ML20206U778
| ML20206U778 | |
| Person / Time | |
|---|---|
| Site: | Prairie Island  |
| Issue date: | 06/13/1986 |
| From: | Richard Anderson, Kapitz J, Rautmann D NORTHERN STATES POWER CO. |
| To: | |
| Shared Package | |
| ML20206U771 | List: |
| References | |
| NSPNAD-8102-A, NSPNAD-8102-A-R04, NSPNAD-8102-A-R4, NUDOCS 8607110135 | |
| Download: ML20206U778 (339) | |
Text
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. Q.j@"%,k g . , UNITED STATES I" g 3
e, g g i y ,f/ j 7, NUCLEAR REGULATORY COMMISSION WASHINGTON, D. C. 20$55
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Docket Nos. 50-282/306 I Mr. D. M. Musoli, Manager - - Nuclear Support Services Northern States Power Company I 414 Nicollet Mall Midland Square, 4th Floor Minneapolis, Minnesota 55401 I
Dear Mr. Musolf:
SUBJECT:
TOPICAL REPORT RELOAD SAFETY EVALUATION METHODS, NSPNAD-8102P, I Rev. 3, (VIPRE-01) PRAIRIE ISLAND NUCLEAR GENERATING PLANT UNIT NOS. 1 AND 2 I We have completed our review of Northern States Power Company's Reload Safety Evaluation Methods, NSPNAD-8102P, Rev. 3, dealing with the thermal margin analysis for the fuel assemblies at the Prairie Island Nuclear Generating
- Plant Unit Nos. I and 2. By letter dated April 19, 1985 and additional l information provided by letters dated November 19, and December 5, 1985, you l I submitted Revision 3 to NSPNAD-8102P describing the applications of the i VIPRE-01 Code in place of COBRA-IIIC/MIT and the WRB-1 Critical Heat Flux
.l correlation in place of W-3 for analyzing the thermal margin for fuel reloads at Prairie Island. We find that the use of the VIPPE-01 thermal hydraulic code and WRB-1 Critical Heat Flux (CHF) correlation with a minimum departure from nucleate I boiling ratio (DNBR) limit of 1.17 for the Prairie Island Units is acceptable. We will require that you follow the EPRI quality assurance I procedure regarding the VIPRE-01 modifications. In addition, the acceptance criterion for the condition IV events appearing in revision 3 of NSPNAD-8102P are to be replaced with the original criterion that would include the DNBR limit of 1.17 for the WRB-1 correlation. Five copies of revision 4 of I I l I ((g
.' (
NSPNAD-8102P containi Dnhe original criterion should be forwarded to the Commissio'n within 30 days _from receipt date of this letter. We consider our review of % e VIPRE.01 Code and the WRB-1 (CHF) complete. A copy of our Safety Evaluation is enclosed. Sincerely, SQ. C L Dominic C. DiIanni, Project Manager I Project Directorate #1 Division of PWR Licensing-A l
Enclosure:
Safety Evaluation cc's: See Next Page I; l I! : I! I I I E I I I! I
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Mr. D. M. Musolf Prairie Island Nuclear Generating Northern States Power Company Plant I cc: Gerald Charnoff, Esq. I Shaw, Pittman, Potts and Trowbridge 1800 M Street, NW Washington, DC 20036 Executive Director Minnesota Pollution Control Agency 1935 W. County Road, B2 Roseville, Minnesota 55113 I . . Mr. E. L. Watzl, Plant Manager Prairie Island Nuclear Generating Plant Northern States Power Company i Route 2 Welch, Minnesota 55089 Jocelyn F. Olson, Esq. Special Assistant Attorney General Minnesota Pollution Control Agency I 1935 W. County Road, B2 Roseville, Minnesota 55113 U.S. Nuclear Regulatory Commission I Resident Inspector's Office Route #2, Box 500A Welch, Minnesota 55089 I Regional Administrator, Region III U.S. Nuclear Regulatory Commission Office of Executive Director for 5 Operations 799 Roosevelt Road Glen Ellyn, Illinois 60137 Mr. William Miller, Auditor Goodhue County Courthouse Red Wing, Minnesota 55066 I I l I 1 8
I
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;....,A UNITED STATES 2 s, g NUCLEAR REGULATORY COMMISSION
$) 7 j wasmucTON, D. C. 20555
' . .Q. . . g SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION g RELATED TO THE VIPRL-01 CODE AND WRB-1 CORRELATION 5 FOR FACILITY OPERATING LICENSE NO3. DPR-42 AND 60 NORTHERN STATES POWER COMPANY PRAIRIE ISLAND UNITS 1 AND 2 DOCKET NOS. 50-282 AND 50-306 , ,
1.0 INTRODUCTION
By letter dated April 19, 1985 (Ref. 1), Northern States Power Company (NSP) submitted Revision 3 to NSPNAD-8102P, " Reload Safety Evaluation Methods for Application to Prairie Island Units," for staff review. Revision 3 includes changes to: (1) replace the thermal-hydraulic hot channel code COBRA-IIIC/MIT with the VIPRE-01 Code, (2) add the WB-1 Critical Heat Flux (CHF) correlation f'r ouse with Westinghouse OFA fuel, and (3) revise the acceptance criteria for Condition IV accidents. Our evaluation with regard to these revisions follows. 2.0 STAFF EVALUATION I 2.1 Use of VIPRE-01/WRB-1 The thermal margin analyses for the Prairie Island Units were done with the COBRA-IIIC/MIT (Ref. 2) subchannel thermal-hydraulic code and the W-3 CHF correlation. In the new reload methods, NSP proposes to use VIPRE-01 (Ref. 3) in place of COBRA-IIIC/MIT and use the WRB-1 CHF correlation in place of W-3. VIPRE-01 is a subchannel thermal-hydraulic code developed by Battelle Pacific Northwest Laboratories under the sponsorship of the Electric Power Research Institute (EPRI). In December 1984, the Utility Group I I
l l 1
; 1 I. for Regulatory Applications (UGRA), which consists of more than 20 utilities, submitted the VIPRE-01 topical reports for staff review.
VIPRE-01 was developed from the COBRA series code including COBRA-IIIC (Ref. 4), COBRA-IV (Ref. 5), COBRA-IIIC/MIT and COBRA-WC (Ref. 6) by I incorporating many features of these codes into one package. The staff review (Ref. 7) has concluded that the VIPRE-01 code is acceptable for PWR application with the following conditions: 5 (1) The application is limited to the heat transfer modes up to critical heat flux. (2) An analysis is made to ensure that the minimum departure from nucleate boiling ratio (DNBR) limit of a CHF correlation used in VIPRE-01 can pred.ict its data base of DNB occurrence with at least I a 95 percent probability at a 95 percent confidence level. (3) Documentation is submitted by each user to provide justification for the modeling assumptions, choice of particular two phase flow j models, correlations and input values of plant specific data, etc. 1 (4) If a profile fit subcooled boiling model which was developed based on steady state data is used.in boiling transients, care should be taken in the time step size used for transient analysis to avoid the Courant number less than 1. (5) Each user should abide by the quality assurance program established 5 by EPRI for the VIPRE-01 code. E Appendix F to NSPNAD-8102P, Revision 3, provides a description of the intended use of VIPRE-01 by NSP, by comparing the analysis results using VIPRE-01 and COBRA-IIIC/MIT, and an analysis of the DNBR limit of the WRB-1 correlation and the NSP thermal margin methodology using VIPRE-01. 5 I I
- E A summary of VIPRE input information is provided in Table F.5 regarding the single phase friction factor, two-phase frictica multiplier, subcooled void and bulk void correlations, turbulent mixing model, transverse momentum factor and crossflow resistance factor. Most of the input values and correlations selected are consistent with those used g
in the previous analysis using COBRA-IIIC/MIT. The heat transfer E correlations with respect to forced convection, and subcooled and saturated nucleate boiling chosen in VIPRE-01 and the reactor core , modeling are also the same as for COBRA-IIIC/MIT as indicated in the NSP response to a staff question (Ref. 8). One difference from the COBRA-IIIC/MIT input is the subchannel turbulent mixing. Turbulent mixing was ignored in the previous NSP analysis due to the limitation g that the COBRA-IIIC/MIT turbulent mixing model is not adequate for W mixing between lumped assemblies. Since VIPRE-01 has the capability of specifying the turbulent mixing parameters for each channel connector, the analysis will consider the turbulent mixing effect between subchannels while ignoring turbulent mixing between lumped assemblies or lumped assemblies and subchannels. A turbulent momentum factor (FTM) of 0.8 will be used in the analysis. The turbulent momentum factor, which has a value from 0 to 1.0, is analogous to the turbulent Prandt1 B 5 number. An FIM value of 0.0 indicates that the turbulent crossflow mixes enthalpy only and not momentum, and an FTM value of 1.0 indicates that it mixes momentum with the same strength as it mixes enthalpy. VIPRE-01 is not very sensitive to the value of FTM and the VIPRE-01 manual recommends that a value of 0.8 be used in the analysis. Therefore, the NSP approach is acceptable. 4 In order to assess the VIPRE-01 capability, Appendix F provides comparisons between the calculations of VIPRE-01 and the approved COBRA-IIIC/MIT code. The analyses were performed for Prairie Island-1, Cycle 9 using the W-3 CHF correlation. For steady-state calculations, comparisons are made between the VIPRE-01 and COBRA-IIIC/MIT analysis. I
1 t results on the core safety limit and the axial offset effect. The core
~
safety limits are the loci of points of thermal power, pressurizer
'p'ressure and inlet temperature for which the minimum DNBR limit is not violated. These core safety limit curves are then used in the g derivation of the over temperature AT (OTAT) trip setpoints to ensure g that the DNBR limit will not be violated for normal operation and the anticipated operational transients.
Since the OTAT setpoints are derived with zero axial offset, an axial offset penalty function f(AI) is used to lower the trip setpoints when highly skewed axial power shapes are encountered. The resulting core I thermal limits (core inlet temperature as a function of thermal power and pressurizer pressure) and the average heat flux for DNB occurrence for various axial offsets are shown in Figures F.1, F.2, F.7 and F.8 with comparisons between the VIPRE-01 and COBRA-IIIC/MIT analysis results. These comparisons show that the VIPRE-01. calculations are either the same as or slightly conservative to the COBRA-IIIC/MIT calculations. l i 1 The transient comparisons between VIPRE-01 and COBRA-IIIC/MIT are made l l for the rod withdrawal at power, turbine trip, 2/2 pump trip, and locked rotor events. The results of minimum DNBR as a function of time are shown in Figures F.3 through F.6. The comparisons show that the calculations from both codes are essentially the same for rod withdrawal i and turbine trip cases and the VIPRE-01 calculations are slightly more l conservative for the pump trip and locked rotor cases. NSP attributed g this conservatism to a slightly higher crossflow out of the hot channel in the VIPRE-01 calculations which leads to higher local quality and therefore lower CHF and DNBR. The overall comparisons show VIPRE-01 to be conservative relative to the approved COBRA-IIIC/MIT code and is therefore, acceptable. I I
l l l The WRB-1 CHF correlation was developed by Westinghouse using the THINC thermal hydraulic code. WRB-1 has previously been reviewed and approved for application to the Westinghouse standard low parasitic fuel and the optimized fuel assembly (Ref. 9, 10, 11). A minimum DNBR limit of 1.17 ; I is acceptable for both the standard R-grid assembly and 0FA fuel g designs. NSP obtained the WRB-1 correlation from Westinghouse and u incorporated it into the VIPRE-01 code. Since the DNBR limit of 1.17 is acceptable for WRB-1 in connection with the THINC code, apalysis must be , done to show that the same limit provides at least the same degree of assurance in the DNB prediction when WRB-1 is used with the VIPRE-01 code. NSP has re-analyzed the Westinghouse CHF test data obtained with bundle geometry representative of the 14x14 OFA design using the VIPRE-01/WRB-1 J package. The results of the analysis for the measured-to predicted CHF ratios (M/P) are shown in Table F.2. Based on these measured CHF to
. predicted CHF (M/P) ratios and the use of standard statistical methods widely used in the industry, a minimum DNBR limit is obtained which would ensure avoidance of LFB with 95 percent probability at a 95 g
percent confidence level. Since this DNER limit is less than 1.17, the y use of 1.17 as the DNBR limit for VIPRE-01/WRB-1 is acceptable. NSP has indicated that the WRB-1 correlation will be applied to the improved 0FA fuel design to be loaded in the Prairie Island units. The improved 0FA is essentially the same as the OFA design except for a six inch natural uranium axial blanket at the top and bottom of the fuel rod in the improved 0FA. Other design features which are important to the CHF behavior such as grid design, grid spacing, pin diameter, etc., remain the same for both the OFA and improved 0FA. Therefore, we conclude that the application of WRB-1 to the improved 0FA is acceptable. The staff has asked what steps will be taken by NSP to assure that only the approved version of VIRPE-01 will be used in licensing analyses. l I I
I . I 6-NSP in its response (Ref. 8) stated that the VIPRE code will be controlled according to NSPNAD Policies and Procedures NAP 5.001A, Revision 4, " Computer Program Control" which covers the use of codes and making modifications to the codes. This computer code control procedure I has been audited by NRC in 1983. NSP provided a brief description of the control procedure which we find to be ecceptable. However, this procedure deals with only the modifications to the code inside NSP. Since VIPRE-01 is an EPRI code developed by PNL to be used by the . utilities belonging to the UGRA, improvements and modifications to the code will be made by the utilities other than NSP and PNL. Therefore, we will require that NSP also abide by the quality control procedures g established by EPRI which the UGRA committed to follow for the VIPRE E code. 2.2 Changes to the Acceptance Criteria of Coedition IV Accidents:
, Revision 3 to NSPNAD-8102P proposes to revise the acceptance criteria for a few Condition IV events including locked rotor, steamline break and control rod ejection events. The proposed revisions are as follows:
(1) Eliminate an existing criterion which states: "The number of fuel rods calculated to experience a DNER of less than 1.3 should not exceed the number which are required to fail in order that the doses due to released activity will exceed the limit of 10 CFR Part 100. This limit is currently the maximum number of failed fuel I rods calculated in the FSAR." The proposed revision also eliminates an equivalent criterion for reload analysis: " number of fuel pins above Fg (DNBR = 1.3) < 207.." I (2) Change the maximum clad temperature limit from 2750*F to 2700*F. I 5 I
The staff has found that the first revision with regards to the elimination of the released activity dose criteria conforming to 10 CFR Part 100 is not acceptable. In a telephone conversation on April 7, 1986, NSP decided that it will withdraw the proposed enanges to the acceptance criteria including the change to the maximum cladding temperature. Since the DNPR of 1.3 in the original acceptance criteria.was for the . W-3 correlation which does not reflect the DNBR limit of 1.17 for WRB-1, the licensee has agreed to re-install the original criteria into Revision 4 to NSPNAD with an exception that the DNBR limit of 1.3 (W-3) will be changed by adding a DNBR limit of 1.17 for WRB-1. We find this to be acceptable.
