ML18092B070
| ML18092B070 | |
| Person / Time | |
|---|---|
| Site: | Salem  |
| Issue date: | 03/14/1986 |
| From: | Blake R, Kent R, Rosenfeld E Public Service Enterprise Group |
| To: | |
| Shared Package | |
| ML18092B069 | List: |
| References | |
| NFU-0039, NFU-0039-R01, NFU-39, NFU-39-R1, NUDOCS 8604070325 | |
| Download: ML18092B070 (205) | |
Text
NFU-0039 Revision l**
March 14, 1986 The Energy People*
SALEM REACTOR PHYSICS METHODS
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NFU-0039 Revision 1 March 14, 1986 SALEM REACTOR PHYSICS METHODS r?a.ikk Date 3-1-4-Bb
- - Prepared by Reviewed by R. A. Bla~e Nuclear Technology Engineer
-;rR.__J.s. kfJ-- Date 3-1~-'B&
Kent Senior Engineer Approved by 24.~
E. S. R0Seneld Date 3-/ ~
Manager - Nuclear Fuel Copy No. 15
NFU-0039 Revision 1 March 14, 1986 ABSTRACT This document is a Topical Report describing the Public Service Electric and Gas Company <PSE&G> reload safety evaluation methods for application to the Salem Units.
The report addresses the determination of calculational uncertainties, the methods for the calculation of cycle specific physics parameters and their comparison to the bounding values used in the accident analyses.
i
1-1 I
NFU-0039 Revision 1 Mar-ch 14, 1986 REVISIONS The following sections of Revision 1 of this topical r-epor-t ar-e new sections which have been added to Revision 0:
a) Section 3.7 b) Sections 4.0 thr-ough 4.6 c) Sections s.o thr-ough 5.14 d) Appendix c e) Appendix 0 f) Appendix E In addition, the r- ef er-ences section has been moved to Section 6.0 in this revision.
Other- changes to Revision 0 of this topical r-epor-t ar-e indicated by solid bars in the right hand margin of the appropriate pages, except for page nu~ber-ing purposes *
- ii
NFU-0039 Revision 1 March 14
- 1986 TABLE OF CONTENTS
1.0 INTRODUCTION
1.0-1 2.0 OVERVIEW OF THE CALCULATIONAL MODEL 2.0-1 3.0 MODEL VERIFICATION AND RELIABILITY DETERMINATION 3.0-1 3.1 Rod Worth Benchmarking 3 .1-1 3.2 Isothermal Temperature Coefficient Benchmarking 3.2-1 3.3 Doppler Coefficient Benchmarking 3.3-1 3.4 Isotopics 3.4-1 3.5 Reliability Factors far Delayed Neutron Parameters 3.5-1 3.6 Power Distribution Benchmarking 3.6-1 3.7 Verification of Transient Power Distribution Simulation Capability 3.7-1
.4.0 GENERAL PHYSICS METHODS FOR SAFETY EVALUATION 4.0-1 4 .1 Power Distribution Analysis 4 .1-1 4.2 Reactivity Coefficients 4.2-1 4.3 Shutdown Margin 4.3-1 4.4 Scram Reactivity 4.4-1 4.5 Effective Delayed Neutron Fraction 4.5-1 4.6 Prompt Neutron Lifetime 4.6-1 5.0 ACCIDENT SPECIFIC SAFETY EVALUATION METHODS 5.0-1
- 5. 1 Uncontrolled Rod Bank Withdrawal from Subcritical Condition 5 .1-1 5.2 Uncontrolled Rod Bank Withdrawal at Power 5.2-1 5.3 Control Rod Mi sa l i gnment 5.3-1 5.4 Dropped Control Rod ) 5.4-1 5.5 Uncontrolled Boron Dilution 5.5-1 5.6 Feedwater System Malfunction 5.6-1 5.7 Excessive Load Increase 5.7-1 5.a Loss of External Load 5.8-1 5.9 Loss of Reactor Coolant Flow - Pump Trip 5.9-1 5.10 Loss of Reactor Coolant Flow -
Locked Rotor 5 .10-1 5.11 Fuel Handling Accident 5.11-1 s.12 Main Steam Line Break
- 5. 12-1 s.13 Control Rod Ejection 5 .13-1 s.14 Loss of Coolant s.14-1 iii
NFU-0039
' ' Revision 1 March 14, 1986 TABLE OF CONTENTS <continued) *
6.0 REFERENCES
6.0-1 APPENDIX A: Statistical Methods for the Determination and Application of Uncertainties A APPENDIX 8: Computer Code Summary Description B APPENDIX C: Sample of Proposed Revisions to the Technical Specifications c APPENDIX D: Sample of Proposed Power Distribution Limit Report D APPENDIX E: Sample of Proposed AFD Limit Calculations E iv
NFU-0039 Revision 1 March 14, 1986 LIST OF FIGURES Figure 2 .o .1 Salem Physics Model 2.0-3 3.3.1 Comparison of Measured and Calculated Doppler Test Parameters 3.3-4 Comparison of EPRI-CELL to Yankee Pu239/Pu240 Isotopic Ratios 3.4-3 Comparisons of EPRI-CELL to Yankee Pu240/Pu241 Isotopic Ratios 3.4-4 Comparisons of EPRI-CELL to Yankee Pu241/242 Isotopic Ratios 3.4-5 Salem Unit 1 and Salem Unit 2 Moveable Incore Detector Locations 3.6-6 Axial Locations of Grids and Detectors 3.6-7 Measured and Calculated Integrated Detector Responses Salem 1 Cycle 4 MAP 1411 3.6-11 Measured and Calculated Detector Responses Salem 1 Cycle 4 MAP 1411 3.6-12 Measured and Calculated Detector Responses Salem 1 Cycle 4 MAP 1411 3.6-13 Measured and Calculated Integrated Detector Responses Salem 1 Cycle 5 MAP 1522 3.6-14 Measured and Calculated Detector Rsponses Salem 1 Cycle 5 MAP 1522 3.6-15 Measured and Calculated Detector Responses Salem 1 Cycle 5 MAP 1522 3.6-16 Measured and Calculated Integrated Detector Responses Salem 2 Cycle 1 MAP 2133 3.6-17 3.6.10 Measured and Calculated Detector Responses Salem 2 Cycle 1 MAP 2133 3.6-18 3.6.11 M~asured and Calculated Detector Responses Salem 2 Cycle 1 MAP 2133 3.6-19 v
NFU-0039
-* Revision 1 March 14, 1986 Figure LIST OF FIGURES <continued)
Page Distribution of Errors XCi,k,m) 3.6-22 Distribution of Errors for Integral X<i, m> 3.6-23 Confidence Levels for XCi,k*m> versus Reactor Power % 3.6-26
- 3. 6. 15 Confidence Levels for XCi,k,m) versus Cycle Exposure 3.6-27 Confidence Levels for XC"i,k,m> versus Axial Height 3.6-28 Confidence Levels for X<i,m> versus Reactor Power 3.6-29 Confidence Levels for XCi,m> versus Cycle Exposure 3.6-30 Measured Increase in FQ 3.7-5 Comparison of Measured and Predicted FQ Increases 3.7-6 Relationship Between FQ<Z> and A.O.
for Full Power, Unrodd~d Conditions 4 .1-23
- 4. 1. 2 The General Relationship Between FQ<Z>
and Axial Offset 4.1-24 Typical Values of T<Z> and DF<Z> 4 .1-25 Effect of Initial Rod Position on SCRAM Reactivity Salem 1. Cycle 7 Beginning of Cycle <Hot Full Power) 4.4-4 Effect of Initial Rod Position on SCRAM Reactivity Salem 1 Cycle 7 End of Cycle <Hot Zero Power) 4.4-5 lCAPP 0) Cycle <N> AFD Engineering Factors D-2 vi
NFU-0039 Revision 1 March 14, 1986
- Table LIST OF TABLES 3.0.1 Reliability Factors and Biases for PSE&G Model Applied to Salem 3.0-2 3.1 ol Dilution Mode Rod Worth Comparisons 3.1-3 3.1.2 Rod Exchange Rod Worth Comparisons 3.1-4 3.1. 3 Rod Worth Reliability Factors 3.1-6
- 3. 2 .1 Measured and Calculated Isothermal Temperature Coefficients 3.2-3 Comparison of Measured and Calculated Doppler Test Parameters 3.3-3 Comparison Between EPRI-CELL and SAXTON Experimental Data 3.4-2 3.6.1 Reactor State Points 3.6-8 3.6.2 Reactor State Points 3.6-9 3.6.3. Reactor State Points 3.6-10 3.6.4 Mean Observed Differences Axial Model Bias 3.6-20 Axial Region Definitions 3.6-21 Confidence Limits for X<i,k,m) Distri-bution by Subgroup 3.6-24 Confidence Limits for X<i,m> Distri-bution by Subgroup 3.6-25 Flux Maps Used for Verification of Transient Power Distribution Simulation Capability 3.7-4 Bounding Initial Conditions for Evaluation of Reactivity Coefficients 4.2-5 Bounding Initial Conditions for Evaluation of Shutdown Margin 4.3-4
- vii
NFU-0039 Revision 1 March 14, 1986 Table LIST OF TABLES <continued)
- Analytical Case Sequence for Evaluation of SOM 4.3-8 4.s.1 Bounding Initial Conditions for Evaluation of fiEFF 4.5-3 Single-Sided Tolerance Factors AS Values of m for 95% Confidence and 95% Probability Tolerance Limits A8 viii
NFU-0039 Revision 1 March 14, 1986
1.0 INTRODUCTION
Sections 2 and 3 of this report describe the Salem reactor physics model and address the qualification and quantifi-cation of reliability factors for application of the model to operations and reload safety evaluations of the Salem Nuclear Reactors.
Sections 4 and 5 of this report describe the methods for the calculation of Salem cycle specific physics parameters and their comparison to the bounding values used in the accident analyses
- A summary description of the computer codes used to model the Salem reactors is given in Section 2.
The qualification of the model is described in Section 3.
Whenever possible, directly observable parameters Cs~ch as rod worths, and incore detector fission rates) are utilized for this qualification. The data used in this evaluation span seven (7) reactor operating cycles. The reactor cycles included are Cycles 1 through 5 for* Salem Unit 1, and Cycles 1 and 2 for Unit 2. In addition, some data from Salem 1 Cycle 6 and Salem 2 Cycle 3 were included in the verification of transient power distribution simulation
- capability in Secti-0n 3.7.
Page 1.0-1
NFU-0039 Revision 1 March 14, 1986 After the measured data to be used in the benchmark process are defined, the model calculations are performed and are compared to measurements. These comparisons are presented in this report as part of the quantification of the PSE&G model calculational uncertainties. and reliability factors.
A statistical approach i~ used to evaluate the uncertain-ties and reliability factors. These uncertainties and reliability factors are con?istent with the model appli-cation procedures and methodology,, The uncertainties and reliability factors are evaluated by direct comparison to experimental data.
ln order to provide a continuing verification of the conservatism of the reliability factors determined herein, ongoing comparisons are made each cycle using statistical methods consistent with those described in this report.
A description of the general physics calculational methods used for safety evaluations is provided in Section 4.
"General methods" are those that apply to several different accident evaluations. "Unique methods" which apply to only I
i
!i a single accident are discussed in Section s.
I Section 4.1 describes a significant improvement over the II traditional approach used to evaluate reload power distributions. The traditional approach is to perform a le I
Page 1.0-2
NFU-0039 Revision 1 .
March 14, 1986
- single Reload Safety Evaluation <RSE> which considers all possible operating scenarios for the cycle operation.
These include continuous load follow operation which can result in skewed core burnup distributions. This tradi-tional analysis is completed, and the Tech. Spec. Limiting Conditions for Operation <LCO's> are set for the entire cycle, prior to reactor startup. In-core flux map measure-ments which are made subsequent to startup are then used to confirm the conservatism of the RSE predictions.
The PSE&G approach is to perform the RSE in several phases
- The first phase is performed prior to the reactor startup, but the remaining phases are performed after startup, and utilize the flux map measurements to re-evaluate the power distribution LCO's on a monthly basis. The implementation of this approach will require changes to the power distri-bution Technical Specifications which are identified in Section 4. These new LCO's have been dubbed "FLEX-SPECS" at PSE&G.
A general description is given in Section 5 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 charac-
- teristics is provided.
Page 1.0-3
NFU-0039 Revision 1 March 14, 1986 Calculations of core physics parameters for the purpose of performing reload safety evaluations requires an intimate knowledge of the safety analyses to which cycle specific comparisons are to be made. Specifically, one must under-stand the manner in which the bounding physics parameters have been used in each of the analyses and the conserva-tisms inherent in the values chosen. In order to acquire such an understanding, Public Service Electric and Gas Company <PSE&G) has developed models for performing various safety analyses for Salem and has performed representative calculations for the incidents of importance for a reload evaluation. The results of these calculations are included in Reference 10. These results demonstrate the PSE&G safety analysis experience and exemplify the expertise required to make the determinations as to whether or not an accident must be reanalyzed and to perform the necessary analyses for a given fuel cycle.
Page 1.0-4
NFU-0039 Revision 1 March 14, 1986
- 2.0 OVERVIEW OF THE CALCULATIONAL MODEL The model used to analyze the Salem Units was constructed using the Advanced Recycle Methodology Program (ARMP) system developed under EPRI sponsorship by UAI. (Reference 1)
A flow diagram for this model is shown in Figure 2.0.1.
The spectral code, EPRI-CELL (ARMP, Part II, Chapter 5),
produces initial nuclide concentrations, depletion and fission product chain data, and tables of microscopic and macroscopic cross sections varying with burn-up for input to the XY diffusion - depletion code, PDQ7/HARMONY (Reference 2 and 3). Lumped absorber* data for PDQ7/HARMONY are generated by a capture fraction matching procedure between PDQ7 and either EPRI-CELL (ARMP, Part I, Chapter 6, Section 4) for burnable poisons or CPM (ARMP, Part I, Chapter 6, Section 3) for control rods. PDQ7/HARMONY*is ;I 1
run both in the full core "(XY) geometry representation and the fuel type (color set) representation. The full core ::1 representation is used for nodal code normalization, local peaking factor generation, and for the establishment of assembly loadi.ng patterns.
In the fuel type (color set) mode, PDQ7/HARMONY supplies input data for PSE&G's nodal code, TRINODE, a derivative of the EPRI-NODE-P program (ARMP, Part II, Chapter 14). The TRINODE program contains improvements over the EPRI-NODE-P program which include input/output changes, execution options, and file management. However, the primary calculational sequence and physics methodology have been preserved from the EPRI-NODE-P program.
It is recognized that the methods used for the construction and application of the Salem model are as much a part of the model definition as are the codes. It is essential, therefore, that the methods u~ed to calculate core safety margins be consistent with those used in the model benchmarking and qualifications process. This is particularly true in the calculation of core power distribution and local peaking factors in which the results are heavily dependent on the methods used to normalize the nodal model.
NFU31/l 8
~ 2.0-1
NFU-0039 Revision 1 March 14, 1986
- 2. 0 OVERVIEW OF THE CALCU.t.ATIONAL MODEL (continued)
The TRINODE model is normalized to the PDQ model. A consistent methodology is used for this normalization throughout the benchmark calculations and will be used in future safety related calculations.
In addition to the main sequence computer codes, a number of auxiliary computer codes are employed to provide a user tailored code package. These auxiliary computer. codes are not basic to the physics methodology, but are vital for automation and transformation of the large volume of calculated and measured parameters required for core~
analysis. The auxiliary computer codes are summarized in Appendix B.
All comparisons to measurement data presented in this topical report are based on TRINODE calculations.
NFU31/1 9 2.0-2 I.
I I .
L_ - _j
NFU-0039 Revision 1 FIGURE 2.0.l March 14, 1986 SALEM PHYSICS MODEL CPM EPRI-CELL
'I EPRI-CELL LUMPED ABSORBER NUPUNCHER PROCEDURE Ir '
PDQ/HARMONY FUEL TYPE (COLOR SET) FULL CORE
- EPRI FIT 'I NORMALIZATION I
REACTION RATES PIN TO BOX SUPER LINK TRI NODE SIGMA PLANT MEASUREMENTS 2.0-3
NFU-0039 Revision 1 March 14, 1986 3.0 MODEL VERIFICATION AND RELIABILITY.DETERMINATION The PSE&G model is benchmarked against Salem Unit 1 measurements made during Cycles 1 through 5, and Salem Unit 2 measurements made in Cycles 1 and 2. This benchmark serves as the basis to* quantify the reliability factors to be used in safety related calculations.
The term reliability factor (RF) is used to describe the allowances (either absolute or relative) to be used in safety related calculations to assure conservatism. The term uncertainty factor is used to describe the actual model precision and is defined as the standard deviation
( ). The reliability factor is always larger than the uncertainty factor.
The term bias is used to describe the statistical difference between an observed or measured distribution and the calculated value.
Table 3.0.1 summarizes the model reliability factors and biases computed as a result of the model benchmark. The remainder of Section 3 is a detailed account of the derivation of these factors. The statistical methods employed are described in Appendix A to this report *
- NFU31/l Page 3.0-1
NFU-0039 Revision 1 March 14, 1986 TABLE 3.0.1 RELIABILITY FACTORS AND BIASES FOR PSE&G MODEL APPLIED TO SALEM PARAMETER RELIABILITY FACTOR
-BIAS
. Rodworth MEAS>600 PCM RF ROD = lS% 0 MEAS<600 PCM RF ROD = lOOPCM 0 TOTALS RF ROD = 10% 0
. Temperature Coefficient Moderator (MTC) RFMTC = 2.1 PCM/°F 0 Isothermal (ITC) RF ITC = 2.1 PCM/°F 0 Doppler RFDC = 10% 0 Doppler Defect RFDD = 10% 0
. Delayed Parameters Neutron B ef f RFB = 4% 0 Jl,
- RFL = 4% 0
. Power Distribution FQ P> .so RFFQ = 0.10 **
P< .so RFFQ = 0.16 -(0.12*P) **
- t. Fo/F Q RFTZ = 8% 0 F t.H P> .30 RFFDH = 0.08 0 P< .30 RFFDH = 0.09 -(P/30) 0
- See Table 3.6.4 NFU31/l Page 3.0-2
NFU-0039 Revision 1 March 14, 1986 3.1 Rod Worth Benchmarking The purpose of this section is to benchmark the PSE&G Salem model to rod worth measurements. This is accomplished by first presenting and qualifying the available measurements and second by computing model reliability factors.
Rod worth measurements have been performed at Salem using two techniques: the boron dilution method and the rod exchange fechnique. (Reference 9)
Boron dilution rod worth measurements were performed on Unit 1 Cycles 1 th.rough 5 and Un_i t 2 Cycles 1 and
- 2. The results of these measurements are summarized on Table 3.1.1 along with PSE&G model predictions.
Rod exchange measurements were performed on Unit 1 Cycles 1, 3, 4, and 5, and Unit 2 Cycle 2. These results, along with model calculations are tabulated 'I on Tables 3.1.2 (a) and (b).
For purposes of model benchmarking, some rod worth measurements are disqualified on the basis of known measurement errors. Measurements disqualified are the boron dilutions for Unit 1 Cycles 1 and 2, and Unit 2 Cycle 1. Additionally, rod exchange measurements for Unit 1 Cycle 1 are disqualified. The basis for this disqualification is measurement errors discovered in dilution measurements made prior to Cycle 3. These errors are due to the effects of spatial flux redistribution caused*by .rod motion during the dilution (Reference 9). Test procedure changes were implemented prior to Cycle 3 measurements to reduce these effects. Since rod exchange measurements use the reference bank dilution measurement to interpret exchange worths, rod exchange measureMents for Unit 1 Cycle 1 are disqualified on the same basis.
Support for the disqualification of dilution measurement made prior to Cycle 3 is available using comparisons to calculated worths. The average difference between measured and calculated rod worths for dilution measurements performed prior to Cycle 3 and those using the improved test procedure are 11%
and 1% respectively. This differ~nce is significant at the 99.9% confidence level, and is attributed to the known measurement errors.
Page 3.1-1 NFU31/l
NFU-0039 Revision 1 March 14, 1986 Rod worth reliability factors were obtained by bounding the results of the comparisons between measured and calculated rod worths. These factors are tabulated on Table 3.1.3.
Comparisons were taken from 7 dilution measurements and 24 exchange measurements spanning 4 reactor cycles, and represent a11 *measurements through Cycle 5 of Unit 1 and Cycle 2 of Unit 2, except those disqualified above.
Using a bounding value for the reliability factor is justi-fied due to its conservatism relative to normal statis-tics. Calculation of reliability factors representing 95/95 confidence levels using normal statistics yields a 91 PCM reliability factor adder for rods worth less than 600 PCM, and 12% reliability factor for rods worth more than 600 PCM. The exception is the reliability factor for rod worth totals, which is computed to be 16%. However, this large value is due to the small sample size of only 5 values. Since the error for rod worth.totals can be no larger than the largest error for the individual rod banks, the reliabilty factor for rod worth totals should be bounded by the maximum observed error for individual banks with worth greater than 600 PCM: 10%. Thus, the factors tabulated on Table 3.1.3 conservatively bound the observed data, and will be used as model rod worth reliability factors.
