ML20215N827
| ML20215N827 | |
| Person / Time | |
|---|---|
| Site: | Millstone  |
| Issue date: | 09/30/1986 |
| From: | Dzenis E WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20215N814 | List: |
| References | |
| TAC-63133, TAC-63198, TAC-64674, NUDOCS 8611070315 | |
| Download: ML20215N827 (63) | |
Text
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RELOAD SAFETY EVALUATION
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MILLSTONE NUCLEAR POWER STATION UNIT 2 - CYCLE 8 REVISION 2 SEPTEMBER 1986 Edited By: N. A. Silva Approved: Ts h ps.c E.A.Dzenis,Manper Core Operations Nuclear Fuel Divisions
"'" ' 8 8611070315 861027 PDR ADOCK 05000336 p PDR
TABLE OF CONTENTS Section Title Page
1.0 INTRODUCTION
AND
SUMMARY
1 1.1 OBJECTIVES 1 1.2 GENERAL DESCRIPTION 3 1.3 ANALYSIS BASIS AND ASSUMPTIONS FOR EXTENDED OPERATION
1.4 CONCLUSION
S 4 2.0 MECHANICAL DESIGN 5 2.1 GENERAL DISCUSSION 5 3.0 THERMAL AND HYDRAULIC DESIGN 6 4.0 NUCLEAR DESIGN 7 5.0 ACCIDENT ANALYSIS 8 1
5.1 INTRODUCTION
AND
SUMMARY
8 5.2 ACCIDENT EVALUATION 8 5.2.1 KINETICS PARAMETERS 9 5.2.2 SHUTDOWN MARGIN 9
.,, 5.2.3 CEA WORTHS - 9 5.2.4- CORE PEAKING FACTORS 9 5.3 INCIDENTS REANALYZED AND EVALUATED 10 5.4 LOCA ANALYSIS ,.
< 10 6.0- REFERENCE,S 11 APPENDIX A - NON-LOCA', SAFETY EVAVATION '~ ' A-1 APPENDI'X B . TECHNICAL SPECIFICATIONS CHANGE PAGES B-1 m,
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LIST OF TABLES Table Title Page 1 Millstone Unit 2 Cycle 8 Core Loading 12 ~
2 Millstone Unit 2 Xinetics Characteristics 13 3 Shutdown Requirements and Margins 14 s
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LIST OF FIGURES Figure Title Page 1 Core Loading Pattern 15 2 Zoned enrichment Fuel Assembly Lattice 16 3 Millstone Unit 2 - Cycle 8 Extended Operation 17 Core Power and Inlet Temperature vs. Burnup
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1.0 INTRODUCTION
AND
SUMMARY
1.1 OBJECTIVES ~
This report presents an evaluation for Millstone Nuclear Power Station Unit 2, Cycle 8, which demonstrates that, with respect to the incidents considered in the Basic Safety Report (2) (BSR), the core reload and a maximum of 17.6 percent steam generator tube plugging will not adversely affect the safety of the plant. The evaluation reported includes the effects of extended operation (via power and temperature reduction) past the predicted nominal full power end-of-life. This evaluation was accomplished utilizing the methodology described in Reference 1.
Based upon the above referenced methodology, only those incidents analyzed and reported in the BSR(2) which could potentially be affected by fuel reload or steam generator tube plugging have been reviewed for the Cycle 8 design described herein. The results of new analyses are included and the justification for the applicability of previous results for the remaining incidents is provided.
1.2 GENERAL DESCRIPTION The Millstone Unit 2 reactor core is comprised of 217 fuel assemblies arranged in the configuration shown in Figure 1. Each fuel assembly has a skeletal structure consisting of five (5) zircaloy guide thimble tubes, nine (9) grids (five assemblies have zircaloy grids, two hundred twelve assemblies have Inconel grids), a stainless steel bottom nozzle, and a stainless steel top nozzle. Typically, one hundred seventy-six fuel rods are arranged in the grids to form a 14x14 array. The fuel rods consist of slightly enriched uranium dioxide ceramic pellets contained in Zircaloy-4 tubing which is l
plugged and seal welded at the ends to encapsulate the fuel.
Nominal core design parameters utilized for Cycle 8 are as follows: i Core Power (Mwt) 2700 System Pressure (psia) 2250 Reactor Coolant Flow (GPM) 340,000*
Core Inlet Temperature (*F). 549 Average Linear Power Density (kw/ft) 6.067 Steam Generator Tube Plugging (%) 17.6**
(based on best estimate hot, densified core average stack height of 136.4 inches) .
The core loading pattern for Cycle 8 is shown in Figure 1. The feed fuel for the Millstone Unit 2, Cycle 8 core will consist of sixteen (16) zoned-enrichment interior feed assemblies, each containing sixty (60) fuel rods at 2.6 w/o U235 and one-hundred sixteen (116) fuel rods at 2.9 w/o U235, and forty eight (48) zoned-enrichment peripheral assemblies, each containing sixty.(60) fuel rods at 2.9 w/o U235 and one-hundred sixteen (116) fuel rods at 3.3 w/o U235. The zoned-enrichment assembly configuration is shown in Figure 2. The feed fuel will replace four (4) Combustion Engineering (CE)
Batch A assemblies, eleven (11) Westinghouse Batch F assemblies, forty-three (43) Westinghouse Batch G assemblies and six (6) Westinghouse Batch H assemblies. In addition, one (1) Westinghouse Batch F assembly and four (4)
Westinghouse Batch H assemblies (these Batch F and H assemblies were removed from the core at the end of Cycle 5 & 6 respectively) will replace four (4)
Westinghouse Batch H assemblies and one (1) Westinghouse Batch F assembly used in Cycle 7. As a result of fuel reconstitution, the fuel rods from five (5)
Westinghouse reload assemblies to be used in Cycle 8 have been replaced in Combustion Engineering (CE) skeletons. Also, seventeen (17) fuel rods have been replaced with stainless steel rods in Cycle 8. The seventeen stainless steel rods are distributed among eight (8) fuel assemblies, with the number of stainless steel rods in each of these assemblies ranging from one to five.
l A summary of the Cycle 8 fuel inventory is given in Table 1.
- Minimum guaranteed safety analysis flow
- Maximum for Each Steam Generator uox .no." 2 I ____ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ .- -
1.3 ANALYSIS BASIS AND ASSUMPTIONS FOR EXTENDED GPERATION Operation of a PWR can be continued beyond the nominal full power end-of-life burnup by reducing average coolant temperature and/or core power level.
Extending the fuel cycle in this fashion provides the additional reactivity necessary to offset the effects of increased fuel burnup. The additional reactivity is gained by a reduction in Doppler feedback and a decrease in the xenon (fission product) nuclide concentration. An additional increase in reactivity is likewise gained by a decrease in the reactor coolant temperature.
The analysis assumptions made by Westinghouse to evaluate Millstone Unit 2, Cycle 8 extended operation can be itemized as follows:
- 1. Cycle 8 is to be extended to a maximum cycle burnup of 1000 MWD /MTV greater than the nominal predicted end of-life.
- 2. Cycle 8 full power operation is to be extended as much as possible by l
core inlet temperature reduction.
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- 3. After the full power extension limit is achieved, operation is to be extended by a combination of power and core inlet temperature i
reduction.
- 4. The HZP coolant temperature is a constant 532*F.
- 5. Peak linear heat generation rate is less than or equal to 15.6 kw/ft.
- 6. Extended operation is based on a minimum guaranteed flow of 350,000 gpm.
These specific assumptions were made to define the limits of plant operation during the Cycle 8 extension. A definition of the plant operational limits is necessary to evaluate all possible limiting core characteristics and to perform the safety evaluation. The specific assumptions noted above are not in conflict with the current Technical Specifications limits or protection systems setpoints.
