ML20199E278

Cycle 12 Core Performance Analysis

ML20199E278
Person / Time
Site:	Vermont Yankee
Issue date:	11/30/1985
From:	Burns K, Chandola V, Heinrichs D VERMONT YANKEE NUCLEAR POWER CORP.
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ML20199E275	List: ML20199E271 ML20199E278
References
YAEC-1507, NUDOCS 8603260006
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n VERMONT YANKEE I CYCLE 12 CORE PERFORMANCE ANALYSIS November 1985 Major Contributors:

K. J. Burns J. D. Robichaud I V. Chandola D. P. Heinrichs K. E. St. John T. A. Schmidt D. M. Kapitz M. A. Sironen I J. Pappas R. J. Weader R. A. Woehlke I Approved by: d R J.Cacc$outi, Manager R actor Phyisics Group

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/ (Daf.e) 7 #d

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Approved by: . h #

i'. A. Be'rgeronganager (D' ate)

Transient Analysis Group Approved by: iM .

is 27/ts S.#P.' Schultz, Manager D (Date)

Nuclear Evaluations and Support Group Approved by: N me t ( w~nat b (I!27 ? ')

A. Husain, Manager (Date)

LOCA Group Approved by: -

II!l7 /88 B.C.Slifer,DirQtor (Date)

Nuclear Engineering Department 8603260006 860313 PDR ADOCK 05000271 p PDR

I DISCLAIMER OF RESPONSIBILITY J

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 Nuclear Regulatory Conunission only.

I With regard to any unauthorized use whatsoever, Yankee Atomic Electric Company, Vermont Yankee Nuclear Power 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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ABSTRACT I This report presents design information and calculational results pertinent to the operation of Cycle 12 of the Vermont Yankee Nucleae 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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I TABLE OF CONTENTS Page DISCLAIMER.................................................. 11 I ABSTRACT.................................................... iii TABLE OF CONTENTS........................................... iv LIST OF FIGURES.... ........................................ Vi LIST OF TABLES.............................................. viii ACKN0WLEDGEMENTS............................................ ix

1.0 INTRODUCTION

................................................ 1 2.0 RECENT REACTOR OPERATING HISTORY............................ 2 2.1 Operating History of the Current Cycle................. 2 2.2 Operating History of Past Applicable Cycles............ 2 3.0 RELOAD CORE DESIGN DESCRIPTION.............................. 6 I 3.1 3.2 3.3 Core Fuel Loading......................................

Design Reference Core Loading Pattern..................

Assembly Exposure Distribution.........................

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6 4.0 FUEL MECHANICAL AND THERMAL DESIGN.......................... 9 4.1 Mechanical Design..............................~......... 9 I 4.2 4.3 Thermal Design.........................................

Operating Experience...................................

9 10 5.0 NUCLEAR DESIGN.............................................. 15 5.1 Core Power Distributions............................... 15 i

5.1.1 Haling Power Distribution....................... 15 5.1.2 Rodded Depletion Power Distribution.............

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I 5.2 5.3 5.4 Core Exposure Distributions............................

Cold Core Reactivity and Shutdown Margin...............

Standby Liquid Control System Shutdown Capability......

16 16 17 6.0 THERMAL-HYDRAULIC DESIGN.................................... 26 6.1 Steady-State Thermal Hydraulics........................ 26 6.2 Reactor Limits Determination........................... 26 i

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I TABLE OF CONTENTS (Continued)

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7.0 ACCIDENT ANALYSIS........................................... 28 7.1 Core Wide Transient Analysis........................... 28 7.1.1 Methodology..................................... 28 7.1.2 Initial Conditions and Assumptions.............. 29 7.1.3 Reactivity Functions............................ 30 7.1.4 Transients Analyzed............................. 32 7.2 Core Wide Transient Analysis Results................... 32 7.2.1 Turbine Trip Without Bypass Transient........... 32 7.2.2 Generator Load Rejection Without Bypass Transient................................ 33 7.2.3 Loss of Feedwater Heating Transient............. 33

'/ . 3 Overpressurization Analysis Results.................... 34 7.4 Local Rod Withdrawal Error Transient Results........... 34 I 7.5 Misloaded Bundle Error Analysis Results................ 37 Rotated Bundle Error............................ 37 I