Page 3.1-2 NFU31/l
NFU-0039 Revision 1 March 14, 1986 TABLE 3. 1. 1 DILUTION MODE ROD WORTH COMPARISONS DATE UNIT/ BANK MEAS CALC DIFFERENCE*
% 7S 12/76 1/1*** D 1107 1030 7.5 c 1183 1005 17.7 B 766 724 5.8 A 1241 1114 11.4 SD 745 681 9.4 SC 1181 1060 11. 4 TOTAL 6223 5614 10.8 12/79 1/2*** D 1041 924 12.7 c 938 846 10.9 B 534 599 -65 A 1163 973 19.5 TOTAL 3676 3342 10.0 8/80- 2/1*** D 1391 1241 12.1 c 1185 1026 15.5 B 1359 1262 7.7 A 501 385 116 SD 750 712 5.3 SC 1052 961 9.5 TOTAL 6238 5587 11. 7 12/80 1/3 D 834 797 4.6 c 960 900 6.7 B 565 600 -35 A 1023 1058 -3.3 TOTAL 3382 3355 0.8 4/82 1/4 D 862 860 0.2 2/83 1/5 D 926 939 -1. 4 7/83 2/2 D 878 835 5.1
- % = ( (M-C}/C}
- 100 for measurements >600 PCM.
/j, = (M-C} for measurements <600 PCM.
- Data disqualified as discussed in text.
Page 3.1-3 NFU31/l
NFU-0039 Revision 1 March 14, 1986 TABLE 3.1.2 ROD EXCHANGE ROD WORTH COMPARISONS DATE UNIT/ BANK MEAS CALC DIFFERENCE CYCLE PCM PCM .M>600 M< 600
-% 6 12/76 1/1*** *D 1107 1030 (7.5) c 825 741 11. 3 B 522 467 55 A 924 858 7.7 SD 469 403 66 SC 351 305 46 TOTAL 4198 3804 10.3 12/80 1/3 *D 834 797 ( 4. 6) c 696 674 3.3 B* 395 450 -55 A 816 789 3.4 TOTAL 2741 2710 -1.1 4/82 1/4 *D 862 860 ( 0. 2) c 596 588 8 B 370 407 -37 A 818 789 3.7 SD 265 316 -51 SC 285 281 4 SB 614 649 -5.4 SA 750 733 2.3 TOTAL 4560 4623 -1. 4 2/83 1/5 *D 926 939 (-1.4) c 613 617 -0.6 B 331 361 -30 A 784 814 -3.7 SD 269 292 -23 SC 317 291 26 SB 769 793 -3.0 SA 735 779 -5.6 TOTAL 4744 4886 -2.9 (continued)
Page 3.1-4 NFU31/l
NFU-0039 Revision 1 March 14, 1986 TABLE 3.1. 2 ROD EXCHANGE ROD WORTH COMPARIOSNS (continued)
DATE UNIT/ BANK MEAS CALC DIFFERENCE*
% !J.
7 /83 2/2 *D 878 835 ( 5. 1) c 770 731 5.3 B 660 603 9.5.
A 252 233 19 SD 299 287 12 SC 292 275 17 SB 787 757 4.0 SA 562 491 71 TOTAL 4500 4212 6.8
- Measurement performed by dilution
- % = ((M-C)/C).
- 100 for measurements >600 PCM
!J. = (M-X) for measurements <600 PCM
- Data disqualified as discussed in text *
- Page 3.1-5
NFU-0039 Revision 1 March 14, 1986 TABLE 3.1. 3 RODWORTH RELIABILITY FACTORS Individual Rod Worth a} Rodworth <600 pcm RFROD = 100 pcm b} Rod worth >600 pcm RFROD = 15%
Total Rod Worth RFROD = 10%
NFU31/l Paqe 3.1-6
NFU-0039 Revision 1 March 14, 1986 3.2 Isothermal Temperature Coefficient Benchmarking The objective of this section is to benchmark the PSE&G model to measured isothermal temperature coefficients (ITC). Based on comparisons between measured and calculated coefficients, a reliability factor for both the isothermal and the moderator temperature coefficient (MTC) is inferred.
A total of 19 ITC measurements are tabulated on Table 3.2.lo These measurements span 7 reactor cycles and range from unrodded conditions to all control banks inserted.
The PSE&G model calculations for ITC are presented on Table 3.2.1 along with the corresponding measurement.
Statistical tests were performed on the comparisons to evaluate normality and pooleability. Normality was demonstrated using the W-test (Reference 8), while
-pooleability was assured using the Bartlett test (Reference 4). The computed standard deviation of the comparisons between measured and calculated ITC's is 0.85 PCM/F.
The observed standard deviation of 0.85 PCM/F ( crossv) is assumed to be made up of three independent components; measurement uncertainty, model calculational uncertainty on moderator temperature coefficient, and, model calculational uncertainty on Doppler temperature coefficient. This relationship is expressed as:
a 2 + cr 2 + a2 = a 2 = (0.85)2 MEAS MTC DC OBSV Since each component is greater than or equal to zero,each component is bounded by the observed error, Therefore, a conservative estimate of the model uncertainty ( cr ) for both the isothermal and moderator temperature coefficients is o.as PCM/F. This is summarized as: 1
/
= 0.85
= 0.85 Page 3.2-1
.NFU31/l 19
NFU-aa39 Revision 1 March 14, 19a6 PSE&G model reliability factors for both ITC and MTC are computed as the product of the standard deviation and the one-sided critical factor (K c) for a 95/95 confidence level u~ing nineteen (19) samples. This product yields reliability factors for ITC and MTC of 2.1 PCM/F.
RFITC = a.as
- 2.42 = 2.1 PCM/F RFMTC =a.as* 2.42 = 2.1 PCM/F
)
Page 3.2-2 NFU31/l
NFU-0039 Revision l March 14, 1986 TABLE 3.2.l MEASURED AND CALCULATED ISOTHERMAL TEMPERATURE COEFFICIENTS ROD POSITION ITC PCML°F UNIT . CYCLE BANK (STEPS) BORON MEAS CALC DIFF l l D 197 1369 -3.51 -3.59 0.00 c 201 1264 -4.11 -4.34 0.23 B 175 1151 -6.17 -6.45 0.28 A 175 1085 -7.85 -9.06 l. 21 SD 175' 965 -11.25 -11. 90 0.65 l 2 D 219 1137 -6.06 -4.45 -1. 61 c 214 1025 -5.79 -5.45 -0.34 l 3 D 219 1258 -3.33 . -3. 26 -0.07 c 206 1157 -4.85 -4.10 -0.75 l 4 D 202 1309 -3.61 -5.29 l. 68 l 5 D 214 1499 -1.52 -2.60 l. 08 2 l D 205 1334 -0.65 -0.59 -0.06 D 188 1329 -0.84 -0.70 -0.14 D 102 1285 -2.68 -1. 89 -0.79 c 184 1197 -4.34 -4.85 -0.51 B 203 1083 -10.53 -9.09 -1. 44 A 198. 955 -10.50 -9.83 -0.67 SD 192 910 -13.48 -13.22 -0.26 2 2 D 218 1362 -4.16 -4.55 0.39 Mean 0.00 Standard Deviation 0.85 Page 3.2-3
- NFU31/l
NFU-0039 Revision 1 March 14, 1986
- 3.3 Doppler Coefficient Benchmarking The objective of this section is to make comparisons between measured and calculated Doppler coefficients and*
establish model reliability factors for Doppler reactivity calculations.
Doppler reactivity coefficient measurements have been performed at the Salem Units using two measurement techniques. Both test procedures require t~e compensation of a reactivity imbalance induced by a reactor power change. The first test procedure balances reactivity using control rods and measures the
- reactivity changes using a reactimeter. This technique was used in Cycle 1 of both Salem Units. The second.
measurement procedure maintains reactivity balance with changes in moderator temperature. The ratio of power change to moderator temperature change is then converted to reactivity using an isothermal temperature coefficient. This technique has been used for Cycles 2 through 5 on Unit 1, and Cycle 2 for Unit 2.
The measurements using rod banks for reactivity control are not used for purposes of model benchmarking. The basis for this disqualification is the large uncertainties associated with reactimeter interpretation for at-power measurements.
The results of all Doppler coefficient measurements performed using the moderator temperature control procedure have been tabulated on Table 3.3.1. This measurement technique requires calculated isothermal temperature coefficients to infer the Doppler coefficient temperature. Since it is the ratio of the changes of these two quantities that is actually measured, this ratio is tabulated along with the inferred Doppler coefficient on Table 3.3.1. The precision associated with each measured ratio has been determined based on the standard deviation of multiple measurements.
Page 3.3-1 NFU31/l
NFU-0039 Revision 1 March 14, 1986 Calculations of the ratio of power to moderator temperature changes have been made using the PSE&G model. Comparisons of the measured and calculated ratio are shown on Table 3.3.l and also Figure 3.3.1 in which the vertical bars represent the measurement precision.
Figure 3.3.1 demonstrates that the measured and calculated ratios typically agree to within the measurement precision, and therefore confirms model capability to calculate these ratios. The scatter in the data shown in Figure 3.3.1 is due primarily to the poor measurement precision.
It is apparent from Figure 3.3.1 that the measurement precision is of the same order of magnitude as the observed differences between measurement and calculation. Thus, the model calculational uncertainty is assumed to be small. For purposes of assigning a model reliability factor for Doppler coefficient (RFDC),
a conserv.ative value of 10% is assumed. The same reliability factor will be assigned to the model for Doppler only power defect (RFDD). Thus:
RFDC RFDD
= 10%
= 10%
NFU31/l Page 3.3-2
- _J
TABLE 3.3.1 Ca.1PARISOO OF MEASURED AND CALCUIATED OOPPLER TEST PARAMETERS UNIT/ ~ NUMBER aD (Li p ) MEAS (Li P) MF.AS/
CYCLE % OF MEAS ECM,7% (Li T )MEAS PRECISION ( LiT)CALC CALC 1/2 39 6 -13.67 -0.95 0.04 -0.86 -0.09
~
93 6 -13.15 -1.39 0.09 -1.41 0.02 Ill l.Q CD 1/3 44 6 -13.31 -0.77 0.02 -0.83 0.06 94 6 -10.91 -1.40 0.17 -1.28 -0.12
.ww I 1/4 43 4 -10. ll -0.90 0.23 -0.84 -0.06 w 99 4 -ll.49 -1.28 0.09 -1.31 0.03 1/5 46 4 -ll.45 -0.66 0.04 -0.70 0.04 :s: ~ z 97 4 -12.01 -0.99 O.ll -1.10 O.ll Ill CD l':r:J 11 <: c 0 I-'* I
- rm o 2/2 98 2 -11. 33 -1.34 0.35 -1.33 -0.01 1-'*0
...... 0 w
.c::. ::s l.O l.O co
°'
NFU-0039 Revision* 1 March 14*, 1986 FIGURE 3.3.l COMPARISON OF MEASURED AND CALCULATED DOPPLER TEST PARAMETERS *
- 1. 7 1.6
- 1. 5
- 1. 4 1.3
~
0 dP 1.2 8
<l
~ 1.1
<l ro (I) lo-I 1. 0
- s(/]
(1j (I)
- 0.9 (m=l) 0.8 0.7 0.6 0.5 ,-.____..____......____...____.______._____..____...._____..____....&.____....
0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 1.5 Calculated 6P/ 6T, %/°F Page 3.3-4
NFU-0039 Revision 1 March 14, 1986
- 3.4 Isotopics Isotopic compositions calculated by EPRI-CELL have been compared with spent fuel isotopic data obtained from Yankee Rowe fuel rods irradiated* beyond 35 GWD/MTU. The reactor representation used for the EPRI-CELL .
benchmarking .in the calculations is described in the ARMP documentation (Part 1, Chapter 1, Section 4.0).
Experimental and analytical isotopic ratios for plutonium from the Yankee Rowe spent fuel are plotted versus accumulated fissions in Figures 3.4.1-3.4.3. The dots are experimental results and the line the EPRI-CELL results.
The agreement between calculated and experimental isotopic ratios is good. The calculated ratios Pu-239/Pu-240 ~nd Pu-240/Pu-241 are within the scatter of the experimental results and the ratio Pu-241/Pu-242 is slightly over-predicted.
Calculated and measured isotopic compositions for M02 fuel (6.6 w/o) irradiated in the Saxton Core II (pellet, rod MY, zone 6 )' are compared in Table 3. 4 .1. The agreement is good for the most important uranium and plutonium isotopes as well as for americium and curium.
The measured burnup ranged from 7 to 22 GWD/MTU. The r~actor representation for the EPRI-CELL benchmarking is described in the ARMP documentation (Part 1, Chapter 3, Section 4.)
Page 3. 4- 1
- NFU31/l
NFU-0039 Revision 1 March 14, 1986 TABLE 3.4.l COMPARISON BETWEEN EPRI-CELL AND SAXTON EXPERIMENTAL DATA NUCLIDE EXPERIMENT EXPERIMENTAL UNCERTAINTY
~PRI-CELL~*l00%
Exp.
ATOM %
U-234 .00465 28.7 .- 3.7 U-235 .574 .9 + 1. 7 U-236 .0355 5.6 - 7.6 U-238 99.386 0 0 Pu-238 .109 2.2 -32.l Pu-239 73.77 0 + 0.6 Pu-240 19.25 .2 - 3.5 Pu-241 6.29 *3 + 4.6 Pu-242 .579 .9 -11. 7 ATOM RATIOS 4
Np-237/ 1.14 x 10- 15 -26.9 U-238 Pu-239/ .04383 .7 + 2.0 U238 Pu-238/ 1. 75 x lo- 3 .4 -17.6 Pu-239 Am-241/ .0123 15 + 2.4 Pu-239 Cm-242/ 1. 05 x lo- 4 10 + 0.4 Pu-239 4 Cm-244/ 1.09- 20 - 3.2 Pu-239 NFU31/l Page 3.4-2
NFU-0039-*
Revision 1 FIGURE 3.4.1 March 14, 1986 COMPARISON OF EPRI-CELL TO YANKEE Pu23-~i/Pu~40 ISOTOPIC RA'l'IOS 10.0 .,.
9.0
\
a.a *.
~
7.0 .
\
6.0
\
<1 s.o ~
~
- 4. 0 ~
~
~ *'
- 3. 0 o* J
- o.o s.o 10.0 15.0 20.0 Accumulated Fissions (barn-cm)-l x io 5 Page 3.4-3 25.0 30.0
NFU-0039 Revision 1 FIGURE 3.4.2 March 14, 1966 COMPARISON OF EPRI-CELL TO YANKEE .Pu24Q/Pu241 ISOTOPIC RATIOS 9.0 a.a 7.0
\I I*
. \
6.0 :
\I*
0 M
~ s.o
\
\ \: *.
r.:*
3.0 2.0
\' .
~ . .
i:-____
1.0 a.a s.a 15.a 2a.o 25.a 3a.
- 5 Accumulated Fissions (barn-cm)-l x ia Page 3. 4.-4
NFU,-0039 Revision.l FIGURE 3.4.3 March 14, 1986.
COMPARISON OF EPRI-CELL TO YANKEE Pu241/Pu242 ISOTOPIC RATIOS
- 10. 0 .. ...
~
- 9. 0
.\
- a. a \
~
1-4
~ 7. o*
N
~
..-*. \
N I\
Q.o r-4
~
N
- > 6. 0
- .*... -~
Q.o
- s. 0 :
""'*.~ .... :
- 4. 0 a.fo.o 5.0 10.0 15.0 20.0 25.0 30.0
-1 5 Accumulated Fissions (barn-cm) x 10
- Page 3.4-5
NFU-0039 Revision 1
- 3.5 March 14, 1986 Reliability Factors for Delayed Neutron Parameters This section deals with determining reliability factors for the effective delayed neutron fraction and the effective neutron lifetime which are values which can be calculated but whose measurement is not practical. In these cases, an argument is made for the general magnitude of the reliability factor without making direct comparisons between measured and predicted values.
The importance of the reliability of the calculated values of the delayed neutron parameters is primarily associated with the core 8eff
- The uncertainties in the calculation of*Seff are composed of several co~ponents, the most important of which are listed below.
- a. Experimental values of 8 , and A. , by nuclide;
- b. Calculation of the spatial nuclide inventory;
- c. Calculation of core average 8 as a flux weighted average over the spatial nuclide inventory;
- d. Calculation of 8eff from the core av*erage as f3 eff =
I* 8
- where I = importance factor.
The experimental determination of the 8 1 s and A.*s are assumed to be accur.ate to within 1 %* The most important nuclide conc~ntrations with respect to core (3 are u238, u 235 , and Pu 39
- Tables 3.4.1 and 3.4.2 indicate that the difference in the calculation of these concentrations is about 1.7% for ECALL. Therefore, components (a) and (b) above are combined as 2.7%.
I~
Page 3.5-1
- NFU31/l
NFU-0039 Revision 1 March 14, 1986
- The uncertainty in the calculation of a core average S depends on the relative flux weighting of the individual assemblies in the core. For demonstration purposes, consider a three region core, each with a different average burnup and average S. This is typical of advanced PWR cycles in that about a third of the core.has seen two previous cycles, a third only one previous cycle and a third is the feed fuel.
Typical regional B 's are given below:
Region 1 (third cycle fuel) B (1) = 0.005 Region 2 (second cycle fuel) B (2) = 0.006 Region 3 (feed fuel) B (3) = 0.007 The effect of errors *in the calculated flux distribution can be evaluated in terms of the effect on the core average ~ As a base case, flux weighting factors (FWF) are all set to 1.0. In this case, the core average B = 0 .006. using a maximum error in the -.
regional flux weighting of 7%, the worst error in the calculation of the core average B is obtained by .
increasing the weight of the Region 1 fuel and decreasing the weight of the Region 3 fuel. The revised S is calculated as follows:
(l)xl.07 = .00535 (2)xl.O = .0060 (3ix0.93 = .00651 B = .00595, which yield a -0.8% error for component (C) above.
/
Page 3.5-2 NFU31/l
NFU-0039
'I Revision 1 !
March 14, 1986 The last uncertainty component, (d), concerns the reduction of core average 13 to obtain 13 eff by using t.he importance factor. Since this reduction is typically about 3% to 4%, an error of 10% in this component would lead to an error in 13 ef f of less than 0.5%.
The sum of the errors for these four factors for ECELL are as follows:
2.7%(a+b) + 0.8%(c) + 0.5%(d) = 4.0%
So the reliability factor for delayed neutron parameters (RFB)is set at 4%.
An argument similar to the delayed neutron parameter argument is applied to the determination of the effective neutron lifetime ( R. *) uncertaLnty. The
-*~
uncertainty components which go into the calculation of R.
- are as follows: *..f:.
- (a)
(b)
Experimental values of microscopic cross sections; Calculation of the spatial nuclide inventory; and (c) Calculation of the core average effective neutron lifetime, R. *, as a flux weighted average over the spatial nuclide inventory which includes the effects of leakages.
Uncertainties for components (a) and (b) are assumed to be the same as described for the calculation of 13 eff, that is, a combination of 1% uncertainty in the experimental determination of nuclear cross sections and 1.7% uncertainty in the determination of
- NFU31/l Page 3.5-3
NFU-0039 Revision 1 March 14, 1986 the spatial nuclide inventory of ECELL. The core average neutron lifetime depends on flux weighting of local absorption lifetimes t*. If a conservative estimate of the error in regional power sharing (7%)
is used in determining the impact on the core average lifetime ( i*), the error in lifetime is on the order of 1.0%. Combining all of these uncertainties linearly results in a total uncertainty of 3.7%. Therefore, a 4% reliability factor (RFL) will be applied to the neutron lifetime calculation when applied to safety related calculations.
Page 3.5-4 NFU31/l
NFU-0039 Revision 1 March 14, 1986 3.6 Power Distribution Benchmarking It is the purpose of this section to quantify the PSE&G Salem model power distribution calculations. This is accomplished by first presenting the measurement data base, followed by a description of the calculational methodology. Second, comparisons are made between the measured and calculated quantities, and lastly, model reliability factors for power distribution calculations are computed.
The primary source of power distribution measurements for Salem Units 1 and 2 is the incore detector system.
This system consists of moveable incore fission chambers which respond to neutron flux. These neutron detectors traverse through instrument guide thimbles which are located at 58 positions throughout the core as shown in Figure 3.6.1. Measurement signals from these detectors are taken at 61 axial positions up the fuel assembly as illustrated in Figure 3.6.2, and are corrected by the on-site process computer to account for detector sensitivity, drift, and background. The corrected signals are then used to compute "measured" power distributions using. analytical data to convert the detector signals to interpreted powers in both instrumented and uninstrumented assemblies.
A total of forty-nine (49) flux maps were chosen for the purpose of benchmarking the PSE&G Salem model.
These flux maps span six reactor-cycles and represent typical steady state operation conditions. These include maps taken at powers ranging from near zero to 100 percent, and cycle exposures of zero to end of cycle, including some coastdown state points. A description of reactor conditions for each flux map chosen is given in Tables 3.6.1 through 3.6.3.
The approach taken to benchmark and qualify -the PSE&G Salem models for power distributions was to compare calculated and measured detector signals. The basis for this is twofold. First, the detector signals Page 3.6-1
- NFU31/l
NFU-0039 Revision 1 March 14, 1986 represent raw measurements and do not include interpretation, unlike "measured" power distributions.