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1.4 CONCLUSION
S From the evaluation presented in this report, it is concluded that the Cycle 8 design does not result in the previously acceptable safety limits for any incident to be exceeded. This conclusion is based on the following:
- 1. Cycle 7 burnup of 10,700 + 1000 MWD /MTU 0 MWD /MTU
- 2. There is adherance to plant operating limitations as given in the Technical Specifications and the revisions given in Appendix B of-this {
report.
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2.0 MECHANICAL DESIGN 2.1 GENERAL DISCUSSION The mechanical design of the Cycle 8 fuel assemblies is essentially identical to that of the Cycle 7 assemblies (5) except that the fuel rod backfill pressure has been reduced. The Westinghouse fuel assemblies are designed to be fully compatible with all resident Millstone 2 fuel assemblies and core ,
components (e.g. adequate clearances for insertion of CEA's, plugging devices, etc.).
Table 1 summarizes pertinent design parameters of the various Westinghouse fuel regions in Cycle 8. The fuel in these regions has been designed according to the fuel performance model in Reference 3. The fuel is designed and operated so that clad flattening will not occur, as predicted by the
, (
Westinghouse model(7) . For all fuel regions, the fuel rod design bases contained in Section 4.2.1.1.1 of Reference 8 will be satisfied, with the exception of the rod internal pressure design basis. The rod internal pressure in the Westinghouse fuel regions will not exceed the nominal primary system coolant pressure during Cycle 8, as stated in the Millstone 2 BSR(2) ,
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3.0 THERMAL AND HYDRAULIC DESIGN A description of the thermal and hydraulic design of the Westinghouse Millstone Unit 2 reload fuel assembly to be utilized in Cycle 8 is given in Chapter 3 of the BSR.
As discussed in the BSR, the Westinghouse fuel assemblies have been designed and shown through testing to be hydraulically compatible with all resident Millstone Unit 2 fuel assemblies. The stainless steel rods in the reconstituted fuel assemblies were treated as heated rods in the THINC ONB analysis. This is conservative since it results in higher subchannel enthalpy predictions.
The DNB analyses for Cycle 8 were performed for a minimum reactor coolant flow rate of 340,000 gpm and a radial peaking factor,r F , f 1.537. This is a decrease in the flow rate and peaking factor assumed in the Cycle 7 and Cycle 6 analyses of 350,000 gpm and 1.565. The reduction in flow has been offset by the reduction in radial peaking factor and this has been confirmed in the Cycle 8 analyses. Appendix B shows the required Technical Specification changes for Millstone Unit 2, Cycle 8. The Cycle 8 analysis takes a partial credit of 3.0% of the net conservatism which exists between convoluting and summing the uncertainties of various measured plant parameters in power space. This partial credit was applied in previous cycles.($)
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4.0 NUCLEAR DESIGN The Westinghouse nuclear design procedures, computer programs, and calculation models utilized in the Millstone Unit 2, Cycle 8 reload design are presented in the BSR.
Figure 3 shows the relative power level and core inlet temperature that are predicted to maintain the core just critical throughout the cycle extension, without changing the soluble baron concentration.
Similar to the Cycle 7 evaluation (5) , Cycle 8 accident simulations take credit for the variable high power trip by terminating accidents 5% above the variable high power trip. Also P values L (see BSR Section 6.0) are computed only if the maximum allowed power density of 21 kw/ft is exceeded.
The Cycle 8 core loading results in a maximum linear heat rate of less than 15.6 kw/ft at all fuel heights at rated power. The safety analysis has I
specifically included the 17 stainless steel rods. Table 2 provides a summary of changes in the Cycle 8 kinetics characteristics compared with the current limit based on the reference safety analysis.(2, 5, 6) It can be seen from the table that all of the Cycle 8 values fall within current limits with an exception in the delayed neutron fraction and maximum differential rod worth noted in Table 2. Table 3 provides the control rod worths and requirements at the most limiting condition during the cycle. The required shutdown margin is based on accident analyses presented in Section 5.0. During extended cycle
. operation, the control rod reactivity requirements are reduced primarily by a reduction in the Doppler defect, xenon (fission product) nuclide concentration, and moderator temperature defect from the nominal end-of-life conditions. However, due to a more negative ASI, the rod insertion allowance increases. The available shutdown rod worth remains relatively constant during the Cycle 8 extension, and the net result 'is a slight increase in shutdown margin from the beginning to the end of the cycle extension period.
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5.0 ACCIDENT ANALYSIS
5.1 INTRODUCTION
AND
SUMMARY
The power capability of Millstone Unit 2 is evaluated considering the consequences of those incidents examined in the BSR(2) , using the previously accepted design basis specified in Section 1.2. It is concluded that, with respect to the incidents considered in the BSR(2) the core reload and a ,
maximum of 17.6 percent steam generator tube plugging will not adversely affect the ability to safely operate at 100% of rated power during Cycle 8.
For the overpower transient, the fuel centerline temperature limit of 4700*F can be accommodated with margin in the Cycle 8 core. The burnup dependent densification model(3'4) was used for fuel temperature evaluations. The
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LOCA limit at rated power can be met by maintaining peak linear heat rates at .
or below 15.6 kw/ft.
5.2 ACCIDENT EVALUATION The effects of the reload and 17.6 percent steam generator tube plugging on the design basis and postulated incidents analyzed in the BSR(2) and updated in the previous reload safety analyses (5,6) were examined. In most cases, it was found that the effects were accommodated within the conservatism of the initial assumptions used in the previous applicable safety analysis. For those incidents which were reanalyzed, it was determined that the applicable design bases are not exceeded. Therefore, the conclusions presented previously are still valid.
A core reload can typically affect accident analysis input parameters in the following areas: core kinetic characteristics, shutdown margin, CEA worths, and core peaking factors. Cycle 8 parameters in each of these areas were examined as discussed below to ascertain whether new accident analyses were required.
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Additionally, it was necessary to assess the impact of 17.6 percent steam generator tube plugging and an associated reduction in the guaranteed thermal design flow rate to 340,000 gpm on the safety analyses documented in the BSR(2) The plugging of steam generator tubes leads to a reduction in the steam generator primary to secondary heat transfer area and a slight reduction in the RCS active volume as well as tne reduced flow. An evaluation or analysis is provided in Section 5.3 (Appendix A) for each of the appropriate transients considered in the BSR(2) , with respect to steam generator tube plugging. Appendix B gives the Technical Specification changes needed to accomodate the 17.6 percent steam generator tube plugging and an associated reduction in the thermal design flow. In all cases it was determined that the applicable design bases are not exceeded. Therefore, the conclusion presented previously are still valid.
E 5.2.1 KINETICS PARAMETERS A comparison of Cycle 8 kinetics parameters with the current limits, established by the BSR and previous reload safety analyses, is presented in Table 2. The maximum delayed neutron fraction and the least negative doppler power coefficient above 30% power exceed the current limit slightly. These parameters were evaluated, and the previous analysis was determined to be adequate.
I 5.2.2 SHUTDOWN MARGIN Changes in minimum shutdown margin requirements may impact the safety analyses, particularly the steamline break and boron dilution accident analyses. Cycle 8 shutdown margin requirements are the same as Cycle 7.
5.2.3 CEA WORTHS Changes in CEA worths may affect shutdown margin. Table 3 shows that the Cycle 8 shutdown margin requirements are satisfied.
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Table 2 shows that the maximum differential rod worth increased from 36.6 to 115 pcm/ inch. This change resulted in the CEA withdrawal from suberitical accident being reanalyzed, as documented in Appendix A.
5.2.4 CORE PEAKING FACTORS l
l l All core peaking factors for Cycle 8 were within the reference cycle limits.