- 3.
SUMMARY
AND CONCLUSIONS The staff has reviewed Revision 3 to NSPNAD-8102P. We find that the use of the VIPRE-01 subchannel thermal hydraulic code and the WRB-1 CHF correlation with a minimum DNBR limit of 1.17 for the Prairie Island units is acceptable. We will also require that NSP abide by the EPRI quality assurance procedure regarding the VIPRE-01 modifications. With regard to the acceptance criteria for the Condition IV events, NSP has decided to withdraw its proposed revisions and has agreed to install the original criteria into Revision 4 to NSPNAD-8102P with an addition of a DNBR limit of 1.17 for the WRB-1 correlation. We find this to be acceptable. Principal Contributor: Hsii Yi Hsiung 4 I l I
1 REFERENCES I 1. Letter from D. Musolf (Northern States Power Company), to Director, Office of Nuclear Reactor Regulation, USNRC " Proprietary (and non-proprietary version) Topical Report Reload Safety Evaluation Methods, NSPNAD-8102P, Rev. 3," April 19,1985.
- 2. Bowring, R. W., and P. Moreno, " COBRA-IIIC/MIT Computer Code Manual "
Prepared by Massachusetts Institute of Technology for EPRI, March 1976.
- 3. EPRI-NP-2511-CCM, "VIPRE-01, A Thermal-Hydraulic Analysis Code for I Reactor Cores," Volumes 1 through 3, EPRI, April 1983, Revision 1, November 1983, Revision 2. July 1985.
- 4. D. S. Rowe, "C0 EPA-IIIC: A Digital Computer Program for Steady-State I and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," Richland, Washington: Pacific Northwest Laboratory, March 1973, BNWL-1695.
I ,
- 5. C. L. Wheeler, et al., " COBRA-IV-I: An Interim Version of COBRA for Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements and ,
Cores," Richland, Washington: Pacific Northwest Laboratory, March 1973, BNWL-1962.
- 5. T. L. George, et al., " COBRA-WC: A Version of COBRA for Single-phase i Multiassembly Thermal-Hydraulic Transient Analysis," Richland, Washington: Pacific Northwest Laboratory, July 1980, PNL-3259.
- 7. Letter from C. E. Rossi (NRC) to J. A. Blaisdell (NUSCO), " Acceptance for Reference of Topical Report, EPRI NP-2511-CCM, VIPRE-01: A Thermal
- i. Hydraulic Analysis Code for Reactor Cores," Volumes 1, 2, 3 and 4, May 1, 1986.
- 8. Letter from D. Musolf (NSP) to Director, Office of NRR, USNRC, " Responses I of NRC Questions on Revision 3 of Topical Report NSPNAD-8102P," November 19, 1985.
I I
l
- 9. WCAP-8762(P), WCAP-8703(NP), "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids,"
July 1976.
- 10. Letter from R. L. Tedesco (NRC) to T. M. Anderson (Westinghouse),
" Acceptance for Referencing Topical Report-WCAP-9401(P)/WCAP-9402(NP),"
May 15, 1981.
- 11. Letter from C. O. Thomas (NRC) to E. P. Rahe, Jr., (Westinghouse)
" Acceptance for Referencing of Licensing Topical Report - -
WCAP-8762(P)/WCAP-8783(NP), Supplement 1, Basis for the Applicability of the WRB-1 Correlation to 15x15 0FA and 14x14 0FA Fuel," June 29, 1984. I _ l
. _i -
5 I I I I I I I
I ? I I i RELOAD SAFETY EVALUATION METHODS FOR APPLICATION TO PRAIRIE ISLAND UNITS NSPNAD-8102-A Revision 4 l June 1986 ,I ~ l i PREPARED BY % O& DATE b//3 $d REVIEWED BY 1W h e u44 OATE b /3 8b APPROVED BY L1 h, M vh^ft DATE [
/3 7b E
lI l Page 1 of 332
I I I E ABSTRACT I This document is a Topical _ port describing the Northern States Power Company (NSP) reload safety evaluation methods for application to the Prairie Island Units. The report addresses the methods for the calculation of cycle specific physics parameters and their comparison to the bounding values used in the accident analyses. In addition, a brief summary is presented of the NSP safety analysis experience and calculational results for Prairie Island. I I I I I I Page 2 of 332 I
'I I I LEGAL NOTICE I This report was prepared by or on behalf of Northern States Power Company (NSP). Neither NSP, nor any person acting on behalf of NSP:
- a. Makes any warranty or representation, express or implied, with respect to the accuracy, completeness, usefulness, or use of any information, apparatus, method or process cisclosed or contained I
in this report, or that the use of any I such information, apparatus, method, or process may not infringe privately owned rights; or
- b. Assumes any liabilities with respect to I the use of, or for damages resulting from the use of, any information, apparatus, method, or process disclosed I. in the report.
5 I Page 3 of 332
TABLE OF CONTENTS Page
1.0 INTRODUCTION
14 2.0 GENERAL PHYSICS METHODS 16 E ) 2.1 Moderator Temperature Reactivity Coefficient 16 2.2 Power Reactivity Coefficient 16 2.3 Doppler Reactivity Coefficient 17 2.4 Boron Reactivity Coefficient 17 2.5 Shutdown Margin 18 2.6 Scram Reactivity Curve 19 2.7 Nuclear Heat Flux Hot Channel Factor 20 2.8 Nuclear Enthalpy Rise Hot Channel Factor 20 2.9 Effective Delayed Neutron Fraction 20 2.10 Prompt Neutron Lifetime 21 3.0 SAFETY EVALUATION METHODS 22 1 3.1 Uncontrolled RCCA Withdrawal from Suberitical 26 l 3.2 Uncontrolled RCCA Withdrawal at Power 32 3.3 Control Rod Misalignment 53 3.4 Dropped Control Rod 56 3.5 Uncontrolled Boron Dilution 81 3.6 Startup of an Inactive Loop 90 3.7 Feedwater System Malfunc' tion 99 3.8 Excessive Load Increase 114 3.9 Loss of External Load 118 3.10 Loss of Normal Feedwater 147 4 Pa,e 4 o, m I E _. I
I TABLE OF CONTENTS Page 3.11 Loss of Reactor Coolant Flow - Pump Trip 153 3.12 Loss of Reactor Coolant Flow - Locked Rotor 167 3.13 Fuel Handling Accident 179 3.14 Main Steam Line Break 181 3.15 Ejected Rod 215 3.16 Loss of Coolant 231 3.17 Fuel Misloading 238 1
4.0 REFERENCES
261 Appendix A - Description of NSP Safety Analysis Computer 262 Progr1m Package Appendix B - NSSS Simulation Methods 266 Appendix C - Thermal-Hydraulic Methodology 270 Appendix D - Fuel Thermal Response Methodology 292 l Appendix E - Containment Analysis Methods 299 Appendix F - VIPRE-01 Thermal Response Methodology 306 8 I .I I 1 5 3 ea,e e of m I
l LIST OF TABLES Table Page 3.4-1 PI 2 Cycle 8 Dropped Rod Results 66 3.4-2 PI 1 Cycle 9 Dropped Rod Results 67 ! l 3.14-1 Steady State Conditions for Hot Channel Analysis of 190 ; Steam Line Break. ; 3.15-1 Comparison of Rod Ejection Maximum Fuel Rod Enthalpies 223 1 and Temperatures. 3.17-1 Summary of Fuel Misloadings Analyzed 243 C-1 COBRA IIIC/MIT Single Channel Model 281 C-2 COBRA IIIC/MIT 1/8 Assembly Model 282 C-3 COBRA IIIC/MIT 1/8 Core Model 283 E-1 Structural Heat Sinks 303 E-2 S. G. Mass and Energy Blowdown 304 F-1 Locked Rotor Failed Pin Analysis 316 F-2 VIPRE-01/WRB-1 4x4 0FA Test Results 317 F-3 Axial Power Shape Study 318 F-4 Time Step Sensitivity Results 319 F-5 VIPRE-01 1/8 Core Model 329 I I I R I Page 6 of 332 I
.I LIST OF FIGURES Figure Page 3.1-1 Uncontrolled Control Rod Withdrawal from Subcritical 32 3.1-2 Uncontrolled Control Rod Withdrawal from Subcritical 33 3.1-3 Uncontrolled Control Rod Withdrawal from Subcritical 34 3.1-4 Uncontrolled Control Rod Withdrawal from Subcritical 35 3.1-5 Uncontrolled Control Rod Withdrawal from Subcritical 36 3.1-6 Uncontrolled Control Rod Withdrawal from Subcritical 37 Peak Heat Flux vs Initial Power Level l 3.2-1 Uncontrolled Centrol Rod Withdrawal from Full Power - 45 Fast Rate 3.2-2 Uncontrolled Control Rod Withdrawal from Full Power - 46 I
Fast Rate l 3.2-3 Uncontrolled Control Rod Withdrawal from Full Power - 47 Fast Rate 3.2-4 Uncontrolled Control Rod Withdrawal from Fu]1 Power - 48 Fast Rate - I 3.2-5 Uncontrolled Control Rod Withdrawal from Full Power - Slow Rate 49 3.2-6 I Uncontrolled Control Rod Withdrawal from Full Power - Slow Rate 50 3.2-7 Uncontrolled Control Rod Withdrawal from Full Power - 51 i Slow Rate l 3.2-8 Uncontrolled Control Rod Withdrawal from Full Power - 52 Slow 3.4-1 PI 2 Cycle 8 Dropped Rod - EOC 68 3.4-2 PI 2 Cycle 8 Dropped Rod - EOC 69 3.4-3 PI 2 Cycle 8 Dropped Rod - EOC 70 3.4-4 PI 2 Cycle 8 Dropped Rod - EOC 71 3.4-5 PI 2 Cycle 8 Dropped Rod - E0C 72 5 Page 7 of 332
I LIST OF FIGURES Figure Page 3.4-6 PI 1 Cycle 9 Dropped Rod - E0C 73 , 3.4-7 PI 1 Cycle 9 Dropped Rod - EOC 74 3.4-8 PI 1 Cycle 9 Dropped Rod - EOC 75 3.4-9 PI 1 Cycle 9 Dropped Rod - EOC 76 3.4-10 PI 1 Cycle 9 Dropped Rod - EOC 77 3.4-11 Functional Block Diagram - Automatic Rod Control System 78 3.4-12 PI 1 Cycle 9 Dropped Rod - EOC - Rod Controller Response 79 3.4-13 PI 1 Cycle 9 Dropped Rod - EOC - Rod Controller Response 80 : 3.5-1 Chemical and Volume Control System Malfunction 86 3.5-2 Chemical and Volume Control System Malfunction 87 3.5-3 Chemical and Volume Control System Malfunction 88 3.5-4 Chemical and Volume Control System Malfunction 89 3.6-1 Start-up of an Inactive Coolant Loop 94 3.6-2 Start-up of an Inactive Coolant Loop 95 3.6-3 Start-up of an Inactive Coolant Loop 96 3.6-4 Start-up of an Inactive Coolant Loop 97 1 3.6-5 Start-up of an Inactive Ccolant Loop 98 3.7-1 Decrease in Feed Water Temperature - BOC Without 104 1 Reactor Control 3.7-2 Decrease in Feed Water Temperature - BOC Without 105 ' Reactor Control 3.7-3 Decrease in Feed Water Temperature - BOC Without 106 Reactor Control 3.7-4 Decrease in Feed Water Temperature - BOC Without 107 Reactor Control 3.7-5 Decrease in Feed Water Temperature - BOC Without 108 Reactor Control Page 8 of 332 I
I LIST OF FIGURES Figure Page
's.7-6 Decrease in Feed Water Temperature - EOC with 109 Reactor Control -
I 1.7-7 Decrease in Feed Water Temperature - EOC with 110 Reactor Control 3.7-8 Decrease in Feed Water Temperature - EOC with 111 . Reactor Control 3.7-9 Decrease in Feed Water Temperature - E0C with 112 Reactor Control 3.7-10 Decrease in Feed Water Temperature - E0C with 113 I Reactor Control 3.9-1 Loss of External Load - BOC with Reactor Control 123 3.9-2 Loss of External Load - BOC with Reactor Control 124 3.9-3 Loss of External Load - BOC with Reactor Control 125 3.9-4 Loss of External Load - BOC with Reactor Control 126 3.9-5 Loss of External Load - BOC with Reactor Control 127 l 3.9-6 Loss of External Load - BOC with Reactor Control 128 3.9-7 Loss of External Load - EOC with Reactor Control 129 3.9-8 Loss of External Load - EOC with Reactor Control 130 3.9-9 Loss of External Load - EOC with Reactor Control 131 3.9-10 Loss of External Load - EOC with Reactor Control 132 3.9-11 Loss of External Load - EOC with Reactor Control 133 3.9-12 Loss of External Load - EOC with Reactor Control 134 3.9-13 Loss of External Load - BOC Without Reactor Control 135 lI l5 3.9-14 Loss of External Load - BOC Without Reactor Control 136 3.9-15 Loss of External Load - BOC Without Reactor Control 137 5 3.9-16 Loss of External Load - BOC Without Reactor Control 138 i j 3.9-17 Loss of External Load - BOC Without Reactor Control 139 Il 1 Page 9 of 332
1 l' LIST OF FIGURES Figure Page 3.9-18 Loss of External Load - BOC Without Reactor Control 140 3.9-19 Loss of External Load - EOC Without Reactor Control 141 3.9-20 Loss of External Load - EOC Without Reactor Control 142 3.9-21 Loss of External Load - EOC Without Reactor Control 143 3.9-22 Loss of External Load - ECC Without Reactor Control 144 3.9-23 Loss of External Load - EOC Without Reactor Control 145 3.9-24 Loss of External Load - EOC Without Reactor Control 146 3.10-1 Loss of Normal Feed Water 150 3.10-2 Loss of Normal Feed Water 151 3.10-3 Loss of Normal Feed Water 152 , 3.11-1 Loss of RCS Flow - 1/2 Pump Trip 159 3.11 Loss of RCS Flow - 1/2 Pump Trip . 160 3.11-3 Loss of RCS Flow - 1/2 Pump Trip 161 3.11-4 Loss of RCS Flow - 1/2 Pump Trip , 162 3.11-5 Loss of RCS Flow - 2/2 Pump Trip 163 3.11-6 Loss of RCS Flow - 2/2 Pump Trip 164 3.11-7 Loss of RCS Flow - 2/2 Pump Trip 165 3.11-8 Loss of RCS Flow - 2/2 Pump Trip 166 3.12-1 Loss of RCS Flow - Locked Rotor 173 3.12-2 Loss of RCS Flow - Locked Rotor 174 3.12-3 Loss of RCS Flow - Locked Rotor 175 3.12-4 Loss of RCS Flow - Locked Rotor 176 3.12-5 Loss of RCS Flow - Locked Rotor 177 3.12-6 Loss of RCS Flow - Locked Rotor 178 8 . I
I LIST OF FIGURES l Fioure Page 3.14-1 Main Steam Line Break - At S.G. Exit with A.C. 191 3.14-2 Main Steam Line Break - At S.G. Exit with A.C 192 3.14-3 Main Steam Line Break - At S.G. Exit with A.C 193 3.14-4 Main Steam Line Break - At S.G. Exit with A.C 194
, 3.14-5 Main Steam Line Break - At S.G. Exit with A.C. 195 3.14-6 Main Steam Line Break - At S.G. Exit with A.C. 196 I 3.14-7 Main Steam Line Break - Downstream of Flow Restrictor with A.C.