Second, the ability of the model to compute) the detector signal requires the same processes as required to compute pin powers. Both calculations require the prediction of the localized fission rate, one in a pin pellet, the other in a fission chamber. The accuracy of the two calculations is essentially the same. The only difference is that there is a small self-shielding or flux depression in the pin which is not in the detector. The impact of this difference on the power distribution reliability factor is assumed to be negligible.
The simulated detector signals are calculated in a manner which is consistent with the calculation of local power peaking factors for the purpose of safety evaluation. The first step is to compute the power distribution under consideration. The resolution used is one node per fuel assembly, with 12 axial levels.
The simulated detector signals are obtained by using the nodal power at each axial level to predict a signal power density for that assembly at that level. This power is then converted to a relative reaction rate.
The conversion factors are calculated for each asembly location as a function of assembly exposure using a two-dimensional, full core PDQ7, fine mesh model. The 12 axial values in each assembly location are then synthesized using a truncated fourier sine series.
Grid flux depressions are then superimposed on the synthesized function using an empirical function designed to match the characteristics of flux depressions measured with in-core fission detectors.
The effect of the grid flux depressions is to raise the flux level in the axial region between grids while depressing the flux in the grid region. Consistency between the above calculations of instrument signals and the calculation of local peaking factors is assured by:
Page 3.6-2 NFU31/l
NFU-0039 Revision l March 14, 1986 A.* Using a common full core PDQ7 model, B. Using a conunon nodal model, and
- c. Using a common procedure to account for axial flux gradients and grid effects.
Typical comparisons of measured and calculated detector signals are shown in Figures 3.6.3 through 3.6.11. The figures are in sets of three and are representative of various core exposures including coastdown conditions.
For each statepoint the first figure of the set presents the differences between the measured and predicted signal integrals for all instrumented locations. The instr~mented core locations are indicated with circles in each of the figures. The second and third figures of each set present axial comparisons in two specific instrumented core ,-.;
locations. The measurements are shown as a solid continuous line over 61 axial levels. The predicted reaction rates are represented as open circles. The two core locations were chosen as typical of regions on the interior of the core and on the core periphery.
In all comparisons, both the predicted and measured reaction rates have been normalized to a core average value of unity for each map.
For purposes of quantifying comparisons, it is .
convenient to define the variable ORR (I,K,M) which represents the difference between measured and calculated detector signals or reaction rates at location I,K, and map M. Thus, DRR(I,K,M) = RRM(I,K,M) - RRC(I,K,M)
Where I = Radial Detector Location K = Axial Detector Location M = Map Index RRM = Measured Detector Reaction Rate RRC = Calculated Detector Reaction Rate
~ NFU31/l Page 3.6-3
NFU-0039 Revision 1 March 14, 1986 An average difference between measured and calculated reaction rates can be computed for each axial level as:
I I DRR(I,K,M)
DRR(K) = I M I I 1 I M where the summation over I is perfo:rmed for each available radial location, and M represents all flux map data except zero power maps. The mean observed differences thus computed are the axial model* bias and are listed in Table 3.6.4. Since it is easier to describe the model uncertainties in terms of deviations relative to the observed bias, a second variable can be defined as X(I,K,M) = (RRM(I,K,M)-RRC(I,K,M)) - ORR(K)
Where X(I,K,M) = The difference between measured and calculated reaction rates adjusted for the observed bias.
All model power distribution behavior can now be characterized by quantifying the difference population X(I,K,M) or the integral of X(I,K,M): X(I,M). This latter quantity is the biased difference between the measured and calculated detector signal integrals.
To better evaluate the behavior of the distributions of X(I,K,M) and X(I,M), the difference population was divided into selected* subgroups. The subgroups were chosen to parameterize the difference behayior as a
- function of axial height, reactor power level, and cycle exposure. These subgroups were defined in a manner to exclude the axial points at grid locations and the upper and lower six axial points.
The axial regions are defined on Table 3.6.5.
The difference population was evaluated for normality using the chi-squared test. This test demonstrates Page 3.6-4 NFU31/1
NFU-0039 Revision 1 March 14, 1986 that most of the subgroups cannot be considered normal. Typical comparisons of the difference population and a normal distribution is illustrated in Figures 3.6.12 and 3.6.13.
As indicated in Tables 3.6.6 and 3.6.7 and Figures 3.6.12 through 3.6.17, 95/95 confidence limits assuming normal statistics and 95/95 confidence limits based on non-parametric statistics are in good agreement. In some cases the non-parametric limit is somewhat lower than the normal limit which simply indicates that the actual (not normal) distribution is slightly more peaked with *fewer samples in the upper (higher M-C values) tail of the di~tribution than is predicted by the normal distribution.
To be conservative, 95/95 confidence limits were evaluated using both normal and non-parametric statistics as described in Appendix A. Confidence limits (95/95) were computed for each subgroup. These
'results are summarized on Tables 3.6.6 and 3.6.7 and Figures 3.6.14 to 3.6.18. Inspection of the figures show that the confidence limits are a function of axial height, reactor power level, and cycle exposure.
Generally, the confidence limits decrease with increased power and exposure. The approach taken to compute PSE&G model reliability factors was to bound the computed confidence limits. Thus, the model reliability factors are:
RFFQ = 0.10 p >
- 50 0.16 - (0.12*P) p <
- 50 RFFDH = 0.08 P> .30 0.09 - (P/30) P< .30 In order to assess the impact of possible dependence among data samples on reliability factors for local peaking factors, the effects of reducing the sample sizes by two thirds (2/3) were evaluated. It was found
.that the reliability factors are relatively insensitive to this reduction in sample size resulting in an increase in reliability factors of approximately .002 units for X(I,K,M) and .004 for X(I,M). As can be seen in Figures 3.6.14 through 3.6.18, the PSE&G reliability factors remain bounding and--are therefore not impacted significantly by possible dependence among data samples.
I Page 3.6-5
NFU-0039 Revision 1 March** 14, 1986 FIGURE 3.6.1 ~
SALEM UNIT 1 AND SALEM UNIT 2 MOVABLE INCORE' DETECTOR LOCATIONS R p N M. L K J H G F E D c B A 28 15 1 4 *3 51 2 10 30 39 52 3 5 36 43 4 11 38 31 24 17 54 14 6 8 6 44 32 16 47 7 23 58 29 46 48 50 49 34 8 57 ' 22 9 56 9 4 1 12 10 33 40 26 21 13 11 12 41 55 7 13 45 35 20 25 14 18 27 42 37 53 2 Page 3.6-6
NFU-0039 FIGURE 3.6.2 Revision 1 AXIAL LOCATIONS OF GRIDS AND DETECTORS March 14, 1986 T
2111.5"11 29.ISH 28.!5!11!1 1§
!i!
AXIAL .GRID 2J.!581 AXIAL DETECTOR LOCATIONS ~ SIGNAL LOCATIONS 'I
~
~
en
~
~
28.~I 29.5!19.
ALL MEASUREMENTS ARE IN INCHES 24.438
- 1,------..(1*
1.243 Page 3.6-7 BOTTOM OF FUEL ROD
NFU-0039 Revision 1 March 14, 1986 TABLE 3.6.l REACTOR STATE POINTS SALEM 1 CYCLE 2 MAP NO. CYCLE EXPOSURE POWER LEVEL D BANK POSITION (MWD/MTU) ( %) (STEPS) 174 0 o.o 228 188 2160 . 100. 0 228 190 3097 100.0 228 194 4382 100.0 228 196 6250 100.0 225 198 7275 82.0 2.06 1201 7945 67.0 218 MAP NO.
(MWD/MTU)
SALEM 1 CYCLE EXPOSURE CYCLE 3 POWER LEVEL
( %)
D BANK POSITION (STEPS) 1300 0 0 212 1313 500 100 228 1315 1040 100 222 1324 3165 99.5 228 1330 4100 97.0 220 1333 5670 97.0 228 1338 7060 96.8 228 1342 8800 75.0 202 Page 3.6-8 NFU31/1
NFU-0039 Revision 1 March 14, 1986 TABLE 3.6.2 REACTOR STATE POINTS SALEM 1 CYCLE 4 MAP NO. CYCLE EXPOSURE POWER LEVEL D BANK POSITION (MWD/MTU) ( %)' (STEPS) 14aa a a.a 211 14a8 10a 84.a 228 1411 56a 1aa.a 228 1412 158a 1ao.a 225 1413 2589 98.6 228 1414 3715 1aa.a 215 1416 3836 1aa.a 228 1417 4998 1aa.a 228 SALEM 1 CYCLE 5 MAP NO. CYCLE EXPOSURE POWER LEVEL D BANK POSITION (MWD/MTU) ( %) (STEPS) 150a 0 a.a 216 15a3 25 47.3 228 1507 140 99.3 228 1509 1391 100.a 228 1512 2531 99.9 226 1517 4662 99.5 228 1520 5444 1ao.o 218 1522 7185 99.9 228 1524 8923 1ao.o 228
- NFU31/1 Page 3.6-9
NFU-0039 Revision l March 14, 1986 TABLE 3.6.3 REACTOR STATE POINTS SALEM 2 CYCLE l MAP NO. CYCLE EXPOSURE POWER LEVEL D BANK POSITION (MWD/MTU) ( %) (STEPS) 2004 0 o.o 206 2102 2435 100.0 222 2115 4677 96.7 228 2120 7386 99.8 228 2122 9196 100.0 224 2127 11755 82. 2 219 2129 13357 82.0 220 2131 14192 82.8 228 2133 15403 82.5 219 SALEM 2 CYCLE 2 MAP NO. CYCLE EXPOSURE POWER LEVEL D BANK POSITION (MWD/MTU) ( %) (STEPS) 2201 0 o.o 220 2203 21 48.6 180 2205 47 72.l 214 2209 292 98.4 228 2210 564 99.0 228 2213 1120 99.0 228 2214 2106 99.1 228 2217 3195 99.2 228
)
NFU31/l Page 3.6-10 I
__ _J
NFU-0039 Revision* 1 March 14, 1986 FIGURE 3. 6. 3 Measured and Calculated Integrated Detector Responses SALEM 1 CYCLE 4 MAP 1411 Absolute Differences Power = 100.0'7..
Exposure = 560 MWD/MTU R .P N M L K J H G F E D c B A Abs. Dif f =
/
' (Meas.-Calc)*lOO 1
/
, ' v .......
\. 4. 1.)
2
~-9 ...I \!*8..) -0.1
~
,. ~
....... II' ....... v
-2.3
... )
~0.5) ' ~o.6
.. .)
,3. 3 )
' 3
/
/ /
6.4 1. 5 -2.l 4
..) I\.. ./ \. .)
'" r 'I
/ '\ /
- 1. 5 5 (2-~ \.1.8.) \.1. 2.)
r
~ ./
r '\ ,
' -0.1'
/
/
- 1. 7 6
~-6 ...I ,!*9,) -1.1
.)
'- .)
v
\. .)
7 '"
-1.2 '
/
-1.8
'\ I'
-1. 7
'\
-1.3 .
'.)
7
'/" \..
.) .I .)
I..
-.../ '\/
/
/
3.2' 2.1' '
'I" / ........ /
-2.3 -2.7 -1. 3 -1.8 :.2.1 -0.5 8 I\.. ~ "- .) ""'-- ~ I\. .J v
I\. .I v
'- ~ '" ..I .J v
II' 3.3' --4.2' 4.0 ' 2.6 "' 9 I\.. ~ I\.. .J I\. .I I\. .J v v I/
-1.4
' -3.3'.)
~2.1 f\.
10 I\.. .J I\.
v II' ....
I\..
4.4'
~
I/
-2.3 I\...
'..I II
'-2. 6 I\.
..J
- 2.9 I\.. ~
11 II'
/
-2.1
- \..
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'.)
12 v v
-..... II' ....... 17 0.4 -1.2 ' .
-3.8~
I\.. ~-3 )
13 I\.. ..I '- ./
v -..... v "' v --..... v --.....
l4 2.9 ~-4,.J 1\.-2 -~ -0.l I\.. .J ~ .J II' .......
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15
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Page 3.6-11
NFU-0039 Revision 1 March 14, 1986 MEASURED AND CALCULATED DETECTOR RESPONSES FIGURE 3.6.4 *--
SALEM 1 CYCLE 4 MAP1411 Legend YE.ASUltED 0 PltEDICTU n-t.a.E=ll>
.....iiiiiiiii~..,;;;;;iiii--...--.......... -: .. *****-********
2 . *****************~---'!IL
- eci***** ******** ... ******************o******* *****************-**o************************ *******-******** ********
ID ID I e e
II a
e ID
- : ~~
I I I ID ID
- ID *****-***************** ********* *******-********~************eaeeo-*****
o.s .................................................... *******-***************** .......... *******-*******
0 0...---...,.--.....----r--~..---.....--.,---...---...,.--....---..----..---~......
0 5 10 15 20 25 30 35 40 ..s 50 55 60 AXIAL POINTS Page 3.6-12
NFU-0039 Revision 1 March 14, 1986
- FIGURE 3.6.5 MEASURED ANO CALCULATED DETECTOR RESPONSES SALEM 1 CYCLE 4 MAP1411 2.5~---..-~----mmll!ll---~....ii--..-----
=1:°"58)~
1 Legend MEASURED nt.a.£: N13 0 PREDICT£D 2 .~........ *....**. ********-*************************-----
LaJ Cl) z 0
a..
~ 1.5 ........ ********-*******-***************** ........ ********-******** **************************-*****************
et:
et:
~
~
LaJ c .
LaJ 1 **..**..*..**.... ;. .*. .. . ..
~ .
ci et:
05 ....... .********-*******-*** ..... *****************-:********-*****************:*********
o.....,..__,._.....,__.....,_.,._.__....._.,_......,.__....._,._.....__.....__.........,.
0 5 10 15 20 25 30 35 -40 45 50 55 60 65 AXIAL POINTS Page 3.6-13
NFU-0039 Revision 1 March* 14'~ 1986 FIGURE 3 .6. 6 Measured and Calculated Integrated Detector Responses SALEM 1 CYCLE 5 MAP 1522 Absolute Differences Power = 99. 7%
Exposure = 7185 MWD/MTU R p N M L K J H G F E D c B A v v *Abs. Dif f =
I\.
0.3 ',)
-o.o I\.
',)
{Meas.-Calc)*lOO 1 v ""'\
0.8 ' /
~o. 9..1 2 I\.. ./
-1. 7
... '/
/
-1.8
\..
',)
- 0.6 3
v ,
0.1'
'\II'
\.
- 1. 7
,,)
'- /
0.6 "
~
4 r .... / 7 0.9 .r ~-6,) " \..0.1 /'
v ' 7 1.2
.)
-o.o
\..
'/
0.9 I\..
'.)
2.2
\..
.)
6 v .... r
-0.4 r-....
.) I\:
0.6
,,)
-1.5
\. ,,)
/
0.3
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/
7 v v v /
1.3
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.. 1. 7
\..
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- 1,J \.
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- 1. 7,,)
/ "\I' "\
~3. 0,) \.0. 7,,)
8 v /
- -o.5
",J
,,L 2,,)
9 v v ....
1.5' -1.5, 0.8 10 I'-. , I\.. I\. ~
tr v II' 1.1' '
1\..-2.~
I/
\.0.9,) - 1\..0.~
' 11 I\. ~
v v 0.7
\..
' ,,)
-0.9
\.. /
- 12 tr v v i/
-1. '9 0.6' '
"\
13
-2.3 ~.l.)
I\.. ./ '- ./ I\. /
v ......, v tr I/
0.7 -1.5
' ~1.9/ 0.3 14 I'- / I\.. ~ I\.. ./
v
~-8,) '
Page 3.6-14
NFU-0039 Revision 1 March 14, 1986 *
- MEASURED AND CALCULATED DETECTOR RESPONSES SALEM 1 CYCLE 5 MAP1522 FIGURE 3.6.7 2.5...--~-~~-~-"91!"'"-""'!"--~~~~-~-~-~---
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0 5 10 15 20 25 30 AXIAL POINTS Page 3.6-15 35 40 45 50 55 60 65
NFU-0039 Revision* 1 March 14, 1986 MEASURED AND CALCULATED DETECTOR RESPONSES FIGURE 3.6.8 SALEM 1 CYCLE 5 MAP1522 2.5------------------~-~-~~--.
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NFU-0039 Revision 1 March 14, 1986
- FIGURE 3.6.9 Measured and Calculated Integrated Detector Responses SALEM 2 CYCLE 1 MAP 2133 Absolute Differences Power 82.5%
Exposure = 154a3 MWD/MTU R p N M L K J H G F E D c B A r Abs. Dif f =
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Page 3.6-17
NFU-0039 Revision 1 March 14, 1986 FIGURE 3.6.10 MEASURED AND CALCULATED DETECTOR RESPONSES SALEM 2 CYCLE 1 MAP2133 2 ********-
1=:~:~~1 Legend MEASURED PltEDlcnD ea.***************"'********-******************o******** *******-*************************** *******-******** ********
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0 5 10 15 20 25 30 35 40 so 55 60 65 AXIAL POINTS Page 3.6-18
NFU-0039 Revision 1 March 14, 1986 FIGURE 3.6.11 MEASURED AND CALCULATED DETECTOR RESPONSES SALEM 2 CYCLE 1 MAP2133 2 ....... 1=:~~1. . . _ _ _ _ _ . ._ - . - - . . . =-~= 0 LiJ
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0 5 10 15 20 25 30 35 40 45 50 55 60 65 AXIAL POINTS Page 3.6-19
NFU-0039 Revision 1 March 14, 1986 TABLE 3.6.4 MEAN OBSERVED DIFFERENCES AXIAL MODEL BIAS AXIAL LEVEL MEAN DIFF AXIAL LEVEL MEAN DIFF I I X (I)* I X (I)
(TOP) 61 .119 31 -.012 60 -.024 30 -.014 59 -.023 29 -.016 58 .005 28 -.062 57 .018 27 -.039 56 .015 26 -.015 55 .018 25 .007 54 -.031 24 .010 53 -.020 23 .009 52 -.021 22 .010 51 -.009 21 .031 50 . .003 20 -.041 49 -.001 19 -.033 48 -.005 18 -.018 47 .011 17 -.009 46 -.016 16 -.003
- 45 -.040 15 -.001 44 -.024 14 -.001 43 -.019 13 .012 42 -.017 12 .029 41 -.023 11 -.024 40 -.033 10 .015 39 -.033 9 .035 38 -.004 8 .052 37 -e069 7 .066 3"6 -.036 6 .068 35 -.009 5 .064 34 -.001 4 .056
. 33 .005 3 .038 32 .007 2 .024 1
- 0~16 I
NFU31/1 Page 3.6-20
NFU-0039 Revision 1 March 14, 1986 TABLE-3.6.S AXIAL REGION DEFINITIONS REGION AXIAL POINTS 1 7 - 10 2 14 - 18 3 22 - 27 4 31 - 35 5 39 - 44 6 48 - 52
- NFU31/l Page 3.6-21
FIGURE. 3.6.12 DISTRIBUTION OF ERRORS X(i,k,m) 0.5 NON PARAMETRIC STATISTICS 95/95 CONF'IDENCE LIMIT OBSERVED 0.4 DISTRIBUTION I /"" NORMAL STATISTICS 95/95 CONFIDENCE LIMIT I
PSEG Rf' f'Q
\
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.."'" ::s. \0 I-'
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<X>
O'I
-3 -2 -1 - 1 2 3 STANDARD ER- UNITS (Z)
- FIGURE .*13 DISTRIBUTION OF ERRORS FOR. INTEGRAL X(i,m) 0.5 NON PARAMETRIC STATISTICS 95/95 CONFIDENCE LIMIT 0.4 OBSERVED DISTRIBUTION !,,--°"" NORMAL STATISTICS 95/95 CONFIDENCE LIMIT I I u
z 0.3 I w
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DISTRIBUTION I I ......
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-4 -3 -2 -1 0 1 2 3 4 STANDARD ERROR UNITS (Z)
NFU-0039 Revision 1 March 14, 1986 TABIE 3.6.6 CONFIIENCE UMITS FOR X(I,K,M) DISTRIBUI'ION BY SUBGROUPS RFACTOR CYCIE AXIAL NUMBER sr. DEV 95L95 CONFIDENCE LIMITS PaiER (%) EXPOOURE (G/T) REGIONS SAMPLEs NOR-1AL NCN - PARAMETERIC 0 ALL 1-6 10075 .075 .125 .139 50< p <70 ALL 1-6 8059 .045 .076 .063 100 ALL 1-6 49573 .036 .059 .063 100 E<2.5 1-6 22966 .041 .067 .072 100 2.5<E<6 1-6 19105 .031 .051 .052 100 6(E 1-6 7502 .028 .047 .040 100 ALL 1 6396 .036 .061 .069 100 ALL 2 7991 .033 .055 .065 100 ALL 3 9587 .034 .057 .069 100 ALL 4 7995 .035 .059 .067 100 ALL 5 9595 .037 .062 .079 100 ALL 6 8009 .039 .066 .073 Page 3.6-24 NFU31/1
NFU-0039 Revision 1 March 14, 1986 TABLE 3.6.7 CONFIDENCE LIMITS FOR X(I,M) DISTRIBUTION BY SUBGROUP REACTOR CYCLE NUMBER ST. DEV 95L95 CONFIDENCE LIMITS POWER (%) EXPOSURE (G/T) SAMPLES NORMAL NON - PARAMETERIC 0 ALL 322 .045 .081 .075 50< p <70 ALL 258 .034 .062 .056 100 ALL 1593 .028 .048 .055 100 E<2.5 739 .033 .057 .066 100 2.5<E<6 614 .024 .042 .042 100 6<E 240 .021 .038 .045 NFU31/l 59
/
Page 3.6-25 NFU31/l
FIGURE 3.6.14 CONFIDENCE LIMITS FOR X(i,k,m) VS REACTOR POWER 3 0.20 0.18 0.16 0~14
~
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'° z 0.10 wW
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0.04 0.02 25 75 ER%
CONFIDENCE LIMITS FOR ~(i,k,m) VS CYCLE EXPOSURE 0.14 0.12 PSEG RffQ 0.10-+----------------------------------
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FIGURE 3.6.16 CONFIDENCE LIMITS FOR X(i,k,m) VS AXIAL HEIGHT 0.12 PSEG RrFQ 0.10-+-----------------------------------
0.08 7 NON-PARAMETRIC STATISTICS 95/95 CONFIDENCE LEVEL ~
- - ---- - - --- - /
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0 25 50 75 100 PERCENT REACTOR POWER %
FIGURE 3.6.18 .