5.3 INCIDENTS REANALYZED AND EVALUATED See Appendix A.
5.4 LOCA ANALYSIS l
The LOCA analysis provided in Reference 9 bounds the impact of the reduction in design flow and the steam generator tube plugging levels, except for extended cperation past the nominal full power end-of-life burnup. Extended operation is based on a minimum guaranteed flow of 350,000 gpm.
Operation at a reduced flow of 340,000 gpm is subject to NRC approval and licensing of the LOCA analyses (9,10) performed to support the reduced flow and increased steam generator tube plugging, in e- .o i 10
6.0 REFERENCES
- 1. Davidson, S. L. (Ed.), et al., " Westinghouse Reload Safety Evaluation I
Methodology", WCAP-9273-A, July, 1985.
- 2. Millstone Unit 2, " Millstone Unit 2 Basic Safety Report", Docket No.
50-336, March, 1980.
- 3. Miller, J. V. (Ed), " Improved Analytical Model used in Westinghouse Fuel Rod Design Computations", WCAP-8785, October, 1976.
- 4. Hellman, J. M. (Ed.), " Fuel Densification Experimental Results and Model for Reactor Operation", WCAP-8219-A, March 1975.
- 5. Letter, Opeka to Miller, Millstone Nuclear Power Station Unit No. 2, Cycle 7 Refueling - Reload Safety Analysis, June, 1985.
- 6. Letter, Opeka to Butcher, Millstone Nuclear Power Station Unit No. 2,
" Follow-up Actions to Amendment No. 90 to Operating License No. DTR-65,"
November 8, 1985.
- 7. George, R. A., et al., " Revised Clad Flattening Model," WCAP-8377 (Proprietary) and WCAP-8381 (Non-Proprietary), July 1974.
- 8. Docket No. STN 50-572, "RESAR-414, The Reference Safety Analysis Report Documenting the Westinghouse 3820 MWt Nuclear Steam Supply System," ,
October 1976.
- 9. Letter, Wilson to Honan, Millstone Nuclear Power Station Unit No. 2,
" Steam Generator Tube Plugging Study," March 18, 1986.
- 10. Letter, Wilson to Honan, Millstone Nuclear Power Station Unit No. 2 " Steam Generator Tube Plugging Study," February 27, 1986.
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TABLE 1 Millstone Unit 2 Cycle 8 Core Loading Initial 80C(3)
Number of Enrichment % Theoretical Burnup Average Region Type Assemblies w/o U235 Density (MWD /MTU) ,
F2(1) 1 3.30 94.9 21000 G2 8 3.19 94.7 23600 G2(1) 4 3.19 94.7 23000 H1 20 2.73 95.2 25600 H1(2) 4 2.73 95.2 12800 ,
H2 44 3.22 94.8 22900 J1 24 2.62/2.91 35.2/95.1 13900 J2 48 2.91/3.29 95.1/95.2 9500 K1 I4) 16 2.6/2.9 95.0/95.0 0 K2(4) 48 2.9/3.3 95.0/95.0 0 (1) Westinghouse fuel reassembled using CE skeletons.
(2) Reinserted from Cycle 6.
l (3) E0L Cycle 7 burnup assumed: 11,700 MWD /MTU.
(4) Region K1 and K2 em ichment and densities are nominal.
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TABLE 2 MILLSTONE UNIT 2 KINETICS CHARACTERISTICS J Current Limit Cycle 8 Most Positive Moderator Temperature Coefficient (ap/*F) x 10 ~4 +0.5 frem 0 to 70% Power +0.5 from 0 to 70% Power <
+0.4 from 70 to 100% Power +0.4 from 70 to 100% Power Most Negative Moderator Temperature
-4 -3.8 -3.8 Coefficient (ap/*F) x 10 , ARI Doppler Temperature Coefficient (ap/*F x 10-5) -1.2 to -1.92 -1.2 to -1.92 ~
Delayed Neutron Lifetime 6,ff (7.) .479 to .634 . G to . W Prompt Neutron Fraction (usec) <32.2 <32.2 Maximum Differential Rod Worth of two CEA control groups moving together or 1 CEA shutdown group 36.6 115.0 at HZP (pcm/in) 1 um ....a i n 13
TABLE 3 SHUTDOWN REQUIREMENTS AND MARGINS MILLSTONE UNIT 2 - CYCLE 8 Control Rod Worth (%ao) Cycle 7 Cycle 8 EOC E0C All Rods Inserted 8.60 8.66 All Rods Inserted Less Worst Stuck Rod 6.95 6.83 o'* ,
(1) Less 10 Percent 6.26 6.15 Control Rod Requirements (%ao)
Reactivity Defects (Combined Doppler, T,yg, Void and Redistribution Effects) 2.58 2.77 Rod Insertion Allowance 0.41 0.45 (2) Total Requirements 2.99 3.22 Shutdown Margin ((1) - (2)) (%Ap) 3.27 2.93 Required Shutdown Margin (%ap) 2.90 2.90 nm .. .. m 14 l
Figure 1 Core Loading Pattern Millstone Unit 2 Cycle 8 A B C D E F GH JK LMNP R S T V W X Y l l l l l l 'l
' u n u u 2
K2 K2 K2 J2 G2 J2 K2 K2 K2 K2 H1 J2 H2 H2 J2 H2 H2 J2 B1 K2 4
K2 32 K1 HI 32 J2 H1 J2 J2 H1 K1 J2 K2 5
K2 El K1 G2 J1 J1 H2 J1 H2 J1 J1 G2 K1 El K2 s
K2 32 H1 31 32 H2 K1 H1 K1 G2 J2 J1 H1 J2 K2 7-K2 H2 J2 J1 H2 J1 H2 H2 H2 J1 H2 J1 J2 G2 K2 s- K2 K2
- J2 J2 H2 J2 H2 K1 H2 J2 H2 J2 J2 H2 H2 K1 H2 t o-K2 K2 11 G2 J2 H1 J1 El H2 H2 F2 H2 H2 Hi J1 Hi J2 G2 12 - K2 K2 J2 H2 J2 H2 K1 H2 J2 H2 J2 G2 K1 H2 J2 H2 J2 14 - y y 15 K2 H2 J2 J1 H2 J1 H2 H2 H2 J1 H2 J1 J2 H2 K2
' El J2 J1 Hi J2 K2 K2 J2 Hi J1 J2 H2 K1 K1 G2 17 K2 H1 K1 G2 J1 J1 H2 J1 H2 J1 J1 G2 K1 El K2 1s K2 J2 K1 El J2 J2 Hi J2 J2 El K1 J2 K2
" K2 Hi J2 H2 11 2 J2 H2 H2 J2 El K2 20 K2 K2 K2 J2 G2 J2 K2 K2 K2 21 K2 K2 K2 K2 Initial Building North Region # Assy w/o U235 F2 1 3.30 G2 12 3.19 HI 24 2.73 H2 44 3.22 J1 24 2.62/2.91 J2 48 2.91/3.29 K1 16 2.6 /2.9 K2 48 2.9 /3.3 15
// //
// %
4 % % '/
MW/V VO44 .I 4444
% '/
// 4
%ZZ%
GM#M WWWh
% b % %
M 2 & 4 444 / '/ %%%%
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FIGURE 2 ZONED-ENRICHMENT FUEL ASSEMBLY LATTICE 16 h..
8 0 Corresponds to Nominal EOL 1.05 1
560 Pcwer 0.85 550
$ \
Temperature y
[ 0.90 540 $
$ k \ !