7.5.1 7.5.2 Mislocated Bundle Error......................... 38 7.6 Control Rod Drop Accident Results...................... 39 7.7 Stability Analysis Results............................. 40 8.0 STARTUP PROGRAM......................,.............. ........ 81 9.0 LOSS-OF-COOLANT ACCIDENT ANALYSIS........................... 82 REFERENCES.................................................. 83 I APPENDIK A CALCULATED CYCLE DEPENDENT LIMITS............... A-1 I

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I LIST OF FIGURES Number Title Page 3.2.1 W Cycle 12 Design Reference Loading Pattern, Lower Right Quadrant 8 4.2.1 W Cycle 12 Core Average Gap Conductance versus Cycle Exposure 13 4.2.2 W Hot Channel Gap Conductance for P8X8R versus Exposure 14 5.1.1 W Cycle 12 Haling Depletion, EOFPL Bundle Average Relative Powers 19 I 5.1.2 W Cycle 12 Haling Depletion, EOFPL Core Average Axial Power Distribution 20 5.1.3 W Cycle 12 Rodded Depletion - ARO at EOFPL, Bundle Average Relative Powers 21 5.1.4 W Cycle 12 Rodded Depletion - ARO at EOFPL, Core Average Axial Power Distribution 22 5.2.1 W Cycle 12 Haling Depletion, EOFPL Bundle Average Exposures 23 5.2.2 W Cycle 12 Rodded Depletion, EOFPL Bundle Average Exposures 24 I

5.3.1 W Cycle 12 Cold Shutdown Delta K in Percent versus Cycle Exposure 25 7.1.1 Flow Chart for the Calculation of- CPR Using the RETRAN/TCPYA01 Codes 46 7.1.2 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL12 " Measured" Scram Time 47 7.1.3 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL12-1000 MWD /ST, " Measured" Scram Time 48 7.1.4 Inserted Rod Worth and Rod Position versus Time From I Initial Rod Movement at EOFPL12-2000 MWD /ST, " Measured" Scram Time 49 Inserted Rod Worth and Rod Position versus Time From I

7.1.5 Initial Rod Movement at EOFPL12. "67B" Scram Time 50 7.1.6 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL12-1000 MWD /ST, "67B" Scram Time 51 7.1.7 Inserted Rod Worth and Rod Position versus Time From Initial Rod Movement at EOFPL12-2000 MWD /ST, "67B" Scram Time 52 I vi I

LIST OF FIGURES Number Title Page 7.2.1 Turbine Trip Without Bypass, EOFPL12 Transient Response versus Time, " Measured" Scram Time 53 7.2.2 Turbine Trip Without Bypass, EOFPL12-1000 MWD /ST Transient Response versus Time, " Measured" Scram Time 56 7.2.3 Turbine Trip Without Bypass, EOFPL12-2000 MWD /ST Transient Response versus Time, " Measured" Scram Time 59 7.2.4 Generator Load Rejection Without Bypass, EOFPL12 Transient Response versus Time, " Measured" Scram Time 62 7.2.5 Generator Load Rejection Without Bypass, EOFPL12-1000 MWD /ST Transient Response versus Time, " Measured" Scram Time 65 7.2.6 Generator Load Rejection Without Bypass, EOFPL12-2000 MWD /ST Transient Response versus Time, " Measured" Scram Time 68 I 7.2.7 Loss of 100 F0 Feedwater Heating, EOFPL12-1000 MWD /ST (Limiting Case) Transient Response versus Time 71 MSIV Closure, Flux Scram, EOFPL12 Transient Response I 7.3.1 versus Time, " Measured" Scram Time 73 7.4.1 Reactor Initial Conditions and Transient Summary for the W Cycle 12 Rod Withdrawal Error Case 1 76 7.4.2 Reactor Initial Conditions and Tr,ansient Summary for the W Cycle 12 Rod Withdrawal Error Casa 2 77 7.4.3 W Cycle 12 RWE Case 1 - Setpoint Intercepts Determined by the A+C Channel 78 7.4.4 W Cycle 12 RWE Case 1 - Setpoint Intercepts Determined by the B+D Channel 79 7.6.1 First Four Rod Arrays Pulled in the A Sequences 80 7.6.2. First Four Rod Arrays Pulled in the B Sequences 80 I