CONFIDENCE LIMITS FOR .X(i,m) VS CYCLE EXPOSURE 0.12 0.10 PSEG RF FAH 0.08
~
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95/95 CON ROEN CE LEVEL_/
0.02 2 4 8 10
NFU-0039 Revision 1 March 14, 1986
- 3.7 VERIFICATION OF TRANSIENT POWER DISTRIBUTION SIMULATION CAPABILITY The objective of this section is to evaluate the analytical capability to predict the differential changes in FQ associated with transient operating conditions. This capability is utilized during plant operation to evaluate the Limiting Conditions for Operation <LCO> associated with Axial Flux Difference
<AFD>, as described in Section 4.1 The approach is to compare measured and predicted values of fractional
- changes in FQ from a steady state reference condition to a transient or perturbed condition. There are two major difficulties associated with this approach. The first is the scarcity of measurements of significantly perturbed flux conditions. A significant perturbation is one with a large fractional change in FQ from the steady-state referehce condition. The second dif-ficulty is the measurement uncertainty associated with flux maps taken during transient conditions.
The majority of flux map measurements are taken near stable, equilibrium conditions in order to minimize the measurement uncertainty. The variation in FQ among these maps Cat a given core burnup and axial level> is on the order of +/-5% or less. This magnitude of varia-tion is too small, relative to the measuremen~ uncer-tainty, to determine reliable differential FQ changes.
The FQ measurement uncertainty assigned to all flux maps i5 8.15% <Tech Spec uncertainty is the product of 1.03 x 1.05 = 1.0815>. For maps taken under steady state, equilibrium conditions, the actual uncertainty is estimated to be on the order of 1 to 2%, which is
- much smaller than the assigned value. However, for maps taken under transient conditions, the uncertainty*
Page 3.7-1
NFU-0039 Revision 1 March 14, 1986 is much larger, and the assigned value is more appropr-iate. The primary reason for the increase in the measurement uncertainty for transient conditions is the length of time required for the measurement process.
All flux maps require 1 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> for the data acqui-sition process. Random variations in core paramete~s, due to the transient conditions during this process can significantly effect the interpretation of core FQ's.
The measurement uncertainty associated with the in-ferred fractional change in FQ between two flux maps, is the statistical sum of the individual uncertainties for each flux map. These individual uncertainties are 1-2% for the equilibrium map, and 8% for the transient map. The statistical sum of these two is on the order of 8%. Therefore, in order to make meaningful compari-sons of fractional FQ ch~nges, the magnitude of the
- measured change should be larger than 8%. This narrows the choice of measurements to those made at reduced power during the startup physics measurements.
Table 3.7.1 describes the flux maps used to infer the I
fractional changes in FQ. Note that with the exception of maps 2201 and 2209, these maps were not used in the statistical evaluations in Section 3.6. These compari-sons therefore represent an independent verification of the model capability.
Figure 3.7*.1 illustrates the comparison of the measured FQ's from one pair of flux maps. Note that only one val*ue of FQ is compared between each pair of grids.
Only increases in FQ from the reference condition were considered.
Page 3.7-2
NFU-0039 Revision 1 March 14, 1986
- Figure 3.7.2 graphically compares the measured and predicted changes in the maximum core FQ's for- all of the maps in Table 3.7.1. The dashed lines displayed on the graph represent the 8% measu~ement uncertainty.
The accuracy of the model is indicated by the line of best fit which is shown in Figure 3.7.2 as a solid line, which has a slope of 1.08. The results in Figure 3.7.2 demonstrate that, to within the scatter of the measurement uncertainty, the model predictions are 8%
too low. Therefore, when the model is used as described in Section 4.1, to predict differential increases in FQ from ~easured reference conditions, an uncertainty of 8% will be applied to the increase.
Additionally, the 8.15% measurement uncertainty wi~l also be applied to the measured values of the reference FQ's. The application of these two uncertainties will therefore bound all of the experimental observations shown in Figure 3.7.2 *
- Page 3.7-3
NFU-0039 Revision 1 March 14, 1986 TABLE 3.7.1 FLUX MAPS USED FOR VERIFICATION OF TRANSIENT POWER DISTRIBUTION SIMULATION CAPABILITY Unit Cycle Map Power Rods(l) A.O. Comment Na. 00 <steps) (%)
1 6 1625 99 228 -4.2 Reference 1603. 45 181 -9.5 1602 45 212 +8.4 1601 44 180 -4.1 1600 2 211 +18.6 2 3 2308 100 228 Reference 2303 47 190 2302 47 200 2301 48 212 +12.4 2300 2 210 +15.4 2 2 2209 98 227 -5.3 Reference 2201 2 110 +33'. 2 2200 2 D=O +21.8 C=O Note ( 1) Rod Position is for bank 0 except as noted.
ARO = 228 steps.
Page 3.7-4
NFU-0039 Revision 1 March 14, 1986
- FIGURE 3.7.1 MEASURED INCREASE IN FQ
- *I I
- i - - - ,___
- Page 3.7-5
NFU-0039*
Revision 1 March 14 , 1986 FIGURE 3.7.2 COMPARISON OF MEASURED AND PREDICTED FQ INCREASES
~~ --~~: -~--1--*---'---'----i---+-----'---+----+---+----+-~-+----+-----
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Page 3.7-6 I
__J
NFU;...0039 Revision 1 March 14, 1986
- 4.0 GENERAL PHYSICS METHODS FOR SAFETY EVALUATION In this section, the general physics calculational methods are described for application to Salem reload safety evaluations and design verification.
The general physics methods are those that apply to several different accident evaluations. Unique methods which apply to only a single accident are discussed in Section 5 on an accident-by-accident basis.
The general physics methods are described for each parameter or *class of parameters below. For each parameter, appropriate definitions are given, bounding initial conditions are descr~bed, and the key aspects rif the analytical procedures are outlined.
The most bounding initial conditions are those conditions that exist at the time of the initiation of the accident.
Therefore, the definition of these initial conditions are constrained to be consistent, or at least conservative with respect to the limits of the Technical Specifications. For all general methods described in this section, the bounding initial conditions are those associated with operational modes 1 and 2, as defined in the Technical Specifications, Page 4.0-1
NFU-0039 Revision 1 March 14, 1 986 with Keff = 1.0, and Tave > 547°F. Other conditions which may be associated with specific accidents are discussed in Section 5 as appropriate.
Page 4.0-2
NFU-0039 Revision 1 March 14, 1986 4.1 POWER DISTRIBUTION ANALYSIS This section describes a significant improvement over the traditional approach used to evaluate reload power distributions. The traditional approach is to perform a single Reload Safety Evaluation CRSE> which considers all allowable operating scenar~os for the cYcle operation.
These include continuous load follow operation which can result in skewed core burnup distributions. This traditional analysis is completed, and the Tech~ Spec.
Limiting Conditions for Operation <LCO's) are set for the entire cycle, prior to reactor startup. In-core flux map measurements which are made subsequent to startup are then used to confirm the conservatism of the RSE predictions.
The PSE&G appr~ach is to perform the RSE in several phases.
The first phase is performed prior to the reactor startup, but the remaining phases are performed after startup, and utilize the flux map measurements to re-evaluate the power distribution LCO's on a monthly basis. The impl~mentation of this approach will require changes to the power distribution Technical Specifications which are identified in Appendix c. These new LCO's have been dubbed "FLEX-SPECS" at PSE&G *
- Page 4.1-1
NFU-0039 Revision 1 March 14, 1986 The key concept of FLEX-SPECS is that the LCO for Axial Flux Difference <AFD> would not contain specific values of the AFD limits. Instead, it would refer to the "Power Distribution Limit Report", which would contain the speci-fie AFD limits for the next 40 EFPD of operation. This report would then be updated at least every 31 EFPD* based on actual in-core flux map measurements, and a safety evaluation performed with the core physics model. The model would utilize the "as-measured" core burnup distribution.
The concept of the "Power Distribution Limit Report" is
_very similar to that of the "Peaking Factor Limit Report" which is currently being used to define the Fxy LCO's for Salem Unit 2. The important difference is that the new Power Distribution Limit Report would be periodically updated based on in-core- flux map measurements and the actual core operating history.
The central theme of the new FLEX-SPECS is the on-going re-evaluation of the core safety margins during operation, using actual measurements, and operating history. The benefits of this approach would be two-fold. First, there would be a much more intimate awareness and understanding This is the same as the current flux map surveillance period.
Page 4.1-2
_J
NFU-0039 Revision 1 March 14, 1986
- of the actual core conditions and safety margins. Second-ly, there would be a significant reduction in the uncer-tainty in the safety evaluation conclusions, due to the increased knowledge of the periodic core measurements.
This reduction in uncertainty would be converted via FLEX-SPECS into increased operating flexibility. This would be in the form of wid~r AFO limits and greater quadrant power tilt limits.
Section 4.1.1 defines the power distribution parameters of interest *
- Section 4.1.2 describes the bounding initial conditions considered in the power distribution analyses.
Section 4.1.3 describes the analytical procedures used to perform the RSE. These pr~cedures are divided into two parts: Pre-Startup, and Post-Startup.
4.1.1 Parameter Definitions The core power distribution parameters of interest are FQ and FoH* _These parameters are not directly observable. Instead, AFO and Tilt are monitored and maintained within limits which have been derived to assure FQ and F0H remain
- within their limits.
Page 4.1-3
NFU-0039 Revision 1 March 14 , 1986 FQ<Z> = Heat Flux Hot Channel Factor F~H = Nuclear Enthalphy Hot Channel Factor A.O. =
Pr = power in the top of the core P8 = power in the bottom of the core AFD = Axial Flux Difference= A.O.*Prel Prel = fractional core power
- PDIL = Power Dependent <Control Rod)
Insertion Limit LCO = Limit Condition for Operation Page 4.1-4 L
NFU-0039 Revision 1 March 14, 1986
- 4.1.2 Bounding Initial Conditions The bounding initial conditions for use in the analysis of the core peaking factors are those associated with the range of permissible values for the key core parameters, and the types of operational scenarios to be co~sidered. These considerations include:
- core power levels
- control rod position
- axial flux difference
- quadrant power tilt
- core depletion scenarios
- care maneuvering scenarios Core Power Level The Technical Specifications far FDH and FQCZ) are
. applicable ta MODE 1 operation. Therefore, RSE analysis of these parameters considers the range of power operation associated with MODE 1 *
- Page 4.1-5
NFU-0039 Revision 1 March 14, 1986 l
Control Rod Position
- The Power Dependent Insertion Limits <LCO> are assumed for the individual rod bank positions. In addition, control rod misalignment from the bank demand position i~ also considered for up to three individual rod control cluster assemblies CRCCA's). The probability that more than three RCCA's will simultaneously be misaligned is *considered to be negligibly small.
Axial Flux Difference The permissible boundaries for AFD are defined as an output from the RSE analysis, and are reported in the Power Distribution Limit Report (Appendix A).
Quadrant Power Tilt The quadrant power tilt ratio is assumed to be 3%
larger than the design value calculated by the core physics model.
- I Page 4.1-6
NFU-0039 Revision 1 March 14, 1986 Core Depletion Scenarios The choice of depletion scenarios to. be considered has an effect on th~ core burnup distribution, which in turn affects the flux distribution. The depletion scenarios considered are:
+ Full Power, unrodded
+ Load Follow
+ Full Power~ rodded
+ Reduced Power, rodded
+ Reduced Power, unrodded The full power. unrodded scenario represents the cu~rent operation. The load follow scenarios include considerations for maximum rod insertion and axial xenon oscillations. The rodded and reduced power scenarios consider the effects of extended* operation with skewed flux distributions ahd subsequent return to full power and/or unrodded operation.
The objective of these depletion scenarios is:
+ to evaluate core peaking factors associated with long term anticipated and postulated power
- operation, and Page 4.1-7
NFU-0039 Revision 1 March 14, 1986 l
+*to create burnup distribution files for use as input to the analysis of short term core maneuvering scenarios, and other care parameters such as shutdown margin and scram reactivity.
Care Maneuvering Scenarios The analysis of the maneuvering scenarios utilizes the various burnup distributions created from the core depletion scenarios. The maneuvering scenarios induce short-term changes in the flux distributions due to variations in care power, rad position, and xenon inventory. The maneuvering scenarios include:
+ power and rad variations
+ Xenon oscillations The power and rod variations consider power maneuvers from 25% to ful 1 power, in which the control rods are used for reactivity and/or AFD control. Both equilibrium and no xenon conditions are considered.
The xenon oscillations are induced at full power with the maximum allowable control rod motion in order to maximize the flux peaking factors.
Page 4.1-8
NFU-0039 Revision 1 March 14, 1986 4.1~3 Analvtical Procedures This section is subdivided into Pre-Startup and Post~
Startup RSE- procedures. As described in Section 4.1, the PSE&G RSE procedure represents a significant improvement over the traditional approach. The essence of this improvement is the periodic re-evaluation of the AFD LCO during operation, based on actual core measurements.
Section 4.1.3.1 describes the Pre-Startup RSE proce-dures. These are very similar to traditional RSE procedures.
Section 4.1.3.2 describe$ the Post-Startup RSE proce-dures. These are similar to the Pre-Startup proce-dures, except that they incorporate the actual core operating history and in-core flux map measure~ents into the analysis. The Post-Startup analysis is performed at least every 31 EFPO, and results in periodic updates to the AFO LCO in the Power Distribution Limit Report.
The implementation of these procedures will require the modifications to the Tech Specs described in Appendix
- c.
Page 4.1-9
NFU-0039 Revision 1 March 14, 1986 4.1.3.1 Pre-Startup RSE Procedures The core physics model is used to evaluate the power peaking factors FDH' and FQ under the bounding conditions described in Section 4.1.2.
This analysis is performed prior to the initial reactor startup following a refueling outage.
The results define the acceptable limits for axial flux difference <AFD) for the first 40 EFPD of cycle operation. These limits are reported in the Power Distribution Limit Report, described in Appendix D. During opera-tion, these limits are periodically updated based on the actual core operating history and in-core flux-map measurements as described in Section 4.1.3.2.
The Pre-Startup RSE analysis is performed in two parts, the core depletion calculations, and then the core maneuvering calculations. Of the several bounding core depletion scenarios described in Section 4.1.2, only the first--the full power, unrodded scenario, is analyzed over the entire cycle length with the Pre-Startup procedures. The others are analyzed for the first 40 EFPD of cycle operation.
Page 4.1-10 The effect
NFU-0039 Revision 1 March 14, 1986 of further depletion is left for periodic re-evaluation using the Post-Startup RSE procedures described in Section 4.1.3.2 below.
The load follow depletion assumes a typical
.load follow maneuver over the first 40 EFPD of cycle operation. The operating principle used for the load follow analysis is tci use control rods for reacti~ity compensation, and let AFD vary unrestrained. This minimizes the plant boration/dilution requirements, and provides the maximum practical flux redistributions due to rod insertion and axial Xenon oscillations.
The larger flux redistributions maximize the effects on the burnup distribution, which in turn affect the core peaking factors.
Usually, the AFD variations during load follow maneuvers are not sufficient to create exce-ssive peaking factors. However, should the peaking factors exceed the allowable limits at any point in the depletion, the AFD limit is appropriately redefined, and the analysis repeated in which the control rod motion is appropriately restricted to maintain AFD within the revised limits. The revised AFD limits are Page 4.1-11
NFU-0039 Revision 1 March 14, 1986 then used with the results of other analyses to define the final AFD limits described in the Power Distribution Limit Report.
The last three depleti,on scenarios listed in Section 4.1.2 are not ant*icipated under normal operation, but represent potentially signifi-cant effects on the core peaking factors. They consider the effects of extended operation under reduced power and/or rodded conditions, w~ich can create significant changes in the burnup distribution. Upon subsequent return to normal operating conditions, the change in the burnup distribution may have a significant effect on the core peaking factors.
Each of thes~ last three scenarios considers the effects of forty full power days of opera-tion under the specified reduced power and/or rodded conditions. As is the case with the load follow scenario, if the peaking factors exceed the allowable limits at any point in the depletion, the AFD limit is appropriately revised, and the analysis repeated in which the control rod motion is appropriately restricted i
Page 4.1-12 L _J i
I
NFU-0039 Revision 1 March 14, 1986 to maintain AFD Just within the revised limits.
The revised AFD limits are then used with the results of* other analyses to .define the final AFD limits described in the Power Distribution Limit Report.
The full power rodded depletion scenario considers the effects of forty EFPD days of operation at full power, with the rods inserted to the PDIL, followed by return to unrodded operation. This type of operation can have significant effects on both the axial and radial burnup and flux distributions. The limiting values of the peaking factor margin and associated AFD results are used in the definition of cycle AFD limits reported in the Power Distribution Limit Report.
The reduced power, rodded depletion scenario I
considers the effects of forty. EFPD days of
) operation at 50% power with the control rods inserted to the PDIL, frrllowed by a return to full power .unrodded operation. The unrodded, reduced power operation is similar, except that the rods are withdrawn during the extended reduced power operation. The results of these Page 4.1-13
NFU-0039 Revision 1 March 14, 1986 last two depletion scenarios can be significant changes in the full power, unrodded AFD, and associated peaking factors. The limiting values of the peaking factor margin and associ-ated AFD values are used in the definition of cycl~ AFD limits reported in the Power Distri-bution Limit Report.
In addition to the analysis of the various depletion scenarios described above, the Pre-Startup RSE procedures include the analysis of
(
the bounding maneuvering scenarios described in Section 4.1.2. These calculations are per-formed over the entire cycle using the burnup distributions derived from the full power, unrodded depletion scenario described above.
The analysis of the maneuvering scenarios is also performed using the burnup distributions derived from the other depletion scenarios over
/ the first 40 EFPD of c~cle operation. The potential effects at later burnups is left for the periodic re-evaluation using the Post-Startup RSE procedures .described in Section Page 4.1-14
NFU-0039 Revision 1 March 14, 1986 The peaking fact~r margin determined from the above depletion and maneuvering scenarios are computed using the following conservatisms and constraints:
+ The calculated values of FNQ<Z> and FNDH are performed full core, in three dimensions using the core physics model.
+ The calculated peaking factors are appropri-ately adjusted for model biases and conser-vatively increase~ by the model reliability factors.
+ The calculated peaking factors ar~ furth~~
increased by 3% to allow for a quadrant power tilt increase.
+ The minimum peaking factor margin is then determined by comparison to the power dependent Limiting Conditions for Operation as described in the current Technical Specification *
- Page 4.1-15
NFU-0039 Revision 1 March 14, 1986 4.1.3.2 Post-Startup RSE Procedures The Post-Startup RSE procedures utilize core measurements, and the core physics model, to re-evaluate and redefine the AFO LCO on a monthly basis during operation. The procedure consists of three components:
+ Core Follow
+ Transient Peaking Factor Analysis
+ AFD Limit Evaluation Core Follow The PSE&G core design engineers follow the operation of the reactor on a daily basis. The objective is to deplete. the core physics model in a manner that will accurately represent the actual th~ee dimensional core burnup distri-bution. Approximately hourly records are main-tained of core power level, control rod posi-tion, and AFO. For the purpose of depleting the core physics model, these data are reduced to sets of average operating periods. For each Page 4.1-16
- NFU-0039 Revision 1 *.i March 14, 1986
- period, average values of power level, control rod position, and AFD are determined. The length of each period depends on the magnitude of variation of the core parameters. During transient conditions, periods of 1 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> are common. During steady state operation, the period can be as long as several weeks.
The power distribution of the core physics model is periodically calibrated to in-core flux map measurements. This calibration is accomplished by comparing the predicted and measured in-core detector signals. A three dimensional set of calibration factors are derived from this comparison. The calibrated model is used to calculate the burnup distributions during the core follow operation.
These burnup distributions are refarred to as the "measured" distributions.
It is noted that the model calibration procedure described above is based on the results of a research and development project sponsored by the Electric Power Research Instit~te. This project is the PWR Power Shape
- Monitoring System <PSMS).
Page 4.1-17 Its objective has
NFU-0039 Revision 1 March 14, 1986 bee~ the development of a PWR technology for an on-line core power distribution monitoring system. A similar BWR-PSMS project has already been completed, and the first implementation has been scheduled at Oyster Creek.
Although PSE&G is not implementing an on-line system at this iime, the methods for mod~l calibration to be used are a direct result of the PWR-PSMS research project.