N N s 0.85 3 530
\ \
w 0.80
\ 520 0.75 510
-200 0 200 400 600 800 1000 ABURNUP PAST NOMINAL EOL (MWD /MTU)
FIGURE 3. MILLSTONE UNIT-2, CYCLE 8 EXTENDED OPERATION CORE POWER and INLET TEMPERATURE vs. BURNUP 17
APPENDIX A ,
MILLSTONE 2 CYCLE 8 NON-LOCA SAFETY EVALUATION FOR 17.6% STEAM GENERATOR TUBE PLUGGING ,
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non..."*" A-1
I. INTRODUCTION The non-LOCA transients have been reviewed to determine the impact of steam generator tube plugging and an associated reduction in the thermal design flow to 340,000 gpm on Millstone Unit 2. The level of steam generator plugging assumed in the non-LOCA analyses / evaluations is a maximum of 17.6%.* The plugging of steam generator tubes leads to a) the reduction of steam generator heat transfer areas between the primary and secondary side, b) a reduction in the thermal design flow rate to 340,000 gpm and, c) a slight reduction in the RCS active volume. Table A.1 provides the design parameters used as the basis for this evaluation. These changes affect the non-LOCA analyses. The following transients were analyzed or evaluated:
- 1. CEA Withdrawal from Subcritical .
- 2. CEA Withdrawal at Power
- 3. RCS Depressurization fi
- 4. Loss of Flow
- 5. CEA Ejection
- 6. Malfunction of a Single Steam Generator
- 7. Steam Line Break
- 8. Boron' Dilution
- 9. Excess Heat Removal Due to Feedwater Malfunction
- 10. Startup of an Inactive Loop
- 11. Excess Load Event
- Maximum for Each Steam Generator sa m ....c.i A-2
- 12. Loss of Load and/or Turbine Trip
- 13. Loss of Normal Feedwater
- 14. CEA Drop
- 15. Single RCP Seized Rotor For each transient analyzed, a brief description of the accident and the effect of steam generator tube plugging is given in the attached report. The assumptions and key input parameters are similar to those used in the BSR(A.1] , except as noted. Based on this evaluation, operation with up to' 17.6% steam generator tube plugging and the resultant reduction in thermal design flow to 340,000 gpm will not result in violation of the safety limits. l II. SAFETY EVALUATION 11.1. CEA WITHDRAWAL FROM A SUBCRITICAL CONDITION DESCRIPTION OF THE ACCIDENT A control element assembly withdrawal incident when the reactor is subcritical results in a uncontrolled addition of reactivity leading to a power excursion. The nuclear power response is characterized by a very fast rise terminated by the negative reactivity feedback of the Doppler coefficient.
The power excursion causes a heatup of the moderator. However, since the power rise is rapid and is followed by an immediate reactor trip, the moderator temperature rise is small. Thus, nuclear power response is primarily a function of the Doppler coefficient.
The reduction in primary coolant flow influences this accident. The reduced primary coolant flow results in a decreased core heat transfer coefficient which in turn results in a faster fuel temperature increase. The fast temperature increase would result in more Doppler feedback thus reducing the 1
um e-no i A-3 l
nuclear peak heat flux excursion, which would partially compensate for the flow reduction. Therefore, the nuclear transient _is only moderately sensitive to the impact of a reduction in primary coolant flow.
METHOD OF ANALYSIS This transient was reanalyzed by two digital computer codes. The TWINKLE IA'33 Code is a multi-dimensional neutron kinetics computer code used to calculate the reactivity transient and hence the nuclear power transient.
The TWINKLE Code was previously used to analyze the CEA ejection transient in the BSR(A.1] . The FACTRAN[A.4] Code is then used to calculate the thermal heat flux transient based on the nuclear power transient calculated by the TWINKLE [A.3] Code. FACTRAN(A.4] also calculates the fuel and clad temperatures. Finally, the THINC( A.5] Code is used to calculate the minimum DNBR during the transient using the heat flux from FACTRAN( A.4] . The principal assumptions employed in the analysis are identical to those t.cesented in the BSR(A.1] ,
PROTECTION SYSTEM The reactor trip was assumed to be initiated by the variable high power level trip. A 10 percent increase was assumed for the variable high power level trip setpoint raising it from the nominal value of 15 percent to 25 percent of full power.
RESULTS The analysis demonstrates that for a 76.7 pcm/sec reactivity insertion rate, the peak core average heat flux achieved was 46.5 percent of nominal, the peak average fuel temperature was 1625'F, and the peak clad inner temperature was 744*F. The results of the analysis are tabulated in Table A.2. Plots of nuclear power, core average heat flux, and hot spot fuel and clad temperatures vs. time may be found in Figures A.1 and A.2. It is concluded that the DNB design basis will be met That is, DNB will not occur on at least 95 percent of the limiting fuel rods at a 95 percent confidence level.
um -esos" A-4
11.2. CEA WITHDRAWAL AT POWER DESCRIPTION OF THE ACCIDENT The CEA withdrawal at power accident is a continuous, uncontrolled CEA withdrawal at power due to a faulty operative action or a malfunction of the reactor instruments. The incident results in an increase in the core heat flux. The heat removal rate of the steam generator lags behind the increase in core heat flux. Thus, the reactor coolant temperature increases. The increase in core heat flux and RCS coolant temperature, if not controlled, could result in.DNB.
METHOD OF ANALYSIS This transient was reanalyzed to determine the effect of the reduced reactor coolant flowrate on the minimum DNBR. The accident was modeled using the LOFTRAN(A.2] computer code for cases of maximum and minimum reactivity feedback.
PROTECTION SYSTEM To protect against DNB, two protection systems were assumed to be functioning in the analysis. The protection systems were the variable high power level trip and the thermal margin / low pressure trip functions.
RESULTS The analysis showed that the accident causes an increase in core heat flux which causes an increase in core average temperature. The reactor was tripped by the high nuclear power level (112% RTP) trip function. The margin to DNB was maintained such that the DNBR limit of 1.30 was not violated. Hence it is
! concluded that adequate protection exists to protect the core against DNB with the reduced thermal design flow rate of 340,000 gpm resulting from the steam I generator tube plugging.
1 ca s-esai: A-5 l
II.3.RCS DEPRESSURIZATION DESCRIPTION OF THE ACCIDENT The most severe core conditions for this accident result from the simultaneous opening of both pressurizer relief valves. The valve opening causes a <
decrease in RCS pressure until the pressure reaches the fluid saturation pressure. During the transient, the reactor coolant temperature decreases slowly, while the pressurizer level increases until a reactor trip occurs. [
The decrease in pressure, if uncontrolled, could result in DNB.
METHOD OF ANALYSIS
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The transient was reanalyzed to determine the effect of an increased steam generator plugging level and the resultant reduction in thermal design flow on the minimum DNBR of the transient. The accident was modeled using the LOFTRAN(A.2] computer code for the case of an RCS depressurization accident subject to the same conservative assumptions presented in the BSR(A.1] ,
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PROTECTION SYSTEM The reactor is tripped by the thermal margin / low pressure trip (TM/LP) function.
RESULTS The results of this analysis indicate that the TM/LP trip provides adequate protection against DNB. The sequence of events for this transient is summarized in Table A.3. When subjected to the most limiting conditions, adequate margin to DNB was calculated. That is, the DNBR was in excess of the limit value, 1.30. Hence, it is concluded that the reactor is adequately l protected against DNB with the 340,000 gpm thermal design flowrate.
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k uon e-e o.i A-6
11.4. LOSS OF FLOW DESCRIPTION OF THE ACCIDENT I
The loss of all reactor coolant pump forced flow may occur as a result of a loss of electrical power to the pumps. The effect of the loss of flow on core conditions is an immediate increase in coolant temperature and pressure. The heatup may result in DNB if the core is not tripped promptly.
METHOD OF ANALYSIS The transient was reanalyzed to determine the effect of a reduced thermal design flowrate on the margin to DNB. The accident was modeled using three computer codes, LOFTRAN(A.2) , FACTRANIA*43, and THINC(A.5] employing the same conservative assumptions presented in the BSR(A.1] . The most limiting transient results when 4 pumps coast down simultaneously.
