ML20116G512
| ML20116G512 | |
| Person / Time | |
|---|---|
| Site: | Point Beach  |
| Issue date: | 07/31/1996 |
| From: | Gross E VECTRA TECHNOLOGIES, INC. |
| To: | |
| Shared Package | |
| ML19311C175 | List: |
| References | |
| PBNP-1C-12, PBNP-1C-12-R01, PBNP-1C-12-R1, NUDOCS 9608080150 | |
| Download: ML20116G512 (236) | |
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DOCUMENT APPROVAL COVER SHEET VECTRA DOCUMENT NUMBER: , __PBNP-IC 12 Technologies Inc. DOCUMEW TITLE: Low and High Pressurizer Pressure Reactor Trio l Instrument Uncertaintv/Setooint Calculation DOCUMk.NT TYPE.
~ sNSNEo" CLENT: WISCONSIN ELECTRIC POWER COMPANY
_ PIPORT
- Ni PROECT: _PBNP Setooint Verification Program PROECT NUMBER: ___0087-00033.303
SUMMARY
DESCRFIlON:
REVISION: DESCRIPTION:
l 0 OriginalIssue l TOTAL NO. OF PAGES 32 plus 1 attachment pages ORIGINATOR: Karen L. De Podesta DATE: 7/17/96 VERIFER: Kim A. Jacklin DATE: 7/17/96 APPROVER: Larry P. Lawrence for Joe Basak DATE: 7/19/96 REVISION: DESCRIPTION:
1 Revised to incorporate setpoint evaluation.
TOTAL NO. OF PAGES 36 plus 7 attachment paans !i ORIGINATOR: /WA f. Lj7r//% DATE: ?/@/4(e VERIFER: t._ Ms% n 1;. DATE: 7/di/J (o APPROVER: hW TMor/1MO 4v /.70 r ah I2. bA.d Cr DATE: ?/M/%
REVISION: DESCRIPTION:
TOTAL NO. OF PAGES ORIGINATOR: DATE:
VERIFER: DATE:
APPROVER: DATE:
PAGE i CONT ON 2 1 Effective 2/94 l
l JUL 311996 i
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l TABLE OF CONTENTS 1.0 OBJECTIVE OF CALCULATION. . . . . . .3 j 2,0 ACCEPTANCE CRITERIA., . . . . . . . . . . . .. . . .3 3.0 ABBREVIATIONS.. . .3 1 I
4.0 REFERENCES
. . .. . . .. 4 l 4.1. General ... . . .
. . . . . . .4 l 1 4.2. Drawings... . . . . .5 j
- 4.3. Procedures.. . .. . . . ..5 4.4. Vendor.. .. . . . . . .. . .6 '
i 4.5. Calculations . .. . . . . . .. .7 5.0 ASSUMPTIONS. . .. . . .7 6.0 DESIGN INPUTS. . . . . . . . 10 6.1. Loop Definitions.. . . . . . . . , 10
, 6.2. High Pressurizer Pressure Reactor Trip Basis .. . . . . . .. . . 10 i 6.3. Low Pressurizer Pressure Reactor Trip Basis . . . . . . . .. . . . . 10 7.0 METHOD AND EQUATION
SUMMARY
.. . . . . . 11 7.1 Block Diagrams.. . . . . . . . . . . . . . .11 7.2. Component Models and Tag Numbers.. . . . . .12 7.3. Environment . . . . . . . . . .. . 12 j 7,4. Sources of Uncertainty.. .. . . .. . . . 13 7.5.. Equation Summary . . . . . . .. . . 14 {
l 7.6. Setpoint Evaluation. .. .. . . . ..... . . 15 8.0 BODY OF CALCULATION.. .. . .. ... . 16 8.1. Device Uncertainties.. . . .. . .. .. . . . 16 i 8.2. Device Uncertainty Notes . .. . . . .. . . 17 j 8.3. Total Loop Error - Normal Conditions... .. .. . . .29 8.4. Total Loop Error - Accident Conditions.... . . . . . . . .... . .. .30 8.5. High Pressurizer Pressure Reactor Trip Setpoint Evaluation.. . . . . . . . . . 31
{ 8.6. Low Pressurizer Pressure Reactor Trip Setpoint Evaluation.. . . . .. . .33 l
9.0 CONCLUSION
S.. . . . .. . . .35 I 10.0 IMPACT ON PLANT DOCUMENTS.. . . .. .. .. .. . .35 j l1.0 OPEN ITEMS.. . .. . . . . .36 12.0 ATTACHMENTS.. .. . . . . . . .. . . . . .36 T
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.I Low and High nessunzer Pressure Reactor Trip Instrurnent Uncertaintv/Setpoint Calculation - PBNP Setpoint Ventication Progr:un
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1.0 OBJECTIVE OF CALCULATION The objective of this calculation is to determine the statistical uncertainty of the Pressurizer I
Pressure instrumentation channels for the Low and High Pressurizer Pressure Reactor Trip l setpoints and evaluate the proposed setpoint changes. ,
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2.0 ACCEPTANCE CRITERIA This calculation will be considered acceptable if the total loop errors are calculated m accordance with the methodology in Ref. G.I. and the results are compared to existing plant l documents. 1 3.0 ABBREVIATIONS l 3.1. AL Analytied Limit 3.2. ATSP -Existag Actual Trip Setpoint 8 I 3,3. AV Allowable Value 3.4. DBE Design Basis Event 3.5. FSAR Final Safety Analysis Report - ,
3.6. HELB High Energy Line Break l 3.7. LBLOCA Large Break Loss of Coolant Accident -
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3.8. LOCA Loss of Coolant Accident ,
j 3.9. M&TE Measurement and Test Equipment 3.10. NSSS Nuclear Steam System Supplier ,
l 3.11. NTSP NominalTrip Setpoint l 3.12. OBE Operating Basis Earthquake l
3.13. PORV Power Operatel Relief Valve 3.14. Press Pressure 3.'15. Pzr Pressurizer 3.16. RAD Radiation Accumulated Dose 3.17. RCS Reactor Coolant System l 3.18 RE Rack Error 3.19. RPS Reactor Protection System ,7 gY 3.20. SELOCA Small Break Loss of Coolant Accident g l 3.21. SP/E Sensor / Process Error 3.22. SRSS Square Root of the Sum of the Squares 3.23. Tech Spec Technical Specifications 3.24. TID TotalIntegrated Dose 3.25. TLE Total Loop Error .
3.26. TS Tecimical Specifications 3.27. URL Upper Range Limit 3.28. Xmtr Transmitter
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I Low and Hiah Pressunzer Pressure Reactor Tnp Instrument Uncertamty/Setpomt Calculation - PBNP Setpoint Venfication Program PAGE I 67) 7/W4(a 2*6(M 1/3 e /% VECTRA JOB NO 0087-00033.303 O KLD 1/l"//% >W 4/17/% Technologhs CALC NO 3 OF 36 REV BY DATE CHECKED DATE loe. PBNP-IC-12 i
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4.0 REFERENCES
Please note that it is the responsibility of the individual revising any of the references listed below to evaluate if the change being made affects this n' tlation.
4.1. General ,
G.I. Point Beach Nuclear Plant Design Guideline DG-101, Instrument Setpoint Methodology, Rev.1 G.2. Point Beach Nuclear Plant Technical Speci6 cations, Section 15.2.3.1.B. and Table 15.3.5-2, latest revisions per PBNP Units I and 2 Technical Specifications " List of Effective Pages, Rev. 96 G.3. Point Beach Final Safety Analysis Report, Sections 7.2.2 and 7.2.3., dated 6/92 G.4. PBNP Setpoint Document STPT 1.4, " Pressurizer Pressure and Level", Rev. 3 G.5. PBNP CHAMPS Records G.6. ASNE Steam Tables, Fourth Edition .
G.7. Equipment Quali6 cation Summary Sheets, Point Beach Nucl'e ar Plant Units 1 and 2, Pages 7.5.A through 7.5.D, all dated 8/14/95 G.8. Point Beach Units 1 and 2 Reactor Protection & ESF Activation Analytical Limit Verification Information Revision, Westinghouse Letter WEP-94-525, dated 1/28/94 G.9. PBNP Condition Report CR 95-109 Evaluation G.10. Wisconsin Electric PBNP DBD-27, Reactor Protection System Design Basis Document, Rev. O G.11. Summary of RPS and ESFAS Functions Actuated, pages 54 and 55 of SECL 95-064, Revision 0, Fax from Rick Kohrt of Wisconsin Electric, dated 7/3/96 (Attachment A)
G.12. PBNP Design Basis Validation Comment Sheet For Attribute Item No. 33, DBD- l 27, Reactor Protection System, Fax from Brian Boysen of Wisconsin Electnc, i dated 8/30/95 (Attachment B) i Low and Hiah Pressuruer Pressure Reactor Trip Instrument Uncertanuty/Setpost Calculation - PBNP Setpoint Venfication Program j _ .
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l G.13. Instrument Society of America Recommended Practice ISA-RP67.04, Part II, Methodologies for the Determination of Setpoints for Nuclear Safety-Related Instrumentation,1994 4.2. Drawings D. I . Foxboro Dwg. BD-10 (Unit 1), Rev. 3, Instrument Block Diagram Pressurizer Pressure Control l
D.2. Foxboro Dwg. BD-10 (Unit 2), Rev. 3, Instrument Block Diagram - Instrument Reactor Protection System, Pressurizer Pressure Control
- D.3. Stone & Webster Dwg. 13754.22-SK-1182, Instrument Installation 2LT-433, 2PT-449 & 2PT-493, Rev. 4 D.4. Stone & Webster Dwg. 13754.22-SK-1198, Instrument Installation 2LT-426, 2PT-420 & 2PT-429, Sheet 1 of 2, Rev. 4 D.5. Stone & Webster Dwg. 13754.22-SK-1198, Instrument Tubing Layout 2LT-426
& 2PT-429, Sheet 2 of 2, Rev. 4 D 6. Stone & Webster Dwg. 13754.22-SK-1201, Instrument Installation 2LT-427 &
2PT-430, Sheet 1 of 2, Rev. 5 D.7. Stone & Webster Dwg. W54.22-SK-1201, Instrument Tubing Layout 2LT-427
& 2PT-430, Sheet 7 9 4 Rev. 5 D.8. Stone & Webster Dwg. 13754.22-SK-1204, Instrument Installation 2LT-428 &
2PT-431, Sheet 1 of 2, Rev. 5 D.9. Stone & Webster Dwg. 13754.22-SK-1204, Instrument Tubing Layout 2LT-428, 2PT-431,2LT-433,2PT-449 & 2PT-493, Sheet 2 of 2, Rev. 5 4.3. Procedures P.I. 11CP-02.001, Rev.1, " Reactor Protection and Emergency Safety Features Analog Quarterly Surveillance Test" P.2. 11CP-02.001BL-1, Rev. 7, " Reactor Protection and Emergency Safety Features Blue Channel Analog Quarterly Surveillance Test" L
Low and Hish Pressunzer Pressure Reactor Trip Insuunent Uncertamty/Wint Calculation - PBNP Setpoint Venfication Program WW 7he/M PAGE I (L-T) W/% VECTRA JOB NO 0087-00033.303 0 KLb '7/11/96 /W 7/I'7h6 T - ' ' :--
CALC NO 5 DAH Inc. PBNP-IC-12 OF 36 REV BY DAH CHEtKED i
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P.3. IICP-02.00lRD-1, Rev. 6, " Reactor Protection and Emergency Safety Features Red Channel Analog Quarterly Surveillance Test" P.4. IICP-02.00lWH-1, Rev. 6, " Reactor Protection and Emergency Safety Features White Chanrlel Analog Quarterly Surveillance Test" P.S. IICP-02.001YL-1, Rev. 5, " Reactor Protection and Emergency Safety Features Yellow Channel Analog Quarterly Surveillance Test" P.6. 2ICP-02.001, Rev.1, " Reactor Protection and Emergency Safety Features Analog Quarterly Surveillance Test" P.7. 21CP-02.001BL-1, Rev. 7, " Reactor Protection and Emergency Safety Features Blue Channel Analog Quarterly Surveillance Test" P.8. 21CP-02.00lRD-1, Rev. 6, " Reactor Protection and Emergency Safety Features Red Channel Analog Quarterly Surveillance Test" P.9. 2ICP-02.001WH-1, Rev. 5, " Reactor Protection and Emergency Safety Features White Channel Analog Quarterly Surveillance Test" P.10. 2ICP-02.00lYL-1, Rev. 5, " Reactor Protection and Emergency Safety Features Yellow Channel Analog Quarterly Surveillance Test" P.11. ICP 4.1H U1, Rev. 3, " Reactor Protection and Safeguards Analog Racks Pressurizer Pressure" P.12. ICP 4.1H U2, Rev. 2, " Reactor Protection and Safeguards Analog Racks Pressurizer Pressure" P.13. IICP-0 4.004-1, Rev. 2, " Event V Test Pressure, RCS Wide Range, RC Hot Leg Pressurizer & RV Head Vent Pressure Instruments Outage Calibration" P.14. 2ICP-04.004-1, Rev.1, " Event V Test Pressure, RCS Wide Range, RC Hot Leg Pressurizer & RV Head Vent Pressure Instruments Outage Calibration" -
4.4. Vendor V.I. Foxboro Composite Books, PBNP Control No. 00623 A1, Rev.12 and PBNP Control No. 00623 A, Rev. 09 Low and High Pressunzer Pressure Reactor Tnp Instrument Uncertamtv/Setpoint Calculation - PBNP Setpomt VenGcation Program i FL7) 77t/6" h '7 / 3c/_9(c VECTRA JOB NO OO87 00033.303 PAGE T> CALC NO 6 O KL'D 7/17'/96' \EU ' 7/17f% 'tr REV BY DA'E CHECKED DATE Inc. PBNP-IC-12 OF 36 i
1 V.2. Foxboro EQ Transmitters Manual, PBNP Control No. 00432, Rev.16 l V.3. Calculation Information Regarding N-El1GM and N-E13GM Transmitters, Fax j
from David Ringland of Foxboro, dated 5/5/94 (Attachment C) l V.4. Foxboro Qualification Test Report of N-E10 Series Transmitters for Class IE Qualification, Q0AACI1, Rev. A 4.5. Calculations C.I. I&C Calculation Book, Section 2.5, Unit I and 2, Pressurizer Pressure High and Low Setpoints, dated 10/28/86 C.2. I&C Calculation Book, Section 4.1, Pressurizer Pressure Transmitters, dated 12/8/75; Water Leg Correction - PT429, 430,431 and 449 (Pressurizer Pressure),
Transmitter Scaling, dated 4/25/88; and Unit 1 Pressurizer Level Transmitter Scaling, dated 8/15/92 C.3. Foxboro N-EllGM Transmitter Drift Calculation, VECTRA Calculation No.
PBNP-IC-13, Rev. 0 .
C.4. Foxboro 66RC-OLA Lead / Lag Module Drift Calculation, VECTRA Calculation No. PBNP-IC-10, Rev. 0 i
C.S. Foxboro 63U-AC Bistable Drift Calculation, VECTRA Calculation No. PBNP-IC-11,Rev 0 l
5.0 ASSUMPTIONS S.b, //
5.1. Please refer to SectionO12.0 for the open items that are to be resolved before y(, t ca!alation is considered to be complete.
5.2. From the PBNP Setpoint Methodology (Ref. G.I.), the statistically derived as- ,
found/as-left drift value includes the effects of M&TE used in past calibrations. To '
maintain the validity of this value, it is assumed the M&TE used to perform future calibrations will be of equivalent accuracy to the M&TE used in the past calibrations on which the as-found/as-left drift data is based.
Low and High Pressurtzer Pressure Reactor Tnp Instrument Uncertatntv/Setpoint Calculation - PBNP Setpoint Venfication Program 1he /56 PAGE
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5.3. It is assumed that the containment ambient temperature at the time of transmitter calibration is approximately 68'F, based on the transmitter calibration information provided in the existing Pressurizer Pressure transmitter scaling calculation (Ref. C.2.). l The minimum temperature in containment during plant operation is assumed to be no less than 60 F.
5.4. According to the PBNP Setpoint Methodology (Ref. G.I.), only mechanical devices experience a permanent output shift due to a seismic event. Ref. G.I. states that the i environmental allowance used will be the uncertainty due to the larger of either the l post accident harsh environment or the seismic effects on the loop devices since the !
events are considered to be independent of one another. Since this calculation only addresses Reactor Trips and not post accident conditions, the seismic effect will not be ;
considered. Ref. G.I. also states that seismic events do not cause safety systems to !
fail and assumes that for seismic events greater than an Operating Basis Earthquake, instrumentation will be recalibrated prior to any subsequent accident; thus, negating 1 any permanent shift that may have occurred due to the seismic event.
5.5. From Ref. G.9., the results of testing performed at PBNP has shown that changing the Plant Process Computer from normal to standby condition causes a change in the instmment bus inverter output voltage. The testing has shown tilat the voltage change (
has an effect only on the Foxboro H-Line current to current converters; therefore, it is assumed the inverter output voltage variations do not affect the other H-Line modules, :
(i.e., bistables or lead / lag modules). l t
5.6. According to Refs. G.3. and G.12., in the Accident Analysis, pressure drops to the l Low Pressurizer Pressure Reactor Trip setpoint of 1775 psia (Analytical Limit) within .
5.6 seconds into the SBLOCA transient. Also from Ref. G.12., a preliminary evaluation of the containment conditions for accident environments was performed by ;
the DBD team using FSAR Figures 14.3.4-8,14 and 15. The containment conditions l at =5.6 seconds were found to be 45 psia and 265'F. However, the FSAR figures represent a Large Break LOCA, and are considered to bound the SBLOCA :
environmental conditions. Therefore, the containment conditions resulting from a ,
LOCA/HELB event for a duration of 5.6 seconds (45 psia and 265'F) are considered to be more severe than the environmental conditions resulting from a SBLOCA event of the same duration. As a result, the maximum temperature which the Pressurizer Pressure transmitters will be subject to is expected to be lower than the peak temperature specified for a LOCA/HELB lasting 5.6 seconds.
Low and Hish nwiins Pressure Reactor Trip Instrument Uncertaintv/Setpoint Calculation - PBNP Setpomt Venfication Program 1 (CD ?/Ac/%56f9r- 7/be/x VECTRA JOB NO OO87 00033.303 PAGE O KLD 7/17/96 A ' KAJ 7/11/% Techselegess CALC NO 8 REV BY DATE CRECKED DATE lar- PBNP IC-12 OF 36
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A temperature effect is provided by the transmitter manufacturer for temperatures up to 250 F. For the purposes of determining the temperature effect for the transmitter, the peak temperature of 250 F is assumed to adequately bound the maximum temperature experienced by the transmitters after a SBLOCA lasting 5.6 seconds. i Furthermore, the plant setpoint will be higher than the Analytical Limit; therefore, the trip will occur sooner than the 5.6 second duration determined in the Accident Analysis. Since the temperature increases as the event duration increases (Fig.14.3.4-15, Ref. G.3.), a shorter duration will result in a lower peak temperature, thereby providing additionaljustification for assuming a lower temperature than 265 F.
5.7. Based on Figure 14.3.4-15 of Ref. G.3., for a LBLOCA, the accumulated gamma dose level is less than 10' RADS until 12 minutes after release. Furthermore, the radiation levels prior to 6 minutes after release are too low to be shown in Figure 14.3.4-15 of Ref. G.3. (i.e., off scale). For the Low Pressurizer Pressure Reactor Trip, a SBLOCA would result in a much lower dose rate than a LBLOCA, and the duration of exposure '
is only 5.6 seconds until the setpoint is reached. Due to the shon duration of possible exposure (5.6 seconds), and since the LBLOCA conditions bound the conditions expected for a SBLOCA, the radiation effect on the transmitters is assumed to be negligible.
5.8. For the Low Pressurizer Pressure Reactor Trip, the duration during which the cabling and splices are subject to higher temperatures than normalis only 5.6 seconds (see !
Assumption 5.6.). It is assumed that increased temperature exposure for a duration of 5.6 seconds is not long enough to affect the conductive propenies of the cabling and splices; therefore, the effects due to insulation degradation are considered to be negligible.
5.9. The values shown in this calculation are linked to a Microsoft Excel spreadsheet file in which the calculated values are determined. The cells are linked such that successively calculated values are dependent on the previously calculated cells, which may cause intermediate values to appear incorrectly rounded. However, the final results are correct with respect to the initial input values.
Low and High Pressmust Pressure Reactor Trip Instrument Uncettaintv/Setpoint Calculation - PBNP Setpoint Venfication Program PAGE I /4T ;?/3C/9/, 4g.- 7/h /_04 YEC' IRA JOB NO 0087 00033.303
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i 6.0 DESIGN INPUTS
) 6.1. Loop Definitions t
i ~ The loop components addressed in this calculation were identified in References D.I.
! through D.9., P.1. through P.14., and V.I. through V.2., and are identified in Sections 7.1 and 7.2. The loops are calibrated from 1700 to 2500 psig (Refs. C.2., P.13., and P.14.).
j 6.2. High Pressurizer Pressure Reactor Trip Basis
{ From Section 7.2.2 of Ret G.3., "The purpose of this circuit is to limit the range of
{ required protection required from the ovenemperature AT trip to protect against reactor coolant system overpressure." Ref G.8. shows the Analytical Limit used in the Safety i Analysis applicable to operation prior to the Unit 2 Steam Generator replacement is 2425 j psia (equivalent to 2410 psig). From Section 15.2.3.1.B.2 of Rc0 G.2., the applicable l Technical Specifications require the High Pressurizer Pressure Reactor Trip setting to be s; l 2385 psig. From Ref G.4., STPT 1.4 shows the associated plant High Pressurizer l Pressure reactor Trip setpoint is 2365 psig. -
l From Ret G.11., the Accident Analyses for the new Unit 2 Steam Generators resulted in j changing the Analytical Limit to 2250 psia (equivalent to 2235 psig) for 2000 psia
{ operation, and back to 2425 psia (equivalent to 2410 psig) for 2250 psia operation.
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6.3. Low Pressurizer Pressure Reactor Trip Basis 4
1 From Section 7.2.2 of Ref G.3., "The purpose of this circuit is to protect against l
i excessive core steam voids." Per Ref G.8., the Analytical Limit used in the Safety Analysis applicable to operation prior to the Unit 2 steam generator replacement is 1775 psia (equivalent to 1760 psig). From Section 15.2.3.1.B.3 of Ref. G.2., the applicable l
i Technical Specifications require the Low Pressurizer Pressure Reactor Trip setting to be 2 i 1790 psig for operation at 2000 psia primary pressure. From Ref G.4., STPT 1.4 shows j j
- the associated plant Low Pzr Press setpoint is 1810 psig for 2000 psia operation. l From Ref G.11., the Accident Analyses for the new Unit 2 Steam Generators did not l i 4
result in changes to the Analytical Limits for the Low Pressurizer Pressure Reactor Trip i provided in Ref. G.8.
I Low and Hish Pressunzer Pressure Reactor Trip li-Lw..u.t i Uncertamtv/Setpoint Calculation - PBNP Setpoint Venfication Program i '?/&/@h '1/ de/n VECTRA JOB NO 0087 00033.303 PAGE 1 KkT) j 0 KLD 7/17/96 f \DJ ' t/11/% T*'f^ CALC NO 10 REV BY DAE CHECKED DATE lac. PBNP.IC-12 OF 36 I
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7.0 METHOD AND EQUATION
SUMMARY
7.1. Block Diagrams The block diagrams shown below represent the components for the Pressurizer Pressure instrument loops and apply to both units. Note that the P-449 loops only consist of the Low Pzr Press Reactor Trip.
lEgh 5
Bistable > Preswber Preswe Reactor Trip Preswaar 0 hw 0 Preswa Low Lea 4 tag Bistable lA Preswner J Pres-.
Reactor Trip Figure 1. Loops P-429,430 and 431 Pwr Supply law Preswho' O !.maalag Bistable A Presurser he Presws Presws Reactor Trip Figure 2. Loop P-449 Low and High Pressunzer Pressure Reactor Trip Iratrument Uncertamtv/Setpoint Calculation - PBNP Setpoint Venfication Program I (4T) ?/5C/4(.f M "7/.3c/SCr VECTRA JOB N0 0087 00033.303 PAGE O KLD 7/17/96 )\ IQlJ #
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7.2. Component Models and Tag Numbers The following table lists the manufacturer, model number, and tag number for each component applicable to the Low and High Reactor Trips shown in Figures 1 and 2 for each of the four loops (these tag numbers are applicable to both Units I and 2).
Table 1. Pressurizer Pressure Instruments Component Model Red White Blue Yellow Channel Channel' Channel Channel Pzr Press Foxboro Transmitter N-El1GM PT-429 PT-430 PT-431 PT-449 Foxboro Power Supply 610AC-O PQ-429 PQ-430 PQ-431 PQ-449 Foxboro Lead / Lag Unit 66RC-OLA PM-429B PM-430C PM-431C PM-449B Low Pzr Press Foxboro Reactor Trip 63U-AC-Bistable OHAA-F PC-429E PC-430H PC-431J PC-449A High Pzr Press Foxboro Reactor Trip 63U-AC- .
Bistable OHBA-F PC-429A PC-430A PC-431 A N/A 7.3. Environment From Refs. G.8. and G.11., the following trips provide protection against the listed accidents and the limits assumed in the safety analyses shown below:
Analytical Limit Sinnal Actuated Event Descriotion (2000 osia coeration) ,
Low Pzr Press Rxt Trip Dropped Rod 1610 psia SBLOCA 1775 psia l High Pu PressRxt Trip Loss of Load 2250 psia Low and High Pressuruer Pressure Reacter Trip Instrument Uncertamtv/Setpomt Calculation - PBNP Setpoint Venfication Program
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The maximum ambient temperature in containment during normal operation is 105 F (Ref.
G.3., Section 7.2.3.). From Assumption 5.3., the containment ambient temperature during calibration is considered to be 68*F and the minimum temperature during normal operation is considered to be 60'F.
The Loss of Load or Dropped Rod would not cause a adverse environment in containment prior to the reactor trip.actuations. A Small Break Loss of Coolant Accident could cause adverse conditions in containment. However, from Assumption 5.6., the maximum temperature at the time of the Low Pressurizer Pressure Reactor Trip is assumed to be 250 F. From Assumption 5.7., the radiation effects for the Low Pressurizer Pressure Reactor Trip are considered to be negligible.
The rack components are located in the control room, which is maintained at 75 10 F and is subject to a mild environment under all plant conditions (Ref. G.3., Section 7.2.3).
I l 7.4. Sources of Uncertainty 4
l The drift values were calculated using as-found/as-left values from plant calibration data, e as described in Ref. G.1., which states that the drift values calculated from as-found/as-left l ' string calibration data include the error effects under normal conditions of drift, accuracy, power supply, plant vibration, calibration temperature, normal radiation, normal humidity, M&TE used for calibration, and instrument readability. Therefore, these error effects will not be considered separately when a drift value is determined from as-found/as-left .
j calibration data. The as-found/as-left data is taken from the surveillance procedures
- performing static verification of the loops and the cale dation of uncertainties will be
' performed with the loops in steady-state condition. The device uncertainties to be considered for normal and adverse environmental conditions include the following (Ref.
- G. I.)
f Sensor Accuracy (sa) i Sensor Drift (sd) i Sensor M&TE (sm) l Sensor Setting Tolerance (sv) j Sensor Power Supply Effect (sp)
! Sensor Temperature Effect (st n and st a)
- Sensor Humidity Effect (sh n and sh a)
Sensor Radiation Effect (sr n and sr a)
Sensor Seismic Effect (ss n and ss a) f Sensor Static Pressure Effect (spe n and spe a) i Sensor Overpressure Effect (ope n and ope a)
Low and Miah Pressuruer Pressure Reactor Trip Instrument Uncertaintv/Setpoint Calculation - PBNP Setpoint Verification Program l u ,
j i /EP W3f/% T%fc7 f7 Bo/D(, VECTRA JOB NO 0087-00033 303 PAGE O KLD 4/17/% ~TM '7/17/96 T- ' ' - C -- CALC NO 13 OF 36 REV BY DATE CHECRED DNE lac. PBNP-IC-12 m -g, . - - - - - --
1 i
Bistable Accuracy (bistAC a)
- Bistable Drift (bistACd) d Bistable M&TE (bistAC m)
- Bistable Setting Tolerance (bistAC V)
Bistable Power Supply Effect (bistACP)
Bistable Temperature Effect (bistACt )
Bistable Humidity Effect (bistACh)
- Bistable Radiation Effect (bistACf)
- Bistable Seismic Effect (bistAC8)
I Lead / Lag Accuracy (UL a)
- Lead / Lag Drift (UL d)
Lead / Lag Setting Tolerance (UL v)
- Lead / Lag Power Supply Effect (UL p)
- Lead / Lag Temperature Effect (UL t)
! Lead / Lag Humidity Effect (UL h) i Lead / Lag Radiation Effect - (UL r) 1 Lead / Lag Seismic Effect (UL s) i Process Considerations (random, independent) (pc n and pc a) -
l Process Considerations (positive bias) (pc pos n and pc pos a) i Process Considerations (negative bias) (pc neg n and pc neg a) 4 i
- 7.5. Equation Summary The total loop error for setpoints and indication is determined in scordance with the i requirements of Ref. G.I. This methodology uses the square rc ' . of the sum of the i squares (SRSS) method to combine random and independent errors, and algebraic
! addition of non-random or bias errors. From Ref. G.I., the general equation for
- combining errors to calculate total loop error is
S 2 2 TLE = 1 dA + B + (C + D)2 iElXl + ZY - IZ J where: A, B = Random and independent uncertainty terms i C, D = Random and dependent uncertainty terms X =
Non-random (urMown direction)
Y = Non-random (positive biases)
- Z = Non-random (negative biases)
Low and High Pressunzer Pressure Reactor Trip Instrument Uncertainty /Setpoint Calculation . PBNP Setpomt Venfication Program l
FAGE
) 1 RG 7/ 84/#- h 4/3c>/5a VECTRA JOB NO 0087 00033.303 O KLD 7/17196-V \Md1 / *//11/96 TV ' .t CALC NO 14 REV BY DATE CHECKED DATE tae. PBNP-IC 12 OF 36 a
1 i
The general equation for total instrument loop error is (Ref. G. I.):
TLE = JA + D + M + V + P + T + H + R + S + SPE + 0PE + PC + B* - B' l l
A = Accuracy Allowance =
(a i2 + a22 ...+ a,2 )
]
l D = Drift Allowance =
(d i2 + d 22 ...+ d,2 ) <
M = M&TE Allowance =
(m,2 + m22 ...+ m,2 )
V =
Setting Tolerance Allowance = (v,2 + y,2 ,,,+ y,2 ) 1 l
P = Power Supply Allowance =
(p,2+p,2.,,+p,2)
T = Temperature Allowance = (t 2 + t2 ' ..+ t,2 )
i H = Humidity Allowance = (h,2 + h 22 . ..+ h,2 ) j l
R = Radiation Allowance = (r 2 + r 22 ...+ r,2 )
S = Seismic Allowance =
(si 2+ s22 . ..+ s,2 ) . j i
SPE = Static Pressure Allowance = (spe i 2 + spe2 2...+ spe,2 ) l l
OPE = Over Pressure Allowance = (ope 2 + ope2 2 + 0P8.2 )
PC = Process Considerations = (pc,2 + pc22 ...+ pc,2 ) l l
B+ = Positive Bias Errors =
+ (pey,2 + pcm22,,,+pe,2)
B- = Negative Bias Errors =
- (pe,,,,2 + pe,,,,2 ,,,+ pe,2) 7.6. Setpoint Evaluation
" The setpoint evaluations are performed in accordance with Ref. G.I. In evaluating i setpoint requirements, a thorough understanding of the relationship between calculated i Nominal Trip Setpoint (NTSP), Actual Trip Setpoint (ATSP), Tech Spec value (TS),
Allowable Value (AV), and Analytical Limit (AL) is required. Refer to Ref. G.I. for a detailed discussion of setpoint determination.
1 Low and High P-u.s Pressure Reactor Trip Instmment Uncertamtv/Setpoint Calculation - PBNP Setpomt Vertfication Program PAGE
! 1 RfX) f/W4f(o h 7/a /o f. VECTRA JOB NO 0087-00033.303 O KLD 7/17/96 Al(41 /7/19/% TV ' .'- CALC NO 15 i
DATI! CHECKED DATE lac. PBNP-IC-12 OF 36 REV BY
8.0 BODY OF CALCULATION 8.1. Device Uncertainties Parameter Uncertainty (% scan)
Sensor Accuracy -
See Note 1.
(sa) 0.224 %
Sensor Drift See Note 2.
(sd) 0.895 %
Sensor M&TE See Note 3.
Sensor Setting Tolerance See Note 4.
(sv) 0.500 %
Sensor Power Supply Effect See Note 5.
Sensor Temperature Effect See Note 6.
Normal Environmental Conditions (stn) ,
1.388 %
Accident Environmental Conditions (st a) 7.284 %
Sensor Humidity Effect See Note 7.
Sensor Radiation Effect See Note 8.
Sensor Seismic Effect See Note 9.
Sensor Static Pressure Effect See Note 10.
Sensor Overpressure Effect See Note 11.
Rack Accuracy See Note 12.
Rack Drift See Note 13.
Bistable 63U-AC Drift (bistAc d) t 0.222 %
Lead / Lag Drift (IJL d) 0.083 %
Rack M&TE See Note 14.
Low and High Pressuruer Pressure Reactor Trip Instrument Uncertamtv/Setpoint Calculation PBNP Setpoint Venfication Program 1 t/,J) Y/tt/f(c h 7/re /M VECTRA JOB NO 0087-00033.303 PAGE 0 KLD 7/17/96 -AM 't/11/96 TecW CALC NO 16 REV BY DATE CHECKED DATE lac. PBNP-IC 12 OF 36
I Parameter Uncenainty (% scan)
Rack Setting Tolerance See Note 15.
Lead / Lag Setting Tolerance (UL v) 0.500 %
Rack Power Supply Effect See Note 16.
Rack Temperature Effect See Note 17.
Rack Humidity Effect See Note 18.
Rack Radiation Effect See Note 19.
Rack Seismic Effect See Note 20.
Process Considerations See Note 21.
Maximum Head Correction Error - Normal (penPos) + 0.026 %
Minimum Head Correction Error - Normal (penneg) -0.069 %
Maximum Head Correction Error - Accident (pca Pos) - + 0.026 %
Minimum Head Correction Error - Accident (pca neg) - -0.130 %
Accident Insulation Resistance See Note 22. j i
l l
8.2. Device Uncertainty Notes !
Note 1. - Sensor Accuracy l 1
The repeatability and hysteresis effects of the transmitter will be included as separate l accuracy terms that are not necessarily included in the as-found/as-left drift analysis results. Vendor information (Ref. V.2.) provides the following specifications:
s rep = 0.100 % span s hyst = t 0.200 % span Low and High Pressunzer Pressure Reactor Trip Instrument Uncenamtv/Setpomt Calculation . PBNP Setpoint Venfication Program 1 JG T) 7/b/#(cW ch
/ -f /ho/3( ,
VECTRA JOB NO 0087-00033.303 PAGE O KLD 7/17/% *-_) \K41 li,;7/% Techsele@s CALC NO 17 REV BY DATE CHECKED DATE Inc. PBNP4C.12 OF 36
l l
Following the PBNP Setpoint Methodology (Ref. G.I.), repeatability and hysteresis I effects are combined as random and independent terms using the SRSS method as follows:
sa = [(s rep)2 + (s hyst)2 ]!/2 sa =
[(0.100 %)2 + (0.200 %)2 ) it2 s- = t 0.224 % span Note 2. - Sensor Drift l
1 To be conservative, the largest recommended drift value for the Pressurizer Pressure transmitters calculated in Ref. C.3. will be used.
sd = t 0.895 % span Note 3. - Sensor M&TE From the PBNP Setpoint Methodology (Ref. G.I.), the M&TE effect is included in the as-found/as-left drift data and is therefore included in the statistical drift value for the Pressurizer Pressure transmitters calculated in Ref. C.3. Therefore, the M&TE error for the transmitters is considered to be zero.
sm = 0.000 % span Note 4. - Sensor Setting Tolerance The setting tolerance for the Pressurizer Pressure transmitters from the calibration procedures (Refs. P.13. and P.14.) is:
sv = i 0.500 % span Note 5. - Sensor Power Supply Effect From Ref. V.I., the AC supply voltage regulation requirement for the Foxboro 610-AC-O loop power supply is 10%. Per Ref. G.9., the instrument bus voltage varied by no more than 2.5 Vrms during instrument bus load testing. This variation is within the 10 %
voltage variation requirement for the loop power supply. In addition, this voltage variation was proven to only have an effect on the current to current converters (see Assumption 5.5.).
