ML20234D813
| ML20234D813 | |
| Person / Time | |
|---|---|
| Site: | Vermont Yankee  |
| Issue date: | 05/31/1987 |
| From: | Paul Bergeron, Cacciapouti R, Sironen M VERMONT YANKEE NUCLEAR POWER CORP. |
| To: | |
| Shared Package | |
| ML20234D749 | List: |
| References | |
| YAEC-1600, NUDOCS 8707070345 | |
| Download: ML20234D813 (100) | |
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1 VERMONT YANKEE CYCLE 13 CORE PERFORMANCE ANALYSIS May 1987 Major Contributors:
V. Chandola J. D. Robichaud B. Y. Hubbard K. E. St. John M. P. LeFrancois T. A. Schmidt J. Pappas R. A. Woehlke R. C. Potter Prepared by: i >: / ^ n '
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M. A.dironen (Da t e')
Nuclear Engineering Coordinator Approved by: AA R. J. Calhiapouti, Manager (Date)
Reactor Physics Group Approved by: O P. A. Beheron, Manager (Date)
Transient Analysis Group Approved by: j JW -
L' I 5 N S. P'. Schultz, Manager ~ (Date)
LOCA Analysis Group Approved by: 7 B.'C. Slif(f, Director (Date)
Nuclear Engineering Department 2810R
DISCLAIMER OF RESPONSIBILITY This document was prepared by Yankee Atomic Electric Company for its .
own use and on behalf of Vermont Yankee Nuclear Power Corporation. This document is believed to be completely true and accurate to the best of our knowledge and information. It is authorized for use specifically by Yankee Atomic Electric Company, Vermont Yankee Nuclear Power Corporation and/or the appropriate subdivisions within the Nu.: lear Regulatory Commission only.
With regard to any unauthorized use whatsoever, Yankee Atcmic Electric Company, Vermont Yankee Nuclear Pcwer Corporation and their officers, directors, agents and employees assume no liability nor make any warranty or , ,
representation with respect to the contents of this document or to its accuracy or completeness.
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A ABSTRACT This report presents design information, calculational results, and operating limits pertinent to the operation of Cycle 13 of the Vermont Yankee Nuclear Power Station. These include the fuel design and core loading pattern ;
descriptions; calculated reactor power distributions, exposure distributions, shutdown capability, and reactivity functions; and the results of safety .
analyses performed to justify plant operation throughout the cycle.
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TABLE OF CONTENTS Page DISCLAIMER....................... .......................... ii l ABSTRACT.................................................... iii TABLE OF C0NTENTS........................-.................. iv LIST OF FIGURES............................................. Vi LIST OF TABLES.............................................. viii ACKN0WLEDGEMENTS............................................ ix
1.0 INTRODUCTION
................................................ 1 2.0 RECENT REACTOR OPERATING HIST 0RY............................ 2 2.1 0? orating History of the Current Cycle................. 2 2.2 Op+ rating History of Past Applicable Cycles............ 2 3.0 RELOAD CORE DESIGN DESCRIPTION.............................. 6
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3.1 Core Fuel Loading.................................. . 6 3.2 Design Reference Core Loading Pattern.................. 2 3.3 Assembly Exp'sure Distribution......................... 6 4.0 FUEL MECHANICAL AND THERMAL DESIGN.......................... 9 4.1 Mechanical Design...................................... 9 4.2 *ihermal Design......................................... 9 4.3 Operating Experience................................... 11 5.0 NUC LE Ai . DES I G N . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 5.1 Core Power Distributions............................... 16 5.1.1 Haling Power Distribution....................... 16 5.1.2 Rodded Depletion Power Distributer.............. 16 5.2 Core Exposure Distributions..................... ...... 17 5.3 Cold Core Reactivity and Shutdown Margin............... 17 5.4 Standby Liquid Control System Shutdown Capability...... 18
- 5 C.5anges in Nuclear Design Methods...................... 18 6.0 THERMAL-HYDRAULIC DESIGN.................................... 28 6.1 Steady-State Thermal Hydraulics........................ 28 6.2 Rea c to r Limi ts De t e rn.ina t ion . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
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TABLE OF CONTENTS (Continued)
Page 7.0 ACCIDENT ANALYSIS........................................... 30 7.1 Transient Analysis..................................... 30 7.1.1 Methodology..................................... 30 7.1.2 Initial Conditions and Assumptions,.............. 31 7.1.3 Reactivity Functions............................ 32 7.1.4 Transients Analyzed............................. 34 7.2 Transient Analysis Results............................. 34 7.2.1 Turbine Trip Without Bypass Transient........... 35 7.2.2 Generator Load Rejection Without Bypass Transient................................ 35 7.2.3 Loss of Feedwater Heating Transient............. 36 7.3 Overpressurization Analysis Results.................... 36 7.4 Local Rod Withdrawal Error Transient Results........... 37 7.5 Misloaded Bundle Error Analysis Results................ 40 7.5.1 Rotated Bundle Error............................ 40 7.5.2 Mislocated Bundle Error......................... 41 7.6 Control Rod Drop Accident Results...................... 42 8.0 LOSS-OF-COOLANT ACCIDENT ANALYSIS........................... 84 9.0 STARTUP PR0 GRAM............................................. 85 REFERENCES.................................................. 86 APPENDIX A CALCULATED OPERATING LIMITS.............. ...... A-1
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i LIST OF FIGURES Number Title Page 3.2.1 VY Cycle 13 Design Reference Loading Pattern, Upper Right Quadrant 8 4.2.1 VY Cycle 13 Core Average Gap Conductance versus Cycle Exposure 14 4.2.2 VY Hot Channel Cap Conductance for BP8X8R versus Exposure 15 5.1.1 VY Cycle 13 Haling Depletion, EOFPL Bundle Average Relative Powers 21 5.1.2 VY Cycle 13 Haling Depletion, EOFPL Core Average Axial Power Distribution 22 5.1.3 VY Cycle 13 Rodded Depletion - ARO at EOFPL, Bundle Average Relative Powers 23 5.1.4 VY Cycle 13 Rodded Depletion - ARO at E0FPL, Core Average Axial Power Distribution 24 5.2.1 VY Cycle 13 Haling Depletion, EOFPL Bundle Average Exposures 25 5.2.2 VY Cycle 13 Rodded Depletion, EOFPL Bundle Average Exposures 26 5.3.1 VY Cycle 13 Cold Shutdown Delta K in Percent versus Cycle Exposure 27 7.1.1 Flow Chart for the Calculation of 6CPR Using the RETRAN/TCPYA01 Codes 49 7.1.2 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL13, " Measured" Scram Time 50 7.1.3 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at E0FPL13-1000 MWD /ST, " Measured" Scram Time 51 7.1.4 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL13-2000 MWD /ST, " Measured" Scram Time 52 7.1.5 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL13 "67B" Scram Time 53 7.1.6 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at E0FPl.13-1000 MWD /ST, "67B" Scram Time 54 7.1.7 Inserted Rod Worth and Rod Position versus Time From Initial Rod Move.aent at E0FPL13-2000 MWD /ST, "67B" Scram Time 55
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LIST OF FIGURES (Continued)
Number Title Page 7.2.1 Turbine Trip Without Bypass, EOFPL13 Transient Response versus Time, " Measured" Scram Time 56 7.2.2 Turbine Trip Without Bypass, E0FPL13-1000 MWD /ST Transient Response versus Time, " Measured" Scram Time 59 7.2.3 Turbine Trip Without Bypass, EOFPL13-2000 MWD /S7 Transient Response versus Time, " Measured" Scram Time 62 7.2.4 Generator Load Rejection Without Bypass, E0FPL13 Transient Response versus Time. " Measured" Scram Time 65 7.2.5 Generator Load Rejection Without Bypass, E0FPL13-1000 MWD /ST Transient Response versus Time, " Measured" Scram Time 68 7.2.6 Generator Load Rejection Without Bypass, E0FPL13-2000 MWD /ST Transient Response versus Time, " Measured" Scram Time 71 7.2.7 Loss of 1000F Feedwater Heating, E0FPL13-1000 MWD /ST (Limiting Case) Transient Response versus Time 74 7.3.1 MSIV Closure, Flux Scram, EOFPL13 Transient Response versus ' lime, " Measured" Scram Time 76 7.4.1 Reactor Initial Conditions and Transient Summary for the VY Cycle 13 Rod Withdrawal Error Case 1 79 7.4.2 Reactor Initial Conditions and Transient Summary for the VY Cycle 13 Rod Withdrawal Error Case 2 80 7.4.3 VY Cycle 13 RWE Case 1 - Setpoint Intercepts Determined by the A+C Channel 81 7.4.4 VY Cycle 13 RWE Case 1 - Setpoint Intercepts Determined by the B+D Channel 82 7.6.1 First Four Rod Arrays Pulled in the A Sequences 83 7.6.2 First Four Rod Arrays Pulled in the B Sequences 83
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LIST OF TABLES Number Title Page 2.1.1 VY Cycle 12 Operating Highlights 3 2.2.1 VY Cycle 11 Operating Highlights 4 2.2.2 VY Cycle 10 Operating Highlights 5 3.1.1 VY Cycle 13 Fuel Bundle Types and Numbers 7 3.3.1 Design Basis VY Cycle 12 and Cycle 13 Exposures 7 4.1.1 Nominal Fuel Mechanical Design Parameters 12 4.2.1 Gap Conductance Values Used in VY Cycle 13 Transient Analyses 13 4.2.2 Peak Linear Heat Generation Rates Corresponding to Incipient Fuel Centerline Melting and 1% Cladding Plastic Strain 13 5.3.1 VY Cycle 13 K-Effective Values and Shutdown Margin Calculation 20 5.4.1 VY Cycle 13 Standby Liquid Control System Shutdown capability 20 7.1.1 VY Cycle 13 Summary of System Transient Model Initial Conditions for Core Wide Transient Analyses 44 7.1.2 VY Cycle 13 Transient Analysis Reactivity Coefficients at Selected Conditions 45 7.2.1 VY Cycle 13 Core Wide Transient Analysis Results 46 7.3.1 VY Cycle 13 Overpressurization Analysis Results 47 7.5.1 VY Cycle 13 Rotated Bundle Analysis Results 47 7.6.1 Control Rod Drop Analysis - Rod Array Full Order 48 7.6.2 VY Cycle 13 Control Rod Drop Analysis Results 48 A.1 Vermont Yankee Nuclear Power Station Cycle 13 MCPR Operating Limits A-2 A.2 Vermont Yankee Nuclear Power Station MAPLHGR Operating Limits for BP8DRB299 A-3
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ACKNOWLEDGEMENTS The authors and principal contributors would like to acknowledge the contributions to this work by the YAEC Word Processing Center. Their l assistance in preparing text for this document is recognized and greatly I
appreciated.