PROTECTION The necessary protection against a loss of forced reactor coolant flow is provided by the reactor coolant pump speed sensor and the low reactor coolant loop flow sensor.
RESULTS The reactor coolant flowrate was assumed to coast down (all pumps) exponentially beginning at time, t=0 sec. The low reactor coolant.flowrote sensor initiated the reactor trip at 0.915 seconds. The minimum DNBR was obtained after the trip signal. The peak fuel average temperature for the hot channel was equal to 1999'F which occurred at 1.8 seconds. Within the same time frame, the nuclear power achieved a maximum value of 102.8% of the nominal power level. A summary of the events for this transient are given in Table A.4. Plots of key parameter results are found in Figures A.3 through A.6. The minimum DNBR calculated was in excess of the limit value. Hence it
(
is concluded that adequate protection exists for the loss of flow event under the reduced flow conditions. ,
u m e
- cein A-7
-11.5. CEA EJECTION DESCRIPTION OF THE ACCIDENT This accident is defined as the mechanical failure of a control rod mechanism pressure housing resulting in the ejection' of a control element assembly (CEA) and drive shaft. . The consequence of this mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage. The core power rise is limited by the Doppler feedback effect, and the transient is terminated by a reactor shutdown following receipt of a high power level reactor trip signal.
A conservative limiting criterion was selected for this accident which applies to all fuel types expected to be inserted in the core in Cycle 8 and subsequent cycles. This criterion is:
Average fuel pellet enthalpy at the hot spot is to be below 200 cal /gm for unirradiated or irradiated fuel.
s -
METHOD OF ANALYSIS The transient was reanalyzed to demonstrate that a reduction in thermal design flowrate does not result in exceeding the 200 cal /gm peak fuel enthalpy.
limit. The analysis was performed using two computer codes, TWINKLE (A.3) and FACTRAN(A.4] ,
PROTECTION SYSTEM The reactor trip was assumed to be initiated by the variable overpower trip at a core nuclear power level of 25 percent for the hot zero power and 112 percent for the full power case.
u m e-e. win A-8
~
54 s Q
_y -
RESULTS - -
The analysis was performed for both zero and full power.- The key input parameters and results are presented _in Table A.5. The criteria discussed above, and presented in the BSR(A.11, are not exceeded. The average pellet enthalpy of the hottest fuel pellet does not exceed the expected clad damage threshold of 200 cal /gm. ,
II'.6. MALFUNCTION OF ONE STEAM GENERATOR DESCRIPTION OF THE ACCIDENT
.The loss of load to one steam generator is assumed to' occur due to the closure of the main steam line isolation valve to one steam generator. As a result, the temperature and pressure in the isolated steam generator, rise until the safety valves lift. Simultaneously, the unisolated steam generator attempts to compensate for the load imbalance resulting in an RCS inlet temperature imbalance. The colder water from the isolated steam generator produces a positive reactivity insertion in the presence of a negative moderator temperature coefficient. A power increase results which normally trips the reactor on a high power level. However, this trip is not assumed to activate in this analysis. The reactor is assumed to trip on low steam generator water level.
'NETH00 0F ANALYSIS This transient was reanalyzed to determine the effect of the reduced reactor coolant flowrate on the margin to ONB. The analysis was performed by modeling the system using the digitaf computer code LOFTRAN(A.2] . The analysis
. employed-the same highly conservative assumptions presented in the BSR(A.1] ,
PROTECTION SYSTEM
.TopredludeDNB,thereactorwasassu_medtotriponlowsteamgeneratorwater ievel'. ,
4.,...L O ,_g
,1 1
RESULTS The loss of load to one steam generator is initiated at t=0 sec. The core inlet temperature and steam pressure in the isolated steam generator loop increase to maximum values of 566 F and 1030 psia while the temperature and
! pressure of the unisolated steam generator loop decreased to 506 F and 497 psia. The influx of the cold water from the isolated steam generator into the core causes the reactivity to increase in.the presence of a negative moderator temperature coefficient. The nuclear power attains.a maximum value of 130.8%
of the nominal power. The reactor was assumed to trip on low steam generator water level. The sequence of events is given .in Table A.6. A minimum DNBR in excess of the limit 1.30 was attained. Hence it is concluded that adequate.
protection against DNB exists for this transient at the reduced thermal design flow conditions. ,
).
11.7. STEAMLINE BREAK 1
DESCRIPTION OF THE ACCIDENT i
l l A steamline break may result from a stuck open safety or relief valve or a ruptured steamline pipe. This results in a rapid depressurization of the 1
steam generators which, in the presence of a negative moderator temperature l coefficient, causes a large positive reactivity insertion to the core via the
! primary side cooldown. As discus' sed in the BSR[A. U , the acceptance criteria for this event is that DNB must not occur following a return to power.
1 r
EVALUATION 3 The impact of increased steam generator tube pluoging level would affect the
( transient primarily due to the. reduced flow, reduced RCS inventory and the reduced. heat transfer coefficient. These impacts would result in changed RCS cooldown and feedback characteristics such that the return to power as presented in the BSRE ^*Il would be more conservative with respect to the lower initial RCS flow. The reduced RCS flow and corresponding reduced heat sam e-eemu A-10 ,
transfer coefficient result in a less severe cooldown of the primary side.
This in turn results in a lower return to power and a less limiting DNBR.
Previous analyses performed for steam generator tube plugging that caused a reduction >in thermal design flow from 370,000 to 350,000 gpm confirmed this.
To provide additional confirmation of the adequacy of the current steamline break analysis, a peaking factor evaluation was performed. The existing transient statepoint was reanalyzed using Cycle 8 peaking factors along with the reduced 340,000 gpm flow resulting from the steam generator tube plugging.
It was found that the DNB design basis was met for this case. Hence it is concluded that adequate protection against DNB exists for this transient at the reduced thermal design flow conditions.
11.8. UNCONTROLLED BORON DILUTION DESCRIPTION OF THE ACCIDENT Reactivity can be added to the core by feeding primary grade water into the ,
RCS by the CVCS (Chemical and Volume Control. System). Boron dilution is a manual CVCS operation. The CVCS is designed to limit, even under various failures, the rate of dilution to provide the operator adequate time to terminate the dilution event.
t EVALUATION h
In order to bound all phases of plant operation, boron dilution during refueling, startup and full power operation was analyzed. The time to reach criticality, delta terit, is calculated by the following expression:
C y Initial delta t erit *Q I" C
critical
,;y L,
non e soosi:
A-11 L
~ . . - .- - - - - -
)
where, 3
V = Active RCS volume (ft )
, Q' = Maximum charging flow (ft3 /sec)
'C initial
= Initial boron concentration
.C critical
= Critical boron concentration The analysis bounding the various phases of plant operation was documented in the BSR(A.1] . ~The analysis computes the time to reach criticality, using the equation above, for each phase of operation and compares that value to the
~
minimum value to reach criticality. In each case a significant safety margin was demonstrated.
The effect of increased steam generator tu.be plugging is a reduction .in the active-RCS volume. The reduction in volume affects only those analyses that assume the RCS to be filled. The change in primary volume is equal to the
( . change in the primary steam generator volume. The absolute relative change in i ~ primary volume is less than 5%.
If the shutdown margin can be maintained with a 5% reduction in primary mass, it can be concluded that the increased level of steam gencrator tube plugging has no adverse effect on tM bcron dilution accident The time to criticality for a 5% reduction in primary mass (or volume) for each mode of operation analyzed in the BSR(A.1] is given in Table A.7. As demonstrated, the time
-to criticality for each event is in excess of the minimum required time.
Therefore, the increased level of steam generator tube plugging will have no adverse impact on the boron dilution accident.