Low and High N us a.r Pressure Reactor Trip Instrument
,s .
Uncertaintv/Setpoint Calculation - PBNP Setpoint Venfication Program 1 4:77 7/kV % 8Tb4N - ~1/h/% VECTRA JOB NO 0087-00033.303 PAGE ,
O KLD 7/17/96d \ FAD 1/17f96 Technolog6ss CALC NO 18 l REV BY DATE CHECKED DATE lac. PBNP-IC-12 OF 36 l f
l
i
) Vendor information for the N-El1GM pressure transmitters (Ref. V.2.) shows that the 10 to 50 mA output model can accept a supply voltage between 60 and 95 Vdc for non-LOCA!HELB applications and between approximately 75 and 95 Vdc for LOCA/HELB 1 applications; however a power supply effect is not specified. Vendor information for the i Foxboro 610A loop power supply (Ref. V.I.) states that the power supply is designed for
! force-balance transmitters and shows the nominal 80 Vdc output can vary from 84 Vdc 2 Vdc at 10 mA to 76 Vdc 12 Vdc at 50 mA. For normal conditions, the voltage supply requirement of the transmitter is met. For the Low Pressurizer Pressure Reactor Trip, the I
power supply output voltage meets the transmitter requirements for the point ofinterest.
I In the event the loop power supply output drops to 74 Vdc at the 50 mA level during .
- accident conditions, any effect that may be seen is not considered to have an effect on the
- Low Pressurizer Pressure Reactor Trip. ,
l 4
sp = 0.000 % span t I
Note 6. - Sensor Temperature Effects l The Pressurizer Pressure transmitters are Foxboro N-EllGM transmitters with an upper ;
span limit of 2000 psi (Refs. G.7. and V.2.). The calibration range is 800 psig, which is
! 40% of the upper span limit. The transmitters have a vendor speci6ed temperature effect
- that must be considered. From information provided by Foxboro (Ref. V.3.), transmitters
{
calibrated between 20% and 50% of the upper span limit may have a zero shift of t 2.500
% of span per 100*F for an operating ambient temperature range of 32 *F to 180 'F and a j 5.000 % of span zero shift per 170'F for a temperature range of 80*F to 250 F. The span effect, regardless of span setting or temperature range is i 1.250 % of span per 100* ,
F change. The zero shift and span change effects are conservatively considered to be dependent random errors, and per the methodology in Ref. G.I., will be added 1 algebraically (straight sum) to determine the overall normal operating temperature effect
, on the transmitters.
I For normal operating conditions, only the temperature change from calibration conditions
- needs to be considered. The calibration values for the transmitter (Ref. C.2.) were
- calculated using standard reference conditions (68'F, atmospheric pressure). The temperature effect on the transmitter at the high end of the normal operating temperature range (105 *F per Ref. G.3.) is calculated using the maximum normal temperature l
j variation of 37*F from calibration.
! xmtr t zero n = (20-50% effect n)(temp change n/100*F) 4
= (2.500 % span)(37'F/100*F) i = 0.925 % span Low and High Pressunzer Pressure Reactor Trip Instrument Uncertaantv/Setpoint Calculation - PBNP Setpoint Venfication Program l
- 1 4 67) W%'/%h r1/3*l00- VECTRA JOB NO 0087-00033.303 PAGE 0 KLD 7/17/% Plod ) M/17/% TM 'f- CALC NO 19 REV BY DATE CHECKED DATE lac. PBNP-IC.12 OF 36
- (
4 i
4 i xmtr t span n = - t (span shift)(temp change n/100*F) j = (1.250 % span)(37'F/100 F) i = 0.463 % span ,
4 .
F j stn
= i(xmtr t zero + xmtr t span) l l = (0.925 % span + 0.463 % span) l l = i 1.388 % span (
l 1 From Assumption 5.6. of this calculation, the maximum temperature at which the Low >
l Pressurizer Pressure Reactor Trip is considered to actuate is 250'F. Based on information l provided by Foxboro (Ref. V.3.), transmitters calibrated between 20% and 50% of the ;
! upper span limit may have a zero shift of 5.000 % of span per 170*F for a temperature !
range of 80*F to 250'F. The span effect, regardless of span setting or temperature range l
! is 1.250 % of span per 100'F change. Taking into account the normal temperature j variation from calibration (68'F) up to 80'F using the speci6ed effects for 32*F to 180*F ;
and the additional error for reaching 250'F before the Low Pressurizer Pressure Reactor ;
Trip occurs, the zero and span effects are calculated as follows:
t {[(20-50% effect nXtemp change w a sov/100'F)] 2 +
1 xmtr t zero a =
- [(20-50% effect aXtemp chang
= t {[(2.500 % spanX12*F/100*F)],e + .
sov wnoy/170 F)] } ;
. [(5.000 % spanX170*F/170'F)],} !
- = t5.009 % span
?
I xmtr t span a = (span shiftXtemp change .a uoy /100*F) l
= (1.250 % spanX182 *F/100*F) j l
- = 2.275 % span !
st, = i(xmtr t zero a + xmtr t span a) l =
! (5.009 % span + 2.275 % span) l = t 7.284 % span i
l Note 7. - Sensor Humidity Effects The normal operating conditions specified by the manufacturer for the transmitters j show there is no operating limit for relative humidity (Ref. V.2.). Information from ,
i the vendor (Ref. V.3.) also states that humidity effects are negligible since the j I transmitters are sealed and are rated for accident conditions.
Low and High N asi Pressure Reactor Trip Instrument .
4 Uncertamty& int Calculation - PBNP Setpoint Venfication Program !
J l XIX) %WM fNV 7ho/04 VECTRA _ JOB NO 0087 00033.303 PAGE O KLD 7/17/% V \lMJ '7/lf/96 T- _ ' . ' -
CALC NO 20 DATE CHECKED DATE lat. PBNP.IC-12 OF 36 REV BY 3
I L . _ . . . , , _.,
sh, = 0.000 % span sh, = t 0.000 % span Note 8. - Sensor Radiation ENect The as-found/as-left drift value for the Pressurizer Pressure transmitters calculated in Ref. C.3. includes the cumulative effect of exposure to radiation. Although the Foxboro N-EllGM transmitters are qualified for adverse environmental conditions, there is a normal radiation effect specified for the Foxboro N-E11GM transmitters.
However, according to vendor supplied information (Ref. V.",.), this radiation effect "can be zero and span adjusted to return to the normal accuracy speci6 cation," or in other words, is corrected by calibration.
I sr, = 0.000 % span Based on Assumption 5.7., the accident radiation effect for the Low Pressurizer Pressure Reactor Trip is considered negligible.
sr, = 0.000 % span Note 9. - Sensor Seismic EKect -
From vendor information (Ref. V.2.), for seismic events less than an OBE, no permanent shift in the instrument input / output relationship occurs.
ss, = t 0.000 % span From Assumption 5.4., seismic events do not cause safety systems to fail and assumes that for seismic events greater than an Operating Basis Earthquake, instrumentation will be recalibrated prior to any subsequent accident; thus, negating any permanent shift that may have occurred due to the seismic event.
ss, = 0.000 % span l
1 Low and High F iunzer Pressure Reactor Trip Instrument Uncertamty/Setpoua Calculation PBNP Setpoint Venfication Program I ECTJ 7/1r/4(f. W '1/ A./6 VECTRA JOB NO 0087 00033 303 PAGE o KLD 7/17/96 ' \ ' kAJ ' 7/lW% T'- '7 CALC NO 21 !
I BY DATE CHECKED DATE tac. PBNP-IC-12 OF 36 REV
I i
Note 10. - Sensor Static Pressure Effect .
- Static pressure effects due to process pressure only apply to differential pressure instruments in direct contact with the process. The process and ambient pressure j effects are considered negligible for the Pressurizer Pressure gage transmitter since j they are scaled units.that have been tested to withstand ambient pressures of 85 psi
! (Refs. V.2. and V.3.) and the maximum ambient pressure expected at the time of the Low Pressurizer Pressure Reactor Trip is 45 psia (see Section 7.3 of this calculation).
In addition, any ambient pressure effects would be in the negative direction (lower j than actual) and would be in the conservative direction with respect to a decreasing
- setpoint (Low Pressurizer Pressure Reactor Trip) . l i
spe, = i 0.000 % span spe, = 0.000 % span 1
t Note 11. - Sensor Overpressure Effect I
The Pressurizer Pressure transmitters are Foxboro Model N-E11GM, Sensor Code E (Ref G.7.), with span limits of 200 to 2000 psig and a maximum overrange limit of 3000 psi (Ref V.2.). The Pressurizer PORVs open at 2335 psig Pressurimer Pressure (Ref I G.2.); therefore, the transmitters would not be subject to overpressure effects.
ope, = 0.000 % span l l
ope, = i 0.000 % span Note 12. - Rack Accuracy From the PBNP Setpoint Methodology (Ref G.I.), when drift error values are derived from as-found/as left calibration data, the resultant drift term includes the effects of accuracy. Linearity, which is a contributor to accuracy, is included in the as-found/as-left drift values calculated in Refs. C.4. and C.S. Also from Ref G.I., when as-found/as-left drift values are used, repeatability and hysteresis are considered to be negligible for rack components unless vendor or other industry experience indicates otherwise.
L/L a = 0.000 % span bistAc a
= 0.000 % span Low and High T-wu.si Pressure Reactor Tno Instrument Uncertainty /Setpoint Calculation - PBNP Setpoint Venfication Program 1 /CCP NW4 14t'T/9r 1/lo/JG VECTRA JOB NO OO87-00033.303 PAGE i I
0 KLD '7/17/96 / \ Khd 1/17?96 T ' ' .' CALC NO 22 REV BY DATE CHECKED DATE lac. PBNP-IC-12 OF 36' i
)
i
l Note 13. - Rack Drift l The drift values are based on the method of calibration performed. The drift analysis results for the lead / lag module and bistable shown below are taken from Refs. C.4. and i C.5. and are for quarterly surveillance intervals, including the 25% extension allowed by the Technical Specifications.
ULd = i0.083 % span bistAc d
= 0.222 % span i
Note 14. - Rack M&TE Effects I From the PBNP Setpoint Methodology (Ref. G.I.), the M&TE effect is included in the as-found/as-left drift data and is therefore included in the statistical drift values for the lead / lag module and bistable calculated in Refs. C.4 and C.S.
UL m = 0.000 % span bistAc m i 0.000 % span Note 15. - Rack Setting Tolerances From Refs. P. I. through P.10., the bistables for the Low and High Press Reactor Trips have one-sided setting tolerances in the conservative direction, whereas the lead / lag module has a two sided setting tolerance. For the purposes of evaluating the setpoints, the bistable setting tolerances are considered to be equal to zero.
L/L v = 0.500 % span bistAc v
= t 0.000 % span Low and High Tm .uer Pressure Reactor Trip Instrument Uncertaintv/Setpoint Calculation - PBNP Setpoint Venfication Program JOB NO 0087-00033.303 PAGE I X/7) h7t/%&v 1/ 3e/.A VECTRA O KLD 7/17/9d ) W '1/l'M)6 Techstes Inc.
CALC NO PBNP-IC.12 23 OF 36 REV BY DATE CHfCKED DATE j
t i
I Note 16. - Rack Power Supply Effect From Ref. V.I., the AC supply voltage regulation requirement for the Foxboro 610-AC-O i loop power supply is 10%. Per Ref. G.9., the instrument bus voltage varied by no more I than 2.5 Vrms during instrument bus load testing. This variation is within the 10 % :
voltage variation requirement for the loop power supply. In addition, this voltage variation was proven to.only have an effect on the current to current converters (see ,
Assumption 5.5.).
i IJL p = i 0.000 % span ,
bistic p = i0.000 % span i
Note 17. - Rack Temperature Effects Section 7.2.3 of Ref. G.3. states that the control room is maintained at 75 10*F and the protective equipment inside the room is designed to operate within design tolerance over
! this temperature range and will perform its protective function in an ambient of 110*F. .
l The instrumentation contained in the control room includes the comgionents downstream l of the transmitters (lead / lag module and bistables); therefore, there is no temperature effect associated with these components.
L/L t = t 0.000 % span bistc t = t 0.000 % span Note 18. - Rack Humidity Effect The rack components are located in the control room, which is subject to a mild environment under all plant conditions. Therefore, the humidity effect for the rack components is considered to be negligible.
IJL h = 0.000 % span bistAc h
= i 0.000 % span l
l l Low and High T.-s Pressure Reactor Trip Instrument Uncertainty /Setpoint Calculation - PBNP Setpoint Ventication Program I XfD -7M/% h 1/2./n VECTRA JOB NO 0087 00033.303 PAGE l 0 KLD 7/17/% V KM ' T/lf/% Tu'- " . " - CALC NO 24 REV 1 BY DATE CHECKED DATE Inc. PBNP-IC-12 OF 36
-, - ~ _ . _ _ _ _ __ _-- - . -
4 e
i i Note 19. - Rack Radiation Effect The rack components are located in the control room, which is subject to a mild environment under all plant conditions and is not a radiologically controlled area.
- Therefore, the radiation effect for the bistable is considered to be negligible. In addition, the as-found/as-left drift values calculated in Refs. C.4. and C.5. would include any cumulative effects of ex,posure to radiation.
4 1
ULr = 0.000 % span i
bistAc r
= 0.000 % span 1
4 Note 20. - Rack Seismic Effect
- l From the PBNP Setpoint Methodology (Ref. G.I.), setsnue or vibration effects on non- l
- mechanical instrumentation (i.e., electronic rack equipment) are considered to be zero unless vendor or other industry experience indicates otherwise. The rack components are a seismically qualified; therefore, would not be expected to experience any vibration effects under normal conditions. Ref. G.I. also assumes that for seismic events greater than an Operating Basis Earthquake, instnamentation will be recalibrated prior to any subsequent accident, thus, negating any permanent shift that may have occurred <iue to the seismic event (Assumption 5.4.).
ULs = 0.000 % span bistAc s
= t 0.000 % span '
j
! Note 21. - Process Considerations
) The uncertainties due to process considerations take into account transmitter calibration
) values, transmitter mounting elevation variations, and sensing line density variations caused by temperature and pressure changes.
4 Ref. C.2. provides a calculation for the Pressurizer Pressure transmitter calibration data.
This calculation developed a water leg correction based on the sensing line being filled with water at 68'F at atmospheric pressure:
4 Water Leg = (292.5")(0.03606 psig/inwc) = 10.55 psig
=
Low and Esh Presset.er Pressure Reactor Tnp Instrument Uncertaintv/Wint Calculation . PBNP Setpoint Venfication Program I )4*O ?/W4(; 45f4 7/4/.9(, VECTRA JOB NO OO87-00033.303 PAGE O KLD 1/17/% "W 7/11/% Te^- ' ' .t CALC NO PBNP.IC.12 25 OF 36 REV BY DAH CHECKED DAR tac.
l 1
i i
4
- The calibration procedures for the Pressurizer Pressure transmitters (Refs. P.13. and d
P.14.) show that all of the transmitters for both Unit I and 2 are calibrated for 9.47 to l 49.47 mA output corresponding to a 1700 to 2500 psig input, which is consistent with the
[ results of Ref. C.2. l l The actual mounting elevations are determined using the dimensions shown in Ref. C.2.
l for Unit 1, and the Unit 2 instmment installation detail drawings (Refs. D.3, through D.9.). The elevations of the transmitter process centerlines for each of the Pressurizer j Pressure transmitters are:
1PT-429: 56.75" + 46' = 50'-8.75" 2PT-429: 5 0'-3 "
l 1PT-430: 48.25" + 46' = 50'-0.25" 2PT-430: 50'-3.5 "
- IPT-431: 60.375" + 46' = 51'-0.375" 2PT-431: 5 0'-9.5" .
1PT-449: 42.25" + 46' = 49'-6.25" 2PT-449: 50'-3.75" l For Unit 1, the lowest transmitter is IPT-449 at 49'-6.25" and the highest is IPT-431 at
!. 5l'-0.375" For Unit 2, the lowest is 2PT-429 at 50'-3" and the highest is 2PT-431 a 50'-
9.5 " -
1 Ref. C.2. indicates that the distance between the Unit 1 Pressurizer upper and lower taps l ;
j during operation is 275" and the distance from the lower tap to the 46' elevation is 63";
{ therefore the Unit I upper tap is at an elevation of 74'-2" during plant operation.
l I The Unit 2 upper tap elevation is shown at 74'-0.75" (Refs. D.5., D.7., and D.9.), which is at cold conditions, since the as-built dimensions were tal:en during refueling. Including l the pressurizer thermal cxpansion value of 1.66" taken from C.2., the upper tap is
- expected to be at an elevation of 74'-2.41" during plant operation.
F Therefore, the actual upper tap to transmitter dimensions vary between 295.750" (IPT-l 449) and 277.625" (IPT-431), rather than the 292.50" used for the transmitter calibrations
] (Ref. C.2.).
- i J l i l l
J
} l 1
Low and High Pressunzer Pressure Reactor Tnp Instrument l j Uncertaintv/Setpoint Calculation - PBNP Setpoint Ventication Program
{- 1 T-@ Grv/h 6 1/ le/90 . VECTRA JOB NO 0087 00033.303 PAGE 1 0 KLD '7/17/% ) \lGJ '7/17796 Techaelosies CALC NO 26 REV BY DATE CHECKED DATE tac. PBNP-IC-12 OF M i
4
1 The conversion factor used in Ref. C.2. for the transmitter calibrations is applicable to water at 68 F and atmospheric pressure. However, during normal operation, the water in the reference leg is compressed water subject to the pressure within the Pressurizer (2000 psia normally, but calibrated for a range of 1700 to 2500 psig, which is equivalent to 1715 to 2515 psia). Under normal plant operation, the temperature of water in the sensing line is considered to range between. the ambient containment temperatures of 60*F and 105 F (see Section 7.3 of this calculation).
For the Low Reactor Trip, the maximum temperature considered is 250*F (see Section 7.3 of this calculation). Although the Reactor Trip is expected to occur within 5.6 seconds of the Small Break LOCA, and the sensing line fluid is not likely to reach this maximum temperature, this calculation will conservatively include the error at 250 F.
The maximum upper tap to transmitter height (IPT-449) is:
max ht diff = 295.750" The minimum upper tap to transmitter height (IPT-431) is:
min ht diff = 277.625" The following equation converts height of water in inches to units of pressure (psig) using the specific volumes for the various temperatures and pressures in the sensing line:
head correction in psig = (ht in inches)/(specific volume in ft'/D
- 1728 in'/ft')
Only the extreme values of tap to transmitter heights, normal temperatures and pressures, and accident temperatures and pressures are evaluated to determine the worst case process errors.
Low and High Pressunzer Pressure Reactor Trip Instrument Uncertaintv/Setpoint Calculation - PBNP Setpoint Venfication Program I Z$ J/:(/Q/f kh M i Ao 08 VECTRA JOB NO 0087-00033.303 PAGE O KLD 7/17/961/ N KAJ 7/17/96 Technologies CALC NO 27 REV BY DATE CHECKED DATE ,
Inc. PBNP4C-12 OF 36
Normal Conditions:
The bounding positive error under normal operating conditions (indicated pressure higher than actual) is seen at minimum normal temperature and maximum pressure (Ref G.6.).
60 F,2515 psia: specific volume = 0.01591 R'/lb max hd correction n =
(295.750")/(0.01591 ft'/lb
- 1728 in'/R')
= 10.759 psig j max hd corr error n = (10.759 psig - 10.55 psig)/800 psig
- 100%
= + 0.026 % span i The bounding negative error under normal operating conditions (indicated pressure lower than actual) is seen at maximum normal temperature and minimum pressure (Ref G.6.).
4 105*F,1715 psia: specific volume = 0.01606 ft'/lb I =
min hd correction n (277.625")/( 0.01606 ft'/lb
- 1728 in'/ft')
= 10.001 psig min hd corr error n = (10.001 psig - 10.55 psig)/800 psig
- 100%
= -0.069 % span Accident Conditions:
Since accident conditions result in a higher than normal temperature, the maximum head correction under normal conditions is the bounding case for accident conditions as well.
max hd corr error a = + 0.026 % span The bounding negative error during accident conditions (indicated pressure lower than actual) is seen at maximum temperature and minimum pressure (Ref. G.6.).
250 F,1715 psia: specific volume = 0.01690 ft'/lb
=
min hd correction a (277.625")/(0.01690 ft'/lb
- 1728 in'/ft')
= 9.508 psig min hd corr error a = (9.508 psig - 10.55 psig)/800 psig
- 100%
= -0.130 % span Low and High Pressunter Pressure Reactor Tnp Instrument Uncertaintv/Setpoint Calculation . PBNP Setpomt Verification Program I 4@ 9/WGla %N 1/ te/pb VECTRA JOB NO OO87-00033.303 PAGE O KLD 7/17/96 V \ K,tJ 7/17/96 Tuhnologies CALC NO 28 BY DATE CHECKED DATE Inc. PBNP-IC-12 OF 36 REV
l l
l Summary of Process Considerations:
l l penpos = max hd corr error n =
+ 0.026 % span penneg = min hd corr error n = -0.069 % span 1
peapos = max hd corr error a =
+ 0.026 % span peaneg = min hd corr error a = -0.130 % span Note 22. - Accident Insulation Resistance EITects From Assumption 5.8., the accident insulation resistance e.ffects are considered negligible for the Low Pressurizer Pressure Reactor Trip, ir. = 0.000 % span 8.3. Total Loop Error - Normal Conditions 8.3.1. High Pressurizer Pressure Reactor Trip Uncertainty Allowances l
The device uncertainty terms determined in Section 8.1 applicable to the High Pressudzer Pressure Reactor Trip are combined into the following Uncertainty Allowances for normal conditions:
Hi Rxr Tnp A =
(sa)2
=
( 0.224 %)2 = 0.050 Hi Rxt Trip D =
(sd)2 +(bistACd)2
( 0.895%)2 + ( 0.222%)2 = 0.850 Hi Rxr Trip V
(sv)2 = ( 0.500%)2 = 0.250 Hi Rxt Trip T =
(stn)2 = ( 1.388%)2 = 1.925 Hi Rxt Trip PC neg = penneg = -0.069 % span 8.3.2. High Pressurizer Pressure Reactor Trip Total Loop Error The Total Loop Error is determined from the Loop Uncertainty Allowances in Section 8.3.1. Combining the random and bias allowances, the TLE equation for the High l Pressurizer Pressure Reactor Trip under normal conditions becomes:
l
- Low and High Pressunzer Pressure Reactor Trip Instrument Uncertaintv/Setpotnt Calculation - PBNP Setpoint Ventication Program i AUT2 J/%/4/. W3 7i;c/M VECTRA JOB NO 0087 00033.303 PAGE O KLD 7/17/96 _A KXJ' 7/17/96 Technologies CALC NO 29 REV BY DATE CHECKED DATE tac. PBNP.IC-12 OF 36
l l
Hi Rxt Trip TLEn = - ( A + D + V + T)*# + PC neg l Hi Rxt Trip TLE n =
-(0.050 + 0.850 + 0.250 + 1.925 ) *# + (-0.069 %)
l l
Hi Rxr Trip TLE n = -1.822 % span =
-14.58 psig 8.4. Total Loop Error- Accident Conditions 8.4.1. Low Pressurizer Pressure Reactor Trip Uncertainty Allowances The device uncertainty terms determined in Section 8.1 applicable to the Low l
Pressurizer Pressure Reactor Trip are combined into the following Uncertainty Allowances for accident conditions:
Lo Rxr Trip A =
(sa)2.= ( 0.224 %)2 = 0.050 Lo Rxt Trip D =
(sd)2 + (UL d) 2 + (bistACd)
= ( 0.895%)2 + (0.083) 2 + ( 0.222%)2 = 0.857 Lo Rxt Trip V = (sv)2 + (ut y)2 = ( 0.500% )2 + ( 0.500%)2 = 0.500 l
Lo Rxt Trip T =
(sta)2 = ( 7.284%)2 = 53.057 Lo Rxt Trip PC pos = = + 0.026 % span pea Pos 8.4.2. Low Pressurizer Pressure Reactor Trip Total Loop Error The Total Loop Error is determined from the Loop Uncertainty Allowances in Section 8.4.1. Combining the random and bias allowances, the TLE equation for the Low Pressurizer Pressure Reactor Trip under accident conditions becomes:
Lo Rxt Trip TLEa = + ( A + D + V + T)*# + PC pos Lo Rxt Trip TLEa = + (0.050+ 0.857 + 0.500 + 53.057) + (0.026 % span) l Lo Rxr Trip TLEa = + 7.406 % span = + 59.25 psig 4 l
Low and High Pressunzer Pressure Reactor Trip Instrument f Uncertaintv/Setpomt Calculation - PBNP Setpoint Ventication Program i C l.77 #Y/Q(; da$.1s 7110/ % VECTRA JOB NO 0087 00033.303 PAGE O KLD 7/17/96 " 'lOJ 7/17/96 Technologies CALC NO 30 REV BY DATE CHECKED DATE tae. FBNP-IC-12 OF M l
l
8.5. High Pressurizer Pressure Reactor Trip Setpoint Evaluation For operation at 2000 psia following the Un'i 2 Steam Generator replacement, the High Pressurizer Pressure Reactor Trip Analytical Limit (AL) is 2235 psig, the proposed Technical Specification (TS) value is 2210 psig, and the proposed Actual Plant Setpoint (ATSP)is 2190 psig. As demonstrated below, the proposed plant setpoint is proven to provide adequate margin with respect to the AL. The negative Total Loop Error calculated in Section 8.3 of this calculation is added to the AL to determine the Nominal
- Trip Setpoint (NTSP). The margin between the calculated NTSP and proposed ATSP is referred to as Setpoint Margin.
l Hi Rxr Trip NTSP = AL - TLE n = 2235 + (-14.58) =
2220.42 psig i Hi Rxt Trip Setpoint Margin = NTSP - ATSP
2220.42 - 2190
30.42 psig l
The allowance from the proposed Tech Spec value to the proposed ATSP is shown I below:
Hi Rxt Trip TS to ATSP Allowance = TS - ATSP = 2210 - 2190 = 20 psig The required allowance between the ATSP and TS value is determiiled by the Allowable i Value (AV), which is defined as the value that the trip setpoint can have when tested periodically. If exceeded, the instrument channel operability is suspect and further evaluation is required. The AV is calculated by adding the magnitude of the negative Rack Error (RE) to the ATSP.
Hi Rxt Trip RE =
- [(bist Ac d)2) v2
=
- [(0.222)2) t/2
=
-0.222 % span
-1.776 psig Hi Rxt Trip AV
ATSP + l RE l = 2190 + l-1.776 l = 2191.78 psig The AV is compared to the proposed Tech Spec value to ensure that the AV is conservative with respect to the TS, the difference being referred to as margin.
Hi Rxt Trip TS to AV Margin = TS - AV = 2210 - 2191.78 =
18.22 psig l
l Low and High Pressuruer Pressure Reacter Tnp Instrument
( Uncertaintv/Setpoint Calculation - PBNP Setpotnt Venfication Program i I h~p -?/@/0/ CMS '7/h /S VECTRA JOB NO OOS7 00033.303 PAGE O KLD 7/17/96" \ KAIJ 7/17/96 Technologies CALC NO 31 REV BY DATE CHECKED DATE lac. PBNP-IC.12 OF 36 j l
1
1 l
l The AV is established to ensure that sufficient margin exists between the ATSP and the AL to account for instmment uncertainties that are not present or measured during periodic testing. This provides assurance that the AL will not be exceeded as long as the AV is satisfied and provides a means to determine unacceptable instrument performance.
The first check calculation performed adds the negative Sensor and Process Errors (S/PE) l to the AL, to account for portions of the loop not validated by surveillance testing of the l rack components.
i l Hi Rxr Trip S/PE = - [(sa)2 + (sd)2 + (sv)2 + (st.)2)ir2 + PC neg
=
l - [(0.224)2 + (0.895)2 + (0.500)2 + (1.388)2)t/2 + (-0.069)
= -1.808 % span
= -14.47 psig l Hi Rxr Trip Check Limit 1 = AL + S/PE
=
2235 + (-14.47) = 2220.53 psig Hi Rxt Trip Check Limit 1 Margin = Check Limit 1 - AV
2220.53 - 2191.78
28.76 psig The second check calculation compares the AV to the limit determined by adding the magnitude of the negative Rack Error (RE) to the calculated NTSP'to determine if the AV is conservative with respect to the second check limit.
Hi Rxt Trip Check Limit 2 =
NTSP + lRE l )
= 2220.42 + l-1.776 l = 2222.20 psig l
Hi Rxt Trip Check Limit 2 Margin = Check Limit 2 - AV
= 2222.20 - 2191.78 = 30.42 psig l Since both check calculations resulted in posi:ive margin to the Allowable Value, the AV is considered to be acceptable. The AV is Jso conservative with respect to the Tech Spec value of 2210 psig.
1 l
l 1
Low and High Pressunzer Pressure Reactor Tnp Instrument Uncertaintv/Setpoint Calculation - PBNP Serpoint Ventication Proeram i f.J -Q W&/4/c)C% 1(10/ 4 VECTRA JOB NO OO87-00033.303 PAGE O KLD 7/17/96~ ) 'KAJ '7/17/96
, Technologh CALC NO 32 DATE CHECKED DATE Inc. PBNP.IC-12 OF 36 i REV BY
l
! l 1
l 8.6. Low Pressurizer Pressure Reactor Trip Setpoint Evaluation l 1 For operation at 2000 psia following the Unit 2 Steam Generator replacement, the Low l l
Pressurizer Pressure Reactor Trip Analytical Limit (AL) is 1760 psig, the proposed Technical Specification (TS) value is 1800 psig, and the proposed Actual Plant Setpoint (ATSP)is 1820 psig. As demonstrated below, the proposed plant setpoint is proven to )
provide adequate margin with respect to the AL for accident temperature effects. The j l
Total Loop Error for accident conditions calculated in Section 8.4 of this calculation is l added to the AL to determine the Nominal Trip Setpoint (NTSP). The margin between the calculated NTSP and proposed ATSP is referred to as Setpoint Margin.
l Lo Rxt Trip NTSP =
AL + TLEa = 1760 + 59.25 = 1819.25 psig l Lo Rxr Trip Setpoint Margin = ATSP - NTSP = 1820 - 1819.25 = 0.75 psig The allowance from the proposed ATSP to the proposed Tech Spec value is shown below:
= 1820 - 1800 = 20 psig Lo Rxr Trip TS to ATSP Allowance = ATSP-TS The required allowance between the ATSP and TS value is determined by the Allowable Value (AV), which is defined as the value that the trip setpoint can have when tested periodically. If exceeded, the instrument channel operability is suspect and further evaluation is required. The AV is calculated by subtracting the applicable Rack Error (RE) from the ATSP.
Lo Rxt Trip RE =
+ {[(Lil d)2 + (bistAC d) ] + [(L/L v)2))t/2
=
+ ([(0.083)2 + (0.222)2] + [(0.5bu)2)):/2
= + 0.553 % span
= + 4.43 psig Lo Rxr Trip AV = ATSP - RE = 1820 - 4.43 = 1815.57 psig The AV is compared to the proposed Tech Spec value to ensure that the AV is conservative with respect to the TS, the difference being referred co as margin.
= 15.57 psig Lo Rxr Trip TS to AV Margin = AV-TS = 1815.57 - 1800 I
l Low and High Pressunzer Pressure Reactor Tnp Instrument
( Uncertaintv/Setpotr.t Calculation - PBNP Setoomt Ventication Program t ,
M c/*L VECTRA JOB N0 0087 00033.303 PAGE 1 2'.C77 t/M&&
O KLD 7/17/96 FKU 7/l7/96 Technologies CALC NO 33 DATE CHECKED DATE Inc. PBNP-IC-12 OF 36 REV BY
i i
l l
l 1
The AV is established to ensure that sufficient margin exists between the ATSP and the AL to account for instrument uncertainties that are not present or measured during periodic testing. This provides assurance that the AL will not be exceeded as long as the ,
l AV is satisfied and provides a means to determine unacceptable instmment performance. '
l The first check calculation performed determines the allowance required for the Sensor and Process Errors (SIPE), portions of the loop not validated by surveillance testing of the rack components. As demonstrated by the positive Setpoint Margin calculated above for the Low Reactor Trip Setpoint, the proposed plant setpoint (ATSP) conservatively allows l for the accident temperature effect on the transmitter. However, the ISA Recommended Practice for Safety Related Setpoints (Ref. G.13.) states that when considering accident environmental effects, it is possible to delete or reduce accident uncertainties from i
calculations based on the timing of the actuation functica (e.g., the trip function is I accomplished long before the environment becomes harsh enough to begin to affect equipment performance significantly).
Based on the qualification test data for the N-ElIGM model transmitter (Ref. V.4.), after the first minute of exposure to the LOCA/HELB environment, the transmitter output error was less than 1% of calibrated span, which funher demonstrates that 5.6 seconds after a SBLOCA (see Assumption 5.6.)is not a long enough period of exposure for the transmitter performance to be adversely affected. Therefore, the check calculation which adds the Sensor and Process Errors to the AL will not include the accident temperature effect for the transmitter.
Lo Rxr Trip S/PE = + {[(sa)2 + (sd)2 + (sv)2 + (st,,)2]v2 + PC pos
=
+ {[(0.224)2 + (0.895)2 + (0.500)2 + (l.388)2)v2 + 0.026
=
+ 1.766 % span
=
+ 14.13 psig to Rxt Trip Check Limit 1 =
AL + S/PE
=
1760 + 14.13 =
1774.13 psig
{
l Lo Rxt Trip Check Limit 1 Margin =
AV - Check Limit 1
=
1815.57 - 1774.13 = 41.4 psig l
Low and Hich Pressurizer Pressure Reactor Tnp Instrument Uncertamtv/Setpomt Calculation - PBNP Setoomt Venficatnn Program i U f/ 7/ W /rh %e/ O(- VECTRA JOB NO 0087-00033.303 PAGE O KLD 7/17/96 \K/J 7/17/96 Technologies CALC NO 34 REV BY DATE CHECKED DATE Inc. PBNP-IC-12 OF 36 I l
I l
The second check calculation compares the AV to the limit determined by subtracting the Rack Error (RE) from the calculated NTSP to determine if the AV is conservative with respect to the second check limit.
3 Lo Rxt Trip Check Limit 2 = NTSP - RE =
1819.25 - 4.43 = 1814.82 psig to Rxt Trip Check Limit 2 Margin = AV - Check Limit 2
=
1815.57 - 1814.82 = 0.75 psig ;
i Since both check calculatior_ .a positive margin to the Allowable Value, the AV is considered to be acceptable N is also conservative with respect to the Tech Spec value of 1800 psig.
l 1
9.0 CONCLUSION
S l The results of the instrument uncertainty calculations and setpoint evaluations performed in Section 8.0 of this calculation indicate the proposed Technical Specification values and plant setpoints for both the High and Low Pressurizer Pressure Reactor Trips provide adequate ,
margin from the Analytical Limits to the proposed plant setpoints, and from the proposed i plant setpoints to the proposed Tech Spec values. The results of this calculation are intended to be used for the Unit 2 cycle following the Steam Generator Replaceinent, during which the i normal operating pressure will be 2000 psia.
