness f rom 36 inches to 15 inches; the dome is 15 inches thick.
- The inside of the dome is 5'5" above the top of the containment dome. Located at the outside of the enclosure building is the plant vent stack, consisting of a light steel frame with steel plates varying in cross-section. The stack carries exhaust air f rom various buildings.
The containment enclosure is designed to control any leakage from the containment structure. To accomplish this, the space between the containment and the enclosure building (approximately 4'6" wide) is maintained at a slight . negative pressure (-0.25" water gauge) during accident conditions by f ans which take suction f rom the containment enclosure and exhaust to atmosphere / through charcoal filters. i There are a number of containment penetrations which are steel components ' that resist pressure. These penetrations are not backed by structural con-crete and include.the following: I e i
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9 l CONTAINMENT TIME OF FAILURE FAILURE MODE WET OR ORY RELEASE SPRAY l REACTOR CAVITY CATEGORY SYSTEM SOURCE TERM Figure 3.1 A schematic representation of source tem calculation. r . .
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Table 3.1 Containment Operating and Design Parameters Parameter Value Normal-Operation Pressure , psig 0.5 Inside Temperature F 120 Outside Temperature F 90 Relative Humidity , % 45 Service Water Temperature , F 80 Refueling Water Temperature F B6 Spray Water Temperature F 88 Containment Enclosure Pressure , inches w.g. -0.25 Desian Conditions Pressure , psig 5'2.0 Temperature , F 296 I Free Volume , ft3 2.7x106 Leak Rate , % mass / day 0.[ Containment Enclosure Pressure , psig -3.5 l
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9
- 1. Equipment hatch,
- 2. Personnel air lock,
- 3. Piping penetrations,
- 4 Electrical penetrations,
- 5. Fuel transfer tube assembly,
- 6. Instrumentation penetrations, and
- 7. Ventilation penetrations.
These components penetrate the containment and containment enclosure shells to provide access, ar.chor piping, or furnish some other opera.ional requirement. All penetrations are anchored to sleeves (or to barrels) which are embedded in the concrete containment wall. Equionent Hatch The equipment . hatch (Fig. 3.2) consists of the barrel, the spherical dished cover plate with flange, and the air lock mounting sleeve. The center-line of the hatch is located at elevation 37'1/2" and an azimuth of 150*. The hatch opening has an inside diameter of 27'5". A sleeve for a personnel air lock, the inside diameter of which is 9'10", is provided at centerline eleva-tion 30'6". Thicknesses of the primary components are as follows: Component Thickness (inches) B rrel 3 1/2 Spherical 1 3/8 Flange 5 3/8 Air lock mounting 1 1/2 sleeve . The equipment hatch cover is fitted with two seals that enclose a space which can be pressurized to 52.0 psig. The flange of the cover plate is at-tached to the hatch barrel with 32 swing bolts,1-3/8 inch in diameter. The barrel, which is also the sleeve for the equipment hatch, is embedded in the
. shell of the concrete containment. The equipment hatch cover can be lifted to clear the opening.
I Inserted into lhetpounting sleeve through the equipment hatch cover is a l
- personnel air lock consisting of two air lock doors, two air lock bulkheads, and the air lock barrel. Signi ficant dimensions of the air lock are as follows: ,
* ; Pa rameter Dimension Inside' Oiameter of Barrel 9'6"
- 1/2" Barrel Thickness Door Opening 6'8" x 3'6" Door Thickness ,3/4" .
Bulkhead Thickness 1-1/8" Each door is locked by a set of six latch pin assemblies, and is designed to withstand the design pressure from inside the containment. To resist the tes.t pressure, each door is fitted with a set of cast clamps. The doors are hinged and both swing into the containnent. Each door is fitted with two seals that e
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are located such that the area between doors can be pressurized to 52.0 psig. The doors are mechanistically interlocked 50 that only one door can be opened at a time. The capability exists for bypassing this interlock to equalize the pressure by use of special tools. The doors may be operated mechanically. Pers-a.nel Ai r Lock The personnel air lock (Fig. 3.3) consists of the air lock doors (2) and the lock barrel. The barrel, which is also the sleeve for the personnel air lock, is imbedded in the shell of the concrete containnent. The centerline of the oarrel is located at elevation 29'6" and an azimuth of 315*. Significant dimensions are as follows: Parameter Dimensions Clear Opening 7'0" 0.D. of Flange on Door 7' 9 1/8" Barrel Thickness 5/8" Cover Thickness 5/E ' The air lock barrel has a door on each end, each of which is designed to withstand the design pressure from inside the containment. The doors are hinged and swing away from the air lock barrel. Each door is fitted with two l seals that are located such that the area between doors can be pressurized to 52.0 psig. The locking device for the doors is a rotating, third ring, breach-type mechanism. These doors are also mechanically interlocked so that only one door can be opened at a time. The capability exists for bypassing this interlock and relieving the internal pressure by use of special tools. The doors may be operated mechanically. . Pioing Penetrations There are two types of piping penetrations: moderate energy and high energy. Moderate energy piping penetrations are used for process pipes in which both the pressure is less than or equal to 275 psi, and the temperature of the process fluid is less than or equal to 200'F. High energy piping pene-trations are used for that piping in which the pressure or temperature exceeds these values. High energy piping penetrations (Fig. 3.4) consist of a section of pro-cess pipe with an integ rally-f orged fluid head, a containment penetration sleeve and, where a pipe whip restraint is not provided, a penetration sliding support inside the containment. The sliding support provides shear restraint while permitting relative motion between the pipe and the support. The annu-lar space between the process pipe and the sleeve is completely filled with fiberglass thermal insulation. The pipe and the fluid head, are classified as ASME III Safety Class 2 (NC), whereas the sleeve is classified as part of the concrete containment, ASME III (CC). The sliding Support inside the contain- * ( ment is classified as an ASME Safety Class 2 component support (NF). Moderate energy piping penetrations (Fig. 3.5) consist of one or more process pipes, the containment penetration sleeve, and a flat circular end . plate. The pipe is classified as ASME 111 Safety Class 2 (NC). The sleeve is classified as ASME !!! Div. 2 (CC). The end-plate is classified as Class MC. l
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4 Table 3.2 gives a list of the containment piping penetrations. Included in this table is the penetration size. All of these piping penetrations are j in the lower portion of the structure. Electrical and instrumentation Penetrations _ Electrical pen:trations (Fig. 3.6) consist of a stainless steel header plate with an attach'ed terminal box, electric 1 modules which are The metallic pressure resisting parts, the sleeve, plate and to the sleeve. stainless steel header plate and carbon steel weld ring we III Safety Class MC components (NE); that portion of the sleeve which is backed by concrete was designed as part of the concrete containment, ASME 111 (CC). for Double silicone and Hypalon 0-rings provide a seal with The header a cavity leakage monitoring between the header plate and the modules. plate is provided with a hole on the outside of the containment to allow for pressurization of the penetration assembly for leakage monitoring. There are a total of 64 electrical penetrations out of which 14 are spare All of these electrical penetrations are below the grade. and 8 are unused. Instrumentation penetrations are of two types -- electrical and fluid. The electrical type is similar in construction to the other electrical pene-trations. The fluid penetrations are similar in construction to the moderate energy piping penetrations. Fuel Transfer Tube Assembly , The fuel transfer tube asse-51y consists of the fuel transfer tube, the e on the reactor side, and the sliding sad-penetration sleeve, the fixed sat die in the fuel storage buildin,. The fuel transfer tube centerline Theisfuel at elevation (-)9'4-1/4" and it has approximately 20" inner diameter. transfer tube wall penetration sleeve, which is embedded in the concrete, has an inside diameter of a5out 25". , j Ventilation Penetrations There are two types of ventilation penetrations -- the containment air purge penetrations (HVAC-1 and HVAC-2) and the containm consist of a pipe sleeve (a rolled and welded pipe section, 36" outer diameter by 1/2" wall thickness) which is flanged at each end with 36" weld neck flanges and, attached to these flanges, the inner and outer isolation valves. Together with the pipe, these valves form a part of the containment pressure boundary. The valves are 36" diameter butterfly valves with fail-safe pneu-liner is matic operators. The weld between the pipe and the
- containment i
! equipped with a leak chase for pressure testing.