Transient Peaking Factor Analysis The core follow effort described above yields a measured burnup distribution and a calibrated core physics model. These are used to perform a peaking factor analysis for transient and normal operating conditions as described below.
This analysis evaluates the FQ and FoH margin for the next 40 EFPD of operation.
At a given time in core life, there exists a well behaved relationship between the variation in the Axial Flux Difference <AFO) and the variation of the maximum values of FQ(Z). A typical example of this relationship is shown Page 4.1-18
NFU-0039 I
Revision 1 , I March 14, 1986
- in Figure 4.1-1, in which the fractional change I
in FQ<Z> is platted as a function of the change in AFO. The changes in FQ<Z> and AFD are relative to a f u 11 power, unrodded, equilibrium xenon reference condition. In this figure, the changes in FQCZ> are plotted for an upper care level, Z=49, and a lower core level, Z=6 <Z=61 for core top,* Z=l for core bottom, see Figure 3.2-2>. Jhe _pl at i1 l ustrate.s t-hat the -frac-tional change in FQ<Z>, at a given level Z, is a linear fun~tion of the change in AFO. This is generally true only when the variation in AFD is due to changes in axial xenon redistri-bution alone, with all other core parameters maintained constant. Other parameters include core power level, control rod bank position, and xenon defect.
The slope of this .linear relationship for full power, unrodded, full xenon defect conditions has been given the name T<Z>;
T <Z> = il F Q <Z ) /F Q <Z) CEqua. 4.1-1)
AFD When the effects of simultaneous changes in other parameters are included, the change in FQ can depart from the linear relationship, T<Z>.
Page 4.1-19
NFU-0039 Revision 1 March 14, 1986 These departures take the form of a scatter band about T<Z> as shown in Figure 4.1-2. This band can be bounded by a line parallel to T<Z>,
but displaced a distance DF<Z> in the positive FQ direction. Typical values of T<Z>, and DF<Z> are shown in Figure 4.1-3. The T<Z> and DF<Z> parameters are computed with the core physics model, using the Post-Startup RSE procedures and the measured burnup distri-butions. The analytical procedures are similar to the analysis scenar*ios evaluated in the Pre-Startup RSE procedures.
The T<Z> analysis is performed for a core burnup very near the burnup at which the reference flux map is to be taken. A full power axial xenon oscillation is .simulated.
This oscillation is induced with control rods and then allowed to progress with the rods withdrawn. The maximum values of ~FQ<Z>IFQ<Z>
are then plotted as a function of the change in axial offset from the equilibrium condition as shown in Figure 4.1-1. These results are then fit to a homogeneous linear function, whose slope defines T<Z>.
Page 4.1-20
NFU-0039 Revision 1 March 14, 1986
- The DFCZ> analysis includes a series of model simulations in which the maximum values of
~FQ<Z>IFQ<Z> are determined for each of the bounding depletion and maneuvering scenarios, over the next 40 EFPD of operation. The analysis techniques are the same as for the Pre-Startup RSE procedures described in Section 4.1.3.1. These results are then plotted along with the TCZ> results as shown in Figu~e 4.1-2.
DF<Z> is then determined as the magnitude of the displacement of the parallel, bounding line
- from the T<Z>-line shown in Figure 4.1-2 at each axial level <Z>.
The DF<Z> values are further increased by 0.03 to allow for a 3% increase in the Quadrant Power Tilt Ratio. This* increas~ is in addition to the maximum tilt value measured by the reference flux map.
The T<Z>, and DF<Z> factors represent ratios of maximum FQ<Z> calculations. The application of model biases and reliability factors to each of these calculated peaking factors results in the cancellation of both the biases and the relia-bility factors. Therefore, the T<Z> and DF<Z>
Page 4.1-21
NFU-0039 Revision 1 March 14, 1986 factors remain independent of the model FQ uncertainty due to their pure differential nature. However, the TCZ)-factors are conservatively increased by 8% based on the results of Section 3.7, and the measured values of the reference FQ's are increased by 8.15% to account for measurement uncertainty.
The TCZ) and DF<Z> factors derived above represent the worst possible differential increase in FQ' which can be associated with a given differential change in AO from the full power, unrodded, equilibrium xenon conditions.
This bounding relationship is representative of the actual core operating conditions, and remains valid for at least 40 EFPD of addi-tional burnup beyond the reference flux map measurement.
The FoH margin is evaluated for each of the T<Z> and DF<Z> analyses described above. The maximum values of FoH are increased by the model reliability factor, and further increased by 3% to allow for the maximum allowed increase in the Quadrant Power Tilt Ratio.
Page 4.1-22
. I
~
NFU-0039 Revision 1 March 14, 1986 FIGURE 4.1.1 RELATIONSHIP BETWEEN F0 CZ> AND A.O.
FOR FULL POWER, UNRODDED CONDITIONS .. i I I
- *'"*rr-
- ~--l-
- Page 4.1-23
NFU-0039 Revision 1 March 14, 1986 FIGURE 4.1.2 THE GENERAL RELATIONSHIP BETWEEN FQ<Z> AND AXIAL OFFSET
:~-* ~
~~~.
- ____-T,_ .. L=:: :_-__
- _~,,,,..
,,£__
~=*
- *-/'.
~ -:~J.!
- ~:-
- -t.-t--*
f-
_JC_*
--~-
- ir**
-+
+-*
Page 4.1-24
NFU-0039 Revision 1 March 14, 1986
- FIGURE 4.1.3 TYPICAL VALUES OF T<Z> AND DFCZ)
~~::~ i~~ ~~i ~ ;;~ .::.:~t~~~ -~~~f~~ --=b ==~~--~~ =~~~:~~~::~*E~:= ~~::=~~=~~ ~:_:~E~ ~;_:;;~~~ ~~~: ~~~~-~~-:: ~~~~~~=
- =--~~::;.: ~-;~~=--==~~ ~=-~=:~-~~~-: ~;-~:-:~. ~~~£=~~=.::-:.~--.:t=:t=-:-r=::t- ~ -- ---~ e:. z.:-:___ ..7 :-::~-~=~=~~t::":- --~~-~:~*::i~.~~~
~ :. :i.~:* ~-~f~:~-.i<f-~:.:~:~.: ~:: ~~~~:it~L~~ ~~:t~:.:.: :~:~.:j-~~~~ ~~j~~r~~i~-~~if~ ~~~~~ ~~ ~=--~,-~*~S:~~ -~~i:~i:~~~t~~;: -~:*~~-~:~: *:~:~~~:-:.:_
- ~~r~
- =-:::c:i:=..,~:-:~:f=:_:._: ; :~:~~=-~ ~:f::~* 0 =t~-::0~~~~~:=l~=-=-=r.:=:-qc:=EE=:f:~:-::~=:==-:~_:=t-:~~~~: 1 :i:~~:_:::
~:-::::: r=::::'ti ~:~.:;:1:--r- 1~=-~:~=:j'~~~~:~== ~~:;,_::r~~~~:X-=:: ==:~~£-= ~F-= =-::~~-:~ ;~:~~~~~i~~~ :::-=~=~L-=-==~~::
- ~::::,:-:::;:sf~:~::~: : ;:: J:::j~~:=J'-::-=::;.~ ~:~~:-t~~~==: ~=~~'J:~:-=.:==¥-=-=~= =~:§§~r~~~= =~: :=~):-:: r~~=,:,:c: -::=-~-= :_=
L:: : : - ~,\E~=::C::-1==:: - : ::::;_; !::=:: =~y~.::;:,:-f~::::'.iJ~=-=~:-::f~~~: ::-=:=~~_:tc.=_==.--: -=--=::=:---~~~:-:.:_::-~~~=:c _::=:=~~:, =-=:: =-==~~: ~~=:~~~
--c:_-=-1=:~:= ::-~:_:c: =:=~ = :_::::: L*:~::-: i:-:-=:~:-:>77:.~r~ ::~*fX:=,_=-==== ~::-:-: -:-'.3-~-::Cj::~~=i:.:=' -~~::~:-: ==:=~:;:*;:f::~:~c::-rc.~~=-~
t"-:'!----r,,'-_..-.J_:-_,_1___-'-*--+!-~:- '°- :-=*= ! : !-~~:::==T:=:-?..,~:i:_Jc::~~=~~~~:~f~2~'-:::~:iiE=:~:::::_: it=::::;~~: ,=:-:-=:~"::i 0
~ ~:--=-* '. ::~-:;::::_- ::-=:f:=~\Cc'-~:.::'-~;~i:*:; ;=-'"1::~.:::~~~~~ ::==-~~:_::: :::.:c::*:f: :__:::: '.'.::::~= -
f
_ . . . ,- I .:.: - -,--*::
1
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! ,- - *-- - -. .:-! -,,~:: ~L==:.=-:-:j:.J::___ :: ::-}?.:'.:-::~t==~~::~~=:::-:.:*::3=='"===-==: .;:;~=f~=:::
~[~_-:====-'..______~...Q2:.._..:~*i._-_ _-~!___-~t~_-_--_-;.._ _;_-=- +i-_-'- -_-~!~-=~:~=-~=- ~=-~/~:*_: -=-~?- =-;~=-~=-~ +ti~=-~=-~ =-:-=-:*:=-;~=-f-=~=-*:~=- ~=~;= -~ =- f.=-:*:=-*~:=-*~=-: =-~:~_=-~ =-: :_ =~=~ =- ~ ~:~=-~ :_~:_:.=-~ =- ~r=-~-:=-= -= -=~=-=~ -=-: =: =:~=-~=-~ -
~-*--*-:---:-*--:--:---:---:--:---:-:---:-:---~i~~:~-,,---,,----;-'---~*--_-7!~~:*~~~~~*-:~~:~+!"_::~~*:.:_*~-~-=+-}~_:~-~-~-:~_*.~~-[_::_.:~_~?-_~~-~~~~E_~==-*~-~:;_~~-~~-t-~~-*~-=-*:~_-=--~;:_~~-:-~~-:~~=~~~f=-~=-~:~~:*~=-*;~~~~-~~~~~*~~;~~-~:.:._=i:
r --L* *t:: =~~.: >:: --~ f ~: * ~ -:*:*: ~ _-<~~~A*~~:~=~~-~ f~: ~~~~ ::. i-: ~ *_: ~ :~: ::t:-~~-;:~~:F:-:*:=-:*~-= r~:~~ ~=~~b: :.:~: :.-:
"*(.U* -**r I*--*********-*-
f-----c---- .-----*** . *- -:-:=:::::_::t=-*::. -**
- I .*r Page 4.1-25
NFU-0039 Revision 1 March 14, 1986 AFD Limit. Analysis
- At a given point in core burnup, TCZ) and DFCZ) parameters can be combined with the results of a flux map measurement to provide a conservative estimate of the relationship of FQ<Z> and A.O.
This relationship can be expressed as follows:
FQ<Z , AO) ~ FQma P <Z) C1+T <Z)
- t:5. AO + OF <Z) J
<Equa. 4.1....:2>
where: FQ<Z,AO) is the maximum FQ at axial level Z, at any axial offset condition, AO.
FQmap and AOmap are the values measured with the in-core flux map.
!:5.AO = AO-AOmap T<Z),OFCZ) are calculated with the core physics mqdel, and based on actual operating history.
Page 4.1-26
NFU-0039 Revision 1 March 14, 1986 The above relationship can be used to evaluate the AO at which FQCZ,AO> will reach FQLCOcz>; the limiting value of FQ.
This value of AO is referred to as AoLCO.
Solving Equation 4.1-2 for AO, and then replacing AO and FQCZ,AO> with the limit-ing values, AoLCO and FQLCO<z>.
FLCO< Z> }
= AOmap +.l Q - DF<Z> -1 TCZ) { Fm~Pcz> .
CEqua. 4.1-2)
- The AO limits are then converted to AFD limits using the appropriate definitions.
A typical example of the AFO Limit Analysis is provided in Appendix E.
This AFD Limit Analysis is repeat~d at least every 31 EFPD based on the periodic in-core flux map measurements, and the
- actual core operating history. The results of each analysis are used to update the Power Distribution Limit Report, which is then used as the new AFD LCO which remains valid for the next 40
- EFPO.
Page* 4. 1-27
NFU-0039 Revision 1 March 14, 1986
- 4.2 REACTIVITY COEFFICIENTS 4.1.2 Parameter Definitions
= Moderator Temperature Coefficient The change in core reactivity associated with a 1°F change in average moderator temperature. Core power level, control rod position, RCS boron, and xenon are h e l d- co n s t an t
- However , f 1 u x an d t em pe r a--
ture distributions are permitted to vary as dictated by the core neutronics and
- thermal hydraulics.
O<P(P) = Power Coefficient The change in core reactivity assocjated with a 1% (of full power) change in core average power level. Rod position, RCS boronp and xenon are held constant. The average moderator temperature is varied according to the Tave vs. Power p~ogram.
However, flux and temperature distri-butions are permitted to vary as dictated by the core neutronics and thermal hydrau-1 i cs *
- Page 4.2-1
NFU-0039 Revision 1 March 14, 19_86
= Doppler Temperature Coefficient The change in core reactivity associated with a 1°F change in the average fuel temperature. The average moderator temperature, boron concentration, rod position and xenon inventory are held constant. However, flux and temperature distributions are permitted to vary as dictated by the core neutronics and thermal hydraulics.
= Boron Coefficient The change in core reactivity associated with a 1 ppm change in core average soluble boron concentration. The core power level, rod position, and xenon inventory are held constant. However, flux and temperature distributions are permitted to vary as dictated by the core neutronics and thermal hydraulics.
Page 4.2-2
- NFU-0039 Revision 1 March 14 , 1986
- 4.2.2 Bounding Initial Conditions The key core parameters which can have the largest effect on the reactivity coeffi-cients are: burnup, power level, and boron concentration. The Bounding initial conditions defined in Table 4.2.1 span the allowable range of these parameters for mode 1 and mode 2 operation, with keff =
1.0. Bounding initial condiiions for other conditions are discussed as appro-priate for the individual accidents in
- Section s.
Reactivity coefficients are evaluated for the bounding *initial conditions shown in Table 4.2.1 for both BOC and EOC condi-ti ans.
The only significant effect of the pres-ence of xenon* i.s to control the critical boron concentration. All cases except the last are evaluated with no xenon. At low power, this creates a conservatively high boron concentration. The lowest critical boron occurs at ful 1 power. The 1ast case
- Page 4 ."2-3
NFU-0039 Revision 1 March 14 , 1 986 at full power is therefore run at equilib-rium xenon conditions. This represents the lowest possible boron .concentration
- Page 4.2-4
NFU-0039 Revision 1 March 14, 1986
- TABLE 4.2.1 BOUNDING INITIAL CONDITIONS FOR EVALUATION OF REACTIVITY COEFFICIENTS
~M' ~P ' ~D' ~B Core Control Rod Xenon Power Configuration 0 PDIL @ 0% none 0 ARO none 50% PDIL @ 50% none 50% ARO none 100% PDIL @ 100% none 100% ARO none 100% ARO equil
- Page 4.2-5
NFU-0039 Revision 1 March 14, 1986 4.2.3 Analytical Procedure The reactivity coefficients defined in Section 4.2.1 are evaluated at the core conditions described in Section 4.2.2.
The analytical procedure is straight-forward, and consists of sets of model calculations in which the core parameters are varied as prescribed in the parameter definitions. All calculations are per-formed in three dimensions with the nodal model
- Page 4.2-6
NFU-0039 Revision 1 March 14, 1986 4.3- SHUTDOWN MARGIN 4.3.1 Parameter Definitions SOM: Shutdown Margin The instantaneous amount of reactivity by which the reactor is subcritical, or would b~ subcritical from its present condition
- assuming all full length rod cluster assemblies <RCA) <shutdown and control) are fully inserted except for the single . ...\
rod cluster assembly of highest reactivity worth which is assumed to be fully with-drawn.
Stuck Rod The rod cluster assembly <RCA) remaining fully withdrawn in the SOM definition
<N-1> Rodworth The amount of negative reactivity change occurring when all* rods, except the stuck rod, move from the fully withdrawn to the fully inserted position *
- Page 4.3-1
--~--------------------------------------------
NFU-0039 Revision 1 March 14, 1986 POIL: Power Dependent Insertion Limits These are limits on the allowed rod insertion for modes 1 and 2. These limits are given in the Technical Specifications.
4.3.2 Bounding Initial Conditions Shut down margin is affected primarily *by core burnup, and choice of the stuck rod.
The burnup effects the rodworth, Doppler defect, and critical boron concentration.
The boron affects the moderator defect.
For mode 1 and 2 operationst the most conservative initial conditions appropri-ate for .SOM evaluation are summarized in Table 4.3.1. These condition are evalua-ted at BOC and EOC. The burnup distribu-tions considered are those derived from the core depletion scenarios described in Section 4.1. The initial power distri-butions considered are those derived from the core maneuvering scenarios described in Section 4.1.
Page 4.3-2
NFU-0039 Revision 1 March 14 , 1986 Since the boron concentration is adjusted to maintain keff = 1.0 for the pretrip conditions, the variations in the xenon inventory provide critical boron varia-tions which have a direct effect on the moderator temperature defect *
- Page 4.3-3
NFU-0039 Revision 1 March 14, 1986 TABLE 4.3.1 BOUNDING INITIAL CONDITIONS FOR EVALUATION OF SHUTDOWN MARGIN Care Control Rod Xenan(l)
Power Configuration Condition 50% PDIL @ 50% none 50% PDIL @ 50% f u 11 power Equil.
100% PDIL @. 100% none 100%
( 1)
PDIL @ 100% f u 1 l power Equi 1.
Baran concentration adjusted ta maintain keff = 1.0 far the pretrip condition Page 4.3-4
NFU-0039 Revision 1 March 14, 1986 4.3.3 Analytical Procedure Calculations of shutdown margin are performed in full core, in three dimen-sions, with the nodal code. The general calculational sequence is straight for-ward, consisting of 3 cases for each condition evaluated. The general case sequence is shown in Table 4.3.2.
As seen from Table 4.3.2, the best esti-mate of shutdown margin is obtained directly from cases 1 and 3. Case 2 is included to provide the N-1 rod worth
<Cases 2 and 3) which is used to compute the uncertainty allowance. This allowance is subtracted from the best estimate value of SOM to yield the final SOM used to compare to the FSAR as discussed in Section s.
The most worthy stuck rod is evaluated by calculating the stuck rod worth for several N-1 core configurations. These calculations are performed in full core Page 4.3-5
NFU-0039 Revision 1 March 14, 1986 with the nodal model. In each cal cu-lation, one rod cluster assembly from a symmetric rod group is withdrawn from the fully rodded condition. Similar calcula-tions are then repeated for each symmetric rod group in the core. The RCCA with the highest worth is then used as the stuck rad. The stuck rod evaluation is per-formed at BOC and EOC, and may result in the identification of different RCCA's for each burnup condition.
The assumptions and conservatisms inherent in the above SOM analytical procedure are summarized as follows:
A. The control rods are assumed to be at the deepest allowable insertion prior to trip <PDIL).
- 8. The most worthy single RCCA is assumed to remain fully withdrawn.
- c. The identification of the most worthy RCCA can change with burnup.
Page 4.3-6
NFU-0039 Revision 1 l
! I March 14, 1986 j I
, I
- o. The maximum change in moderator and fuel temperature is assumed to occur between the pretrip and post-trip cond it i ans.
E. No changes in RCS boron concentration, or xenon inventory are permitted.
Page 4.3-7
NFU-0039 Revision 1 March 14, 1986 TABLE 4.3.2 ANALYTICAL CASE SEQUENCE FOR EVALUATION OF SOM CASE CORE ROD TAVE TFUEL POWER CONFIGURATION 1 ( 1) p (4) PDIL @ pl T <P )( 2 ) T <p ) (2) 1 m 1 f 1 2 0% PDIL @ pl 547°F( 3 ) 547°F( 3 )
3 0% N-1 rods in 547°F< 3 > 547°F( 3 )
( 1)
Note Xenon and boron held constant at a value to maintain keff = 1. 0 for case #1 (2)
These are normal at power temperatures.
(3) These are normal zero power conditions.
(4) P is the core power for the bounding initial c6ndition in Table 4.3.1.
Page 4.3-8
l NFU-0039 *~ I Revision 1 '
March 14, 1986
- 4.4 SCRAM REACTIVITY 4.4.1 Parameter Definitions 6p <t>scRAM: Scram Reactivity The amount of negative reactivity inserted into the core, following a reactor trip, as a function of time after rod release, assuming the single most reactive rod remains in its pretrip position.
Bounding Initial Conditions The 6p <t>scRAM is affected primarily by core burnup, choice of the stuck rod, the initial control rod position, and axial flux shape.
For Mode 1 and 2 operation, the bounding initial conditions appropriate for 6pCt>scRAM evaluation are 0% and 100%
power, the most negative anticipated axial flux shape, and the rod falling from the fully withdrawn position. The initial burnup, and power distributions considered, are those derived in Section 4.1. These evaluations are performed at
Page 4.4-1
NFU-0039 Revision 1 March 14, 1986 The choice of power levels assures that the entire allowed operating power range is bounded. The negative axial flux shape decreases the amount of inserted rod worth in the early portion of the scram curve.
The first one second of the scram curve is the most important portion with respect to the effect on the various accidents. The most negative axial flux shape at the time of trip is that which can be induced with an allowable xenon condition. For full power, this would be the condition consistent with the most negative allowed axial offset.