10.0 IMPACT ON PLANT DOCUMENTS The following plant documents are to be reviewed to determine if they are affected by this calculation:
Licensing Documents:
Point Beach FSAR Point Beach Technical Specifications Design Basis Documents:
Point Beach Reactor Protection System DBD Plant Documents:
PBNP Setpoint Document PBNP EOPSTPT Low and High Pressunzer Pressure Reactor Tnp Instrument p Uncertamtv/Setpomt Calculation . PBNP Setoomt Ventication Program I X.LT) Mc/4/6'h. 4/ h/J re -
VECTRA JOB NO 0087-00033.303 PAGE O KLD 7/17/96 J \'LO 7/17/96 Technologies CALC NO 35 REV BY DATE CHECKED DATE Inc. PBNP IC-12 OF 36
. Calculations:
PBNP I&C Calculation Book Procedures:
PBNP ICP Procedures PBNP RESP Procedures Other: i l
EQ Sununary Sheets l
11.0 OPEN ITEMS 11.1. This calculation will be updated to include the Low Pressurizer Pressure Safety Injection uncenainty, the indication uncertainty, and to reflect the setpoint evaluations l after the new Technical Specification values have been approved.
{ l i
l I
12.0 ATTACHMENTS 12.1. Attachment A - Ref. G.11. (2 pages) :
12.2. Attachment B - Ref. G.12. (2 pages) 12.3. Attachment C - Ref. V.3, (3 pages) ;
1 i
l t i i l 1
4 e
1 4
I 1
i
'i l
Low and High Pressunzer Pressure Reactor Tnp Instrument Uncertamtv/Setpoint Calculation - PBNP Setpoint Ventication Progt:un j 7/?c/Off "W s.
I Efi/ -1/1 e/ 9f, VECTRA JOB NO 0087 00033.303 PAGE O KLD 7/17/96 W 7/17/96 Tuhnologia tac.
CALC NO PBNP-IC-12 36 OF 36 q
l REV BY DATE CHECKED DATE l
... .. .. ... .. .. . . . . . . . . . . . . . . . . . . .... ........,.,s-..ns pest it* Far Nora 7671 M~
Juw f M ia M /M o ,
u GT~., hwou "*" /h M~e csw g fj c' 4'f)6~ SECL 95-064 ma y P?ge S Revision o p4p j)n,,, yf gq
~~ hafC hk of 7$
'"ffV.2)-9Pk.4!?V.!
ATTACIDIENT 3 - Table 5 Samimary of RPS and ESFAS Functions Actuated FSAR Event RPS or ESFAS Signal (s)
Scenen Description Analysis Delay Acmated Setpoint (sec) 14.1.1 Rod Withdrawal froan Power range high neutron flux Subentical reactor trip (low setting) 35 % 0.5 14.1.2 Rod Withdrawal at Power range high acutron flux Power reactor trip (high setting) 118% 0.5 Over*vm e AT reactor trip Table 7 Note 1 14.1.3 Dmppai RCCA low poihi pressure 1860 paia (2250 paia) 2.0 reactor trip 16 to psia (2000 pain) 2.0 (No8e 5) 14.1.4 Boron Dilution None NA NA 14.1.5 Startup of an Inactive None NA NA 180P
~
14.1.6 Feedwater Malfunction None NA NA 14.1.7 Excessrve load increase None NA NA 14.1.8 losa of Flow and Iacked RCP bus undervoltage reactor trip Note 2 1.5 Rocor I.aw RCS loop flow remeter trip 87 % 2.0 14.1.9 Ioss of Load High pressarizer prcasare reactor 2425 paia (2250 psia) 2.0 trip 2250 psia (~000 psia) 2.0 Overtamparansre AT reactor trip Tabic 7 Low-low So water level reactor trip 0% NRS (Note 3) 2.0 14.1.10 I. css of Nermal Feedwater L-Iow SGWL rescrer trip 343.2 inches (Noce 7) 2.0 L-low SGWL Iv1 AFW pump san 343.: Inches (Nuce 7) 300.0 14.1.11 I.oss of AC Power L law SGWL Ivl reactor trip 343.2 inches (Note 7) 2.0 L low SGWL Iv! AFW pump start 343.2 inches (Note 7) .500.0 14.2.5 hmHnc Brest Cort High-high steam flow setpoint 200% (Note 6) NA pm Low steam praesure SI seepom 335 psia NA 1.aw pressariant pnssure SI setpomt 1700 psia NA p,.-H., isolation delay 7.0 (from coincidenm of SI and HHSF serpoints)
Feedwasar isolados delay (frem SI seepoint) 17.0 Safety injecuon pump start (from SI setpoint) (Note 4) 17.0 SI purnp at full speed (from SI putnp start) 10.0 14.2.6 Rod Ejection Power range high neutron tiux 35% (low secing) 0.5 reactor trip (Iow and high settings) 118% (high setting) 0.5 ATTACHMENT [ SHT l il CALC NO.:
, . . , . . . . . . . . .. .. ".....i...t ...a.... e. . s 6th .N. 9.*.060 sag *
, , .m .g wssr Honot:sz PEOPSmTARY Ct.W sc SECL 95-064 Revision 0 Page 55 of 78 ATTACHMENT 3 - Table 5 (continued)
Summary of RPS and ESFAS Funenons Actuated Ta Me M Vntes
- 1. He modelling of the ovettempersmre AT reactor trip includes a time constant (first order lag) of 4.0 seconds for the measurement of the vessel Tavg and AT. His lag accounts .
for the response of the RTDs, the RTD electronic fittar (if any), the RTD bypass piping )
Ouid transport delay, and the RTD bypass piping hestup thenna! lag. In addition, a straight delay of 2.0 seconds is n=u which accounts for electronics delay, reactor trip breakers opening, and RCCA grrpper release.
- 2. RCP bus undervoltage only credited for the complete loss of flow (CLOF) event, ne RCP bus voltage is not mndeled in the LOFTRAN calculations, thus there is no specitTc seepoint assumed in the CLOF analysis. The analysis assumption is that 6s RCCAs begin to full 1.5 seconds following loss of power to the RCPs.
- 3. Nenher the ov ==p =-e AT nor the high pressurizer pressure reactor trip was reached i for the loss ofload analysis case assummg ..~. N- ... initial RCS Tavg, maxnnum initial RCS pressure, .. l... . rea.ddy. feedback, and manmum pressurizer pressure relief.
Immedimely following the mrbine trip, the nuclear power decreased rapidly due to the negative mad *ratar temperamre coefficient and reached a quasi ceady-state condnion at ,
appronmaraly 305 power. h-+ar trip med on the 'uw-low SG water !cvel signal ]
(assumed setpoint 322.6 inches above the mbesheet) at appra*=#f 130 seconds after i the turbine trip.
]
- 4. He 17-second delay for SI pump start includes 2.0 seconds for signal processing and electronics delay, and 15.0 seconds for the starmp and loading of the diesel generators (offsite power is !
- ==ad not available).
- 5. The Dropped RCCA analysis usee genene statepoims based on a low pressurizer pressure trip ,
setpoim which is 390 psi below nominal pressure. This results in the analysis values listed for the :
I Dropped RCCA event for 2000 psia and 2250 psia operation.
- 6. The high-high steam flow setpoint was increased to 200% of nominal steam flow to accommodate full-power operation at the low end of the allowed Tavg window. As Tavg decreases, the steam pressure will decrea;. . sad the volumetric steam flow rac will increase, thus increasing the measured AP which is compared to the high-high steam flow setpoint. The safety-analysis high-high steam flow seapoint was increased to 200% to allow for an increase in the actual secpoint in case there was a problem with margin to the setpoint for full power operation at a reduced Tavg.
- 7. The low-low steam generator water level setpoint of 343.2 'mcbes corresponds to 10% NRS with the lower narrow-range level tap relocated to an elevation 322.6 inches above the tubeshecc.
ATTACHMENT _ A SHT 1 /
2 CALC NO.:
P7.;A/F4C -/ Z -
~PBNP DESIGN BASIS VALIDATIO For Attribute item No. 3.3N COMMENT CBD- 27 Rev _ Final Draft t
DED Trtle _pcAf70p PAOTcCT)r)N SYSTcM A *bute to be Vsfidstef: _
m Oetermine whether environmental qualification is rm:
MSLB. based on whether the accident signal.
essary for RPS harsh environm primary trio sensors during ent occurs quicidy enough to attoct thetnosensor's Com-ents/Cenetusions Per OED Worksheet 2.3.10 the only primary trfp pararn only the RPS trip function and not the PAM ,
t functionpressure. P e intent of his attribute validation discussion is limited to LOCA environment or an MSLS event. effects on thB e pressurizer low pressure transmitter s Per conversations temperature per the PSNP with FSAM John analysisHinck is the
, of Wisconsin small b Electric th e most lirniting EQ condition for containment Therefore, this wit! be the only reak LOCA condition with a 4*evaluatedshows line break (Ref 11. Thist appen =5.6 seconds after the break.
the PBNP definition of a harsh environment: . Per the CA procedures manual $ef 2) the!
e '
A harsh temocrature environment is defined as any area j temperature in excess of 130*F occurs for local operation or repair of equignent
. This is estimated where a to sustained be the upper t
(more Smi thanfortenhum min O
psi. This value is estimatedcto pressure transients. maybe see anear the upper rapid pressure limit rise exceeding 5 o o
uman capability to withstand rapid adverse A harsh radiation environment is defined as any are degrading affects in organic polymers. Ther NRC staff O ated to be al the lower threshold value for A harsh themical spray environment exists only irtside cont iso endo o
of a F , ;0s - NaOH solution (1982 ppm H.-Sc - NaOH) at a pH of 7 51 3 a A harsh steam relat!ve created by a DBA.humidity environment is assumed to be 100% ;
relative humidity LWsinci due to A review of the FSAR figures for accident environments f parameters for the accident.
or the containment $ef 31 shows the calcula 4
Form OBCP 4 5.3 ATTACHMENT SHT I 15 CALC NO.: *] O
.,_ 7.5 N ? d C'/2-
... , , . . . - .. .. ...... ., ---...., n ., c . . . i n. 4 . ,r.:. . :, g , y 7,,-
Initial conditions in containment are assumed to be the f Itowing:
Pressure: 14.7 PSIA Temperature: 90*F Conditions at 5.6 seconds after the break are taken from the Reference 3 figures are as follows:
Pressure: 45 PSIA L31 Cd g "8 Temperature: 265'F j
.C, 5SL.cc4 Tc validata the chance in conditions in containment at the pressure transmitters and to determine the affect on these transmitters further analysis is required. This anadysis could mciude wakdowns to locate the IPansmitter and break locations, thermal gradient calculations, review of EQ report and vendor mformanon to determine affects of environmental conditions on transmitter performance. The oci.noment qualification summary sheet (Ref 4) shows information for EO of RG 1.97 function. This sheet would also require 4
revision for addition of the RPS function.
2 The team is aware of the review and analysis being done at other stations to rarnove equipment from the EQ i
program. It is the teamis opinion that based on the time recuired to perform the trip functron, that these l transmitters could be removed from the EQ program. Informal conversecons with Foxboro has confirmed Ehat test data exist with the vencor that the transmitter will not be affected by a temperature rise for 5-10 minutes.
- However, to provide objective evidence vertfying this opinion. an analysis should be completed. This is
- outside tne scope of the va!!dation effort for the DBD.
As such this attribute will remain open. l References )
l 2.
i 3. FSAR Figures 14.3.4-8,14 and 15 9,, kres.L @ !
- 4. Eculoment Qualification Summary Sheet, dated January 1,1991 d
Ct\PTIEACW\tXT\tDSC00\3-3.we i
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4 4
i 4
i
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( !
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Form DBOP 4 5.3
- ATTACHMENT SHT 31 2' Pace 2 o S'
CALC No.:
i f?&AP'SC -/2 4
a se..a w . ".a.
l FACSIMILE DATE: MAY 5, 1994 To:
VECTRA TECHNOLOGIES, INC.
1330 BUTTERFIELD ROAD ,
SUITE 550 DOWNERS GROVE, IL. 60515 KAREN DEPODASTA FAX # 708-512-8660 PHONE # 708-512-8659 FROM:
THE FOXBORO COMPANY 33 COMMERCIAL STREET D.3347/B52-2K FOXBORO, MA. 02035 DAVID R. RINGLAND '
FAX # 508-549-6580 PHONE # S08-S49-6333 FILE:
VECTRA - CIBCULATION,INFORMATION FOR N-E10 SERIES TRANSMITTERS.
SUBJECT:
YOWIAERMILE TO OUR MR. F. BONFANTI, DATED MARCIF29P,;R 1994.
PAGES:
THREE INCLUDING THIS PAGE.
COPIES: F. BONFANTI, CH1-01 R. SCHWANTIES CH1-01 l
l l
ATTACHMENT b SHTI/
l CALC NO.:
78MP-2 ~/1
Resnonses to Questions from Vectra Recardine N-E11 Transmitters and N-E13
References:
PSS9-1B1A (1984) and FOXBORO Qualification Document QOAAC11 4
- 1. Accuracy expressed as a +/- % does include the combined i i
Each of the aforementioned characteristics has a specif limit. ~
specification. All specification are in % of Span.The specified l 1
1 i
i 2. The performance characteristics in Question / Response 1 ;
{ are measured at " Reference Operating Conditions" and .
i performance at " Normal Operating Conditions" includes
. the influences of Ambient Temperature Effects, etc.
i Accuracy under ambient temperature changes does affect
! the zero and span of the transmitter. The other i characteristics should not at other than reference change,conditions.
operating but are not specified
! Using the example of an N-E11DM transmitter the Ambient Temperature Effects are specified as follows:
Zero Shift Span Settings,4 of USL per 100F-Change ,% of Span Above per 170F Change i
80%
MBA2 (32 to 180F) 100% +/-1%
_ (80 to 250F)_
50% 80% +/~2%
{ 20% 50% +/-1.5% +/-34
- +/-2.5% +/-5%
j Span change: +/-1.25% per 100F Example:
i N-E11DM-IIB, USL:200 psi Calibrated Range: 0-100 psi i
! 100 psi = 50% of USL 200 psi therefore the
{
Ambient Temperature Effects are :
- +/-2.5%/100F or +/-5%/170F Relative Humidity Effects
- Negligible P us humi has no of e .
/
I *
- 3. The " Normal TIDof3.5x1y0diation" Specification of +/-0.5% for a i specification. FOXBORO Qualification Report QOAAC11, rads gamma is Sect.IV.,
i Pg. IV-25 does show graphically several other i
lower radiation levels for an N-E11GM transmitter which is i
)
similar to an N-E11DH transmitter. The radiation effects can be zero and span adjusted to return to the normal accuracy specification.
i j
i i
i ATTACHMENT C/ SHT11 CALC N04 77;N7 GC-lZ- -
E
. . , - - - - , - - . - - e- , - s
tv <-as-:, m c:53 Acn PA2e 54 e5 0 snes cessa p,a ;
- 4. The Seismic DBE performance specifications of +
were sat as goals in the transmitter qualification program and we did not attempt to determine a threshold response spectrum. an transmitter Reviewing N-EllGM the(F1)qualification data for the similar USL, betteritperfor=ance is possible that spans above this setting do havewhose span set specifications.
- 5. The I4CA/HIL3 Output Shifts of +/-8% at 25% of USL and
+/-3% at USLUsing these lLaits. can be theinterpolated N-E11GM from for span the settings in-between the span setting was 40% of USL thus the +/-8% qualification test be selected. spec. would Usines the Ambient Temperature Effects table as follows the so to 80% span settings can be developed:
Span Settings, output Shift, Span *
% of USL % of Span output Shift, Above Ubto ist 3 Hrs.
Setting % of Span 80% Ratio 50 to 80%
100% +/-3% Ref.
50% 80% -----
20% 1.5 +/- 4.5%
50% +/-8% 2.5 Adding the 50 tomargin 80% ofwe would USL specify the output Shift at +/-5% for settings.
- The IcCA/HELB is an event similar to ambient temperature effects and the use of the normal ambient temperature specifications is justified in deriving an error ratio for the 50 to 80% settings and applied to the USL specification.
Note:
The N-E11GM and N-E11DH transmitters have the same Ambient Temperature Effects as stated above, for specificationSeries other N-E10 differsTransmitters and must bethe manner of reviewed individually.
ATTACHMENT SHT3l.$.
CALC NO.:
T%v7-sc -iz-TC':.,L 3, ]T
e_ .. m . -
ATTACIIMENT FINAL SAFETY ANALYSIS REPORT MARK-UPS 4
s 9
i a
e I
TABLE 4.1-1 4
REACTOR COOLANT SYSTEM DESIGN PARAMETERS AND PRESSURE SETTING Total Primary Heat Output, MWt /f2u, y- 1 ;;,5-Total Primary Heat Output, Btu /hr fed 2. Ji.1&T' x 10' Number of Loops 2 Coolant Volume (liquid), including pressurizer g (Unit 2) 6000 (Unit I volume, at full power (60% full), ft* bl Yb l Total Reactor Coolant Flow, Ib/hr 66.7 x 10' I I
Pressure i
g.11g ?
i Design Pressure 2485 Operating Pressure (at pressurizer) 1985 or 2235 z 100 Safety Valves 2485 Power Relief Valves 2335
! Pressurizer Spray Valves (open) 2260 j High Pressure Trip 5 2385 '
Low Pressure Trip 1 1790 or 1 1855 ,
Hydrostatic Test Pressure (Cold) 3110 l h
l 4
- 5, 425 psig when the Overpressurization Mitigating System is activated i
j
__. .. .~ _
. TABLE 4.1-4 STEAM GENERATOR DESIGN DATA i
Unit 2 Unit 1 Model , W 4 tf 7 44F Number of Steam Generators 2 2 Design Pressure, Reactor Coolant /
Steam, psig 2485/1085 2485/1085 Reactor Coolant Hydrostatic Test pressure (tube side-cold),psig -
4M6 3/6 7 3106 i Design Temperature, Reactor Coolant / Steam,'F 650/556 650/556 Reactor Coolant Flow, gpa 89,00/85200 89,000 l Total Heat Transfer Surface {
Area, ft' M 47/C0 43,467 Heat Transferred, Btu /hr ,2 60/ W 10' 2591 x 10*
Steam Conditions at Full Load, Outlet Nozzle:
Steam Flow, lb/hr 3,2 7 - 3.31 x 10' 3.31 x 10' Steam Temperature, 'F g, p -Jim ^f26,2 521.2 Steam Pressure, psia C/2 ~ .82T 677 821 )
Feedwater Temperature, at 1005 Load, 'F MM30,0 435.7 j Overall Height, ft-in. ,j3-F 6- 69_ -// 63-1.6 Shell 00, upper / lower, in. /66,4 M67MP /2.7, 8 166/127
. Shell Thickness, upper / lower, in.3*h/4-75 2 s/ 3 FLM/LM-2,7T Number of U-Tubes M .7477 3214 U-Tube 00, in. 0.875 0.875 Tube Wall Thickness, (nominal),
in. 0.050 0.050 Number of Manways/ID, in. 4/16 3/16 Number of Handholes/ID, in, g f$6 6/6 Inspection Ports /ID, in. p/r 2[9 1/3 (1 of 2)
TABLE 4.1-4 (continued) ggg.- Unit 2 Unit 1 1518.5 "Wt Zero Power 1518.5 MWt Zero. Power neactor Side Coolant Water Volume, ft' gff/ M f9/ 925 925 Orimary Side fluid Heat Content, Stu 2 6 -2 K 7 53 i+:99 x 10' 'A
x 10' 24.99 x 10' 24.42 x 10' iecondary Side Water Volume, jggg_gyg gg ft' 46fH- -38i9- 1714 2877
'econdary Side Steam Volume, ygyg _yjpg. j 7pg ft' -ee9fr +759- 2967 1804 scondary Side Fluid Heat 3 7_ gg 7y g Content, Btu -45.00s 10' 50- x 10' 45.80 x 10' 75.50 x 10' (2 of 2)
~
j TABLE 4.1-8 l THERMAL AND LOADING CYCLES i
h ansient Condition -
Desion Cycles *
- 1. Plant heatu;, at 100*F per hour 200 (5/yr) 1 2. Plant cocidown at 100*F per hour 200 (5/yr)
. 3. Plant loading 'at 5% of full power per minute 14,500 (1/ day)
- 4. Plant unloading at 5% of full power per minute 14,500 (1/ day)
- 5. Step load increase of 10% of full power 2,000 (1/ week) l (but not to exceed full power) j 6. Step load decrease of 10% of full power 2,000 (1/ week)
- 7. Step load decrease of 50% of full power 200 (5/ year)
- 8. Reactor trip and attendent temperature 400 (10/ year) transients
- 9. Hydrostatic test pressure 3110 psig 5 (pre-operational) temperature 100'F
- 10. Hydrostatic test pressure 2485 psig yt/[(post-operational) j temperature 400*F
/ 3)f. Steady state fluctuations - The reactor coolant average temperature for purposes of design is assumed to increase and decrease a maximum of 6*F in one minute. The corresponding reactor coolant pressure variation is less than 100 psig. It is assumed that an infinite number of such fluctuations will occur.
60
- Estimated for equipment design purposes gyear life) and not intended to be an accurate representation of actual transient or to reflect actual operating experience.
te -uw,.4 / g7
//,
7 es44st P,wo,.y70psy)
(3:-
IES tb Sewdmy b-ixecry l@N N5b
l l
l l
the control room. A small continuous spray flow is provided to assure that
) the pressurizer liquid is homogeneous with the coolant and to prevent excess cooling of the spray piping.
During a negative pressare surge caused by an increase in plant load, flashing of water to steam and generation of steam by automatic actuation l! of the heaters keep the pressure above the minimum allowable limit.
Heaters are also energized on high water level during positive surges to
}
heat the subcooled surge water entering the pressurizer from the reactor l coolant loop.
i i
i- The pressurizer is constructed of carbon steel with internal surfaces clad f with austenitic stainless steel. The heaters are sheathed in austentic stainless steel. .All nozzle safe ends (forgings) in the top and bottom j heads and the nozzles of the pressurizer safety valves may have been j furnace sensitized during the fabrication sequence. Subsequent non-
! destructive examination showed no degradation in integrity of the i materials.
The pressurizer vessel surge nozzle is protected from thermal shock by a l thermal sleeve. A thermal sleeve also protects the pressurizer spray '
l nozzle connection.
I l' Pressurizer - Sunnort Structure j The pressurizer is supported on a heavy concrete slab spanning the concrete I
shield walls of its compartment. The pressurizar is a bottom-skirt
- supported vessel.
1 i Steam Generators Each loop contains a vertical shell and U-tube steam generator. A steam i generator of this type is shown in Figure 4.2-4. Principal design param-eters are listed in Table 4.1-4.
4 i
i 4.2-5 June 195 i.
_ _ _ _ _ _ - _ - . _ _ ~- - - - - - - - - -
1 i
Reactor coolant enters the inlet side of the channel head at the bottom of l the steam generator through the inlet nozzle, flows through the U-tubes l
l to an outlet channel, and leaves the generator through another bottom nozzle. ;
l The inlet and outlet channels are separated by a partition. Primary side i manways are provided to permit access to the U-tubes. This permits steam generator tubes to be periodically inspected and allows defective tubes to be repaired or plugged in accordance with approved procedures.
Feedwater to the steam generator enters just above the top of the U-tubes through a feedwater ring. The water flows downward through an annulus between the tube wrapper and the shell and then upward through the tube bundle where part of it is converted to steam.
The steam-water mixture from the tube bundle passes through a steam swirl vane assembly which imparts a centrifugal motion to the mixture and separates the water droplets from the steam. The water removed by the swirl vane combines with the feedwater for another pass through the tube ;
bundle.
{
The steam rises through additional separators which limit the moisture content of the steam to one fourth of one percent or less under all design load conditions.
(Alky G90 The steam generator is constructed primarily of carbon steel. The heat -
transfer tubes are 3 1-4 The interior surfaces of the channel heads and nozzles are clad with austenitic stainless steel, and the side of the tubesheet in contact with the reactor coolant is clad with insenet. The tube to tubesheet joint is welded. ,
g The two primary nozzle safe ends per generator became furnace sensitized at the weld metal buttering during the fabrication sequence. Subsequent non-destructive examination showed no degradation in the integrity of the materials.
4.2-6 June 1992
l l
1 4
i i
The evaluation of Westinghouse steam generator tubesheets is performed 2
according to rules of the ASME Boiler and Pressure Vessel Code for Nuclear Vessels,Section III,1968 Edition Article 4 - Design. The design criteria
] encompasses steady-state, transient, and emergency operations as specified in the Equipment Specification. Due to the complex nature of the tube-k tubesheet-shell-head structure, the analysis of the tubesheet required the application of results of related research programs (such as the design data on perforated plates resulting from PVRC programs) and the utilization i
of current techniques in computer analysis, the application of which is verified by comparison of analytical and experimental results for related j equipment.
I The Westinghouse analysis of the steam generator tubesheets is included as part of the Stress Report requirements for Class A Nuclear Pressure
! Vessels. The evaluation is based on the stress and fatigue limitations j outlined in Article 4 Design of Section III. The stress analysis tech-niques utilized include all factors considered appropriate to conservative i
determination of the stress levels utilized in evaluation of the tubesheet complex. The analysis of the tubesheet complex includes the effect of all
- appurtenances attached to the perforated region of the tubesheet considered j appropriate to conservative analysis of stress for evaluation on the basis f
of Section III stress limitations. The evaluation involves the heat conduction and stress analysis of the tubesheet, channel head, secondary 4
- shell structure for particular steady design conditions for which Code l stress limitations are to be satisfied, and for discrete points during l transient operation for which the temperature / pressure conditions must be known to evaluate stress maxima and minima for fatigue life usage. In I addition, limit analyses are performed to determine tubesheet capability to j sustain emergency operating conditions for which elastic analysis does not
] suffice. The analytic techniques utilized are computerized and significant stress problems are verified experimentally to justify the techniques where 4 possible.
l 4
Generally, the analytic treatment of the tube-tubesheet complex includes l determination of elastic equivalent plate stress within the perforated.
l region from an interaction analysis utilizing effective elastic constants j appropriate to the nature of the perforation array. For the perforated 4.2-7 June 191
l l I
region of the tubesheet, the flexural rigidity is based on studies of j
behavior of plates with square hole arrays utilizing techniques such as I those reported by O'DonnellW, Mahoney*, Leacoe*, and others. Similarly,
! 1 stress intensity factors are determined for square hole arrays using the I j
combined equivalent plate interaction forces and moments applied to results
{ of photoelastic tests of model coupons of such arrays as well as i
verification using computer analysis techniques such as " Point Matching" or j
" Collocation . The stress analysis considers stress due to synnetric temperature and pressure drop across the tubesheet divider lane.
I j
The fatigue analysis of the complex is performed at potentially critical I regions in the complex such as the junction between tubesheet and channel i
) head or secondary shell as well as at many locations throughout the l perforated region of the tubesheet. For the holes for which fatigue evaluation is done, several points around the hole periphery are considered i j
to assure that the maximum stress excursion has been considered. The i fatigue evaluation is computerized to include stress maxima-minima j
)
excursions considered on the intra-transient basis. ;
i j
The evaluation of the tube-to-tubesheet juncture of Westinghouse PWR System j steam generators is based on a stress analysis of the interaction between l tube and tubesheet hole for the significant thermal and pressure transients ;
i that are applied to the steam generator in its predicted histogram of cyclic operation. The evaluation is based on the numerical limits i specified in the 1968' Edition of the ASME Boiler and Pressure Vessel Code, i Section III, Nuclear Vessels.
i j Of importance in the analysis of the interaction system is the behavior of
~
the tube hole, where it is recognized that the hole behavior is a function j of the behavior of the entire tubesheet complex with attached head and shell. Hence, the output of the tubesheet analysis giving equivalent plate l stresses in the perforated region is utilized in detemining the free j boundary displacements of the perforation to which the tube is attached.
l Analysis of the juncture for the fillet-type weld utilized in the West-l inghouse steam generator design has been made with consideration of the effect of the rolled-in joint in the weld region as well as with the 1
4.2-8 June 1992 1
conservative assumption that the tube flexure relative to the perforation is not inhibited with the rolled-in effect.
The major concern in fatigue evaluation of the tube weld is the fatigue strength reduction factor to be assigned to the weld root notch. For this reason, Westinghouse has conducted low-cycle fatigue tests of tube material
)
samples to detemine the fatigue strength reduction factor and applied them to the analytic interaction analysis results in accordance with the accepted t'echniques in the Nuclear Pressure Vessel Code for Experimental Stress Analysis. The fatigue strength reduction factor determined '
4 therefrom is not different from that reported in the well known paper on the subject by O'Donnell and Purdy*. An actual tubesheet joint contained in a tubesheet has been successfully tested experimentally under themal transient conditions much more severe than that achieved in anticipated power plant operation.
f i 1 A wide range of computational tools are utilized in these solutions including finite element, heat conduction, and thin shell computer solu-i tions. In addition, analysis techniques have been verified by photoelastic model tests and strain gaging of prototype models of an actual steam !
j generator tubesheet.
1 Finally, in order to evaluate the ultimate safety of the structural com-
] plex, a computer program for detemining a lower-bound pressure limit for the complex based on elastic-plastic analysis has been developed and applied to the structure. This was verified by a strain gage steel model l of the complex tested to failure.
l In all cases evaluated, the Westinghouse steam generator tubesheet complex
! meets the stress limitations and fatigue criteria specif.ied in Article 4 of i the Code as well as emergency condition limitations specified in the Equipment Specifications or anticipated otherwise.
In this way, the tube-tubesheet integrity of a Westinghouse steam generator l is demonstrated under the most adverse conceivable conditions resulting i
from a major breach in either the primary or secondary system piping.
4.2.g June 199 9
Steam Generator - Sunoort Structure Each steam generator is supported on a structural system consisting of four vertical support columns and two (upper and lower) support rings. The 3
vertical columns, which are pin connected to the steam generator support j feet, serve as vertical restraint for operating weights, pipe rupture, and
)
i l seismic considerations while permitting movement in the horizontal plane. l
) The support rings, by using a combination of pins, stops, guides, and
! snubbers, prevent rotation and excessive movement of the steam generator in l l any plane. Thermal expansion is permitted in the support rings by a key i arrangement. !
i i ~ !
a f Unit 2 - Steam Generator Tube Sleevina i
j l ;
i !
Point Beach Nuclear Plant Unit 2 steam generator tubes were sleeved during , !
i l 1983 as a preventive maintenance program to minimize the likelihood of 1 l primary-to-secondary leakage and the resultant adverse affect on unit '
i reliability and availability. The initial sleeving program included the l l 1500 tube central region of the hot leg of each steam generator, the region I l in which intergranular attack and stress corrosion cracking of the outside l
{ diameter of the tubes is most likely to occur due to contaminant
{
j concentration. The internal sleeves, which provide a new primary side l l pressure boundary, span the tubesheet thickness plus the maximum l
l , experienced height of sludge and are made of a material which was j
] l specifically chosen for its inherent resistance to the corrosion
, experienced in the local steam generator environment.
l i
j Maintenance through selective sleeving of tubes on the cold leg side of the Unit 2 steam generators was performed in 1987, 1988, and 1989. The total l l l number of sleeves installed for these three years is 118 and 776 sleeves int l the A and 8 steam generators, respectively. This sleeving was performed to
! address observed cold lea thinning af the tube _s just above the tubeshe '
j -
I Unit 1 - Steam Generator Replacement i
i a
Both Unit I steam generators lower assemblies were replaced during 1984.
l The. performance of the replacement lower assemblies matches the performance
! 4.2-10 June 1992 l
L_____________. __ _ . . . . _
)
i of the original lower assemblies. However, several design features that do i
not alter the perfor1sance parameters are included in the design. A comparison of the original Westinghouse Model 44 steam generators and the replacement Westinghouse Model 44F steam generators is provided in Table !
j 4.1-4. The design features of the Model 44F steam generator lower j
assemblies and modifications made to the moisture separator equipment of l the upper assemblies provide improved thermal hydraulic performance, provide improved access to the tube bundle, and reduce the potential for secondary side corrosion. The original steam generator lower assemblies, f which were removed, are stored onsite in a shielded, restricted access
! structure awaiting future disposition.
i
%,t2 -swa Rekt k J ysU Reactor Coolant Pumos '
i 3
Each reactor coolant loop contains a vertical single stage centrifugal pump which employs a controlled leakage seal assembly. A view of a controlled leakage pump is shown in Figure 4.2-5 and the principal design parameters
{ for the pumps are listed in Table 4.1-5. The reactor coolant pump estimated performance and NPSH characteristic are shown in Figure 4.2-6.
)[
The performance characteristic is common to all of the higher specific j speed centrifugal pumps and the ' knee" at about 45% design flow introduces
- no operational restrictions since the pumps operate at full flow.
l The motor-impeller can be removed from the casing for maintenance or l inspection without removing the casing from the piping. All parts of the pumps in contact with the reactor coolant are austenitic stainless steel or j equivalent corrosion resistant materials. l i
1 l The pump employs a controlled leakage seal assembly to restrict leakage ;
i along the pump shaft, as well as a secondary seal which directs the controlled leakage out of the pump, and a third seal which minimizes the leakage of water and vapor from the pump into the containment atmosphere.
! l
} A portion of the high pressure water flow from the charging pumps is 1 injected into the reactor coolant pump between the impeller and the l controlled leakage seal. Part of the flow enters the Reactor Coolant l' System through a labyrinth seal in the lower pump shaft to serve as a 4.2-11 June 199 1
1 _ __ _ _ _ _ . _ _ _ - _
Unit 2 - Stamm Generator Reo1= cement Both Unit 2 steam generators have been replaced. Whereas the Unit I replacement project !
changed out only the lower assemblies, the Unit 2 replacement steam generators (RSGs) !
consisted of the complete vessel, i.e. both the lower and upper assemblies. De RSGs are Westinghouse Model A47 and are similar in design and functionally the same as the orginal Westinghouse Model 44 steam generators. Design data of the replacement generators for Unit 2 is provided in Table 4.1-4. De RSGs have design features which provide additional i resistance to known degradation mechanisms and which support their reliability and maintainshility.
1 I
l
M i
4 i -
- e 90 c H Ad@g j 14 - SAFETY ANALYSES l
This section evaluates the safety aspects of either Unit 1 or Unit 2 of the
)
i plant, demonstrates that either or both units can be operated safely and that
! exposures from credible accidents do not exceed the guidelines of 10 CFR 100.
i i
This section is divided into three subsections, each dealing with a different i behavior category:
Core and Coolant Boundary Protection Analysis. Section 14.1
! The abnormalities presented in Section 14.1 have no off-site radiation con-I sequences.
b l Standby Safety Features Analysis. Section 14.2
{ The accidents presented in Section 14.2 are more severe than those discussed j in 14.1 and may cause release of radioactive material to the environment.
1
- Ruoture of a Reactor Coolant Pine. Section 14.3 -
I The accident presented in Section 14.3, the rupture of a reactor coolant pipe, j is the worst case accident and is the primary basis for the design of i engineered safety features. It 1: shown that even the consequences of this accident are within the guidelines of 10 CFR 100.
l -
j .
I Parameters and assumptions that are common to various accident analyses are j described below to avoid repetition in subsequent sections.