197 3.14-8 Main 5 team Line Break - Downstream of Flow Restrictor 198 with A.C. 3.14-9 Main Steam Line Break - Downstream of Flow Restrictor 199 with A.C. 3.14-10 Main Steam Line Break - Downstream of Flow.Restrictor 200 with A.C. I 3.14-11 Main Steam Line Break - 06wnstream of Flow Restrictor 201 l with A.C. 3.14-12 Main Steam Line Break - At S.G. Exit Without A.C. 202 3.14-13 Main Steam Line Break - At S.G. Exit Without A.C. 203 3.14-14 Main Steam Line Break - At S.G. Exit Without A.C. 204 3.14-15 Main Steam Line Break - At S.G. Exit Without A.C. 205 3.14-16 Main Steam Line Break - At S.G. Exit Without A.C. 206 3.14-17 1 Main Steam Line Break - Downstream of Flow Restrictor without A.C. 207 3.14-18 Main Steam Line Break - Downstream of Flow Restrictor 208 I without A.C. 3.14-19 Main Steam Line Break - Downstream of Flow Restrictor 209 I without A.C. 3.14.20 Main Stesm Line Break - Downstream of Flow Restrictor 210 l without A.C. I I Page 11 of 332
I LIST OF FIGURES Figure Page 3.14.21 Main Steam Line Break - Downstream of Flow Restrictor 211 without A.C. 3.14-22 Main Steam Line Br?ak - 247 lbm at 1100 PSIA with A.C. 212 I 3.14-23 Main Steam Line Break - 247 lbm at 1100 PSIA with A.C. 213 3.14-24 Main Steam Line Break - 247 lbm at 1100 PSIA with A.C. 214 , 3.15-1 Rod Ejection - HZP, BOL 225 3.15-2 Rod Ejection - HZP, BOL 226 3.15-3 Rod Ejection - HFP, BOL 227 i 3.15-4 Rod Ejection - HFP, BOL 228 3.15-5 Rod Ejection - HZP, EOL 229 I 3.15-6 Rod Ejection - HFP, EOL 230 l I 3.17.1 Misloaded Bundle - Case A 244 3.17.2 Misloaded Bundle - Care B 245 3.17.3 Misloaded Bundle - Case C 246 3.17.4 Misloaded Bundle - Case D 247 3.17.5 Misloaded Bundle - Case E 248 ; 3.17.6 Misloaded Bundle - Case F 249 3.17.7 Misloaded Bundle - Case G 250 3.17.8 Misloaded Bundle - Case H 251 1 3.17.9 Misloaded Bundle - Case I 252 3.17.10 Misloaded Bundle - Case J 253 l 3.17.11 Misloaded Bundle - Case K 254 E i 3.17.12 Misloaded Bundle - Case L 255 3.17.13 Misloaded Bundle - Case M 256 j 3.17.14 Misloaded Bundle - Case N 257 Page 12 of 332 li i I
I LIST OF FIGURES Fiqure Page 3.17.15 Misleaded Bundle - Case 0 258 3.17.16 Misloaded Bundle - Case P 259 3.17.17 Misloaded Bundle - Case Q 260 C-1 COBRA IIIC/MIT 1/8 Core Model Channel Layout 285 C-2 COBRA IIIC/MIT 1/8 Core Model Channel Layout 286 C-3 COBRA IIIC/MIT Axial Power Shape 287 C-4 COBRA IIIC/MIT Axial Flow vs Channel Length 288 C-5 COBRA IIIC/MIT Mass Flux Convergence 289 C-6 COBRA IIIC/MIT Enthalpy Convergence 290 C-7 COBRA IIIC/MIT Crossflow Convergence 291 D-1 Locked Rotor - Clad Temperature vs Time 298 E-1 Containment Fan Cooler Heat Removal Rate 305 F-1 Thermal Overtemperature Limits 321 F-2 Thermal Overtemperature Limits 322 F-3 Prairie Island 1 Cvele 9 - Turbine Trip 323 F-4 Prairie Island 1 Cycle 9 - Fast Rod Withdrawal 324 F-5 Prairie Island 1 Cycle 9 - 2/2 Pump Trip 325 F-6 Prairie Island 1 Cycle 9 - Locked Rotor 326 F-7 Axial Power Profile Study 327 F-8 Axial Power Profile Study 328 F-9 Prairie Island 1 Cycle 9 - Turbine Trip 329 F-10 Prairie Island 1 Cycle 9 - Slow Rod Withdrawal 330 F-11 Prairie Island 1 Cycle 9 - Fast Rod Withdrawal 331 F-12 Prairie Island 1 Cycle 9 - 2/2 Pump Trip 332 1 Page 13 of 332 8
I
1.0 INTRODUCTION
This report addresses the methods for the calculation of Prairie Island cycle specific physics parameters and their comparison to the bounding values used in the accident analyses. A brief description of the general physics calculational procedures is reviewed in Section 2. General methods are described for each of the key physics parameter.: of interest in reload safety evaluations. Cycle specific physics calculations and their comparisons to the g safety analyses are described for each accident in Section 3. The W specific applications of the reliability factors descr' bed in Reference 1 are also presented in this section. A general description is given in Section 3 of each of the accidents that are sensitive to physics parameters and is therefore of concern for a reload evaluation. For each accident, a discussion of the general input assumptions, consequences and sensitivities to various physics characteristics is provided. Calculations of core physics parameters for the purpose of performing I reload safety evaluations requires an intimate knowledge of the safety analyses to which cycle specific comparisons are to be made. Specifically, one must understand the manner in which the bounding g physics parameters have been used in each of the analyses and the W conservatisms inherent in the values chosen. In order to acquire such an understanding, Northern States Power (NSP), in conjunction with Nuclear Associates International (NAI), has developed models for performing various safety analyses for Prairie Island and has per-formed representative calculations for the incidents of importance for ' a reload evaluation. A summary of the results of these calculations g is included in Section 3 to demonstrate NSP safety analysis experience W and to exemplify the expertise required to make the determinations as to whether or not an accident must be re-analyzed and to perform the necessary analyses for a given fuel cycle. Page 14 of 332 8
I I A determination of those analyses which are affected by a reload design has been performed, Section 3.0 identifies the analyses which I require review and itemizes the physics parameters that change necessitating an analysis review. The specific bounding values for each analysis are provided in the cycle specific Reload Safety ] Evaluation Report utilizing the most up-to-date analysis. R 3 Appendix A gives an overview of the computer code package that is used to simulate the transients and accidents listed in this report. A I discussion of computer code uncertainties is also included in this section. Appendix B gives a description of the DYN0DE-P computer code which is used to simulate the transient response of the Nuclear Steam Supply System (NSSS). I Appendix C gives a description of the COBRA IIIC/MIT computer code which is used to simulate the thermal-hydraulic response of the hot . coolant channel. A discussion of the NSP thermal margin methodology is also included in this appendix. I Aopendix 0 gives a description of the T00DEE2 computer code which I is used to simulate the thermal response of the hot fuel rod and associated coolant channel under transient conditions. A discussion of the NSP fuel thermal response methodology is also included in this appendix. Appendix E gives a description of the CONTEMPT-LT/026 computer code which is used to simulate the transient response of the containment. I A discussion of the NSP containment analysis methodology is also included in this appendix. Appendix F gives a description of the VIPRE-01 computer code which is to replace COBRA-IIIC/MIT for simulating the thermal hydraulic response of the hot channel. A discussion of the NSP thermal margin methodology using VIPRE-01 is also included. I Page 15 of 332 5
I 2.0 GENERAL PHYSICS METHODS In this section the general physics calculational methods are de-scribed for application to reload safety evaluations for Prairie Island. Cycle specific calculations, the application of reliability factors, biases and comparisons to the safety analyses are discussed in Section 3 for each accident considered. Reference 6 contains detailed procedures for calculating the cycle specific parameters for each accident. 2.1 Moderator Temperature Reactivity Coefficient, og Definition: a M is the change in core reactivity associated with a 1 F change in average moderator temperature at constant average fuel temperature. Calculations of a gare performed in three dimensions with the nodal model (1). The average moderator temperature is varied while the independent core parameters such as core power level, control rod position and RCS boron concentratien are held constant. Dependent core parameters such as power distribution and moderator temperature distribution are permitted to vary g as dictated by the changes in core neutronics and thermal- 5 hydraulics. The average fuel temperature is held constant and no changes in nodal xenon inventory are permitted. 2.2 Power Reactivity Defect, ap p Definition: app is the change in core reactivity associated g with a change in core average power level. W Calculations of appare performed in three dimensions with the Page 16 of 332 I
I I nodal model (1). Core power is varied while all other independent parameters such as rod position and RCS baron I concentration are held constant. Dependent core parameters such as power distribution, average fuel and moderator temperatures, and moderator temperature distribution are permitted to vary as dictated by the changes in core neutronics and thermal-hydraulics feedback. No changes in nodal xenon inventory are permitted. I 2.3 Doppler Reactivity Cnefficient, a0 Definition: a is the change in core reactivity associated D with a 1 F change in average fuel temperature at constant average RCS moderator temperature. aD is computed as the difference between power defect, App, and the moderator coefficient, aM' as shown below. Ap - AT p M "M "D " AT p 2.4 Boron Reactivity Coefficient, a B Definition: aB is the change in reactivity associated with a IPPM change in core average soluble boron con-centration. I Calculations of aB are perf rmed in three dimensions with the nadas model (1). The core average boron concentration is varied while the independent core parameters such as core power level and control rod position are held constant. Dependent core parameters such as power distribution and moderator temperature distribution are permitted to vary as dictated by the changes in core neutronics and thermal-hydraulics. No changes in nodal I xenon inventory are permitted. Page 17 of 332
0 I 2.5 Shutdown Margin, SDM Definition: SDM is the amount of reactivity by which the core g would be subcritical following a reactor trip, assuming the most reactive control rod is stuck out W of the core and no changes in xenon or RCS boron concentration. Case #1 - At power condition with rods at the power dependent insertion limits. l Case #2 - Hot zero power condition with all rods in except *he stuck rod. No changes in xenon or boron are assumed. Case #3 - HZP with rod position from Case #1. The dependent I core parameters such as power distribution and temperature distribution are permitted to vary as dictated by the changes in core neutronics and thermal-hydraulics feedback. All spatial effects and I a rod insertion allowances are explicity accounted for in each calculation. The SDM is computed as the change in core reactivity between case 1 and 2. This value is conservatively reduced using Case #3, model reliability factors RF9 (Reference 1), and biases. These factors are applied to the inserted rod worth, I, l the temperature defect and the Doppler defect. l'l Il II I Page 18 of 332 I l I1
I I 2.6 Scram Reactivity Curve, Ap(t) SCRAM 1 I ! Definition: ap(t) is the rod worth inserted into the core as SCRAM a function of time after rod release. The most reactive rod is assumed to remain fully withdrawn. The independent core parameters such as power level, RCS boron I concentration and xenon inventory are held constant during the insertion. The dependent parameters such as flux distribution are I permitted to vary as dictated by the changes in core neutronics. A conservatively slow scram curve is generated by making the following assumptions:
- 1. The integral of the scram curve is based on an initial rod position at or below power dependent insertion limits.
I 2. The shape of the scram curve is based on an initial rod position of full out. The rods are assumed to move uniformly together. This provides the longest possible delay to significant reactivity insertion.
- 3. The positional insertion dependence is converted into a time dependent function using empirical data relating rod position to time after rod release. The empirical data is normalized such that the total time to rod insertion is equal to or greater than the innits defined by the Technical Specifications.
- 4. The xenon distribution is that which causes the minimum shutdown margin.
- 5. Instantaneous radistribution of flux is assumed to occur during the rod insertion, Page 19 of 332 5
I 2.7 Nuclear Heat Flux Hot Channel Factor, F g Definition: F is the maximum local fuel rod linear power g
. density divided by the core average fuel rod g linear power density. g Calculations of F are g based on three dimensional power distributions obtained with the nodal model (1) coupled with local peak pin to assembly power ratios obtained from the quarter core PDQ model (1). Model reliability factors and biases are used to increase Fg to a conservative value.