[
For dilution at power, it is necessary that the time to lose shutdown margin i be sufficient to allow identification of the problem and termination of the dilution. As in the dilution during startup case, the RCS volume reduction due to steam generator tube plugging must be considered. With the reactor in 5 manual control, the slight reduction in RCS volume has no effect on the effective reactivity insertion rate. Therefore the previously safety margins are maintained.
no x e-. o ' A-12
11.9'. EXCESS HEAT REMOVAL DUE TO FEEDWATER MALFUNCTION
-This' excess ~ heat removal due to a feedwater malfunction transient was-not analyzed since it is bounded by the limiting cooldown event, the steam line break.
11.10. STARTUP OF AN INACTIVE REACTOR COOLANT PUMP As discussed in the BSR(A.1] , the startup of an inactive reactor coolant pump transient was not analyzed since the Technical Specifications do not i permit operation with less than four reactor coolant pumps.
l-l 11.11. EXCESSIVE LOAD EVENT The excessive load transient was not analyzed since'it is bounded by the limiting cooldown event, the steam line break. The conclusions presented in the BSR[A.1) remain valid.
11.12. LOSS OF i_0AD AND/0R TURBINE TRIP f
l DESCRIPTION OF THE ACCIDENT l
The result of a loss of load is a core power level which momentarily exceeds the secon.dary system power extractic, causing an increase in RCS pressure and water temperature.
{
EVALUATION The. impact.of increased levels of steam generator tube plugging on the loss of load analysis would be principally due to the reduced RCS flow and the decreased RCS mass inventory. Two cases, analyzed for beginning of life and end of life conditions, are presented in Section 5.3.5 of.the BSR[A.1] ,
The BSR[A.1) analysis results in a peak pressurizer pressure of 2573 psia following' reactor trip and a minimum DNBR which never decreases below its initial value. A reduction in loop ~ flow and RCS mass inventory will result in m x e-e m i A-13
a more rapid pressure rise than is currently shown. The effect on peak RCS pressure will be minor, however, since the reactor is tripped on high pressurizer pressure and the time to trip will be decreased which will result in a lower total energy input to the coolant. The initial margin to DNB may be reduced, however, as in the analysis presented in the BSR[A.1] , the DNBR will increase throughout the transient and always remains above the limit value. Thus, the conclusions presented in the BSR[A.1] remain valid.
11.13. LOSS OF NORMAL FEEDWATER
.l DESCRIPTION OF THE ACCIDENT The loss of normal feedwater accident results from a failure of the secondary system to remove heat from the core. To prevent damage to the core, an auxiliary feedwater system is employed. If the auxiliary feedwater system were not used, the primary system water would continue to heat and expand into the pressurizer-causing the pressurizer pressure to increase. A significant loss of RCS inventory could lead to core damage.
EVALUATION The effects of reducing core flowrate would result in a more rapid rise or heat up of the primary system which increases the volume of water in the pressurizer. Sensitivity studi,es have been performed which conservatively estimate that the expected increase in pressurizer volume would be negligible for a plugging level of 17.6% and a reduction in thermal design flow to 340,000 gpm. Thus, adequate margin to filling the pressurizer still exists.
Therefore, it was concluded that no additional analysis was required and the conclusions presented in the BSR[A.1] remain valid.
~~~
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A-14
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a h.1 DESCRIPTION OF THE ACCIDENT %; M. : . :,
, y.
..64 .':-; } -
g.~
g In this accident, the control element assembly is dropped into the core at 4it,h,cif
.m.n ,
G. power. As a result, the reactor becomes suberitical and the neutron flux w.* :-(:/;,j,.
n .r . .
D. level decreases. If the reactor is not tripped, the mismatch between the core 4. .. .-
l [ power level and, turbine power causes a decrease in the RCS water tempera-
'M,4.Q[
i.. ture. If a negative moderator temperature coefficient is assumed, the .d.f
.r .
[ temperature decrease adds positive reactivity to the core. If the core f . ,v... b ' ,;
it. . . . . - .
9 returns to power, the power distribution may cause the thermal limits to be y J.i
- b. ; exceeded. f[%[.~:w .:. %
?'
~
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EVALUATION kk~C.D W. t 7 ,Q v .> .,. ,.
5 ~- The effect of decreasing the thermal design flowrate and increasing the degree '.: i.~ .y.M[
- e. , 1 ~ . . ..
pg of steam generator tube plugging is a reduction in the RCS temperature and a ,2
.[.
slight power increase. This transient was originally analyzed in the :,
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N..
BSR[A*1] Subsequently, it was reanalyzed to evaluate the effect.of a 5% ~ %". y' 9"
~
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.:.s..
reduction in thermal design flow (from 370,000 to 350,000 gpm). The results of 3p;y%;yf. ,. .
h[ r.
that analysis were compared to those of the BSR[A.1] , and showed a small
< [k?.h sc
decrease below its initial value. It was concluded that the BSR[A*1] M'
.C" a .
.y" [T , [." -
~1 analysis was more limiting. Similarly it is expected that the additional
,y .
reduction in flow to 340,000 gpm will yield a decrease in the RCS minimum gg(@.
i' f temperature and pressure.' [ P-O "'
cip p,;:
?]';,
h[ .c; . l' *.5 To provide additional confirmation of the adequacy of the current analysis, a %.
h..C . . f ;. .V-.
'i peaking evaluation was performed. The existing statepoint was reanalyzed using "'
%0+
-a Cycle 8 peaking factor along with the reduced 340,000 gpm flowrate resulting i
.; from the steam generator tube plugging. The results showed that the DNB design @;bh^-
5 basis was met. Hence, it is concluded that the existing BSR[A*1] analysis 7 -
sufficiently bounds this accident for the Cycle 8 reload with a reduced
- s. ;:.. _Ah Q;@. p. - -
3 gy;M%
M. thermal design flowrate of 340,000 gpm. - %f y k. . M .
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11.15. SINGLE RCP SEIZED ROTOR DESCRIPTION OF THE ACCIDENT o The most severe Locked Rotor Accident is due to an instantaneous seizure of a reactor coolant pump rotor at 100% power. The seizure causes a reactor trip on a low flow signal.
EVALUATION Following the trip signal, the residual heat in the fuel rods transfers to the coolant causing the coolant to expand. Simultaneously, the reduced flow decreases the heat transfer in the steam generator. As a result of the -
expansion of the coolant in the RCS combined with the reduced heat transfer in the steam generator, flow surges into the pressurizer. This insurge causes an ..
increase in RCS pressure. The pressure increase activates the automatic spray system and opens the relief valves.
'The reduced thermal design flow rate could affect several events of the transient. The areas that need to be addressed are the peak RCS pressure relative to the mechanical design limit and the peak clad temperature.
The impact on the locked rotor analysis of increased steam generator tube plugging will be primarily due to the reduced flow. These impacts will not affect the time to DNB since DNB is conservatively assumed to occur at the beginning of the transient. The flow coastdown in the affected loop due to -
the locked rotor is so rapid that the time at which the reactor trip low flow setpoint is reached is essentially identical that of the most recent analyses.
It is estimated that the peak pressure will not increase above the previous ._ - z value'due to reduced power, however the maximum calculated value was 2778 psia {g.g .'~$
based on an initial RCS pressure of 2250 psia plus 30 psia uncertainty. This 2
- is significantly below the pressure at which vessel stress limits are exceeded. 1 c:'"%if.
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g.,Wg 44 [;;$
re.
n:
sn e-esm: A-16 -h[M N W
B e a E 1 3ia
_ . . . . . . . EIMERB"i
The reduction in the primary side steam generator tube volume would result in less than a 5 percent reduction in primary mass which decreases the heat capacity of the RCS by the same amount. This would not result in higher peak -
temperatures or pressures since the peak values are reached in considerably less than one-loop transport time constant.