1 Steadv State Errors
{
i
). For most accidents which are DNB limited, nominal values of initial conditions i are assumed. The allowances on power, temperature,,and pressure are j determined on a statistical basis and are included in the limit DNBR, as
.) described in WCAP-ll397 (Reference 1). This procedure is known as the f
, 14-1 June 1992 a
QC (Nb dk N i
j ' Revised Thermal Design Procedure," and is discussed more fully in i Section 3.2.2.
For accidents in which the Revised Thermal Design Procedure is not employed, the initial conditions are obtained by , adding the maximum steady state errors to rated values. The following conservative steady state errors were assumed f
in the analyses:
i i 1. Core Power '12 percent allowance for l
calorimetric error
! 2. Average Reactor Coolant 14*F allowance for controller
. Temperature deadband and measurement error j 3. Pressurizer Pressure t30 pounds per square inch (psi)
- allowance for steady state l fluctuations and measurement error i
Tables 14-1 and 14-2 sussiarize initial conditions and computer codes used in the accident analyses, and show which accidents employed a DNB analysis using i the Revised Thermal Design Procedure (RTDP).
i i Power Distribution '
l i
The transient response of the reactor system is dependent on the initial power distribution. The nuclear design of the reactor core minimizes adverse power i.
distribution through the placement of control rods and operating instructions,
]
j Power distribution may be characterized by the radial peaking factor (F.,,) and j the total peaking factor (F ). The peaking fac, tor limits are given in the l Technical Specifications.
i' For transients which may be DNS limited, the radial peaking factor is of l importance. The radial peaking factor increases with decreasing power level i due to rod insertion. This increase in F.,, is included in the core limits i illustrated in Figure 14-1. All transients that may be DNS limited are assumed to begin with a F.,, consistent with the initial power level defined in j the Technical Specifications. -
f The axial power shape used in the DNB calculation is discussed in Section 3.2.2.
, 14-2 June 1992
. go cHAJ4E5 The radial and axial power distributions described above are input to the THINC code as described in Section 3.2.2.
For transients which may be overpower limited, the total peaking factor (F )
is of importance. All transients that may be overpower limited are assumed to begin with plant conditions, including power distributions, which are consistent with reactor operation as defined in the Technical Specifications.
For overpower transients which are slow with respect to the fuel rod thermal time constant (for example, the Chemical and Volume Control System malfunction which results in a decrease in the boren concentration in the reactor coolant, lasting many minutes, and the excessive increase in secondary steam flow incident which may reach equilbrium without causing a reactor trip), the fuel rod thermal evaluations are performed as discussed in Section 3.2.2. For overpower transients which are fast with respect to the fuel rod thennal time constant (for example, the uncontrolled rod cluster control assembly bank withdrawal from suberitical and rod cluster control assembly ejection incidents which result in a large power rise over a few seconds), a detailed fuel heat transfer calculation must be performed. Although the fuel rod thermal time constant is a function of system conditions, fuel burnup and rod -
power, a typical value at beginning-of-life for high power rods is approximately five seconds.
Reactivity Coefficients Assumed in the Accident Analyses The transient response of the reactor system is dependent on reactivity feedback effects, in particular the moderator temperature coefficient and the Doppler power coefficient. These reactivity coefficients and their values are discussed in detail in Section 3.2.1.
In the analysis of certain events, conservatism requires the use of large reactivity coefficient values, whereas in the analysis of other events, conservatism requires the use of small reactivity coefficient values. Some analyses such as loss of coolant from cracks or ruptures in the Reactor Coolant System do not depend on reactivity feedback effects. The justification for use of conservatively large versus small reactivity co-14-3 June 1992 s
.- 90 c4AMR efficient values is treated on an event-by-event basis. In some cases conservative combinations of parameters are used to bound the effects of core life, although these combinations may represent unrealistic situations.
Rod Cluster Control Assembiv Insertion Characteristics The negative reactivity insertion following a reactor trip is a function of the position versus time of the rod cluster control assemblies and the variation in rod worth as a function of rod position. With respect to accident analyses, the critical parametar is the time of insertion up to the dashpot entry or approximately 85 percent of the rod cluster trav'el.
The rod cluster control assembly position versus time assumed in accident analyses is shown in Figure 14-2. The rod cluster control assembly insertion time to dashpot entry is taken as 2.2 seconds.
Figure 14-3 shows the fraction of total negative reactivity insertion versus normalized rod position for a core where the axial distribution is skewed to the lower region of the core. An axial distribution which is skewed to the.
lower region of the core can arise from an unbalanced xenon distribution.
This curve is used to compute the negative reactivity insertion versus time .
following a reactor trip which is input to all point kinetics core models used in transient analyses. The botton-skewed power distribution itself is not l input into the point kinetics core model.
There is inherent conservatism in the use of Figure 14-3 in that it is based j on a skewed flux distribution which would exist relatively infrequently. For cases other than those associated with unbalanced xenon distributions, significant negative reactivity would have been inserted due to the more favorable axial distribution existing prior t'o trip.
The normalized rod cluster control assembly negative reactivity insertion versus time is shown in Figure 14-4. The curve shown in this figure was obtained from Figures 14-2 and 14-3. A total negative reactivity insertion following a trip of 4 percent AK is assumed in the transient analyses except where specifically noted otherwise. This assumption is conservative with respect to the calculated trip reactivity worth available. For Figures 14-2 14-4 June 1992 s
go cdA4465 and 14-3, the rod cluster control assembly drop is normalized to 2.2 seconds, unless otherwise noted for a particular event.
Reactor Trio A reactor trip signal acts to open the two series trip breakers feeding power to the control rod drive mechanisms. The loss of power to the mechanism coils causes the mechanisms to release the control rods, which then fall by gravity into the core. There are various instrumentation delays associated with each tripping function, including delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to trip is defined as the time delay from the time that trip conditions are reached to the time the rods are free and begin to fall. The time delay assumed for each tripping function is given in Table 14-3.
Reference is made in Table 14-3 to overtemperature and overpower AT trip points shown in Figure 14-1. Figure 14-1 presents the allowable reactor coolant loop average temperature and AT for the design flow and power dis-tribution, as described in Section 3.2.2, as a function of primary coolant i pressure. The boundaries of operation defined by the overpower AT trip and l
the overtemperature AT trip are represented as " Protection Lines" on this .
diagram. The protection lines are drawn to include all adverse instrumen-tation and setpoint errors so that under nominal conditions a trip would occur well within the area bounded by these lines. The utility of this diagram is in the fact that the limit imposed by any given DNBR can be represented as a line. The DN8 lines represent the locus of conditions for which the DN8R equals the limit value (1.33 for the thimble et11 and 1.33 for the typical cell). All points below and to the left of a DN8 line for a given pressure have a DNBA greater than the limit value. The diagram shows that DNB is prevented for all cases if the area enclosed with the maximum protection lines is not traversed by the applicable DNBR line at any point.
The area of permissible operation (power, pressure, and temperature) is bounded by the combination of reactor trips: high neutron flux (fixed setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint);
overpower and overtemperature AT (variable setpoints).
14-5 June 1992 s
l I
l cHA#3G;E.5}
[ 40 The limit value, which was used as the DNBR limit for all accidents analyzed with the Revised Thermal Design Procedure (see Table 14-1), is conservative compared to the actual design DNBR value required to meet the ONB design basis .
as discussed in Section 3.2.2.
i The difference between the limiting trip point assured for the analysis and the nomal trip point represents an allowance for instrumentation channel error and setpoint error. Nominal trip setpoints are specified in the plant Technical Specifications.
Instrumentation Drift and Calorimetric Errors - Power Rance Neutron Flux I i
The instrumentation drift and calorimetric errors used in establishing the power range high neutron flux setpoint are represented in Table 14-4.
The calorimetric error is the error assumed in the detemination of core thermal power as obtained from secondary plant measurements. The total ion chamber current (sum of the top and bottom sections) is calibrated (set equal) to this measured power on'a periodic basis.
The . secondary power is obtained from measurement of feedwater flow, feedwater inlet temperature to the steam generators and steam pressure. High accuracy instrumentation is provided for these measurements with accuracy tolerances much tighter than those which would be required to control feedwater flow.
Plant-to-Plant Interaction ,
The safety evaluation of a two unit plant, where two reactors are situated in close physical proximity on the same site, sharing certain facilities and l operated as combined power producing units, requires that the safety assessment treat the plant as a two unit facility rather than as two indi-l vidual single unit facilities. However, for the reasons discussed below, the l nature of the two unit plant design confines the location of a reactor fault I
condition to one of the two units at any time (with the exception of possible f faults arising in the electrical grid system to which both units are connected, and these have no off-site radiation consequences). Thus, for the 14-6 June 1992 l ,
l
)
lp0 G H AMk two unit plant, the potential consequences of each and every credible reactor 6
fault condition are no different than those for a single unit plant.
Possible sources of interaction between the two units are discussed below:
Sharina of Systems i
As noted in Sections 1, 9, 10 and 11, all or part of certain systems (e.g.,
Chemical and Volume Control System, Waste Disposal System) are shared by the two units. A functional evaluation of the components of those systems which l
, are shared by the two units is given in Appendix B. i
)
The plant is provided with a control room which is common to both units.
j Physical separation of control panels in the control room essentially ella-
,4 inates interaction of the control systems of the two units.
l The two units are connected to the same external electrical grid, and it is therefore possible that the following transients could affect both units simultaneously:
l 1
l
- 1. Loss of external electrical load (Section 14.1.9) I
! 2. Loss of all AC power to the station auxiliaries (Section 14.1.11) !
I
- The design is such that the occurrence of either of these two transients, in both units simultaneously, can be accommodated without an unsafe condition arising in eithe unit.
Except for the electrical grid conditions noted above, all systems which are t
shared by both units are designed such that a shared system can neither cause a simultaneously unsafe condition in both units, nor propagate an accident condition, which may arise in one unit, to the other unit.
i j Physical Proximity
}
- The positioning of the two units in close physical proximity introduces no possibility of external interaction. For each unit, the integrity of all systems whose functions are necessary to maintain the safety of the reactor is 1
, 14-7 June 1992
s O (M ensured by the nature of the design: e.g., through separation of redundant components such as wiring, and missile shielding both inside and outside the l l
, containment.
Thus, with the exception of the electrical faults already noted, the two unit l
plant precludes by the nature of its design, any possibility of either (a) j simultaneous occurrence in both units of fault conditions having a cosmon ori-gin, (b) the propagation from one unit to the other unit of a fault condition.
I j In addition, it is not considered credible that both units could develop i unrelated accidents, either of the same or a different nature simultaneously. !
Thus, the criteria for plant design require the capability to deal with the
- affected unit while maintaining safe control of the other unit. Although t'xse criteria do not directly imply that the other unit must be shut down 3110 wing the occurrence of an accident condition in one unit, the two unit f 1
plant design includes the capability to meet all safety criteria in the affected unit, and simultaneously shut the second unit down and maintain it at l hot shutdown, if required. In fact, continued on-line operation of the f adjacent unit enhances the assurance of a continuous supply of electrical j power for the engineered safety features of the affected unit.
\
In a two unit plant, the overall design of each unit represents no essential ,
departure from the current design of the unit which comprises a single unit {
pl ant. Thus, the methods and techniques for the safety assessment of a single unit plant are directly applicable to a two unit plant. Further, since both
! units of a two unit plant are nearly identical, the safety assessnent
- (presented in this section for a single unit) is equally applicat,le to either l unit.
l I
Comeuter Codes Utilized Summaries of some of the principal computer codes used in transient analyses are given below. Other codes, in particular very specialized codes in which the modeling has been developed to simulate one given accident, such as those used in the analysis of the primary system pipe rupture (Section 14.3), are sunnarized in their respective accident analyses sections. The codes used in l the analyses of each transient have been listed in Table 14-1.
I 14-8 June 1992
. _\
c t4 AJ M5 \
FACTRAN
{ 90 FACTRAM calculates the transient temperature distribution in a cross section of metal clad 00, fuel rod and the transient heat flux at tt.s surface of the cladding using as input the nuclear power and time-dependent coolant parameters (pressure, flow, temperature, and density). The code uses a fuel model which exhibits the following features simultaneously:
- 1. A sufficiently large number of radial space increments to handle fast transients such as rod ejection accidents. l
- 2. Material properties which are functions of temperature and a 1 sophisticated fuel-to-cladding gap heat transfer calculation, f i
.)
i 3. The necessary calculations to handle post-DNB transients: film boiling I heat transfer correlations, Zircaloy-water reaction and partial melting i of the materials.
i FACTRAN is further discussed in Reference 2.
I LQU.B85 f
j The LOFTRAN program is used for studies of transient response of a PWR system l to specified pert % ations in process parameters. LOFTRAN simulates a i multiloop system by a model containing reactor vessel, hot and cold leg l piping, steam generator (tube and shell sides) and the pressurizer. The 1
4 pressurizer heaters, spray, and relief and safety valves are also considered l in the program. Point model neutron kinetics, and reactivity effects of the moderator, fuel, boron, and rods are included. The secondary side of the l
j steam generator utilizes a homogeneous, saturated mixture for the thermal j transients and a water level correlation for indication and control. The
] Reactor Protection System is simulated to include reactor trips on high 4
neutron . flux, overtemperature AT, overpower AT, high and low pressurizer
{ pressure, low flow, and high pressurizer level. Control systems are also j simulated including rod control, steam dump, feedwater control, and j pressurizer pressure control. The Emergency Core Cooling System,. including l the accumulators and upper head injection, is also modeled.
1 14-9 June 1992 i -
~ _ . ___ _ _ _ _ _ _. _ - ~ - - . _ - --
1 1
i .
d i
j.
j l LOFTRAN is a versatile program which is suited to both accident evaluation and j control studies as well as parameter sizing.
1 i LOFTRAN also has the capability of calculating the transient value of DNBR based on the input from the core limits illustrated in Figure 14-1. The core
! limits represent the minimum value of DNBR as calculated for typical or thimble cell.
f LOFTRAN is further discussed in Reference 3.
i
! TWIKLE L
! The TWINKLE program is a multi-dimensional spatial neutron kinetics code, I which is patterned after steady state codes presently used for reactor core
! Sesign. The code uses an implicit finite-difference method to solve the two- .
group transient neutron diffusion equations in one, two and three dimensions. l l
The code uses six delayed neutron groups and contains a detailed multi-region )
fuel-cladding-coolant heat transfer model for calculating pointwise Doppler l 3 900 >
j and moderator feedback effects. The code handles up t % patial points, and perfoms its own steady state initialization. Aside from basic cross-section data and thermal-hydraulic parameters, the code accepts as input basic -
driving functions such as inlet temperature, pressure, flow, boron concentration, control rod motion, and others. Various edits are provided, e.g., channelwise power, axial offset, enthalpy, volumetric surge, pointwise power, and fuel temperatures. !
l The TWINKLE code is used to predict the kinetic behavior of a reactor for I transients which cause a major perturbation in the spatial neutron flux distribution.
l TWINKLE is further discussed in Reference 4. l E
The THINC Lode is described in References 47 and 48, of Section 3.2.2.
14-10 June 1992
1 1
i I l 90 GNAdQE References - Section 14 -
l 1. Friedland, A. J., Ray, S., " Revised Thermal Design Procedure," WCAP-
- 11397, February 1987.
- 2. Hargrove, H. G., 'FACTRAN - A Fortran-IV Code for Thermal Transients in a U0, Fuel Rod," WCAP-7908, June 1972.
- 3. Burnett, T. W. T., et al., "LOFTRAN Code Description," WCAP-7907, June j 1972.
j 4. Risher, D. H. Jr. and Barry, R. F., " TWINKLE - A Nulti-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary), and WCAP-8028-A (Non-Proprietary), January 1975.
4 5
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(
Table 14-1 Summary of initial Conditions and Computer Codes Used ,
Reedor Vessel Vessel Average RCS Pressure, psia Aoddent CcImpider DNB RTDP InidalCore Power.
% of 1518.5 MWL Coolant Flow, gpm Coolant Temp.*F Codes used Coneladon 0 81880 547 1970 Unconisaged RCCA Wilhdrawallsom a TWBELE W-3 No Subesiucal Condson FACTRAN WRB-1 THINC 100 181800 573.9 and 557 2000arxi 2250 Uncontsosed RCCA Wlhlrewal et Power LOFTRAN WAB-1 Yes 80 563.1 and 553 2000 and 2250.
10 549.7 and 548 2000 and 2250 100 181600 573.9 2500 RCCA Dsop LOFTRAN WRB-1 Yes 100(power) NA 573.9 (power) 1970 (power)
Chemical and Volume Conisol NA NA NA 5483 (stadup) 1970 (stadup)
Syelem Medundlon 5 (eterke) 140(refuebng) 15 (refusing) 0 (reluoIng) 12 81400 550.2 1970 Stadupof anhadive LOFTRAN NA No Reedor Coolert Loop FACTRAN THB4C 100 181800 573.9 2000 pankneviin Fe Endialpy NA NA NA 100 181800 573.9 2000 Fnmaalve Lead Inaeems LOFTRAN WRB-t Yes 100 188800 573.9 2000 and 2250 Lees ellandfruddne Tsip LOFTRAN WRB-1 Yes ,
l~T8,000 5774 aunt Suf.1 2Z80 Lossof NosmalFeedwated LOFTRAN NA MO t O 2.
m .u w . -AC 0 178000 547 2000 Steamine lhoek LOFTRAN W4 No 100 181800 573.9 2000 Loss ol Flow LOFTRAN WRB1 Yes FACTRAN 102 178000 577.9 2280 I t arhad Rotor LOFTRAN NA No FACTRAN 81800 547 970 Rod Flacaing TWINKLE NA No O 577.9 FACTRAN l 02- 178000
_ - _ _ _ _ _ ._.___._.-.___._.___.___.__.________.___._.-_______________.__.____..____________._______.__.___m__ _ _ _ . . - _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ . _ _ . . _ _ - . . _ _ _ . - _ - . . _ _ _ . . _ _ . _
I Table 14-2 t Nominal Values of Pertinant Plant Parameters for Non-LOCA Accident Analyses Parameter Max T-ave. RTDP hfa T-ave. non-RTDP hiin T-ave. RTDP Min T-ave. non-RTDP
'Ihermal Output of NSSS (MWI) 1524.5 1524.5 1524.5 1524.5 Maximum Core Power (MWI) 1518.5 1518.5 1518.5 1518.5 Reactor Coctant Flow Per Loop (gpm) 90900 89000 90900 89000 -
Reactor Coolant System Pressure (psia)* 2000 or 2250 2000 or 2250 2000 or 2250 2000 or 2250 Vessel Coolant Average Temperature ('F)** 573.9 573.9 557.0 557.0 Core Inlet Temperature (*F) 545.5 (2000 psia) 545.0 (2000 psia) 528.1 (2000 psia) 527.5 (2000 psia) 545.3 (2250 psia) 544.8 (2250 psia) 527.9 (2250 psia) 527.3 (2250 psia)
Steam Generator Tube Plugging level (%) O to 10 0 to 10 0 to 10 0 to 10 Steam Generator Outlet Pressure (psia) 806 (0% SGTP) 806 (0% SGTP) 686 (0% SGTP) 686 (0% SGTV) 780 (10% SGTP) 780 (10% SGTP) 663 (10% SGTP) 663 (10% SGTP)
Assumed Feedwater Temperature at Steam Generator inlet (*F) 430.0 430.0 430.0 430.0 Average Core licas Flux (Bau/hr-ft2) 185850 185850 185850 185850
- Accident analyses support plant operation at either 2000 psia or 2250 psia.
- Accident analyses support a range of full power T-avg from 557 'F to 573.9 'F.
bi ____.
I Table 14-3 l
. Trip Points and Time Delays to Trip Assumed in Accident Analyses Limiting Trip Point Limiting Trip Point Time Assumed in Analysis Assumed in Analysis Delay Trio Function f.o.r 2000-osia Ooeration for 2250-osia Ooeration (seconds)
Power range high neutron flux, high setting i18 % 118 % 0.5 Power range high neutron flux, low setting 35 % 35 % 0.5 Overtemperature AT Variable see Variable see 6.0" Figure 14-1 Figure 14-1 i
Overpower AT Variable see Variable see 6.0* j Figure 14-1 Figure 14-1 High pressuruer pressure 2250 psia 2425 psia 2.0 Low pressuruer pressure 1775 psia jib 0 psia 2.0 Low reactor coolant flow (from loop flow detectors) 87% loop flow 87% loop flow 1.0 Undervoltage trip 68% nommal 68% nommal 1.5 ** i Turbme Trip NA NA 2.0 Low-low steam generator l water level lO '[e 10*[, 2.0
- Total time delay (including RTD bypass loop fluid G-spes delay effect, bypass loop piping thermal capacity, RTD time response, and trip circuit, channel electmaics delay) from the time the temperature difference in the coolant loops exceeds the trip setpoint until the' rods are free to fall.
- The undervoltage setpoint is not explicitly modeled in the safety analysis. The 1.5-second celay includes the delay for the RCP bus voltage to fall below the setpoint following 'mstantaneous foss of power to the bus.
J
Figure 14-1 (sheet 1 of 2)
Illustration of Overtemperature and Overpower Delta-T Protection For 2000-psia Operation 70 -
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20 560 570 580 590 600 610 620 630 T-avg (deg. F) 4 Core Limits OTDT Setpoint OPDT Setpoint A
Figure 14-1 (sheet 2 of 2)
Illustration of Overtemperature and Overpower Delta-T Protection For 2250-psia Operation 70 -
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/
Figure 14-3 Normalized Reactivity Worth vs. Rod Position 1 ,
0.9 -
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4 4
i Figure 14-4 Normalized Trip Reactivity vs. Time 4
i 1
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No CHMQES' 14.1 CORE AND COOLANT BOUNDARY PROTECTION ANALYSIS 14.1.1 UNCONTROLLED RCCA WITHDRAWAL FROM A SUBCRITICAL CON 0! TION An RCCA withdrawal incident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCAs resulting in a power excursion. Such a transient could be caused by a malfunction of the reactor control or rod control systems. This could occur with the reactor subcritical, at hot zero power or at power. The "at power" case is discussed in Section 14.1.2.
Although the reactor is normally brought to power from a subcritical condition by means of RCCA withdrawal, procedures for the initial startup following refueling call for boron dilution. The maximum rate of reactivity increase in the case of boron dilution is less than that assumed in this analysis (Section 14.1.4).
The rod cluster drive mechanisms are wired into preselected banks, and these bank configurations are not altered during core life. The rods are therefore physically prevented fros' withdrawing in other than their respective banks.
Power supplied to the rod banks is controlled such that no more than two banks can be withdrawn at any time. Additionally, with the Bank Selector Switch in either the Automatic (AUTC) or Manual (MAN) position, the banks can be withdrawn only in their proper withdrawal sequence. The rod drive mechanism is of the magnetic latch type and the coil actuation is sequenced to provide variable speed rod travel. The maximum reactivity insertion rate is analyzed in the detailed plant analysis assuming the simultaneo_u_s wit _hd_rawa_l of the j combination of the two control banks with the maximum combined worth at j i maximum speed.
l The neutron flux response to a continuous reactivity insertion is char-
{ acterized by a very fast rise terminatad by the reactivity feedback effect of j the. negative Doppler coefficient. This self limitation of the power excursion
- is of primary importance since it limits the power to an acceptable level ,
j during the delay time for protective action. Should a continuous RCCA
]- withdrawal accident occur, the transient will be terminated by the following l automatic features of the reactor protection system:
I 14.1.1-1 June 1992 i
. ~_ _ __ . __ _ _ .
. I tJO cH AMES
- 1. Source range high neutron flux reactor trip.
l Actuated when either of two independent source range channels indicates a .
flux level above a preselected manually adjustable setpoint. This trip j function may be manually blocked only after an intermediate range flux l channel indicates a flux level above a specified level. It is I automatically reinstated when both intermediate range channels indicate a flux level below a specified level.
- 2. Intermediate range high neutron flux reactor trip.
Actuated when either of two independent intermediate range channels indicates a flux level above a preselected manually adjustable level. j This trip function may be manually blocked only after two out of four power range channels are reading above approximately 10 percent of full power and is automatically reinstated when three of the four channels !
l indicate a power level below this value.
- 3. Power range high neutron flux reactor trip (low setting).
Actuated when two out of the four power range channels indicate a power 1evel above approximately 25 percent of full power. This trip function may be manually blocked when two out of the four power range channels l
indicate a power level 'above approximately 10 percent of full power and is automatically reinstated only after three out of the four channels indicate a power level below this vafue. ,
4
- 4. Power range high neutron flux reactor trip (high setting).
Actuated when two out of the four power range channels indicate a power i
level above a preset setpoint. This trip function is always active.
- In addition, control rod stops on high intemediate range flux level (one of two) and high power range flux level (one of four) serve to discon-tinue rod withdrawal and prevent the need to actuate the intermediate j range flux level trip and the power range flux level trip, respectively.
14.1.1-2 June 1992
h 1
i 40 CRAAGES l Method of Analysis j
1 i The analysis of the uncontrolled RCCA bank withdrawal from suberitical accident is perfcreed in three stages: first an average core nuclear power transient calculation, then an average core heat transfer calculation, and
! finally the DNBR calculation. The average nuclear power transient with
- respect to time calculation is perfomed using a spatial neutron kinetics code, TWINKLE, which includes the various total core feedback effects, i.e., j Doppler and moderator reactivity. The FACTRAN code is then used to calculate
! the themal heat flux transient, based on the nuclear power transient
] calculated by TWINKLE. FACTRAN also calculates the fuel and cladding
! temperatures. The average heat flux is next used in THINC, References 47 l and 48, (Section 3.2.2) for transient DNBR calculation, l
i Plant characteristics and initial conditions are discussed in Section 14. In j order to give conservative results for a startup accident, the following assumptions are made.
, 1. Since the magnitude of the nuclear power peak reached during the initial f part of the transient for any given rate of reactivity insertion is i, strongly dependent on the Doppler coefficient, conservatively low (lcwest
- absolute magnitude) values are used.
5
- 2. Contribution of the moderator reactivity coefficient is negligible during l the initial part of the transient because the heat transfer time between
- the fuel and the moderator is much longer'than the nuclear flux respo'nse !
- time. However, after the initial nuclear flux peak, the succeeding rate
] of power increase is affected by the moderator reactivity coefficient. A
! highly conservative value is used in the analysis to yield the maximum i peak heat flux.
4
! 3. .The reactor is assumed to be at hot zero. power. This assumption is more j conservative than that of a lower initial system temperature. The higher l initial system temperature yields a larger fuel-water heat transfer coefficient, larger specific heats, and a less negative (smaller absolute magnitude) Doppler coefficient, all of which tend to reduce the Doppler
- 14.1.1-3 June 1992 1
i
'w-
I feedback effect thereby increasing the neutron flux peak. The initial effective multiplication factor is assumed to be 1.0 since this results
'in maximum neutron flux peaking and, thus, the most severe nuclear power transient.
- 4. Reactor trip is assumed to be init'iated by power range flux (low setting). The most adverse combination of instrument and setpoint errors, as well as delays for trip signal actuation and RCCA release, is taken into account. A 10 percent increase is assumed for the power range flux trip setpoint, raising it from the nominal value of 25 percent to 35 percent. Since the rise in the neutron flux is so rapid, the effect of errors in the trip setpoint on the actual time at which the rods are released is negligible. In addition, the reactor trip insertion characteristic is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn position.
- 5. The maximum positive reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two sequential control banks having the greatest combined worth at maximum speed (45 inches / minute).
, 6. The most limiting axial and radial power shapes, associated with having the two highest combined worth sequential banks in their highest worth position, are assumed for DNS analysis.
- 7. The initial power level was assumed to be below the power level expected for any shutdown condition (10d of nominal power). The combination of highest reactivity insertion rate and lowest initial power produces the highest peak heat flux.
- 8. One reactor coolant pump is assumed to be in operation. This lowest initial flow minimizes the resulting DNBR.
Results % khroo IN I+!"3 Figures 14.1.1-1 and 14.1.1-2 show the transient behavior for the uncontrolled RCCA bank withdrawal with the accident terminated by reactor trip at 35 14.1.1-4 June 1992
percent nominal power. The reactivity insertion rate used is greater than that calculated for the two highest worth sequential control banks, both assumed to be in their highest incremental worth region. Figure 14.1.1-1 shows the neutron flux transient.
The energy release and the fuel temperature increases are relatively small.
The thermal flux response, of interest for departure from nucleate boiling Jr. g considerations,isshownon[ Figure 14.1.1-1. l The beneficial effect of the inherent thermal lag in the fuel is evidenced by a peak heat flux less than 19.l.l-~2 g
i the full-power nominal value. The minimum DNBR at all times remains above the J. limit value and there is a high degree of subcooling at all times in the core.
1p Figure 14.1.1-2lshows the response of the hot spot average fuel and cladding
(
~
temperature. The average fuel temperature increases to a value lower than the b nominal full-power value.
The calculated sequence of events for this accident is shown in Table 14.1.1-1. With the reactor tripped, the plant returns to a stable condition.
The plant may subsequently be cooled down further by following normal plant shutdown procedures.
i Conclusion In the event of a RCCA withdrawal accident from the subcritical condition, the core and the reactor coolant system are not adversely affected. The minimum departure from nucleate bciling ratio remains above the limit value and thus, no fuel or clad damage is predicted.
d J
i i
14.1.1-5 June 1992
J i
~
TABLE 14.1.1-1
! TIME SEQUENCE OF EVENTS FOR UNCONTROLLED RCCA WITHDRAWAL FROM l A SUBCRITICAL CONDITION i
, Time of Each Event
- Event (Seconds)
. i
. Initiation of uncontrolled rod withdrawal, 0
) k l00 pcm/second reactivity insertion rate,jfrom 10 ' of nominal power 4 i
Power range high neutron flux low setpoint reached T,Q Peak nuclear power occurs E.(p Roes ,e,in to fa,, i.to co,e em.,
Peak heat flux occurs 10. (a Minimus DNBR occurs 10.(a
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, 3,a l t- ph 0 4 8 12 16 20 2h 26 Time, Seconds TIGUEZ 14.1.1-4
1 I
a j
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10 4 i i 0 5 10 15 20 25 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Uncontrolled RCCA Bank Withdrawal from Subcritical Nuclear Power Transient Figure 14.1.1-1
i 2
b
.7 _
e .s -:
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E -
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0 l l l l 0 5 10 15 20 25 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Uncontrolled RCCA Bank Withdrawal from Subcritical Core Heat Flux Transient Figure 14.1.1-2
a i
4 H.: so : Fu.s c.rnwun.
~~~
Hot Spot Fuel Average 3000 _
m__ _
n -
6 _
s e
200o -- -
E : es e . / 's s
.s. 15o0 -_ s s - s s
g, -
~
s E 1000 ~~
e . '% .
F _
~~~
o . . . .
o, 5 to 15 to 25 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 1 Uncontrolled RCCA Bank Withdrawal from Subcritical Hot-Spot Fuel Temperatures Figure 14.1.1-3
1 r
90 CHA M S 1 14.1.2 i
UNCONTROLLED RCCA WITHDRAWAL AT POWER '
An uncontrolled RCCA withdrawal at power results in an increase in core heat i
flux. Since the heat extraction from the steam generator remains constant, !
1 there is a net increase in reactor coolant temperature. Unless terminated by
{
manual or automatic action, this power mismatch and resultant coolant '
temperature rise would eventually result in DNB. Therefore, to prevent the possibility of damage to the cladding, the Reactor Protection System is designed to terminate any such transient with an adequate margin to DNB.
The automatic features of the Reactor Protection System which prevent core damage in a rod withdrawal accident at power include the following:
1
- 1. Nuclear power raage instrumentation actuates a reactor trip if two out of the four channels exceed an overpower setpoint.
- 2. Reactor trip is actuated if any two out of four AT channels exceed an overtemperature AT setpoint. This setpoint is automatically varied with power distribution, temperature and pressure to protect against DNB. l
- 3. Reactor trip is actuated if any two out of four AT channels exceed an overpower AT setpoint. This setpoint is automatically varied with temperature to ensure that the allowable full power rating is not exceeded.
- 4. A high pressure reactor trip, actuated from any two out of three pressure channels, is set at a fixed point. This set pressure will be less than the set pressure for the pressurizer safety valves.
- 5. A high pressurizer water level reactor trip, actuated from any two out of three level channels, is actuated at a fixed setpoint. This affords additional protection for RCCA withdrawal accidents.
i The manner in which the combination of overpower and overtemperature AT trips provide protection over the full range of reactivity insertion rates is illustrated in Section 14. Figure 14-1 represents the possible conditions of reactor vessel average temperature and AT with the design power distribution 14.1.2-1 June 1992
l l
1 r I go cRAd4E.5
+
in a two-dimensional plot. The boundaries of operation defined by the overpower AT trip and the overtemperature AT trip are represented as
" protection lines' on this diagram. These protection lines are drawn to .
include all adverse in:trumentation and setpoint errors, so that under nominal conditions trip would occur well within the area bounded by these lines. A maximum steady state operating condition for the reactor is also shown on the i figure.
The utility of the diagram just described is in the fact that the operating limit imposed by any given DNB ratio can be represented as a line.on this coordinate system. The DNB lines represent the locus of conditions for which the DNBR equals the limit value (1.33 for the thimble cell and 1.33 for the typical cell). All points below and to the left of this line have a DNS ratio greater than this value. The diagram shows that DNB is prevented for all cases if the area enclosed within the maximum protection lines is not traversed by the applit.able DNB ratio line at any point.
) The region of permissible operation (power, pressure and temperature) is completely bounded by the combination of reactor trips: nuclear overpower (fixed setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint); overpower and overtemperature AT (variable setpoints). These trips 4
are designed to prevent overpower and a DNB ratio of less than the limit value, j Method of Analysis j -
- Uncontrolled rod cluster control assembly bank' withdrawal is analyzed by the l LOFTRAN code. This code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, '
steam generator, and steam generator safety valves. The code computes pertinent plant variables, including temperatures, pressures, and power level.
The core limits, as illustrated in Figure 14-1, are used as input to LOFTRAN to determine the minimum departure from nucleate boiling ratio during the transient. This accident is analyzed with the Revised Thermal Design Procedure as described in Reference 49, Section 3.2.2. Plant characteristics and initial conditions are discussed in Section 14.
14.1.2-2 June 1992
l l
l l
l
)
In order to obtain conservative values of departure from nucleate boiling ratio, the following assumptions are made: !
l-l A.cc
- 1. Initial Conditions - Initial reactor power, reactor coolant average temperatures, and reduced reactor coolant pressure (2000 psia) are
( '
assumed to be at their nominal values. Uncertainties in initial M uA* I conditions are included in the limit DNBR as described in Reference l 49, Section 3.2.2. I I
- 2. Reactivity Coefficients - Two cases are analyzed.
- a. Minimum Reactivity Feedback - A positive (5 pcm/*F) moderator coefficient of reactivity is assumed, corresponding to the beginning of core life. A variable Doppler power coefficient with core power is used in the analysis. A conservatively small (in absolute magnitude) value is assumed.
- b. Maximum Reactivity Feedback - A conservatively large positive moderator density coefficient and a large (in absolute ;
magnitude) negative Doppler power coefficient are assumed. l
- 3. The rod cluster control assembly trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position. l l
- 4. The reactor trip on high neutren flux is assumed to be actuated at a
~
conservative value of 118% of nominal full power. The overtemperature AT trip includes all adverse instrumentation and setpoint errors; the delays for trip actuatien are assumed to be the maximum values. No credit was taken for the other expected trip ]
functions.
- 5. The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the combination of the two control banks having the maximum combined worth at maximum speed.
14.1.2-3 June 1992
f 1
The effect of rod cluster control assembly movement on the axial core power distribution is accounted for by causing a decrease in the overtemperature AT
, trip setpoint proportional to a decrease in margin to DNB.