2.8 Nuclear Enthalpy Rise Hot Channel Factor, F AH I Definition: F is the ratio of the integral of linear power aH along the rod on which minimum DNBR occurs to the core average integral rod power. Calculations of FAH are based on three-dimensional power distributions obtained with the nodal model (1) coupled with the g local peak pin to assembly power ratio obtained from the quarter 5 core PDQ model (1). Model reliability factors and biases are used to increase F 3g to a conservative value. 2.9 Effective Delayed Neutron Fraction, S,ff Definition: S,ff is the core effective total delayed neutron g fraction. 5 Nodal values of S, for core delay neutron group i, are determined by weighting the delayed neutron fraction from each fissile isotope by the local fission sharing as determined from CASMO II. This local result is then power weighted using the nodal 3D power distribution. The importance factor I, applied as a .97, accounts for the effects of reduced fast fissioning, increased resonance escape, and decreased fast leakage by the delayed neutrons. S,ff is the product of 6 and I, where S=ISj . Page 20 of 332 I
I I 2.10 Prompt Neutron Lifetime, t* Definition: t* is the average time from the emission of a prompt neutron in ffssion to the absorption of the neutron somewhere in the reactor. 1/v ! t* = j 2 x,I, + 08 1 IJK I S /2VEL S = source 1/v = IJK VEL = neutron velocity I I S IJK I S vI f / K-1,I, = X L XL = eigenvalue IJK l I I S l IJK I S vIf [(Q,-AH) WH + (Q -AV) K WV] 082= IJK 2 I S where: 6, = number of exposed radial faces per node QK = number of exposed axial faces per node AH = horizontal albedo AV = vertical albedo WH = horizontal neutron leakage probability WV = vertical neutron leakage probability Page 21 of 332
I 3.0 SAFETY EVALUATION METHODS This section addresses the evaluation of the cycle specific physics parameters with respect to the bounding values used in the safety analyses. Specific methods are described for each accident or transient by which the determination is made as to whether or not any re-analysis is required. For each accident or transient the following material is described:
- a. Definition of Accident - a brief description of the causes and consequences. <
- b. Accident Analysis - a brief description of the methods employed and discussion of the sensitive physics parameters. Included is a list of the acceptance criteria.
- c. NSP Safety Analysis Experience - a brief summary of the NSP w calculational experience and results of the comparisons of their
- models to the Prairie Island Final Safety Analysis Report (Ref-erence 2).
- d. Cycle Specific Physics Calculations - a description of the specific physics calculations performed each cycle for the purposes of a reload safety evaluation.
E
- e. Reload Safety Evaluation - A description of the comparisor.s of the cycle specific physics characteristics and the bounding values used in the safety analysis. Specific applications of the model reliability factors and biases are also addressed. Biases and reliability factors are to be applied in the following form: ,
E I Page 22 of 332 I .
I
~-
I
- Moderator Temperature Coefficient o g j Apply in a conservative direction as follows:
I aM * "M (MODEL) + Bg RF M B g moderator temperature coefficient bias (pcm/ F) RFM = Moderator temperature coefficient reliability factor (pcm/ F) I - Doppler Coefficient a0 Apply in a conservative direction as follows:
"D * "D(MODEL)*(1 RFD )
RFD = 0 ppler coefficient relative reliability factor Note: Doppler bias defined as zero.
- Boron Reactivity Coefficient aB I Apply in a conservative direction using the following form:
8 "B
- B (MODEL) * (1+BB )*(1 RFB )
l B B
= B r n c efficient bias 1
RFB = Boron coefficient reliability factor
- Scram Reactivity apScram(t)
Apply in a conservative direction as follows: ApScram = apScram(t) (MODEL)*(1 RF R )*(1+B R ) l l Ra,e n of m l
I I - Nuclear Heat Flux Hot Channel Factor (Fg ) Apply in a conservative direction as follows: I F g = (F (MODEL)+RFpg+8pg)*(1+T) g RFpg = nuclear heat flux hot channel factor reliability Bpg = nuclear heat flux hot channel factor bias T = Technical Specification Tilt Limit - Nuclear Enthalpy Rise Hot Channel Factor (FAH) Apply in a conservative direction as follows: FAH " ( AH(MODEL)+RFFAH+BFAH)*(1+T) RFFAH = nuclear enthalpy rise hot channel factor reliability I B FAH
= nuclear enthalpy rise hot channel factor bias T = Technical Specification Tilt Limit I - Effective Delayed Neutron Fraction: 8,ff Apply in a conservative direction as follows:
Oeff*0eff(M del)*(1 RF3 ) RF g = S,ff relative reliability factor I Page 24 of 332 I I,
I ;
- Prompt Neutron Lifetime (t )
Apply in a conservative direction as follows: t =t (MODEL)*(1 RFe *) RFt * = relative prompt neutron lifetime reliability factor (no units) i NOTE: t* bias defined as zero.
- Rod Worth (ApR)
Apply in a conservative direction as follows: (apR) * (A P R)(MODEL)*(1 RF R )*(1+BR ) I BR = rod worth relative bias RFR = r d worth relative reliability factor i The specific numerical values assigned as the bounding values for each accident for purposes of performing the Prairie Island reload safety I evaluations will be presented in the cycle specific Reload Safety Evaluation Report. If an accident or transient requires re-analysis because anyone of the i cycle specific physics parameters exceeds the current bounding value, the re-analysis will be performed utilizing the transient analysis I methodology as described herein for that specific event and which has been qualified by the presented results. I I eageasof m y I
I 3.1 Uncontrolled RCCA Withdrawal from a Subcritical Condition 3.1.1 Description of the Accident . An uncontrolled addition of reactivity due to uncontrolled withdrawal of a Rod Cluster Control Assembly (RCCA) results in a power excursion. The nuclear power response is characterized by a very fast rise terminated by the reactivity effect of the negative fuel temperature coefficient. After the initial energy release, the reactor power is reduced by this inherent g feedback and the accident is terminated by a reactor g trip. Due to the small amount of energy released to the ' core coolant, pressure and temperature excusions are minimal during this accident. 3.1.2 Accident Analysis The uncontrolled RCCA withdrawal from a suberitical condition is analyzed using a dynamic simulation , incorporating point neutron kinetics, including delayed neutrons and decay heat; fuel, clad, and gap heat , conduction; and channel coolant thermal-hydraulics. The j reactivity effects due to moderator and fuel temperature : effects, as well as that due to control rod insertion g after trip, are included. 5, The core is assumed initially to be at hot zero power, HZP. Power is supplied to the RCCA drive mechanisms ! such that no more than two banks may be withdrawn f simultaneously. The maximum reactivity insertion due to the rods are therefore conservatively assumed as that g, due to two banks of maximum worth moving simultaneously W) at maximum speed through the region of highest differential worth. Page 26 of 332 I
A, I The magnitude of the power peak reached during the transient is strongly dependent upon the Doppler reactivity coefficient for a given rate of reactivity
.. insertion. A value conservatively small in absolute magnitude, which generally occurs at Beginning of Cycle (80C), is assumed for the accident analysis. The magnitude of the power spike is relatively insensitive to the value of moderator temperature reactivity coefficient chosen. The least negative value, occurring I at BOC, maximizes the calculated consequences of the accident.
I In calculating reactivity due to control rod insertion by reactor trip, the most adverse combination of instrument and setpoint errors and time delays is assumed. The power range - low range trip setpoint is I assumed to be 10*.' (of full power) above its nominal value. The most reactive rod is~ assumed to stick in the fully withdrawn position when the trip signal is actuated. I As long as the reactivity insertion remains small ccmpared to 6,ff, the total delayed neutron yield, the I :hortest reactor period during the transient will remain large compared to t*, the mean neutron lifetime. In this case, the transient core power response is relatively insensitive to the value of t* and is determined predominately by the yields and decay constants of the delayed neutron precursors. The postulated initial core pressure and temperature !re I conservatively taken as the minimum and maximum, respectively, that are consistent with the assumed rod and power configurations. I I Page 27 of 332 I
I, The results of the analysis are compared to the following acceptance criteria:
- a. The maximum power density in the fuel must be less than that at which center line melting or other modes of fuel failure occur. ,
- b. The minimum departure from nucleate boiling ration Il '
l (DNBR) calculated using the W-3 correlation must be greater than 1.30. El 3.1.3 NSP Safety Analysis Experience NSP has analyzed the uncontrolled RCCA withdrawal from a subcritical condition transient using input assumptions consistent with the Prairie Island FSAR (Reference 2). The rrodels described in Appendix A were used to analyze g the case of a rapid (8.2 x 10 -4 Ak/sec) RCC assembly 5 withdrawal from subcritical. The results of these cal-culations are compared to Figures 14.1-3, 14.1-4, and 14.1-5 of Reference 2 in Figures 3.1-1-3.1-5. The NSP model predicts higher fuel, clad, and coolant temperatures than those of Reference 2, however, the NSP fuel model is consistent with the Doppler and moderator g' 5 reactivity coefficients used so that the nuclear power and core average heat flux compares well with the Reference 2 results. A sensitivity study showing the effect of initial power Il ' level on peak heat flux was performed and the results E are compared to Figure 14.1-2 of Reference 2 in Figure 3.1-6. 5l ll I ;
E I This study was only run at one reactivity insertion rate, i.e. 8.2E-4 Ak/sec, the same insertion rate that I was used to generate Figures 3.1-1 to 3.1-5, however, the results compare well to the FSAR results. 3.1.4 Cycle Specific Physics Calculations I These calculations are performed at the most limiting core conditions found during the cycle, e.g. the point I in time, and the power level, that gives the most conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine the limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. for each parameter. I
- a. Moderator Temoerature Coefficient, a M
Calculations of a Mare performed in accordance with the general procedures described in Section 2.0.
- b. Doppler Temperature Coefficient, a0 I Calculations of a Dare performed in accordance with the general procedures described in Section 2.0.
I c. Scram Reactivity Curve, ApSCRAM I Calculations of the scram reactivity curve are performed in accordance with the general procedures I described in Section 2.0. I 1 I eae. 2e ef 352 I
I
~
- d. Effective Delayed Neutron Fraction, 6,ff I
Calculations of 6,ff are performed in accordance with the general procedures described in Section 2.0.
~
Maximum Reactivity Insertion Rate, ap/At I e. The assumption is made that two banks of highest worth will be withdrawn simultaneously at maximum speed. This value requires two components. First, the maximum withdrawal speed is required in inches per second and secondly, the maximum differential reactivity insertion per inch for two maximum worth rod banks moving in 100% overlap. This result is obtained by first calculating the g two banks which have the maximum worths. These two 3 banks are then moved simultaneously at HZP. A rod worth reliability factor and bias is applied to the integral worth by position. I, I! I I I I Page 30 of m 8 i I
I I 3.1.5 Reload Safety Evaluations Each of the physics parameters calculated above are adjusted to include the model reliability factors, RF$ and biases, B 9 . These adjusted values are the cycle specific parameters which are then compared to the bounding values assumed in the safety analysis. The cycle specific parameters are acceptable if the following inequalities are met: CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. a ' "M (least negative M + RFM+BM bounding value)
- b. * (1 - RFD ) 5 "D (least negative I
a 0 bounding value)
- c. ApSCRAM(t)*(1 - RFR ) 2 ApSCRAM(t) (bounding)
* (1 + BR)
I d. S,7f * (1 - RFg ) 2 S,ff(minimum) a i e. Ap/At * (1 + RFggg3) s Ap/At (bounding)
* (1 + BR)
The integral of the bounding value of the scram curve, ApSCRAM(t), is taken as that rod worth required to produce the shutdown margin assumed in the safety i analysis for the most limiting cycle specific core I conditions. l I t I I Page 31 of 332 5
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s.3 , .;, . .. . . . .;. . .. .., ... .. .g. .. .g. .. . .g... ..
- s. . . .
. . . .. as .. .. . . ... , .. .
- s. .
- s. .
l 0.01- , ", , . , , . i i 5 10 15 20 25 30 35 40 45 50 l 0 Seconds l. , Page 32 of 332 ! ) E E W W W W W W W Fge 3g . g g g g g g g .
um een am num use amm amm ese an em nas em aus man me um as me as i FSAR Rod Withdrawal DNP133 From Subcritical F.S.A.R. . . J Core Average Flux vs Time 3 0.8 .
- e., :.
j e . . .
- s. : .
. s. . .
0.7 - -
+ : - -
+---- - :- -
;- --- =-- -
,. '.s.--
s : :
. s .
+
l 0.6 - *=- --
's- - - - -
i - -
?
a
- s. .-
i . . .
! .a s. : -
as 0.5 - - - - -- - - - - - - - --
,c,
*:6 .
g - . . E 8 : :s
-.ss i
. i.
O . . s : Z "*- -- .
. - .. s .
C 0.4 - -
- . v-s i -
o 8 i
's. *:.
. . i.
+- .
O . g as a
% : ,%..}
UL g,3 - .,' . .:. .
.....} .
.....+ . ..
*e 1 . . . .
--+
0.2 - ,
- i. - -+- i-- i. -
-?~-.*.*--+.-----
i i i i i . . . . i , -
+-
j 0.1 - n' - + -
=- - - -
+ -
+:
l
~"
s' . ! 00- i i i i e i i - i e 9 10 11 12 13 14 15 16 17 18 19 20 Seconds i F10ure 3.1-2 Page 33 of 332 . I
4 I i l i
- FSAR Rod Withdrawal oneisa l From Subcritical -
, F.S.A.R. . .
. i j
Average Clad Temperature vs Time 620 . 4 i . . . .
-l I -:--"" : :
610-i : i : i 600- -- - -
- i " * - "
i *" --* - 1 . . 1 : O ggg_ . : .~ \ .,-s...
, s : . .
- i - . -
J: . I c
- j O -
- 6. a j r a s :- : i : i i :
; O . .
LA. 580- - --- - e- 's"g .
; -~
I cm . . .: . (-) : : : . :. :. e
- s,: . . .
O 6. 8 I *
- ~
CD *- n 570- ,
-'c %
i. 5"
- O *
\
e . . . 1 e : * . : : . : sg - E s .