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1.0 INTRODUCTION
This report provides information to support the operation of the Vermont Yankee Nuclear Power Station through the forthcoming Cycle 13. In this report, Cycle 13 will frequently be referred to as the Reload Cycle. The preceding Cycle 12 will frequently be referred to as the Current Cycle. The refueling between the two will involve the discharge of 136 irradiated fuel bundles and the insertion of 136 new fuel bundles. The resultant core will consist of 136 new fuel bundles and 232 irradiated fuel bundles. Some of the irradiated fuel was present in the reactor in Cycles 10 and 11, as well as the Current Cycle. These cycles will frequently be referred to as Past Cycles.
The new fuel bundles differ from the irradiated bundles in the following ways: 1) the average bundle enrichment has been raised from 2.89 to 2.99, 2) the cladding now has a barrier; and 3) the pellet diameter has been increased. The use of this new bundle type will provide a transition to longer cycles.
This report contains descriptions and analyses results pertaining to the mechanicci, thermal-hydraulic, physics, and safety aspects of the Reload Cycle.
The Cycle 13 MCPR operating limits and the MAPLHGR operating limits for the new fuel bundles are given in Appendix A.
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l 2.0 RECENT REACTOR OPERATING HISTORY 2.1 Operating History of the Current Cycle The currently operating cycle is Cycle 12. To date, the Current Cycle has been operated smoothly at, or near, full power with the exception of normal maintenance, sequence exchanges, and a schedule coastdown. The operating history highlights and control rod sequence exchange schedule of the Current Cycle are found in Table 2.1.1.
2.2 Operating History of Past Applicable Cycles M
The irradiated fuel in the Reload Cycle includes some fuel bundles initially inserted in Cycles 10 and 11. These Past Cycles operated smoothly at, or near, full power with the exception of normal maintenance, sequence _
exchanges, and coastdown to the end of cycle. The highlights of the Past Cycles are found in Tables 2.2.1 and 2.2.2. The Past Cycles are described in detail in References 1 and 2.
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TABLE 2.1.1 VY CYCLE 12 OPERATING HIGHLIGHTS Beginning of Cycle Date June 30, 1986 End of Cycle Date August 8, 1987*
Weight of Uranium As-Loaded (Short Tons) 74.44 Beginning of Cycle Core Average Exposure (MWD /ST) 9820 End of Full Power Core Average Exposure (MWD /ST) 16370 End of Cycle Core Average Exposure (MWD /ST) 17720*
Number of Fresh Assemblies 120 Number of Irradiated Assemblies 248 Control Rod Sequence Exchange Schedule:
Sequence Date From To September 6, 1986 A2-1 B1-1 November 1, 1986 B1-1 Al-1 December 20, 1986 Al-1 B2-1 February 7, 1987 B2-1 A2-2 April 4, 1987 A2-2 B1-2
- Projected Dates and Exposures.
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TABLE 2.2.1 VY CYCLE 11 OPERATING HIGHLIGHTS Beginning of Cycle Date August 6, 1984 End of Cycle Date September 21, 1985 Weight of Uranium As-Loaded (Short Tons) 74.25 Beginning of Cycle Core Average Exposure (MWD /ST) 10,418 End of Full Power Core Average Exposure (MWD /ST) 16,733 End of Cycle Core Average Exposure (MWD /ET) 18,283 Capacity Factor While Operating (%) 89.2 Number of Fresh Assemblies 104 Number of Irradie.ed Assemblies 264 h Control Rod Sequence Exchange Schedule:
Seg-tence Date From To October 24, 1984 Al-1 B2-1 December 15, 1984 B2-1 A2-1 February 2, 1985 A2-1 B1-1 March 23, 1985 B1-1 Al-2 May 18, 1985 Al-2 B2-2 4
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TABLE 2.2.2 VY CYCLE 10 OPERATING HIGHLIGHTS Beginning of Cycle Date June 17, 1983 End of Cycle Date June 15, 1984 Weight of Uranium As-Loaded (Short Tons) 74.13 Beginning'of Cycle Core Average Exposure (MWD /ST) 10,463 End of Full Power Core Average Exposure (MWD /ST) 17,185 .
End of Cycle Core Average Exposure (MWD /ST) 17,806 Capacity Factor While Operating (%) 93.6 Number of Fresh Assemblies 108 Number of Irradiated Assemblies 260 Control Rod Sequence Exchange Schedule:
Sequence Date From To August 12, 1983 Al-1 B2-1 October 1, 1983 B2-1 A2-1 November 5, 1983 A2-1 B1-1 December 17, 1983 B1-1 Al-2 January 23, 1984 Al-2 B2-2 March 3, 1984 B2-2 A2-2 April 16, 1984 A2-2 B1-2 2810R
3.0 RELOAD CORE DESIGN DESCRIPTION 3.1 Core Fuel Loading The Reload Cycle core will consist of both new and irradiated assemblies. All the assemblies have bypass flow holes drilled in the lower l
l tie plate. Table 3.1.1 characterizes the core by fuel type, batch size, and first cycle loaded. A description of the fuel is found in Reference 3.
3.2 Design Reference Core Loading Pattern The Reload Cycle assembly locations are indicated on the map in Figure 3.2.1. For the sake of legibility only the upper right quadrant is shown.
The other quadrants are mirror images with bundles of the same type having nearly identical exposures. The bundles are identified by the reload number in which they were first introduced into the core. If any changes art made to I the loading pattern at the time of refueling, they will be evaluated under 10CFR50.59. The final loading pattern with specific bundle serial numbers will be supplied in the Startup Test Report.
3.3 Assembly Exposure Distribution The assumed nominal exposure on the fuel bundles in the Reload Cycle design reference loading pattern is given in Figure 3.2.1. To obtain this exposure distribution, Past Cycles were depleted with the SIMULATE model [4,5) using actual plant operating history. For the Current Cycle, plant operating history was used through January 6, 1987. Beyond this date, the exposure was ,
accumulated using a best-estimate rodded depletion analysis to End of Full Power Life (EOFPL) followed by a projected coastdown to End of Cycle (EOC).
Table 3.3.1 gives the assumed nominal exposure on the Current Cycle and the Beginning of Cycle (BOC) core average exposure that results from the ,
shuffle into the Reload Cycle loading pattern. The Reload Cycle EOFPL core -
average exposure and cycle capebility are provided.
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TABLE 3.1.1 VY CYCLE 13 FUEL BUNDLE TYPES AND NLHBERS Fuel Reload Cycle Possible Designation Designation Loaded Number Bundle ids Irradiated P8DFB289 R9 10 8 LY4XXX P8DPB289 R10 11 104 LY6XXX, LY7XXX P8DPB289 R11 12 120 LY7XXX, LYCXXX ,
New ,
BP8DRB299 R12 13 136 LYJXXX NOTE: XXX stands for the last three digits of the bundle serial number.
I TABLE '.3.1 DESIGN BASIS VY CYCLE 12 AND CYCLE 13 EXPOSURES Assumed Current Cycle Core Average Exposure End of Cycle 12 17.72 GWD/ST Assumed Reload Cycle Core Average Exposure Beginning of Cycle 13 8.43 GWD/ST Haling Calculated Core Average Exposure at End of Full Power Life, Cycle 13 16.80 GWD/ST Cycle 13 Exposure Capability S.37 GWD/ST 4
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l VERMONT YANKEE I CYCLE 13 l BOC BUNDLE AVERAGE EXPOSURES 23 25 27 29 31 33 35 37 39 41 43
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l I I I R10 R10 R10 BUNDLE ID 44 19.02 19.66 19.98 EXPOSURE (GWD/ST) R9 - P8DPB289, Reload 9 R10- P8DFB289, Reload 10 R12 R11 R11 R10 R11- P8DPB289, Reload 11
.00 10.22 10.30 18.74 R10 R12 R11 R12 R11 R10 R9 40 16.58 .00 9.84 .00 10.40 18.97 20.57 R12 R11 R12 R11 R12 R12 R11 Rio 38
.00 8.03 .00 9.50 .00 .00 9.99 17.96 '
R11 R12 R10 R12 R11 R12 R11 R11 R9 ,
36 7.04 .00 17.96 .00 10.36 .00 8.61 9.93 20.49 R10 R11 R12 R11 R12 R11 R12 R12 R10 34 18.41 9.78 .00 9.85 .00 B.52 .00 .00 19.19 R10 R12 R10 R12 R10 R12 R11 R12 R11 13.57 .00 15.84 .00 17.12 .00 10.42 .00 10.40 R12 R11 R12 R1) R12 R11 R12 R11 R12 R10
.00 7.14 ,00 7.11 40 9.79 40 9.41 .00 18.92 R11 R12 R10 R12 R10 R12 R10 R12 R11 R11 R10
- 28 .
8.58 .00 17.86 .00 15.97 .00 17.97 .00 9.81 10.23 19.88 '
R10 R11 R12 R11 R12 R11 R12 R11 R12 R11 R10 NP.
- 26 17.6+ 7.08 .00 7.15 .00 9.73 .00 8.04 .00 10.31 19.61 R10 R10 R11 R12 R10 R10 R11 R12 R10 R12 R10
- 24 17.03 17.57 8.57 .00 15.51 18.47 7.01 .00 16.66 .00 19.12 FIGURE 3.2.1 VY CYCLE 13 DESIGN REFERENCE LOADING PATTERN, UPPER RIGHT QUADRANT e
I 4.0 FUEL MECHANICAL AND THERMAL DESIGN 4.1 Mechanical Design All fuel to be inserted into the Reload Cycle was fabricated by the General Electric Company (GE). The major mechanical design parameters are given in Table 4.1.1 and Reference 3. Fuel rod design changes have been incorporated in the Reload Cycle fuel design. The new fuel pellet diameter is increased and is fabricated to a higher nominal fuel pellet density with less in-reactor densification. In addition, the new fuel cladding incorporates a barrier on the internal surface designed to reduce the effects of pellet-cladding interaction. Detailed descriptions of the fuel rod mechanical design and mechanical design analyses are provided in Reference 3. These design analyses remain valid with respect to the Reload Cycle operation.
Mechanical and chemical compatibility of the fuel assemblies with the in-service reactor e rironment is also addressed in Reference 3.