! Results
' r
, Figures 14.1.2-1 and 14.1.2-2 show the response of neutron flux, pressure ,
average coolant temperature, and departure from nucleate boiling ratio to a 4
rapid rod cluster control assembly withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the start of l the accident. Since this is rapid with respect to the thermal time constants j- of the plant, small changes in T , and pressure result, and a large margin to DNB is maintained. Lab The response of neutron flux, pressure, average coolant temperature, and DNBR l for a slow control rod assembly withdrawal from 10% power is shown in Figures
{ 114.1.2-3 and 14.1.2-4. Reactor trip on overtemperature AT occurs after a j longer period, and the rise in temperature and pressure is consequently larger -
than for rapid rod cluster control assembly withdrawal. Again, the minimum DNBR is greater than the limit value, i.
I Figure 14.1.2-5 shows the minimum departure from nucleate boiling ratio as a 1
function of reactivity insertion rate from initial full-power operation for I the minimum an' d maximum reactivity feedback cases. It can be seen that two
{ reactor trip channels provide protection over the whole range of reactivity
- insertion rates. These are the high neutron flux and overtemperature AT trip channels. The minimum DNBR is never less than the limit value.
l -
Figures 14.1.2-6 and 14.1.2-7 show the minimum departure from nucleate boiling ratio as a function of reactivity insertion rate for rod cluster control assembly withdrawal incidents starting at 60% and 10% power respectively. The A results are similar to the 100% power case, except that as the initial power is decreased, the range over which the overtemperature AT trip is effective is increased. In neither case does the departure from nucleate boiling ratio t
fall below the DNBR limit value.
a i 14.1.2-4 June 1992 1
-e-
l In the referenced figures, the shape of the curves of minimum departure from i
nucleate boiling ratio versus reactivity insertion rate is due both to reactor core and coolant system transient response and to protection system action in initiating a reactor trip. I
'Figoc< 14.1.1-5 W E D y -
Referring to tb .... r_::t i . ' t, '; i;'. :n; :: ":,-.- c.:.: ,, for example, it is noted that: l I
- 1. For high reactivity insertion rates (i.e., between -100 pcm/ second l and -15 pcm/second), reactor trip is initiated by the high neutron flux trip. The neutron flux level in the core rises rapidly for these insertion rates, while core heat flux and coolant system temperature lag behind due to the thermal capacity of the fuel and coolant system fluid. Thus, the reactor is tripped prior to significant increase in heat flux or water temperature with resultant high minimum departure from nucleate boiling ratios during the transient. Within this range, as the reactivity insertion rate decreases, core heat flux and coolant temperatures can remain more nearly in equilibrium with the neutron flux; minimum DNBR during the transient thus decreases with decreasing insertion rate.
- 2. With further decrease in reactivity insertion rate, the over-temperature AT and high neutron flux trips become equally effective in terminating the transient.
i l
The overtemperature AT reactor trip circuit initiates a reactor trip when measured coolant trip AT exceeds a setpoint based on measured reactor coolant system average temperature and pressure. It is important in this context to note, however, that the average temperature contribution to the circuit is lead-lag compensated in order to decrease the effect of the themal capacity of the reactor coolant system in response to power increases. g{ {
For reactivity insertion rates between 15 pcm/second and -7 pcm/
second, the effectiveness of the overtemperature AT trip increases (in tems of increased minimum departure from nucleate boiling ratio) due to the fact that, with lower insertion rates, the power 14.1.2-5 June 1992
4 increase rate is slower, the rate of rise of average coolant temperature is slower, and the system lags and delays become less significant.
+$
- 3. For reactivity insertion rates less than /second, the rise in reactor coolant temperature is sufficiently high so that the steam generator safety valve setpoint is reached prior to trip. Opening these valves, which act as an additional heat load on the reactor
, coolant system, sharply decreases the rate of rise of reactor coolant system average temperature. This causes the overtemperature l AT trip setpoint to be reached later with resulting lower minimum !
departure from nucleate boiling ratios. l res 14.1.2-5, 14.1.2-6, and 14.1.2-7 illustrate minimum departure from I i
nucleate boiling ratio calculated for minimum and maximum reactivity feedback.
The calculated sequence of events for this accident is shown in Table
[
14.1.2-1.
2 Conclusions l
4 In the unlikely event of an at power (either from full power or lower power levels) control rod bank withdrawal incident, the core and reactor coolant I l system are not adversely affected since the minimum value of DNB ratio reached is in excess of the DNB limit value for all rod reactivity rates. Protection is provided by nuclear flux overpower and overtemperature AT. Additional 4 protection would be provided by the high pressurizer level, overpower AT, and the high pressure reactor trip. The preceding sections have described the
} effectiveness of these protection channels.
1 14.1.2-6 June 1992
Insert "A"
- 1. Initial Conditions - Cases are analyzed for all combinations of the following initial conditions: three initial power l
levels (100%, 60%, and 10%); high and low pressure operation; and minimum and maximum nominal RCS average temperature. Uncertainties in the initial conditions are included in the limit DNBR as described in Reference 49, Section 3.2.2.
Insert "B" Figure 14.1.2-1 shows the response of neutron flux, DNBR, pressurizer pressure, pressurizer water volume, vessel T-avg. and vessel AT to a rapid rod cluster control assembly withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the start of the accident. Since this is rapid with respect to the thermal time constants of the plant, small changes in T-avg and pressure result, and a large margin to
.I DNB is maintained.
The response of neutron flux, DNBR, pressurizer pressure, pressurizer water. volume, vessel T-avg, and vessel AT for a slow control rod assembly withdrawal from 10% power is shown in 1
{
1 Figure 14.1.2-2. Reactor trip on overtemperature AT occurs after a longer period, and the rise in temperature and pressure is consequently larger than for rapid rod cluster control assembly withdrawal. Again,.the minimum DNBR is greater than the limit I l
value.
l l
Figure 14.1.2-3 shows the minimum departure from nucleate boiling l
ratio as a function of the reactivity insertion rate for the three initial power levels (100%, 60%, and 10%), minimum and maximum
, reactivity feedback, high and low pressure operation, and minimum and maximum nominal RCS T-avg. It can be seen that the high neutron flux and overtemperature AT trip channels provide
}.
protection over the whole range of reactivity insertion rates.
The minimum DNBR is never less than the limit value.
j 4
4 TABLE 14.1.2-1 l
, TIME SEQUENCE OF EVENTS FOR
, UNCONTROLLED RCCA WITHDRAWAL AT POWER 4
Time of Each Event
- f11D1 (Seconds)
- Case A:
l l
Initiation of uncontrolled rod cluster 0 I
control assembly withdrawal at full power and maximum reactivity insertion rate 1 (100 pcm/sec) i I Power range high neutron flux high trip point reached /tb Rods begin to fall into core /, h Minimum departure from nucleate boiling ratio occurs kf Case B:
l 4
Initiation of uncontrolled rod cluster control 0 assembly withdrawal at 10% power and at a small reactivity insertion rate (3 pcm/sec)
Overtemperature AT reactor trip signal initiated f!
Rods begin to fall into core ll he 4
Minimum departure from nucleate boiling ratio occurs 157 f[f f
}
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Point Beach Nuclear Plant Units 1 and 2 Rod Withdrawal at Power 100% Power, Minimum Feedback 100 pcm/second Figure 14.1.2-1 (sheet 1 of 3)
l 2500 2400 --
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, 1100 -
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Point Beach Nuclear Plant Units 1 and 2 Rod Withdrawal at Power 100% Power, Minimum Feedback 100 pcm/second Figure 14.1.2-1 (sheet 2 of 3)
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Point Beach Nuclear Plant Units 1 and 2
. Rod Withdrawal at Power 100% Power, Minimum Feedback 100 pcm/second Figure 14.1.2-1 (sheet 3 of 3)
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- a f f i 4 i 6 i 0 20 40 00 80 100 120 140 100 180 200 Time (secondils) 5 _
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- 1 i i , . . .f i
. . t.
0 20 40 00 to 100 120 140 180 100 200 Time (seconda)
Point Beach Nuclear Plant Units 1 and 2 Rod Withdrawal at Power 10% Power, Minimum Feedbac.'
3 pcm/second Figure 14.1.2-2 (sheet 1 of 3)
asco n
j 230o --
~
& m0--
w -
3 e 2100 --
E 2000
/
n.
n.
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t 1o00 -:
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Point Beach Nuclear Plant Units 1 and 2 Rod Withdrawal at Power 10% Power, Minimum Feedback 3 pcm/second Figure 14.1.2-2 (sheet 2 of 3)
l I
l i
- 800 E -
s 570 -- )
2 4
w seo --
l a -
e ss0 --
e 540 --
e -
e e 530 --
m- -
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500 l l l l l l ;
l l O M 40 to 80 100 120 140 160 180 200 Time (seconds) a0 n--
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a _
e
- 3. 50 - -
Q -
e
= 40 - -
b -
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e e -
e
> 20 --
10 -~
' ' i . . , . , , , , , , ,
o i i i i i . , , ,
0 20 40 80 to 100 120 140 100 100 200 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Rod Withdrawal at Power 10% Power, Minimum Feedback 3 pcm/second Figure 14.1.2-2 (sheet 3 of 3)
Rod Withdraw:.1 ct Pcwer 100% Power, Minimum Feedback I
Figure 14.1.2-3 (Sheet 1 of 6) 2.8 2.6 -
I 2.4 -
CC .
o 2.2 -
~
E - - - - - - ' ' ' ' -
E 2 -
.E -
l.8
..a......---....
1.6 -
I . . I ,,,1 i I ,,,1 1 , t . ...
g,4 0.1 0.2 0.5 1 2 5 10 20 50 100 Reactivity Insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg
___.._.__-_.-.____.__---_--_____.x_ - . _ - - _ . _
1 Rod Withdrawal at Power 100% Power, Maximum Feedback Figure 14.1.2-3 (Sheet 2 of 6) 2.8 2.6 -
2.4 - - ~
oc 's g s
, 2.2 ..................
g .... s , __
........ s B /
2 .....,,,,..\
.g e
.......,. . , ~.
... \....................
.........,,,,- I i
. - . , ~
..-..-.....~._,~...I .......,.....
,,g - .,_,_...
u
~. , , . .
1.6 ' -
~.*,, -
l.4 ' '
O.1 0.2 0.5 1 2 5 10 20 50 100 Reactivity insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg
Rod Withdrawal at Power 60% Power, Minimum Feedback Figure 14.1.2-3 (Sheet 3 of 6) 4 3.8 -
3.6 -
3.4 -
3.2 -
P g 3 -
A 2.8 -
B 8
g 26 -
.g g1A 2.2 ;-.=...............................,'
=.
.~ n ' - -
2 1.8 ~
1.6 -
~._.
' ' ' ' ' ' ' ' ' ' ' ' l I ' ' ' ' ' '
1.4 '
O.1 0.2 0.5 1 2 5 10 20 50 100 Reactivity Insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg a.-._ - . . . _ _ --
Rod Withdrawal at Power 60% Power, Maximum Feedback Figure 14.1.2-3 (Sheet 4 of 6) 4 3.8 -
3.6 -
3.4 ~""""""" """ _
~7 32 ?...'..-
d.
p 4 3 ---- .
..s..sN...~s.
C -
2 .8 -
.'\
s .
- g 26 7 ,s\
3 2 ._ - -
,,,...s ' ~ ~
.4
. ... ~
2.2 -
' s 2 -
' s.s s, 1.8 -
1.6 -
~
l . I . ,,,1 1 I , ,,,1 g,4 1 , , I , ,,,
0.1 0.2 0.5 1 2 5 10 20 50 100 Reactivity Insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg
Rod Withdrawtl ct PI,wer 10% Power, Minimum Feedback Figure 14.1.2-3 (Sheet 5 of 6) 4 3.8 .-
3.6 ?
3.4 r 3.2 }-
@3 2'8 : -
E .
a 2.6 ,_ ,............................f'.'_.~.,J. s........
.E g . , - c, .....,
.- 2 4 _..._._. ~.
y ~.. ~ .
- ...'s ..
2.2 ~.s4 . . su . s "*
s N. #. .
2 _ ., , .
, s.
1.8 l.6 ~
t i I . , , , I I i . 1 . . . .
g4 ,
I 2 3 5 10 20 30 50 100 Reactivity Insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg
Rod Withdrawal at Power 10% Power, Maximum Feedback Figure 14.1.2-3 (Sheet 6 of 6) 4 3.8 ,
3.6 -
3.4 -
3.2 ~
ct 3 -
a 2., ;_ ..---....
~, ....
g .- ~
a *.
,e 2.6 ......... ................ ...... ................................ ........... ... ................ **.
.a 2.4
,e 2.2 -
2 -
1.8 r l.6 -
, , , , i , , , I , , , ,
y,4 5 10 50 100 Reactivity Insertion Rate (pcm/second) low pressure, high T-avg low pressure, low T-avg high pressure, high T-avg high pressure, low T-avg
} !
10 CR/bl9ES I
i j 14.1.4 CHEMICAL AND VOLUME CONTROL SYSTEM MALFUNCTION s 1 I
j Positive reactivity can be added to the core with the Chemical and Volume Control System by feeding reactor makeup water into the Reactor Coolant System via the reactor makeup control system. The normal dilution procedures call for a limit
' )
on the rate and magnitude for any individual dilution, under strict '
administrative controls. Boron dilution is a manual operation. A boric acid blend system is provided to pemit the operator to match the concentration of
- reactor coolant makeup water to that existing in the coolant at the time. The Chemical and Volume Control System is designed to limit', even under various postulated failure modes, the potential rate of dilution to a val 6e which, after indication through alarms and instrumentation, provides the operator sufficient time to correct the situation in a safe and orderly manner.
There is only a single, common source of reactor makeup water to the reactor l coolant system from the reactor makeup water storage tank, and inadvertent dilution can be readily terminated by isolating this single source. The operation of the reactor makeup water pumps which take suction from this tank provides the only supply of makeup water to the reactor coolant system. In order for makeup water to be added to the reactor coolant system, the charging pumps must be running in addition to the reactor makeup water pumps.
The rate of addition of unborated water makeup to the reactor coolant system is limited to the capacity of the CVCS pumps. This limiting addition rate is l 181.5 gpm. For totally unborated water to be delivered at this rate to the reactor coolant system at pressure, three charging pumps must be operated.
Normally one charging pump is operated in manual and one pump is operating in the automatic mode, responding to pressurizer level changes.
The boric acid from the boric acid tank is blended with the reactor makeup water in the blender and the composition is detemined by the preset flow rates of boric acid and reactor makeup water on the reactor makeup control. Two separate operations are required. First, the operator must switch from the automatic makeup mode to the dilute mode. Second, the start button must be depressed.
Omitting either step would prevent dilution. This makes the possibility of inadvertent dilution very small.
14.1.4-1
)
i I i
j cRAdQ ES,
[ 90 Information on the status of the reactor coolant makeup is continuously available to the operator. Lights are provided on the control board to indicate the !
l
" operating condition of pumps in the chemical and volume control system. Alarms are actuated to warn the operator if boric acid or domineralized water flow rates deviate from preset values as a result of system malfunction. An additional
.' alarm is available to warn the operator of a potential dilution condition.
. 4 4
l To cover all phases of plant operation, boron dilution during refueling, startup, and power operation are considered in this analysis. i Method of Analysis and Results
. Dilution Durino Refuelina During refueling the following conditions exist:
1.
One residual heat removal pump is running to ensure continuous mixing in the reactor vessel, 2.
The valve in the seal water header to the reactor coolant pumps is closed, j
- 3. The valves on the suction side of the charging pumps are adjusted for addition of concentrated boric acid solution.
4.
The boron concentration of the refueling water is at least 1800 ppe, correspondig to a shutdown margin at least of 5% Ak with all control rods in; periodic sampling ensures that this concentration is maintained, and
- 5. Neutron sources are installed in the core and BF, detectors connected to instrumentation giving audible count rates are installed to provide direct monitoring of the core. Currently, neutron source assemblies are not >
utilized in Unit I and Unit 2.
A minimum active water volume in the reactor coolant system of 1971 ft* is considered. This corresponds to the volume necessary to fill the reactor vessel up to the midplane of the nozzles plus the volume of one RHR train. This ensures mixing via the residual heat removal loop.
The maximum dilution flow of 121 gpm and uniform mixing are also considered.
Administrative procedures limit the charging flow available during this condition. The maximum dilution flow assumes the single failure, such that two pumps are delivering maximum flow. The actual amount of reactor makeup water 14.1.4-2 June 1993
delivered to the suction of the charging pumps would be determined by the l position of FCV-111 which is normally set at no more than 40 gpm. At the full open position, FCV-111 would pass approximately 100 gpm.
i I
The operator has prompt and definite indication of any boron dilution from the l audible count rate instrumentation. High count rate is alarmed in the reactor containment and the main control room. The count rate increase is proportional to the inverse multiplication factor.
The boron concentration must be reduced from 1800 ppm to approximately 1400 ppm before the loss of all shutdown margin. This would take at least 30.1 minutes.
This is ample time for the operator to recognize the audible high count rate signal and isolate the reactor makeup water source by closing valves and stopping the reactor makeup water pumps.
Dilution Durino Cold Shutdown
~} This analysis of the boron dilution event used conservative RCS volumes (i.e.,
A reduced or effectively reduced RCS volumes for the applicable primary system I flowpath) . Mixing of the diluting water (boron free) and the RCS water was j assumed to take place in the vessel downcomer and proceed in a " wave front fashion" through the rest of the RCS. Specific calculations were performed for each of the applicable primary flowpaths (i.e., reduced or effectively reduced j RCS volumes). These calculations determined the boron concentration as a function of time at the core inlet. The boron concentration as a function of time for the two conditions of reduced RCS volume were compared. The most i limiting condition is determined by the lowest concentration achieved by diluting the reduced RCS volume for 15 minutes. The reduced RCS volume, where the primary system is in a half-pipe condition and on residual heat removal, is the resulting limiting condition.
A parametric study for the limiting condition was performed to determine concentration at a function of time. The parameters, which were variables in the parametric study, are:
i
- 1. number of charging pumps - 1, 2, or 3
- 2. number of RHR pumps - 1 or 2 14.1.4-3
i 4
- 3. initial boron concentration - 2000 ppa through 200 ppe in 100 ppa increments.
The number of charging pumps affects the dilution rate and the RCS flowrate. The number of RHR pumps affects the RCS flowrate and the RCS volume. The RCS volume
' is affected by the number of RHR pumps because.with one RHR pump, only one RHR heat exchanger is used. Whereas with two RHR pumps, two RHR heat' exchangers are used. The increase in volume from one RHR pump to two RHR pumps results in the fact that the transit time of the wavefront around the RCS flowpath is greater than one-half the transit time when using only one RHR pump. The initial baron concentration affects the change in concentration (i.e., initial concentration minus concentration after 15 minutes) because dilution is a non-linear function.
The parametric study was performed using all combinations of the three variables.
The change in concentration was computed for all cases. The concentration change was multiplied by the differential boron worth. The boron worth was a value derived from the nuclear design reports for the Point Beach Nuclear Plant Units 1 and 2. The resulting product is the reactivity addition for 15 minutes of dilution at the specific parameter values. The results of the parametric study '
show that the dilution event with one RHR pump is slightly more limiting than with two RHR pumps.
This analysis was performed to determine the shutdown margin which is necessary to prevent criticality from an inadvertant boron dilution event with a reduced RCS volume and a duration of 15 minutes. The shutdown margin which is necessary to prevent criticality is equal to reactivity addition from dilution or 1% Ak/K whichever is greater. The required shutdown margin is illustrated in Figure 14.1.4-1. This figure shows the required shutdown margin versus initial concentration for 1, 2, and 3 charging pumps in operation. Also shown on this figure is the most limiting BOL shutdown margin versus concentration.
In order to use this figure, one determines the required shutdown margin based on the RCS boron concentration and the number of charging pumps in use. Then, one detensines the actual shutdown margin using the most limiting BOL shutdown margin versus concentration or some other appropriate means available to the operator. If the required shutdown margin is less than the actual shutdown margin, then criticality cannot occur for at least 15 minutes during a dilution.
14.1.4-4
I
\
' j l
i f
j As an example, if the baron concentration is 1300 ppe and one charging pump is l' to be operational, the required shutdown margin is 1.062k/K and the actual shutdown margin is 1.87Ek/K. Therefore, criticality cannot occur within 15 minutes of the initiation of dilution. However, if boron concentration were 1300 )
ppa and two charging pumps are to be operational, then the required shutdown j margin is 2.01Mk/K and the actual shutdown margin is 1.87Ek/K. Therefore, subcritical conditions will not be maintained during 15 minutes of diluting.
i It is concluded, from Figure 14.1.4-1, that a prudent practice would be to limit charging operation to one charging pump in service while the plant is in a reduced volume condition and provide an alarm to alert the operator that charging has begun and a potential dilution event is in progress. !
The above conclusion is implemented by administratively (procedurally) limiting charging operations to one pump in service and by installing a limit switch on the valve for the reactor water makeup pump. The limit switch will activate an alarm in the control room on the alars status board whenever the valve is not ,
closed. The alarm light has a message similar to " POTENTIAL DILUTION IN ,
PROGRESS". This resolution will provide substantially more than 15 minutes warning to the operator. This is because it will take approximately five minutes for the wavefront to reach the core inlet after charging has been initiated and the alarm has been sounded.
Dilution Durino Startuo I
Prior to refueling, the reactor coolant system is filled with borated water from the refueling water storage tank. Core monitoring is by external BF, detectors.
Mixing of reactor coolant is accomplished by operation of the reactor coolant pumps. Again the maximum dilution flow (181.5 gpa) is considered. The volume of reactor coolant is approximatelkS179 ft*'which is the volume of the reactor coolant system excluding the pressurizerQhe volume has been calculated taking into account steam generator tube plugging. High source level and all reactor trip alarms are effective. d SilVy &, 3 The minimum time required to reduce the reactor coolant boron concentration to 1600 ppa, where the reactor could go critical with all rods at the insertion limits,isabout]l8Jainutes. Once again, this should be more than adequate 14.1.4-5
{
l !
1
! I
- i i,
j
. time for operator action to the high count rate signal, and temination of
{ dilution flow.
j In any case, if continued dilution occurs, the reactivity insertion rate and i
consequences thereof are considerably less severe than those associated with the
{ uncontrolled rod withdrawal analyzed in Section 14.1.1, uncontrolled RCCA j Withdrawal from a Subcritical Condition.
l j Dilution at Power l For dilution at power, it is necessary that the time to lose shutdown margin be
! sufficient to allow identification of the problem and temination of the
~
i
{ dilution.
As in the dilution during startup case, the RCS volume reduction due j
to steam generator tube plugging ts considered.
The effective reactivity l
j addition rate is a function of the reactor coolant temperature and boron !
concentration.
The reactivity insertion rate calculated is based on a i
i conservatively high value for the expected boron concentration at power (1500 ppm) as well as conservatively high charging flow rate capacity (181.5 gpe). The
! reactor is assumed to have all rods at the insertion limits in either automatic !
or manual control. With the reactor in manual control and no operator action to terminate the transient, the power and temperature rise will cause the reactor jy.]
to reach the reactor protection (i.e., OTAT, high nuclear flux) trgetpoint resulting in a reactor trip. After reactor trip there is at least for operator action prior to return to criticality. The boron diluti]17.0 minutes on transient in this case is essentially the equivalent to an uncontrolled rod withdrawal at power.
The maximum reactivity insertion rate for a boron dilution transient is conservatively estimated rates analyzed for uncontrol w to (led rod / ithdpcm/sec and'is rawal at power. Prior to reaching the reactor protection trip, the operatol will have received an alam on overtemp-erature AT and turbine runback. d
- 8 With the reactor in automatic control, a boron dilution will result in a power and temperature increase such that the rod controller will attempt to compensate by slew insertion of the control rods. This action by the controller will result in rod insertion limit and axial flux alams. The minimum time to lose the 1 percent AK shutdown margin required at beginning-of-life would be greater than r h einutes. The time would be significantly longer at and of life due to the low initial boron concentration and 2.77 percent AK shutdown margin.
d-4 M.1 14.1.4-6 June 1991
i i
e 40 C4A44Es f 4
Conclusions i
Because of the procedures involved in the dilution process, an erroneous dilution 4
is not considered credible. Nevertheless, if an unintentional dilutien of boron 9
in the reactor coolant does occur, numerous alarms and indications are available i
to alert the operator to the condition. The maximum reactivity addition due to
! the dilution is slow enough to allow the operator to determine the cause of the addition and take corrective action before the required shutdown margin is lost.
l l
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14.1.4-7
1 l
fNoCHAM m .
14.1.7 EXCESSIVE LOAD INCREASE INCIDENT An excessive load increase incident is defined as a rapid increase in steam generator steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The reactor control system is designed to accomodate a 105 step load increase and/or a 55 per minute ramp load increase (without a reactor trip) in the range of 15 to 100% full power.
Any loading rate in excess of these values may cause a reactor trip actuated by the reactor protection system. If the load increasa exceeds the capability of the reactor control system, the transient is te'ainated in time to prevent
, DNBR less than the limiting value, by a combinat'on of the nuclear overpower trip and the overpower-overtemperature AT trips, as discussed in Section 7.
An excessive load increase incident could result from either an administrative violation such as excessive loading by the operator or an equipment malfunction such as steam bypass control or turbine speed control.
For excessive loading by the operator or by system demand, the turbine load limiter keeps maximum turbine load below 1005 rated load.
During power operation, steam bypass to the condenser is controlled by reactor coolant condition signals, i.e., abnormally high reactor coolant temperature indicates a need for steam bypass. A single controller malfunction does not l
cause steam bypass; an interlock is provided which blocks the control signal to the valves unless a large turbine load decrease or a turbine trip has occurred.
Method of Analysis This accident is analyzed using the LOFTRAN code. The code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, steam generator safety valves, and feedwater system. The code. computes pertinent plant variables, including temperatures, pressures, and power level.
14.1.7-1 June 1992
_ - . _ ~ - - - - - - - - . . - .
i i
i j Four cases are analyzed to demonstrate the plant behavior following a 10%
[ step-load increase from rated load. These cases are as follows:
1 .
- 1. Reactor control in manual with minimum reactivity feedback.
- 2. Reactor control in manual with maximum reactivity feedback.
- 3. Reactor control in automatic with minimum reactivity feedback.
- 4. Reactor control in automatic with maximum reactivity feedback.
For the minimum reactivity feedback cases, the core has the least negative moderator temperature coefficient (0 pcm/*F) of reactivity and the least negative Doppler only power coefficient; therefore, the least inherent transient response capability. For the maximum reactivity feedback cases, the moderator temperature coefficient of reactivity has its most negative value and the most negative Doppler only power coefficient. This results in the largest amount of reactivity feedback due to changes in coolant temperature. ,
A conservative limit on the turbine valve opening is assumed, and all cases are studied without credit being taken for pressurizer heaters. This accident is analyzed with the Revised Themal Design Procedure as described in Reference 1 Section 14. Plant characteristics and initial conditions are as discussed in Section 14.1. Initial reactor power, pressure, and RCS temperatures are assumed 3 to be at their nominal values. Uncertainties in initial conditions are included g+-
in the limit DNBR, as described in Reference 1, Section 14. ,
~7 Results i 14,1 1 -t a d N.1,7 -3 Nt Figures 14.1.7-1 throuah 14.1.7-4 311ustrate the transient with the reactor in U the manual control mode. For the beginning-of-life case, there is a slight power rQ increase, and the average core temperature shows a large decrease. This results I in a departure from nucleate boiling ratio that. increases above its initial h '
value. For the end-of-life, manually controlled case, there is a much larger -
increase in reactor power due to the moderator feedback. A reduction in departure from nucleate boiling ratio is experienced, but the departure from nucleate boiling ratio remains above the limit value. FiguresQ4.1.7-5through 14.1.7-2 June 1992 l
l y ;'...he illustrate the transient when the reactor is assumed to be in the automatic control mode. Both the beginning-of-life and the end-of-life cases show that core power increases, thereby reducing the rate of decrease in coolant average temperature and pressurizer pressure. For both the beginning-of-life and the end-of-life cases, the rinimum departure from nucleate boiling ratio remains i above the limit value. Yhe calculated sequence of events is shown in Table l 14.1.7-1.
The excessive load increase incident is an overpower transient for which the fuel temperatures rise. When a reactor trip does not occur, the plant reaches a new equilibrium condition at a higher power level corresponding to the increase in steam flow.
Conclusions it has been demonstrated that, for an excessive load increase, the minimum departure from nucleate boiling ratio during the transient will not be below the limit value. j l
l i
14.1.7-3 June 1992
TABLE 14.1.7-1 [
TIME SEQUENCE OF EVENTS FOR EXCESSIVE LOAD INCREASE INCIDENT 1
l l
Q3g Event Time (seconds) i
- 1. Beginning of core life, manual reactor control 10% step lead increase 0 l
Steady-state conditions reached (approximate) 150
- 2. Beginning of core life, automatic reactor control 10% step lead increase 0 Steady-state conditions reached (approximate) 200
- 3. End of core life, manual reactor control 10% step load increase O Steady-state conditions reached (approximate) 100
- 4. End of core life, automatic reactor control 10% step load increase 0 Steady-state conditions 1
reached (approximate) 100 i
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- Excessive Loed incressa EOL, Auto Control 4
Figure 14.1.7 4 i
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14.1.8 LOSS OF REACTOR COOLANT FLOW Nhb Flow Coastdown Accidents A loss of coolant flow incident can result from a mechanical or electrical failure in one or more reactor coolant pumps, or from a fault in the power supply to these pumps. If the reactor is at power at the time of the inci-dent, the immediate effect of loss of coolant flow is a rapid increase in coolant temperature. This increase could result in departure from nucleate boiling (DNB) with subsequent fuel damage if the reactor is not tripped I promptly. The following trip circuits provide the necessary protection against a loss,of coolant flow incident and are actuated by:
- 1. Low voltage on pump power supply bus;
- 2. Pump circuit breaker opening (low frequency on pump power supply bus opens pump circuit breaker); or i
- 3. Low reactor coolant flow.
These trip circuits and their redundancy are futher described in Section 7.2, Reactor Control and Protection System.
Frequency decay for both reactor coolant pumps during full power operation is the most severe credible loss-of-coolant flow condition. For this condition reactor trip together with flow sustained by the inertia of the coolant and rotating pump parts will be sufficient to prevent fuel failure, reactor coolant system overpressure and prevent the DNB ratio from going below the limit value.
Method of Analysit The following loss of flow cases are analyzed:
- 1. Loss of two pumps from a reactor coolant system, heat output of 1518.5 MWt with two loops operating; and
- 2. Loss of one pump from a reactor coolant system, heat output of 1518.5 MWt with two loops operating; and 14.1.8-1 June 1994
- 3. Reactor coolant pump underfrequency event for both pumps with frequency decay rate of 3 Hz/sec, heat output of 1518.5 MWt with two loops operating. (see reference 1)
The third case represents the worst credible coolant flow loss. The first two cases are less severe, with the second case being the least severe. Loss of one pump above 50% of full load is assumed to cause a reactor trip by a low flow signal. For the third case, flow decreases with the frequency to the low reactor coolant pump bus frequency setpoint (57.5 Hz) where the pumps will trip. Reactor trip for the third case is caused by a low flow signal.
The normal power supplies for the pumps are the two buses connected to the generator, each of which supplies power to one of the two pumps. When a generator trip occurs, the pumps are automatically transferred to a bus supplied from, external power lines. Therefore, the simultaneous loss of power to all reactor coolant pumps is a highly unlikely event.
Following any turbine trip, where there are no electrical faults which require tripping the generator from the network, the generator remains connected to the network for approximately one minute. Since both pumps are not on the same bus, a single bus fault would not result in the loss of both pumps.
These transients are analyzed by three digital comr7ter codes. First, the LOFTRAN code is used to calculate the loop :ac core flow during the transient, the time of the reactor trip based on the cal tulated flow, the nuclear power transient, and the primary system pressure and temperature transients. The FACTRAN code is then used to calculate the heat flux transient based on the nuclear power and flow from LOFTRAN. Finally, the THINC code is used to calculate the minimum DNBR during the transient based on the heat flux from FACTRAN and the flow from LOFTRAN.
Initial Ooeratino Conditions I O Initial reactor power, pressure, and RCS temperature are assumed to be at yg i their nominal values (reduced pressure operation). Uncertainties in initial conditions are included in the limit DNBR as described in Reference 49, i
14.1.8-2 June 1994 j
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t'.. f..;....., 9::" r"--+ " '9 _;. .. wicaw. ^ . ;.
f ,:.. :d :f thi: :::ti=
Reactivity Coefficients A conservatively large absolute value of the Doppler-only power coefficient is used. The total integrated Doppler reactivity (power defect) between 0% and 100% power is assumed to be 0.016Ak. 'h: "-t r'r:;x..g rc:-+ "+-d + ha O_
-: ;- g 7 1 . nm.,.en.frici.n+ _e _
The lowest absolute magnitude of the moderator temperature coefficient (5 } ACE-pcm/*F) is assumed, since this results in the maximum core power during the pT b 11 L
initial part of the transient, when the minimum departure from nucleate boiling ratio is reached.
g Flow Coastdown The flow coastdown analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance and the pump characteristics and is based on high estimates of system pressure losses.
No single active failure in the plant systems and equipment which are necessary to mitigate the effects of the accident will adversely affect the consequences of the accident during the transient. A conservatively evaluated overall heat transfer coefficient has been used in the analysis.
Freauency Decay Event Cases The RCS design conditions assumed for the underfrequency event correspond to PBNP operation with an effective (i.e. sleeved and/or plugged) uniform steam generator tube plugging level of up to[13% for Unit I and 14% for Unit 2. /The initial reactor coolant average temperaturefis 573.g*F and initial pressurizer pressure is 2000 psia.
Jo*
/ , Sr Ods I and 2.
14.1.8-3 June 1994 l
Results l Reactor coolant flow coastdown curves for a loss of both pumps are shown on Figure 14.1.8-1. Figures 14.1.8-2 and 14.1.8-3 show the nuclear flux,
! pressurizer pressure, the average channel heat flux, and the hot channel heat flux response for the two pump loss. Figure 14.1.8-4 shows the DNB ratio as a function of time for this case. The minimum WRB-1 DNB ratio is reached 2.9 seconds after initiation of the incident.
Figures 14.1.8-5 through 14.1.8-7 show the transient for loss of one pump with both loops operating and Figure 14.1.8-8 shows the DNB ratio as a function of h
gjh time for.this case. The minimus DNB ratio occurs 3.8 seconds after initiation k b) of the transient. Table 14.1.8-1 sunnarizes the sequence of events for the transient.
The transient response of the RCS for the underfrequency event is shown in Figures 14.1.8-9 through 14.1.8-12. Figure 14.1.8-9 shows the reactor vessel flow coastdown. Figure 14.1.8-10 shows the power fraction and pressurizer -
pressure. Average and hot channel heat flux data is shown in Figure 14.1.8-11. Minimum WRB-1 DNB ratio occurs at 4.6 seconds into the transient (Figure 14.1.8-12). The calculation sequence of events is shown on Table d 14.1.8-2.
Conclusions Since DNB does not occur in any loss of coolant flow incident, there is no cladding damage and no release of fission products into the reactor coolant.
Therefore, once the fault is corrected, the plant can be returned to service in the normal manner. The absence of fuel failures would, of course, be l
verified by analysis of reactor coolant samples. )
l locked Rotor Accident A hypothetical transient analysis is performed for the postulated instan-taneous seizure of a reactor coolant pump rotor. Flow through the reactor I coolant system is rapidly reduced, leading to a reactor trip on a low-flow signal. Following the trip, heat stored in the fuel rods continues to pass 14.1.8-4 June 1994
4 5
into the core coolant, causing the coolant to heat up and expand. At the same time, heat transfer to the shell side of the steam generator is reduced, first because the reduced flow results in a decreased tube side film coefficient and then because the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero upon plant trip).