The containment on-line purge penetrations each consist of a pipeAsleeve short l (a rolled and welded pipe section, 8" o.d. by 1/2" The wall thickne containment, and a 3/4" valve and test connection is attached to it. I O
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Table 3.2 Containment Liner Penetrations a Penetration Penetration Numbers Service Size X-1 to X-4 Main steam line 30" X-5 to X-8 Main feedwater 18" X-9, X-10 RHR pump suction 12" X-11 to X-13 RHR to safety injtction 8" X-14 to X-15 Containment building spray 8" X-16, X-18 Containment on-l'r.e purge 8" X-17 Hydrogenated vent header 2" X-20 to X-23 CCW supply and return 12" X-24 to X-27 Safety injec-icn 4" X-28 to X-31 CVCS to pump seal injection 2" X-32, X-34 Drain line 3",2" X-33, X-37 CVCS 3" X-35, X-36, X-40 RCS test / sample control 1" or smaller
- X-52, X-71, X-72 X-38 Combustible gas control 10" X-39 Spent fuel pool cooling 2" - ,
X-43, X-47, X-50 Instrumentation lines I) -c ( X-57 X-60, X-61 From containment recirculation sump 16" X-62 Fuel transfer tube 20" X-63 to X-66 Steam generator blowdown 3" , X-67 Service air 2" HVAC-1,2 Containment purge supply / exhaust lines 36" X-19, X-41, X-42 , _ - X-44 to X-46, X-48 Spare '?/ X-49, X-51, X-58 X-59, X-68 to X-70 O w
l i I j OUTBOARD CONTAINMENT WALL _ INBOARD i J STAINLESS STEEL LINER PLATE HEADERPLATE] THERMAL INSULATING Y I BOX MOUNTING RING l GASKET
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i 21 ends of this resulting assembly are welded to 8" weld neck flanges which are through-bolted to the inner and outer isolation valves. These valves are 8" dianeter butterfly valves having fail-safe pneumatic operators. The weld be-tween the pipe sleeve and the contaiment liner is equipped with a leak chase for pressure testing. These on-line purge penetrations are very similar to those for 36" lines shown earlier. 3.2.3 Leakaoe Rate Calculation Under severe accident conditions the pressure inside the containment quickly builds up in the range of 75 to 200 psi. At these pressures, any leakage through the contaiment holes will essentially be choked. The }[ leakage under choked flow condition is given as (Ref.10): k+1 W= k(k1) A III ' where W = discharge rate (kg/s), A = leak area (m2 ), P = absolute pressure (N/m2 ), p = mixture density (kg/m3 ), and k = ratio of specific heat at constant pressure to that at constant volume. For air and water vapor mixture, k - 1.3. If the mixture density is expressed by perfect gas law P (2) p=g where R = gas constant, and T = the absolute temperature, Then Eq.(1) becomes k+1 I W=jk(k1) A (3) The mass of mixture can be written as H = Vp or, (4) M=h . Equations (3) and (4) where V is the free mixture volume in the contaiment. can be combined to get the leakage rate, in terms of mass fraction, as I
22-l )k+1 vTA (5)
=kk(k ,
Note that the leakage rate, when expressed in tems' of mass fraction, depends only on the leakage area. For Seabrook-1, using V = 2.704x10 6 ft3 and T = 296 F. Eq.(5) gives Leakage Rate = 0.721 Ai n W/o Per hour (6) 2 where iA n is the leakage area in in . Alternately, Leakage Rate = 17.3 Ai n W/0 Per day. (7) The essentially intact e contaiment leakage of 0.2 w/o per day, thus, corresponds to an eq ent leakage area of 0.012 in2 (or, an equiva-lent hole o' 1/8-in diameter). A leakage area of 4 to 10 in would 2 correspond to the leakage rate of 2.9 to 7.2 w/o per hour. In other words, it will take2 about 14 hours to leak the entire cont?nt to the envirornent through a 10-in hole. 3.2.4 Containment Failure Model 3.2.4.1 Leak-Before-Failure During accident sequences involving core damage, the containmer.t struc-ture will be exposed to pressures and temperatures beyond those used in the design basis accident (DB A) . Response of the contalment building to these severe conditions is evaluated in SSPSA by employing, for the first time, a leak-before-failure model. In this model allowance is v.ade for continuous leakage from the contaiment to the surroundings. This mode of contaiment ' failure is temed local failure. The containment leakage can occur at many locations and discontinuities such as mechanical and electrical penetrations, personnel lock, equipment hatch, fuel transfer tube, welds, and in between the liner and concrete. Depending upon the size of leakage area and the accident sequence, local failures may gradually relieve pressure thereby gross con-taiment failure may be averted. ! The leak-before-failure model is a realistic one. The extent of leakage l and the health consequences must, however, be carefully studied. In order to explain this issue, it is observed that traditionally probabilistic risk as-sessment is made by using what is temed a threshold model. In the threshold model, the containment is considered intact until the internal loading equals
- or exceeds a pressure threshold (which may also be temperature dependent), at which it is deemed to have suffered a failure (gross). If the internal lo.ad-ing is below this threshold value, the contaiment is considered intact and hence the risk is quite low. In the leak-before-falture model, the release of activity, which is considerably small compared with that for the gross fail,are mode, must be considered in health consequences. However, such leakages can i
potentially prevent the internal pressure from approaching the threshold value l and thus a catastrophic or gross failure may be avoided. l l
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i l 3.2.4.2 Classification of Failure The SSPSA report has clas,1fied containment failures in three categories: l
. Conta'nment Failure Category A. Includes containment failures that develop a small leak that iTsubstantially larger than the leak ac-ceptable from an intact containment, but not large enough to arrest
.the pressure' rise in the containment. Category A failures thus cause an early increase in the rate of leakage of radionuclides over the de-sign basis leak rate but pressurization of the contairrtent continues until either a category B or C contaiment failure cccurs.
The intact containment is defined as the one in which leakage is lim- g Ic ited by tie Technical Specification value. For Seabrook-1, this value 4 t is 0.2 w/o per day at the calculated peak accident pressure of approx- / imately 47 psig. Note that the S " "mly has used 0.1 volume perf- g cent per day for this leakage, Althoug w we most recent
... y prior. cited both 0.1 volume per-amendment dated August 1984, the cent and 0.1 w/o per day. The 10CFR50, Appendix J mandates the allow-able leakage to be quoted as w/o per day. The higher value noted here is based on Amendment 53, August 1984.*
. Containment Failure Category B. Includes failure modes that develop a large enougn leak area so tnat the pressure in the containment no longer increases. The time during which a substantial fraction of the radionuclide source term is released is longer than approximately 1 to 2 hours. Category B f ailures include self-regulating failure modes where the leak area is initially small but increases with pressure so that it becomes sufficient to terminate the pressure rise before a category C contaiment failure occurs.
The definition of " substantial" fraction is unclear.
. Containment Failure Category C. Includes those containment failure mcdes tnat dev.elop a large leak area. A large fraction of the total radionuclide source tem is released over a period of .lesse 1/
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hour. All gros,s failure modes are included in category C. Mathematica51y, these three failure categories can be expressed in tems l of leakage areas as follows: Type A g/ ADBA %AA ANP AB'< Ap Type B (8) j e Type C
#('_ g/e AC where. .
y[~ jy[" ADBA = leakage area corresponding to th,e technical specification limit for containment leakage, ineered Safety 7
*Inere appears to be substantial update / changes in t Features flow diagram, including arringements of moto oper ed valves and I bypass lines, which may substantially change the,ffequency o g events. BN!. . .
however/is not reiiewi,ng this part of SSP,SA.,/ e
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3,1 I.l Anp = leakage area not large enough to arrest
, j\, \ gj pressurizat, ion, and f\ h Ap = leakage area sufficient to release 100 w/o
- h. \ in one hour.
A( s The leakage area required to release a subst al fraction of the radio-OA nuclide source term in approximately an hour an > can be computed using Eq. 9 Assuming one-hundred percent turnover as a substantial fraction in one (6). 2 or about pp. hour i q. (6) E gives theany Therefore, required containment leakageleak area area to be in equal excesstoof138 1 ftinwill be 2 - yf k.;\ 1 ft . fined as a gross _.contaiment failure .(Catecorvi)- This estimate of the ead' /
#\ area is factor two too high from t'he value stated in uri The leakage area required to arrest containment pressurization is in the range of 4 to 10 square inches, the lower value being more representative of A leak wet sequences and the upper value is representative of dry sequences.
area of about 6 square inches will result in the release of about 100 w/o of Tne upper bound leak area for Type A failure ac-ivity is taken as 4 in . in a dag (see Eq. 7).This corresponds to release of the radioactive source tem (1007, turnover ir about 36 hours. The Category B leak area is, thus, in the range of 4 in} to 1 ft , Figure 3.8 is a pictorial representation of these, 2 leakage categcries. 3.2.5 Containment Pressure Capacity , 3.2.5.1 Concrete Containment T1e Seabrook PSA has examined failure modes for the containment structure itself, the steel liner, all penetrations, equipment and personnel lock hatch-i es, and the secondary ontainment. The containment structure includes the cylindrical wall, the hemispherical dome, the base slab and the base slab and containment wall junction. The most critical membrane tension was found to occur in the cylinder in the hoop direction. The median pressure which causes yield of both the liner steel and the reinforcing bars was found to be approx-imately 157 psi, with a coefficient of variation of 0.084. The ultimate hoop load in .ylinder is 216 psig. The contaiment wall is, tnus, assumed to fail at this pressure. At pressures beytad this, very large irreversible defoma-tions occur which will cause cracks in the reinforced concrete but the lossThe of integrity of the pressure boundary may not occur until the liner tears. compiled radial defonnations of the containment wall are shown in Figure 3.9. Note that the radial strain at the expected failure pressure of 216 pst is 4.77. ( Ar/ r) . The hemispherical dome was calculated to yield at a slightly higher pres-sure (163 psig). The failure pressure is predicted at 223 psig. The median pressure for flexural f ailure of the base slab is 400 psig, with a logarithmic standard deviation of 0.25. However, the shear mode of failure is more restrictive. For this mode, the median failure pressure is estimated in SSpSA as 323 psig, with a logarithmic standard deviation of 0.23. Although the uncertainty for failure of the base slab is large, the probability of failure is small because the median capacities are high. Thus', f ailure of the base slab is not considered to be a critical failure mode and
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127-an estimation of leak areas was, therefore, not considered for this mode of failure. Secondary stresses in the cylindrical portion of the contalment occur at discontinuity such as at the base slab containment wall junction, at the springline, and where the amount of reinforcing changes. The flexural yield at the base of the cylinder occurs at 175 psi. At higher pressures, a plastic hinge foms with considerable cracking of the concrete. These cracks, how-ever, 'are small enough so as not to threaten the integrity of the liner. The loss of integrity of the liner is not expected until a median pressure of 408 psi is reached. Thus, the failure of the base slab and containment wall junc-tion is not limiting. In sumary, the containment wall is expected to undergo significant de-fomation (=4.7% Ar/r) prior to its f ailure at 216 psig. At this pressure, Type C-(i.e., gross) f ailure occurs.