4.2 Thermal Design The fuel thermal effects calculations were performed using the FROSSTEY computer code [6,7). The FROSSTEY code calculates pellet-to-cladding gap conductance and fuel temperatures from a combination of theoretical and empirical models which include fuel and cladding thermal expansion, fission gas release, pellet swelling, pellet densification, pellet cracking, and fuel and cladding thermal conductivity.
The thermal effects analysis included the calculation of fuel temperatures and fuel cladding gap conductance under nominal core steady state and peak linear heat generation rate conditions. Figure 4.2.1 provides the core average response of gap conductance. These calculations integrate the responses of individual fuel batch average operating histories over the core average exposure range of the Reload Cycle. The gap conductance values are weighted axially by power distributions and radially by volume. The core-wide gap conductance values for the RETRAN system simulations, described in Sections 7.1 through 7.3, are from this data set at the corresponding exposure statepoints.
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The gap conductance values input to the hot channel calculations (Section 7.1) were evaluated for the given fuel bundle type as a function of the assembly exposure. The calculation assumed a 1.4 chopped cosine axial power shape with the peak power node running at the maximum average planar linear heat generation rate (MAPLHGR) limit defined in Reference 8 and Table A.2. Figure 4.2.2 provides the hot channel response of gap conductance. In Figure 4.2.2, " planar exposure" refers to the exposure of the node operating at the MAPLHCR limit. Gap conductance values for the hot i
channel analysis were extracted from Figure 4.2.2 using the limiting bundle exposure of any minimum critical power ratio (MCPR) limiting bundle within the exposure interval of interest. The SIMULATE rodded depletion (Section 5.1.2) provides predictions of both limiting MCPR and the associated bundle exposure for the entire cycle.
Table 4.2.1 provides the core average and hot channel gap conductance values used in the transient analyses (Section 7.1). The values for gap conductance are higher than those calculated in previous cycles for the following reasons: 1) the diametral gap of the fresh fuel has been reduced as .
a result cf an increase in pellet diameter, 2) the as-fabricated pellet density has been increased which results in a decrease in the in-reactor fuel -
pellet densification, thereby causing a smaller reduction in the irradiated pellet-cladding gap, and 3) the specification for surface roughness for the fuel pellet and the fuel cladding has been revised. These changes improve the heat transfer capability of the fuel which result in reduced fuel pellet temperatures, fission gas releases, and fuel rod internal pressures.
Fuel rod local linear heat generation rates (LHGR) at fuel centerline ,
incipient melt and 1% cladding plastic strain as a function of local axial segment exposure for the peak gadolinia concentrations used in Vermont Yankee ?
fuel bundles were calculated. These values are displayed in Table 4.2.2. , ;
Initial conditions assumed that fuel rods operated at the local segment power level of the maximum allowable LHGR prior to the power increase.
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4.3 Operating Experience All irradiated fuel bundles scheduled to be reinserted in the Reload Cycle have operated as expected in Past Cycles of Vermont Yankee. Off-gas measurements in the Current Cycle indicate that a small number of fuel rod failures may have occurred. Vermont Yankee is planning to identify the failed rods during the outage.
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TABLE 4.1.1 NOMINAL FUEL MECHANICAL DESIGN PARAMETERS Fuel Types i Irradiated New l
Fuel Bundle
- Bundle Type P8X8R BP8X8R Vendor Designation (Table 3.1.1) P8DPB289 BP8DRB299 Initial Enrichment, w/o U-235 2.89 2.99 Rod Array 8X8 8XS Fuel Rods per Bundle 62 62 Fuel Channel Material Zr-4 Zr-4 Wall Thickness, Inches 0.080 0.080
- Complete bundle, pellet, and rod descriptions are found in Reference 3.
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TABLE 4.2.1 GAP CONDUCTANCE VALUES USED IN VY CYCLE 13 TRANSIENT ANALYSES Cycle Exposure Core Average Hot Channel Hot Channel (1)
Statepoint Gap Conductance Bundle Exposure Gap Conductance (MWD /ST) (BTU /Hr-Ft - F) (MWD /ST) (BTU /Hr-Ft - U F)
BOC13 1910 9447(2) 4850 E0FPL13-2000 MWD /ST 3130 8096 4450 E0FPL13-1000 MWD /ST 3290 9363 4820 EOFPL13 3475 9447 4850 NOTE (1) Hot channel gap conductance values are all derived for the BP8DRB299 fuel type.
(2) Between BOC13 and E0FPL13-2000 MWD /ST, the highest exposure limiting hot channel bundle is once-burned.
TABLE 4.2.2 PEAK LINEAR HEAT GENERATION RATES CORRESPONDING TO INCIPIENT FUEL CENTERLINE MELTING AND 1% CLADDING PLASTIC STRAIN (1) 0.0 w/o Gd230 3.0 w/o Gd23 0
Exposure Melt 1% Cp Melt 1% Cp (MJD/MT) (kW/ft) (kW/ft) (kW/ft) (kW/ft)
Fuel Type P8x8R P8DFB289 0 24.0 24.0 22.5 24.0 25,000 24.0 24.0 21.5 22.5 50,000 24.0 17.5 20.0 14.5 0.0 w/o Gd230 4.0 w/o Gd23 0
Exposure Melt 1% Cp Melt 1% Cp (MVD/MT) (kW/ft) (kW/ft) (kW/ft) (kW/ft)
Fuel Type BP8x8R BP8DRB299 0 24.0 24.0 21.5 24.0 25,000 24.0 24.0 20.5 20.5 50.000 23.5 15.5 19.0 12.0 NOTE (1) Peak linear heat generation rates shown are minimum bounding values to the occurrence of the given conditien.
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5.0 NUCLEAR DESIGN 5.1 Core Power Distributions The Reload Cycle was depleted using SIMULATE [4] to give both a rodded depletion and an All Rods Out ( ARO) Haling depletion.
5.1.1 Haling Power Distribution The Haling depletion serves as the basis for defining core reactivity characteristics for most transient evaluations. This is primarily because its flat power shape has conservatively weak scram characteristics.
The Haling power distribution is calculated in the ARO condition. The Haling iteration converges on a self-consistent power and exposure distribution for the burnup step to EOFPL. In principle, this should provide the overall minimum peaking power shape for the cycle. During the actual cycle, flatter power distributions might occasionally be achieved by shaping with control rods. However, such shaping would leave underburned regions in the core which would peak at another point in time. Figures 5.1.1 and 5.1.2 give the Haling radial and axial average power distributions for the Reload Cycle.
5.1.2 Rodded Depletion Power Distribution The rodded depletion was used to evaluate the mislocated bundle error and the rod withdrawal error because it provides initializing rod patterns and it makes more realistic predictions of initial CPR values. It was also used in the rod drop worth and shutdown margin calculations because it burns the top of the core more realistically than the Haling.
To generate the rodded depletion, control rod patterns were developed which give critical eigenvalues at each point in the cycle and peaking similar to the Haling calculation. The resulting patterns were frequently more peaked than the Haling, but were below expected operating limits. However, as stated above, the underburned regions of the core can exhibit peaking in excess of the Haling peaking when pulling ARO at EOFPL. Figures 5.1.3 and 5.1.4 give the APO at EOFPL power distributions for the Reload Cycle rodded depletion.
2810R
Note, in Figure 5.1.4, that the average axial power at ARO for the rodded depletion is more bottom peaked than the Haling (Figure 5.1.2). The rodded depletion would result in better scram characteristics at E0FPL.
5.2 Core Exposure Distributions The Reload Cycle exposures are summarized in Table 3.3.1. The projected BOC radial exposure distribution for the Reload Cycle is given in Figure 3.2.1. The Haling calculation produced the EOFPL radial exposure distribution given in Figure 5.2.1. Since the Haling power shape is constant, it can be held fixed by SIMULATE to give the exposure distributions at various mid-cycle points. BOC, EOFPL-2000 MWD /ST, EOFPL-1000 MWD /ST, and EOFPL exposure distributions were used to develop reactivity input for the core wide transient analyses.
The rodded depletion differs from the Haling during the cycle because
\
the rods shape the power differently. However, rod sequences are swapped frequently and the overall exposure distribution at end of cycle is similar to the Haling. Figure 5.2.2 gives the E0FPL radial exposure distribution for the Reload Cycle rodded depletion.
5.3 Cold Core Reactivity and Shutdown Margin The cold K ff with ARO and the cold K ff with All Rods Inserted (ARI) at BOC were calculated using the SIMULATE code [4,5] and are shown in Table 5.3.1. df is the amount of K with ARO minus the cold critical K df excess core reactivity. K df with ARI minus the K,gg with ARO is the, worth of all the control rods.
The cold critical eigenvalue K,gg was defined as the average calculated critical eigenvalue minus a 957, confidence level uncertainty. Then all cold results were normalized to make the critical K d f
- i E *""# l"* ' 9 "81 to 1.000.
2810R _ _ _ _ _
Technical Specifications [8] state that, for sufficient shutdown margin, the core must be suberitical by at least 0.25% AK + R (defined below) with the strongest worth control rod withdrawn. Again, using SIMULATE, a search was made for the strongest worth control rod at various exposures in the cycle. This is necessary because rod worths change with exposure on adjacent assemblies. Then the cold K,gg with the strongest rod out was calculated at B0C and at the end of each control rod sequence. Subtracting each cold K,ff with the strongest rod out from the cold critical K,7g eigenvalue defines the shutdown margin as a function of exposure. Figure 5.3.1 shows the results. Because the local reactivity may increase with exposure, the shutdown margin (SDM) may decrease. To account for this, and other uncertainties, the value R is calculated. R is defined as Ry plus R.
2 R is the difference between the cold K,ff with the strongest rod out at BOC and the maximum cold K eff with the strongest rod out in the cycle. R is 2 a measurement uncertainty in the demonstration of SDM associated with the manufacture of past control blades. It is presently set at .07% AK. The shutdown margin results are summarized in Table 5.3.1.
5.4 Standby Liquid Control System Shutdown Capability The shutdown capability of the Standby Liquid Control Systen. (6LCS) is designed to bring the reactor from full power to cold, AR0, xenon free shutdown with at least 5% AK margin. Using SIMULATE [4], the ppm of boron was adjusted until the Kgg reached the cold critical K gg minus .05.
Each case assumed cold, xenon-free conditions, with All Rods Out. The cycle was searched to find the most reactive point in the cycle. This analysis found that the plant would be suberitical by 5% AK at the worst point in cycle with less than the 800 ppm of boron required by VY Technical Specifications [8]. Table 5.4.1 lists the amount of boron concentration and the corresponding shutdown capability of the SLCS.