The rapid expansion of the coolant in the reactor core, combined with the i
reduced heat transfer in the steam generator causes an insurge into the 3 pressurizer and a pressure increase throughout the reactor coolant system.
The insurge into the pressurizer compresses the steam volume, actuates the I automatic spray system, opens the power-operated relief valves, and opens the pressurizer safety valves, in that sequence. The two power-operated relief valves are designed for reliable operation and would be expected to function properly during the accident. However, for conservatism, their pressure-reducing effect is not included in the analysis.
Method of Analysis Two digital computer codes are used to analyze this transient. The LOFTRAN code is used to calculate the resulting loop core and flow transients i following the pump seizure, the time of reactor trip based on loop flow transients, nuclear power following reactor trip, and to determine peak
- pressure. The thersal behavior of the fuel located at the core hot spot is investigated using the FACTRAN code, which uses the core flow and nuclear j
power calculated by LOFTRAN. The FACTRAN code includes a film boiling heat I
transfer coefficient.
4 One case is analyzed: one RCP coasting down, one locked rotor. At the j beginning of the postulated locked rotor accident (i.e., at the time the shaft 4
in one of the reactor coolant pumps is assumed to seize), the plant is assumed
- to be operating at 102 percent of NSSS Thermal Design Power, with maximum 3
steady state pressure and maximum steady state coolant average temperature.
Then, peak pressure is evaluated; the initial pressure is conservatively estimated as 30 psi above nominal pressure (2250 psia) to allow for errors in the pressurizer pressure measurement and control channels. This is done to obtain the highest possible rise in the coolant pressure during the transient.
To obtain the maximum pressure in the primary side, conservatively high loop 14.1.8-5 June 1994
1 l
1 i
l N.I.$-
A pressure drops are added to the. calculated pressurizer pressure. The pressure responseshowninFigure[14.1_.8-18}istheresponseatthepointinthereactor coolant system having the maximum pressure.
Evaluation of the Pressure Transient - After pump seizure, the neutron flux is rapidly reduced by control rod insertion effect. Rod motion is assumed to begin one second after the flow in the affected loop reaches 87% of nominal fl ow. No credit is taken for the pressure-reducing effect of the pressurizer relief valves, pressurizer spray, steam dump or controlled feedwater flow after plant trip.
Although.these operations are expected to occur and would result in a lower peak pressure, an additional degree of conservatism is provided by ignoring their effect.
_ A%
The pressurizer safety valves are full open at 2575 psia, and their capacity .L for steam relief is 288,000 lb/hr. " D, Evaluatien of Decarture from Nucleate Boiline in the Core Durino the Accident
- For this accident, departure from the nucleate boiling is assumed to occur j in the core, and therefore, an evaluation of the consequence with respect to fuel rod thermal transients is performed. Results obtained from analysis of this hot spot condition represent the upper limit with respect to cladding temperature and zirconium-water reaction. In the evaluation, the rod power at the hot spot is assumed to be 2.Qi times the average rod power (F,=2.65) at
~
the initial core power level.
2+ N N 2.fo Film Boilino Coeffit.in.1 - The film boiling coefficient is calculated in the FACTRAN code using the 21 shop-San'd berg-Tong film boiling correlation. The fluid properties are mluated at film tesperature, which is the average between the wall and bulk temperatures. The program calculates the film coefficient at every time step, based on the actual heat transfer conditions l at the time. The neutron flux, system pressure, bulk density, and mass flow rate as a function of time are used as program input.
For this analysis, the initial values of the pressure and the bulk density are used throughout the transient, since they are the most conservative with 14.1.8-6 June 1994
respect to cladding temperature response. For conservatism, departure from nucleate boiling is assumed to start at the beginning of the accident.
Fuel - Claddino Gao Coefficient - The magnitude and the time dependence of the heat transfer coefficient between fuel and cladding (gap coefficient) have a pronounced influence on the thermal results. The larger the value of the gap coefficient, the more heat is transferred between the pellet and the cladding.
Based on investigations of the effect of the gap coefficient on the maximum cladding temperature during the transient, the gap coefficient is assumed to increase from a steady-state value consistent with an initial fuel temperature
/10,000 Btu per hour-square foot *F at the initiation of the transient.
Thus, thq large amount of energy stored in the fuel because of the small initial value is released to the cladding at the initiation of the transient.
Zirconium-Steam Reaction - The zirconium-steam reaction can become significant above a cladding temperature of 1800*F. The Baker-Just parabolic rate equa-tion shown below is used to define the rate of the zirconium-steam reaction:
d(w') = 33.3 x 10' exp (45,000)( M o M h a. --1 P O~
dt -
1.986T Lj$*$OO ukW where:
w = amount reacted (mg/cm2) t = time (seconds)
T = temperature (*F). ( hould k Mvh The reaction heat is 1510 cal /gs.
l Results 14 1.5- 4 i
)
b Figure [14.1.8-17] shows the core flow -d N "- :
M..0 h ;h; . the nuclear power and maximum pressure j transientsand loop flow tr channel and hot channel heat flux transients = hn: ' "r e l' ' "-9 d dad k cladding temperature transient.h ;im '- ";-e l' L" 20.*The results of these calculations are summarized in Table 14.1.8-3. The sequence of events is
~ ~
shown in Table 14.1.8-4.
4 I N. \, S '""1 d IN.l.$~f 14.1.8-7 June 1994
i t]o cgA4 @ 3 Conclusions Since the peak reactor coolant system pressure reached during any of the transients is less than that which would cause stresses to exceed the faulted condition stress limits of 3120 psi, the integrity of the primary coolant system is not endangered. '
Since the peak cladding surface temperature calculated for the hot : pot during the more severe transient remains considerably less than 2700*F and the amount of zirconium-water reaction is small, the core remains in place and intact with ,
no consequential loss of core cooling capability.
References l
)
- 1. Goldberg, G., Westinghouse Electric Corporation, letter to E. J. Lipke, NPD, " Wisconsin Electric Power Company Point Beach Units 1 and 2 Final Reports for RCP Bus Frequency Decay Analysis," WEP-91-196, August 19, 1991.
- 2. Friedland, A. J., Ray, S., " Revised Thermal Design Procedure,"
WCAP-11397-P-A (Proprietary), WCAP-ll397-A (Non-Proprietary), April,1989.
14.1.8-8 June 1994
i INSFAT "A" Initial reactor power, RCS temperature, and pressure are assumed to be at the most limiting nominal conditions, i.e. 100% power, maximum RCS temperature, and reduced pressure operation. Uncertainties in initial conditions are included in the DNBR limits as described in Referenc's 2.
INSERT =m" The most-positive moderator temperature coefficient limit for full-power operation (0 pcm/*F) is assumed, since this results in the ==v4== core power during the initial part of the transient, when the minimum DNBR is reached.
INSERT "C" Figure 14.1.8-1 shows the reactor coolant flow coastdown curves for a loss of both pumps. Figure 14.1.8-1 also shows the nuclear flux, RCS pressures, average channcA heat flux, and hot channel heat flux transients for the coastdown of two pumps. The corresponding transients for the coastdown of one pump, and for the underfrequency event, are shown in Figures 14.' 8-2 and 14.1. 8-3 respectively. Table 14.1.8-1 sn=narizes the sequence of events for these transients.
INSERT *D" The lift pressure of the pressurizer safety valves is assumed to be 4% above the nominal set pressure of 2500 psia. 'once this lift pressure is reached, an additional delay of I second is assumed to account for the clearing of the water in the pressurizer safety valve loop seals. The safety valve steam relief capacity is 288,000 lbm/hr per valve.
TABLE 14.1.0-1 Lo w of. Forced Reactor Coolant Flow Time Sequence of Events Time h Event (seconds 1 Complete loss of forced reactor coolant flow Both operating pumps lose power and begin coasting down 0.0 Reactor coolant pump '
i undervoltage trip setpoint reached {
0.0 J Rods begin to drop 1.5 Partial loss of reactor coolant flow (two loops operating, one pump I coasting down) j Coastdown begins 0.0 )
i Low flow reactor trip 1.5 '
Rods begin to drop 2.5
! Underfrequency event Frequency decay begins and 1 pumps begin to decelerate 0.0 Underfrequency setpoint reached; RCPs trip and l
begin coasting down 0.8 l
Low RCS flow reactor trip setpoint reached 2.1 Rods begin to drop 3.1 Min 4== DNBR occurs 4.8 l
Locked rotor Rotor on one pump locks 0.0 Low RCS flow reactor trip setpoint reached 0.03 Rods begin to drop 1.03 t
hv4== RCS pressure e:. curs 3.6 Maximum cladding temperature occurs 4.0 m
1 l
l
TABLE T14.1.8-3j - j N, I. T-d SIM1ARY OF LIMITING RESULTS FOR LOCKED ROTOR ACCIDENT Maximum Reactor Coolant System Pressure 2744
&DT i /,
(psia) g g4p g af).;:,
Maximum Cladding Temperature (*F) 3}$3 at Core Hot Spot
'#1.3 Zr-H,0 Reaction at Core Hot Spot
(% by weight) -
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Point Beach Nuclear Plant Units 1 and 2 Complete Loss of Flow
(
l (2/2 RCP Coastdown)
Figure 14.1.8-1 (sheet 1 of 3)
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Figure 14.1.8-1 (sheet 3 of 3) l
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Point Beach Nuclear Plant Units 1 and 2 Partial Loss of Flow (1/2 RCP Coastdown)
Figure 14.1.8-2 (sheet 1 of 3)
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Point Beach Nuclear Plant Units 1 and 2 Partial Loss of Flow (1/2 RCP Coastdown)
Figure 14.1.8-2 (sheet 2 of 3)
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. . . . . i i i i 0 1 2 3 4 5 s 7 s a 10 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Partist Loss of Flow (1/2 RCP Coastdown)
Figure 14.1.8-2 (sheet 3 of 3)
1 l
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. 2 -
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Point Beach Nuclear Plant Units 1 and 2 RCP Bus Underfrequency (2/2 RCP Deceleration)
Figure 14.1.5-3 (sheet 1 of 3)
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i i . . . . . . .
0 1 2 3 4 5 s 7 8 9 10 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 RCP Bus Underfrequency (2/2 RCP Deceleration)
Figure 14.1.8-3 (sheet 2 of 3)
l J
l 1
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a n O s :
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Point Beach Nuclear Plant Units 1 and 2 RCP Bus Underfrequency (2/2 RCP Deceleration)
Figure 14.1.8-3 (sheet 3 of 3)
1.2 .
=, .
r 3 1- -
=
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J
- Point Beach Nuclear Plant Units 1 and 2 i
RCP Locked Rotor Figure 14.1.8-4 (sheet 1 of 4)
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n
< ~~~
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sa .
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/
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teco -.
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o 1 2 3 4 5 s 7 8 9 10 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 RCP Locked Rotor Figure 14.1.8-4 (sheet 2 of 3)
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US
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0 1 2 3 4 5 6 7 8 9 10 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 RCP Locked Rotor Figure 14.1.8-4 (sheet 3 of 3)
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2 200o -
a _
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p .
8 1500 --
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o 5
- 10o0 -
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- u. so --
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Point Beach Nuclear Plant Units 1 and 2 RCP Locked Rotor l Figure 14.1.8-4 (sheet 4 of 4) {
i
i i
4 i
' \
! - go cHAAGES) i 14.1.9 LOSS OF EXTERNAL ELECTRICAL LOAD The plant is designed to accept a 55 loss of electrical load while operating at full power or a complete loss of load while operating below 50% power without ]
actuating a reactor trip. Tae automatic steam bypass system with 40% steam dump capacity to the condenser is able to accommodate this load rejection by reducing the transient imposed upon the reactor coolant system. The reactor power is reduced to the new equilibrium power level at a rate consistent with the capability of the rod control system. Should the reactor suffer a complete loss of load from full power, the reactor protection system would automatically actuate a reactor trip.
The most likely source of a complete loss of load on the nuclear steam supply system is a trip of the turbine-generator. In this case, there is a direct reactor trip signal derived from either the turbine autostop oil pressure or a closure of the turbine stop valves, provided the reactor is operating above 50%
power. Reactor temperature and pressure do not increase significantly if the steam bypass system and pressurizer pressure control system are functioning properly. Wutheplant behavior is evaluated for a complete loss of load
~
from full power without a direct reactor trip, primarily to show the adequacy of 7hdressure relieving devices and also to show that no core damage occurs. The Wa'cGr coolant system and ste.2 system pressure relieving capacities are d2 signed to ensure the safety of the plant without requiring the automatic rod control, pressurizer pressure control, and/or steam bypass control systems. 1 l
Method of Analysis The total loss of load transients are analyzed by employing the detailed digital computer program LOFTRAN. The program simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves.
The program computes certinent plant variables, including temperatures, pressures, and power level.
14.1.9-1
l I
4 1
1 l
] In this analysis, the behavior of the unit is evaluated for a complete loss of j steam load from full power without direct reactor trip, primarily to show the j j adequacy of the pressure-relieving devices and also to demonst ate core I'
protection margins.
7
! This accident is analyzed with the Revised Thermal Design Procedures in WCAP-i 11397, Reference 1, Section 14. Plant characteristics and initial conditions are discussed in Section 14.
Initial Goeratina Conditions - The ini+M reactor power and reactor coolant system temperatures ar9 ass % at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-11397.
@M
}
The initial reactor coolant system pressure of 2000 and 2250 psia are assumed to ss,4 "A" determine the minimum DNBR and peak pressure respectively. The initial reactor
- coolant pressure of 2000 psia results in the minimum margin-to-core protection limits at the initiation of the transient. "'
- l 1
Moderator and Donaler coefficients of Reactivity - The loss of load accident is analyzed with both maximum and minimum reactivity feedback. The maximum feedback l cases assume a large negative moderator temperature coefficient and the most f negative Doppler power coefficient. The minimum feedback cases assume positive j moderator temperature coefficient and the least negative Doppler coefficient.
I j Reactor Control - From the standpoint of the maximum pressures attained, it is j conservative to assume that the reactor is in manual control.
1
~
{ Steam Release - No credit is taken for the operation of the steam dump system or j steam generator power-operated relief valves. The steam generator pressure rises j to the safety valve setpoint, where steam release through safety valves limits l secondary steam pressure at the setpoint value.
j Pressurizer Sorav and Power-Ocerated Relief Valves - Two cases, for both maximum and minimum feedback, are analyzed.
j$ a. Full credit is taken for the effect of pressurizer spray and power-j operated relief valves in reducing or limiting the coolant pressure.
i
- 14.1.9-2 ii
- ~ -
1 1
, l 1
INSERT *A" Initial ooeratine conditions - The initial core power, i
reactor coolant temperature, and reactor coolant a.
pressure are assumed to be at the most limiting f nominal values. The DNBR calculations are performed using the Revised Thermal Design Procedure
! (WCAP-11397), in which the uncertainties in the initial conditions are included in the DNBR limit value. -For the peak RCS pressure and peak steam generator pressure calculations, uncertainties of 2%,
d 1 30 psi, and 4 *F are applied in the most limiting direction to the initial core power, reactor coolant i
l
} temperature, and reactor coolant pressure.
J i i i
i ,
4
- b. No credit is taken for the effect of pressurizer spray and power- ;
operated relief valves in reducing or limiting the coolant pressure. f Safety valves are operable. l Feedwater Flow - Main feedwater flow to the steam generators is assumed to be lost at the time of loss of external electrical load.
Reactor trip is actuated by the first reactor protection system trip setpoint reached, with no credit taken for the direct reactor trip on turbine trip. l Results The transient responses for a total loss of load from full power operation are shown for four cases--two cases for minimum reactivity feedback and two cases for maximum reactivity feedback illustrated in Figures 14.1.9-1 through[14.1.9-8.M Ikeh The figures reflect the limiting case of 2250 or 2000 psia operation. However the differences between each transient for the two pressures are minor.
Fl yf L 14 d.9-} ShDJS M bAntJ ( rgsp ntt
" ;::-- ' * ' - ' I' . '. . ; .:.n tM
+--"--t n.,,...... for the total loss of steam load with minimum reactivity feedback, assuming full credit for the f pressurizer spray and pressurizer power-operated relief valves. No credit is j taken for the steam dump.
The reactor is tripped by the overtemperature AT trip. The minimum departure from nucleate boiling ratio is well above the limit value. The pressurizer safety valves are not actuated.
kW4. M-)""2 @oJE
'i..... ;4.;.; ; .. 4 ;;.1.; ; .:.. the response for the total loss of steam load with a large negative moderator temperature coefficient. All initial plant parameters are the same as those in Figure /14.1.9-1, M l' ' ." b The DNBR increases thoughout the transient and never drops below its initial value. Pres-surizer relief valves and steam generator safety valves prevent over-pressurization in primary and secondary systems, respectively. .The pressurizer safety valves are not actuated for this case.
In the event that feedwater flow is not terminated at the time of turbine trip for this case, flow would continue under automatic control with the reactor at 14.1.9-3
3
)
l l
l a reduced power. The operator would take action to terminate the transient and bring the plant to a stabilized condition. If no action were taken by the ;
operator, the reduced power operation would continue until the condenser hot well !
was emptied. A low steam generator water level reactor trip would be generated along with auxiliary feedwater initiation signals. Auxiliary feedwater would then be used to remove decay heat with the results less severe than those presented in Section 14.1.10, Loss of Nonnal Feedwater. )
Pao <c M,iFM she, W c total loss of load accident 5:: i ; i::- t % y assuming the plant to be initially operating at full power, with no credit taken for the pressurizer spray, pressurizer power-operated relief valves, or steam dump. The reactor is tripped on the high pressurizer pressure signal. =F s... ,,...; ; .... .,.;.; ; A i _ . . . . n . -. T. . ; ;.....! x t: it 2000 v . . . .,. . . . . . : . . . . . . . . . :" m , . . . . b
[ r..; p n.... u .u...o . The neutron flux increases slightly until the reactor is tripped. The departure from nucleate boiling ratio increases throughout the transient. In this case, the pressurizer safety valve is actuated.
d Ry,e. A\.M 36us T:,_...;0..0-7..~1,...;!i;thetransient[withmaximumfeedback it 22" ,"*
9 ,..:. and all other assumptions being the same as those in ";n. M. .;-; ..~ i Fi3ua.
7 M.;.s-6. Again, the departure from nucleate boiling ratio increases throughout N i .3.- 4, the transSnt, and the pressurizer safety valves are actuated.
N ,j,9 3 I The calculated sequence of events for these four cases is shown in Table g ej% f 14.1.9-1.
Conclusions Results of the analyses show that the plant design is such that a total loss of external electrical load without a direct or immediate reactor trip presents no hazard to the integrity of the reactor coolant system or the main steam system.
Pressure-relieving devices incorporated in the two systems are adequate to limit the maximum pressures" within the design limits.
The integrity of the core is maintained by operation of the reactor protection system; i.e., the departure from ntcleate boiling ratio is maintained above the limit value.
14.1.9-4
t j
4 4
TABLE 14.1.9-1 TIME SEQUENCE OF EVENTS FOR 4 LOSS OF EXTERNAL ELECTRICAL LOAD 4
1 Time of Each Event i
(
, gig Eyv1D1 (Seconds)
- a. With pressurizer Loss of electrical load 0 a
control (minimum
' feedback)
Overtemperature AT reactor
,' trip reached g ?.9 Rod begins to drop g j,k Initiation of release from SG safety valves g /,2.(
Minimum departure from nucleate boiling ratio occurs /[. )e ,
Sc5 e ur [S O i
i b. With pressurizer Loss of electrical load 0 )
5 control (maximum
- feedback)
[ Peak yS p;...
, f ;ier pressure ag
! occurs l' fInitiationofreleasefrom g / Z')r
. I., (. SG safety valve
~
b 'g h_l*._'s_EW!h. _ _..' .b? ...
! level reactor trip point l reached S.3 i Rods begin to drop
[0 /O.,I Minimum departure from nucleate boiling ratio occurs
- 4 i
(1 of 2) i
i TABLE 14.1.9-1 (continued)
Time of Each Event hit lyEl (Secondsi
- c. Without Loss of electrical load 0 Pressurizer control (minimum feedback) ,
High pressurizer pressure reactor trip point reached gCb/
Rods begin to drop M I.b Initiation of release from SG safety valves y /[,0 Sc5 Peak psesseetrer pressure occurs 13 # / O .0 Minimum departure from nucleate boiling ratio occurs *
- d. Without Loss of electrical load 0 pressurizer control *
(maximum feedback)
High pressurizer pressure reactor trip point reached [gg Rods begin to drop >t$ft I.h Peaka~ pes " ' r pressure occur's JPr[ /O.I Initiation of release from SG safety valves g /[,6 Minimum departure from nucleate boiling ratio occurs *
- DNBR does not decrease below its initial value.
(2 of 2)
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Point Beach Nuclear Plant Units 1 and 2 Loss of Electrical Loed With Pressure Control Minimum Reactivity Feedback Figure 14.1.9-1 (sheet 2 of 3)
i I
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Point Beach Nuclear Plant UnMs 1 and 2 Loss of Electrical Load With Pressure Control Maximum Reactivity Feedback Figure 14.1.9-2 (sheet 2 of 3)
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Point Beach Nuclear Plant Units 1 and 2 Loss of Electrical Load Without Pressure Control Minimum Reactivity Feedback Figure 14.1.9-3 (sheet 1 of 3)
Core Average Core W 610 .
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Point Beach Nuclear Plant Units 1 and 2 Loss of Electrical Load Without Pressure Control Minimum Reactivity Feedback Figure 14.1.9-3 (sheet 2 of 3)
I Preeeurtaer Pressure I
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Preneurtaar Pressure RCP Outlet Pressure 2000 -
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aw w e 2400 -- / N 3w - /
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0 10 20 30 40 50 80 70 80 90 100 Time (seconds) 1300 s 1200 g "_ -
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NO CHAM 14.1.10 LOSS OF NCRMAL FEEDWATER A loss of nomal feedwater (from a pipe breik, pump failure, or valve mal-function) results in a reduction in capability of the secondary system to remove the heat generated in the reactor core. If the reactor is not tripped during this accident primary plant-damage could possibly result from a sudden loss of heat sink. If an alternate supply of feedwater were not supplied to the plant, residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer occurs, and significant loss of water from the reactor coolant systes could conceivably lead to core damage.
The following provides the protection in the event a loss of nomal feedwater occurs.
- 1. Reactor trip on low-low water level in either steam generator
- 2. Reactor trip on steam flow-feedwater flow mismatch coincident with low water level in either st.eam generator.
- 3. Two motor driven auxiliary feedwater pumps (200 gpa each) which are -
started on
- a. Low-low level in either steam generator -
- b. Opening of both feedwater pump circuit breakers c Any Safety Injection signal
- d. Manually
- 4. One turbine driven pump (400 gpe) which is started on
- a. Low-low level in both steam generators
- b. Loss of voltage on both 4160 volt busses
- c. Manually -
The motor driven auxiliary feedwater pumps are supplied by the diesel if a loss of outside power occurs. The turbine-driven pump utilizes steam from the secondary systems and exhausts the steam to the atmosphere. The auxiliary 14.1.10-1 June 1992
_~ . _ . . - . - - . - - - - - . - . _ . - - _ - . . . - - _ _ _ _ - ..
l 1
i i
i a
i l
j pumps take suction directly from the condensate storage tank for delivery to the steam generators. .
I q.,,
- The above units provide considerable backup in equipment and control logic to g ensure that reactor trip and automatic auxiliary feedwater flow will occur {g l
following any loss of normal feedwater including that caused by loss of AC -
( yg i power. Js '5 t ,
M i C +-+% -
o Method of Analysis
])
i A detailed analysis using the LOFTRAM code is performed in order to obtain the t.E 45- -
3
] plant transient following a loss of normal feedwater. The simulation c 7. #f -
=3
- l describes the plant thermal kinetics, RCS including the natural circulation, 6*
j pressurizer, steam generators, and feedwater system. The digital program M 4
computes pertinent variables, including the steam generator level, pressurizer 3 l ,
i water level, and reactor coolant average temperature. C .5 L
. T j JL l 'C The following assumptions were made:
j
}s"g y },
rA d ,4cW
- 1. The S... is initially operating at 1025 of 1518.5 MWt.
- 2. Core residual heat generation is based on the 1979 version of ANS-5.1 14 1g {w
~~
j k
, (Reference 1) plus two standard deviations for uncertainty. ANSI /ANS-5.1-197g is a conservative representation of the decay heat release O b"'
l
- '8*
/\ l
! 3. Auxiliary feedwater flow at a rate of 200 gpa is split between the two h j l L steam generators one minute after the incident.
j l
! m ;;; e fu n :.. ... e n e t: u s. .: ; .;. _ ..;.... "
l 4.h 7h assv.md skm pendsc- madeIt om qqp:(un;}j),
j d
- 5. Secondary system steam relief through the safety valves. O -(F [dn // g),
$ IRElh 4
l The calculated sequence of events for this event is Itsted in Table 14.1.10-1.
dFigure 14.1.10-1]sh the plant parameters following a loss of normal 4 7; m 14.1.10-1 ad M.).lM 14.1.10-2 June 1992
1
'M ab) h Gh I .J M 2 . i
(
l feedwater accident with the assumptions listed labove.) Low-low level signal in either steam generator initiates the reactor tr'ip. The reactor trip then initiates the turbine trip. Following the reactor and turbine trip from full g load, the water level in the steam generators
- falls due to the reduction of steam generator void fraction and because steam flow through the safety valves j 'q continues to dissipate the stored and generated heat.
]N e1 Upon the initiation of the low-low level signal, the auxiliary feedwater pumps are automatically started. The pumps will supply auxiliary feedwater to both steam generators within one minute, reducing the rate of water level decrease. [q g.
v w It is possible that a loss of normal feedwater initiated by a seismic event %5 ,
could also result in the interruption of the normal source of auxiliary
]#}
-i -
feedwater from the Condensate Storage Tanks because the Condensate Storage Tanks are not classified as seismic class I. The plant operators would be f
4%G alerted to this problem by the receipt of low suction pressure alarms on the auxiliary feedwater pumps. Switchover to the alternate source of seismically qualified auxiliary feedwater, the Service Water System, can be accomolished by the operators in five minutes or less. A calculation demonstrates that a h
]f 19
- f,,
'five minute delay in the initiation of auxiliary feedwater flow to the steam generators does not degrade the capability of the steam generators to remove decay heat because minimum steam generator water inventory does not drop below theminimumwaterlevelcalculatedintheanalysisdocumentedanddescribedinj
'this section. The calculation assumes that one auxiliary feedwater pump is (
alsooutofserviceorfailed.) ;
The capacity of the auxiliary feedwater system ts such that the water level in the steam generators does not recede below the lowest level at which '
sufficient heat transfer area is available to dissipate core residual heat t]ithout water relief from the RCS relief or safety valves. From Figures 14.1.10-1 it can be seen that at no time is there water relief from the pressurizer.
Conclusion l
The loss of normal feedwater does not result in any adverse condition in the core, because it does not result in water relief from the pressurizer relief or safety valves.
14.1.10-3 June 1994
/
1 TABLE 14.1.1'0-1 TIME SEQUENCE OF EVENTS FOR LOSS OF NORMAL FEEDWATER FLOW INCIDENTS ,
T k Csace 25I 11tal (s ds) O^b I OnN E Main feedwater flow stops 1% go 10 Low steam generator water level trip 6h 6f -dC Rods begin to drop 6 6't 'f L Peak water level in pressurizer occurs 71I jgl@ J9ff
)w6steamgenerator/beginsto 12g0 b SO receive auxiliary feedwater Cold auxiliary feedwater is 2 0 @)O delivered to the steam generator Core decay heat decreases to - d auxiliary feedwater heat removal capacity
12 -
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Point Beach Nuclear Plant 4
Units 1 and 2 Unit 1 (Model 44F SG)
Imes of Normal Feedwater 14110-1 .
est 1 of 8)
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Point Beach Nuclear Plant Units 1 and 2 Unit 1 Model 44F SG) loss of N(ormal Feedwater I i
Figure 14 1.10-1 l (sheet 2 of 6)
tm 700 ein .
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Point Beach Nuclear Plant Units 1 and 2 Unit 1 Model 44F SG)
Imse of N(ormal Feedwater Figwe 14110-1 (sheet 3 of 6)
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)
l Unit 1 Model 44F SG) !
Ioss of N(ormal Feedwater !
1 Figure 14110-1 l (sheet 4 of 6) l l
l 1
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Point Beach Nuclear Plant Units 1 and 2 Unit 1 (Model 44F SG) loss of Normal Feedwater Figure 14110-1 (sheet 5 of 6)
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\-
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, Unit 1 (Model 44F SG) .
less of Normal Feedwater l 14110-1 6of6) 1
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Time (seconds) l Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG)
Imse of Normal Feedwater Figure 14110-2 (sheet 1 of 6)
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Ims of Normal Feedwater eet o6
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700
=
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Point Beach Nuclear Plant Units 1 and 2 l
Unit 2 (Delta-47 3G)
Ioss of Normal Feedwater Mgure 141.10-2 (sheet 3 of 6)
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG)
Loss of Normal Feedwater Figure 14110-2 (sheet 4 of 6)
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG) 14ss of Norinal Feedwater 14110-2 eet 5 of 6)
l l
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG)
Ioss of Normal Feedwater I Figure 14110-2 (sheet 6 of 6)
NO CHAMkbb REFERENCES - Section 14.1.10
- 1. "American National Standard for Decay Heat Power in Light Water Reactors,"
ANSI /ANS-5.1 - 1979, August 1979. j l
- 2. Burnett, T. W. T., et. al., "LOFTRAN Code Description " WCAP-7907-P-A (Proprietary), WCAP-7907-A (non-Proprietary), April 1984. ,
I j
i e
l 4
c NO CAA44ES
. 14.1.11 LOSS OF ALL AC POWER TO THE STATION AUXILIARIES i In the unlikely event of a complete loss of all auxiliary AC power; the i turbine will be tripped and there will be a loss of power to the station l auxiliaries. The sequence below is described for the unit following a turbine j trip.
- 1. Plant vital instruments are. supplied by the emergency power sources.
j
- 2. As the steam systes pressure subsequently increases, the steam system l power relief valves are automatically opened to the atmosphere. Steam bypass to the condenser is not available because of loss of 1[he ..
- circulating water pumps.
~
1
- 3. As the steam flow rate through the power relief valves may not be sufficient, the steam generator self-actuated safety valves say
- temporarily lift to augment the steam flow until the rate of heat
- dissipation is sufficient to carry away the sensible heat of the fuel and
] coolant above no-load temperature plus the residual heat produced in the reactor.
] 5. As the no-load temperature is reached, the steam systes power relief 1 valves are used to dissipate the residual heat and to maintain the plant
! at the hot shutdown condition.
The steam turbine driven auxiliary feedwater pump is automatically started by the loss of AC power on the buses that supply power to the Main Feedwater
! Pumps. The turbine utilizes steam from the secondary system to drive the
! feedwater pump to deliver sakeup water to the steam generators. The turbine driver exhausts the secondary steam to the atmosphere. The motor driven i auxiliary feedwater is supplied by power from the diesel generators. The l pumps take suction directly from the condensate storage tank for delivery to
.the_ steam generators. The auxiliary feedwater system insures feedwater supply j of more than 200 gpa upon the loss of power to the station auxiliaries, since I
the steam turbine driven auxiliary feedwater pump has a capacity of 400 gpa l
and the motor driven auxiliary feedwater pumps have a capacity of 200 gpa i each.
14.1.11-1 June 1992
1 a
1 i
I The steam driven feedwater pump can be tested at any time by admitting steam s to the turbine driver. The motor driven auxiliary feedwater pumps also can be N tested at any time. The auxiliary feedwater* control valves and power relief '5 0- l valves can be operationally tested whenever the plant is at hot shutdown and the remaining valves in the system are operationally tested when the turbine p
.2 M j driver and pump are tested. .
j i
Method of Analysis '
9 - } s-l A detailed analysis using the LOFTRAN code is performed in order to obtain the $.51bh >
plant transient following a loss of all AC power to the station auxiliaries. *0g i
l The simulation describes the plant thermal kinetics, RCS including the natural circulation, pressurizer, steam generators, and feedwater system. The digital
< j t,, ;
1 l 5 f. J program computes pertinent variables, including the steam generator level, ,5 pressurizer water level, and reactor coolant average temperature.
L 3~ f .:.
E
, ?r
- The following assumptions are made:
YeMht" ie {$ l
- /. The-thn4 is initially operating at 102% of 1518.5 MWt. ,
g J, Core residual heat generation is based on the 1979 version of ANS-5.1 (Reference 1) plus two standard deviations for uncertainty.
AN51/ANS-5.1 - 1979 is a conservative representation of the decay heat 4
release rates.
4
" Auxiliary feedwater flow at a rate of 200 gpa is split between the two steam generators one minute after the incident.
u m._ ;x.. .- .m M . : n .. n : .. ., in en:.. . '-
"Th a.ssuused Shew $tner$ /webN M NN
- 0' N I r, Secondary system steam relief through the self-actuated safety valves.
6, After normal steam generator level is established, auxiliary feedwater flow is controlled to maintain the water level.
The assumptions used in the analysis are similar to the loss of normal feedwater (14.1.10) except that power is assumed to be lost to the reactor 14.1.11-2 June 1992
j coolant pumps at the time of reactor trip plus an appropriate delay time (2 sec. for reactor trip and 2 sec. for loss of power for a total of 4 sec.).
Results The calculated sequence of events for this accident is listed in Table .
14.1.11-1. The transient responsa of the RCS following a loss of AC power is shown in Figures 14.1.11-1 and 14.1.11-2.
The first few seconds after the loss of power to the reactor coolant pumps will closely resemble the simulation of the loss of reactor coolant flow event (14.1.8), where core damage due to rapidly increasing core temperatures is ~
prevented by promptly tripping the reactor. After the reactor trip, stored and residual decay heat must be removed to prevent damage to either the RCS or the core.
The results of the analysis show that the natural circulation flow available is sufficient to provide adequate core decay heat removal following reactor trip and RCP coastdown. -
~
Conclusion The loss of AC power to the station. auxiliaries does not cause any adverse '
condition in the core, since it does not result in water relief from the pressurizer reitef or safety valves.
}
14.1.11-3 June 1992
- _ . . . - . _ . _ ._ -= .,-.-_ - . =_ -
i i
TABLE 14.1.11-1 l .
TIME SEQUENCE OF EVENTS FOR LOSS OF 0FFSITE POWER INCIDENTS
- I Time TW C5,te d s) l hant fseconds) g,,4 l On& z,.
p
- Main feedwater flow stops 10 10 Low steam enerator 6[1 O 6I
, water leve trip Rods begin to drop 6[1 N Reactor coolant pumps begin to coastdown 6g1 b1 Peak water level in 8)d ENh i pressurizer occurs
! steamgenerator/beginsto 1[.0 O
3 receive auxiliary feedwater Cold auxiliary feedwater is 2[0 N delivered to the steam generato Core decay heat decrea~ses to -480 SM l auxiliary feedwater heat removal / -
capacity 1
- Nonemergency AC power to station' auxiliaries is lost at 69 seconds. I l
1 i
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1 i
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Unit 1 (Model 44F SG)
Ices of AC Power 14.1.11-1 et 2 of 6)
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=1 e 1 2 3 4 10 10 10 10 10 10 Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Unit 1 (Model 44F SG)
Ioss of AC Power Figure 14.1.11-1 (sheet 3 of 6)
2600 .