- r . s- . . : : :
i 580- - - s:- W' - - - : n. s ' ,. n:. . si : i *.. -
~~~'" ~~~~ws;.;,h. a. ;
l i 550-i a .5 l ----: _m . 540- 4 a i i
- 8
' 8 0 5 10 15 20 25 30 35 40 45 50 i
Seconds j Pa e 35 of 332 l M M M M M M M M MFl0g 1M g g g g g g g
as as as
~
amm een am am uma ese ama ma uma em um um uma asa me am FSAR Rod Withdrawal DNP133 .l From Subcritical l F_S_A_R_ . .
! Average Fuel Temperature vs Time i
1200 - - -
. i i . . : .
1
- l - -
; : i l .
i -
; 1100- -
1 : : 1 : : : : : : ) 2 3g00 .. .... .. .. . .. , . .
. . ~ , . . . . ..
4
-r . :. . :. :. :. .
(3 . 4 - - c -
, ()
L ggg. .
.i...
.,p*. . .. .
+
+ . .
.. .g.
1 i r - i a . e . i ! 1 :. .s g
- :8
! m :. s : ( .-,s g
, ( 800- - - - ; ---
p- -- v g .-- + .
+ -- -- -- - -:
- - - + -- ,- -- - + - -
i a's . e e o i 6
- ss l
c .
=
. s s
i t s . 700- - R.- 8
- -- ^'
s g
- - - - 1. - - - -- :- - - - +-- .
- a: s :. :. . .
g, g . .
. ,: .g: . : :
si :- : - i l a: s: 1',, -
- ' ~ 4 ;; -- - - -
600- : - ei ' r: . i l:
~ ~ ; . ._ . - - -:- -- . -. _- _ . . .; -
l . : : : : : : 2 l 500- , , , , , . ! O 5 10 15 20 25 30 35 40 45 50 i I Seconds t Figure 3.1-5
I FSAR Rod Withdrawal 3 oyuoo, From Subcritical
- F8AR Peak Heat Flux vs initial Power Level l
0.8 . . . . l i '. : . : :. :. : :. - i N:-% . f - i 0.7 - -- -- -
% N.
-i --
-i l-~~~-i-~---
i . N : i i . i !
. N. g :
N
. - - - - -i 0.6 - - - - . --
b" - -- " ; - - - " "..--" ", - - - - - --"
-- -' %: N-
~
~
. M .- :
i i i ! - as 0.5 - -
- N !
C : : - : : : . - a . :. :. :. . : . : r- . . . . . 1 C : . .
. O -
i z 0.4 -
~ ~ - - - - " -
C
. ":"*"~".-"""':""--
4
.C).
o : :. : .
- K"" W --"y a3 : : -
[ 0.3 - - - 3 i- " "-" " i i . . . . . . 0.2 - -
- - -.~~---.-
i . . : . : .
- i : i i -
. : i
, i ; : : :
O.1 - - Q = Q(Nominal)x10**N
-i.--
- i
; 0.0 - . . . . . . . . . . . .
1
-13 -12 -11 -10 -9 -8 -7 -6 -5 -4 -3 -2 -1 \
N l FI ure 3.1-6 Page 37 of 332 W W W W W W W M W M M M W W M M M M
I I 3.2 Uncontrolled RCCA Withdrawal at Power I 3.2.1 Definition of Accident An uncontrolled RCCA withdrawal at power results in a gradual increase in core power followed by an increase in core heat flux. The resulting mismatch between core power and steam generator heat load results in an increase in reactor coolant temperature and pressure. I The fuel in the reactor core could eventually encounter departure from nucleate boiling if the power excusion were not checked by the reactor protection system. Depending on the initial power level and rate of reactivity insertion, the following trips serve to prevent fuel damage or overpressurization of the coolant system: high nuclear power, over temperature and over power AT, high pressurizer level, and high pressurizer pressure. For the more rapid rates of reactivity I insertion, the maximum power reached during the transient will exceed the power at the time the trip setpoint is reached by an amount proportional to the . insertion rate and the time delay associated with trip I circuitry and rod motion. 3.2.2 Accident Analysis The uncontrolled RCCA withdrawal at a power condition is analyzed using a dynamic simulation incorporating point neutron kinetics, reactivity effects of moderator, fuel i I and rods, and decay heat. A simulation of the reactor vessel, steam generator tube and shell sides, l pressurizer, and connecting piping is required to evaluate the coolant pressure and core inlet temperature response and their effect on core thermal margins. The reactor trip system, main steam and feedwater systems, and pressurizer control systems are also included in the 1 model. This model calculates the response of the l 5 Page 38 of 332 I
1 I I tve age core channel thermal-hydraulic conditions and I 5: eat generation and is coupled to a detailed model of tne hot channel. This latter model calculates the departure from nucleate boiling ratio (DNBR) as a g function of time during the accident. W In order to maximize the peak power during the transient, the fuel and moderator temperature coefficients used in the analysis are the least negative likely to be encountered. The least negative Doppler and moderator coefficients are normally encountered at BOC. The reactivity reduction due to reactor trip is calculated by considering the most adverse combination of instrument and setpoint errors and time delays. The rate of reactivity insertion corresponding to the trip of the RCC assemblies is calculated assuming that the most reactive assembly is stuck in the fully withdrawn position. Since the reactivity insertion rate determines which I l protective system function will initiate termination of the accident, a range of insertion rates must be l considered. Relatively rapid insertion rates result in ) reactor trip due to high nuclear power. The maximum I rate is bounded by that calculated assuming that the two highest worth banks, both in their region of highest incremental worth, are withdrawn at their maximum speed. Relatively Slow rates of reactivity insertion result in a slower transient which is terminated by an I I Page 39 of 332 I I
I overtemperature A T trip signal, or in some cases, a high pressurizer pressure signal. The minimum rate which need be considered in the analysis is determined by reducing the reactivity insertion rates until the analysis shows no further change in DNBR. The accepted criteria for this accident are that the maximum pressures in the reactor coolant and main steam systems do no exceed 110% of design values and that I cladding integrity be maintained limiting the minimum DNB ratio greater than 1.30. 3.2.3 NSp Safety Analysis Experience I NSP has analyzed the uncontrolled RCCA withdrawal from a full power condition transient using input assumptions I consistent with the Prairie Island FSAR (Reference 2). I The models described in Appendix A were used to analyze the following two control rod withdrawal transients from full power:
-4
- Fast rate (8.2 x 10 AK/sec)
-5
- Slow rate (3.0 x 10 Ak/sec).
The transient response of the NSSS for the fast rate case is compared to Figures 14.1-6 and 14.1-7 of Reference 2 in Figures 3.2 3.2-3. The reactor trip is generated on high neutron power for this case. The I T AVE and pressure responses are slightly more severe using the NSP models, however, the NSP models show the same trends as the FSAR results. I ! l E Page 40 of 332 ll
I The corresponding NSSS results for the slow rate case are compared to Figures 14.1-8 and 14.1-9 of Reference 2 in Figures 3.2 3.2-7. The NSP results predict a slower power ramp and corresponding by slower pressure and T increases. A reactor trip is generated on AVE overtemperature AT for this case. The NSP model uses a dynamic simulation of the setpoint generator and predicts a trip at a slightly lower power level than Reference 2. Figures 3.2-4 and 3.2-8 show the transient hot channel DNBR comparisons to Figures 14.1-7 and 14.1-9 of Refer-ence 2 for the fas* and slow rate cases respectively. The NSP hot channel DNBR analyses were computed using I both a single closed channel model, a multichannel 1/8 assembly model, and an 1/8 core lumped subchannel model. It is believed that the single channel model and the 1/8 E
, assembly model will provide a better comparison to the ,
5 FSAR, however, the 1/8 core model is more accurate and will be used for licensing analyses. A detailed description of the NSP Thermal Margin methodology is given in Appendix C. The NSP single channel model predicts a MDNBR of 1.726 g for the fast withdrawal case and 1.641 for the slow 3, withdrawal case as compared to 1,63 and 1.36 from Reference 2 respectively. Note that the initial MDNBR of the NSP nodel is slightly higher, ( ~.038 ), tending to bias the results throughout the transient. 3.2.4 Cycle Specific Physics Calculations These calculations are performed at the most limitino core condit. ions found during the cycle, e.g., the point Page 41 of 332 I
~~
E in time and the power level that gives the most conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine these limiting conditions accounting for I the effects of control rods, xenon, power level, temperature, etc. for each parameter. .
- a. Moderator Temperature Coefficient, og I Calculations of a Mare performed in accordance with the general procedures described in Section 2.0.
- b. Doppler Temperature Coefficient, aD Calculations of a0 are performed in accordance with the general procedure described in Section 2.0.
- c. Scram Reactivity Curve, ApSCRAM(t)
Calculations of the scram reactivity curve are performed in accordance with the general procedures described in Section 2.0. I d. Nuclear Entahalpy Rise Hot Channel Factor, F AH The maximum core F AH is assumed to remain within the current limits as defined in the Techniczl Specifications for allowable combinations of axial offset and power level. For Prairie Island, the continuous surveillance of the power distribution is accomplished with the excore detectors using the Exxon PDC-Ila(3) scheme. I Page 42 of 332 I
I, I I
- e. Effective Delayed Neutron Fraction, S,ff Il l Calculations of S,ff are performed in accordance I l with the general procedures described in Section 2.0.
- f. Maximum Reactivity Insertion Rate,4p/At Calculations similar to those described in Section 3.1.4 (c) are performed at the full power, and constant equilibrium xenon conditions.
3.2.5 Reload Safety Evaluations I Each of the physics parameters calculated above are I adjusted to include the model reliability factors, RF9 , and biases, B9 . These adjusted values are the cycle specific parameters which are then compared to the g bounding values assumed in the safety analysis. The 3 cycle specific parameters are acceptable if the following inequalities are met: I I I
~
I I I Page 43 of 332
E CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS I a. a M + RFM+Bg s ag (least negative bounding value)
- b. a D*(1-RFD) 5 "D (least negative bounding value)
- c. ApSCRAM(t)*(1-RFR005) 2 ApSCRAM(t) (bounding)
* (1 + BR )
- d. (F3g+RFFAH+BFAH)(1+T) s Technical Specifications
- e. 6,f7 * (1 - RF3) 2 6,7f(minimum)
- f. Ap/At * (1 + RFR005) s Ap/At (bounding)
* (1 + BR)
The integral of the bounding value of the scram curve, ApSCRAM(t), is taken as that rod worth required to produce the shutdown margin assumed in +' a safety analysis for the most limiting cycle * .cific core conditions. I I I ; l I I l 1 Page 44 of 332 I
I, i i .. I l FSAR Rod Withdrawal
- DNP134 Fast Rate from HFP
.F.S.A.R. . .
l Absoluto Power vs Time 130- . . . . s.s
-- :- ws -. -----: -
120- -- ' -
-1 -
t .: . s s i g . . .
, g . .
2--
.' - - -i- -
,- ' t --- ----- --- .
'l - - - -- -- -- - -
110 - - s-- .- . . . .
, g ;
#
- t , ,
1' # g .. i .
' - - - +--- - . - -
+ ----. - -- -+ - .
100- - -
.l. -
t g i
. . . t . .. . .
i . . . e . 1
. t . .
t . .
. g. .
i @ . .. . .......tt.... ... .... . .. .. . . . . l 8 . ..
. g..
i _C . 8 . . . g . E . . - l O
~ + -.. - . --
.. q. - - --- . - - - . - . . . + . - . - . - . . . . . . . - - - + . ..
-:.-..+--
70- . - --
. . g 4
j
& . . g . . . .
-- - + -------;----
--+- -;- -
; g - -- -- - -- ; - - - --
-+
----.*t-s -- - - - l.--- - - - - ;.-- -- - . . .
o 60- . . . . i U . t . t .
. . . t . . .
.....g.. ...;. .... .... .. : . :.. .... ..... . ;
O,, 50 .. t t .
. . t . . .
t -- 40- ---- -- - - . ---- -
. - -- ----- -.-- -=- ' . . ga g . . . .
e.s. j .. .
.-------~:-------: - --
30- '. - -- ----:--- -- t- -
- - - - "t-.-
...w
......4..~"'~w.t..*****"t."-
20- - -- - + - - - --~~+-- -
--t."--- -
"i." - -- -
t.-
---t.-
S..==.....f.......... 10 .. .
, s s l
D- s s s
. s s 8 9 10 2 3 4 5 8 7 0 1 Seconds 4
i . Page 45 of 332 M M M M M M M M M U*E W m m e m _ . . _ . _ _ - mm _ _ . ,
i, as aan mm um a m, en sus sua em um me ses mas sus as em
! e. , FSAR Rod Withdrawal DNP134 Fast Rate from HFP F.. S.A.R. . .
i Pressurizer Pressure va Time 2270 .
, gggg- .. . . .
<1 . . .
. . , s .
2250- - - - - + - - -:- .
- t.
i :. . s : .
. , g .
j ;--- - - - , - - - - - - - -- - - 2240- ?.--- i --
- i.- ,,- ,,- -
-[.,
i 2235- - I-s- s
- - !- I
--i , ,o .
s . . s..: 1 . : : . I 2220- - ---
.s
. g :.
. .- . . s .
---"s------ : - - - + v- "l- --"
"l- -
2210- -- N
- T. - --
. . . . s. .
+.
s . L 2200-
--;------+- - ---
*-t.---- -* - -
. -- .. sg
.s
. *g .
4 - . . g . . . , 2190- - -
- - -i- .
- -i- -- - i- .
"*w %
- -i. -
. . . . . g. .
. . 4. s
! 2180- -- -- i 2170- -- -
. -- --- .- - - -t-.
- - - - :. - - -- - - .t- - :- - - --
.t- - - -~ %." , - - - "t- -
.,..g...,.
\. .. . . j 2160- +- -
-l" - -- -t--- -- - l" - - -
-t" -
+: ' r".*-
- - - - + - -
-t-l :. -
i g i
. 4
'r" 2160- - -
----i.---
g g 1 . . . . 2140- . . . . - i . . . i 0 1 2 3 4 5 6 7 8 9 10 Seconds I l Page 46 of 332 F10ure 3.2-2
,i. )
! FSAR Rod Withdrawal DNP134
- Fast Rate from HFP t' F. S. .A.R. . .
f Vessel Average Temperature vs Time
- 5 74 . . ..
i 1 573- -- 572-( 3
....,.e------- + . . . . . -
1
.....--h--- .