5.5 Changes in Nuclear Design Methods The comycter code, CASMO-1 [5), had input limitations that did not allow adequate modelling of the new fuel type. Therefore, CASMO-2 [10] was 2810R
l l
, )
i used to generate the cross sections for all fuel types used in Cycle 13. '
j CASMO-1 and CASMO-2 use the same' cross section library and are neutronically cquivalent with one exception. The reflecting boundary condition for the ]
two-dimensional bundle calculation has been revised in CASMO-2. CASMO-2 more !
i cccurately handles cases with high flux gradients close to the problem boundary.
Before utilizing the methodology of CASMO-2 with SIMULATE, Vermont !
Yenkee Cycles 9 through 11, and most of the current Cycle 12 were benchmarked. The CASMO-2/ SIMULATE model compared to the CASMO-1/ SIMULATE rodel as follows: The average hot eigenvalue went up by +.0009. The average of the standard deviations went from !.00240 to .00246. When comparing both models to actual plant Traversing Incore Probe (TIP) readings, the average cbsolute error for the four cycles increased from 1.53% to 1.58%. The cold
} critical data for the four cycles was also benchmarked. The 95% confidence, minimum cold critical eigenvalue went from .9926 to .9935. As can be seen by these results, the differences between the CASM0-2 and CASMO-1 were examined in detail and found to have minimal impact on the analyses.
2810R
f TABLE 5.3.1
~
'VY CYCLE 13-K gg, VALUES AND SHUTDOWN MARGIN CALCULATION o
BOC eK gg - Uncontrolled 1.1220 BOC eK gg - Controlled .9705 Cold Critical K gg e Eigenvalue 1.0000 BOC.Keff - Controlled With~ .9887 Strongest Worth Rod Withdrawn..
Cycle Minimum Shutdown Margin Occurs at 1.13% AK BOC With Strongest Worth Rod Withdrawn R1 , Maximum Increase in Cold K egg .00% AK With Exposure TABLE 5.4.1 VY CYCLE 13 STANDBY LIQUID CONTROL SYSTEM SHUTDOWN CAPABILITY ppm of Boron Shutdown Margin 718 .5.0% AK 800 6.6% AK i
ll 2810R 1
l _- _-_ ____________---____ ___ - _ - _
VERMONT YANKEE CYCLE 13 HALING DEPLETION EOFPL BUNDLE AVERAGE RELATIVE POWERS ,
23 25 27 29 31 33 35 37 39 41 43 i
R10 RIO R10 BUNDLE ID 44
.464 426 .371 EOFPL RELATIVE POWER R9 - P8DPB289, Reload 9 R10- P8DPB289, Reload 10 R12 R11 R11 R10 R11- P8DPB289, Reload 11 R12-BP8DRB299, Reload 12 42
.870 .754 .677 .536 R10 R12 R11 R12 R11 Rio R9 40
.578 1.073 .907 .926 .736 .546 .412 R12 R11 R12 R11 R12 R12 R11 R10 38 1.21B 1.083 1.185 1.008 1.066 .965 .707 .491 R11 R12 R10 R12 R11 R12 R11 R11 R9 36 1.138 1.276 1.017 1.252 1.064 1.143 .905 .708 .412 R10 R11 R12 R11 R12 R11 R12 R12 R10 34 1.016 1.148 1.318 1.160 1.290 1.103 1.143 .965 .541 R10 R12 RIO R12 R10 R12 R11 R12 R11 1.072 1.340 1.108 1.354 1.075 1.290 1.064 1.066 .736 R12 R11 R12 R11 R12 R11 R12 R11 R12 R10 30 1.350 1.231 1.378 1.237 1.354 1.160 1.253 1.010 .927 .534 R11 R12 R10 R12 R10 R12 R10 R12 R11 R11 R10
- 28 1.174 1.349 1.102 1.378 1.106 1.318 1.01B 1.185 .908 .678 .371 R10 R11 R12 R11 R12 R11 R12 R11 R12 R11 R10 4
- 26 1.002 1.165 1.349 1.231 1.341 1.148 1.277 1.083 1.073 .753 .427 R10 R10 R11 R12 R10 R10 R11 R12 R10 R12 R10 l - 24 l
.947 1.003 1.175 1.350 1.073 1.016 1.139 1.218 .877 .870 .462 l I
FIGURE 5.1.1 ,
VY CYCLE 13 HALING DEPLETION, EOFPL BUNDLE AVERAGE RELATIVE POWERS a
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VERMONT. YANKEE CYCLE 13 RODDED DEPLETION EOFPL BUNDLE AVERAGE RELATIVE POWERS 23 25 27 29 31 33 35 37 39 41 43 I I R10 Rio R10 SUNDLE ID
- 44
.452 415 .359 EOFPL RELATIVE POWE't R9 - P8DPB289, Reload 9 RID- P8DPB289, Reload 10 R12 R11 R11 R10 R11- PSDPB289, Reload 11 H12-BP8DRB299, Reload 12 42 R10 R12 R11 R12 R11 R10 R9 40
.865 1.056 .889 .908 .719 .532 .400 R12 R11 R12 R11 R12 R12 R11 R10 38 1.210 1.070 1.170 .990 1.046 .944 .689 .475 R11 R12 Rio R12 R11 R12 R11 R11 R9 36 1.140 1.278 1.013 1.241 1.048 1.121 .884 .600 .399 R10 R11 R12 R11 R12 R11 R12 R12 R10 34 1.027 1.157 1.328 1.156 1.280 1.084 1.121 :944 .527 R10 R12 R10 R12 R10 R12 R11 R12 R11 32 1.099 1.370 1.123 1.367 1.074 1.280 1.047 1.045 .718 R12 R11 R12 R11 R12 R11 R12 R11 R12 R10 30 1.403 1.266 1.414 1.257 1.367 1.155 1.239 .990 .906 .618 R11 R12 R10 R12 RIO R12 R10 R12 R11 R11 R10
- 28 1.224 1.407 1.138 1.418 1.123 1.324 1.009 1.167. .888 .659 .358 R10 R11 R12 R11 R12 R11 R12 R11 R12 R11 R10
- 26 1.044 1.215 1.407 1.269 1.371 1.152 1.271 1.065 1.082 .733 .413 R10 R10 R11 R12 R10 RIO R11 R12 RIO R12 R10
- 24
.982 1.042 1.221 1.402 1.007 1.020 1.131 1.202 .859 .848 .448 FIGURE 5.1.3 VY CYCLE 13 RODDED DEPLETION-ARO AT EOFPL, BUNDLE AVERAGE RELATIVE POWERS r
5 2
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VERMONT YANKEE :
CYCLE 13 HALING DEPLETION EOFPL BUNDLE AVERAGE EXPOSURES 23 25 27 29 31 33 35 37 39 41 43 l l
l- I I R10 R10 R10 BUNDLE ID 22.92 23.26 23.10 EXPOSURE (GWD/ST) R9 - P8DFB289, Reload 9 R12 R11 R11 R10 R10- P80PB289, Reload 10 R11- P8DPB289, Reload 11 7.20 16.57 16.01 23.26 l
R10 R12 R11 R12 R11 R10 R9 j 40 l 23.98 8.88 17.49 7.66 16.60 23.57 24.04 I R12 R11 R12 R11 R12 R12 R11 R10 ,
36 !
10.07 17.16 9.80 17.99 8.82 7.98 15.95 22.09 j i
R11 R12 R10 R12 R11 R12 R11 R11 R9 l 36 16.63 10.56 26.54 10.36 19.33 9.45 16.24 15.90 23.96 R10 R11 R12 R11 R12 R11 R12 R12 R10 34 26.97 19.46 10.90 19.62 10.67 17.82 9.45 7.98 23.75 R10 R12 R10 R12 R10 R12 R11 R12 R11 24.60 11.09 25.18 11.20 26.18 10.67 19.38 8.82 16.61 R12 R11 R12 R11 R12 R11 R12 R11 R12 R10 30
, 11.17 17.52 11.40 17.54 11.20 19.58 10.36 17.92 7.67 23.43 1
R11 R12 R10 R12 R10 R12 R10 R12 R11 R11 RIO I - 28 i
18.48 11.16 27.15 11.40 25.30 10.90 26.55 0.61 17.46 15.95 23.01 l R10 H11 R12 R11 R12 R11 R12 R11 R12 R11 R10 l
- 26 26.09 16.90 11.16 17.52 11.09 19.41 10.56 17.17 8.88 16.65 23.20 RIO R10 R11 R12 R10 RIO R11 R12 R10 R12 R10
- 24 33.02 26.02 18.47 11.17 24.55 27.03 16.60 10.07 24.04 7.19 23.01 FIGURE 5.2.1 VY CYLCE 13 HALING DEPLETION, EOFPL BUNDLE AVERAGE EXPOSURES
1 l
l l
VERMONT YANKEE I I
CYCLE 13 RODDED DEPLETION EOFPL BUNDLE AVERAGE EXPOSURES q l
l 23 25 27 29 31 33 35 37 39 41 43 I I l RIO R10 R10 BUNOLE ID 44 23.02 23.45 23.35 EXPOSURE (GWD/ST) R9 - P8DPB289, Relsac. 9 l R10- P8DPB289, Reload 10 i R12 R11 R11 R10 R11- P8DPB289, Reload 11 l R12-BP8DRB299, Reload 12 42 6.88 16.79 16.35 23.47 l
R10 R12 R11 R12 R11 R10 R9 40 24.08 8.65 17.74 7.45 16.72 23.66 24.21 R12 R11 R12 R11 R12 R12 R11 R10 9.68 17.35 9.56 18.19 8.50 7.61 16.17 22.35 i R11 R12 R10 R12 R11 R12 R11 R11 R9 1 1
36 !