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LIST OF TABLES Number Title Page 2.1.1 VY Cycle 11 Operating Highlights 3 2.2.1 VY Cycle 10 Operating Highlights 4 2.2.2 VY Cycle 9 Operating Highlights 5 3.1.1 VY Cycle 12 Fuel Bundle Types and Numbers 7 3.3.1 Design Basis VY Cycle 11 and Cycle 12 Exposures 7 4.1.1 Nominal Fuel Mechanical Design Parameters 11 4.2.1 Cap Conductance Values Used in VY Cycle 12 Transient Analyses 12 4.2.2 Peak Linear Heat Generation Rates Corresponding to Incipient I Fuel Centerline Melting and 1% Cladding Plastic Strain 12 5.3.1 VY Cycle 12 K-Effective Values and Shutdown Margin Calculation 18 5.4.1 VY Cycle 12 Standby Liquid Control System Shutdown Capability 18 7.1.1 VY Cycle 12 Sumary of System Transient Model Initial Conditions for Core Wide Transient Analyses 41 I 7.1.2 VY Cycle 12 Transient Analysis Reactivity Coefficients at Selected Conditions 42 7.2.1 VY Cycle 12 Core Wide Transient nalysis Results 43 7.3.1 VY Cycle 12 Overpressurization Analysis Results 44 7.5.1 VY Cycle 12 Rotated Bundle Analysis Results 44 7.6.1 Control Rod Drop Analysis - Rod Array Pull Order 45 7.6.2 VY Cycle 12 Control Rod Drop Analysis Results 45 I A.1 Vermont Yankee Nuclear Power Station Limiting Cycle 12 MCPR Results A-2 A.2 Ver1nont Yankee Nuclear Powet Station Technical Specification MCPR Operating Limits A-3 I

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I ACKNOWLEDGEMENTS I The authors and principal contributors would like to acknowledge the contributior.s to this work by R. C. Potter and the YAEC Word Processing I Center. Their assistance in preparing figures and text for this document is recognized and greatly appreciated.

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1.0 INTRODUCTION

1 This report provides information to support the operation of the Vermont Yankee Nuclear Power Station through the forthcoming Cycle 12. In this report, Cycle 12 will frequently be referred to as the Reload Cycle. The preceding Cycle 11 will frequently be referred to as the Current Cycle. The refueling between the two will involve the discharge of 120 irradiated fuel bundles and the insertion of 120 new fuel bundles. The resultant core will consist of 120 new fuel bundles and 248 irradiated fuel bundles. Some of the irradiated fuel was present in the reactor in Cycles 9 and 10, as well as the Current Cycle. These cycles will frequently be referred to as Past Cycles.

This report contains descriptions and analyses results pertaining to the mechanical, thermal-hydraulic, physics, and safety aspects of the Reload Cycle (Cycle 12).

The cycle-dependent operating limits calculated for the Reload Cycle are bounded by the Vermont Yankee plant Technical Specifications. Both are given in Appendix A.

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I 2.0 RECENT REACTOR OPERATING HISTORY I 2.1 Operating History of the Current Cycle I The currently operating cycle is Cycle 11. The Current Cycle had a slow startup because of various mechanical problems. During the remainder of the cycle, the reactor operated smoothly at, or near, full power with the exception of normal maintenance, sequence exchanges, and a few unplanned reactor trips. The operating history highlights and control rod sequence exchange schedule of the Current Cycle is found on Table 2.1.1.

2.2 Operating History of Past Applicable Cycles The irradiated fuel in the Reload Cycle. includes some fuel bundles initially inserted in Cycles 9 and 10. These Past Cycles operated smoothly I at, or near, full power with the exception of normal maintenance, sequence exenanges, a few unplanned reactor trips, 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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l TABLE 2.1.1 VY CYCLE 11 OPERATING HIGHLIGHTS I Beginning of Cycle Date August 6, 1984 End of Cycle Date September 20, 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) 16,760.

End of Cycle Core Average Exposure (MWD /ST) 18,283.