57f. l..
. . . g. .
4
- . . - s.:. .
m : : :. .: .. r : : . . : : . C 570- -
' c. :. .
i m .
. , y. :
w
.c : - -
1 qs : .
,q s,, 5sg. .. .
... w. .
{ . . . . . m . . :. m . . , i . .
- i e
u En 568- - --- e .
.g
. . . . . g
; "'s i - - --
---i 567-
? :-
1, . . . 566- - -
. . . g .
l 58C. UW
. . . e. . ! 564- , .
2 3 4 5 6 7 8 9 10 0 1
' Seconds W me __ - ___
e _ e e _ e __ a e B' $8 m -- -_ a - y^"47 Q32 g
W W W M M M M M M W M M M
.. A CBR099 Sin 013 X CDR0091/8 Assembly FSAR Rod Withdrawal o COR0201/8 Coro Fant nato from MFI' FSAR W-3 Of1D Ratio vs Timo 3 ._ -.. _._- -... _.. . . _ . . _ . _ _
I 2,75 . :. . . . . :. . . .
^'
/.
2.50- :- - - - - - - =- - - - -
--4---- - - -
y' - - O
.;/
! b
; cJ
[I' /
/.
[]
^
I r0 / j z 2.2 5 - -
- -- - :- - - - -+----- --- -
---+- --
/ -
D- - ' o a.
/ A o
/ X 3 *
/ O I
2- - - - - - - - - - - - -- - - - - - - - - - - - - - - - - -- - - - - - -:--- :- - 9 u o o o "
~ ._ _ra ~ ~q< o go .
e is h Q o O ,G
~
1.7G - - -- - - 4- - - - - - -@o -b--- f . .
- : O --[A -
1.50- t i i i - i . . O 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 5.5 i Seconds Figure 3.2_4 Page 48 of 332 1
I,
. u ..
FSAR Rod Withdrawal DNP135 Slow Rate from HFP FS.A.R. . . Absolute Power vs Time 120- .
....a...- s : .
- s..
- : ....w i :
--j .... - --
;- ~' . ;
110 - - --- s.
- i:
- . . g. .
# " '_ i : :.
- i. i
- si :
100 . :. s;.. s. i i
- . :. e a . . . . e g- .
; 90- - - -- - - - - --
s
. s, .
l : :. - i :. : p 80- -- -
- : : : : :s .
ca : . . : : : :s - -
- s .
.C 70 .. . .
, E . o .
- :s
- 60- -- - -
- s c .
. : :. : :. :. :s
.g m .
O 50 . ..
.. . .. . . . . . . .. .. . . t... . ...
m . , ct :.
- : s, . .
40- - - - -- - -
. -- -- . . 3 --
1 . . . na QU.
..A...
g s . 6 1 20
...9 % ... ..
a, . . . . . j
' 10- - -
. 0- s s s e e s s a s s s 1 O 5 10 15 20 25 30 35 40 45 50 55 60 Seconds re -5 pane 40 or 332 W W W W M M M M M
FSAR Rcid Withdrawal DNP135 Slow Rate from HFP F. .S.A.R. . . Pressurizer Pressure vs Time 2375- . 1 . 4 .
-i.
- i. r.
.l. .* . ,2.- m g350 . .. .+ .
.i. . .
. i. .
~M..- .
e . . .
*,..,.,.g
- s-
,*.*,p--
2325- - - -
+
+.j,s****; + -*- - -
2300- - - + .
. i
- i. .
i o . . 2275r - -- - -+ .
- : - -+ . ,-*--, t--
,# r 2250- -- --
- ----,**4, : *
- - .~ -
s-
-e
-.._;;,..; ,, ..,,;, *.. +. - . s.
- gggs.
+
. .. . .n-....:. .
+. .-i,-. .. .+ .
. . g .
. . . g to 2200- -- -
~~:- .
sg : i j n, . i . . . . . . .
; -:- s+ ;
2175- - - . - --+------: - - - - .+ - .- + - + -.
- .+ - - .
I' . .
- : .: i. .: -
.- .- . s. .
2150- - - -
. . .ig .
i
- - + - - ;- - +-+-
2125-
- - -- - - + -. - -- - - -l- -
.-- -----l.~~ . .
. . . . . . g .
s
. . . t .
- + - - : -+ -
+- +
2100-
. . . . . g .
. sg . .
s. 2075 . .. ... .. . .. . . . . .;. . . ... .
.4...
4
. . g
. . . . . g
+
.: - - - - +- -+
, 2050- -
+- .
- --+ .
- - - - +. - -v'
. . s 2025- . .
0 5 10 15 20 25 30 35 40 45 50 55 80 Seconds Floure 3.2-6 Page 50 of 332 I
r i .. 1 FSAR Rod Withdrawal DNP135 i Slow Rate from HFP F. _S_A_R. . . 1, i Vessel Average Temperature vs Time 582 . . i . . . . .
. . . .e,=
e s .
*:- - s- . - -
j 580- - - -
- t .
. a .
1
. +# l .
t . ,. s .
. g .
j . .
- .9; 1,
$7g. . . . .-
,r . : . . ..
g. J . *
, , s .
,,,g,a $.6 y . . , .
+t :
- -- - - - - + - " --; - "+ - - - - -
- l. *"a * * "
l.. - . t . 516.- . . .
,,,P
. . . t .
. g .
t
+4 . .
. . . g i
- +
" - - + - +-- .--
"s. m o "'. -
-.'-",--",*P.p, i 5 74 - -
- 6 -
J: . . : , . . . s t . l C .
. . t O .
-g y
512- + -
"i" T--
+- --
T. - T-" - T. t ? - l .c -
-- - ----...s.... .
T.. s g . I C3 . .. . . . . .
. . . s i 1 .
. . . . . . . t "1"
~*
j (f) .
- + - - - :
-"-1" "- - -
.l -
~~
t-. i W 510- --
. 4 p
g Q
.g.~
.g.. .~g
..- .g.
n -: .g.
. ... . . .g.t~ .
Q ggg. .. .. . .. . : .
~.
t
. .. . . . . t g
. g "t-
,i 566- " t-t-
- "t i
.t-t t
. . . g i
g I. . . .
- -~:- : - 'g-
-- ~ : "----:. - " - - " -:. - - - : -:-
564- - - i, . t.
- . . - -~: : - -
i 562- - - 560- . . .
. 60
. 55 10 15 20 25 30 35 40 45 50 0 5 Seconds FI uro 3.2-7 Page 51 of 332 M M M M M M m
M M M M M M M M M M M M M M M M A CBR100 Ginolo X CDR1001/8 Assembly FSAR Rod W.thdrawal i [] CDR0301/0 Coro Slow Rate from HFP
.F,S. An -- _
W-3 DND natio vs Timo ' l 2.3-- -- - I
. r -
1 2.2- - - - --- - - i-- -
]
- . ( .
2,j . .. . . . . . . . . . . . .................J
. . .. . .. ..... j . . .. .........g.... . . j ...... .. . ..
. :. ( ,
2- - --- - - -- - - - : -- - - -- - - - - -- O - - i; 1.9 { y- -- - --
- ----:----*--;-~~~'--:-------*
c3 .
- I :
C p.h - -
- i. : :
kQ 1.8 - - ;; -
- 4)
-{-
I - - -- l: O 6 .: m - 6 J 1.7 - - - -
*w--
f ---------~~~--h---- - - -}i- -
. A. O -.A
% e j,g .
......................................g............................
- g .
- : : s .
t : s : -
- 1. 5 - - -- -
- - -~ ~ -
--y--~~: .
- 3. 4 . . .. .... .. .. .. ..
3 . . ... ..........!...........................
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0 10 20 30 40 50 60 Seconds Figure 3. 2-8 Page 52 of 332
3.3 Control Rod Misalignment 3.3.1 Definition of Accident In the analysis of this accident, one or more rod cluster control assemblies (RCCA) is assumed to be statically misplaced from the normal or allowed position. This situation might occur if a rod were left behind when inserting or withdrawing banks, or if a single rod were to be withdrawn. Full power operation under these conditions could lead to a reduction in DNBR and is subject to limitations specified in the plant Technical Specifications. 3.3.2 Accident Analysis I In the analysis of misaligned control rods, F AH will be determined for the most limiting configuration. In g general, the worst case is that with Bank D fully 3 inserted, except for a single withdrawn assembly, since Bank D is the only bank which may be inserted at full power. In practice, multiple independent alarms would alert the operator well before the postulated conditions are approached. The limiting value of F 3g is input to a steady state thermai-hydraulic subchannel calculation to determine the departure from nucleate boiling ratio (DNBR). This calculation assumes the most adverse combination of steady state errors applied to core neutron flux level, coolant pressure, and coolant temperature at the core inlet. The acceptance criteria for this accident are that the I DNBR calculated using the W-3 correlation is not less than 1.3 and that fuel temperature and cladding strain Page 53 of 332
l limits consistent with the acceptance criteria of Standard Review Plan 4.2 are not exceeded. 3.3.3 NSP Safety Analysis Experience I NSP has analyzed the control rod misalignment accident using input consistent with the Prairie Island Final Safety Analysis Report (Reference 2). Using the methods described in Appendix A, the control rod misalignment I incident was analyzed using a hot channel factor (FAH) of 1.92. The DNBR obtained was 1.470 using a multichannel,1/8 assembly hot channel model which is in agreement with the Reference 2 result of greater than 1.38. . 3.3.4 Cycle Specific Physics Calculation These calculations are performed at the most limiting core conditions found during the cycle, e.g.', the point in time and the power level that gives the most conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. for each parameter. I The nuclear enthalphy rise hot factor (FAH) is calculated for this accident consistent with the procedure described in Section 2.0. The maximum F I" AH a control rod misalignment at full power is calculated with Bank 0 at the full power insertion limit (FPIL) and one rod cluster of Bank D fully withdrawn. I I I Page s4 of 332 'I
3.3.5 Reload Safety Evaluations The F calculated above is conservatively adjusted to , AH account for the model reliability factor, RFFAH, and bias, B Additionally, a further adjustment is made FAH. ' to account for the maximum initial quadrant tilt condition (T) allowed by the Technical Specifications. The resultant FAH is then compared to the value used in the safety analysis as follows: CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS Il' s FAH(Rod Misalignment) (FAH+RFFAH+BFAH)(1+T) E I I . I I' I I I l' I: Page 55 of 332 I
I 3.4 Control Rod Drop 3.4.1 Definition of Accident In the analysis of this accident, a full-length RCCA is assumed to be released by the gripper coils and to fall I into a fully inserted position in the core. A dropped rod cluster control assembly (RCCA) typically results in a reactor trip signal from the power range E negative neutron flux rate circuitry. The core power distribution is not adversely affected during the short interval prior to reactor trip. I The drop of a single RCCA may not result in a reactor trip. The calculated consequences for this event are dependent upon whether the reactor is being operated in an automatic or manual mode. For operation in the manual mode, the plant is brought back to full power with an assembly fully inserted and a reduction in core thermal margins may result because of a possible I increased hot channel peaking factor. If a rod drop event occurs when the reactor is in the automatic mode, the reactor control system responds to both the reactor power drop, as seen by the excore detectors, (mismatch between turbine power and reactor power) and the decrease in the core average temperature and attempts to I restore both quantities to their original values. This restoration of reactor power by the reactor control system may result in some power overshoot. This power overshoot combined with the possible increased hot channel peaking factor (due to the inserted RCCA) will cause a reduction in the core thermal margin. I I eaee se ef 232 I
I 3.4.2 Accident Analysis The analysis for the control rod drop accident is divided into two parts; (1) determination of whi,ch dropped rods will trip the negative flux rate scram system and thus require no further analysis and (2) the determination of the consequences of the transient for rods which do not cause a trip. For the part (1) analysis of which rods will trip the negative flux rate scram system, a dynamic simulation is performed using the DYN0DE-P code (Appendix B) using the following input parameters and assumptions. Dropped rod specific physics parameters determined I , using static methods described in section 2.0 (listed in section 3.4.5.1). The relative tilt as seen by the excore detectors is evaluated by correlating the core edge power densities to the relative excore detector readings. A nominal negative flux rate trip setpoint of 5 : percent reactor total power (RTP) with a time - l constant of 2 seconds; with uncertainty, a setpoint l of 6.9 percent RTP with a time constant of 2 seconds l is used. Two out of four (with worst failure) excore detector rate trip system logic. . Only single RCCA drops were examined. Multiple drops could result from a single failure, however, it was i found that only the lowest worth single RCCA drops did not cause a flux rate trip and therefore multiple drops l were not examined. If this situation were to change in future cycles, multiple drops would be examined. ; 3 age 57 of 332 , II: 1
I I For the part (2) analysis of the transient consequences, , l a dynamic simulation is performed, for those rods which ! do not cause a trip, using the DYN00E-P code (Appendix B) f including the rod control system. The following input I parameters and assumptions are used. l Dropped rod specific physics parameters determined I using static methods described in section 2.0 (listed in section 3.4.5.2).
- The relative tilt as seen by the excore detectors is evaluated by correlating the core edge power densities to the excore detector readings. The I detector tilt is conservatively assumed constant throughout the transient.