To provide additional confirmation of the adequacy of the current locked rotor analysis, a peaking factor evaluation was performed. The current statepoint was analyzed using Cycle 8 peaking factors, along with the reduced 340,000 gpm flowrate resulting from the steam generator tube plugging. The number of rods in DNB predicted to occur by this calculation remained less than the current limit value of 3.0 percent.
III. CONCLUSION This evaluation has discussed the effects on the non-LOCA licensing basis analyses of 17.6% steam generator tube plugging level. The primary effects are: a reduction in the thermal design flow to 340,000 gpm, a reduction in the primary to secondary heat transfer area, and a slight reduction in the RCS mass inventory. The results of this evaluation demonstrate that the Millstone ,
Unit 2 non-LOCA licensing basis analysis conclusions, as presented in the -
BSR[A.1] , remain valid with a maximum of 17.6% steam generator tube plugging. - -
N uex.-eesi:
A-17
REFERENCES A.1. Millstone Unit 2, " Millstone Unit 2 Basic Safety Report", Docket No.
50-336, March, 1980.
A.2. Burnett, T. W. T., et al, "LOFTRAN Code Jescription", WCAP-7907, June, 1972.
A.3. Risher, D. H., Jr. and Barry, R. F., " TWINKLE - A Multi-Dimensional Neutron Kinetics Computer Code", WCAP-7979-P-A, January, 1975 (Proprietary) and WCAP-8028-A, January, 1975 (Non-Proprietary).
,A.4. Hunin, C., "FACTRAN, A FORTRAN IV Code for Thermal Transients in a UO Fuel Rod", WCAP-7908, June, 1972.
2 A.S. THINC code discussion in RESAR-3S, Section 4.4.3.2.
A.6. Letter, Wilson to Honan, Millstone Nuclear Power Station Unit No. 2,
" Steam Generator Tube Plugging Study," March 18, 1986.
s2ox - .o.'2 A-18
D TABLE A.1 ,
DESIGN PARAMETERS USED IN THE MILLSTONE UNIT 2 STEAM GENERATOR TUBE PLUGGING EVALUATION Power Level 2700 MWth HFP Inlet Temperature 549 deg F Thermal Design Flow 340,000 gpm Primary Pressure 2250 psia Steam Generator Pressure 883 psia Feedwater Temperature 432 deg F g
Steam Generator Tube Plugging Level 17.6%
L srex -swai:
A-19
TABLE A.2 SEQUENCE OF EVENTS FOR THE ROD WITHDRAWAL FROM SUBCRITICAL EVENT Time (sec) Event Value Initiation of Uncontrolled Rod Withdrawal 76.7 pcm/sec Reactivity f Insertion Rate from 10E-13 of 0.0 Nominal Power 9.09 Variable Power Range High Neutron Flux Setpoint Reached 25% RTP k 358% of Nominal I
9.20 Peak Nuclear Power ;
i 9.59 Rods Begin to Fall _
10.51 Peak Clad Temperature 744*F 10.68 Peak Core Average Heat Flux 46.5% of Nominal 11.31 Peak Average Fuel Temperature 1625 F 11.86 Peak Fuel Center Temperature 2334*F-k s2m .-.mi:
A-20 L_.
TABLE A.3 SEQUENCE OF EVENTS FOR THE RCS DEPRESSURIZATION ACCIDENT MAXIMUM REACTIVITY INSERTION RATE b
Event Time (sec) l Opening of two p essurizer Relief Valves 0.09 Reactor Trip 36.42 ,
Minimum DNBR 36.50 MINIMUM REACTIVITY INSERTION RATE Event Time (sec)
Opening of two pressurizer Relief Valves 0.05 Reactor Trip 36.69 _
Minimum DNBR 37.00
(
um e-eeoei A-21
TABLE A.4 SEQUENCE OF EVENTS FOR THE LOSS OF FLOW ACCIDENT Event Time (sec)
Loss of Power to all Pumps, Coastdown Begins 0.0 Reactor Coolant Pump Low Speed Setpoint Reached 0.99
{
(91.5 percent of nominal, 340,000 gpm) .. - , -
CEA's Begin to Drop 1.64 Minimum DNBR Occurs 3.1 9
sax e-""
A-22 1
l
TABLE A.5.
PARAMETERS AND RESULTS OF THE CEA EJECTION ANALYSIS Parameter HZP HFP Power Level, percent of nominal 0 102 Ejected Rod Worth, pcm 650 230 Delayed Neutron Fraction, percent 0.55 0.55 Feedback Reactivity Weighting 2.73 1.30 Trip Reactivity, pcm' 2.55 4.23 Fg before rod ejection --
2.64 Fg after rod ejection 19.1 5.70 Number of operating reactor coolant pumps 2 4 Results Max. pellet average temperature, deg F 2979 4008 Max. centerline temperature, deg F 3416 4961 Max. clad temperature, deg F 2202 2234 Max. pellet melting, percent 0.0 6.18 Max. fuel stored energy, cal /gm 122.8 174.7 um e-sweir A-23
TABLE A.6 SEQUENCE OF EVENTS FOR THE MALFUNCTION OF A SINGLE STEAM GENERATOR Event Time (sec)
Isolation of one steam generator 0.0 Safety valves open on isolated S/G 5.2 High. power trip reached (1.12%) 10.7 Low steam generator level trip 20.0 CEA's begin to fall 20.9 Minimum DNBR occurs 21.2 l
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cold shutdown Shutdown margin lost 19 15 4"Px R i ;'
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.@ APPENDIX B T i .;4 l:-p7.li e-y...,. , , 3..;
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A TECHNICAL SPECIFICATIONS CHANGE PAGES S TM .' 1:_,<' .
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Margin Safety Limit - Four Reactor E. 3
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f Coolant Pun.ps Operating .
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TABLE 2.2-1 REACTOR PROTECTIVE INSTRINENTATION TRIP SETP0lHT Ll?ilTS
$ TRIP SETPOINT ALLOWABLE VALUES E FUNC NAL UNIT Not Applicable Not Applicable
- 1. Man (actorTrip r a p Z 2. Power Level- gh u s
< 9.6% above THERHAL POWER, 1 9.7% abov ERMAL POWER, with f.
FourReactorCooktPumps with a minimum setpoint of a minlm
^
Operating y N < 14.6% of RATED THERt1AL THERMA
.7%
OWER~< 14.7% of RATED of and a maximum of RATED THERMAL POWER.
f F0WER, and a maximum of s,\ < 106.6% of RATED THERMAL
~<1 T:
F0WER. O l \sN E.
N -.
- 3. Reactor Coolant Flow - Low (1) \N M 1
Four Reactor Coolant Pumps > 91, of reactor co nt > 90.1% of reactor coolant flow m
Tiow w 4 pumps o rating *. with 4 pumps operating *.
L Operating
> 830 rpm > 823 rpm
- 4. Reactor Coolant Pump ,
Speed - Low Pressurtzer Pressure - High < 240,Vpsia 1 2408 psia
- 5. _
75 psig < 5.23 psig
- 6. Containment Pressure - High ,.f
" p
/ > 500 psfa '-> 492 psia
- 7. Steam Generator Pressure - ~
f.
Low (2) (5) p./ N.
r
> 36.0% Water Level - each > 3h{% Water level - each steam E 2 8. Steam Generator Water ~
genera'to E Level - Low (5) team generator B.
. gg
/
M /
Trip setpoint adjusted to not Trip setpot t adjusted to not n
." 9. Local Power 0 niity - High (3) exceed the lim lines of w exceed the Ilmit lines of Figures 2.2-1 an .2-2 (4). ?
" Figures 2.2-1 and 2.2-2 (4). -
- ./ $
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- esign Reactor Coolant f. low with 4 pumps operating is 350,033 qpm.