Capacity Factor While Operating (%) 89.2 Number of Fresh Assemblies 104 Number of Irradiated Assemblies 264 I

Control Rod Sequence Exchange Schedule:

Sequence Date From To_

October 24, 1984 Al-1 B2-1

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December 15, 1984 February 2, 1985 A2-1 B1-1 March 23, 1985 B1-1 Al-2 May 18, 1985 Al-2 B2-2 I -

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I TABLE 2.2.1 I VY CYCLE 10 OPERATING HIGHLIGHTS I Beginning of Cycle Date June 17, 1983 End of Cycle Date June 15, 1984 I 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 I Control Rod Sequence Exchange Schedule:

Sequence Date From T J

I August 12, 1983 Al-1 B2-1

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October 1, 1983 B2-1 A2-1 November 5, 1983 A2-1 B1-1 B1-1 I

December 17, 1983 Al-2 January 23, 1984 Al-2 B2-2 March 3, 1984 B2-2 A2-2 April 16, 1984 A2-2 B1-2 I 4-I

I TABLE 2.2.2 I VY CYCLE 9 OPERATING HIGHLIGHTS I Beginning of Cycle Date December 1, 1981 End of Cycle Date March 5, 1983 Weight of Uranium As-Loaded (Short Tons) 74.15 l 1

Beginning of Cycle Core Average Exposure (MWD /ST) 9192.

End of Full Power Core Average Exposure (MWD /ST) 16595.

End of Cycle Core Average Exposure (MWD /ST) 18137.

Capacity Factor While Operating (%) 90.9 Number of Fresh Assemblies 120 Number of Irradiated Assemblies 248 I Control Rod Sequence Exchange Schedule:

Sequence Date From To_

January 28, 1982 Al-1 B2-1

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March 13, 1982 B2-1 A2-1 April 24, 1982 A2-1 B1-1 June 10, 1982 B1-1 I

Al-2 July 24, 1982 Al-2 B2-2 September 11, 1982 B2-2 A2-2 October 30, 1982 A2-2 B1-2 I

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I 3.0 RELOAD CORE DESIGN DESCRIPTION 3.1 Core Fuel Loading )

I The Reload Cycle core will consist of both new and irradiated I

assemblies. All the assemblies have bypass flow holes drilled in the lower 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 I The Reload Cycle assembly locations are indicated on the map in Figure 3.2.1. For the sake of legibility only the lower 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 are made to 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 with 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 March 26, 1985. 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 I 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 capability are provided.

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I TABLE 3.1.1 VY CYCLE 12 FUEL BUNDLE TYPES AND NUMBERS Fuel Reload Cycle Possible Designation Designation Loaded Number Bundle ID's Irradiated P8DPB289 R8 9 36 LJTXXX, LJZXXX P8DPB289 R9 10 108 LY4XXX P8DPB289 R10 11 104 LY6XXX, LY7XXX New P8DPB289 R11 12 120 LY7XXX, LYCXXX NOTE: XXX stands for the last three digits of the bundle serial number.

TABLE 3.3.1 DESIGN BASIS VY CYCLE 11 AND CYCLE 12 EXPOSURES Assumed Current Cycle core Average Exposure End of Cycle 11 ,

18.27 GWD/ST Assumed Reload Cycle Core Average Exposure Beginning of Cycle 12 9.85 GWD/ST Haling Calculated Core Average Exposure at End of Full Power Life, Cycle 12 16,85 GWD/ST Cycle 12 Exposure Capability 7.00 GWD/ST I  ;