The most conservative initial rod position is usually the all~rods-out *<ARO>
position. However, there are some situations in which the minimum scram reactivity during the first one second of rod travel results when the rods are initially at the power dependent insertion limits <PDIL). These conditions are illustrated in Figures 4.4.1 and 4.4.2.
Figure 4.4.1 shows a typical situation where the ARO initial rod position is more Page 4.4-2
NFU-0039 Revision 1 March 14, 1986
- limiting than the PDIL initial rod position. Figure 4.4.2 shows a typical situation where the PDIL initial rod position is more limiting.
Figure 4.4.2 also illustrates a typical situation where the cycle specific SCRAM curve crosses the FSAR curve for the later times after rod release. Such crossings of the FSAR curve at times later than one*
second after rod release are not signi-f icant because fast accidents are essen-tially over by this time and slow acci-dents are even less sensitive to SCRAM reactivity.
The analytical procedures consider initial conditions with rods withdrawn as well as at the PDIL *
- Page 4.4-3
NFU-0039 Revision 1 March 14, 1986 FIGURE 4.4.1 EFFECT OF INITIAL ROD POSITION ON SCRAM REACTIVITY 11 SALEM 1 CYCLE 7 BEGINNING OF CYCLE (HOT FULL POWER) j : I
- - - ARO, A.O. = 3.82--'-,~--~~~~~...;...-~~~-'------" i.
- o - ARO, A.O. = 11.5~.>-+i~-+-'----'-~---T---'---l-~_..:._a-...;_------l i l.
.......+--- ----- PDIL, A.O. = 9.2J..-;.~----,----;----,.---;-_;_.,~-+--'-rlr-_.;.--+ U/J 10 i i :f
!
- fi.
i ; i i 1i I !* :
I If ! !I I I
9 :.* I I .I ilh i. I
- .. i I
rr 1 I;;.: :* ! .,
I 8
- t. I
. ; f*. !
- t. .
r~-~ r i i. .f ** I* ~: * ::: : !. , i 7 I i i i i ~:T I*
i ! : I !
I I I i I
6 I I *~ ! ! i I I
! ! ! i l ' I i
. ,* ... J_:
!.. ;~ _,.,~~""'". ' -.-::K '\.f'\. '-:i\..K:;\;.:i\.~~~::: ;: .!...
0
_;... -~r-"L--~--~. :.... "- "(i"\. ":"'-"'-i'\. '\J'( \..]\.. \. ':'-N:'f ** .- [::
0 0.4 0.8 1. 2 1. 6 2.0 2.4 2.8 TIME AFTER ROD RELEASE, SECONDS Page ~ ~* 4-4
- NFU-0039 Revision 1 March 14, 1986 FIGURE 4.4-2 EFFECT OF INITIAL ROD POSITION ON SCRAM REACTIVITY SALEM 1 CYCLE 7 END OF CYCLE
- L~: / t~::~~"F-=: :~"~~- 'HP'J;: ZE~O! POWE~) I*
/I*
i I !* I
~'-*-'~*-*~~ARO, A.O.= 72.0 I *! ;
-*.,....*--;-- ---- PDIL, A. 0. = 5 4. 0 f i f i 10~*-**~!-+-.:..,.,.,...,,........,...,,...,..,...,.,.,,.,?:"'.'Tl~~c:-r..,.,..,..-r.....,.--;--:-+---!-+-~ 1 J'.
-'---Hr.--t--i"-~,---j----:----t L !
- -*l .* i': * .,. : ; :_,y~7 : ':L' '::'t::_ =** ~
0 .. l ' I
- r. i i I j'. i i I
j:.
i i ~-
9
! ::*. '
- L_::A:_::::+:: ::.:~:- . ; :. !: Ki:"\0.i .. -~
T:.:' T * '.:: [" :: i'-' .*. :: -*. ** ,, . *., * ::: '-.f...-'j... f). ~": .** .-- f 1
8 ._,;__1-..:---i.~*-*-1-~-*~**b:~'=r~c:~-*~,:~'f'_-+-~t.~*~-~-*~*~-*~t.-'~* ~*"'"!--~*~--~-~**:.-~*~._~:-..~-~*~ c;:-*~~::1_*~.. *~*:--l1 1:*.
- i:::: ::';f-~°' ~:~:r~:::: ,:::~., : L --:T. *** .. ,_, ,::.lJ:. >J "'-~~~r=:o' ,,,.:, I
~~::-~= =** ;~~~-~T~~=--E - -~~-~ -}.:=~~;~~~=~ -:~~~f:~: _:_: :-~.=: ~:f~;~~=-:i::J:==c~ -- - ..--:--=~::.~.:=:
- c;. >t~='>::-:.:g: ::.:.-1~:: :: :- .. :. . ::; .. :=.t:<'F.!~f: ~:'.'!s;~~~f,,::-=:f'*
.... : .. :1..
. , .I
.: i'
- I.
II . :. r-..!1'-,,;'-.~1-~: '°'
i'\. :----~--~---
4 r*
3 '
- * ':. t:::: ~=~t::::E'V.:=_J .. ~= [; * ** !- i * ~~°".J"°'~*~)xf ,: * *
-, r: .,. . :::: :: :~_:(:~~..::E=:c;EE~: :;:;J: ... t ; !. Af.\.;'\.l"J~l~'S~"l * =; Y 2 1* : .. , .. i_:* A~:~=1~: :-::l: ::~' . f:. !. .* iAJ"-:"\,.KT'\~~:\KL:* *' 1:
. I ! . :. v ~j:f=;.i :::~£=--= ~y;;: :~:y ... ! .. ::* : :* : L'\;~f"-KX1~h"Sl?\L: :~::i.:
0 0.4 0.8 1. 2 1. 6 2.0 2.4 2.8 TIME AFTER ROD RELEASE, SECONDS Page 4.4-5
--- --~---
NFU-0039 Revision 1 March 14, 1986 4,4.2 Analytical Procedures The ~p(t)SCRAM analysis considers the limiting initial power and b~rnup distributions derived in Section 4.1. For each condition analyzed, the care thermal distribution is held constant at the initial condition, but the flux distribution is allowed to vary in response to the rod insertion.
This assures the slowest possible initial reac-tivity rate.
The calculations are performed in full core, in three dimensions, using the core physics model.
In each analysis, the change in core reactivity is first calculated as a function of rod position, as the N-1 rods move from full out to full in. This is then converted to time using the position versus time relationship assumed in the most recent safety analysis.
Page 4.4-6
NFU-0039 Revision 1 March 14, 1986 4.5 EFFECTIVE DELAYED NEUTRON FRACTION 4.5.1 Parameter Definitions fii =delayed neutron fraction for ith neutron energy group.
fi = rni sum over i-groups I = The importance factor, which accounts for the effects of reduced fast fissioning, increased resonance escape, and decreased fast leakage by the decayed neutrons. I = 0.97 4.5.2 Bounding Initial Conditions The bounding conditions are associated with core burnup and power distribution.
j The core burnup affects the local fission sharing and fii-values. The core power distribution affects the core average neff through the flux weighting effects. As an example, then tend to decrease as core Page 4.5-1
NFU-0039 Revision 1 burnup increases.
March 14, 1986 The core average B therefore increases as the axial flux shape skews towards either the top or bottom of the core, where the fuel burnup is lower.
Generally speaking, however, the burnup effects are much larger than the flux redistribution effects.
The bounding initial conditions for evaluation of Beff are summarized in Table 4.5.1. These calculations are performed at both BOC and EOC.
)
Page 4.5-2
NFU-0039 Revision 1 March 14, 1986
- TABLE 4.5.1 BOUNDING INITIAL CONDITIONS FOR EVALUATION OF fiEFF Power Rods Xenon
<steps) Condition 100 228 Equilibrium
- 0 228 None I
I .
- Page 4.5-3
NFU-0039 Revision 1 March 14, 1986 4.5.3 Analytical Procedures For each evaluation, the appropriate power and burnup distributions are obtained'from the depletion and maneuvering scenarios described in Section 4.1.
The local fission sharing by isotope and fuel type is determined from two-dimen-sional PDQ calculations as a function of local burnup.
Nodal values of fii are then determined by weighting the local delayed neutron fractions from each fissile isotope by the local fission sharing associated with that fuel type and nodal burnup.
The core average fi is determined as the nodal power weighted average of the nodal fi's. The importance factor I=0.97 is then applied to obtain the core fieff for the core condition under investigation.
Page 4.5-4
NFU-0039 Revision 1 March 14, 1986
- 4.6 PROMPT NEUTRON LIFETIME 4.6.1 Parameter Definitions
. t* = the average time from the emission of a prompt neutron in f issian ta the absorption of the neutron somewhere in the reactor.
Ia = care average, one group macroscopic absorption cross section *
- V = average neutron speed.
.[
4.6.2 Bounding Initial Conditions
~* is primarily sensitive to core burnup.
Calculations are therefore performed at BOC and EOC for the unrodded, full-power depletion scenario.described in -
Section 4.1 *
- Page 4.6-1
NFU-0039 Revision 1 March 14 , 1986 4.6.3 Analytical Procedures For a given core burnup, i
- is calculated as:
where the ratio of keff/koo is taken from the nodal model, and Ia is obtained from the two dimensional PDQ.
Page 4.6-2
NFU-0039 Revision 1 March 14, 1986
- 5.0 SAFETY EVALUATION METHODS This section addresses the evaluation of the cycle specific physics parameters with respect to the bound-ing 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 reanalysis is required. For each accident or transient the following material is described:
- Definition of Accident - a brief description of the causes and consequences.
- Accident Specific Methods - a description of the physics methods whi~h are unique to each acci-dent. General physics methods are described in Section 4.
+ Reload Safety Evaluation - a description of the comparisons of the cycle specific physics charac-teristics and the bounding values used in the safety analysis. Specific applications of the model reliability factors and biases are also addressed *
- Page 5.0-1
NFU-0039 Revision 1 March 14, 1986
- 5.1 UNCONTROLLED ROD BANK WITHDRAWAL FROM A SUBCRITICAL CONDITION s.1.1 Description of the Accident An uncontrolled addition of reactivity due to uncontrolled withdrawal of a Rod Bank results in a power excursion. The nuclear power response is characterized by a very fast rise terminated by the reactivity effect of the negative Doppler temperature coefficient. After the initial energy release, the reactor power is reduced by this inherent feedback, and the accident is terminated by a reactor trip. Due to the small amount of energy released to the core coolant, pressure and temperature excursions are minimal during this acci-dent.
s.1.2 Accident Specific Methods All of the cycle specific physics para-meters of importance to this accident are calculated with the General Physics Methods described in Section 4.0, except
- the Maximum Reactivity Insertion Rate.
Page S.1-1
NFU-0039 Revision 1 March 14, 1986 Maximum Reactivity Insertion Rate, ~p/~t In order to compare with the reactivity insertion rate assumed by the safety analysis for -uncontrolled rod withdrawal faults, 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 banks moving in 100% overlap. During normal operation, the trip breakers that supply power to the rod drive mechanisms are open until entry into Mode 2 ~eactor condi-tions. By current.Technical Specifica-tions, the rod insertion limit assures that at most only bank D can be fully inserted, bank C can be about mid-core, and bank B nearly withdrawn. For censer-vatism, and also to allow for potential future changes in the Technical Specifi-cations, the calculation of the maximum reactivity insertion rate will be per-formed with the assumption that control Page S.1-2
NFU-0039 Revision 1 March 14, 1986 banks O, C, and Bare fully inserted, and all other banks are withdrawn. Therefore, the choice of the two banks of maximum worth banks must .be made from these three banks.
The calculation is performed full core, in three dimensions with the core physics model. The two banks chosen are withdrawn in 100% overlap with the third bank either fully withdrawn, or fully inserted, whichever maximizes the reactivity insertion rate.
Calculations are performed at BOC and EOC at zero power conditions *
- Page S.1-3
NFU-0039 Revision 1 March 14, 1986 s.1.3 Reload Safety Evaluations CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. 0<M + RFMTC i O<M (least negative bounding value) b * ()( 0 * ( 1 - RFOC) i 0<
0 (least negative bounding value)
- c. fieff * <1 - RFB) 2. fie ff (minimum)
- d. tip /tit * + RFROO) i t.p /ti t <bounding)
(1
- e. t.p <t) SCRAM*< 1-RFROO) 2. tip ( t) SCRAM
<bounding)
Page 5.1-4
- I
__J
NFU-0039 Revision 1 March 14. 1986 5.2 Uncontrolled Rod Bank Withdrawal at Power 5.2.1 Definition of Accident An uncontrolled Rod Bank withdrawal at power results in an 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. The fuel in the reactor core could I
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 the rate of reactivity insertion, the following trips serve to prevent fuel. damage or over-pressurization of the coolant system: high power range neutron flux, core coolant ~T, high pressurizer 1 evel, and high pressurizer pressure. For the more rapid rates of reactivity 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 circuitry and rod motion.
Page 5.2-1
NFU-0039 Revision 1 March 14, 1986 s.2.2 Accident Specific Methods All of the cycle specific physics parameters of importance to this accident are calculated with the General Methods desc~ibed in Section 2, except for the maximum reactivity inser-tion rate.
Maximum Reactivity Insertion Rate,6p/6 t Calculations similar to those described in Section s.1.2 are performed at the full power, equilibrium xenon condition at BOC and EOC. The initial rod position is assumed to be the full power rod insertion limit.
S.2.3 Reload Safety Evaluations Each of the physics paramete~s calculated above are adjusted to include the model reliability factors and biases. These adjusted values are the cycle specific parameters which are then compared to the bounding values assumed in the safety analy-sis. The cycle specific parameters are acceptable if the following inequal.ities are met:
Page 5.2-2
NFU-0039 Revision 1 March 14 , 198q CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. 6p/6t * <1 - RFRQO) i 6p/6t (bounding)
- b. 6p(t>SCRAM*<l-RFROD> 2. 6p <t>scRAM<bounding)
<Maximum reactivity feedback)
- c. ~D * (1 + RFOC) 2. ~D <most negative bounding value)
- d. ~M - RFMTC 2 ~M <most negative bounding value)
Minimum reactivity feed back)
- e. ~O * <1 - RFDC> i ~
0 <least negative bounding value)
- f. ~M + RFMTC ~ ~M <least negative bounding value)
- Page S.2-3
NFU-0039 Revision 1 March 14, 1986 5.3 Control Rod Misalignment Definition of Accident In the analysis of this accident, one or more rod cluster control assemblies is assumed to be statically misplaced from th~ nor~al 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 reduc-tion in DNBR and is subject to limita-tions specified in the .plant Technical Specifications.
Accident Specific Methods An RCCA misalignment is considered as part of the General Physics Methods for Mode 1 operation.
Page S.3-1
NFU-0039 Revision 1 March 14, 1986 Reload Safety Evaluations The FoH calculated above is conserva-tively adjusted to account for model reliability factors, RFFOH. Addition-ally, a further adjustment is made to account for the maximum initial quadrant tilt condition <T> allowed by the Technical Specifications. The resultant FoH is then compared to the value used in the safety analysis as follows:
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
<FoH+RFFDH><l+T) i FoH<Rod Misalignment)
Page 5.3-2
- NFU-0039 Revision 1 March 14, 1986 5.4 Dropped Control Rod 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 into a fully inserted position in the core.
A dropped rod cluster control assembly CRCCA> typically results in a reactor trip signal from the power range nega-tive neutron flux rate circuitry. The core power distribution is not adversely affected during the short interval prior to reactor trip. The drop of a single RCCA assembly may or may not result in a reactor trip. If the plant is brought to full power with an assembly fully inserted, a reduction in core thermal margins may result because of a possible increased hot channel peaking facto~.
Accident Specific Methods FoH is calculated under dropped rod conditions in full core, and in three Pag~ 5.4-1
NFU-0039 Revision 1 March 14, 1986 dimensions with the core physics model.
The initial rod position is either unrodded, or rods at the PDIL, whichever maximizes the FoH* An analysis is performed for all symmetrically unique, single dropped rods, at both BOC and EOC. Xenon transients following the dropped rod are considered in a manner consistent with the time and power reduction requirements for Quadrant Power Tilt Ratio and Rod Misalignment.
Reload Safety Evaluation The nuclear enthalpy rise factor F 0H calculated above is conservatively adjusted to account for calculational uncertainties. This is further in-creased to account for the Tec*hnical Specifications allowance for quadrant tilt <T). The resultant value is then compared to the FOH assumed in the safety analysis for the dropped rod to demonstrate conservation. The compari-son is as follows:
Page S.4-2 J
NFU-0039 Revision 1 March 14, 1986 CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
<F OH+RFFDH> <1 +T) FoH <dropped rod)
Page 5.4-3
NFU-0039 Revision 1 March 14, 1986
- s.s Uncontrolled Boron Dilution s.s.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 condi-tions. Dilution during refueling or startup is assumed to be recognized and terminated by operator intervention before loss of shutdown margin. At power, with the reactor in automatic control, the power and temperature increase from boron dilution results in the insertion of the RCC assemblies and a decrease in shutdown 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 reactor in manual control, the power and temperature rise due to boron dilution would eventually result in an overtemperature dT reactor trip if the operator d1d not Page S.S-1
NFU-0039 Revision 1 March 14, 1986 intervene; After such a trip, the operator is expected to isolate the unborated water source and initiate boration procedures.
5.5.2 Accident Specific Methods All of the physics parameters of importance to this accident are calcu-lated with the General Physics Methods described in Section 4.
5.5.3 Reload Safety Evaluation All the cycle specific parameters discussed above are adjusted to include model r~liability factors and biases.
These results are then compared to the bounding values assumed in the ~afety analysis. The cycle specific p~rame-ters are acceptable if the following inequalities are met:
Page S.5-2
NFU-0039 Revision 1 March 14, 1986
- CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS SOM <bounding)
SOM 2.
ocM + RFMTC i ocM <least negative bounding value) i oc (least negative oco
- Cl - RFOC> 0 bounding value)*
<FoH+RFFOH> Cl+T> i F~l:I <Technical S ecifications)
- Page S.S-3
NFU-0039 Revision 1 March 14, 1986
- 5.6 Feedwater System Malfunction 5.6.1 Definition of Accident Two classes of accidents are to be con-sidered under this classification: Those that result in a decrease in feedwater temperature and those that result in an increase in feedwater flow. Either con-dition 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 coef-ficients and/or control system action. For the case of a *decrease in feedwater tern-perature, the worst accident which may be postulated involves opening the bypass valve which diverts flow around the feed-water heaters. For the case of an increase in feedwater flow rate, the worst accident which may be postulated involves the full opening of a 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
- Page S.6-1
NFU-0039 Revision 1 March 14, 1986 5.6.2 Accident Specific Methods All of the physics parameters of importance to these accidents are calculated with the General Physics Methods described in Section 4.
5.6.3 Reload Safety Evaluation Each of the physics parameters calculated above are adjusted to include the model reliability factors and biases. These adjusted values are the cycle specific parameters which are then compared to the bounding values assumed in the same analysis. The cycle specific parameters are acceptable with. regard t*o feedwater malfunction transients if the following inequalities are met:
Page 5.6-2
___j
NFU-0039 Revision 1 March 14, 198.6 I
. I I
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. oc 0 * <1 - RFDC> oc 0 (least negative bounding value) ocM - RFMTC 2. ocM <most negative bounding value)
- c. ( FOH+ RFF DH) <1+T> FoH <Technical Specifications)
- Page S.6-3
NFU-0039 Revision 1 March 14, 1986 1
- 5.7 Excessive Load Increase 5.7.1 Definition of Accident An excessive load increase accident is defined as a rapid increase in steam gener~tor steam flow that causes a power mismatch between core heat generation and secondary side load demand.
The ensuing decrease in reactor coolant temperature results in a core power increase due to fuel and moderator feedback and/or control system action. Only steam flow increases within the capability of the turbine control valves are considered here; larger flow increases are considered in connection with main steam line I
rupture accidents <Section S.13).
S.7.2 Accident Specific Methods .
All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 4 *
- Page S.7-1
NFU-0039 Revision 1 March 14, 1986 5.7.3 Reload Safety Evaluations.
- Each of the physics parameters calculated above are adjusted to inclu~e the model reliability Factors and biases. 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 with regard to excessive load increase transients if the following inequalities are met:
CYCLE SPECIFIC PARAMETERS SAFETY ANALYSIS PARAMETERS
- a. ~D * <1 - RFOC) ~
0 <least negative bounding value)
- b. ~M - RFMTC ~M <most negative bounding value)
FoH <Technical Specifications)
Page 5.7-2
NFU-0039 Revision 1 March 14, 1986
- 5.8 Loss of External Load 5.8.1 Definition of Accident The most likely source of a complete loss of loa~ is a turbine-generator trip. Above approximately 10% power, a turbine trip generates ~ direct reactor trip which is signaled from either of two diverse inputs:
release of autostop oil or stop valve closure. If credit is taken for a steam bypass system and pressurizer control system, there is no significant increase in reactor coolant temperature or pressure.
To provide a conservative assessment of the accident, however, no credit is taken for direct reactor trip, steam bypass actua-tion, or pressurizer pressure control.