16.65 10.09 26.54 10.06 19.51 9.19 16.51 16.12 24.13 i l
R10 R11 R12 R11 R12 R11 R12 R12 R10 j l
34 1 26.97 19.46 10.37 19.77 10.40 18.06 9.19 7.61 23.84 l R10 R12 R10 R12 R10 R12 R11 R12 R11 j 32 I 84.30 10.38 25.05 10.67 26.20 10.39 19.55 8.50 16.73 ;
R12 R11 R12 R11 R12 R11 R12 R11 R12 R10 9.96 17.15 10.59 17.39 10.65 19.73 10.08 18.12 7.47 23.65 R11 R12 R10 R12 Rio R12 R10 R12 R11 R11 R10
- 2B 17.86 9.94 26.57 10.46 25.09 10.45 26.64 9.5E 17.72 16.30 23.28 R10 R11 R12 R11 R12 R11 R12 R11 R12 R11 R10
- 26 25.75 16.41 9.94 17.07 10.29 19.51 10.26 17.41 8.67 16.90 23.42 R10 R10 R11 R12 R10 R10 R11 R12 R10 R12 R10
- 24 24.99 25.79 17.98 9.98 24.29 27.19 16.84 9.8C 24.20 6.92 23.13 FIGURE 5.2.2 VY CYCLE 13 RODDED DEPLETION, EOFPL BUNDLE AVERAGE EXPOSURES 0
0 _
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6.0 THERMAL-HYDRAULIC DESIGN-The thermal-hydraulic evaluation of the Reload Cycle was performed using the methods described in the following section. ;
I 6.1 Steady-State Thermal Hydraulics i 1
Core steady-state thermal-hydraulic analyses were performed using the FIBWR [11,12] computer code. The FIEWR code incorporates a detailed geometrical representation of the complex flow paths in a BWR core, and explicitly models the leakage flow to the bypass region. FIEWR calculates the core pressure drop and total bypass flow for a given total core flow. The power distribution, inlet enthalpy, and geometry are presumed known and are supplied to FIBWR. The power distribution is derived by SIMULATE [4]. Core pressure drop and total leakage flow predicted by the FIBWR code were used in setting the initial conditions for the system transient analysis model.
6.2 Reactor Limits Determination The objective for normal operation and anticipated transient events is to maintain nucleate boiling. Avoiding a transition to film boiling protects the fuel cladding integrity. Based on Reference 3, the fuel cladding integrity safety limit for Vermont Yankee is a Lowest Allowable Minimum Critical Power Ratio (LAMCPR) of 1.07. Operating limits are specified to l maintain adequate margin to onset of the boiling transition. The figure of merit utilized for plant operation is the Critical Power Ratio (CPR). This is l defined as the ratio of the critical power (bundle power at which some point within the assembly experiences onset of boiling transition) to the operating l
l bundle power. Thermal margin is stated in terms of the minimum value of the l
l critical power ratio, MCPR, which corresponds to the most limiting fuel assembly in the core. Both the transient (safety) and normal operating thermal limits in terms of MCPR are derived based on the GEXL correlation as described in Reference 13.
2810R 3
q Vermont Yankee Technical Specifications [8] limit the operation of the Reload Cycle. fuel to a linear heat generation rate (LHGR) limit of 13.4 kW/ft. _The basis for a MLHGR of 13.4 kW/ft can be found in Reference 3.
l f
l l
l I
3810R
?
-_-___L
1 1
7.0 ACCIDENT ANALYSIS 7.1 Transient Analysis Transient simulations are performed to assess the impact of certain ,
transients on the heat transfer characteristics of the fuel. The figure of I 1
merit used is the Critical Power Ratio (CPR). It is the purpose of the analysis to determine the minimum critical power ratio (MCPR) such that the l safety limit is not violated for the transients considered. ;
l 7.1.1 Methodology 1
1 The analysis requires two types of simulations. A system level simulation is performed to determine tne overall plant response. Transient core inlet and exit conditions and normalized power from the system level calculation are used to perform detailed thermal-hydraulic simulations of the l fuel, referred to as " hot channel calculations". The hot channel simulations provide the bundle transient ACPR (the initial bundle CPR minus the MCPR experienced during the transient).
l The system level simulations are performed with the model documented in Reference 14 The hot channel calculations are performed with the RETRAN [15] and l TCPYA01 [16] computer codes. The GEXL correlation [13} is used in TCPYA0i to evaluate critical power ratio. The calculational procedure is outlined bel,w.
The hot channel transient ACPR calculations employ a series of
" inner" and " outer" iterations, as illustrated by the flow chart in Figure 7.1.1. The outer loop iterates on the hot channel initial power ;
level. This is necessary because the ACPR for a given transient varies with Initial Critical Power Ratio (ICPR). However, only the ACPR corresponding j l
to a transient MCPR equal to the safety limit (i.e., 1.07 + ACPR = ICPR) is I appropriate. The approximate constancy of the ACPR/ICPR ratio is useful in these iterations. Each outer iteration requires a RETRAN hot channel run to I
2810R l
1
- - _ _ . __ ______________________________O
calculate the transient enthalpies, flows, pressure and saturation properties at each time step. These are required for input to the TCPYA01 code. TCYPA01 is then used to calculate a CPR at each time step during the transient, from which a transient ACPR is derived. The hot channel model assumes a chopped cosine axial power shape with a peak / average ratio of 1.4.
The inner loop iterates or. the hot channel inlet flow. These iterations are necessary because the RETRAN hot channel model calculates the entrance loss coefficient when the initial power level, flow, and pressure drop are given. The pressure drop is assumed equal to the core average l pressure drop, and the flow is varied for a given power level until the calculated entrance loss coefficient is correct. FIBWR (11, 12] is utilized to estimate the correct inlet flow for a particular power level and pressure drop.
I t 7.1.2 Initial Conditions and Assumptions The initial conditions for the system simulations are based on maximum turbine capacity of 105% of rated steam flow. The corresponding reactor f conditions are 104.5% core thermal power and 100% core flow. The core axial power distribution for each of the expcsure points is based on the 3-D ;
SIMULATE predictions associated with the generation of the reactivity data (Section 7.1.3). The core inlet enthalpy is set so that the amount of carryunder from the steam separators and the quality in the liquid region outside the separators is as close to zero as possible. For fast pressurization transients, this maximizes the initial pressurization rate and predicts a more severe neutron power spike. A summary of the initial operating state used for the system simulations is provided in Table 7.1.1.
I Assumptions specific to a particular transient are discussed in the section describing the transient. In general, the following assumptions are made for all transients: l l
- 1. Scram setpoints are at Technical Specification [8] limits.
I 2810R l i
- 2. Protective system logic delays are at equipment specification limits.
- 3. Safety / relief valve and safety valve capacities are based on Technical Specification [8] rated values.
- 4. Safety / relief valve and safety valve setpoints are modeled as being at the Technical Specification [8] upper limit. Valve responses l are based on slowest specified response values.
- 5. Control rod drive scram speed is based on the Technical
. Specification limits. The analysis addresses a dual set of scram speeds, referred to as the " Measured" and the "67B" scram times.
" Measured" refers to the faster scram times given in Section !
3.3.C.1.1 of the Technical Specifications [8]. "67B" refers to the slower scram times given in Section 3.3.C.1.2 of the Technical Specifications [8].
I 7.1.3 Reactivity Functions 1
The methods used to generate the fuel temperature, moderator density, and scram reactivity functions are described in det' ail in Reference 17. The method is outlined below.
A complete set of reactivity functions, the axial power distribution, and the kinetics parameters are generated from base states established for EOFPL, EOFPL-1000 MWD /ST, E0FPL-2000 MWD /ST, and BOC exposure statepoints.
These statepoints are characterized by exposure and void history distributions, control rod patterns, and core thermal-hydraulic conditions.
The latter are consistent with the assumed system transient conditions provided in Table 7.1.1.
l l The BOC base state is established by shuffling from the previously i l
defined Current Cycle endpoint into the Reload Cycle loading pattern. A criticality search provides an estimate of the BOC critical rod pattern. The 1
2810R '
l i
EOFPL and intermediate core exposure and void history distributions are
. calculated with a Haling depletion as described in Section 5.2. The E0FPL state is unrodded. As such, it is defined sufficiently. However, E0FPL-1000 MWD /ST and E0FPL-2000 MWD /ST exposure statepoints require base control rod patterns. These are developed to be as " black and white" as possible. That.
is, beginning with the rodded depletion configuration, all control rods which are more than half inserted are fully inserted, and all control rods which are less than half inserted are fully withdrawn. If the SIMULATE calculated parameters are within operating limits,'then this configuration becomes,the base case. If the limits are exceeded, a minimum number of control rods are adjusted a minimum number of notches until the parameters fall within limits.
Using this method, the control rod patterns and resultant power distributions are established which minimize the scram reactivity function and maximize the core average moderator density reactivity coefficient. For the transients analyzed, this tends to maximize the power response.
At each exposure statepoint, reactivity function table sets are' {
produced for the 12 core-volumes of the Vermont Yankee RETRAN model. The fuel I temperature (Doppler) data set is generated by fixing the power distribution while varying the fuel temperature associated with that power. A moderator density table set is generated specifically for each transient type. The moderator density reactivity functions for-the subcooling transient are l generated by quasi-statically varying the inlet subcooling only. The moderator enthalpy source distribution is allowed to iterate with the i calculated nuclear power until equilibrium is reached. The moderator density reactivity functions for the pressurization transients are generated by quasi-statically varying the core pressure. A series of calculations are l performed for various inlet moderator temperatures. The moderator enthalpy source distribution is fixed at the base state case.
In order to qualitatively compare the core reactivity characteristics between different base configurations, core average reactivity coefficients at selected conditions are provided in Table 7.1.2. Calculated point kinetics parameters for RETRAN are also provided.
2810R _ - - --
The reactivities versus scram insertion are calculated at constant, pretransient moderator conditions. These are fitted to yield highly detailed scram reactivity curves. The curves are combined with the appropriate rod position versus time data to generate the final RETRAN scram reactivity functions. Figures 7.1.2 through 7.1.4 display the inserted rod worths and rod positions as functions of scram time fer the " Measured" scram time analysis. Figures 7.1.5 through 7.1.7 display similar curves for the "67B" scram time analysis.
7.1.4 Transients Analyzed Past licensing analysis has shown that the transients which result in the minimum core thermal margins are:
i 1
- 1. Generator load rejection with complete failure of the turbine j bypass system. I I
- 2. Turbine trip with complete failure of the turbine bypass system. ]
l 1
- 3. Loss of feedwater heating.
The "feedwater controller failure" (maximum demaud) transient is not a limiting transient for Vermont Yankee, because of the plant's 110% steam flow f bypass system. Past analyses have shown this transient to be considerably less limiting than any of the above for all exposure points. Brief descriptions and the results of the transients analyzed are provided in the following section.