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TADLE 2.2-1 REACTOR PROTECTIVE INSTRUMENTATION TRIP SETPOINT LIMITS
{
TRIP SETPOINT ALLOWABLE VALUES FUNCTIONAL UNIT Manual Reactor Trip Not Applicable Not Appitcable 1.
z Z 2. Power Level-High Four Reactor Coolant Pumps 1 9.6% above THERHAL POWER, 1 9.7% above THERMAL POWER, with l
Operating with a minimum setpoint of a minimum of i 14.7% of RATED THERMAL POWER, and a maximum of
< 14.6% of RATED THERHAL -
F0WER, and a maximum of i 106.7% of RATED THERMAL POWER.
< 106.6% of RATED THERMAL F0WER.
- 3. Reactor Coolant Flow - Low (1)
Four Reactor Coolant Pumps > 91.7% of reactor coolant > 90.1% of reactor coolant flow m Tiow with 4 pumps operating *. with 4 pumps operating *.
i Operating
> 830 rpm > 823 rpm
- 4. Reactor Coolant Pump Speed - Low Pressur1zer Pressure - Hfgh 1 2400 psia i 2408 psfa 5.
Containment Pressure - High 1 4.75 psig 1 5.23 psfg 6.
> 500 psia > 492 psta a 7. Steam Generator Pressure - .
2 Low (2) (5)
?. .
> 36.0% Water Level - each > 35.21 Water Level - each steam m 8. Steam Generator Water _
generator Level - Low (5) steam generator P
Local Power Density - High (3) Trip setpoint adjusted to not Trip setpoint adjusted to not
- 9. exceed the limit lines of exceed the limit Ilnes of Figures 2.2-1 and 2.2-2 (4). Figures 2.2-1 and 2.2-2 (4).
- Design Reactor Coolant f.luw with 4 pumps operating is 340,000 gpn
. _ . - . . . . ._.- .. .. - . . . - _ . ~ , _ . . ._ . . . . _ _ - . . . _ -
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MILLSTONE - UNIT 2 3/4 2-8(a) Amendment No. 38.52473.9D.91. 99 . . , .
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l r \ I l 1 I II ,
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, j Q O.8 -- ( 1. 5 5 0. 0. 8) -g j ,g
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~ (1.675, 0.675) --
(1,838, 0.6 5)T y f ,
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< NY e
1 0.4 w
J m
a 3 0.2 M O; ACCEPT ABLE OPER ATION REGION
=
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P 0.0 g 1.52 1.58 1.64 1.70 1.76 1.82 1.88
- T T R-o F
r: F xy (F x (1 + Tq) : F r xy x (1 + Tq))
=
4 TD 5
FIGURE 3.2-3b Total Radial Peaking Factor vs. Allowable Fraction of RATED THERMAL POWER w
to p e -
4 f ff June 19, 1985
')
POWER DISTRIBUTION LIMITS _
T r s OTAL INTEGRATED RADIAL PEAKING FACTOR -F L TING CONDITION FOR OPERATION 3.2.3 hecalculatedvalueofFf,definedasFT7= F 7(1+Tq ) shal be limited s to < l.
- 5. ,
APPLICABIL Y: MODE 1*.
ACTION:
With Ff > 1.565, ithin 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> either:
L POWER to bring the combini[ ion of THERMAL POWER
- a. Reduce TH and FT7 to w hinthelimitsofFigure[2-3bandwithdrawthe full length C to or beyond the Lo6g Tem Steady State Insertion Limits of Spec cation 3.1.3.6; or f
. /
.c b.
Be in at least SURVEILLANCE REQUIREMENTS
/ H0NTANDBY.
4.2.3.1 The provisions of Speci ca ion 4.0.4 are r.ot applicable.
pression rFT=F(1+T)andFfshall 4.2.3.2 F r shall be calculatet[ by the r q be detemined to be within following int,ervals:
j fts limit at t
- a. Prior to oper tion above *i0 percen of RATED THERMAL POWER after each fuel lo ing,
- b. At least nce per 31 days of accumulated operation in' MODE 1, and
\
- c. Withi our hours if the AZIMUTHAL POWER MLTq(T ) is > 0.020.
T 4.2.3.3 F r 11 be detemined each time a calculatio of Fr is required by using the core detectors to obtain a power distributto map with all full Limit for the length CE at or above the Long Tem Steady State Insert existin Reactor Coolant Pump Combination.
T required 4.2. 4 T qshall be determined each time a calculation of F 7 T
an the value of T qused to determine F7shall be the measured v ue of Tq.
v *5ee Special Text Exception 3.10.2. s Amendment No. 38,52,79,9 99 MILLSTONE - UNIT 2 3/4 2-9 I
a ) b i Juna 19, 1985 i
\
POWER DISTRIBUTION LIMITS T
TOTAL INTEGRATED RADIAL PEAKING FACTOR - F r LIMITING CONDITION FOR OPERATION 3.2.3 The calculated value of FT, defined as FT = F (1+Tq ) sh'all be limited 7 7 7 to i 1.537, 1 APPLICABILITY _: MODE 1*.
ACTION: l With FT> 1.537, within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> either: g 7
- a. Reduce THERMAL POWER to bring the combination of THERMAL POWER and Ff to within the limits of Figure 3.2-3b and withdraw the full length CEAs to or beyond the Long Term Steady State Insertion Limits of Specification 3.1.3.6; or
- b. Be.in at least HOT STANDBY.
SURVEILLANCE REQUIREMENTS i
4.2.3.1 The provisions of Specification 4.0.4 are not applicable.
4.2.3.2 Ff shall be calculated by the expression Ff = F r(1+Tq) and FT7 shall
- be determined to be within its limit at the following intervals:
- a. Prior to operation above 70 percent of RATED THERMAL POWER after each fuel loading,
- b. At least once per 31 days of accumulated operation in MODE 1, and
- c. Within four hours if the AZIMUTHAL POWER TILT (T q ) is > 0.020.
4.2.3.3 Fr shall be determined each time a calculation of FTr is required by using the incore detectors to obtain a power distribution map with all full length CEAs at or above the Long Tenn Steady State Insertion Limit for the existing Reactor Coolant Pump Combination.
4.2.3.4 Tq shall be determined each time a calculation of F T7is required and the;value of T qused to determine F Tshall be the measured value of T .
7 q
- See Special Text Exception 3.10.2.
MILLSTONE - t, NIT 2 3/4 2-9 Amendment No. M,52,77,99, 99 i ]
L_____.____
l l
w kdW w i k [Mowi n p Dscimbsr 30, 1 3 TABLE 3.2-1 DNB MARGIN LIMITS Four Reactor l Coolant Pumps l Para \me _r Operating Cold Leg Tempe e 1 549'F Pressurizer Pressu t 2225 psia
- Reactor Coolant Flow Rate t 350,000 g }
AXIAL SHAPE INDEX Figure 3 7-4
/
- Limit not applicable during eithe .a ERMAL POWER ramp increase in excess of 5% of RATED THERMAL POWER'^ er minute or a THERMAL POWER !
step increase of greater than 10% . RATED THERMAL POWER.
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MILLSTONE - UNIT 2 3/4 2-14 knendmentNo.38,JJ,yg, 90
T ABL E 3.2-1 DNB MARGIN LIMITS Four Reactor Coolant Pumps Parameter Operating Cold Leg Temperature 5,549'F Pressurizer Pressure > 2225 psia
- Reactor Coolant Flow Rate > 340,000 gpm l AX1AL SHAPE INDEX Figure 3.2-4
' Limit not applicable during either a THERMAL POWER ran.p increase in excess of 5% of RATED THERMAL POWER per minute or a THERMAL POWER step increase of greater than 10% of RATED THERMAL POWER.
s MILLSTONE - UNIT 2 3/4 2 14 Amendment No.