- 3. 2. 5.2 Liral The elongation capacity of the steel liner is computed by neglecting the f riction forces between the liner and the concrete. The possibility that the liner stresses and strains could be different between two different pairs of tees was, however, considered. The SSPSA computed an elongation of 8.1 per .
cent under uniaxial conditions, or an elongation of 4.7 percent under plane strain conditions can be achieved without fracture. This would ensure integ-Additionally, the rity of the liner until fracture of the reinforcing bars. leakage of the containment at penetrations is considered likely before hoop failure of the liner occurs. 3.2.5.3 Penetrations At all major penetrations, the contaiment wall is thickened and addi-tional reinforcement is provided to resist stress concentrations. None Instead, of the meridional or hoop reinforcing bars are teminated at penetrations. they are continued around the penetrations, thus ensuring that excess hoop and meridional capacity is available. Table 3.2 lists all piping penetrations. As the contaiment pressure increases beyond its yield val (157 psi), large radial deformations begin to occur. Tnis induces stresses in the pipes by relative displacements between the contaiment wall and the pipe whip re-straints. Therefore, the most critical penetrations are the areas where the pipe is supported close to the penetration. Also, stronger and stiffer pipes develop higher forces at the penetrations for a given relative displacement. The SSPSA study selected the following penetrations for investigation as being among the lines most likely to fail: Penetration X-23 12" schedule 40 carbon steel (also X-20 to X-22 by ~ similarity) .
~
Penetration X-26 4" schedule 160, stainless steel (also X-24, X-25, X-27)
-___ .~,_.,._.,--..,-..-._y.-._, . _ , . . - , _ _ , _ _ . , - , _ _ , _ . _ . . , , , - _ - . - , . - , - . - - - , _ . . . - . - .
Penetration X-71 1" - multiple pipe penetration (also X-72 and possibly* others) 18" main feedwater schedule 100, Penetration X-8
- (also X-5 to X-7) carbon steel Fuel Transfer Tube Convoluted Rellows j
The probability of failure at these penetrations was computed by (a) l establishing a pressure-displacement relation, (b) estimating the failure probability as a function cf radial displacement and then (c) combinin two. The vertical displacement due to meridional strains is small 2 (Fig. 3.9). (less than 3 in.hes) and hence its impact on the penetrations was ignored. Since most of these. penetrations are in the lower part of the containment, the radial displacements experienced by them due to plastic deformation of contairvaent would also be small. . The multiple penetration (X-71 and X-72) would not fail even for the most - unfavorable forces which these pipes could sustain. For penetrations X-23 and i
' X-26, the most likely location for failure is at the partial penetration fil- -
let welds which join the pipe to the end plate. When failure ofThe this gap weld oc-between curs, the pipe remains in the hole provided in the end plate. the pipe and the end plate is likely to remain small unless the pipe wall buckles. Exact gap size is hard to compute.- The SSPSA appears to use a uni-form gap size of 0.04 in., and 0.10 in. as median and upper estimate tively. i X-26 (as well as X-24, X-25, and X 27) penetrations are shown in Table k 3.3.as___ 2 The dian f ailure ressure for X-23 penetration, t a n n s .e is higher than the hoop failure ressu e (216 psig) of These leak areas, therefore, are not expected to devel-he contairunent wall. Op. l Penetration X-26 is expected to fail at a median pressure of 166 psig.
- f' The combined leak area for all safety injection penetrations 2 is obtained by independently adding individual median leak area of 0.5 in ,
Penetrations X-71 and X-72 are not likely to contribute to the overall leak area, as stated earlier. The main feedwater lines (penetrations K-5 to X-8) are 18-in. diameter, Schedule 100 pipes. The failure mode of most concern isAtfailure a median of the flued pressure Mhe d due to axial loads in the pipe at the penetration.each one of these pene org/180 psig,2 area of 50 in each. Since all four2 of these can f ail independent of each Although the failure of a single s'uch other, the total leak area is 200 in . ~ penetration can be considered as Type 8 failure,15 all four main feedwater f penetrations were to fail simultaneously the resulting leakage will be of Type *
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C. The fuel transfer tube is fixed to an elevated floor inside the contain-ment. As the pressure in the containment increases, the containment wall . moves outwards and thereby exerts pressure on the bellows. The most pertinent e - = - - - ---e .--.,,,---,..--.,,v.- ----=---emir-,- -w-,---m. m- - . - , , . - - - - - - arwwe g-y- -+-r-e'y--,~,ww
c' . i Table 3.3 Leak Area Estimates for Mechanical Penetrations Median Median Line Penetration Leak Area Failure Pressure Size inz psig 6.0 >216 12" X-20 to X-23 CCW Supply and Return 2.0 166 4" X-24 to X-27 Safety Injection X-71 and X-72 Negligible -
< 1" Sample / Control 200 180 18" d$
X-5 to X-8 Main Feedwater Fuel Transfer Tube 3 172 X-16, X-18 See Text 8" ' On-line Purge HVAC-1,2 See Text 36" Containment Purge
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bellows f ron the viewpoint of containment leakage is the one inside the con-tainment (EP-2). Three potent,ial f ailure modes, in their order of decreasing
' probability of failure, considered are (a) f ailure due to overall buckling of the bellows, (b) failure due to local buckling within the convolute, and (c) f ailure due to meridional 2 bending strains. The SSPSA hasThis estimated median is a Type A leak area of about 3 in at a pressure of about 172 psig. ,
failure. There are two sets of containment penetrations which are cpen to the The on-line penetrations (X-16 and contat ment atmcsphere on the inside. X-18) are the 8-inch purge suction and discharge lines and contaiment purge suction and discharge lines (HVAC-1 and 2) are the 36-inch lines. Each one of these four lines has two containment isolation valves, oneoperated inside and one butter-outside the contaiment. All eight valves are pneumatically (usually ethylene fly val ses. At elevated temperatures, the seal material propylene) on thes: valves may deteriorate and lose its sealing function. Any deposition of radioactive aerosols could further deteriorate the sealant materi al . Consideri .g sealant degradation due to temperature alone, ethylene propylene seal life (Ref.10) is 5 hours, 40 mts, or 20 mts if exposed to 400, 500 or 600 F, respectively. b e p, a nafrow crack leak . In the event of the f ailure of the sealant to leak into the space path may develop :nd containment atmosphere may on valves are closed from between the two isolation valves. Since the isol the contaircent isolation signal system, the leakage of containmer,t atmosphere to the enviroment can occur only if theThe sealant of the outer time duration contaiment elapsed before iso-this lation butterfly valve also fails. happens can be significantly long (of the order of hours). The $$PSA has es-timated it to be long compared to the contalment failure by other causes. The SSPSA study, therefore, has disregarded this release path. The available leakage area due to sealant degradation has been estimated (Ref.10) by assuming an equivalent clearance of 1/16 inchThis betweengivesvalve disc a total and body for ' low' and 1/8 inch for 'high' estimates. ' leakage area of 17 in z as low value and 34 int as high value. As noted ear- f lier, the outer butterfly valves must also experience high temperaturesThe SSPSA prior study to a through release path. This leak area is of Category B. has argued that such a leak path is not likely to result prior to a gross containment failure (Category C). Electrical penetrations can f ail primarily due to overheating of the pot-ting compound. Tte SSPSA study has concluded that the failure of electrical penetrations is not expected to make a significant contribution to containment failure for any accident sequence. This conclusion, appears justified for the wet case, but, for the dry case, it is based on their estimate of slow over-A heating of the potting compound. Such a calculation, similar to the should be made to check this assessment. problen of vent / purge line butterfly isolation valve failure, is beyond/ the scope of this work and hence it was not done. f ai) either sili-The equipment hatch and personnel lock penetrations can(generally due to pressure loading or degradation of the sealant materialThe structura cor.e). ly. The sealant material can degrade at high temperatures typical of a
severe accident. According to the 0-Ring Handbook (see Ref.10), silicone can survive for twenty hours when exposed to 500 F temperature. Furthemore, the personnel air lock is a double door system so even if the sealant around one door were to become ineffective, substantial time delay would be required to make the second sealant also ineffective. It, thus, appears that the equip-ment hatch and personnel lock penetrations do not contribute significantly to Type B failure. 3.2.5.4 Containment Failure Probability
- The calculation of the probability of containment failure as a function of the pressure is quite involved. The method used and results reported in t5e SSPSA study seem reasonable except for the impact of all four main feed-water lines failure. The SSPSA has categorized the f ailure of X-8 (one of the four main feedwater lines) penetration as Type S since anticipated leak area is 50 in 2. It appears to us that when one such penetration fails, the remain-ing three will also fall at nearly the2 same pressure of 180 psig (195 psia).