1 1
l 7.2 Transient Analysis Results l The transients selected for consideration were analyzed at exposure points of EOFPL, EOFPL-1000 MWD /ST, and E0FPL-2000 MWD /ST; the loss of i 1
feedwater heating transient was also evaluated at BOC conditions. The l transient results reported in Table 7.2.1 correspond to the limiting bundle type in the core. ,
1 l
2810R
l
.i j
k 7.2.1 Turbine Trip Without Bypass Transient (TTWOBP) l 1
The transient is initiated by a rapid closure (0.1 second closing time) of the turbine stop valves. It is assumed that the steam bypass valves, which S normally open to relieve pressure, remain closed. A reactor protection system
.l L signal is generated by the turbine stop valve closure switches. Control rod j drive motion is conservatively assumed to occur 0.27 seconds after the start of turbine stop valve motion. The ATWS recirculation pump trip is assumed to occur at a setpoint of 1150 psig dome pressure. A pump trip time delay of 1.0 second is assumed to account for logic delay and M-G set generator field collapse. In simulating the transient, the bypass piping volume up to the valve chest is lumped !nto the control volume upstream of the turbine stop valves. Predictions of the salient system parameters at the three' exposure points are shown in Figures 7.2.1 through 7.2.3 for the " Measured" scram time analysis.
7.2.2 Generator Load Re_iection Without Bypass Transient (GLRWOBP)
The transient is initiated by a rapid closure (0.3 seconds closing time) of the turbine control valves. As in the case of the turbine trip transient, the bypass valves are assumed to fail. A reactor protection system signal is generated by the hydraulic fluid pressure switches in the ;
acceleration relay of the turbine control system. Control rod drive motion is 1
conservatively assumed to occur 0.28 seconds after the start of turbine )
control valve motion. The same modeling regarding the ATWS pump trip and bypass piping is used as in the turbine trip simulation. The influence of the accelerating main turbine generator on the recirculation system is simulated by specifying the main turbine generator electrical frequency as a function of time for the M-G set drive motors. The main turbine generator frequency curve is based on a 100% power plant startup test and is considered representative for the simulation. The system model predictions for the three exposure ]
points are shown in Figures 7.2.4 through 7.2.6 for the " Measured" scram time analysis, j l
2810R i
7.2.3 Loss of Feedwater Heating Transient (LOFWH) l A feedwater heater can be lost in such a way that the steam extraction 1
line to the heater is shut off or the feedwater flow bypasses one of the 1 4
heaters. In either case, the reactor will receive cooler feedwater, which will produce an increase in the core inlet subcooling, resulting in a reactor i
power increase. )
The response of the system due to the loss of 100 F of the feedwater heating capability was analyzed. This represents the current licensing l assumption for the maximum expected single heater or group of heaters that can be tripped or bypassed by a single event.
Vermont Yankee has a scram setpoint of 120% of rated power as part of the Reactor Protection System (RPS) on high neutron flux. In this analysis, t no credit was taken for scram on high neutron flux, thereby allowing the L reactor power to reach its peak without scram. This approach was selected to provide a bounding and conservative analysis for events initiated from any :
power level.
The transient response of the system was evaluated at several exposares during the cycle. The transient evaluation at E0FPL-1000 MWD /ST was found to be the limiting case between BOC to E0FPL. The results of the system response to a loss of 100 F feedwater heating capability evaluated at EOFPL-1000 MWD /ST as predicted by the RETRAN code are presented in Figure 7.2.7.
7.3 Overpressurization Analysis Results
)
l i
Compliance with ASME vessel code limits is demonstrated by an analysis '
of the Main Steam Isolation Valves (MSIV) closing with failure of the MSIV position switch scram. EOFPL conditions were analyzed. The system model used 1 1
is the same as that used for the transient analysis (Section 7.1.1). The
{
initial conditions and modeling assumptions discussed in Section 7.1.2 are I applicable to this simulation.
I 2810R l 4
1 1
)
j The transient is initiated by a simultaneous closure of all four- I MS1Vs. A 3.0 second closing time, which is the Technical Specification [8]
minimum, is assumed. .A reactor scram signal is generated on APRM high flux.
Control rod drive motion is conservatively assumed to occur 0.28' seconds after reaching'the high flux setpoint. The system response'is shown in Figure 7.3.1 for..the " Measured" scram time analysis.
The maximum pressures at the bottom of the reactor vessel calculated.
for the " Measured" scram time analysis and for the "67B" scram time' analysis are given in Table 7.3.1. These results are within the allowable code limit 4 of 10% above-vessel design pressure for upset conditions, or 1375 psig. {
~.
7.4 Local Rod Withdrawal Error Transient Results i The rod withdrawal error (RWE) is a local core transient caused by an operator erroneously withdrawing a control rod in the continuous withdrawal mode. If the core is operating at its operating limits for MCPR and LHGR at the time of the error, then withdrawal of a control rod could increase both local and core power levels with the potential for overheating the fuel.
There is a broad spectrum of core conditions and control rod patterns which could be present at the time of such an error. For most normal situations it would be possible to fully withdraw a control rod without exceeding 1% clad plastic strain or violating the CPR based fuel cladding integrity safety limit.
To bound the most severe of postulated rod withdrawal error events, a portion of the core MCPR operating limit envelope is specifically defined such that the cladding limits are not violated. The consequences of the error depend on the local power increase, the initial MCPR of the neighboring locations and the ability of the Rod Block Monitor-(RBM) System to stop the 4 withdrawing rod before MCPR reaches 1.07.
The most severe transient postulated begins with the core operating !
according to normal procedures and within normal operating limits. The
~ ~
2810R l
l operator makes a procedural error and attempts to fully withdraw the maximum worth control rod at maximum withdrawal speed. The core limiting locations are close to the error rod. They experience the spatial power shape transient as well as the overall core power increase. )
The core conditions and control rod pattern are conservatively modeled for the bounding case by specifying the following set of concurrent worst case assumptions:
- 1. The rod should have high reactivity worth. This is provided for by analysis of the core at several exposure po.ints around the core peak reactivity. The test patterns are developed with xenon-free conditions. The xenon free condition and the additional control rod inventory needed to maintain criticality exaggerates the worth ;
of the withdrawn control rod substantially when compared to normal operation with normal xenon levels.
I
- 2. The core is initially at 104.5% power and 100% flow.
I I
- 3. The core power distribution is adjusted with the available control l rods to place the locations within the four by four array of bundles around the error rod as close to the operating limits as possible.
4 Of the many patterns tested, the pattern with the highest SCPR (
1 results is selected as the bounding case. 1 The Rod Block Monitor System's ability to terminate the bounding case is evaluated on the following bases: )
- 1. Technical Specifications [8] allow each of the separate RBM j channels to remain operable if at least half of the Local Power Range Monitor (LPRM) inputs at every level are operable. For the interior RBM channels tested in this analysis, there are a maximum of four LPRM inputs per level. One RBM channel averages the inputs 2810R l l
from the A and C levels; the other channel averages the inputs from the B and D levels. Considering the inputs for a single channel, there are eleven failure combinations of none, one and two failed
-) i i
LPRM strings. The RBM channel responses are evaluated separately j at these eleven input failure conditions. Then, for each channel l taken separately, the lowest response as a function of error rod-position is chosen for comparison to the RBM setpoint.
- 2. The event is analyzed separately in each of the four quadrants of the core due to the differing LPRM string physical locations relative to the error rod.
Technical Specifications require that both RBM channels be operable during normal operation. Thus, the first channel calculated to intercept the
-RBM setpoint is assumed to stop the rod. To allow for control system delay times, the rod is assumed to move two inches after the intercept and stop at the following notch.
The analysis is performed using SIMULATE [4]. Necessary properties of that model for use in this analysis are:
- 1. Accurate bundle power calculation as shown by the PDQ and gamma scan comparisons.
comparisons.
l l
- 3. Accurate control rod worths and core power coefficient as shown by the consistent core eigenvalues.
l Two separate cases are presented f rom numerous explicit SIMULATE analyses. The reactor conditions and case descriptions are shown in Figures 7.4.1 and 7.4.2. Case 1 analyzes the bounding event with zero xenon at the most reactive point in the cycle for the worst case abnormal rod pattern configuration. Case 2 is the worst of the 104.5% power conditions'modeled 2810R ._. . _ . ._ .. _ _ . _ _ _ . . _ . .. .. .. ..
i l
I with more normal control rod patterns and equilibrium xenon. The transient i
results, the ACPR and maximum linear heat generation rate (MLHGR) values, are also shown in Figures 7.4.1 and 7.4.2. The ACPR values are evaluated such that the implied operating limit MCPR equals 1.07 + ACPR. This is done by conserving the figure of merit (ACPR/ initial CPR) shown by the SIMULATE calculations. The use of this method provides valid ACPR values in the analysis of normal operating states where locations near the assumed error rod l are not initially near the MCPR operating limit.
1 Case 2 is the worst of all the rod withdrawal transients analyzed from 104.5% power, full flow and normal rod pattern conditions. Case 2 is bounded by Case 1 with substantial MCIR margin. The Case 1 RBM channel responses are shown in Figures 7.4.3 and 7.4.4. They also show the control rod position at the point where the weakest RBM channel response first intercepts the RBM setpoint. For this same bounding case, the operating limit ACPR envelope component versus RBM setpoint is taken from Figure 7.4.1. The same figure shows the resultant LHGR assuming the limiting bundle is placed on the operating limit of 13.4 kW/ft prior to the withdrawal. The calculation includes the 2.2% power spiking penalty. The limiting bundle MLHGR demonstrates margin to the 1% plastic strain limit given the low exposure of the bundle. High exposure bundles which have low 1% plastic strain limits are never limiting.
7.5 Misloaded Bundle Error Analysis Results 7.5.1 Rotated Bundle Error The primary result of a bundle rotation is a large increase in local j I
pin peaking and R-factor as higher enrichment pins are placed adjacent to the surrounding wide water gaps. In addition, there may be a small increase in reactivity, depending on the exposure and void fraction states. The R -f actor 1
increase results in a CPR reduction, while the local pin peaking factor increase results in a higher pin LHGR. The objective of the analysis is to .
l insure that in the worst possible rotation, the LHGR and CPR safety limits are not violated with the most limiting monitored bundles on their operating limits.
a ir .. n.
E________________
To analyze the CPF response, rotated bundle R-factors as a function of exposure are developed by adding the largest possible AR-factor resulting from a rotation to the exposure dependent R-factors of the properly oriented bundles [13). Using these rotated bundle R-factors, the MCPR values resulting from a bundle rotation are determined using SIMULATE. This is done for each control rod sequence throughout the cycle. The process is repeated with the K-infinity of the limiting bundle modified slightly to account for the increase in reactivity resulting from the rotation. For each sequence, the MCPR for the properly oriented bundles is adjusted by a ratio necessary to place the corresponding rotated CPR on its 1.07 safety limit. The maximum of these adjusted MCPR's is the rotated bundle operating limit.