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I I VERMONT YANKEE CYCLE 12 80C BUNOLE AVERAGE EXPC;URES R9 R9 RIO R9 R10 R11 R9 R11 R9 R11 R9 18 43 18.98 9 78 18.60 8.03 0 00 17.26 0 00 15 40 0.00 16.31 - 22 R9 RIO R11 R10 R8 RIO R11 R10 R11 R10 R9 18 89 10 04 0 00 9 21 20 33 6.97 0.00 8 55 0 00 10 02 18 00 - 20 R10 R11 49 R11 R10 R9 R10 R11 R9 RIO RB 9.79 0.00 17 73 0 00 8 50 18.62 10 36 0.00 15.85 10.01 23 42 - 18 R9 R10 Rtt R9 R11 R10 R11 RIO R11 RB 19 13 9.30 0.00 17.78 0 00 9.74 0.00 9.38 0.00 21 85 16 RIO R8 R10 Rtt R9 R11 R9 All R9 8 02 20.36 8 55 0.00 18.52 0 00 18.70 0.00 18.21 14 R11 R10 R9 Rio R11 RIO R11 R11 R9 0.00 6 98 18 60 9 84 0.00 10 07 0 00 0 00 16.23 12 R9 R11 R10 R11 R9 R11 R9 R10 R8 16.f 3 0.00 10.33 0 00 18 68 0.00 16.15 9 82 21.11 10 R11 RIO R11 RIO R11 R11 RIO R8 0.00 8.54 0.00 9 41 0.00 0 00 9.78 23 68 8 R9 R11 R9 R11 R9 R9 r.8 15.23 0.00 15.72 0 00 16.18 16 30 21 24 6 R8 - P8DPB289, RELOAD 8 0 00 9 99 9 97 21.73 R9 - P8DPB289, RELOAD 9 - 4 R10 - P8DPB289, RELOAD 10 R9 R9 R$ 8UNOLE 10 R11 - P8DPB289, RELOAD 11 18 25 18.03 23 39 EXPOSURE (OWO/ST) 2 I 23 I

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35 37 39 41 43 FIGURE 3.2.1 VY CYCLE 12 DESIGN REFERENCE LOADING PATTERN, LOWER RIGHT QUADRANT I 4.0 FUEL MECHANICAL AND THERMAL DESIGN I 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. 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 environment is also addressed in Reference 3.

I 4.2 Thermal Design The fuel thermal- effects calculations were performed using the FROSSTEY computer code [6-7]. The FROSSTEY code calculates pellet-to-clad gap conductance and fuel temperatures from a combination of theoretical and empirical models which include fuel and cladding thermal expansion, fission gas release, pellet swellini;, pellet densification, pellet cracking, and fuel and cladding thermal conductivity.

The thermal effects analysis includ'ed 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.

l 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 MApLHGR limit defined in Reference 8. Figure 4.2.2 provides the hot channel response of gap l

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conductance. In Figure 4.2.2, " planar exp3rura" refcrs to ths cxpecurs of th3 node running at the MAPLHGR limit. Gap conductance values for the hot channel analysis were extracted from Figure 4.2.2 using the maximum bundle exposure of any 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).

Fuel rod local linear heat generation rates at fuel centerline incipient melt and 1% clad plastic strain as a function of local axial segment exposure for the 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 linear heat generation rate (13.4 kW/ft) prior to the power increase.

4.3 Operating Experience All irradiated fuel bundles scheduled to be reinserted in the Reload

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Cycle have operated as expected in Past Cycles of Vermont Yankee. Off-gas measurements in the current Cycle were at normally low levels indicating that no fuel failures were present.

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TABLE 4.1.1 NOMINAL FUEL MECHANICAL DESIGN PARAMETERS FUEL TYPE P0X8R Vendor Designations (Table 3.1.1)

New P8DPB289 Irradiated P8DPB289 Fuel Pellets Fuel Material (sintered Pellets) UO2 and UO2 + Gd2 03 Initial Enrichment, w/o U-235 2.89 Pellet Density, % theoretical 96.5 Pellet Diameter, inches 0.410 Fuel Rod Active Length, inches 150.0 Fuel Rod Pitch, inches 0.640 Diametral Gap (cold), inches 0.009 Fill Gas Helium Fill Gas Pressure, psig (See Ref. 3)

Cladding Material Zr-2 Outside Diameter, inches "

0.483 Thickness, inches 0.032 Inside Diameter, inches 0.419 Fuel Channel Material Zr-4 Inside Dimension, inches 5.278 Wall Thickness, inches 0.080 Fuel Assembly Fuel Rod Array 8x8 Fuel Rods per Assembly 62 1 Spacer Grid Material Zr-4

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GAP CONDUCTANCE VALUES USED IN VY CYCLE 12 TRANSIENT ANALYSES 1

Cycle Exposure Core Average Hot Channel Hot Channel Statepoint Gap Conductance Bundle Exposure Gap Conductance (MWD /ST) (BTU /Hr-Ft - F) (MWD /ST) (BTU /Hr-Ft - F)

BOC12 1095 11029(1) 2140 EOFPL12-2000 MWD /ST 1410 6806 1850 EOFPL12-1000 MWD /ST 1475 7777 2050 1 EOFPL12 1485 8855 2100 NOTE (1) Between BOC12 and EOFPL12-2000 MWD /ST, the highest exposure limiting hot channel bundle is once-burned.