Ext re detectors averaged response (with worst failure) as input to the rod controller. I - Technical Specification control bank insertion. I - Rod controller withdrawal deadband ignored. The calculation is performed assuming full power with the most adverse combination of steady state errors applied to core neutron flux level, coolant pressure, and coolant temperature at the core inlet. Least negative moderator and Doppler coefficients are used to maximize the transient power overshoot, j since the reactor temperature increases over its l initial value. j i Page 58 of 332 I
I Sensitivity studies were performed on rod worth, rod drop time, and initial rod speed error bias. It was found that the largest rod worth, the slowest rod drop time, and the largest initial rod speed error bias caused the worst transient response. For each rod, three xenon conditions are examined, equilibrium, highly positive and highly negative axial offset. If these three cases do not produce similiar results, i.e. reactor trip or no trip, a sensitivity study is performed to find the highest worth for that particular rod which will not cause a trip. Bounding values for excore tilt (Part I analysis) and for excore tilt and peaking factors (Part 2 analysis) are used. Beginning and end of cycle conditions were examined. End of cycle conditions result in the maximum rod worth not to cause a reactor negative flux rate trip (part 1 analysis) and thus result in the highest power overshoot in the part 2 analysis. Beginning of cycle conditions result in lower transient MDNBR due to the higher peaking factors. For the case with the rod controller in manual, the core is assumed to return to hot full power conditions under manual control. 1 The DNB response during a dropped rod transient is ; evaluated using the methodology described in Appendix C. 1 This methodology is modified slightly since the dropped rod transient causes core power redistribution and I yl bence loss of core symmetry. A full core COBRA IIIC/MIT model is used, rather than a 1/8 core model as in Appendix C, to account for the non-symmetric core radial power distribution. Dropped rod specific peaking factors and power distributions are obtained using the Page 59 of 332 I
I l physics methods described in section 2.0, and are conservatively assumed constant throughout the transient. The acceptance criteria for the accident are that the DNBR calculated using the W-3 correlation is not less I than 1.3 and that fuel temperature and cladding strain
. limits consistent with acceptance criteria of Standard Review Plan 4.2 are not exceeded.
3.4.3 NSP Safety Analysis Experience NSP has analyzed the control rod drop accident using input I consistent with the Prairie Island Final Safety Analysis Report (Reference 2), i .e. manual rod control . A hot channel factor (FAH) f 1.62 is used with a multichannel 1/8 assembly hot channel COBRA model (Appendix C). The DNBR obtained was 1.934 which is in good agreement with the Reference 2 result of " greater than 1.9." Prairie Island Unit 2 Cycle 8 and Unit 1 Cycle 9 have been analyzed using the methodology described in section , 3.4.2. A summary of the part (1) results is shown on ! tables 3.4-1 and 3.4-2 for PI 2 Cycle 8 and PI 1 Cycle 9 1 I respectively. The worst case results, with respect to MDNBR, for part (2) of the analyses, are shown in Figures 3.4-1 through 3.4-5 and 3.4-6 through 3.4-10 I for PI 2 Cycle 8 and PI 1 Cycle 9 respectively. I The transient response is similiar for both cycles. The initial rod drop causes a drop in reactor power and is followed by a corresponding drop in vessel average temperature and pressure. The automatic rod controller (Figure 3.4-11) responds to the sensed mismatch between I vessel average and reference temperature and between turbine and reactor power and starts to withdraw the control bank rods. The excore detectors magnify the power mismatch signal. The rod controller continues to I Page 60 of 332
I pull the control bank rods until the rod speed error signal drops below (in absolute magnitude) the lockup temperature g (-1.0 *F). When the rods stop, the rod speed error begins 5 to increase (in absolute magnitude) and then slowly decreases as the power and temperature approach a steady state. If the increase in rod speed error, after a rod pull, is sufficient to reach the deadband temperature, (-1.5 F) then a second rod pull is initiated and the cycle continues until a new steady state condition is reached. For the PI 1 Cycle 9 analysis, four separate rod pulls are initiated with a total of 134 pcm being added by the automatic rod controller. Figure 3.4-12 shows the actual response of the automatic rod controller including all of the functions shown in Figure 3.4-11 for PI 1 Cycle 9. Figure 3.4-13 shows the calculated response of the controller vari:bles without simulating the controller circuitry. The automatic circuitry causes an additional three rod pulls (29 pcm), after the initial pull, and thus causes a power overshoot. The results for PI 1 Cycle 9 and PI 2 Cycle 8 show transient MDNBRs of 1.586 and 1.913, respectively. g The large variation in MDNBR between the two cycles 5 is due to the variation in Fg (dropped rod). The core average transient responses are very similar. 3.4.4 Cycle Specific Physics Calculations These calculations are performed at the most limiting core conditions found during the cycle, e.g. the point in time that gives the most conservative value of the parameter in question. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. Page 61 of 332 1 I
I
- a. Moderator Temperature Coefficient, g Calculations of M are performed in accordance with the general procedure described in section 2.0.
I b. Doppler Temperature Coefficient, D Calculations of D are performed in accordance with the general procedure described in section 2.0.
- c. Nuclear Heat Flux Hot Channel Factor, F g
I The nuclear heat flux hot channel factor, FQ, I is calculated for all possible dropped rods, consistent with the procedure described in section 2.0. Each rod is dropped at full power,
- d. Nuclear Enthalpy Rise Hot Channel Factor, F 3g I The nuclear enthalpy rise hot channel factor, FAH, is calculated for all possible dropped rods, I consistent with the procedure described in section 2.0. Each rod is dropped at full power, equiliorium xenon and with highly skewed power shapes (both positive and negative axial offsets).
The FAH is then censervatively adjusted for allowed quadrant tilt. I e. Effective Delayed Neutron Fraction, S,ff The value of S,ff is calculated in accordance with the general procedure described in section 2.0. Cycle specific calculations are performed for each dropped rod condition. i 1 I ease e2 ef 222
I
- f. Promot Neutron Lifetime, ta The value of E,ff is calculated in accordance with the general procedure described in section 2.0.
Cycle specific calculations are performed for each dropped rod condition.
- g. Dropped Rod Worth, apDROP Calculations of the dropped rod worth are performed with the nodal model in three dimensions. Cycle i specific calculations are performed for each dropped rod. All calculations are performed with the control rods at the FPIL. The worth of the dropped rod includes considerations of equilibrium as well as highly skewed power shapes (both positive and negative axial offsets).
- h. Control Bank Worth, apCONTROL Calculations of the integral worth of Bank D ,
(control bank) are performed with the nodal model in three dimensions. Cycle specific calculations are performed for those dropped rods and sets of conditions which have been calculated not to cause a reactor negative flux rate trip.
- i. Excore Tilt 1
Calculations of the relative tilt as seen by the I excore detectors are performed by correlating he core edge power densities based on the distance squared between the edge assemblies and the detector. A 2% allowed tilt (Technical Specifications) is included in the calculations. Page 53 of 332 I'
I I 3.4.5 Reload Safety Evaluation l Each of the physics parameters calculated are adjusted to include the model reliability factors, RF , and 9 biases, 8 9, (Reference 1). These adjusted values are used in both parts of the rod drop analysis. I 3.4.5.1 Part 1 Flux Rate Trip Analysis The adjusted physics parameters are calculated for all possible dropped rods, consistent with the procedure described in Section 2.0. These parameters are used to determine the spectrum of rods which will not trip the I reactor when dropped. The physics parameters should be adjusted as follows. CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS l b. ag-RFM-BM 2 aM (m st negative bounding value) I a. aD*(1+RFD ) *
"D (m st ne:ative bounding value)
- c. B eff * (1 + RFg ) s 6,ff (maximum)
I d. t* * (1 + RFg ) s t* (maximum)
- e. ApDROD * (1 - RFR ) 2 ApDROD(bounding)
- f. Excore Tilt + T s Excore Tilt I
I - I Page 64 of 332 !I l - .
E 1 3.4.5.2 Part 2 Transient Consequences The adjusted physics parameters are calculated, for those rods which do not cause a negative flux rate trip, consistent with the procedure described in section 2.0. These parameters are used to determine the transient response following a rod drop. The physics parameters should be adjusted as follows. CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. ag+RFg+BM ' "M (least negative bounding value)
- b. ' "0 (least negative aD (1-RFD )
bounding value)
- c. (FQ+RFpg+8pg) * (1+T) s FQ (maximum) s FAH (maximum)
- d. (FAH+RFFAH+0FAH)*(1+T)
- e. 8,ff * (1 - RFg ) 2 S,ff (minimum)
- f. t* * (1 - RFg) 2 t* (minimum)
- g. ApDROD * (1 + RFR ) s ApDROD (maximum)
- h. ApCONTROL * (1 + RFR ) $ ApCONTROL (maximum)
- 1. Excore Tilt - T 2 Excore Tilt I
I l Page 65 of 332 I
I TABLE 3.4-1 PI 2 CYCLE 8 DROPPED ROD RESULTS , I Xe Condition Rod Ap(pcm) Tilt Trip Xe = Eq, BOC F-6 200 1.02419 yes D-4 167 1.04791 yes C-7 129 1.07656 yes C-5 156 1.07331 yes B-6 122 1.08180 yes Xe = +AO, BOC F-6 200 1.02417 yes 0-4 169 1.04797 yes C-7 118 1.07190 yes I C-5 B-6 158 123 1.07337 1.08184 yes yes Xe = -AO,BOC F-6 200 1.02411 yes I D-4 167 1.04749 yes C-7 143 1.08120 yes C-5 157 1.07261 yes I Xe = Eq,EOC B-6 F-6 124 208 1.08124 1.02437 yes yes D-4 194 1.05035 yes I C-7 C-5 119 178 1.06842 1.07323 no yes B-6 146 1.08410 yes Xe = +AO,EOC F-6 206 1.02433 yes 0-4 200 1.05092 yes I C-7 C-5 B-6 95 182 148 1.05831 1.07331 1.08364 no yes yes I Xe = -AO,E0C F-6 D-4 C-7 203 190 161 1.02409 1.04886 1.07872 yes yes yes C-5 176 1.07098 yes I B-6 150 1.08252 yes I I I Pe0e e6 or 332 I
I, TABLE 3.4-2 PI 1 CYCLE 9 l DROPPED RCD RESULTS I l Xe Condition Rod ap(pcm) Tilt Trip Xe = Eq, BOC F-6 169 1.02380 yes D-4 164 1.04866 yes a C-7 130 1.08198 yes 1.07396 yes E C-5 144 B-6 133 1.09074 yes Xe = +AO, B0C F-6 169 1.02379 yes 0-4 167 1.04891 yes C-7 115 1.07512 yes E C-5 145 1.07382 yes g B-6 134 1.09036 yes Xe = -AO,80C F-6 168 1.02371 yes D-4 163 1.04817 yes C-7 148 1.08799 yes C-5 144 1.07308 yes B-6 134 1.08993 yes Xe = Eq,EOC F-6 203 1.02416 yes D-4 191 1.05049 yes l C-7 124 1.07169 yes a C-5 170 1.07379 yes B-6 141 1.08388 yes Xe = +AO,EOC F-6 203 1.02410 yes D-4 196 1.05109 yes C-7 96 1.05984 no C-5 175 1.07361 yes B-6 143 1.08290 yes Xe = -AO,EOC F-6 197 1.02387 yes 0-4 186 1.04887 yes C-7 170 1.08263 yes C-5 170 1.07151 yes B-6 147 1.08238 yes I Page 67 of 332 l I
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1 Page 79 of 332 l
f DYNODE-P Pl-1 CYCLE 9 s.... .ac...,.ii.,si ..i Dro ed Rod. EOC DNP017/84 s_. ,_, _. .a_ p. . . , e,,,, S.ns d Temperatur. Error Sensed Error 4- _ 2- /\- - i- - - - --
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0 50 100 15 0 200 250 300 350 400 Seconds f igur e 1. l 11 Page 80 of 332 W W W M M M M M M M M M M M M m m m m
I I 3.5 Uncontrolled Boron Dilution I 3.5.1 Definition of Accident The accident considered here is the malfunction of the chemical and volume control system in such a manner as to deliver unborated water at the maximum possible flowrate to the reactor coolant system under full power conditions. Dilution during refueling or startup is assumed to be recognized and terminated by operator intervention before loss of shutdown margin. With the I reactor in automatic control, the power and temperature increase from baron dilution at power results in the insertion of the RCC assemblies and a decrease in shut-down margin. Rod insertion limit alarms would alert the operator to isolate the source of unborated water and initiate boration prior to the time that shutdown margin was lost. With the reacto'r in manual control, the power I and temperature rise due to boron dilution would eventually result in an overtemperature AT reactor trip if the operator did not intervene. After such a trip, the operator is expected to isolate the unborated water source and initiate boration procedures. 3.5.2 Accident Analysis I The system transient response to an uncontrolled boron dilution is simulated using a detailed model of the plant which includes the core, reactor vessel, steam generators, pressurizer, and connecting piping. The model also includes a simulation of the charging and the letdown systems, pressurizer control systems, and the I reactor protection systems. Reactivity effects due to the fuel and moderator feedbacks, coolant boron concentration, and control rod motion after trip are included in the analysis. This model I Page 81 of 332 I
I provides the transient response of average core power, reactor coolant pressure, and coolant temperature at the core inlet which are applied as forcing functions to a thermal-hydraulic simulation of the hot channel. The hot channel model uses the W-3 correlation to calculate the departure from nucleate boiling ratio in the hot channel. The reactivity due to boron dilution is calculated by assuming the maximum possible charging flow and minimum reactor coolant volume and taking into account the , effect of increasing boron worth as dilution continues. The core burnup and corresponding boron concentration , are selected to yield the most limiting combination of moderator temperature coefficient, Doppler temperature coefficient and spatial power distribution. This is normally the 80C condition. The minimum shutdown margin allowed by the Technical Specifications is conservatively assumed to exist prior to the initiation l of the transient. The maximum time delay is assumed to . exist between the time the trip setpoint is reached and the rods begin to move into the core. The most reactive rod is assumed to remain in its fully withdrawn position . after receipt of the trip signal. The acceptance criteria for this accident are that - pressures in the reactor coolant system and main steam system do not exceed 110% of the respective design pressures, and that fuel ciad integrity is maintained by limiting the minimum DNBR greater than 1.3. 3.5.3 NSP Safety Analysis Experience NSP has analyzed a chemical and volume control system malfunction resulting in a decrease in the baron concentration of the reactor coolant. The analysis was Page 82 of 332 I
I performed using the model described in Appendix A with input consistent with the FSAR (Reference 2). The results are compared to those presented in Section 14.1.4 of Reference 2. Sensitivity studies indicate that critical input parameters in an analysis of the boron dilution accident are the moderator temperature I coefficient, the baron worth coefficient, and the parameters used in the overtemperature AT trip set point algorithm. The NSSS and hot channel transient response calculated by the NSP model are shown in Figures 3.5-1 to 3.5-4. No corresponding transient results are given in Reference 2, however, reactor trip on overtemperature AT I was stated to occur at 78 seconds. The trip time calculated using the NSP model was 77 seconds, also on overtemperatura AT, with a 6.0 second delay. From Figure 14.1-10 of Reference 2, the minimum DNBR corresponding to this rate of reactivity insertion is 1.365. The NSP multichannel 1/8 assembly hot channel I model result is 1.633. 3.5.4 Cycle Specific Physics Calculations These calculations are performed at th most limiting core conditions found during the cycle, eg., the point in time and the power level that gives the most I conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. for each parameter. I I Page e3 ef 332 I
Il
- a. Moderator Temperature Coefficient, aM I' i Calculations of aM are performed using the methods j described in Section 2. Cycle specific calculations for this accident are made at unrodded full power.