Any depressurization due to a 50-in hole is not likely to be fast enough to reduce the containment pressure substantially prior to the failure of the three remaining penetrations. Assuming that all four main feedwater lines f all at 180 psig, an equivalent leak area of 200 in2 will result. This fail-ure, therefore, should be classified as Type C. The impact of this change on the containment failure probability numbers will be to reduce the rate for Type 8 with a corresponding increase in Type C. The total failure rate is not likely to change. Estimated containment failure fractions are compared with the ESPSA results in Fig. 3.10. 3.2.5.5 Containment Enclosure
/
The containment enclosure building is designed to withstand 3.5 psipres-sure dif ference between the enclosure and the environment. During nomal operation, the internal pressure is about -0.25 inches of water gauge. The SSPSA study has calculated its pressure capacity to range from more than 1 psid to 10 psid. In view of relatively strong primary containment, the role of the secondary containment is important primarily for Type 8 failures of the primary contaiment. In the event of Type C failure, the secondary enclosure building might not play any significant role as far as the source term calcu-lation is concerned. 3.3 Definition of Plant Damage States and Contaiment Response Classes The grouping of accident sequences into plant damage states proceeds from the premise that the broad spectrum of many plant failure scenarios can be discretized into a manageable number of representative categories for which a , single assessment of core and contalment response will represent the response ' of all the individual scenarios in that category. , The plant damage states classify events in accordance to the following three parameters: ,
- 1. Initiating Events "A" - Large Loss of Coolant Accident "S" .Small Loss of Coolant Accident "T" - Transient e
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- 2. Timing of Core Melt and Conditions at Vessel Failure
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- 3. Availability of Containment Systems "C" - Long-Tem Containment Spray Cooling "4" - Long-Tem Spray Recirculation, No Cooling "1" - Isolation Failure er Bypass Figure 3.11 gives the definition of the plant damage states and their re-spective frequencies listed in Table 3.4 as used in the SSPSA risk model.
These damage states are categcrized in a matrix of eight physical conditions in the containment (numerals (1) to (8)) and six combinations of containment safety function availability (letters A to F) for a total of 48 potential plant damele states. A nint5 damage state type has been defined for accident sequences involving steam generator tube ruptures. Figure 3.11 indicates that only 39 plant damage states can be identified as credible sequences.
~~
From the viewpoint of containment response, many of the plant damage states can be grouped into contalment classes. The classes defined in Table 3.5 are differentiated primarily according to spray availability. The f re-quency of each contaiment class is the sum of the frequencies of the plant states included therein. Annual plant state frequencies calculated by the applicants for both in-ternal and external events were reviewed by the Lawrence Livemore National Laboratory' and were found acceptable. Table 3.6 presents. the calculated contaiment class frequency estimates for internal events, fires, floods and truck crashes; moderate and severe seismic events. In order to comprehensively assess the risk from seismic events, it is necessary to make separate consequence calculations for those accidents which are initiated by earthquakes severe enough to impair evacuation. For this purpose, the seismic frequency estimates are divided into two estegories in Table 3.6. The seismic events with instrument peak ground acceleration below 0.5g can be binned with internal events, fires, ficods and truck e. rashes. Seismic events'with acceleration greater than 0.50g are judged to impair evac- M untion, and must be treated separately in the consequence analysis. These contaiment response classes (or ant damage states) are the starting point for the containment event breejnalysis and they define the link or interf aces with the plant analysis. g 3.4 Contalment Event Tree and Accident Phenomenology , An important step towards the development of the contaiment matrix in. volves the quantification of branch point probabilities in the contain.sent event tree. These probabilities depend heavily on the analyses of degraded ~ and core melt phenomenology and the contalment building response described in Appendix H of the SSPSA.5 9 6
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Plant 5: ate _ Represe nts 2A AEC 4A TEC.SEC 2C/6C AE4 4C/8C TE4 10 AE 20/60 AL 30/70 SE.TE/TEFW 40/80 St.TL 2E/6E AECI 4E/8E TEC1 1F V 2F/6F AEI 3F/7F SEI 4F/EE SLI Figure 3.15 Definitions of the plant damage states used in SPSS.
a . - l Table 3.4 Frequencies of Occurrence of the Plant Damage States Frequency Frequency Plant Dama9e (eventsper Plant Damage (eventsper State reactor year) State reactor year) 10 .3.03(-7) 6A 3.41(-7) IF 1.89(-6) 6C 3.57(-10) 1FA 6.10(-11) 60 2.49(-7) 1FP 8.52(-7) 6E 5.30(-14) 2A 1.85(-6) 6F 2.08(-16) 2C 1.91(-9) 6FA 1.11(-11) 20 2.53(-7) 6FP 1.34(-12) 2E 1.40(-13) 70 7.06(-5) 2F 1.06(-13) 7F 3.55(-8) 2FA 3.10(-11) 7FP 1.09(-5) 2FP 1.58(-10) 8A 4.50(-5) 30 1.94(-5) 8C 4.29(-8) 3F 5.00(-7) 80 5.51(-5) 3FP 6.21(-6) 8E 5.02(-11) - 4A 1.28(-5) 8F 1.02(-10) 4C 1.65(-7) 8FP 1.95(-7) 40 2.79(-6) 9A 7.51(-10) 4E 2.24(-11) 9C 3.62(-13) 4F 2.25(-13) 90 9.09(-9) 4FP 1.18(-7) , TOTAL 2.30(-4) NOTE: Exponential notaticn is indicated in abbreviated form; i.e., 3.03(-7) = 3.03 x 10-7
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Table 3.5 Containment Response Class Definitions Class Plant State Represents 1 10 AE 2 2A/6A 4A/8A AEC, TEC, SEC 3 2C/6C 4C/8C AE4, TE4, SE4 4 30/7D SE, TE TEFW 5 20/60, 4D/8D AL, SL,, TL 6 1F, 2F, 3F, 4F, 6F, V 7F, 8F 7 2E/6E 4E/8E AECI, TEC1 8 IFP, 3FP/7FP Small leaks w/o RWST 9 2FP/6FP,4FP/8FP Small leaks w/ RWST 10 IFA, 2FA/6FA Aircraft crashes 11 9A V2(SGTR) 12 9C V2(SGTR) 13 90 V2(SGTR) l l g 1 k
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Table 3.6 Containment Class Mean Frequenciest Frequency (per reactor year) Containment Internal, Fires, Internal Response Class Floods and Truck Seismic <0.5g Seismic >0.5g Total Seismic and Crashes External 1 1.08E-7 - 1.95E-7 1.95E-7 3.03E-7 2 5.70E-5 1.54E-6 1.24E-6 2.78E-6 6.0E-5 3 1.80E-7 1.91E-8
- 1.91E-8 1.99E-7 4 8.60E-5 1.85E-6 2.27E-6 4.12E-6 9.0E-5 5 5.50E-5 1.10E-6 1.76E-6 2.86E-6 5.8E-5 ,
6 1.80E-6 1.66E-7 3.93E-7 5.59E-7 2.4E-6 7 8
- 5.29E-6 1.25E-5 1.79E-5 1.79E-5 9
- 1.12E-7 2.40E-7 3.52E-7 3.52E-7 10 * - - -
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- Indicates frequencies less than.10-8 yr I.