TABLE 4.2.2 PEAK LINEAR HEAT GENERATION RAfrS CORRESPONDING TO INCIPIENT FUEL CENTERLINE MELTING AND 1E CLADDING PLASTIC STRAIN (1) 1 0.0 w/o Gd 0 * "

23 23 Exposure Melt 1% (p Melt 1% (p (MWD /MT) (kW/ft) '(kW/ft) (kW/ft) (kW/ft)

Fuel Type P8x8R ,

O 21.5 21.5 21.5 21.5 25,000 21.5 21.5 20.5 21.5 50,000 21.5 17.0 .19.5 14.5 NOTE (1) Peak linear heat generation rates shown are minimum bounding values to the occurrence of the given condition.

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m m m- M M M M M M m M M M M M M M M M VY CYCLE 12 FUEL PERFORMANCE CORE AVERAGE GAP CONDUCTANCE

^ 1500 W /

n 1450 l

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m 155o

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W

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O 1250 7 8

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1200

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• 115 0 W

M nw e l w

3 1050 1 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 5.5 6 6.5 7 CYCLE EXPOSURE (GWD/ST)

FICURE 4.2.1 ,

VY CYCL.E l2 CORE AVERACE CAP CONDifCTANCE VERSlfS CYCLE EXPOSITRE

m M M M M M M M m m M M M M m m m M M VERMONT YANKEE - HOT CHANNEL GAP CONDUCTANCE P8X8R FUEL -- GAP CONDUCTANCE VS EXPOSURE 1.4 CH0PPED COSINE AXtAL POWER SHAPE WITN PEAK AT MAPLHo#

2250

-/N l 7 N

[ 2000 f x I

% N N

k #

1750 w /

i N g 1500 /

% D '

i O /

1250 f k

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3

- 1000 /

0 0

3:

750 , , ,

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 PLANAR EXPOSURE (GWD/ST) b 1 2 3 4 $ $ 7 b h 10 11 12 13 14 15 16 f7 18 19 2O BUNDLE EXPOSURE (GWD/ST)

FIGURE 4.2.2 VY 110T CllANNEL CAP CONDl'CTANCE FOR P8X8R VEttSUS EXPOSURE

5.0 NUCLEAR DESIGN lI a.1 Core Power Distributions I The Reload Cycle was depleted using SIMULATE (4) to give both a rodded depletion and an All Rods Out (ARO) Haling depletion. 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 l conservatively weak scram characteristics. The rodded depletion was used to evaluate the misloaded bundle error and the rod withdrawal error. This is because of the more realistic predictions it makes of initial CPR values. It was also used in the rod drop worth and shutdown margin calculations because I it burns the top of the core more realistically than the Haling.

5.1.1 Haling Power Distribution The Haling power distribution is calculated in the All Rods Out (ARO) condition. The Haling iteration converges on a self-consistent power and exposure shape 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 I 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 To generate the rodded depletion, control rod patterns were developed

~w hich gave critical eigenvalues at each point in the cycle and gave peaking l

similar to the Haling calculation. The resulting patteres were frequently I more peaked than the Haling, but were below expected operating limits.

l 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 ARO at EOFPL power distributions for the Roload Cycle rodded depletion. Note in Figure 5.1.4 that the average axial power at ARO I .

I fcr th] r:dd:d d:plc,ticn 10 mora b3ttom p;rk:d than tha H211ng (Figura 5.1.2). The rodded depletion would result in,better scram characteristics at EOFPL.

I 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 EOTPL I 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 of power shaping by the rods. 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 EOFPL radial exposure distribution for the Reload Cycle rodded depletion.

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I 5.3 Cold Core Reactivity and Shutdown Margin The cold K,gg with ARO and the cold K g with All Rods Inserted I (ARI) at BOC were calculated using the SIMULATE code [4,5) and are shown in Table 5.3.1. K,gg with ARO minus the cold critical K gg is the amount of excess core reactivity. K g with ARI minus the K gg with ARO is the worth of all the control rods.