~
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Point Beach Nuclear Plant Units 1 and 2 Unit 1 (Model 44F SG) loss of AC Power 14.1.11-1 4 of 6) l
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Point Beach Nuclear Plant Units 1 and 2 Unit 1 (Model 44F SG)
Imse of AC Power Mgure 14111-1 (sheet 5 of 6) l 1
l t
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1.2 9 -
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Deid-47 SG) 1mes of AC Power Figure 14111-2 (sheet 1 of 6)
r 1
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Imes of AC Power 14111-2 2 of 6)
I h
700 s
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Time (seconds)
Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG)
Imse of AC Power
> 14111-2 l (sheet 3 of 6) 1 l
1 1
1 2600 .
m 2400 --
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s i e 2200 - -
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10 10 10 10 10 10 Time (seconds) 1 1100 -
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w.s :
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500 --
400 ", '
-1 0 1 2 3 4 10 10 10 10 10 10 l
l Time (seconds) l Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG) loss of AC Power Figure 14111-2 i (sheet 4 of 6) i l
1400 7 -
1200 --
1 -
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R 1000 --
- n. .
~
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600 -
400 l 'l 'l
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600 --
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG) loss of AC Power Figure 14 1.11-2 (sheet 5 of 6)
I
160000 .
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10 0 0 0 0 --
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Point Beach Nuclear Plant Units 1 and 2 Unit 2 (Delta-47 SG)
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1 l
q l
@ CNAdbN l
REFERENCES - Section 14.1.11 l
- 1. "American National Standard for Decay Heat Power in Light Water Reactors,' ANS!/ANS-5.1 - 1979, August 1979. l f
o 2. Burnett, T. W. T., et.al,, "LDFTRAN Code Description." WCAP-7907-P-A (Proprietary), WCAP-7907-A (non-Proprietary), April 1984.
e i
s
1 !
t i
! l 14.2.4.4 Radiolocical Consecuences of a Steam Generator
] Tube Ruoture Accident i
! This section presents an evaluation of the offsite consequences of a
} steam generator tube rupture accident. ^" ' * " - - - - - - " " ---
l ..u, - .. . s _. ... .n .za__, w .. . ... :,_u,. __, _,,
. s_ _ _ _ -
! Assumotions: The following assumptions were used in the analysis of the off-site consequences:
! ^
[M/} ?
ine equilier primary coo T. activity equivaient oa r l 6 j factive fuel. ef. Table 9. (See note 1 page 14.2. 9) i i l
- ,2 . With prima o secondary age assumed r to the po ulated I
! accident, the uilibrium a vity in the s ndary sys is f l ected assumin e followin N N, i .
- a. Pr ry to seconda eakage is e ly distribut 'n both steam nerators. '
]
unit vol. ca i I b The iodine p tion factor % Amount unt todine/
iod vol. liquid %
l assumed to be 1 in steam g rators and blowdown tank l
Amo n odine/ unit aas 1 The ine partition tor , Amount i
/ unit vol. id
- ~
i s ass to be 10~' in ensers. .
- d. The ondar stem activity val process' sisted of l lease, blowd tank radioac e dec air ejector l
j venting, blowd tank liquid arge.
- e. T lowdown ra from th as generator s continuous.
j Addi 1 paramet used t alculate th quilibrius .
l re present n Table
! activiti 'n the sec ary syst
! .2.4-1. k i 14.2.4-5 1
i____ --
l I
Insert 1 (on FSAR page 14.2.4-5 where marked)
- 1. Both pre-accident and accident initiated iodine spikes are analyzed. For the pre-accident iodine spike it is assumed that a reactor transient has occurred prior to the steam generator tube rupture and has raised the RCS iodine concentration to 60 pCi/gm of dose equivalent (DE) I-131. For the accident initiated iodine spike , the reactor trip associated with the steam generator tube rupture creates an iodine spike in the RCS which increases the iodine release rate from the fuel to the RCS to a value 500 times greater than the release rate corresponding to the maximum equilibrium RCS Technical Specification concentration of 1.0 Ci/gm of DE I-131. The duration of the accident initiated iodine spike is 1.8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />.
- 2. The noble gas activity concentration in the RCS at the time the accident occurs is based on a fuel defect level of 1.0%. 'Ihis is approximately equal to the Technical Specification value of 100/E pCi/gm for gross radioactivity.
- 3. The iodine activity concentration of the secondary coolant at the time the steam generator tube rupture occurs is assumed to be equivalent to the Technical Specification limit of 1.2 Ci/cc of I-131.
- 4. The amount of primary to secondary steam generator tube leakage in the intact steam generator is assumed to be equal to the Technical Specification limit for a single steam generator of 500 gallons / day.
- 5. No credit for iodine removal is taken for any steam released to the condenser prior to reactor trip and concurrent loss of offsite power.
- 6. An iodine partition factor in the steam generators is used as follows:
0.1 (curies I/gm steam + curies I/gm water)
- 7. All noble gas activity carried over to the' secondary side is assumed to be immediately released to the outside atmosphere.
1 i
I equil rium co entrati of iod) and no e gas in the se dary's e versu assumed imary to condary akag ates j
with inary lant ac vities sociated h 1% de tive 1 e giv in Fig es 14.2. -1 thro 14.2.4- For a en lea ra , if a cool t acti y is 1 s than th equiva t 1%
acti ty, se ndary tivity ld be respondin lowe S.
/ Thirty minutes after the postulated tube rupture accident the pressure between the faulted steam generator and the primary
) ,ygoj system is equalized. Approximately'ob,405 lbs. of reactor j coolant is discharged to the secondary side of the faulted steam j
g 3 0 { generator. Also, approximateW,;^ lbs. of steam is released
] to the atmosphere via the ruptured steam generator during the time interval.
j 8/ Auxiliary feed water is available during the accident.
i I#f. Six hours after the accident the residual heat removal system is placed into operation.
i
//f. Six hours after the accident no further activity is released to. l
- the environment.
i
/2 /. .
i i
The atmospheric dispersion factor (X/Q) at the site boundary
- (1200 meters) and at the boundary of the low population zone I (9000 meter) are
l l
X/Q (sec/m8) 0-2 HR 2-6 HR 1200 meter 2.3x10" duIsteP 9000 meter 2.6x10-' 1.3x10**
O./ Breathing rate used to calculate the thyroid dose for the accident is 3.47x10" m8/sec.
14.2.4-6
i Consecuences:
4th pr ary oolan activi asso lateo tn d facti fuel et pr ary se ndary leakag rate f .3 gps, act ities i di etive odin and ble ga s rel sed o r vari time riod as rator nha e' 'tra e civ in T2 e la.2. -22 a.e e sta The thyroid demas and body doses at the site boundary and the '
Nndary of the low population zone are given in Table 14.2.4#7, is J .ih ed m . ry osa acti is ess an 46 uivai eti y ea a corr on gdosk (u 1, he e a yst l . wha b 1 r.f Because cine u uoie in ter, cc icerapi eparacio tai o r the st generat . The factive econtant tion fac r is fun ion of , steam up rate, nd ste generato iquid ma as well a the par ion coef ient as ntione in the as tions e equi rium act ities li d above, furth partitio actor of 10 is ass in the in con er. s is ba d upon an ine iarti n fact of 10" less fr experi nts per rmed in Ca a d Rus ."**'
< The t roid an whole b doses p sented i able 2.4-3 a given kbot condit s of 'h leadh both i ren= t en e :1;::14 o the public as a result o a team generator tube rupture _n'.d =t M .4.: -t " are less than the permissible limits of 10 CFR Part 100.
14.2.4.5 Multiole Tube Ruotures l l
l A much larger dose, e.g., whole body dose of 25 res at the exclusion radius, can only result from the rupture of sufficient steam generator tubes to cause fuel cladding failure.
Operating experience with steam generators of the type used in this plant has not shown significant numbers of single gross and immediate tube failures. Small leaks in a single tube which caused erosion type damage to adjacent tubes have been reported, but did not cause a rupture of the adjacent tubes. Thus, if a single tube failure were postulated, it is probable that adjacent tubes would not 14.2.4-7 s
l be damaged but any adjacent failure would be an erosion-caused leak rather than a sudden gross failure.
To perform a rigorous analysis of the flow dynamics of blowdown through multiple tube ruptures, one must understand and define mathematically the physical configuration of the ruptures. Because no reasonable mechanism exists for the multiple ruptures, it is instead just as meaningful to analyze the consequences of a pipe rupture, equivalent in terms of discharge rate to various multiples of the single tube discharge rate.
l Such an analysis reveals that the core cooling system will prevent clad damage l for break discharge rates equal to or smaller than that resulting from a broken pipe between 4 in. and 6 in. in diameter. The discharge rates which bracket the onset of cladding damage correspond to 18 and 40 times the discharge from a single severed steam generator tube. Actually, the ratio would be much larger owing to the fact that the discharge from a tube failure will be limited by the back pressure in the steam generator. Ultimately, the tube discharge would terminate when the reactor coolant system and the steam generator reached pressure equilibrium. The operator can initiate cooldown through the unaffected steam generator. I These conclusions are based on single-failure mode performances of the core cooling system. The core does not become uncovered by the calculated quiet level in those cases where cladding damage is found to be prevented.
The incredibility of multiple simultaneous tube failures is supported by the l following reasoning:
- 1. At the maximum operating internal pressure the tube wall sees only about 1530 psi compared with a calculated bursting pressure in excess of 11,100 psi based on ultimate strength at design temperature.
1
- 2. The above margin applies to the longitudinal failure modes, induced by hoop j stress. This failure mode is the least likely to cause propagation of j failure tube-to-tube. An additional factor of two applies to ultimate
- pressure strength in the axial direction tending to, resist double-ended failure (total factor of 14.6).
14.2.4-8
1 l
\
l
- 3. Failures induced by fretting, corrosion, erosion, or fatigue are of such a nature as to produce tell-tale leakage in substantial quantity while ample metal remains to prevent severance of the tube (a small fraction of the original tube wall section) as indicated by the margin derived in 2 above.
Thus, any incipient failures that would develop to the point of severe leakage requiring a shutdown for plugging or repair, in accordance with Section 15.3.1 of the Technical Specifications, would happen long before the large safety margin in pressure strength is lost.
Note 1: It should be n ed that the inary lant tivity as dth}his alysis differs 'th the maxi coolan activ limits o echniki Sp ification 15.3. . The Techn al Spe icati limits we firsh -
impos on Unit 1 by Confirmato Order ted N ember 30, 9, and lat applied to Unit as well. O pril 1983, e NRC issu icense ndments 71 an 76 for Uni 1a 2, r actively, in rporatin these activit inits in t Poin Beach chnical Spec ications. These activit limits are sed o a par tric evalua n conduc by the NRC o pical sites d are nserv ve or Point each Nuc r Plant. The echnical sp ficat limi e ure that e resulti 2-hour dose the site bo ary owing i as genera tube rup e does not e ed an approp ately all ,
fracti of 10 C 100 limit The specif< tions are i tica the Stan rd Techn 1 Specift ions for We nghouse Pre urize<
er React s, NUREG- 52, Revisi 2.
14.2.4-9
REFERENCES - Section 14.2.4
- 1. L. C. Watson, A. R. Bancroft and C. W. Howlke, " Iodine Containment by Dousing in NPD-11" AECL-ll30 Atcmic Energy of Canada Limited, Chalk River, Ontario, October 27, 1960.
- 2. M. A. Styrikovich, O. I. Martynova, K. Ya, Katkovskaya, I. Ya, Dubrovskii, and I. N. Smirnova, " Transfer of Iodine from Aqueous Solutions to Saturated i Vapor," Atomriaya Energiya, Vol. 17, No. 1, pp. 45-49, July 1964. l I
- 3. O. J. Mendler, "Thefkffect of Steam Generator Tube Uncovery on Radiciodine Release," WCAP-13132, January 1992. l
- 4. R. C. Jones, U.S. NRC, Letter to .L. A. Walsh, WOG, " Westinghouse Owners Group - Steam Generator Tube Uncovery Issue," March 10, 1993, Attachment to WOG-93-066 dated March 31, 1993.
June 1994
4 I
! TABLE 14.2.4-J THYROID DOSES AND WHOLE BODY DOSES STEAM GENERATOR TUBE RUPTURE ACCIDENT j
l A. wid fu-sm'4fM&: :
E tt ^" E't: "r_:r ft: d:::: _ r> :" 9' '
l
! 0 - 2 HOUR 0 - S HOUR
! DOSE AT SITE BOUNDARY OOSE AT _PZ THYROID M BODY THYROID M BODY i REM !!In i
) ze.1 +.+ w o 2.s 1.o w '-
! 4Wegee ; . U.# " . ;.,; M +weneP*
W1 8,f?' A::: h h ?r 5: :l W h " d':
j B. t ^ " ~ ; t. " ... . l _c.f _c._ .:t 1 i
i 0 - 2 HOUR 0 - 6 HOUR DOSE AT SITE BOUNDARY 00SE AT LPZ i a- r
- THYR 010 m BODY THYROID WN BODY
! !!IB REM iL"7 442XM~* 21 f.ox1*
- W W :.:J.? +retztet TE:
- 1. el Def t = 1%
P mary to onda sak R e=. G
NOCHAr)465j 14.2.5 RUPTURE OF A STEAM PIPE 14.2.5.1 General A rupture of a steam pipe is assumed to include any accident which results in I an uncontrolled steam release from a steam generator. The release can occur due to a break in a pipe line or due to a valve malfunction. The steam release results in an initial increase in steam flow which decreases,during the accident as the steam pressure falls. The energy removal from the Reactor Coolant System causes a reduction of coolant temperature and pressure. With a negative moderator temperature coefficient, the cooldown results in a reduction of core shutdown margin. If the most reactive control rod is assumed stuck in its fully withdrawn position, there is a possibility that the core will become critical and return to power even with the remaining control rods inserted. A return to power following a steam pipe rupture is a potential problem only because of the high hot channel factors which may exist when the most reactive rod is assumed stuck in its fully withdrawn position.
- Assuming the most pessimistic combination of circumstances which could lead to power generation following a steam line break, the core is ultimately shut down by the boric acid in the Safety Injection System. g, Theanalysisofasteampiperuptureisperformedtodemons$te'that:
- 1. With a stuck rod and minimum engineered safety features, the core remains in place and essentially intact so as not to impair effective cooling of the core.
- 2. With no stuck rod and all equipment operating at design capacity, insignificant cladding rupture occurs.
Although DNS and possible cladding perforation (no cladding melting or zirconium-water reaction) following a steam pipe rupture are not necessarily unacceptable, the following analysis, in fact, shows that no DNB occurs for any rupture assuming the most reactive rod stuck in its fully withdrawn position.
14.2.5-1 June 1992
( l i
~
l
. . l l
e yo C%AQS 1
The following systems provide the necessary protection against a steam pipe rupture:
l
- 1. Safety Injection System actuation on:
- a. Two out of three pressurizer low pressure signals,
- b. Two out of three low pressure signals in any steam line.
- c. Two out of three high containment pressure signals. !
l
- 2. The overpower reactor trips (neutron flux and AT) and the reactor trip l 1
occurring upon actuation of the Safety Injection System.
l
- 3. Redundant isolation of the main feedwater lines. Sustained high I feedwater flow would cause additional cooldown, thus, in addition to the l
normal control action which will close the main feedwater valves, any safety injection signal will rapidly close all feedwater control valves, i
1 trip the main feedwater pumps, and close the feedwater pump discharge l valves. Additional isolation is provided by tripping the condensate and l heater drain tank pumps on a high containment pressure safety injection signal to help prevent over-pressurization of the containment for ruptures inside containment.
- 4. Trip of the fast acting steam line isolation valves (designed to close in less than 5 seconds with low flow) on:
- a. One out of the two high steam flow signals in that steam line in l coincidence with any safety injection signal. (Dual set points are provided, with the lower set point used in coincidence with two out of four indications of low reactor coolant average temperature.)
, b. Two out of three high - high containment pressure signals.
Each steam line has a fast closing isolation valve and a check valve. These four valves prevent blowdown of more than one steam generator for any break location even if one valve fails to close. For example, for a break upstream 14.2.5-2 June 1992
of the isolation valve in one line, closure of either the check valve in that line or the isolation valve in the other line will prevent blowdown of the other steam generator. -
Steam flow is measured by monitoring dynamic head in nozzles inside the steam pipes. The nozzles (16 in. I.D. vs a pipe diameter of 28 in. I.D.) are located inside the containment near the steam generatorLand also serve to limit the maximum steam flow for any break further downstream. In particular, the nozzles limit the flow for all breaks outside the containment,.J The. Unit 1}-
steam generators contain a steam nozzle flow limiting device which is designed to limit the steam generator depressurization rate by restricting the steam flow during any postulated steam line break accident. Y 14.2.5.2 Method of Analysis Orhh d de b d l The analysis of the steam pipe rupture has been performed to determine:
- 1. The core heat flux and reactor coolant system temperature and pressure resulting from the cooldown following the steam line break. The LOFTRAM code has been used. ,
.[.,,
- 2. The thermal and hydraulic behavior of the core follsing a steam line t" ik. A detailed thermal and hydraulic digital computer code, THINC, i
been used t Octermine if DN8 occurs for the core conditions computed f
(1) above. ,
- 3. The off-site consequences of the steam line break accident which include consideration of the additional secondary loop activity resulting from a steam generator tube leak prior to the accident.
i The following assumptions are made:
- 1. A 2.77% shutdown reactivity from the rods at no load conditions. This is the end of life design value including design margins with the most reactive rod stuck in its fully withdrawn position. Operation of rod cluster control assembly banks during core burnup is restricted in such a ,
way that addition of positive reactivity in a secondary system steaia 14.2.5-3 June 1992
l l
)
release accident will not lead to a more adverse condition than the case analyzed.
- 2. The negative moderator temperature coefficient corresponding to the end of life core with all but the most reactive rod inserted. The variation of the coefficient with temperature and pressure has been included. The k versus temperature at (I000 psia} corresponding to the negative.aedepe4es.R.,
~
fp j temperature coefficient used is shown in Figurell4.2.5-1.] In computing
,g g l the power generation following a steam line break, the local reactivity I l feedback from the high neutron flux in the region of the core near the stuck control rod has been included in the overall reactivity balance.
The local reactivity feedback is composed of Doppler reactivity from the j high fuel temperatures near the stuck control rod and moder'ator feedback from the high water enthalpy near the stuck rod. For the cases analyzed where steam generation occurs in the high flur regions of the core the effect of void forisation on the reactivity has been included. The effect of power generation in the core on overall reactivity is a function of the core temperature, pressure, and flow and thus is different for each
- case studied. The curves assume end of life core conditions with all l rods in except the most reactive rod which is assumed stuck in its fully l withdrawn position.
l ,
1
- 3. Minime capability for injection of 2,000 ppe boric acid solution j corresponding to the most restrictive single failure in the safety injection system. The emergency core cooling system consists of three systems: 1) the passive accumulators, 2) the low head safety injection
'(residual heat removal) system, and 3) the high head safety injection'
, systes. Only the high head safety injectionfand the passive accumulators 15 Quodeled for the steanline break accident analysis.
The modeling of the safety injection systee in LOFTRAN is described in
- Reference 1. The flow corresponds to that delivered by one safety l injection pump delivering its full flow to both RCS cold legs.
M For the cases where offsite power is assumed.The sequence of events in the safety injection system is the following: After the generation of the safety injection signal (appropriate delays for instrumentation, 14.2.5-4 June 1992
- . =- .- -. -~ -
)
i gn clgd(M; AS @ AS A I S~~~SC C0"
@luo.,ce (v diesel cynuM g QAMuf2 AMM locad M Ob i
-ytte55c d Softh i]jtrb80" Cjuyrne
/
J logic, and signal transport [ included), the appropriate valves begin to I operate and the high head safety injection pump starts. Ten seconds later, the valves are assumed to be in their final position and the pump i is assumed to be at full speed. The volume containing unborated water is swept into the core before the 2,000 ppe borated water reaches the core. This delay, described above, is included in the modeling. I In cases where offsite power is not available, an additional 15 second ! delay is assumed to start the diesel generators and to load the necessary safety injection equipment onto them.
- 4. Design value of the steam generator heat transfer coefficient including-i i
allowance for fouling factor. l ! 5. Power peaking factors corresponding to one stuck RCCA and nonuniform core } inlet coolant temperatures are determined at end of core life. The l l coldest core inlet temperatures are assumed to occur in the sector with ! l the stuck RCCA. The power peaking factors account for the effect of the
- j local void in the region of the stuck RCCA during return to power phase l j following the steamline break. This void in conjunction hithe large j
negative moderator coefficient partially offsets the ef , tefthestuck j RCCA. The power peaking factors depend upon the core peer, temperature, } pressure, and flow, and thus are different for orth case studied. ' ! 700f" Q . 6. E'70 rd '-"'- : c' in.L .in; --f[ initial plant conditions have been { considered in determining the core power and reactor coolant system l transient.
- a. Complete severance of a pipe outside the containment,l downstream of f--p h NM
~ r
# tt: .^. . _ __ .. ;; xx_ .' at initial no load conditions ritt "
- " ....a. ; _.. ;. m
- .
i outs'ide ) f
- b. Complete severance of a pipe 4ao4de the containment lat the outlet of
[O...na,_ th :t-- ;--- :t: at initial no load conditions -'" xt:f f: ::-: ".
-afh dn k 7
i k j j 14.2.5-5 June 1992
~
- c. Case (a) above with loss of outside power simultaneous with the steam bmak.
- d. Case (b) above with the loss of outside power simultaneous with the steam break,
- e. A break equivalent to steam release through one steam generator safety valve with outside power available. -
C. @ Case (a) above with only one loop in service.
. Case (b) above with only one loop in service.
t ,___,_,_<_.._m __,. ___ ,___ ,_ ___ ,__ CL i The cases above assume initial hot shutdown conditions with the rods inserted (except for one stuck rod) at time zero. Should the reactor be just critical or operating at power at the time of a steam line break the reactor will be tripped by the normal overpower protection system when the power level reaches a trip point. 5 Following a trip at power the reactor coolant system contains Iore stored energy than at no load, the average coolant temperature is higher than at no load and there is appreciable energy stored in the fuel. Thus, the additional stored energy is removed via the cooldown caused by the steam line break before the no load conditions of reactor coolant system teoperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analyses which assume no load conditions at time zero. l 14.2.5.3 Results l l The results presented are a conservative indication of the events which would I occur assuming a steam line rupture. The worst case assumes that All of the [ following occur simultaneously. I 14.2.5-6 June 1992
l l l 1
- 1. Minimum shutdown reactivity margin equal to 2.77%.
l
- 2. The most negative moderator temperature coefficent for the rodded core at end of life.
- 3. The rod having the most reactivity stuck in its fully withdrawn position.
l
- 4. One safety injection pump fails to function as designed.
A. Care Power and Reactor coolant System Transient l l Figures 14.2.5-2 through 14.2.5-4 show the reactor coolant sys,tes transient qq and core heat flux following a steam pipe rupture (complete severance of a jf pipe) outside the containment, downstreast of the flow measuring nozzle at initial no load conditions with two loops in operation. The break assumed is f """ the largest break which can occur anywhere outside the containment either h upstream or downstream of the isolation valves. Offsite power is assumed M, available such that full reactor coolant flow exists.1 The transient shown assumes the rods inserted at time 0 (with one rod stuck in its fully withdrawn
=p l position) and steam release from only one steam generator. Should the core be -
g critical at near zero power when the rupture occurs, the initiation of safety 7 l injection by low steam line pressure will trip the reactor. Steam release Q l from at least one stesa generator will be prevented by either the check valve N ' l or by automatic trip of the fast acting isolation valve in the steam line by 2 l the high steam flow signal in coincidence with the safety injection signal. Even with the failure of one valve, steam release is limited to no more than u $$jA five seconds for one steam generator while the second generator blows down. (The steam line isolation valves are designed to be fully closed in less than M C*f f fivs seconds following receipt of closure signal under low flow conditions. w nlith the high flow existing during a steam line rupture, the valves will close considerably faster.) w%y 9 W N $ w i 1 c 16 t u o u .. The core becomes critical with the rods inserted (assuming one stuck rod) at l ! 58.1 seconds. Boron solution at 2,000 ppe reaches the reactor core from the hf
$ ga l
I safety injection system (initiated automatically by the low steam line b M& 'j
'~
pressure) at 28.4 seconds which includes the delay required to clear the I 14.2.5-7 June 1992
i 1 l l 1 l l l l INSERT 'A' I I l Figures 14.2.5-1 through 14.2.5-4 show the reactor coolant system transient and core heat flux following a steam pipe rupture for
)
each of the cases considered. A maximum break area of 1.4 ft2 13 j assumed which is the effective cross-sectional area of the steam nozzle flow limiting devices on the Unit 1 and Unit 2 steam ' generators. The reactor coolant pumps are assumed to be available l l such that full reactor coolant flow exists. This is a more l limiting condition for the return-to-power transient than if the RCPs are assumed to lose power and coast down, due to the enhanced transport of cold water from the steam generator outlet to the core, and the reduced enthalpy rise feedback effects. i
safety injection system lines of unborated water. The peak core heat flux is 4.4% of 1518.5 Wt. Figures 14.2.5-8 through 14.2.5-10 show the case of a steam line rupture at the exit of a steam generator at no load with two loops in operation. The sequence of events is similar to that described above for the rupture outside the containment. The peak core heat flux is 12.95 of 1518.5 Mt. Figures 14.2.5-5 through 14- and 14.2.5-11 through 14.2.5-13 show the responses for the cases asfuming a loss of outside power at the time the safety injection signal is generated which then results in a reactor coolant system flow coastdown. The safety injection system delay time includes the
. time required to start a safety injection pump on the diesel. Only one diesel is assumed to start. Credit is taken for only the safety injection flow entering the cold leg lines. The peak powers are 6.35 and 13.55.
Q Figures 14.2.5-15 through 14.2.5-17 and Figures 14.2.5-18 through 14.2.5-20 show the steam pipe rupture-cases assuming only one loop in operation. The time sequence of events is similar to that described above for the two loops in operation cases. Peak core heat flux is 17.8% and 25.3% of 1518.5 Wt. Figure 14.2.5-14 shows the transient results for a steam flow of 247 lb/sec at 1100 psia from one steam generator with two loops in operation. Figure 14.2.5-21 shows the same transient with one loop in operation. The assumed steam release is typical of the capacity of any single steam dump, relief or safety valve. Safety injection is initiated automatically by low pressurizer pressure. Baron solution. enters the RCS, providing sufficient negative reactivity to maintain the reactor below criticality. The transient is quite conservative with respect to cooldown, since no credit is taken for the energy stored in the system metal other than that of the fuel elements or the energy stored in the other staan generator. Since the transient occurs over a period of about five minutes, the neglected stored energy is likely to have a significant effect in slowing the cocidown. 14.2.5-8 June 1992
No Cha$cs - REFERENCES - SECTION 14.2.5
- 1. Burnett, T. W. T., et al., "LOFTRAN Code Description," WCAP-7907-P-A, April 1984
- 2. Akers, J. J., et al., WCAP-12602, " Report for the Reduction of SI System Baron Concentration, Point Beach Nuclear Units 1 and 2,* September 1990.
1 i i i 1 I i 4 1 4 i f i
,1 i
June 1994
Table 14.2.5-1 I g)$ Rupture of a Steam Pipe L TABL E \ Analysis Assumptions and Sequence of Events PBNP Unit Affected Unit 1 Unit 1 Unit 2 Unit 2 Steam Generator Model 44F 44F Delta-47 Delta-47 Number of Loops in Service 2 1 2 1 Initial shutdown margin, %Ak 2.77 2.77 2.77 2.77 i Rupture occurs in main steamline, sec 0.0 0.0 0.0 0.0 j High-High steam flow setpoint reached, sec 0.1 0.1 0.1 0.1 l i Low steam pressure SI setpoint reached, see 1.5 1.5 1.5 1.5 Steamline isolation occurs, sec 8.5 NA 8.5 NA Low pressurizer pressure SI setpoint reached, see 10.2 15.7 10.1 15.6 Feedwater isolation occurs, sec . 18.5 32.7 18.5 32.6 Safety injection pump at full speed, see 28.5 42.7 28,5 42.6 Core returns to criticality, see -60 -80 -60 -80 Boron reaches core, sec ~70 -80 -70 -80 Time of may1=nn core heat flux, see -130 -160 -130 -160 wn4 == core heat flux, fraction of nominal 0.166 0.123 0.164 0.120 1 4 l l
..m-4 Loop 1 Loop 2 1200 _
4
- I
_ 1000 -- l e I l n. soo - e l a.
- s. _
l g soo - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ ] _
- n. .
e 5 400 - =. l i -
- e. _
_ l m - i
- 2o0 -:
o a a 6 . o So 1o0 150 2o0 250 Time (sec.onds) Loop 1 Loop 2 4 4 sooo -
= 1 4 - !
m o e 25o0 - ! e 3 a
- 2o00 --
a - 3 .e - e - . E 15o0 -: , 3 -i _o _i
- tooo -
E :I e _l e - E Soo -: I
- I
- i, ,
o . . . . o So too 15o 2co 25o Time (seconds) 4 Point Beach Nuclear Plant f Units 1 and 2 Rupture of a Steam Pipe 4 Unit 1 (two loops in service) Figure 14.2.5-1 (sheet 1 of 4)
2500
~
lit g 2000 - ~ e 3 e 1500 -- e _ E -
- n. -
w . 1000 -- r - 3 - e _ e - E
- n. s00 -
0 . . . . 0 50 100 150 200 250 Time (seconds) s00 u m-- I
?.!
0 -
> =0 --.
6 2 s - 3 - t N 288 -: r - 3 - e . l a. 100 -- l {
- .\ .
0 0 50 100 150 X. 200 250 Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a steam Pipe Unit 1 (two loops in service) Figure 14.2.5-1 (sheet 2 of 4)
1 i Loop 1 Loop 2 ^ j
- u. 1 a s00 .
I e - t : l 7 sso -- - m : I e w 500 -: , e
- s ,
s 1 E 450 -h _
's s e -
s F
, 4ao -:
~~ I e :
e - _ sso -: ' e a e soo -: _
$ 250 -b ze ;
e noo . . . . E O 50 10o 150 200 250 Time (seconds) Loop 1 . .4 e c - ] o - c 3 = o - c - 2 - U - ~
! .2 --
M - a g -
.1 -
Z - e 6 o U - 0 . . . . 0 50 100 150 20o 250 Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (two loops in service) Figure 14.2.5-1 (sheet 3 of 4)
l l l
)
too E - c. c. w c 15c -- o - z e - . 2 - c e - j e 1m -- i o - c e _ o u _ a o 1 m u- - w o - 1 0 _ , o . . . . . o So too 15o 20o 25o a' Time (seconds) 1000 _ 1 Boo -: E o-: i o - n.
- : I > 50c -:
2 : 2 - 3 1oo0 -: a - e - 1500 -b 2 - o : o .acco -: 2500 -:
.aoco . . , ,
o 50 1oo 15o 20o 25o Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (two loops in service) Figure 14.2.5-1 (sheet 4 of 4)
Loop 1
~~~
Loop 2 1200 - t 1 m 1000 , i 5 h ! o.
- 1 s00 - l e
a., s - e - e soo -- - .e. -
- n. -
~
E a-e - l e m - 200 - - l 0 i i i i 0 50 100 150 200 250 I Time (seconds) Loop 1
- " - " ~
Loop 2 3000 - 2500 - U. 3 a - 2000 -- e - e : E 1500 -- B - .o. - S 1000 -
- i E : '
s - e - e 500 -: 0 . . . i 0 50 100 150 200 250 Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (one loop in service) Figure 14.2.5-2 (sheet 1 of 4)
-. . --= . _ _ . . - -_ -. - _.
1 i 1
- 2500 l l.
e 2000 - - c. e 4 . i 3 2 e 1500 -- . e - i u e -
- n. -
. s. - e - N 1000 - - g s - e _ e - E
- n. 500 --
0 . . . . 0 50 100 150 200 250 j Time (seconds) l 1 i s00 i l u , no-- - l 1 - , s - i o -
> 300 --
6 - 2s - 3 - m 2o0 -- 2 - t - z - e - 100 -
- n. _
, . . , , , . , , , . , , , . m. , _ . _ . m ._
o . . . m. 0 50 100 150 200 250 I Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (one loop in service) Figure 14.2.5-2 (sheet 2 of 4)
l l Loop 1 u. s e s00 _ 3 sso -_
! -\ $u 500 -5\ -
e -
- n. -
E 4s0 -g 7 e - l H - i 400 -: l e : c - sso -: e e
~
e 300 -: - e
$ 250 -3 3,
l - , , , , , . , , , , , , , , , . . , , , l . 200 i . i , l E O 50 100 150 i 200 250 Time (seconds) Loop 1 i I l _ .4 1 e c - E o - c ,- - l o - c - o - i E -
.E. .2 --
? u - s - E -
.5 -
2 - 6 e - o U ~ 0 . . . . 0 50 100 150 200 250 Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (one loop in service) Figure 14.2.5-2 (sheet 3 of 4) l
l l i 200 l 1 E l g . . w
- c. _ >
1 e 150 -- ! 2 -
; - i k _ l c
e - o e 100 -- o - O _ i c - !, o w - o a 50 -- e w o - O _ 0 . , , 0 50 100 150 200 250 Time (seconds) 1000 _ 500 -: n E O-: o - c.
> -s00 -: = -
i
! 1000 -b e -
e - o 1500 -b w o : O 2000 --
.asco --
- - , , , i , , , . , , . , . , , , , , , , , , .
0 50 100 150 200 250 ' Time (seconds) Point Beach Nuclear Plant Units 1 and 2 Rupture of a Steam Pipe Unit 1 (one loop in service) Figure 14.2.5-2 (sheet 4 of 4)
3 Loop 1
, Loop 2 1200 .
1
. 1000 -"'
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f40 CHMGS 14.2.6 RUPTURE OF A CONTROL R00 MECHANISM HOUSING-RCCA EJECTION In order for this accident to occur, a rupture of the control rod mechanism housing must be postulated creating a full system pressure differential acting on the drive shaft. The resultant core thermal power excursion is limited by the Doppler reactivity effects of the increased fuel temperature and terminated by reactor trip actuated by high nuclear power signals. A failure of a control rod mechanism housing sufficient to allow a control rod to be rapidly ejected from the core is not considered credible for the following reasons:
- 1. Each control rod drive mechanism housing is completely assembled and shop-tested at 4100 psi.
- 2. The mechanism housing are individually hydrotested to 3105 psig as they are installed on the reactor vessel head to the head adapters, and checked during the hydrotest of the completed reactor coolant system.
- 3. Stress levels in the mechanism are not affected by system transients at power, or by the thermal movement of the coolant loops. Homents induced by the design earthquake can be accepted within the allowable primary working stress range specified by the ASME code, Section III, for Class A components.
- 4. The latch mechanism housing and rod travel housing are each a single length of forged type-304 stainless steel. This material exhibits excellent notch toughness at all temperatures that will be encountered.