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The $$PSA containment event tree uses the twelve top events identified in Table 3.7 as major phenomenajogical phases which could occur with respect to the formation and 10,ca,t of, core debris. These processes are grouped into four phases folloy j(cident initiation (1) phenomena occurring while the core is still e; (2) phenomena occurring while the core is located be'ow tne lower gri ate but is still in the reactor vessel; (3 nena occurring with the core debris located in the reactor cavity a d onthe n-tainment floor; and (4) the phenomena involving long-term coolin con-tainment and/or basemat penetration. 3.5 Containment Matrix (C-Matrix) Tne twelve top events in the Seabrook contaiment event tree are summar-ized in Table 3.7. A negative response at any of the five nodes (4. 8,10 [j 11, and 12) in the contalment event tree results in the failure of the con-tainment bui'cing by a variety of f ailure modes. Each of these failure modes results in a particular radiological release category. For those paths that do not have a negative response at any of the five nodes, the path will even-tually result in no f ailure of the containment. The conthiment event tree thus links the plant damage states to a range of possible contaiment failure modes via the various paths through the tree. For a given tree, each path ends in a conditional probability (CP) of occurrence, and these cps should sum to unity. The quantification of an event tree is the process by which all the , paths are combined to give the conditional probabilities of the various release categories. In S$PSA, fourteen release categories are used for the quantification as sumarized in Tab'e 3.8. Note that two of these release categories (namely, SS and 6) correspond to intact / isolated containment. Fission product release for this category would, therefore, be via nomal leakage paths in the containment (3rd enclosure) bdiding, which can be,4ifNS ferent-depending on availability of the enclosure building v'entilatio f fil-prati,on system. Table 3.9 sets forth a simplified containment matrix (C-matrix) for the Seabrook plant using the contaiment response class definitions discussed in Section 3.3, and the release category definitions given in Table 3.8. In arriving at the C-matrix of Table 3.9 all of the very low probability values were disregarded. This is shown 7 to be insignificant to the risk estimate. The present assessment of contaiment response for Seabrook plant is not based upon independent confimatory calculations of accident progression and contaiment response. Instead the knowledge gained from review of similiar risk studies for other 1 ,3," , pressurized water reactors with large dry containments is used to guide this assessment. The mode and timing of contalment failure cannot be calculated with a great degree of accuracy. Judgements must be made about the nature of the dominant phenomena and about the magnitude of several important parameters. Furthemore, the codes and methods used for these calculations are approximate + and do not model all of the detailed phenomena. Fortunately, risk measured in personal exposure is not sensitive to minor variations in fatture mode /and. timing. It is important, however, te pre;erly characterize the major attri-butes of f ailure mechanisms; (1) whether the failure is early or late. (2) whether it is by overpressurization, bypass, or basemat melt-through and, (3), whether or not radionuclide removal systems are effective, d 0 1
Table 3.7 Accident Phase and Top Events for the Seab, rook Containment Event Tree Accident Phase Top Event Initiator 1 Plant State Debr in Vessel 2 Debris Cooled in Place 3 No H2 Burn 4 Containment Intact Debris in Reactor Cavity 5 Debris Dispersed from Cavity 6 Debris Cooled
. 7 No H2 Burn 8 Contair. ment Intact -
Long-Term Behavior 9 No Late Burn 10 Containment Shell intact 11 Basemat Intact Failure Mode 12 Benign Containment Failure * (Small Leak) i - t 9 6 i i O 1 i
Table 3.8 Release Categories Employed in the Seabrook Station Risk Model Release Category Release
- Group Category Definition 55 Containment intact / isolated with enclosure Containment air handling filtration working.
Intact / Isolated , 55 Same as 55 but with enclosure a r handling filtration not working. 52 Early containment leakage with late over-pressurization failure and containment building sprays working. 37 Same as 52, but with containment building spray not working. TN Same as 37, but with an additional vaporiza-tion compenent of the source term. 53 Late overpressurization failure of the con. - Long-Term tainment with no early leakage and contain-Containment ment building sprays working. Failure - 3T Same as $3, but with containment building sprays not working. 3Ti Same as 37, but with an additional vaporiza-tion component of the source term. - 54 Basemat penetration failure, sprays operating 3Ti Containment basemat penetration failure with ' containment building sprays not working and additional vaporization component of the source term. 56 Containment bypass or isolation f ailure with containment building sprays working. 3ri same as $6, but with containment building sprays not working and an additional vaport. Early ration component of the source term. Containment Failure /sypass 51 Early containment failure due to steam exp1'o- ! sion or hydrogen burn.with containment building sprays working. . l 3T Same as $1, but with containment building j sprays not working.
*S denotes applicability to Seabrook Station; number corresponds with contain-l ment failure mode; bar denotes nonfunctioning of containment building sprays; 1
and V denotes achievement of sustained elevated core debris temperatures and l associated vaporization release. l l
~ . Table 3.9 Sirrplified Centainment Matrix for Seabrook Release Category Class 51 S2 53 55 56 R U 37if M M N E7.c 1 . 0.60 0.40 2 0.01 0.99 3 1.0 4 0.89 0.11 5 1.0 6 1.0 7 1.0 8 1.0 9 1.0 10 1.0 11 1.0 12 1. 0 13 1.0 a e 9' s e t e 9
The assessment of the containment response and failure mechanisms is based on the general understanding of that accident phenomenology and the con-tainment design characteristics discussed earlier. The phenomena of interest
- 7 may be sunnarized as follows:
Early Failure (51. Yl") which n result f rn a steam explosion or an early'$yh [ t orogen burn ts celleved to e unlikelf . Altnough explosions in the reactor' vessel lower plenum are probable, the resulting mechanical energy w3uld . be limited by the fraction of the core which could participate in a sing"le 7ex-plosion and by the efficiency of the process. In recent PRA revi44s e we have assigwd a conditional probability of 10-6 to steam explosion induced contaiment failure. This probability lead: to the conclusion that, steam ex-plosions would have a negligible effect on risk, and consequently, the appit-cants 5x10-* value is not included in the simp 1tfied C-matrix. The conditional probability for an early containment failure 'due to e$- ? ternal mnts (i.e.; aircraft crashes) is assigned 1 in the SSPSA as shown.in Table @ This simply indicates that an aircraft crash ir.to the contalment ' isassumedtofaillthecontainmentstructurewithcertainty.
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Early containment fa'ilure could also conceivably result from direct heat-ing due to a rapid dispersal of the core debris throughout contaiment in the
- form of aerosols. The dispersal could only be caused by the high primary sys , ~
tem pressures that may exist at vessel failure for a number of transient se-quences (recent calculations 11 indicate that there exists a propensity for - l establishment of natural convection pattern inside the reactor vessel and the , hot leg; which can cause rapid heatup of the ACS boundaries possibly leading to failure and depressurization prior to bottom head melt through, thus elim-inating, high pressure ejection sequences). The aerosols could rapidly pres-c p surize containment by direct heat exchange and concomitant chemical reac-tions. Scoping calculations perfomed by the Containment Loads Working Gecup (CLWG) showed that a very severe challenge to the contalment integrity result provided 25 percent However, no consensus could be reached among the CLWG analysts as to tne cred < 4 ibility of this parameter value, and this failure mode is still speculative. Furthemore, the configuration of the Seabrook lower cavity would tend to re- T duce the dispersal of core debris beyond the reactor cavity boundaries. - For the reasons outlined above as well as the high containment failure pressure for Seabrook, it is concluded that early overpressure failure has a very low itkelihood. h Early containment Leakage ($2. 37. TN) without gross failure of containment builoing is expecteo to occur for nonisolated steam generator tube rupture ilable (52), for large break LOCA sequences
, , event with containment sprays ava with RWST injection in the absence of sprays (T2), and for dry cavity sequences with a vaporization release (377). , .
There seems to be a basic inconsistency in assigning plant damage sts'tes 3 Specifica1*y, large brear to this failure mode as defined in the C-matrix. LOCA sequences with RWST injection in the absence of contalwent sprays are. expected to lead to an 51 failure mode with 100% probability (see 37 below),- l t i f i
y i' ?.
" ' ' . -43 while they are also' assigned to Tl with 100% probability. This can be correct only if the initiatre and the sequences are indeed different, but at this time k cannot resolve the inconsistency.
t similarly, the significance of contaiment functions on steam generator /J
't.uhe }upturp sequences is not at all obvious.
# date Overpressuriza*. ion Failure (53, IT, TN) can occur due to steam produc.
YRn in a wet cavity or not.condensable gas production as a result of core-con-
' crete interactior, for. a dry cavity situation. For sequences in which early and intermediate failure is not expected to occur, and for which contaiment sprays are inoperabic, fatture is expected to be a certainty.
~
l The conditional probability for a late overpressurization failure with a
., vaporization release. (dry cavity) is shown to be 0.60. This results from the relative competition between the late overpressure failure and the basemat
" penetration (TW) for accident sequences without the containment sprays.
The failure time for the late ovtrpressurization failure mode is much o ntai nme nt .1,3," longer than previously calculated for other large dry This is as a result of the very high failure pressure for
- te Seabrook con-tainment. As a consequence of this high contalment failure pressure (median pressure of 211 for wet and 187 psia for dry
- sequences) it is difficult to
- challenge the contaiment integrity by any conceivable event.
i Hydr: gen deflagration early in the accident sequence or later after vessel failure when steam condensation occurring as a result of reactivation of sprays (due to regaining of ac power), or other natural heat sink mecha-nisms, which can produce a deinerted atmosphere is not expected to challenge the contalrunent integrity. I The impact of changes in the contaiment failure distribution discussed
- in 3.2.5.4 is not significant for late failures. '
t j Basemat Penetration Failure ($4, TIV) can only result in the absence of con-4 tainment neat removal system (sprays) for a dry cavity. A 26-inch high curb surrounds the reactor cavity that prevents the entry of water into the cavity i i unless the full RWST has been injected. The conditional probability of the i basemat melt though is always less than the late overpressurization failure, l particularly for Seabrook with the natural bed rock femation directly under the basemat foundation. Therefore, the basemat penetration failure probabil-l ities are conservatively assigned. h l No Failure (55. TT) would result for all sequences with full spray operation. I The raciological releases are thus limited to the design basis leakage with ! essentially negligible of f-site consequences. ! Contalment isolation Failure ($6. 569 is represented by an 8-inch diameter purge line. Ine accident sequences where the containment is either/not ( 'For dry sequences, only primary system water inventory is available in the ! , contai nment. In this case, the contalment atmosphere becomes superheated ! and, at f ailure, the temperature can exceed 700'F. l
isolated or bypassed (Event V) are assigned a conditional probability of unity to this release category. An 1.nterfacing systems LO'CA (Y sequence) results from valve disc rupture or disc failing open for series check valves that nomally separate the high pressure system. This event results in a LOCA in which the reactor coolant bypasses the containment and results in a loss-of-coolant outside the cor.tain-me nt . Furthemore, the concurrent assumed loss of RHR and coolant make-up capability leads to severe core damage. In the SPSS, three possible inter-facing systems LOCA sequences have been found and discussed. These are
- 1. Disc rupture of the check valve in the cold-leg injection lines of the RHR.