'Io determine the MLHGR resulting from a rotation, the ratios of the maximum rotated bundle local peaking factor to the maximum properly oriented bundle local peaking are determined for the expected range of exposure and void conditions. The maximum of this ratio is applied to the 1.HGR operating limit of 13.4 kW/ft. This maximum rotated bundle LHCR is, in addition, modified to account for the possible reactivity increase resulting from the rotation. It is also increased by the 2.2% power spiking penalty.
l The results of the rotated bundle analysis are given in Table 7.5.1.
Given the low exposure of the bundle, there is sufficient margin to the 1%
plastic strain limit.
7.5.2 Mislocated Bundle Error Misloading a high reactivity assembly into a region of high neutron importance results in a location of high relative assembly average power. ,
Since the assembly is assumed to be properly oriented (not rotated), R-factors l used for the misloaded bundle are the standard values for the fuel type. l The analysis uses multiple SIMULATE [4] cases to examine the effects of explicitly dislocating every older interior assembly in a quarter core with a fresh assembly. Because of symmetry, the results apply to the whole core.
Edge bundles are not examined because they are never limiting, due to neutron I leakage. l
?
l 2810R
The effect of the successive dislocations is examined for every control rod sequence throughout the cycle. For each sequence, the MCPR for the ;
properly loaded core is compared to the misloaded core at the misloaded location. The MCPR for the properly loaded core is adjusted by a ratio necessary to place the mislocated assembly on the 1.07 safety limit. The maximum of these adjusted MCPRs is the mislocated bundle operating limit.
Using the above procedure, all possible dislocations result in calculated operating limits well below that set by the rotated bundle analysis. This makes the mislocated bundle analysis less limiting than the rotated bundle analysis given in Section 7.5.1.
7.6 Control Rod Drop Accident Results The control rod sequences are a series of rod withdrawal and banked I withdrawal instructions specifically designed to minimize the worths of individual control rods. The sequences are examined so that, in the event of the uncoupling and subsequent free fall of the rod, the incremental rod worth is acceptable. Incremental rod worth refers to the fact that rods beyond Group 2 are banked out of the core and can only fall the increment from all in to the rod drive withdrawal position. Acceptable worth is one which produces a maximum fuel enthalpy less than 280 calories / gram.
Some out-of-sequence control rods could accrue potentially high ]
worths. However, the Rod Worth Minimizer (RWM) will prevent withdrawing an out-of-sequence rod, if accidentally selected. The RWM is functionally tested before each startup.
1 The sequence in the RWM will take the plant from All Rods In (ARI) to well above 20% core thermal power. Above 20% power even multiple operator errors will not create a potential rod drop situation above 280 calories per sram [18, 19, and 20]. Below 20% power, however, the sequences must be examined for incremental rod worth. This is done throughout the cycle using l the full core, xenon-free SIMULATE model (4].
2810R
)
Both the A and B sequences were examined. It was found that the highest worth rod was the first rod of the second group. Any of the first 1 1
four rod arrays shown in Figures 7.6.1 and 7.6.2 may be designated as the first group pulled. However, a specific second group must follow as Table 7.6.1 illustrates. For added conservatism, the highest worth rod in the second group was deliberately assigned to be the first rod pulled. This assures that in any sequence the worths will always be less than those calculated here. The results of the calculations, as presented in Table 7.6.2, fit under the bounding analysis.
Beyond Group 2, procedures [21] apply which severely reduce the rod incremental worths. Therefore, the xenon-free, hot standby worths are much less than the cold, xenon-free worths [9].
l l
l 1
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TABLE 7.1.1 VY CYCLE 13
SUMMARY
OF SYSTEM TRANSIENT MODEL INITIAL CONDITIONS FOR TRMVSIENT ANALYSES Core Thermal Power (MWth) 1664.0 Turbine Steam Flow (% NBR) 105 Total Core Flow (10 61bm/hr) 48.0 !
Core Bypass Flow (1061bm/hr) 5.3 Core Inlet Enthalpy (BTU /lbm) 520.9 Steam Dome Pressure (psia) 1034.7 Turbine Inlet Pressure (psia) 986.0 Total Recirculation Flow (10 61bm/hr) 23.4 Core Plate Differential Pressure (psi) 18.5 Narrow Range Water Level (in.) 35 Average Fuel Gap Conductance (See Section 4.2) 2810R
TABLE 7.1.2 l 1
~i j
VY CYCLE 13 TRANSIENT ANALYSIS REACTIVITY COEFFICIENTS AT SELECTED CONDITIONS Cycle Exposure Point (MWD /ST) 4 Calculated Parameter EOFPL E0FPL-1000 EOFPL-2000 BOC Axial Shape Index(l) -0.0852 -0.2013 -0.2176 -0.1371 Moderator Density Coefficient 21.18 21.90 23.44 19.65 (Subcooling), d/Au(2)
Pressure = 1050 psia Subcooling = 30 BTU /lbm Moderater Density Coefficient 23.31 23.36 24.82 (3)
(Pressurization), d/Au Pressure = 1050 psia Inlet Enthalpy = 520 BTU /lbm l
Fuel Temperature Coefficient -0.292 -0.294 -0.297 -0.270 ;
at 11300F, d/0F !
Effective Delayed 0.005421 0.005505 0.005571 0.006115 Neutron Fraction i
Prompt Neutron Generation 41.46 41.48 40.72 38.24 1 Time in Microseconds P -P Notes: (1) Axial Shape Index (ASI) = p p T B (2) Au = change in density, in percent (3) Pressurization transients are not calculated at BOC l I
l 2810R TABLE 7.2.1 VY CYCLE 13 TRANSIENT ANALYSIS RESULTS Feak Peak Avg.
Prompt Power Heat Flux (Fraction of (Fraction of Transient Exposure Initial Value) Initial Value) ACPR Turbine Trip EOFPL 2.732 1.236 .20 Without Bypass,
" Measured" EOFPL-1000 2.155 1.152 .14 Scram Time E0FPL-2000 1.409 1.023 .03 Turbine Trip E0FPL 3.026 1.278 .25 Without Bypass, "67B" EOFPL-1000 2.518 1.208 .20 Scram Time EOFPL-2000 1.760 1.077 .07 Generator Load EOFPL 2.695 1.224 .19 Rejection Uithout Bypass, E0FPL-1000 2.116 1.141 .13
" Measured" Scram Time E0FPL-2000 1.333 1.003 .01 Generator Load EOFPL 3.101 1.280 .25 Rejection Without Bypass. EOFPL-1000 2.565 1.210 .20 "67B" Scram Time EOFPL-2000 1.742 1.065 .06 Loss of 1000F E0FPL
- 1.213 1.205 .17 Feedwater Heating EOFPL-1000 1.222 1.214 .18 j
EOFPL- 2000 1.223 1.216 .18 BOC 1.211 3.203 .17 I
2810R
TABLE 7.3.1 j VY CYCLE 13 OVERPRESSURE 2ATION ANALYSIS RESULTS q i
Maximum Pressure at Reactor j Conditions Vessel Bottom (psia) ]
" Measured" Scram Time 1265 "67B" Scram Time 1282 l
l TABLE 7.5.1 VY CYCLE 13 ROTATED BUNDLE ANALYSIS RESULTS 1
Resulting ;
Initial MCPR Resulting MCPR LHGR (Kw/ft) ;
i 1.20 1.07 17.20 f I
i l
l I
2810R
l TABLE 7.6.1 CONTROL ROD DROP ANALYSIS - ROD ARRAY PULL ORDER l
l The order in which rod arrays are pulled is specific once the choice of first group is made.
First Group Second Group Successive Group 1 Pulled is: Pulled Must Be: Is Banked Out Arrey 1 Array 2 Array 3 or 4 Array 2 Array 1 Array 3 or 4 Array 3 Array 4 Array 1 or 2 l Array 4 Array 3 Array 1 or 2 TABLE 7.6.2 VY CYCLE 13 CONTROL ROD DROP ANALYSIS RESULTS Maximum Incremental Rod Worth .79% AK Calculated Cold, Xenon Free Bounding Analysis Worth for Enthalpy 1.30% AK Less than 280 Calories per Gram (References 18, 19, and 20) l I
I l
j 3810R
Choose ICPR
~
I f Estimate Power a If Estimate Flow With FIBWR If RETRAN Flow initialization Run If is Loss No Revise Coefficient Correct? Flow Yes If RETRAN/TCPYACI Hot Channel Run if l
Choose Has ACPR No Converged? New ICPR Yes STOP TICURE 7.1.1 FLOW GIAkt FOR THE CALCULATION OF SCPR USING THE RETRAN/TCPYA01 CODES VY CYCLE 13 - MST, EOFPL 9
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F100KE 7.1.2 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL13, " MEASURED" SCRAM TIME 1
l VY CYCLE 13 - MST, ECPPL-1 8
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FIGURE 7.1.3 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL13-1000 MWD /ST
" MEASURED" SCRAM TIME
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! FIGURE 7.1.4 l INSERTED ROD WORTH AND ROD POSITIO'N VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL13-2000 MWD /ST j
" MEASURED" SCRAM TIME ,
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FIGURE 7.1._5 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL13, "67B" SCRAM TIME i
____._____.._.______________O
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FIC'JRE 7.1.6 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL13-1000 MWD /ST "67B" SCRAM TIME
VY CYCLE 13 - 678 SCRAM, E0FPL-2 9
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INITIAL COND1TIONS 43 39 16 28 .16 35 24 10 10 24 31 16 40 38 40 16 27- 10 0 0 10 28 23- 28 38 38 19 - 10 0 0 10 15 16 40 38 40 16 {
11 24 10 10 24 l 07 16 28 16 0'
, i i i 02 06 10 14 18 22 26 30 34 38'42 Core Therrrel Power = 1664 Wt Core Average Pressure = 1042 pelo Core Flow =
48 Mlb/hr Initial WCPR = 1.287 Cycle Exposure = 8240 Wd/st initial MLHCR = 13.4 kw/ft Xenon Free RWE Control Rod = 26-27 TRANSIENT
SUMMARY
RBM Rod MLHCR Seteeint Positten ACER (kw/f ti 104 10 0.13 13.9 105 10 0.13 13.9 106 12 0.16 14.4 107 14 0.18 15.0 !
108 16 0.21 15.6 !