I The cold critical eigenvalue K,gg was defined as the average calculated critical eigenvalue minus a 95% confidence level uncertainty. Then all cold results were normalized to make the critical K,gg eigenvalue equal to 1.000.

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 I I

I search was made for the strongest worth control rod at various exposurcs in the cycle. This is necessary because rod worths change with exposure. Then the cold K gg with the strongest rod out was calculated at BOC cnd at the end of each control rod sequence. Subtracting each cold Kg with the strongest rod out from the cold critical K,gg eigenvalue defines the shutdown margin as a function of exposure. Figure 5.3.1 shows the result.

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 R plus R . R is the difference 2

between the cold K,gg with the strongest rod out at BOC and the maximum cold K with the strongest rod out in the cycle. R is a measurement g 2 uncertainty in the demonstration of SDM. 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 System (SLCS) is designed to bring the reactor from full power to cold, ARO, xenon free shutdown with at least 5% AK margin. Using the boron concentration search option in SIMULATE [4), the ppm of boron was adjusted until the K,gg reached the cold critical K g minus .05. This case assumed cold, xenon free conditions, with All Rods Out at the most ' reactive time in the cycle. The criticality search found that the plant would be suberitical by 5% AK at the worst point in time 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.

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I TABLE 5.3.1 VY CYCLE 12 K gg VALUES AND SHUTDOWN MARGIN CALCULATION l

BOC eK gg - Uncontrolled 1.1186 i BOC eK gg - Controlled .9714 l Cold Critical K gg e Eigenvalue 1.0000 BOC Keff - Controlled With .9905 Strongest Worth Rod Withdrawn I Cycle Minimum Shutdown Margin occurs at BOC With Strongest Worth Rod Withdrawn 0.95% AK I R1 , Maximum Increase in Cold K egg With Exposure

.00% AK I

TABLE 5.4.1 VY CYCLE 12 I STANDBY LIQUID CONTROL SYSTEM SHUTDOWN CAPABILITY ppm of Boron -

ShutdownMarkin I 693 5.0% 6K 800 7.1% AK I

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I I VERMONT YANKEE CYCLE 12 HALINO DEPLET!ON EOFPL BUNDLE AVERADE RELATIVE POWERS I

R9 R9 RIO R9 R10 R11 R9 R11 R9 R11 R9 0 94 0.98 1 14 1.05 1 20 1.33 1.09 1.26 0 96 0.93 0 62 - 22 R9 R10 R11 R10 RB R10 R11 RIO R11 RIO R9 0 99 1 14 1.30 1 19 1 04 1 22 1 30 1.14 1 11 0 81 0.47 - 20 RIO R11 R9 R11 RIO R9 R10 R11 R9 R10 RB 1 14 1.30 1.11 1 33 1 21 1 07 1 16 1 21 0 89 0.71 0.37 - 18 R9 RIO R11 R9 R11 RIO R11 RIO R11 R8 1.05 1.19 1 33 1 12 1 33 1 18 1 26 1 05 0 93 0.52 16 R10 R8 RIO R11 R9 R11 R9 R11 R9 1 20 1 04 1 21 1 33 1 09 1 27 0 99 1.06 0 69 14 R11 RIO R9 RIO R11 R10 R11 R11 R9 1 34 1 23 1 08 1 18 1.28 1 11 1 13 0 96 0 56 12 R9 R11 RIO R11 R9 R11 R9 R10 R8 1 11 1 31 1 16 1.26 1 00 1 13 0.84 0'.71 0.41 IC R11 RIO R11 RIO R11 R11 RIO R8 1.27 1 16 1 22 1.06 1.06 0 96 0.71 0.44 8 R9 R11 R9 R11 R9 R9 R8 0 97 1.12 0 89 0.94 0.69 0.56 0 41 6 R11 RIO R10 RB R8 - P8DPB289, RELOAD 8 0.94 0.82 0 71 0.52 R9 - P8DPB289, RELOAD 9 4 I R9 0 53 R9 0.48 R8 0.37 BUNDLE 10 RIO - P8DPB289, RELOAD 10 R11 - P8DPB289, RELOAD 11 RELATIVE P0hER 2 l l l l l l 23 26 27 29 31 33 35 37 39 41 43 I

I Figure 5.1.1 VY CYCLE 12 HALING DEPLETION, EOFPL BUNDLE AVERAGE RELATIVE POWERS I

M M M M M M M C M M M M M M M VY CYCLE 12 CORE AVERAGE AXI AL POWER DISTRIBUTION TAKEN FROM THE HALING CALCULATION TO EOFPL 15 14 13 12

~'

11- /  %

~

5 / s 3: to g c.9 O

0.8 k.