- 2. Disc rupture of the two series motor-operated valves in the normal RHR hot-leg suction.
- 3. Disc rupture of the motor-operated valve equipped with a steam mount-ed limit switch and " disc failing open while indicated closed" in the other motor-opercted valve in the nomal RHR hot-leg suction.
For the V-sequence, the core melts early with a low RCS pressure and a dry reactor cavity at vessel melt-through. The containment sump remains dry . anc recirculatior, is not possible. The core and con;aiment phenomenology used to arrive at the split frac-tions for the containment event tree and thus the C-matrix are in general agreement with the other previously NRC reviewed studies ,3, for PWRs 1 with large dry contaiments. Furthemore, the claimed unusually high strength of the Seabrook centaiment reduces the impact of sensitivity ' caused by uncer-tainties in the severe accident progression. However, should the claimed stre gth of the containment be reduced to levels comparable to some of the other large dry containments, the impact of uncertainties may become signifi-cantly more pronounced, as discussed in our review of the HPSS-3.7 3.6 Release Cateoory Frecuencies Based on the cantainment class frequencies in Table 3.6 and the contain- - ment failure matrix of Table 3.9, the release frequencies were computed and are summarized in Table 3.10. M/[/ Table 3.10 indicates that on1 light ,f the release categories dominate the total release frequency. Tables 3.11 and 3.12 set forth the contribution to core melt frequency
- from the various contairrsent response classes and release categories, respec-tively. It is seen that containment classes 2, 4, and 5 cominate the core melt f requency whil ,% the release categories S5 (containment intact),13 and 3Ti doninate the sou c tem frequency. ,
1 l l I f
+ , . , - - - - - - . . . - - -.-,----..-------,,.----.,----,-..n- .
Table 3.10 Frequency of Dominant Release Categories (yr-1) Internal, Fires. Ficods and Truck . Internal and Category Crashes Seismic <0.5g Seismic >0.59 External 53 7.50E-7 3.45E-8 2.69E-7 1.05E-6 55 5.64E-5 1.52E-6 1.23E-6 5.92E-5
- 1.12E-7 2.40E-7 3.52E-7 lCf lCi 5.50E-5 1.10E-6 1.76E-6 5.79E-5
- 1.25E-5 1.78E-5 52V 5.29E-6 3CEI 7.66E-5 1.65E-6 2.14E-6 8.04E-5 1907 9.50E-6 2.04E-7 3.27E-7 1.0E-5 337 1.80E-6 1.66E-7 3.93E-7 2.36E-6 O
O 4 H 9 3
l Table 3.11 Contribution of Containment Response Classes to the Total Core Melt Frequency I
! Internal, Fires, . internal Containment Floods and Truck and Class Crashes Seismic <0.5g Seismic >0.59 Total Seismic External
! .1 - - - - <0.01 2 0.25 <0.01 <0.01 0.01 0.26 3 - - - - <0.01 4 0.37 0.01 0.01 0.02 0.39 , 5 0.24 - 0.01 0.01 0.25 6 0.01 - - - 0.01 7 8
- 0.025 0.055 0.08 0.08 g.33 * * * *
- l
=
e
'h t 4 .
- r Table 3.12 Release Category Frequency as a Fraction of Core Melt Frequency Release Internal Fires.
Internal and l' Category Floods and Truck ' Crashes Seismic <0.59 Seismic >0.59 Total Seismic External
<0.01 i <0.01 <0.01 <0.01 <0.01 S3 55 .
' $0 .25 <0.01 <0.01 0.01 0.26
* <0.01 <0.01 <0.01 <0.01 i 52
; 3Ci 0.24 <0.01 <0.01 0.01 0.25 ,
- 0.03 0.05 0.08 0.08
! S2V 0.33 0.01 0.01 0.02 0.3S J 53V e
.' <0.01 <0.01 0.04 S4V 0.04 <0.01 {l 0.01 <0.01 <0.01 <0.01 0.01 56V i s I i I 1 i l . I I I I l
4. ACCIDENT SOURCE TERMS In this chapter the approach utilized in the SSPSA to determine the f raction of fission products originally in the core which can leak to the out-side environment will be outlined. The fission product source to the environ-ment as calculated by this approach for each release category will also be discussed. . 4.1 Assessment of Severe Accident Source Terms As in the Reactor Safety udy (RSS)l3 the CORRAL-II code was used in the )( SSPSA for determining fission product leakage to the environment. This code takes input from the thermal-hydraulic analysis carried out for the contain-ment atmosphere. In addition, it needs the time-dependent emission of fission products. The fission products were assumed to be release'd in distinct phases as suggested in the RSS, namely, the Gap, Melt, and Vaporization phases. The time dependence of these phases is determined by the timing of core heatup, primary system f ailure, and core / concrete interaction. The methods used in , the SSPSA differ from the RSS methods in the following ways:
- 1) The treatment of iodine was changed and iodine was treated as cesium iodide. This was accomplished by merely using the same fraction of core inventory released for both the cesium group and the iodine group,
- 2) Leakage releases are represented by a multi-puff model,
- 3) An uncertainty analysis was carried out in which it was attempted to "
account for shortcomings in the RSS methods. In general, the net result of the SSPSA analysis was to reduce the fractional release of particulate fission products. This will be discussed in more de-tail later. In all, fourteen releases were dete. mined ranging from contain-ment bypass sequence to the no-fall sequence as shown in Table 3.8. These release categories were evaluated by considering the containment failure mode, the availability of the spray system, and whether or not the cavity was wet or dry. Table 4.1 shows the point-estimate releases as deter-mined by the methods outlined above. Containment failure mode S1 corresponds to a gross f ailure of the containment, resulting from a steam explosion, early pressure spikes, or early hydrogen burns. Failure mode S2 represents a loss of containment function early in the accident sequence. This loss of function takes the form of an increase in the leak rate to 40% per day where it stays until the containment fails due to overpressurization. Failure mode 53 repre-sents a late overpressurization f ailure of the containment driven by decay heat or late hydrogen burns. Failure mode S4 represents a basemat melt-through, SS represents no containment failure and the leak rate is limited to the containment design basis leak rate. Finally, failure mode 56 represe.nts sequences where the containment is failed or bypassed as part of the initiat-ing event. The second parameter considered in defining the source term is the avail-ability of sprays. This is determined by the plant damage states. Those
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, - v - --. - , . . - - - - . -
o < . . release categories with operating sprayj systems are designated 51 to 56, while those with spray systems,not operating are designated'5T to T6. The third and final parameter considered in differentiating between source terms distinguishes between wet and dry cavities. In the case of dry cavities a vapo,rization release due to core / concrete interactions will occur, while for wet cavities the core debris is assumed to be quenched or the water in the' cavity will scrub the vaporization release thus effectively reducing the release to zero. The release categories which include a vaporization re-lease include a "/" in their designation as shown in Table 3.8. From the point of view of risk it was found 5 that 5TI, 5T, M, and T67 were dominant either for acute or latent health effects. In view of this re-sult these four categories will be considered in more detail.
, Release categories TI and M have late overpressurization failure modes, with no spray systems operating and differ only in the omission or inclusion of a vasorization release, respectively. The containment at Seabrook is cal-culatec to fall at a median pressure of 211 psia for wet sequences and 187 psia for dry sequences. At this pressure a gross failure is expected result-ing in a puff release of approximately 0.5 hr release duration. From Table 4.1 it is seen that the IT and 577 sequences fail at 27.2 hrs and 91.5 hrs, respectively. These failure times are several hours later than was cal-culated for Indian Point, Zion, and Millstone-3. The primary reason for the later failure in this case is due to the superior strength of the containment structure. Table 4.2 compares the 5'3, 577 release parameters with similar parameters for the o ,her three reactors mentioned above. Note that a fair comp arison should set (Ol+1) equal to (Cs-Rb), since iodine was treated as Cs!. It is seen that I, Cs, and Ba groups for 5T are approximately half the other releases, while the Te, Ru, and La groups are low by approximately l an order of magnitude. This difference is due to the latter failure time, I
allowing more time for settling and the absence of a vaporization release, which dominates the release of Te, Ru, and La. A similar comparison for the
'5TI release indicates a uniform reduction of approximately an order of magni-tuce for all species. The reduction is entirely due to the late failure time for this sequence.