The joints between the latch mechanism housing and head adapter, and between the latch mechanism housing and rod travel housing, are threaded joints reinforced by canopy type rod welds. I Nuclear Desian Even if a rupture of a RCCA drive mechanism housing is postulated, the operation of a plant utilizing chemical shin is such that the severity of an ejected RCCA is inherently limited. In general, the reactor is operated with-14.2.6-1
i I 40 cHAJ4E5 l the RCCA's inserted only far enough to permit load follow. Reactivity changes ; caused by core depletion and xenon transients are compensated by boron changes. Further, the location and grouping of control RCCA banks are ] selected during the nuclear design to lessen the severity of a RCCA ejection accident. Therefore, should a RCCA be ejected from its normal position during I
- full power operation, only a minor reactivity excursion, at worst, could be
- expected to occur.
l However, it may be occasionally desirable to c;nate with larger than normal i insertions. For this reason, a rod insertion limit is defined as a function of power level. Operation with the RCCA's above this limit guarantees adequate shutdown capability and acceptable power distribution. The position of all RCCA's is continuously indicated in the control room. An alarm will i occur if a bank of RCCA's approaches its insertion limit or if one RCCA deviates from its bank. Operating instructions require boration at the low-low alarm. Reactor Protection The reactor protection in the event of a rod ejection accident has been described in Reference 4. The protection for this accident is provided by high neutron flux trip (high and low setting). These protection functions are described in detail in Section 7.2 of the FSAR. Effects on Ad_iacent Housinas Disregarding the remote possibility of the occurrence of a RCCA mechanism housing failure, investigations have shown that failure of a housing due to either longitudinal or circumferential cracking would not cause damage to adjacent housings. %ver, even if damage is postulated, it would not be expected to lead to a more severe transient, since RCCA's are inserted in the core in symmetric patterns, and control rods immediately adjacent to worst ejected rods are not in the core when the reactor is critical. Damage to an adjacent housing could, at worst, cause that RCCA not to fall on receiving a trip signal; however, this is already taken into account in the analysis by assuming a stuck rod adjacent to the ejected rod. 14.2.6-2
Limitina Criteria This event is classified as an ANS Condition IV incident. Due to the extremely low probability of a RCCA ejection accident, some fuel damage could be considered an acceptable consequence. Comprehensive studies, both of the threshold of fuel failure and of the threshold or significant conversion of the fuel thermal energy to mechanical energy, have been carried out as part of the SPERT project by the Idaho Nuclear Corporation. Extensive tests of U0, zirconium clad fuel rods representative of those in pressurized water reactor type cores have demonstrated failure thresholds in the range of 240 to 257 cal /gs. However, other rods of a slightly different design have exhibited failures as low as 225 cal /gs. These results differ significantly from the TREAT results, which - indicated that this threshold decreases by about 10% with fuel burnup. The cladding failure mechanism appears to be melting for zero burnup rods and brittle fracture for irradiated rods. Also important is the conversion ratio of themal to mechanical energy. This ratio becomes marginally detectable above 300 cal /gm for unirradiated rods and 200 cal /gm for irradiated rods; catastrophic failure (large fuel dispersal, large pressure rise) even for irradiated rods did not occur below 300 cal /ge. In view of the above experimental results, criteria are applied to ensure that there is little or no possibility of fuel dispersal in the coolant, gross lattice distortion, or severe shock waves. These criteria are:
- a. Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal /gm for irradiated fuel.
- b. Average cladding temperature at the hot spot below the temperature M at which cladding embrittlement may be expected (2700*F). 4 )I O
- c. Peak reactor coolant pressure less than that which could cause .I 5 stresses to exceed the faulted condition stress limits. I..E OL 14.2.6-3
INSERT *A*
- a. Average fuel pellet enthalpy at the hot spot below 200 cal /gm (360 Btu /lbm) for irradiated fuel. This bounds non-irradiated fuel which has a slightly higher enthalpy 1Lmit.
- b. Peak reactor coolant pressure less than that which would cause stresses to exceed the faulted condition stress limits.
- c. Fuel melting limited to less than the innermost ten percent of the fuel pellet at the hot spot, even if the average fuel pellet enthalpy is below the limits of criterion (a) above.
.I 1 I i, b i i l i l ~
! l i 3 1
- I f l
- d. Fuel meltine will be limited to less than ten percent of the fuel t
volume at the hot spot even if the average fuel pellet enthalpy is l below the limits of criterion (a) above. j i Method of Analysis The calculation of the transient is performed in two stages, first an average l core calculation and then a hot region calculation. The average core ! calculation is performed using spatial neutron kinetics methods to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator density reactivity. l Enthalpy and temperature transients in the hot spot are determined by adding a multiple of the average core energy generation to the hotter rods and ; performing a transient heat-transfer calculation. The asymptotic power distribution calculated without feedback is pessimistically assumed to persist throughout the transient. Averaae Core Analysis The spatial kinetics computer code, TWIRLE (Reference 4 in Section 14), is used for the average core transient analysis. This code solves the two group neutron diffusion theory kinetic equation in one, two or three spatial o dimensions (rectangular coordinates) for six delayed neutron groups and up to 8h 2000 spatial points. The computer code includes a detailed multiregion, h transient fuel-cladding-coolant heat transfer model for calculation of pointwise Doppler and moderator feedback effects. In this analysis, the code i is used as a one dimensional axial kinetics code, since it allows a more realistic representation of the spatial effects of axial moderator feedback I and RCCA movement. However, since the radial dimension is missing, it is still necessary to employ very conservative methods (described in the following) of calculating the ejected rod worth and hot channel factor. Further description of TWIKLE appears in Section 14. Not Soot Analysis In the het spot analysis, the initial heat flux is equal to the nominal times the design hot channel factor. During the transient, the heat flux hot 14.2.6-4
1 1 i i QO chb 5 channel factor is linearly increased to the transient value in 0.1 second, the time for full ejection of the rod. Therefore, the assumption is made that the hot spots before and after ejection are coincident. This is very conservative, since the peak after ejection will occur in or adjacent to the assembly with the ejected rod, and prior to ejte'. ion the power in this region will necessarily be depressed. The hot spot analysis is performed using the detailed fuel-and cladding transient heat transfer computer code, FACTRAN (Reference 2 in Section 14). This computer code calculates the transient temperature distribution in a cross section of a metal clad UO, fuel rod, and the heat flux at the surface of the rod, using as input the nuclear power versus time and the local coolant conditions. The zirconium-water reaction is explicitly represented, and all , material properties are represented as functions of temperature. A conservative pellet radial power distribution is used within the fuel rod. FACTRAN uses the Dittus-Boelter or Jens-Lottes correlation to determine the film heat transfer before DNB, and the Bishop-Sandburg-Tong (BST) correlation ; to determine the film boiling coefficient after DNS. The BST correlation is conservatively used assuming zero bulk fluid quality. The DNS ratio is not calculated, instead the code is forced into DNB by specifying a conservative DNS heat flux. The gap heat transfer coefficient can be calculated by the code; however, it is adjusted in order to force the full power steady-state j temperature distribution to agree with the fuel heat transfer design codes. Further description of FACTRAN appears in Section 14. Systes Overnressure Analysis Because safety limits for fuel damage specified earlier are not exceeded, there is little likelihood of fuel dispersal into the coolant. The pressure surge may therefore be calculated on the basis of conventional heat transfer from the fuel and prompt heat' generation in the coolant. The pressure surge is calculated by first performing the fuel heat transfer calculation to determine the average and hot spot heat flux versus time. Using this heat flux data, a THINC (Section 3.2.2) calculation is conducted to determine the volume surge. Finally, the volume surge is simulated in a plant-14.2.6-5
s 4 1 No cdA44ES ' > transient computer code. This code calculates the pressure transient taking into account fluid transport in the reactor coolant system and heat transfer to the steam generators. No credit is taken for the pressure reduction caused by the assumed failure of the control rod pressure housing. e i Calculation of Basic Parameters Input parameters for the analysis are conservatively selected on the basis of 4 values calculated for this type of core. The more important parameters are discussed below. Table 14.2.6-1 presents the parameters used in this analysis. Eieeted Rod Worths and Hot Channel Factors The values for ejected rod worths and hot channel factors are calculated using either three dimensional static methods or by a synthesis method employing one dimensional and two dimensional calculations. Standard nuclear design codes are used in the analysis. No credit is taken for the flux flattening effects of reactivity feedback. The calculation is performed for the maximum allowed ' bank insertion at a given power level, as determined by the rod insertion limits. Adverse xenon distributions are considered in the calculation. 4 , Appropriate margins are added to the ejected rod worth and hot channel factors to account for any calculational uncertainties, including an allowance for nuclear power peaking due to densification. Power distributions before and after ejection for a " worst case" can be found in Reference 4. During plant startup physics testing, ejected rod worths and power distributions are measured in the zero and full power rodded configurations and compared to values used in the analysis. It has been found that the ejected rod worth and power peaking factors are consistently overpredicted in the analysir. Reactivity Feedback Weichtino Facton i 1 The largest temperature rises, and hence the largest reactivity feedbacks occur in channels where the power is higher than average. Since the weight of 14.2.6-6
i i i } i i a region is dependent on flux, these regions have high weights. This means j that the reactivity feedback is larger than that indicated by a simple channel ) analysis. Physics calculations have been carried out for temperature changes ! with a flat temperature distribution, and with a large number of axial and radial temperature distributions. Reactivity changes have been compared and f
- effective weighting factors determined. These weighting factors take the form j of multipliers which when applied u single channel feedbacks correct them to
! effective whole core feedbacks for the appropriate flux shape. In this l analysis, since a one dimensional (axial) spatial kinetics method is employed, l axial weighting is not necessary if the initial condition is made to match the ojected rod configuration. In addition, no weighting is applied to the l j moderator feedback. A conservative radial weighting factor is applied to the { transient fuel temperature to obtain an effective fuel temperature as a i function of time accounting for the missing spatial dimension. These j weighting factors have also been shown to be conservative compared to three dimensional analysis (Reference 4). j Moderator and Donaler Coefficient i . The critical boron concentrations at the beginning of life and end of life are adjusted in the nuclear core in order to obtain moderator density coefficient curves which are conservative compared to actual design conditions for the I plant. As discussed above, no weighting factor is applied to these results. ! J l The Doppler reactivity defect is determined as a function of power level using 4 l a one dimensional steady-state computer code with a Doppler weighting factor Z l of 1.0. The Doppler defect used is given ini the main text of Section 3.2.1. >g . The Doppler weighting factor will increase under accident conditions, as g discussed above. * { % i j Delaved Neutron Fraction. 8_,, i 2 Calculations of the effective delayed neutron fraction (p ,,) typically yield
- values no less than 0.705 at beginning of life and 0.50% at end of life for
! the first cycle. The accident is sensitive to p if the ejected rod worth is i i equal to or greater than p as in zero power transients. In order to allow for i 14.2.6-7
p c8AWS s reload cycles, pessimistic estimates of p of 0.49% at beginning of cycle and 0.43% at end of cycle were used in the analysis. Trio Reactivity Insertion 1 i The trip reactivity insertion assumed is given in Table 14.2.6-1 and includes 2 the effect of one stuck RCCA. These values are reduced by the ejected rod 3 reactivity. The shutdown reactivity has been simulated by dropping a rod of the required worth into the core. The start of rod motion occurs 0.5 second ] after the high neutron flm: trip point is reached. This delay is assumed to t consist of 0.2 second for the instrument channel to produce a signal, 0.15 l second for the trip breaker to open and 0.15 seccnd for the coil to release
- the rods. A curve of trip rod insertion versus time is used which assumes
- that insertion to the dashpot does not occur until 2.2 seconds after the start l l of fall. The choice of such a conservative insertion rate means that there is
- over one second after the trip point is reached before significant shutdown
- reactivity is inserted into the core. This is a particularly important conservatism for hot full power accidents.
1 l Reactor Protection i Reactor protection for a rod , ejection is provided by high neutron flux trip (high and low setting). These protection functions are part of the reactor trip system. No single failure of the reactor trip system will negate the protection functions required for the rod ejection accident, or adversely affect the consequences of the accident. Results l Cases are presented for both beginning and end of life at zero and full power. j
- 1. Beoinnina of Cvele. Full Power l 1
0 Control bank U u anused to be inserted to its insertien limit. The ) worst ejected rod worth and hot channel factor are conservatively I calculated to be 400 pcm and 4.5 respectively. The peak hot spot 14.2.6-8
1 i l l t 2H9% 'F 1 l cladding average temperature is[2350*F The peak hot spot fuel center temoerature reaches melting, and is conservatively assumed at 4900*F. However, melting is restricted to less than 10% of the pellet.
- 2. Beainnina of Cvele. Zero Power l For this condition, control bank D is assumed to be fully inserted and banks B and C are at their insertion limits. The worst ejected rod is located in control bank D and has a worth of 790 pcm and a hot channel l U*O tor of @ The peak hot spot ' cladding temperature] reaches 260l*F,f-i> 2(af3 *F the fuel center temperature is[3661*F. h '
E c,)aM afi
- W/*d}L
- 3. End of Cycle. Full Power 3 (,3'3 *p:. .}.gg hyt j I
Control bank D is assumed to be inserted to its insertion limit. The ^ ejected rc' worth and hot channel factors are conservatively calculated
, to be 420 pcm and 5.69 respectively. This results in a peak cladding 22 % 'F' average temperature of l230l*F./ The peak hot spot fuel temperature ,
reaches melting conservatively assumed at 4800*F. However, melting is restricted to less than 10% of the pellet.
- 4. End of Cycle. Zero Power The ejected rod worth and hot channel factor for this case are obtained assuming control bank D to be fully inserted and banks C and B at their insertion limit. The results are 850 pcm and 13.0 respectively. The peak cladding average and fuel center temperatures are Mand 3478'F respectively.
MIT b 43RsM op A summary of the cases presented above is given in Table 14.2.6-1. The nucleaf power and hot spot fuel and cladding temperature transients Jfor the% worst casesi are presented in Figures 14.2.6-1 through [14.2.6-2hbeginning of} A life, full power and beginnira of itfe, zero power).l J
'-+ m. a. 6- +
For all cases, reactor trip occurs very early in the transient, after which the nuclear power excursion is terminated. As discussed previously, the reactor will remain subcritical following reactor trip. l 14.2.6-9 June 1991 p
1 2Y 9 'P t , I cladding average temperature is[2350*F. The peak hot spot fuel center temperature reaches melting, and is conservatively assumed at 4900*F. l However, melting is restricted to less than 10% of the pellet.
- 2. Beainnina of Cycle. Zero Power For this condition, control bank D is assumed to be fully inserted and banks 8 and C are at their insertion limits. The worst ejected rod is
, located in control bank D and has a worth of 790 pcm and a hot channel N* factor of Q The peak hot spot ' cladding temperaturelreachesl260l*F,9 2(af 3 *F thefuelcentertemperatureis}3661*F. h N (,,)xd <$s' eVt/**)L
- 3. End of Cvele. Full Power 3(,53 *p. Q by t.
Control bank D is assumed to be inserted to its insertion limit. The ejected rod worth and hot channel factors are conservatively calculated
, to be 420 pcm and 5.69 respecti_vely. This results in a peak cladding MM 'F' aver *9* ta=P*rature of 1230l*F.) The peak hot spot fuel temperature ,
reaches melting conservatively assumed at 4800*F. However, melting is restricted to less than 10% of the pellet.
- 4. End of Cvela. Zero Power The ejected rod worth and hot channel factor for this case are obtained assuming control bank D to be fully inserted and banks C and B at their j insertion limit. The results are 850 pcm and 13.0 respectively. The !
peak cladding average and fuel center temperatures are hand l3478'F l respectively. 2M2 'F F -+ 3 gqq .y A summary of the cases presented above is given in Table 14.2.6-1. The , nucleaf power and hot spot fuel and cladding temperature transients [for theh worst cases are presented in Figures 14.2.6-1 through [14.2.6-h(beginning ofK life, full power and beginning of life, zero power). l J I 1-+19.a.co-+ For all cases, reactor trip occurs very early in the transient, after which l the nuclear power excursion is terminated. As discussed previously, the reactor w ll remain subcritical following reactor trip. 14.2.6-9 June 1991
4 t-_ _ , --,me 4 Aa. A. _4ms4.eA w.. s e ,+a-.- .y en p.,am m:_e m_.a,. - __. __ , l l 1 i So cAAA % . The ejection of an RCCA constitutes a break in the reactor coolant system, j located in the reactor pressure vessel head. The effects and consequences of ! loss of coolant accidents are discussed in Section 14.3. Following the RCCA ejection, the operator would follow the same emergency instructions as for any other loss of coolant accident to recover from the event. 3 Fission Product Release 4 j It is assumed that fission products are released from the gaps of all rods entering DNS. In all cases considered, less than 15% of the rods entered DNB
- based on a datailed three-dimensional THINC analysis.
Pressure Surae ] A detailed calculation of the pressure surge for an ejection worth of one ) dollar at beginning of life, hot full power, indicates that the peak pressure l does not exceed that which would cause stress to exceed the faulted condition { stress limits. Since the severity of the present analysis does not exceed the
- " worst case" analysis, the accident for this plant will not result in an excessive pressure rise or further damage to the reactor coolant system.
1 l Lattice Deformations i 2 A large temperature gradient will exist in the region of the hot spot. Since
- the fuel rods are free to move in the vertical direction, differential
. expansion between separate rods cannot produce distortion. However, the j temperature gradients across individual rods may produce a differential } expansion tending to bow the midpoint of the rods toward the hotter side of
- the rod. Calculations have indicated that this bowing would result in a 3 negative reactivity effect at the hot spot since Westinghouse cores are j undermoderated, and bowing will tend to increase the undermoderation at the j hot spot. Since the 14 x 14 fuel design is also undermoderated, the same effect would be observed. In practice, no significant bowing is anticip&ted.
l since the structural rigidity of the core is more than sufficient to withstand the forces produced. Boiling in the hot spot region would produce a net flow away from that region. However, the heat from the fuel is released to the water relatively slowly, and it is considered inconceivable that crossflow 4 14.2.6-10
tJo cri A% ES l will be sufficient to produce significant lattice forces. Even if massive and rapid boiling, sufficient to distort the lattice, is hypothetically postulated, the large void fraction in the hot spot region would produce a reduction in this ratio at the hot spot. The net effect would therefore be a negative feedback. It can be concluded that no conceivable mechanism exists for a net positive feedback resulting from lattice deformation. In fact, a small negative feedback may result. The effect is conservatively ignored in the analysis. Conclusions Conservative analyses indicate that the described fuel and cladding limits are not exceeded. It is concluded that there is no danger of sudden fuel dispersal into the coolant. Since the peak pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the j reactor coolant system. The analyses have demonstrated that the fission } product release, as a result of a number of fuel rods entering 0NB, is limited to less than 15% of the fuel rods in the core. The position with regard to l fission product release is, therefore much better than for the double ended j coolant pipe break, (the maximum hypothetical accident) for which over 70% of j the rods are assumed to release fission products. j I t l 4 3 l . i ! 14.2.6-11 l 1 !
REFERENCES - Section 14.2.6
- 1. Tong " Post DNB Heat Transfer" WCAP 7247
- 2. Redfield, J.A., " CHIC-KIN -- A Fortran Program for Intermediate and Fast Transients in a Water Moderated Reactor," WAPD-TM-479, January, 1965.
- 3. Barry, R. F., "The Revised LEOPARD Code - A Spectrum Dependent Non Spatial Depletion Program," WCAP-2759 (1965).
- 4. " Power Distribution Control of Westinghouse PWR" WCAP 7208 (1968).
- 5. Conway and Hein, Journal of Nuclear Materials (15.1),1965.
- 6. Ogard & Leary, "High Temperature Heat Content and Heat Capacity of Uranium Dioxida - Plutonium Dioxide Solid Solutions," LA-DC-8620.
s gocHA%s]. 1 1 1
I l l-l I I
)
l TABLE 14.2.6-1 I l i Input Parameters and Results of the Rod Cluster Control Assembly Ejection Accident Analysis 1 l 4 i Perarnecer BOL-HZP BOL-HFP EOL-HZP EOL-HFP l Initial core power level, l percent of 1518.5 MWt 0% 102% 0% 102% Ejected rod worth, pcm 790 400 950 420 Delayed neutron fraction 0.0049 0.0049 0.0043 0.0043 l Doppler reactivity defect l (absolute value), pcm 1000 1000 900 900 i Doppler feedback reactivity weighting 2.071 1.2 1.885 1.3 Trip reactivity, %Ak 2.0 4.0 2.0 4.0 Fq before rod ejection NA 2.5 NA 2.5 Fq after rod ejection 11.0 4.5 10.0 5.69 Number of operational pumps 1 2 1 2 Maxi mun fuel pellet l average temperature, 'F 3293 4047 2964 3749 Maximum fuel center - temperature, *F 3653 >4900 3249 >4800 l l' Maximum cladding average temperature, *F 2653 2498 2392 2294 ) Maximum fuel stored energy, calories / gram 138.3 176.8 122.2 161.4 Maximum fuel melt nil 3.5% nil 0.5%
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I i 14.3.3 CORE AND INTERNALS INTEGRITY ANALYSIS i N Internals Evaluation A 7pggy i t
- The forces exerted on reactor internals and core, following a loss-of-coolant l accident, are computed by employing the M digital computer program developed for the space-time-dependent analysis of multi-loop PWR plants.
Desian Criteria l The criteria for acceptabili'ty are that the core should be coolable and intact ! following a pipe rupture up to and including a double ended rupture of the ) reactor coolant system. This implies that core cooling and adequate core shutdown must be assured. Consequently, the limitations established on ti)e internals are concerned principally with the maximum allowable deflections and/or stability of the parts. l Critical Internals Uneer Barrel The upper barrel deformation has the following limits: To assure reactor trip and to avoid disturbing the RCC guide structure, the barrel should not interfere with any guide tubes. This condition requires a stability check to assure that the barrel will not buckle under the accident loads. RCC Guide Tubes The RCC guide tubes in the upper core support package have the following allowable limits. Tests on guids tubes show that when the transverse deflection of the guide tube becomes significant, the cross section of the RCC guide tube changes. A allowable transient maximum transverse deflection of 1.0 in. has been established for the blowdown accident. Beam deflections above these limits produce cross section changes with increasing delay in 14.3.3-1
e I l 1 u ~ 3 l j scram time until the con 71 rod will not scram due to interference between the rods and the guide. , The no loss of function limit is established as l ) 1.75 in. With a maximum transier.t transverse deflection of 1.75 in., the } cross section distortion will not exceed 0.072 in., after load removal. This ' ) cross section distortion allows control rod insertion. For a maximum ! i transient transverse deflection of 1.0 in., a cross section distortion not in excess of 0.035 in. is anticipated. ! j Fuel Assamhlies The limitations for this case are related to the stability of the thimbles at the upper end. During tho' accident, the fuel assembly will have a vertical j displacement and could touch the upper package subjecting the components to j dynamic stresses. I ' The upper end of the thimbles shall not experience stresses above the buckling compressive stresses because any buckling of the upper and of the thimbles will distort the guide line and could affect the fall of the control rod. 4 j Unner Facir== l i j The maximum allowable local defomation of the upper core plate where a guide l ' tube is located is 0.100 in. This defomation will cause the plate to contact the guide tube since the clearance between plate and guide tube is 0.100 in. This limit will prevent the guide tubes from being put in compression. In ! order to maintain the straightness of the guide tube, a maximum allowable i i total deflection of I' for the upper support plate and deep beam has been established. The corresponding no loss of function deflection is above 2 in. l 4 Allowable Stress criteria l The allowable stress criteria fall into two categories dependent upon the { nature of the stress state: membrane or bending. A direct state of stress j (Membrane) has a uniform stress distribution over the cross section. The j allowable (Maximum) membrane or direct stress is taken to be equal to the j stress corresponding to 0.2 of the unifom material strain or the yield
- strength, whichever is higher. For unirradiated 304 stainless steel at i 14.3.3-2
O j operating temperature, the stress corresponding to 20% of the unifom strain 3 is: 1 N (S.),,,,,,,,, = 39500 psi l i j For irradiated materials, the limit stress is higher. l For a bending state of stress, the strain is linearly distributed over a cross j section. The average strain value is, therefore, one half of the outer fiber j strain where the stress is a maximum. Thus, by requiring the average strain j to satisfy an allowable criterion similar to that for the direct state of I t stress, the outer fiber strain may be 0.4 times the uniform strain. The j maximum allowable outer fiber bending stress is then taken to be equal to the l stress corresponding to 40% of the unifom strain or the yield strength, ! whichever is higher. For unirradiated 304 stainless steel at operating \ ; temperature, we obtain from the stress strain curve: i ! (5,),,,,,,,,, - 50,000 psi
- i
\
i For combinations of membrane and bending stresses, the maximum allowable ! stress is taken to be equal to the stress corresponding to the maximum outer i 1 fiber strain not in excess of 405 uniform strain and average strain not in l i ' i excess of 205 uniform strain. i 1 l Blowdown and Force Analysis ! gloudoun Model AMLTIFlas. EGGIM is a digital computer program for calculation of local fluid pressure, flow, and density transients that occur in the reactor primary coolant systems during a loss of coolant accident. This program applies to the subcooled, transition, and saturated two-phase blowdown regimes. This is in contrast to programs, such as WiAM,"' which are applicable on1'y to the subcooled region and which, due to their method of solution, could not be extended into the region in which large changes in the sonic velocities and fluid densities take place. 14.3.3-3
1 N L TIFLAR 4t00Mbe is based on the method of characteristics wherein the resulting set of ordinary differential equations, obtained from the laws of conservation of mass, somentusq, and energy, are solved numerically uti.lizing a fixed mesh in i both space and time. I Although spatially one-dimensional conservation laws are egloyeo, the code
)
can be applied to describe 3-dimensional system geometries through the use of the equivalent piping networks. Such piping networks may contain any number of pipes or channels of various diameters, dead ends, branches (with up to six pipes connected to each branch), contractions, expansions, orifices, pumps, and free surfaces (such as in a pressurizer). All types of the system losses (such as friction, contraction, expansion, etc.) are considered. 'a This h{oresh* Is yredede$ k 1 Connarison With Ernerimental Dat
&N g l 2 predictions have been compared with data obtained by ips Petroleum ny from their LOFT Semi-Scale and % Scale bl experiments.
An example of these coup s is shown in Fi.gur4I14.3.3-1 which illustrates
~
I the pressure history in the bl .ipe-for the Semi-Scale test #522. This j is a ' bottom blowdown test frog the' Bet No.1" geometry with initial i uniform fluid conditj.ons-61268 psia and 445'F. f sseenthattheBLOOWN-2,/ 1 digital /computer program gives good agreement in both t i i ooled and thaj j saturated blowdown regimes. ; ! ~ j 4 { Force Model
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l Sk0094
#pto evaluates the pressure and velocity transients for a maximum of j E .. M ations throughout the system. These pressure and velocity transients i are stored as a perunnent tape file and are made available to the progrant
[4 7444 a*/ FORCEf,which utilize \ a detailed geometric description in evaluating the j loadings on the reactor internals. l Vertis AI j Each reactor component for which g force calculations are required is designated j as an element and assigned an element number. F [ elements are calculated suusing the V""'"I effects of:g orces acting u
]
4 ' 14.3.3-4
I i i j . The pressure differential across the element
- 2. Flow stag (ation on, and unrecovered orifice losses across the element i
- 3. Friction losses along the element
} MM 7'//lMX l Input to the code, in addition to the 4h00tAkt pressure and velocity j transients, includes the effective area of each element on which acts the , force due to the pressure differential across the element, a coefficient to I j account for flow stagnation and unrecovered orifice losses, and the total area g of the element along which the shear forces act. i Vertical Ereitation i l Structural Model and Method of Analysis l beamsk b6 l Theresponseofreactorinternals[:omponentsduetoanexcitationp complete severance of a e .-6 .., pipe is analyzed. Assuming a deshts - ! .anded. pipe break occurs in a very short period of time, the rapid drop of i l pressure at the break produces a disturbance which propagates along the j primary loop and excites the internal structure. i l The internal structure is simulated by a multi-mass system connected with springs and dashpots representing the viscous damping due to structural and i impact losses. The gaps between various components, as well as coulomb type of friction, is also incorporated into the overall model. Since the fuel elements in the fuel assemblies are kept in position by friction forces originating from the preloaded fuel assembly grid fingers, any sliding that occurs between the fuel rods and assembly is considered as coulomb type of friction. A series of mechanical models of local structures have been developed and analyzed so that certain basic nonlinear phenomena previously mentioned could be understood. Using the results of these models, a final eleven-mass model is adopted to represent the internals structure under vertical excitation. Figure 14.3.3-2 is a schematic representation of the internals structures. The eleven-mass model is shown in Figure 14.3.3-3. A comparison between Figure 14.3.3-2 and 14.3.3-3 shows the parallel between the plant and the model. The modeling is conducted in such a way that uniform 14.3.3-5
I I I Insert A (on FSAR page 14.3.3-5 where marked) In addition to the vertical forces calculated by FORCE 2, the horizontal forces on the vessel, core barrel, and thermal shield are calculated by LATFORC. The horizontal forces are calculated by summing the lateral force components around the vessel, core barrel, and thermal shield, based on the pressure differential across each section, multiplied by the area of each section. This is done at ten different elevations. The total lateral force is calculated by summmg the forces over the ten elevations. l l l l l l i
i i i masses are lumped into easily identifiable discrete masses while elastic j elements are represented by springs. A legend for the different masses is
- given in Table 14.3.3-1.
The masses are readily recognized as Items WI j through W11. The core barrel and the lower package are easily discernable, j The fuel assemblies have been segregated into two groups. The majority of the j fuel mass, W4, is indirectly connected to the deep beam structure represented j by mass W8. There is also a portion of the fuel mass, W6, which connects j through the long columns to the top plate. j The stiffness of the top plate panels is represented by KS. The hold down spring, K1, is bolted-up between { the flange of the deep beam structure and the core barrel flange with the l preload, Pl. After proloading the hold down spring, a clearance, G1, exists j i between the core barrel flange and the solid height of the hold down spring. i Within the fuel assemblies, the fuel elements W4 and W6 are held in place by i frictional contact with the grid spring fingers. Coulomb damping is provided ! in the analysis to represent this frictional restraint. 1 i : j The analytical model is also provided with viscous tems to represent the j structural damping of the elastic elements. The viscous dampers are { represented by C1 through C11. 4 i j Restrictions are placed on the displacement amplitudes by specifying the free j travel available to the dynamic masses. Available displacements are designated by symbols G1 through GS. } The displacements are tested during the solution of the problem to see if the ) available travel has been achieved. When the limit of travel has been l attained, stops are engaged to arrest further motion of the dynamic masses. ! The stops of snubbers are designated by the symbols $1 through $11. 1 l contact with the snubbers results in some damping of the motion of the j model. The impact damping of the snubbers is represented by the devices
- D1 through 011.
f g During the assembly of the reactor, bolt-up of the closure head presets the { spring loading of the core barrel and the spring loading on the fuel j assemblies. Since the fuel assemblies in the model have been segregated into two groups, two preload values are provided in the analysis. Preload j j 14.3.3-6
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i i l j values P1, P3, and P5 represent the hold down spring preload on the core barrel and the top nozzle springs preload values on the fuel assemblies. The fomulation of the transient motion response problem and digital computer j programming have been perforised. The effects of an earthquake vertical ) excitation are also incorporated into the program. f ! In order to program the multi-mass system, the appropriate spring rates, l weights, and forcing function for the various masses were determined. The
- spring rates and weights of the reactor components are calculated separately l for each plant. The forcing functions for the masses are obtained from the j ,
FORCE 7 program described in the previous section. It calculates the transient j forces on reactor internals during blowdown using transient pressures and I fluid velocities. 4 j For the blowdown analysis the forcing functions are applied directly to tlie j various internal masses. i ; i i i For the earthquake analysis of the reactor internals, the forcing function, i which is simulated earthquake response, is applied to the eulti-mass systes at the ground connnections (the reactor vessel). Therefore, the external j excitation is transmitted to the internals through the springs at the ground I connections. l Y Rasults l s, led 18l1 l Analysis was performed for vertut$sur4 Pnp 6ere opening time, and for hot leg l and cold leg breaks. The response of the structure to these excitations j indicates that the vertical' notion is irregular with peaks of very short duration. The deflections and action of some of the reactor components are l j limited by the solid height of springs as is also the hold down spring located j above the barrel flange. 1 1 The internals behave as a nonlinear system during the vertical oscillations l I produced by the blowdown forces. The nonlinearities are due to the Coulomb f frictional forces between grids and rods, and to gaps between components causing discontinuities in force transmission. The frequency response is i i I 14.3.3-7 4
j i 1 consequently a function not only of the exciting frequwncies in the system, i but also of the amplitude. i frequencies ig the system. Different break conditions excite different This situation can be seen clearly when the f response under blowdown forces is compared with the one due to vertical seismic acceleration. Under seismic excitation, the system behaves almost i linearly because component notion is not sufficient to cause closing of the various gaps in the structure of slippage in the fuel rods. i Under certain blowdown excitation conditions, the core moves upward, touches l the core plate, and falls down on the lower structure causing oscillations in { all the components. During the time that the oscillations occur and, depending on its initial position, the fuel rods slide on the fuel assembly, { The response shows that the case could be represented as two large vibrating { masses (the core and the barrel), and the rest of the system oscillates with
- respect to the barrel and the core.
l Damping effects have also been considered; it appears that the higher 5 frequencies disappear rapidly after each impact of slippage. The results of the computer program give not only the frequency response of j the components, but also the maximum impact force and deflections. From these j results, the stresses are computed using the standard " Strength of Material" i formulas. The impact stresses are obtained in an analogous manner using the maximum forces seen by the various structures during impact. l i i l i i i i i l 14.3.3-8 l
References - Section 14.3.3 (1) 5. Fabic: " Computer Program WiAM for Calculation of Pressure, Velocity, j and Force Transients in Liquid Filled Piping Networks, " Kaiser Engineers Report No. 67-49-R (November 1967). i Mh4 l l 1 I i i : i i 1 i l 1 i ! ) I 1 1 i, .I 1 i l I
i l L Insert B (on FSAR chapter 14.3.3 References where marked) j (2) K. Takeuchi: " MULTIFLEX, A FORTRAN-IV Computer Program for Analyzing 1 { Dermal-Hydraulic-Structure System Dynamics," WCAP-8708-PA, WCAP-8709-A (Non-Proprietary), September,1977. l l I I l 1 i i I I i i
1 l l TABLE 14.3.3-1 MULTI-MASS VIBRATIONAL MODEl-DEFINITION OF SYMBOLS W1 - Core Barrel K1 - Hold Down Spring W2 - Lower Package K2 - Lower Package Major W3 - Fuel Assemblies Major K3 - Top Nozzle Springs Major W4 - Fuel Rods Major K5 - Top Nozzle Springs Minor W5 - Fuel Asemblies Minor K7 - Short Columns I , W6 - Fuel Rods Minor K8 - Upper Core Plate l W7 - Core P1 ate & Short Column K9 - Long Columns W8 - Deep Beam K10 - Top Plate W9 - Core P1 ate & Long Columns K11 - Core Barrel l W10 - Top Plate (Ctr.) 1 W11 - Core Barrel l Snubbers Imoact Damoers S1 - Core Barrel Flange D1 - Barrel Flange
, S2 - Hold Down Spring D2 - Hold Down Spring S3 - Top Nozzles Bars, Major D3 - Top Nozzle Bars, Major S4 - Pedestal Bars, Major D4 - Pedestal Bars, Major S5 - Top Nozzles Bars, Minor DS - Top Nozzle Bars, Minor 56 - Pedestal Bars, Minor D6 - Pedestal Bars, Minor 3 57 - Top Nozzle Bumpers, Major D7 - Top Nozzles, Major l_1 S8 - Top Nozzle Bumpers, Minor D8 - Top Nozzles, Minor 59 - Pedestals, Major D9 - Pedestal, Major 510 - Pedestals, Minor DIO - Pedestal, Minor S11 - Deep Beam Flange D11 - Deep Beam Flange Structural Damners Clearances t
C1 - Hold Down Springs G1 - Hold Down Spring C2 - Lower Package G3 - Fuel Rod Top, Major C3 - Top Nozzle, Major G4 - Fuel Rod Bottom, Major C5 - Top Nozzle, Minor G5 - Fuel Rod Top, Minor C7 - Short Columns G6 - Fuel Rod Bottom, Minor C8 - Upper Core Plate G7 - Fuel Assembly Major C9 - Long Columns G8 - Fuel Assembly Minor C10 - Top Plate C11 - Core Barrel Preloads P1 - Hold Down Spring P3 - Top Nozzle Springs Major PS - Top Nozzle Springs Minor
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