FIGURE 7.4.1 REACTOR INITIAL CONDITIONS AND TRANSIENT
SUMMARY
FOR THE VY OYCLE 13 ROD WITHDRAWAL ERROR CASE 1 1
]
INITIAL CONDITIONS. l 43 39 42 42 35 31 26 16 16 26 !
27-23- 20 4 4 20 19 -
15 26 16 16 26 11 07 -
42 42 03 --
i i i 02 06 10 14 18 22 26 30 34 38 42 Core Thermal Power = 1664 Wt Core Average Pressure = 1042 psic Core Flow =
48 Wlb/hr Initial WCPR = 1.414 Cycle Exposure - 4160 Wd/st initial WLHGR = 12.1 kw/f t l Equiller lurn Xenon RWE Control Rod 23 TRANSIENT
SUMMARY
RBM Rod MLHGR Seteeint Position M (kw/f t)
I 104 12 0.07 12.8 105 12 0.07 12.8 106 16 0.11 14.0 107 20 0.14 14.7 108 26 0.17 14.6 FIGURE 7.4.2 REACTOR INITIAL CONDITIONS AND TRANSIENT
SUMMARY
FOR THE VY CYCLE 13 ROD WITHDRAWAL ERROR CASE 2
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43 3 39 2 1 '1 2 35 4 3 4 3 4 31 1 2 2 1 27 - 3 4 3 4 3' 23 - 1 2 1 1 2 1 10 - 3 4 3 4 3 15 1 2 2 1 11 4 3 4 3 4 07 2 1 1 2 03 3 1 l 1 02 06 10 14 18 22 26 30 34' 38 42 RGURE 7.6.1 RRST FOUR ROD ARRAYS PULLED IN THE A SEQUENCES, 43 3 3 39 2 1 2 35 3 4 4 3 31 2 1 2 1 2 27 - 3 4 3 3 4' 3 23 - 1 2 1 2 1 10 - 3 4 3 3 4 3 15 2 1 2 1 2 11 3 4 4 3 07 2 1 2 03 3 3 I I I 02 06 10 14 18 22 26 30 34 38 42 FIGURE 7,6,2 RRST FOUR ROD ARRAYS PULLED IN THE B SEQUENCES
8.0 LOSS-OF-COOLANT ACCIDENT ANALYSIS The results of the complete evaluation of the loss-of-coolant accident for Vermont Yankee, as documented in Reference 22, provide the required support for the operation of the Reload Cycle. The MAPLHGR limits for the new feel type, as a function of average planar exposure, are provided in Table A.2 of Appendix A.
2810R , ,
I
i 1
9.0 STARTUP PROGRAM 1
I l
Following refueling and prior to vessel reassembly, fuel assembly position and orientation- sill be verified and videotaped by underwater <
l television.
l l
l The Vermont Yankee Startup Program will include process computer data checks, shutdown margin demonstration, in-sequence critical measurement, rod q scram tests, power distribution comparisons, TIP reproducibility, and TIP symmetry checks. The content of the Startup Test Report will be similar to >
l that sent to the Office of Inspection and Enforcement in the past [23].
(
)
l i
i 2810R ,
i l
1
l REFERENCES l
- 1. M. A. Sironen and R. C. Potter, Vermont Yankee Cycle 10 Summary Report, YAEC-1438, September 1984.
- 2. R. C. Potter, Vermont Yankee Cycle 11 Summary Report, YAEC-1513, February 1986.
- 3. General Electric Standard Application for Reactor Fuel (GESTARII),
NEDE-24011-P-A-8, GE Company Proprietary, May 1986, as amended.
l 4. D. M. VerPlanck, Methods for the Analysis of Boiling Water Reactors Steady State Core Physics, YAEC-1238, March 1981.
- 5. E. E. Pilat, Methods for the Analysis of Boiling Water Reacters Lattice Physics, YAEC-1232, December 1980.
- 6. S. P. Schultz and K. E. St. John, Methods for the Analysis of Oxide Fuel Rod Steady-State Thermal Effects (FROSSTEY) Code /Model Description Manual, YAEC-1249P, April 1981.
- 7. S. P. Schultz and K. E. St. John, Methods for the Analysis of Oxide Fuel Rod Steady-State Thermal Ef fects (FROSSTEY) Code Qualification and Application YAEC-1265P, June 1981.
S. Appendix A to Operating License DPR-28 Technical Specifications and Bases for Vermont Yankee Nuclear Power Station, Docket No. 50-271.
- 9. A. A. F. Ansari, et al., Vermont Yankee Cvele 9 Core Performance Analysis, YAEC-1275, August 1981.
- 10. M. Edenius, A. Ahlin, H. Haggblom, CASMO-2: A Fuel Assembly Burnup Program, STUDSVIK/NR-81/3, Proprietary Studsvik Report.
- 11. A. A. F. Ansari, Methods for the Analysis of Boiling Water Reactors:
Steady-State Core Flow Distribution Code (FIBWR), YAEC-1234, December 1980.
- 12. A. A. F. Ansari, R. R. Gay, and B. J. Gitnick, FIBWR: A Steady-State Core Flow Distribution Code for Boiling Water Reactors - Code Verification and Qualification Report EPRI NP-1923, Project 1754-1 Final ,
Report, July 1981. l
- 13. General Electric Company, GEXL Correlation Application to BWR 2-6 Reactors, NEDE-25422, GE Company Proprietary, June 1981. )
- 14. A. A. F. Ansari and J. T. Cronin, Methods for the Analysis of Boiling Water Reactors: A Systems Transient Analysis Model (RETRAN), YAEC-1233, April 1981.
I
- 15. EPRI, RETRAN - A Program for One-Dimensional Transient Thermal-Hydraulic i Analysis of Complex Fluid Flow Systems, CCM-5, December 1978.
l l
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l
_ ______________D
- 16. A. A. F. Ansari, K. J. Burns, and D. K. Beller, Methods for the Analysis of Boiling Water Reactors: Transient Critical Power Ratio Analysis (RETRAN-TCPYA01), YAEC-1299P, March 1982.
- 17. J. M. Holzer, Methods for the Analysis of Boiling Water Reactors Transient Core Physics, YAEC-1239P, August 1981.
- 18. C. J. Paone, et al., Rod Drop Accident Analysis for Large Boiling Water Reactors, NEDO-10527. March 1972.
- 19. R. C. Stirn, et al., Rod Drop Accident Analysis for Large Boiling Water Reactors Addendum No. 1, Multiple Enrichment Cores With Axial Gadolinium, NEDO-10527, Supplement 1, July 1972.
- 20. R. C. Stirn, et al., Rod Drop Accident Analysis for Large Boiling Water Reactor Addendum No. 2 Exposed Cores, NED0-10527, Supplement 2, January 1973.
- 21. D. Radcliffe and R. E. Bates, " Reduced Notch Worth Procedure", SIL-316, November 1979.
- 22. Loss-of-Coolant Accident Analysis for Vermont Yankee Nuclear Power Station, NED0-21697, August 1977, as amended.
- 23. Letter, FVY 86-92, dated October 6, 1986, R. W. Capstick to T. E. Murley, Regional Administrator, " Cycle 12 Startup Test Report".
)
1 2810R __ __ __________-__ _ _____-_.
APPENDIX A l
I CALCULATED OPERATING LIMITS The MCPR limits appropriate for the Reload Cycle are calculated by adding the calculated ACPR to the safety limit LAMCPR of 1.07. This is done for each of the analyses in Section 7 at each of the exposure statepoints. ;
For an exposure interval between statepoints, the highest MCPR limit at either end.is assumed to apply to the whole interval.
Table A.1 provides the highest calculated MCPR limits for the Reload Cycle for each of the exposure intervals for the various scram speeds and for the various rod block lines. Ca Table A.1, as in the Technical Specifications, End of Cycle (EOC) is understood to mean End of Full Power I
Life (EOFPL).
The MAPLHGR limits, as a function of average planar exposure, for the new fuel type are provided in Table A.2.
A-1 2810R
)
2
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t i
m .
i RL 888882557 57 2557 57 2 P 222 223 222223222223 . .
C g .
M n 1 1 1 1 1 1 1 1 1 11 1 1 1 1 1 1 1 i e l
_ t a b a
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O n i
n w
o T T T T T T h
/ / / / / / s D D D D D D W W W W W W n G G G G G G o i
1 1 1 1 1 1 t .
e - - - - - - a n g TCCTCCTCCTCCTCCTCC u o
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a O eR W W W e r G ooG ooG ooG ooG ooGt too
- I TS l
c e yr 2 t t 2
t t 2
t t 2 2 t t 2
t t h t
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RL E
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x EWWEWWEWWEWWEWWEWW o
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GC d e l o
WG n ON t 21 t 21 t 21 t 21 t 21 t 21 PI - - - - - - - - - - - - i e 1 T CCCCCCCCCCCCCCCCCC r m l g
RA OOOOOOOOOOOOOOOOOO e n A
AR BEEBEEBEEBEEBEEBEE t i EE e s E LP CO d L r B U NR e o A P r f T EC a 1
EM s 0 K t N3 n 0 A1 d Y o i y
E R o .
r r r r r r ps b TL e e NC l e e e e t n OY oe t t t t t t eo d MC rm t t t t t t si e R t i e . e . e . e . e . e . pa t s E nT a V o b O.1 b O. 2 b O. 1 b O. 2 b O.1 b O. 2 i c ri e
r Cm r C.1 ro C.1 ro C.1 ro C.1 ro C.1 ro C.1 a o . t f c er L L . L L . L . LC i n gc l Cl Cl Cl Cl Cl ) c i aS an an an an an an . M e B p e r ua3 ua3 ua3 qh ua3 ua3 ua3 qh qh RS r e qh qh . qh v Et3Et 3Et3Et3Et 3Et 3 ( a l
A ra s oc i t
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TABLE A.2 VERMONT YANKEE NUCLEAR POWER STATION MAPLHGR OPERATING LIMITS FOR BP8DRB299 MAPLHGR for BP8DRB299 Average Planar Exposure Two-Loop Single-Loop PCT 0xidation (mwd /St) Operation Operation _F_ Fraction 200 10.70 8,8 2030 0.019 1,000 10.80 8.9 2037 0.019 5,000 11.40 9.4 2093 0.023 ;
10,000 12.20 10.1 2178 0.029 15,000 12.30 10.2 2198 0.031 20,000 12.20 10.1 2193 0.031 25,000 11.70 9.7 2139 0.026 ;
35,000 10.60 8.8 1972 0.028 41,900 9.40 7.8 1800 0.012 l
l 1
1 A-3 2810R