Another important consideration is the increased rate of release due to an increase in the leak area prior to attaining gross failure conditions. This can also impact the radionuclide transport mechanisms inside the contain-ment due to changes in the containment thermal hydraulic conditions. Release category Ifi is associated with early contai nment failure in which the contairnent function is compromised by incre3 sing the leakage area in such a way that the leak rate increases from 0.1% per day to 40% per day. 3 This release rate is not enough to prevent an ultimate overpressurization failure. This release is modeled as a multi-puff
- release. The first puff corresponds to the release up to the time when vaporization starts (mel t+ga s) . The secon pu cludes the period of vaporization release .and ,
the third puff is equi to an overpressurization failure at the time of catastrophic containne failure. In this model the duration of the melt i , I
' N N f .
Table 4.2 Late Overpressurization Failure Comparison Millstone-3 7 Zion / Indiani " Indian 3 Seabrook5 Point Study point 33r jagr g.7 TMLB' 2RW Xe 9.0(-1) 1.0 9 (-1) 9.6(-1) 1.0 10+1 1.2(-1) 2.4(-2) 1.5(-1) 1.05(-1) 9.3(-2) Cs-Rb 1.2(-1) 2.4(-2) 3.0(-1) 3.4(-1) 2.6(-1)
~
Te-Sb 2.2(-2) 3.0(-2) 3.0(-1) 3.8(-1) 4.4(-1) Ba-Sr 1.5(-2) 2.6(-3) 3.0(-2) 3.7(-2) 2.5(-2) Ru 4.4(-3) 2.3(-3) 2.0(-2) 2.9(-2) 2.9(-2) La 4.4(-4) 3.9(-4) 4.0(-3) 4.9(-3) 1.0(-2) T (release) 27.2 81.5 20 (hrs) T (duration) 0.50 0.50 0.50 0.50 (hrs) Energy 300E7 300E7 540E6 150E6 (Btu /hr) e e 4 e
release is s n to be 3.5 hours. vaporization release 7.2 hours and the re-maining ase 78.0 hours. It is not clear that the melt release in this case is 5.3 hours, however, it does not seem to be unreasonable. A 7 2 hour duration or the vaporization release is not consistent with the RSS,13 which only allows 2 hours for this phase. Finally, it is not clear how the 78 hours for tr.e last phase was detern.ined. The release duration for a single puff, ' which is ;he sum of the above three phases leads to a release time of 88.7 hours whic" seems extraordinarily long. Our recommendation would be to reduce 2Q these times to be more consistent with RSS methods (see Table 4.7). The total rele:se of fission products from the sequences can be compared to the M-4 release determined for the Millstone-3 study. This comparison is made in Table 4.3. It is seen that, once adjustments are made for the dif-ferent ways in which iodine is treated, the 57i release is approximately half the M-4 re12ase. Without the benefit of a calculation, it is difficult to judge whether the differences are reasonable. However, a possible reason for this reduction is the credit taken for the enclosure building surrounding the /C actual containment building. This feature is unique to the Seabrook contain-ment structure. Release category T6V has binned into it an isolation failure corresp:nd-i ng to an 8" diameter breach in containment and the interf acing LOCA (V-sequence). This sequence is also represented by a multi-puff release. In
- this case as in the previous case, the total release time is long compared to acceptable limits of the RSS 13 consequence mocal. Our recommendation would be /f to reduce these times to more reasonable values (see Table 4.7).
The release fraction can be compared (Table 4.3) to the M 4 release from M Millstone-3, PWR-2 for the RSS and the V-sequence from the RSSMAP study for Surry.15 Except for the iodine group, it is seen that the release fracticns are comparable. If the iodine group were set equal to the cesium group value, it is seen that tre value for S6V would be the lowest release fraction. 4.2 Source Term Ur. certainty Analysis In this section we will briefly describe the uncertainty analysis carried out for the four dominant accident sequences and, where possible compare the fission product leakage to the en'vi ronment to more mechanistic determina-tions. There are two contributors to the uncertainty in release characteriza- ' tion. First, uncertainty in time parameters which are influenced by:
- 1) Prediction of key event times, and
- 2) The mix of accident sequences binned into a release category.
Second, uncertainties in release fractions, which are influenced by: -
- 1) Analysis methods and data, and +
- 2) Uncertairties in timing of key events.
O r- -. _ . _ _ _ ._ . -, . _ . _ _ . . . . . , _ _ ._ _ _ _ _ , . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
e Table 4.3 Comparison of Releases for Failure to Isolate Centainment and the By-Pass Sequence Seabrook 5 Millstone-3 7 RSS13* RSSMAP15 T2Y T67 M-4 PWR-2 V-Sequence l.0 9.7 (-1) 9.0(-1) 1.0 1.0 Xe 0l+1 3.1(-1) 4.3(-1) 2.0(-1) 7.0(-1) 4.8(-1) Cs-Rb 3.1(-1) 4.3(-1) 6.0(-1) 5.0(-1) 7.9(-1) Te-Sb 3.2(-1) 4.0(-1) 5.0(-1) 3.0(-1) 4.4(-1)
~
Ba-Sr 2.4(-2) 4.8(-2) 7.0(-2) 6.0(-2) 9.0(-2) Ru 2.5(-2) 3.3(-2) 5.0(-2) 2.0(-2) 4.0(-2) La 4.2(-3) 5.3(-3) 7.0(-3) 4.0(-3) 6.0(-3) 2.2 2.2 0.20 2.5 2.5 T (release) (hrs) 88.7 14 2.0 1.0 1.0 T (duration) . (hrs) 4E6 70E6 20E6 0.5E6 Energy (Btu /hr) 140E6
*The same as M1A release category in Millstone-3.7 .
.. . 2,o o
l e
a a , , l The above principles were used to determine source term multipliers which ; roduct leakage to the environment. A probabil-would give a rangewith ity ,is associated of fission each p" source term, and for later overpressurization j j f ailure modes (53, 53V, and 577) the following discrete probability distribu- < tion is used, i .e., Subcategory Probability 5T-a .02 IT-b .08 IT-c .30 3T-d .60 This indicates, for example that there is an 81 confidence level that 5T-b correctly defines the source term for the 3]I release category. The results of this analysis for the overpressurization failure modes is: Particulate P.elease Factor (nultiplier) Probability 37 33Y 57[
.22 .63 .17
.02 .07
.08 .071 .22
.024 .065 .02
.30 .007
.60 .0071 .021 From this table it is seen that for the most likely release, i.e., "d", the reduction fac ors of the source term are substantial.
The first two releases can be compared to releases published in BMI-2104 Volume V (Surry) for the TMLB'.-c and AB-c sequences. These two sequences - correspond to late containment failures and are both binned into 53 From and S3V this sequences. A comparison of these sequences is shown on Table 4.A. table it is evident that for the volatile species Xe, Cs, and I, the release categories 3]I and 537 bracket or exceed the mechanistic estimates carried out in BMI-2104 for both TM:B' and AB sequences. However, for the less volatile species Te, Ba, Re, and La, the release of the AB sequence is the only one bracketed or superseded by the IT and 577 releases. The release f ractio n detennined for the TMLB' sequence is higher than all the if and S3v releases. This discrepancy is primarily due to the comparatively early failure time. It is felt that agglomeration and settling would reduce the source for the TMLB' No comparitive sequence to values close to those reported for f3'and 53V. sequence for S2V was analyzed in BMI-2104. I In the case of the 337 release category a different probability distribu-tion was used. This change reflects the break location, which initiates the 9 l
l Table 4.4 Comparison of AB-c and TML8'-c (BMI-2104) to UV and U P Release Fractions Release Probability Release ' Category Time (hrs) Xe Cs I Te 8a Ru La
$3v-a .02 28 1.0 1.5(-2) 1.5(-2) 1.9(-2) 1.6(-3) 1.5(-3) 2.5(-4)
$3V-b .08 36 9.0(-1) 5.3(-3) 5.3(-3) 6.6(-3) 5.7(-4) 5.1(-4) 8.6(-5)
S3V-c .30 54 8.0(-1) 1.6(-3) 1.6(-3) 2.0(-3) 1.7(-4) 1.5(-4) 2.5(-5) , 4 4
$3v-d .60 89 7.0(-1) 5.0(-4) 5.0(-4) 6.3(-4) 5.5(-5) 4.8(-5) 8.2(-6) ,
l U-a .02 22 1.0 2.6(-2) 2.6(-2) 4.9(-3) 3.3(-3) 9.7(-4) 9.7(-5) U-b ,.08 28 9.0(-1) 8.5(-3) 8.5(-3) 1.6(-3) 1.1(-3) 3.1(-4) 3.1(-5) j U-c .30 34 8.0(-1) 2.9(-3) 2.9(-3) 5.3(-4) 3.6(-4) 1.1(-4) 1.1(-5) E-d .60 53 7.0(-1) 8'5(-4)
. 8.5(-4) 1.6(-4) 1.1(-4) 3.1(-5) 3.1(-6)
TMLB'-c - 12 1.0 2.8(-3)- 6.0(-4) 8.5(-2) 1.7(-2) 2.4(-5) 4.3(-4) j AB-c , 24 1.0 4.8(-5) 4.7(-5) 4.0(-5) 4.9(-5) 2.4(-7) 3.6(-5)
.)