Under these assumptions, both secondary and primary pressures increase rapidly, and a reactor trip is generated by the high pressurizer pressure signal. This accident is primarily of concern from the standpoint of demonstrating the adequacy of overpres-surization protection, since the hot Page S.8-1
NFU-0039 Revision 1 March 14 , 1 986 channel DNBR increases <or decreases only slightly) during the accident.
5.8.2 Accident Specific Methods
- All of the physics parameters of importance to this accident are calculated with the Genera1 Physics Methods described in Section 4.
5.8.3 Reload Safety Evaluations Each of the physics parameters calculated above are adjusted to* include the model reliability factors and biases. These adjusted values are the cycle specific parameters which are then compared to the bounding v~lues assumed in the safety analysis. The 'Cycle specific parameters are acceptable with regard to excessive load increase transients if the following inequalities are met:
Page 5.8-2
NFU-0039 Revision 1 March 14, 1986 CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- c. ~p<t>scRAM *(1-RFROD>l ~p<t>scRAM
<bounding>
- Page S.8-3
NFU-0039 Revision 1 March 14, 1986
- 5.9 Loss of Reactor Coolant Flow - Pump Trip 5.9.1 Definition of Accident The accident considered here is the simultaneous loss of electrical power ta all four of the reactor coolant pumps. As a result of loss of driving head supplied by the pumps, the coolant flow through the core begins to decrease. This decrease in flow rate is retarded by the hydrau-lie inertia of th* fluid itself and the fly-wheels on the pump motors. The reactor is tripped by any one of several diverse and redundant signals which monitor coolant flow conditions. This trip results in a power reduction before the thermal-hydraulic condi-tions in the core approach those which could result in damage to the fuel. Loss of power to two of the pumps with all pumps initially operating may also be considered, but the j
consequences are less severe than* for the four pump trips. Seizure of the reactor coolant pump shaft is considered in Section s.10 .
- Page S.9-1
NFU-0039 Revision 1 March 14, 1986 5.9.2 Accident Specific Methods All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 4.
5.9.3 Reload Safety Evaluations Each of the physics parameters calculated above are adjusted to include the model reliability factors and biases. These adjusted values are the cycle specific parameters which are then
- compared to the bounding values assumed in the safety analys~s. The cycle specific parameters are acceptable with regard to excessive load increase transients if the following inequali-ties are met:
~!
Page S.9-2 I
J
NFU-0039 Revision 1 March 14, 1986 CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. O< D * ( 1 + RF 0 > 2. ()(0 <most negative bounding value)
- b. O<M + RFM i ()(M <1 east negative bounding value) c* ~p(t)SCRAM*<l-RFROD> 2. ~P(t)SCRAM<bounding)
- 1-,
.\":
d* <FoH+RFFDH> <l+T) i F DH. <Techn i ca 1 Specifications)
Page S.9-3
NFU-0039 Revision 1 March 14, 1986 s.10 Loss of Reactor Coolant Flow - Locked Rotor s.10.1 Definition of Accident The accident postulated is the instantaneous seizure of the rotor of a single reactor coolant pump. Flow through the affected loop is rapidly reduced, leading to a reactor trip initiation due to low flaw. The sudden decrease in core flow while the reactor is at power results in a degradation of core heat transfer. Departure from nucleate boiling is assumed ta occur in the analysis for this accident. However, the calculated peak clad surface temperatures reached do not result in loss of fuel integrity nor consequential loss of core cooling capabil-ity.
s.10.2 Accident Specific Methods All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 1.
Page 5 .10-1
NFU-0039 Revision 1 March 14~ 1986 s.10.3 Reload Safety Evaluations Each of the physics parameters calculated above are adjusted to include the model reliability factors and biases. 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 with regard to excessive load increase transients if the following inequalities are met:
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. ()(0 * (1 + RFOC) 2 ()(0 <most negative bounding value)
- b. O<M + RFOC i O<M <least negative bounding value)
- c. 6p(t)SCRAM<bou~ding) 6p(t)SCRAM *(1-RFROO> L.
- d. <FoH + RFFOH)Cl+T) FoH <Technical Specifications)
Page 5. 10-2
NFU-0039 Revision 1 March 14, 1986
- s.11 Fuel Handling Accident s.11.1 Definition of Accident The accident considered is the sudden release of the gaseous fission products held in the voids between the pellets and cladding of one fuel assembly. The activities associated with this accident would be released either inside the Containment Building or the Auxiliary Building.
A high radiation level on the Auxiliary building ~
vent monitor would automatically activate the special ventilation system with subsequent absolute and charcoal filtration. In calculating the offsite exposure from the accident, however, it is assumed that the activity is discharged to the atmosphere at ground level from the Auxiliary Building since this maximizes the offsite doses.
S.11.2 Accident Specific Methods All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 4 *
- Page s.11-1
NFU-0039 Revision 1 March 14, 1986 s.11.3 Reload Safety Evaluations
- The FQ calculated above is conservatively adjusted to include reliability factors and biases. This value is then compared to the value assumed in the accident analysis. The comparison is acceptable if the following inequality is satisfied:
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
<FQ+RFFQ)*<l+T) FQ <maximum bounding this accident)
Page 5.11-2
NFU-0039 Revision 1 March 14, 1986
- 5.12 Main Steam Line Break s.12.1 Accident Description The accident considered here is the com-plete severance of a pipe inside contain-ment at the exit of the steam generator with the p1ant initially at no load condi-tions and all four reactor coolant pumps running. The complete severance of a p{pe outside the containment, down-stream of the steam flow measuring nozzle Cat the same plant conditions> is also considered and has similar results. In both cases the resulting uncontrolled steam release causes a rapid reduction in reactor coolant temperature and pressure as the secondary side is depressurized. If the most reac-tive RCC assembly is assumed stuck in its fully withdrawn. position, there is a possibility that the core will become critical and return to power due to the negative moderator coefficient. A return to power is potentially a problem mainly because of the high hot channel factors Page s.12-1
NFU-0039 Revision 1 March 14, 1986 which exist with a stuck RCC assembly. The core is ultimately restored to a subcriti-cal condition by boric acid injection via the emergency Core Cooling System. The zero power case is considered because the stored energy of the system is at a minimum and steam generator secondary inventory is at a maximum under these conditions, thus increasing the severity of the transient.
Page 5.12-2
NFU-0039 Revision 1 March 14, 1986
- s.12.2 Accident Specific Methods All *of the physics parameters of *importance to this accident are calculated with the General Physics Methods, except Keff<T>*
Core keff versus moderator temperature is calculated for zero power conditions with all rods in, and the most reactive rod stuck out, at 1000 psia, assuming an initial moderator temperature of 547°F.
s.12.3 Reload Safety Evaluations Each of the physics parameters calculated above are adjusted to include the model reliability factors and biases. These results are then compared to the bounding
'values assumed in the safety analysis. For Keff versus temp~rature during cooldown, the reliability factors are applied to the
- calculation of the moderator temperature Page 5.12-3
NFU-0039 Revision 1 March 14, 1986 coefficient prior to the determination of Kerr* The cycle specific parameters are acceptable with regard to excessive load increase transients if the following inequalities are met:
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS Keff <T> bounding SOM L SOM <bounding)
~O * <1 - RFOC) i ~
0 <least negative bounding value)
Page S.12-4
NFU-0039 Revision 1 March 14, 1986 5.13 Control Rod Ejection 5.13.1 Definition af Accident This ac~ident is ~ostulated to result from the unlikely failure of a control rod mechanism pressure housing followed by ejection of an* RCC assembly by the reactor coolant system pressure. If a rod inserted in a high worth region of the core were to be ejected, the rapid reactivity insertion and unfavorable power distribution which would result might cause localized fuel rod damage.
5.13.2 Accident Specific Methods All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 2, except for the following:
Maximum EJected Rod Worth, 6p EJECT The ejected rod worth is calculated with the nodal physics model in full core, and in three dimensions. The initial rod*
Page 5.13-1
NFU-0039 Revision 1 March 14 , 1986 position is assumed ta be the power depen-dent insertion limits. Calculations are performed at BOC and EOC, full power and zero power. In each case, the core thermal distribution is held constant at the pre-accident condition, and the flux is allowed ta redistribute without feedback in response ta the positive reactivity inser-tion of the ejected rod. No changes in xenon or boron are permitted.
Heat Flux Hot Channel Factor~ FQ The maximum FQ is determined from the same cases described above for t:,,p EJECT CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
- a. °'D * <1 - RFDC) °'D <least negative bounding value) b* l\ff* ( l+RFB) l\ff <minimum)
- c. t:,,pEJECT*(l+RFROD) ~- t:,,pEJECT <bounding)
- d. t:,,p (t)SCRAM*(l-RFR00)2. t:,,p (t) SCRAM( bounding)
FQ (bounding)
Page 5.13-2
NFU-0039 Revision 1 March 14, 1986
- 5.14 Loss of Coolant Accident 5.14.l Definition of Accident The loss of coolant accident is defined as the rupture of the reactor coolant system piping or any line connected to the system, up to and including a double-ended guil-lotine rupture of the largest pipe.
Ruptures of small flow area would cause coolant expulsion at a rate which would allow replacement at the same rate via the charging* pumps and an orderly shutdown would be possible. A larger rupture would result in a net loss of reactor coolant inventory and a decreasing pressurizer water level and pressure. When the pres-surizer low pressure trip setpoint is reached, both a safety injection signal and a reactor trip occurs followed by the injection of borated water into the reactor i coolant system, isolation of the normal feedwater, and initiation of auxiliary feedwater supply. When the reactor coolant system depressurizes to 600 psia, the
- nitrogen bubble in the accumulator tanks Page 5.14-1
NFU-0039 Revision 1 March 14, 1986 expands, forcing additional water into the reactor coolant system. For large breaks, void formation in the core coolant during the initial blowdow~ phase results in almost immediate power reduction down to decay heat levels.
5.14.2 Accident Specific Methods All of the physics parameters of importance to this accident are calculated with the General Physics Methods described in Section 2.
CYCLE SPECIFIC SAFETY ANALYSIS PARAMETERS PARAMETERS
/1p <UsCRAM*(l-RFROO) l. /1p(t)SCRAM <bounding)
<FQ+RFFG)*(l+T) s FQ <Technical Speci-fication Limits)
(j SOM 2 SOM <bounding)
Page*s.14-2
NFU-0039 Revision 1 March 14, 1986
6.0 REFERENCES
- 1. Advanced Recycle Methodology Program <ARMP) System Documentation CCM-3 Research Project 118-1, September 1977.
- 2. Pfeifer, c. J., "POQ-7 Reference Manual" 11"9 WAPD-TM-947(L), Westinghouse Electric Corporation, February 1971.
- 3. Breen, R. J., o. J. Marlowe, and c. J. Pfeifer,
- "HARMONY: System for Nuclear Reactor Depletion Computation," WPAD-TM-478, Westinghouse Electric Corporation, January 1965.
- 4. Wa 1 pol e , R. E. , Myers , R. H. , "Prob ab i 1 it y and Statistics for Engineers and Scientists", Mac Millan Publishing Company, New York, 1978.
- s. Owen, o. 8., "Factors for One-Sided Toler:-ance Limits and for Variables Sampling Plans", SCR-607, Sandia Corporation, March 1963. <Available from office of Technical Services, Department of Commerce, Washington, O.C.)
- Page 6.0-1
NFU-0039 l
Revision 1 March 14, 1986
- 6. USNRC Regulatory Guide 1.126, "An Acceptable Model and Related Statistical Methods for the Analysis of Fuel Densification.", March 1978.
- 7. Somervi 11 e, P. N., "Tables for Obtaining NonParametric Tolerance Limits, Annals of _Mathematical Statistics 29, 599 (1958).
- 8. Assessment of the Assumption of Normality <Employing Individual Observed Values), ANSI NlS.15-1974.
- 9. Safety Evaluation of the PSE&G Rod Exchange Methodology, NFU-004, Revision 2, August 22, 1984. -*
- 10. Accident Analysis Methods for Application to Salem Nuclear Units, NFU-0033, Revision 0, March 14, 1986.
Page 6.0-2
NFU-0039 Revision 1 March 14, 1986
- APPENDIX A STATISTICAL METHODS FOR THE DETERMINATION AND APPLICATION OF UNCERTAINTIES
- NFU31/l A
NFU-0039 Revision 1 March 14, 1986
- APPENDIX A STATISTICAL METHODS FOR THE DETERMINATION AND APPLICATION OF UNCERTAINTIES The purpose of using statistical methods is to compute the value X such that there is a 95% probability at the 95% con¥ictence level that X will be conservative with respect to X (true value)Rwhen applying the calculational methods to safety related reactor analyses.
The first step is to determine whether or not a distribution is normal. If it is, the methods described in Sectiori A.1 are used. If the distribution cannot be*
treated as normal, but the distributions are known, then the methods described in Section A.2 are used *
- NFU31/l A 1
NFU-0039 Revision 1 March 14, 1986 A.l Application of Normal Distribution Statistics Treatment of Measurement and Calculational Uncertainties Comparison of measured and calculated reactor parameters include the effects of both the measurement and calculational uncertainties. Methods used in this report to isolate the calculational uncertainties are described below in terms of the following definitions:
XT = true reactor parameter XM = measured reactor parameter XC = calculated reactor parameter eM = x M - XT = measurement error ec = XC - XT = calculation error eMC = XM - Xe = observed differences
µi = ei =-mean error (i = M, C, or MC)
If eM and ec are independent, then the following relationships exist. (Not~ that these relationships apply for non-normal distributions as well).
NFU-0039 Revision 1 March 14, 1986 2
= 2 + 2 (A-1) a MC ac crM
= µ - (A-2)
µ MC M These equations can be solved for c<G and~* Once
~~t:n~h~~ ~:~ ~=l~~!~t~~ =~~~yh~~~~~~~:!ism to future calculations of reactor parameters, XR, as follows:
The factor K is defined as described in Table A.l to provitle a 95% probability at the 95%
confidence level that X is conservative with
,respect to the true valfie, XT.
Alternatively, as done in most instances this report, it can be noted that since each term in equation (A-1) is greater than or equal to zero each term is bounded by the variance in the observed differences. Thus the calculational uncertainty can be conservatively estimated as the uncertainty in the observed differences between measured and calculated values:
2 2 ac = (A-3) er MC or ac = a MC (A-4)
In the later alternative, once crMC and µMC are calculated from historical data they can be used to apply conervatism to future calculations of reactor parameters, XR as follows:
XR = Xe + µMC + RF (A-5) where RI:', = K C a MC The quantity µM~ is used as a bias on the calculated parameter and K is as defined above. The term RF is called the reliability factor as described below *
- A 3 NFU31/l
1----
NFU-0039 Revision 1.
March 14, 1986 Reliability Factors It is the objective to define reliability factors which are to be used to increase or decrease calculated results to the point where there is a 95%
probability at the 95% confidence level that they are conservative with respect to actual parameters.
For any given application, we are only concerned with one side of the component~ that is, if the calculated value is too large or too small. We may therefore use one-sided tolerance limits based on normal distributions to find a Kc which will give a 95% probability at the 95% confidence level to the reliability factor defined by RF = Kc* a c Numerical values of Kc for various sample sizes used to calculated crc are provided on Table A.l
{Reference 5)
- NFU31/l 66 A4
NFU-0039 Revision 1 March 14, 1986 TABLE A.l SINGLE-SIDED TOLERANCE FACTORS N Kc 2
26.26 3 7.66 4 5.14 5 4.20 6 3.71 7 3.40 8 3.19 9 3.03 10 2.91 11 2.82 12 2.74 15 2.57 :*),
20 2.40 25 2.29 30 2.22 40 2.13 60 2.02 100 1.93 200 1.84 500 1. 76 00 1.645 N = Number of samples used to
- calculate ac
- NFU31/l 67 AS
NFU-0039 Revision 1 March 14, 1986 A. 2 Application of Non-Normal Distribution Stat.istics This section documents the procedure used to determine the value X such that there is a 95%
confidence level thatRX~ will be conservative relative to the actual value (X ) when the distribution of x is not assumea to be a normal distribution. The approach taken is consistent with non-parametric methods given in reference 6 and 7.
In general, the procedure requires the ordering of N samples taken from a continuous but unknown distribution function. The statistic "m" is determined such that, at the 95% confidence level, 95% of the population lies between the rth smallest and the sth largest value in the ordered N samples, where m = r + s. The statistic m can be determined from Table A.2 (Reference 7). Since, for any given application we are usually only concerned with one side of the component, one-sided tolerance limits are required. Therefore, for upper one-sided tolerance limits, r is set to zero, and m = s.
This procedure has been implemented to obtain reliability factors using the following steps.
First, the mean error µMC = eMC was determined, where eMc =
- XM - Xe* (See
.Section A.l for definitions). Next, the population of N errors e - µ was computed, and the resulting di~tribu~lon ordered. Using Table A.2, the mth value of the error distribution defines error eR' for which, at the 95% confidence level, 95% of the error distribution will be less than
.eR.
A 6 NFU31/l
NFU-0039 Revision 1 March 14, 1986 Once e and µ are calculated from R MC historical data they can be used to apply conservatism to future calculations of the reactor parameter, XR, as follows:
XR = XC + + RF (A-6)
µMC -
where RF = e
- R The term RF is the reliability factor which provides the desired 95% probability at the 95% confidence level for the computed parameter x.
A 7 NFU31/l
NFU-0039 Revision 1 March 14, 1986 TABLE A.2 Values of m for 95% Confidence and 95%
Probability Tolerance Limits Number of Observations (n) m so 55 60 1 65 1 70 1 75 1 80 l 85 l 90 l 95 2 100 2 110 2 120 2 130 3 140 3 150 3 170 4 200 5 300. 9 400 13 500 17 600 21 700 26 800 30 900 35 1000 39 For n> 1000 Increase m by 4 for each additional 100 observations NFU31/l 70 A8
_J
NFU-0039 Revision 1 March 14, 1986 APPENDIX B .
COMPUTER CODE
SUMMARY
DESCRIPTION
- B
NFU"".'0039 Revision 1 March 14, 1986
- APPENDIX B COMPUTER CODE
SUMMARY
DESCRIPTION COMPUTER CODE DESCRIPTION CPM CPM is a multigroup two-dimensional collision probability code for*
depletion and branch calculations for a single assembly.
Reference 1 EPRI-CELL EPRI-CELL computes the space, energy and burnup dependence of the netitron spectrum within cylindrical cells of light water reactor fuel rods. It is used to generate cross sections for PDQ on a ECDATA file
- Reference 1 INTEGRAL INTEGRAL edits PDQ files to obtain pin and assembly powers. Pin to box ratios are then input to TRINODE.
NUPUNCHER NUPUNCHER prepares HARMONY cross section tables from cross section data on an ECDATA file.
Reference 1 PDQ7/ PDQ7/HARMONY is a nuclear reactor HARMONY analysis program which solves the neutron diffusion equations and performs depletion calculations.
Reference 2,3
- NFU31/l B 1
NFU-0039 Revision 1 March 14, 1986 COMPUTER CODE DESCRIPTION SHUFFLE SHUFFLE is the same as EPRI-SHUFFLE and will read a PDQ7 concentration file and write a new updated concentration file. It is used to simulate* assembly movement between cycles.
Reference 1 SIGMA SIGMA ~alculates the predicted detector reaction rates using nodal power distribution and PDQ7 detector reaction rate to assembly power factors. The predicted detector reaction rates are then compared to measured detector reaction rates
- TRI NODE TRINODE is a modified version of the EPRI-NODE-P computer code program.
Modifications are summarized as follows:
a) Automated file management b) User triendly input c) Rod search for constant axial offset control d) Separate BP reactivity insertion equations e) Flexible edit options TAU TAU is a computer code used to compute statistics from residual reaction rates generated by SIGMA
- NFU31/1 B 2
NFU-0039 Revision 1 March 14 , 1986
- APPENDIX C SAMPLE OF PROPOSED REVISIONS TO THE TECHNICAL SPECIFICATIONS AXIAL FLUX DIFFERENCE TILT 6.9.12 POWER DISTRIBUTION LIMIT REPORT 3/4.2 BASES
- c
NFU-0039 Revision 1 March 14, 1986 3/4.2 POWER DISTRIBUTION LIMITS AXIAL FLUX DIFFERENCE <AFD>
LIMITING CONDITION FOR OPERATION The indicated AXIAL FLUX DIFFERENCE <AFD> shall be maintained within the AFD maneuvering band. The upper and lower limits of this band ~hall be determined periodically~
during operation, based on in-core flux map measurements, and the actual operating history.
APPLICABILITY: MODE 1 ABOVE 50% RATED THERMAL POWER*
ACTION:
- a. With the indicated AXIAL FLUX DIFFERENCE outside of the above limits:
- 1. Either restore the indicated AFD to within the maneuvering band limits within 15 minutes, or
- 2. Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within 30 minutes and reduce the Power Range Neutron Flux-High Trip Setpoints to ~55% of RATED THERMAL POWER within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
- b. THERMAL POWER shall not be increased above 50% unless the indicated AFD is within the maneuvering band.
- See Special Test Exception 3.10.2 SALEM - UNIT 1 314 2-1 C-1
NFU-0039 Revision 1 March 14, 1986 POWER DISTRIBUTION LIMITS SURVEILLANCE REQUIREMENTS 4.2.1.1 The indicated AXIAL FLUX DIFFERENCE shall be determined to be within its limits during POWER OPERATION above 15% of RATED THERMAL POWER by:
- a. Monitoring the indicated AFD for each OPERABLE excore channel
- 1. At least once per 7 days when the AFD Monitor Alarm is OPERABLE, and
- 2. At least once per hour for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after restoring the AFD Monitor Alarm to OPERABLE status.