- b. Doppler Temoerature Coefficient, aD Calculations of aD are performed in accordance with the procedure described in Section 2.
I
- c. Boron Concentration Reactivity Coefficent, a B
Calculations of ag are performed using methods de-scribed in Section 2. Cycle specific calculations for this accident are made at unrodded full power.
- d. Shutdown Margin, SDM I
For refueling and startup modes (cold), the shutdown margin is calculated directly with all rods in rather than with one stuck rod, consistent with the assumptions made in the safety analysis. Calculations of SDM are performed using methods described in Section 2. I I I I Page 84 of 332 I! ! I'
I
- e. Nuclear Enthalpy Rise Hot Channel Factor, F AH 1
The maximum core F is assumed to remain AH within the current limits as defined in the Technical Specifications for allowable com-binations of axial offset and power level. For I Prairie Island, the continuous surveillance of the power distribution is accomplished with the excore detectors using the Exxon PDC-iia (3) scheme. 3.5.5 Reload Safety Evaluation All the cycle specific parameters discussed above are adjusted to include model reliability factors, RF j , and
. biases, B . These results are then compared to the bounding values assumed in the safety analysis. The cycle specif,1c parameters are acceptable if the following inequalities are met:
I CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. ag+RFM+Bg s aM (least negative bounding value)
- b. aD*(1-RFD ) ' "0 (least negative bounding value)
- c. 2 B (m st negative B (1+RF B )*(1*08 )
I d. Refueling and Cold 2 bounding value) SDM (bounding) Startup Conditions SDM (ARI) At Power Conditions SDM 2 SDM (bounding)
- e. s (FaH+RFFAH+BFAH)(1+T) FAH (Technical Specifications)
I Page 85 of 332 I
i 4 FSAR Boron Dilution 1.6 E-5 dK/sec DNP127 Absoluto Power vs Time i 120 . : . . i 11 0 - - - - 100 .
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i , .. . 70 80 90 100 10 20 30 40 50 60 0 Seconds Page 86 of 332 Figure 3.5-1 m M M M m _ m _ m m m m m M M _- M M m m m M --
man sus aus uma sus i aus ese am num nas aus sus as sus sum um aus FSAR poron Dilution ' 1.6 E-5 dK/sec i DNP127 I Pressurizer Pressure vs Time 2310- A -
- : . i . . ! -
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1 - i . : : : - - 1 2170- .. ., , , , . 0 10 20 30 40 50 60 70 80 90 100 l Seconds , Page 87 of 332 i Floure 3.5-2
e . I FSAR Boron Dilution 1.6 E-5 dK/sec DNP127 Vessel Average Temperature vs Timo 580- - : 678- . i.- 576- - - r : i . 574- - - -- sa
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som em aus mas sus sus, mas em aus sus sus : sus mas sus sus um e*
- FSAR Boron Dilution t
1.6 E-5 dK/sec A CBR086 W-3 DNB Ratio vs Time 2 -
+
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3.6 Startup of an Inactive Coolant Loop 3.6.1 Definition of Accident Since there are no isolation valves or check valves in the Prairie Island reactor coolant system, operation of the plant with an inactive loop causes reversed flow through that loop. If there is a thermal load on the steam generator in the inactive loop, the hot leg coolant in that loop will be at a lower temperature than the core inlet temperature. The startup of the pump in the idle loop results in a core flow increase and the injection of cold water into the core, followed by a rapid reactivity and power increase. The resulting increase in fuel temperature limits the power rise due to Doppler feedback. Above 10% rated power, however, the reactor protection system prevents operation with en inactive loop, and consequently the temperature g differential in an inactive loop would be small enough E to minimize the accident consequences. Furthermore, the Prairie Island Technical Specifications do not permit operation with a reactor coolant pump out of service except during low power physics testing. 3.6.2 Accident Analysis The system transien*. -asponses to an inactive loop startup is simulated tsing a detailed model which includes the core, reactor vessel, steam generators, main steam and reactor coolant piping, and the plant control and protection systems. This model calculates the time-dependent behavior of the average core power, coolant pressure, and core inlet flow and temperature which are supplied as forcing functions to c model of the hot channel for calculation of DNBR. I Page 90 of 332 I) 1
I The accident is analyzed using the most negative moder-ator temperature coefficient calculated to occur during the cycle. This is normally the EOC condition. No credit is taken for reactivity reduction caused by reactor trip. The reactor is initially assumed to be operating at 12% of rated power with reverse flow through the inactive loop. This includes a 2% uncertainty for calibration error above the 10% power setpoint for single loop operation in the reactor protection system. The assumption of this high initial power level is conservative, since it maximizes the temperature difference between the hot leg and cold leg in the inactive loop. The most adverse combination of initial coolant pressure and core inlet temperature is chosen to minimize the margin to core DNB limits. I The acceptance criteria for this accident are that the maximum pressures in the reactor ccalant and main steam systems do not exceed 110% of design values and that cladding integrity be maintained limiting the minimum DNB ratio greater than 1.30. 3.6.3 NSP Safety Analysis Experience I NSP has analyzed the inactive loop start-up accident using the models and methods described in Appendix A. The results obtained are compared to the results presented in Section 14.1-5 of Reference 2. The flow in both loops was linearly ramped to the I nominal value in 10 seconds as stated in Reference 2. Figures 3.6-1 to 3.6-5 provide a comparison of NSSS transient response to Figures 14.1-14, 14.1-15(a) and Page 91 of 332
14.1-15(o) of Reference 2. The results of the NSP model compare well with those of Reference 2. The same transient was run using a dynamic flow simulation of the RCS pumps. The NSP model shows that , l' the inactive pump will reach full speed in approximately 22 seconds as compared to 20 seconds as stated in . Section 14.1.5 of Reference 2. 3.6.4 Cycle Specific Physics Calculations These calculations are performed at the most limiting core conditions found during the cycle, eg., the point in time and the power level that gives the most conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. for each parameter.
- a. Moderator Temperature Coefficient, o g Calculations of a Mare perf rmed in accordance with the general procedures described in Section 2.
Specific calculations for this accident are E performed for hat zero power, rodded, no xenon E. conditions. The model bias, B,is included in the calculations.
- b. Doppler Temperature Coefficient, a0 Calculations of aD are performed in accordance with the general procedures described in Section 2.
Page 92 of 332 I
- c. Nuclear Enthalpy Rise Hot Channel Factor, F AH The maximum core F AH is assumed to remain within the current limits as defined in the Technical Specifications for allowable combinations of axial offset and power level. For Prairie Island the continuous surveillance of the power distribution is accomplished with the excore detectors using the Exxon PDC-iia (3) scheme.
3.6.5 Reload Safety Evaluations Each of the physics parameters calculated above are conservatively adjusted to include the model reliability factors, RF4 , and biases, B 9. These adjusted values are the cycle specific parameters which are then compared to the bounding values assumed in the safety analysis. The cycle specific parameters are acceptable if the follow-I ing inequalities are met: CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. ag+Bg-RFM 2 aM (m st negative bounding value)
- b. 5 O (least negative aD (1-RFO )
bounding value)
- c. (FAH+RFp3g+BFAH)(1+T) s Technical Specifications I
Page 93 of 332
l l I e. FSAR Idle Loop Start-up DNP124 (I i i
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i Delta Core Inlet Temperature vs Timo 5
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F1 uro 3.6-1 Page 94 of 332 M M - _ - _-_- E ._ __. - _ _ . .-. -__ E M . .._... ~ -. M - M .. E
m e um up mas sus mas aus mas sus e e mim FSAR Idle Loop Start-up owi24 10 Sec Linear Acceleration . F. .S.A.R. . . l Delta Core Average Temperature vs Time i i 1 . . . .
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! Floure 3.6-2 Page 95 of 332
i 1
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1 Floure 3.6-5 Page 98 of 332 W M M M M M M M M M M M M M W W M W .-- W
I 3.7 Feedwater System Malfunction 3.7.1 Definition of Accident Two classes of accidents are to be considered under this classification: Those that result in a decrease in feedwater temperature and those that result in an I increase in feedwater flow. Either condition will result in an increased heat transfer rate in the steam generators causing a decrease in the reactor coolant temperature and an increased core power level due to negative reactivity coefficients and/or control system action. For the case of a decrease in feedwater temperature, the worst accident which may be postulated I involves opening the bypass valve which diverts flow around the feedwater heaters. For the case.of an increase in feedwater flow rate, the worst accident
- which may be postulated involves the full opening of a 6 feedwater control valve. For this case, sustained high feedwater flow rate would ultimately result in a reactor trip due to high steam generator water level.
I 3.7.2 Accident Analysis The feedwater system malfunction transient is analyzed using a dynamic simulation which includes core kinetics and heat transfer, reactor vessel and coolant piping, steam generators, pressurizer, and control systems. I Pertinent variables obtained from the NSSS simulation are then applied as forcing functions to a separate thermal-hydraulic model of the core hot channel which calculates DNBR. I Two cases are analyzed. The first case is for a reactor without automatic control and with the least negative I moderator coefficient. This is normally the BOC condition. This represents the situation where the Page 99 of 332 I
I I reactor has the least inherent transient response capability. In this case, the core power slowly increases due to Doppler reactivity effects until the core power level again matches the load demand and a new g steady state is achieved. The reactor does not trip. E The core power increase is caused by a conlant temperature decrease which has the effect of increasing the margin to DNB. The second case analyzed assumes that the reactor auto-matic control system responds to the decreasing coolant temperature and matches reactor power to load demand. , A conservatively large (in absolute value) negative moderator temperature coefficient is assumed to exist. This is normally the EOC condition. This case results in a somewhat higher final core power level than the y uncontrolled case without moderator feedback; this in turn results in a net decrease in DNBR but the decreased coolant temperature again maintains a significant margin above the 1.3 limit. The core neutronic characteristics which exert a I significant influence on the calculated results of this transient are the Doppler and moderator reactivity coefficients. The most negative Doppler temperature E coefficient calculated to occur during the cycle is used 9 in the analysis to maximize the power increase. For such slow rates of reactivity addition as are encountered, the transient response is insensitive to the value of t*, the thermal neutron lifetime. Trip reactivity insertion characteristics are not relevant, since the reactor does not trip. The acceptance criteria for the feedwater system malfunction transient are that cladding integrity be Page 100 of 332 I l I
I I maintained by limiting the minimum DNBR to be greater than 1.3 and that maximum pressure in the reactor I coolant and main steam system not exceed 110% of the design pressure. 3.7.3 NSP Safety Analysis Experience I Not all classes of feedwater system malfunction transient have been analyzed. Specifically, the transient that was analyzed was the opening of the feedwater heater bypass valve. The NSP safety analysis I experience in this area is represented by the analysis of a decrease in feedwater temperature transient using the model described in Appendix A. This calculation has been performed using input consistent with the Prairie Island FSAR (Reference 2). The models used correspond to BOC conditions without I control and EOC conditions with control. Figures 14.1-16 and 14.1-17, 14.1-18 and 14.1-19 of Reference 2 were used to obtain forcing functions of the transient feedwater-steam flow, and feedwater enthalpy for the two cases respectively. The response of the NSSS for the BCC case is compared to , Figures 14.1-16 and 14.1-17 of Reference 2 in Figures 3.7-1 to 3.7-4. The hot channel transient ONBR was computed using the NSP 1/8 assembly multichannel model l and is compared to Figure 14.1-16 of Reference 2 in Figure 3.7-5 for the BOC case. The EOC comparisons to Figures 14.1-18 and 14.1-19 of Reference 2 are shown in Figures 3.7-6 through 3.7-10. The NSP models predict the same trends throughout the transient as the FSAR results. Page 101 of 332 I , l
I 3.7.4 Cycle Specific physics Calculations These calculations are performed at the most limiting core conditions found during the cycle, eg., the point in time and the power level that gives the most conservative result with respect to the acceptance criteria. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc, for each parameter,
- a. Moderator Temperature Coefficient, "M Calculations of ag are performed in accordance with the general procedures described in Section 2.0.
Cycle specific calculations are performed to g determine the least negative aM at full power E conditions and the most negative aM under all operating conditions. The model bias is included in these calculations,
- b. Doppler Temperature Coefficient, a 0
Calculations of 0 are perf rmed in accordance with the general procedures described in Section 2.0. Cycle specific calculations for this accident are performed as a function of power level over the full operating range from 0-100'.' power.
- c. Nuclear Enthalpy Rise Hot Channel Factor, F 3g The maximum core F AH is assumed to remain within the current limits as defined in the Technical Specifications for allowable combinations of axial offset and power level. For Prairie Island, the continuous surveillance of the power distribution Page 102 of 332 I
l is accomplished with the excore detectors using the Exxon PDC-iia (3) scheme. 3.7.5 Reload Safety Evaluations i Each of the physics parameters calculated above are adjusted to include the model reliability factors RF, and biases, B9 . These adjusted values are the cycle specific parameters which are then compared to the I bounding values assumed in the same analysis. The cycle specific parameters are acceptable with regard to feedwater malfunction transients if the following inequalities are met: CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. Case 1 g + RFM+OM S "M (least negative bounding value)
Case 2 "M - RFg+Bg 2 g (most negative boundingvalue)
- b. 2 a0 (m st negative I aD (1-RFD )
bounding value) s FAH (Technical Specifi-
- c. (FaH+RFFAH+8FAH)(1+T) cations)
I I I I Page 103 of 332 I
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