l V-sequence. This break could be either in the hot-leg (b release subcategory) J or the cold-leg (c release subcategory). This sequence is modeled as multi-puff release and each puff is treated separately. In this comparison only the sum of the release will be considered, since no adequate method of analyzing a multi-puff release is readily available. Table 4.5 shows a comparison be-tween the totals of tre various 567' releases and two V-sequence releases com-puted for Surry and published in BMI-210c. One of the V-sequences is " dry," implying no water in the path of the release and the other is " wet," implying that the release passes through 3 feet of water before entering the atmo-sphere. From this comparison it can be seen that all the releases, except Cs for the " dry" V-sequence, are bracketed by the 3I7 releases. 4.3 Recommended Source Terms The severe accident source terms used in the Seabrook Probabilistic Safe-ty Study reviewed in the previous sections, are aimed at the multi-puff con-sequence model present in the CRACIT computer code. In order to make these source terms useful to the NRC staff for evaluation with the CRAC code, total releases must be used as sucmarized in Table 4.6. Furthermore, the suggested source terms of Table 4.6 together with their release category characteristics given in Table 4.7 have been adjusted to more closely represent our assessment j:7 of the severe accidents based upon the RSS methodology. It must also be noted that the suggested source term for the Steam Gener-ator Tube Rupture (SGTR) sequence is assumed to be one-tenth of the source term for the event V (567). This is believed to be a conservative estimate and can be used in the absence of a more specific mechanistic calculation. The suggested source terms can be used to estimate the number of health and economic effects (consequences) in.the population surrounding the Seabrook Station due to radioactive atmospheric releases as a result of core melt acci-dents. The resulting consequences together with the frequency of radiological releases will enable the establishment of the severe accident risk at the Sea- ! brook site considering the double-reactor unit effect. 1 0 e 9 O e
- , - - , - - , - , - - , , , - ,,m,, ,-n- , - ,,.-,,,.,-na, , - , , , - - - - - , - ,+
,,-- - ,v._. , - , - - -
v - -
,-w----m,, , , - , , , . , , - - , .
l l . l Table 4.5 Comparison of 56v (sum) to V-sequence (Surry) 0 Release Fractions Release Probability l Category .Xe Cs i Te Ba Ru La i 56v-a .02 .97 4.3(-1) 4.3(-1) 4.06(-1) 4.2(-2) 3.32(-2) 5.3(-3) 56v-b .45 .97 2.95(-1) 2.95(-1) 1.36(-1) 3.53(-2) 1.52(-2) 2.0(-3) i
- 56v-c .45 .97 1.295(-1) 1.295(-1) 3.2(-2) 1.593(-2) 5.2(-3) 5.3(-4) 0' 56v-d .08 .97 5.2(-2) 5.2(-2) 1.3(-2) 6.6(-3) 2.0(-3) 2.2(-4)
~
V (dry) ,
- 1.0 5.52(-1) 1.99(-1) 1.2(-1) i i
V - 1.0 1.04(-1) 3.84(-2) 2.5(-2) I (submerged) l i
- Individually not reported.
~.
i
)
, .. . 3 Table 4.6 BNL-Suggested Source Terms Release Category Xe 01 1-2* Cs Te 8a Ru La 51 0.94 -
0.023 0.023 0.24 0.0033 0.41 9.8E-5 52 0.89 - 2.1E-5 2.1E-5 4.4E-6 2.9E-6 8.8E-7 8.8E-8 53 0.90 7E-3 1.E-7 1.E-7 1.9E-8 '1.3E-B 3.8E-9 3.8E-10 55 0.0091 - 3.5E-8 3.5E-8 6.1E-9 4.0E-9 1.2E-9 1.2E-10 56 0.90 - 3.6E-3 3.6E-3 6.7E-4 4.4E-4 1.3E-4 1.3E-5 lCI 0.94 - 0.75 0.75 0.39 0.093 0.46 2.8E-3 , 3Cf 0.90 - 0.31 0.31 0.057 0.038' O.011 1.1E-3 3CFI 1.0 - 0.31 0.31 0.32 0.034 0.025 4.2E-3 i ICI 0.90 - 0.12 0.12 0.022 0.015 4.4E-3 4.4E-4 lCfi 1.0 - 0.024 0.024 0.030 2.6E-3 2.3E-3 3.9E-4 S4V 1.0 - 0.058 0.058 0.072 6.2E-3 5.4E-3 9.1E-4 , SS 0.014 7E-4 5.2E-7 5.2E-7 9.5E-8 6.3E-8 1.9F-8 1.9E-9 S6v 0.97 - 0.43 0.43 0.40 0.b48 0.033 5.3E-3 56v-d 0.90 - 0.043 0.043 0.040 4.8E-3 3.3E-3 5.3E-4
- Elemental iodine not used, all iodine treated as CsI.
**S6v-d release is 1/10th of the 56v values. ,
e e
59-
*. \
\
Table 4.7 BNL-Suggested Release Characteristics for Seabrook Release Release Release Release Warning
- Release Time Duration Energy Height Time Category (hr) (hr) (Btu /hr) (m) (hr) 51 1.9 0.5 140E6 10 0.35 52 2.6 1.0 0.5ES 10 1.05 S3 66.0 0.5 250E6 10 63 SS 1.9 10 n/a 10 0.35 56 4.5 4 0.5E6 10 0.50 TE 1.4 0.5 520E6 10 0.30 TE 27 10 10E6 10 26 Tf7 35 10 25E6 10 35 TI 27 0.5 250E6 10 26 Tiv 81 0.5 450E6 10 76 54v 50 0.5 250E6 0 49 4.3 24 10E6 10 0.30 56v 2.5 1.0 0.5E6 10 1.0
~~
56V-d 2.5 1.0 0.5E6 10 1.0
- Warning time is defined as the time after core melt starts to the time of radiological release.
D d O cm 9
, ,. - - - - , , ,.,,~__.-..-_7-_ _ _ _ _ _ _ _ _ _ . __- ___.c_ .. -- __ _ , ,, _ - - , ,.
.>e . ,
- 5.
SUMMARY
AND CONCt.USIONS The purpose of this report is to describe the technical review of the Seabrook Station Probabilistic Safety Assessment and to present an assessment of containment perfomance, and radiological source term estimates for severe core melt accidents. The 'contaiment response to severe accidents is judged to be an important f actor in mitigating the severe accident risk. There is negligible probabil-ity of prompt contaiment failure or failure to isolate. Failure during the first few hours after core melt is also unlikel Most core melt accidents would be effectively mitigated by contaiment spr y operation. Our assessment of the containment failure characteristics indicate that, there is indeed a tendency to fail contaiment through a realistic benign mode compared with the traditional gross failures. t[ The point-estimate release fractions usedlin the SSPSA are comparable in magnitude to those used in the RSS. In those' cases where comparisons can be made to the more mechanistic source tem study carried out by the Accident Source Tem Program Of fice (ASTFO) and reported in BMI-2104 it was found that the SSPSA releases were either higher than op' for the most part similar to the recent release fractions. It was also foupd that the energy of release was somewhat higher in the SSPSA than for other/ existing studies.
./
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o uAs h-i423 fel em Tjy'- [ viss w - m o c(ruD SS?SA hlem Type "7. *f C rn 7s jl t? I. Gem, VM ER LLJ SH l l lt. G rab'n \ % O (, (, - N l Ovngnuar< or & . tbg l 3.ra,ac 4 # o n I
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l i
- 6. REFERENCES
- 1. " Zion Probabilistic Safety Study," Commonwealth Edison Company (September 1981).
- 2. " Lime ri ck Probabilistic Safety Study." Philadelphia Electric Co.
(September 1982).
- 3. " Indian Point Probabilistic Safety Study," Power Authority of the State of New York and Consolidated Edison Company (March 1982).
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- 8. J. Ga rrick, et al., "Seabrook Station Probabilistic Safety 5.
Assessment," Pickard, Lowe and Garrick, Inc., PLG-0300 (December 1983).
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12,1984).
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R. O. Wooten and H. Avei, " MARCH: Meltdown Accident Response Character-8. istics - Code Description and User's Manual," BMI-2064, NUREG/CR-1711 (1980). An Improved Model for Molte n
- 9. J. F. Puis, et al . , "CORCON-Mod 1:
Core / Concrete Interactions," SAND 80-2415 (1981).
- 10. 8. E. Miller, A. K. Agrawal, and R. E. Hall, "An Estimation of Pre-Exist-ing Containment Leakage Areas and Purge and Vent Valve Leakage Areas Re-sulting from Severe Accident Conditions," A-3741,11/15/84 (D. aft report dated August 1984) transmitted via letter to V. Noonon, June 29, 1984.
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taiment Loads Working Group, NUREG-1079 (Draft 1985). WASH-1400,
- 13. " Reactor Safety Study," U.S. Nuclear Regulatory Commission, NUREG-75/014 (October 1975). ,
14 " Preliminary Assessment of Core Melt Accidents at the Zion and Indian Point Nuclear Power Plants and Strategies for Mitigating Their Ef fects," NUREG-0850, Vol.1 (November 1981). Kolb, et al., " Reactor Safety Study Methodology Application
- 15. G. S.
Program: Oconee #3 PWR Plant," NUREG/CR-1659/2 of 4 i
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