- b. Monitoring and logging the indicated AXIAL FLUX DIFFERENCE for each OPERABLE excore channel at least once per hour for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> and at least once per 30 minutes thereafter, when the AXIAL FLUX DIFFERENCE Monitor Alarm is inoperable. The logged values of the indicated AXIAL FLUX DIFFERENCE shall be assumed to exist during the interval preceding each logging.
4.2.1.2 The indicated AFD shall be considered outside of its limits when at least 2 of 4 or 2 of 3 OPERABLE excore channels are indicating the AFD to be outside the limits of Specification 3.2.1.
4.2.1.3 The upper and lower limits of the AFD maneuvering band shall be *determined a) From the POWER DISTRIBUTION LIMIT REPORT for use during
~ower escalation following a refueling .outage through the first 31 EFPD of operation; b) Determined from in-core flux map measurement within 15 EFPD of first exceeding 50% power following a refueling outage, and at least once every 31 EFPO ther~after. The provisions of specification 4.0.4 are not applicable.
4.2.1.4 The determination of the upper and lower limits of the AFD maneuvering band shall utilize T<Z> and DF<Z> engineering factors which are:
a) Obtained from the Power Distribution Limit Report for core burnups not greater than 31 EFPD, or b) Derived from core physics analysis of the power distribution behavior using the actual operating history up to the d a t .
of the flux map used to determine the AFD limits.
C-2
NFU'."'"0039 Revision 1 March 14, 1986
- 4.2.1.s The upper and lower limits of the AFO maneuvering band shall be determined from flux map measurements as follows:
Upper AFD limit = { AOM + [ll AOCZ> J 1 east }
- pre l positive}
Lower AFD limit = { AOM + [ll AO< Z> J least
- prel negative AOM =measured axial offset from fl.ux map.
FL<Z>
AO<Z> = .1 T<Z> { F~<Z>
Q DF<Z> }
- F~<Z> = 2.32 x KCZ) as defined in Spec. 3.2.2 for P=l.
F~<Z> = maximum measured values of FQ<Z> from* flux map, without uncertainties N
FMQz> = FQE x FQU x FQ<Z>
FQE = 1. 03 t accounts for manufacturing tolerances FQU. = 1. 05, accounts for measurement uncertainty TC Z> , OF ( Z) = AFD Engineering factors* identified in Specification 4.2.1.4.
= THERMAL POWER ASSOCIATED WITH AFD LIMIT RATED THERMAL POWER C-3
NFU-0039 Revision 1 March 14 , 1986 POWER DISTRIBUTION LIMITS HEAT FLUX HOT CHANNEL FACTOR-FQCZ>
LIMITING CONDITION FOR OPERATION 3.2.2 F~<Z> shall be limited by the following relationships:
F8<z> ~ [2.32J CK<Z>J for P > o.s p
F8<Z) ~ [4.64] CK<Z>J for P ~ 0.5 wher~ p = THERMAL POWER
--R-AT""""E......D
___T__H"""'"E--R-MA_L_P__O_W__E__R and K<Z> is the function obtained from Figure 3.2-2 for a given core height location.
APPLICABILITY: MODE 1 ACTION:
with F8<Z> exceeding its limits:
- a. Reduce THERMAL POWER at least 1% for each 1% F~<Z) exceeds the limit within 15 minutes and similarly redure the Power Range Neutron Flux-High Trip Setpoints within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />; POWER OPERATION may proceed for up to a total of 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />; subsequent POWER OPERATION may proceed provided the Overpower T ~rip Setpoints have been reduced at least 1%
for each 1% F <Z> exceeds the limit. The Overpower ~T Trip Setpoint reduc~ion shall be performed with the reactor in at least HOT STANDBY.
- b. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POWER, above the reduced limit required bM a. above; THERMAL POWER may then be increased provided FQCZ) is demonstrated through incore mapping to be within its limits.
C-4
NFU-0039 Revision 1 March 14 , 1986
- POWER DISTRIBUTION LIMITS SURVEILLANCE REQUIREMENTS The provisions of Specificati~n 4.0.4 are not applicable.
FM~<Z> shall be determined to be within its limits using flax map measurements at least once per 31 EFPO.
F8 = F~<Z> x FQE x FQU where FNQ<Z>, FQE' FQU are defined in Specification 4.2.i.s.
C-5
NFU-0039 Revision 1 March 14, 1986 POWER DISTRIBUTION LIMITS QUADRANT POWER TILT RATIO LIMITING CONDITION FOR OPERATION 3.2.4 THE QUADRANT POWER TILT RATIO shall not exceed 1.03.
APPLICABILITY: MODE 1 ABOVE 50% OF RATED THERMAL POWER*
ACTION:
- a. With the QUADRANT POWER TILT RATIO determined to exceed 1.03
- but ~ 1.09:
- 1. Within 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />s:
- a. Either reduce the QUADRANT POWER TILT RATIO to within its limit, or
- b. Reduce THERMAL POWER at least 3% from RATED THERMAL POWER for each. 1% of indicated QUADRANT.
POWER TILT RATIO in excess of 1.0 and similarly reduce the Power Range Neutron Flux-High Trip Setpoints within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
- 2. Verify that the QUADRANT POWER TILT RATIO is within its limit within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after exceeding the limit or reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within the next 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and reduce the Power Range Neutron-Flux-High Trip setpoints to ~ 55% of RATED THERMAL POWER within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
- 3. Identify and correct the cause of the *out of limit condition prior to increasing THERMAL POWER; subsequent POWER OPERATION above 50% of RATED THERMAL POWER may proceed provided that the QUADRANT POWER TILT RATIO is verified within its limit at least once per hour for 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> or until verified acceptable at 95% or greater RATED THERMAL POWER.
- b. With the QUADRANT POWER TILT RATIO determined to exceed 1.09 due to misalignment of either a shutdown or control rod:
- 1. Reduce THERMAL POWER at least 3% from RATED THERMAL POWER for each 1% of indicated QUADRANT POWER TILT RATIO in excess of 1.0, within 30 minutes.
- See Special Test Exception 3.10.2.
SALEM - UNIT 1 3/4 2-11 C-6
- NFU-0039 Revision 1 March 14, 1986
- POWER DISTRIBUTION LIMITING CONDITION FOR OPERATION <Continued)
- 2. Verify that the QUADRANT POWER TILT RATIO is within its limit within 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> after exceeding the limit or reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within the next 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and reduce the Power Range Neutron Flux-High Trip Setpoints to ~ 55% of RATED THERMAL POWER within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />;
- 3. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POWER; subsequent POWER OPERATION above 50% of RATED THERMAL POWER may proceed provided that the QUADRANT POWER TILT RATIO is I verified within its limit at least once per hour for 12 **I hours or until verified acceptable at 95% or greater RATED THERMAL POWER.
c* With the QUADRANT POWER TILT RATIO determined to exceed 1.09 due to causes other than the misalignment of either a shutdown or control rod: )-\*
- r. Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and reduce the Power Range Neutron Flux-High Trip Setpoints to ~ 55% of RATED THERMAL POWER within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
- 2. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POWER; subsequent POWER OPERATION above 50% of RATED THERMAL POWER may proceed provided that the QUADRANT POWER TILT RATIO is verified within its limit at least once per hour for 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> or until verified at 95% or greater RATED THERMAL POWER.
SURVEILLANCE REQUIREMENTS 4.2.4 The QUADRANT POWER TILT RATIO shall be determined to be within the limit above 50% of RATED THERMAL POWER by:
- a. Calculating the ratio at least once per 7 days when the alarm is OPERABLE.
- b. Calculating the ratio at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> during steady state operation when the alarm is inoperable *
- C-7
NFU-0039 Revision 1 March 14, 1986 POWER DISTRIBUTION LIMITING CONDITION FOR OPERATION <Continued)
- c. Using the movable incore detectors to determine the QUADRANT POWER TILT RATIO at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> when one Power Range Channel is inoperable and THERMAL POWER is > 75 percent of RATED THERMAL POWER.
C-8
NFU-0039 Revision 1 March 14, 1986 POWER DISTRIBUTION LIMIT REPORT 6.9.1.12 The Axial Flux Difference <AFD> limits shall be provided to the NRC Regional Administrator with a copy to the Director of Nuclear Reactor Regulation, Attention: Chief, Core Performance Branch, u. s.
Nuclear Regulatory Commission, Washington, o.c. 20555, at least 60 days prior to each cycle's initial criticality unless otherwise approved by the Commission by letter
- C-9
NFU-0039 Revision 1 March 14, 1986 3/4.2 POWER DISTRIBUTION LIMITS BASES The specifications of this section provide assurance of fuel integrity during Condition I <Normal Operation> and II <Incidents of Moderate Frequency) events by: Ca> maintaining the minimum DNBR in the core L 1.30 during normal operation and in short term transients, and Cb> limiting the fission gas release, fuel pellet temperature and cfadding mechanical properties to within assumed design criteria. In addition, limiting the peak linear power density during Condition I events provides assurance that the initial conditions assumed for th LOCA analyses are met and the ECCS acceptance criteria limit of 2200 0F is not exceeded.
The definitions of hot channel factors as used in these specifications are as follows:
F8<z> Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod at core elevation Z divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods.
FNH Nuclear Enthalpy Rise Hot Channel Factor, is defined *as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power.
3/4.2.1 AXIAL FLUX DIFFERENCE <AFO>
The limits on AXIAL FLUX DIFFERENCE assure that the FM<Z> upper bound envelope of 2.32 times the normalized axial peaking ~actor is not exceeded during either normal operation or in the event of xenon redistribution following power changes.
The AFD limits are periodically re-evaluated during plant operation. This evaluation is based on flux map measurements and incorporates the effects the actual operating history. This provides additional assurance that the hot channel factor FM<z>,
remains within ~ts limits. The TCZ) and DF<Z> engineering f~ctors are determining from the actual operating hisMory, and represent a conservative estimate of the changes in F ~during the allowable operating maneuvers at the burnup conditio~ of the core.
C-10
NFU-0039 Revision 1 March 14 , 1986 POWER DISTRIBUTION LIMITS BASES Provisions for monitoring the AFD are derived from the plant nuclear instrumentation system through the AFD Monitor Alarm. A control room recorder continuously displays the auctioneered high flux difference and the maneuvering band limits as.a function of power
~evel. A first alarm is received any time the auctioneered high flux difference exceeds the maneuvering band limits.
3/4.2.2 and 3/4.2.3 HEAT FLUX AND NUCLEAR ENTHALPY HOT CHANNEL AND RADIAL PEAKING FACTORS F8<Z) and F~H The limits on heat flux and nuclear enthalpy hot channel factors ensure that 1) the design limits on peak local power density and
. minimum DNBR are not exceeded and 2l in the evant of a LOCA the peak ~
fuel clad temperature wi 11 not exceed the 2200 F ECC.S acceptance criteria limit.
- Eac~ of these hot channel factors are measurable but will normally only be determined periodically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient to insure that the hot channel factor limits are maintained provided:
- a. Control rod in a single group move together with no individual rod insertion differing by more than + 12 steps from the group demand position.*
- b. Control rod groups are sequenced with overlapping groups as described in Specification 3.1.3.5.
- c. The control rod insertion limits of Specifications 3.1.3.4 and 3.1.3.5 are maintained.
- d. The axial power distribution, expressed in terms of AXIAL FLUX DIFFERENCE, is maintained within the limits.
The relaxation in F~~ as a function of THERMAL POWER allows changes in the radial power shape for all permissible rod insertion limits. F~H will be maintained within its limits provided conditions a through a above, are maintained *
- C-11
NFU-0039 Revision 1 March 14, 1986 POWER DISTRIBUTION LIMITS BASES When an Fa measurement is taken, both experimental error and manufacturing !olerance must be allowed for. 5% is the appropriate allowance for a full core map taken with the incore detector flux mapping system and 3% is the appropriate allowance for manufacturing tolerance.
When FN is measured, experimental error must be allowed for and 4% is the a~~ropriate allowance for a full core map taken with the incore detection system. The specified limit for F~~ also contains an 8% allowanc~ for uncertainties which mean that norma'I' operation will result in F6 ~ s 1.55/1.08. The 8% allowance is based on the following consideratio~s.
- a. abnormal perturbations in the r~dial power shape, suchMas from rod mi_salignment, effect F6 H more directly than F Q'
- b. aMthough rod movement has a direct influence upon limiting
_F c:i -to within i~s limit, such control is not readily a v a i l .
abTe to limit F6 H, and -
- c. errors in prediction for control power shape deteMted during startup physics tests can be compensated for in F by r~stricti~g axial flux distributions. This compen~ation for F6 H is less readily available.
3/4,2.4 QUADRANT POWER TILT RATIO-The quadrant power tilt ratio limit assures that the radial power distribution satisfies the design values used in the power capability analysis. Radial power distribution measurements are made during startup testing and periodically during power operation.
The limit of 1.03 at which corrective action is required provides ONB and linear heat generation rate protection ~ith x-y plane power tilts. Adequate margin for a 3% increase in the Quadrant Power Tilt Ratio is included in the ~eload Safety Evaluation, and also in the determination of -the AFD limits as described in Specification 4.2.1.4, The two hour time allowance for operation with a tilt condition greater than 1.03 but less than 1.09 is provided to allow identifi-cation and correction of a dropped or misaligned rod. In the event sHch action does not correct the tilt, the margin for uncertainty on F 0 is reinstated by reducing the power by 3 percent from RATED THERMAL POWER for each percent of tilt in excess of 1.0.
C-12
NFU-0039 Revision 1 March 14 , 1986
- APPENDIX 0 SAMPLE OF PROPOSED POWER DISTRIBUTION LIMIT REPORT .
- D
NFU-0039 Revision 1 March 14, 1986
- POWER DISTRIBUTION LIMIT REPORT This report is provided for Salem Unit , Cycle N , in accordance with Paragraph 6.9.1.12 of the Salem UnTI _Nuclear Plant Technical Specifications.
AFD Limits The upper and lower limits of the AFD maneuvering band are:
Core Lower Upper Power AFD AFD Limit Limit Limit 100% -20% +12%
75% -20% +12%
50% -20% +12%
These limits are appropriate for use during power escalation of Cycle
_rf_, following the refueling outage, through the first 31 EFPD of operation *
- T<Z>. DF<Z> Engineer Factors The T<Z> and DF<Z> AFD engineering factors for BOC are shown in Figure
- 1. These factors are based on the actual cycle <N-1> operation through shutdown, and are appropriate for use in cycle <N> from startup through 31 EFPO.
0-1
NFU-0039 Revision 1 March 14, 1986 FIGURE 1 CYCLE <N> AFD ENGINEERING FACTORS i.:~~; ~~~~: *: ~~~~ ~=~~:-:~~ *.~~*~t:*.~J=:=F-~ ~:~i~":: ~~~~f~::. -~==-?~~:~~- =-:=:*r=.=~t=*+/-~:* ~ ~?,:-:~~ ~7~.=i"::~~1 =~~~~-~-~ ~?~~-=.:?: ~i~=-~
~~~~~-~-~ ~~~::~~ .:*~~~-~ ~*:*~~ ~- .f:. ~~~:~.:-~~~~:~~~~:~ ~~;~~~-~~~ ~~j~-~~-~~~~~~~~~ ::~*~i~~::::~~~~?~~E*:~ :::~~-~=~~-~ ~~~1~~*~* ~-~~~~.:=<b~:-~ !~~~:t ~-~ ~-~ ~-=~~: :: ~-:*~~~; ~
- c: :==-~ :. h;;:::*! - =: . _
- l~:~:~;::.j=::T::: =>>:. :=::=:~~~:: ::.c:f==-:::::r~:: ~:::::)=:::µ=::: ~~::-~=c~:i::~::::= :-<: :[:::::=
.. : t::-~t::'jl:(:~~J:::::: ~ :::::::: :::~:=: J::_x '~: !:=::1:=:.: ::::Y:: :-::::~~~~~:~~~~~ ::~~~::: ~~~~,~:~~~b~:~=-:= ~:-::>::r~:~~::-=:
- - . r~='~'*:~~-;:::-::r==~~-- J::.:_,i*:>: :::::!:,. 1:-:~~~'"=l=~L=:=:~:,~: :=:~+/-'~:~ ~~~::c:~:~-:=:i::.:-=:S::~~=--=-~~~:=:-=-i:~=~:::b:-:::=~c=:-1~:.~~:~J::::i='
l:c~~=:
I.: ~:
- .:.:.::.::: I.. .
~~..:. ::***1 *-
I::: '. .... ~c:
i..:._:__*. _ .~i._***-~----***-L--r+---: ; ...
I- . ~~-- ' I I .I l....:::.:.
i
~--!-;-----------1-~--*-
- ------*-------- ~-------
L-----------------------*----~-------'f'-;:-,_-'_-.:..:-l___.*___-_-..~....~,-:-: -;~-;:-~*;-!~'-'~-~: .:~: .:.~:.:;~=-i~t-=-:;=.:.:-:.: ~: .1:c:. ~:. :.r:.:~:.:.*~:-'~-' - r-.....1- - - ;*-~:- ~=-_:+-:l- ~-.~-_-:.: _~: : : = -~ : .:~=:-~.: :~: .:.:-;=-_-_=-i-~: .c: .:.~. .:~;.: :i~=i~,
.. r- : 1-- - :.:::-*r:~.~- -* ***t :-* -- *r* **-:*L~-:::***l.:~~=:.~:=i:: ~:::::.:-:::::::.:::1
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- I . .. :*:.:!: ..:l
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-~------------
)* .. -* --* ....
- !. . .t.
. .... *: t.:-:: :: :: :'
I
- _. ___ :_____. ---~,_1._~*-~~_i:_-~-t-l:_.:-_::_::-t._**-*-_,**__,::--l*~IC'--:_'--*::..:._'-'-i-'---'-'---*-+1-_._~*c__:-_*._,.!_:;_*:_:*_**.....,'_**_....::.:....;..:"_.:....:c=::+=.:.:.:.::'-'--1
~~-:-:-t:--~ l: : ' ... . ~::~~:~t:*:c*.~: .f.: :. : 1:=<~: .:-::-= 1:::
i 40 .SQ . 21::): : !. . :. Io=
j-...:...---i-1--*r--:-:-:--r*r.iilf:::::----=-it-:-:--:--t-:::-:-:---:-f:"-:-:-::-:-:j-::::-;::::::-t:"-:---:::--:=t===---:-t:--:c~t-:-:--:-:::-:-:-r~~ihti~f-:---'-:-:--f::':::::'-=i I !**:::,~. I*-- .. i 0-2
NFU-0039 Revision 1 March 14, 1986 APPENDIX E SAMPLE OF PROPOSED AFD LIMIT CALCULATIONS E
NFU-0039 Revision 1 March 14, 1986 AFD LIMIT ANALYSIS
- 1. FLUX MAP MEASUREMENT; MAP #
Power _ _ _ _ _%
Burnup _ _ _ _ _ MWO/MTU Axial Offset _ _ _ _ _ %
- 2. DEFINITIONS F~(Z) = measured value of maximum FQ at axial level z. with no uncertainties applied
= measured FQ with uncertainties applied F8<Z) = F~ <Z)
- Fu FE = 1. 03 Fu= 1.os F~<Z) = LCO on maximum FQ
= 4. 34 x K <Z) <Prel < 0.5)
K(Z) defined in Tech Specs E-1 I
NFU-0039 Revision 1 March 14, 1986 TCZ) Engineering factors defined in the Peaking Factor DF<Z) Limit Report T<Z) includes 8% allowance for calculational uncertainty DFCZ) includes 3% allowance for increase in the quadrant power tilt ratio.
F~<Z)
FQ-Margin = { F8<n
-1 - OF< Z) }
~AO<Z) = *FQ-Margin/T(Z)
E-2
NFU-0039 Revision 1 March 14, 1986
- 3. CALCULATIONS IN CORE SIGMA F~<Z> F~<Z> DF<Z> FQ-MARGIN T<Z> Ll AOC Z>
z _z_
5 57CTOP>*0.930 1.9542 .105 o.9963 .0178 55.9%
12 51 1.483 2.2098 .109 o.3811 .0162 23.5%
14 48 1.543 2.2214 .111 0.3287 .0147 22.4%
19 43 1.668 2.2504 .116 o.2332 .0120 19.4%
I*
22 40 1. 726 2.2678 .118 0 .1959 .0099 19.8%"
30 32 1.829 2.3142 .114 0.1513 .0024 63.0%
36 26 1.837 2.3200 .101 0. 1619 -.0039 -41.5%
39 23 1.842 2.3200 .094 0 .1655 -.0068 -24.3%
45 17 1.792
- 1. 745 2.3200 2.3200
.081
.076 0.2136 0.2535
-.0114
-.0132 l-18.7%
-19.2%
I**
48 14 53 9CBOT> 1.555 2.3200 .070 0.4220 -.0157 -26.9%
least positive
- least negative
- Upeer Limit =
=
1 ADM +
-4.2% + [+19.4%]}
least
[.!l AOJ positive
}
x pre 1 1.0 = +15.2%
Lower Limit =
=
i ADM +
[.!lAOJ
-4.2% + [-18.7%]
1east negative
}*
}
- Prel 1.0 = -22.9%
- E-3