4 0.7 m \

> 0.6

/

e1- 0.5

\

0.4 0.3 0.2 -

0.1 0.0 }

O 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18'19 20 21 22 23 24 25 PLANT AND SIMULATE AXI AL NODE (24= TOP)

FIGURE 5.1.2 VY CYCLE 12 HALING DEPLETION, EOPPL CORE AVERAGE AXIAL POWER DISTRIHUTION

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VERMONT YANKEE I CYCLE 12 RODDED DEPLETION EOFPL BUNDLE AVERAGE RELATIVE POWERS I

R9 R9 R10 R9 RO R11 R9 R11 R9 Rt1 R9

.958 .998 124 1063 1206 1344 22 1.097 1263 .957 .925 .518 R9 R10 R11 R10 R8 Rio Rt1 R10 R11 R10 R9 1002 1159 1327 1207 2.039 1222 1309 iM3 20 1106 .804 .465 I R10 1156 R11 1325 R9 it26 R11 1351 R10 1 211 R9 1070 R10 1853 R11 1208 R9

.880 R10

.703 R8

.359 8 I R9 1055 R10 1201 R11 1347 R9 1121 Rt1 U30 R10 M76 R11 1249 R10 1045 R11

.924 R8

.510 16 I R10 1207 R8 1037 R10 1207 R11 1330 R9 1083 R11 1269 R9

.985 Rt1 1055 R9

.679 "

I R11 1349 Rio 1224 R9 1074 Rio 1176 Rt1 1270 Rio 110 7 Rit '

it20 Rt1

.952 R9

.557 E I -

R9 1710 R11 1314 Rio 1157 R11 1252 R9

.987 Rif 1121 R9

.828 R10

.705 R8

.401 E I R11 1271 RIO tMS R11 1212 R10 1048 R11 1056 Rt1

.954 R10 706 R8

.431 8 I R9

.967 R11 U14 R9

.885 R11

.927.

R9

.681 R9

.558 R8

.400 6 I .931 . 8 11 .706 .5t2 R8 - P8DPB289, RELOAD 8 R9 - P8DPB289, RELOAD 9 R10 - P8DPB289, RELOAD 10 I

R9 R9 R8 BUNDLE ID R11 - P8DPB289, RELOAD 11

.525 .470 .363 RELATIVE POWERS I

, 2, 25 2, 29 3, , , .T , 6, .3 FIGURE 5.1.3 VY CYCLE 12 ROODED DEPLETION - ARO AT EOFPL, BUNDLE AVERAGE REL ATIVE POWERS M M M M M M M M M M M M M M M M M M M VY CYCLE 12 CORE RVERAGE RXIRL POWER DISTRIBUTION RODDED DEPLETION -- RLL RODS OUT RT E0FPL 1.5

1. 4 1.3

/ N 1.2 x 1* 1 -%

b 1.0 s

5 N a- 0.9 / \

i d ** / N .

=

8 5

g 0.7

( .

N

[ 0.6 5 0.5

(

0.4 0.3 0.2 0.1 0.0 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 PLANT AND SIMULRTE RXIRL NODE (24-TOP)

FIGURE 5.1.4 VY CYCLE 12 RODDED DEPI,ETTON-ARO AT EOFPI,. CORE AVERAGE AXIAI, POWER DI ST Rillt!TI ON

I I VERMONT YANKEE CYCLE 12 HALING DEPLETION EOFPL BUNDLE AVERAGE EXPOSURES I

R9 R9 R10 R9 RIO R11 R9 R11 R9 R11 R9 25 02 25 87 17.75 25 97 16 40 9 32 24 92 8.80 22 11 6.51 19 98 ~22 R9 R10 R11 R10 R8 RIO R11 R10 R11 RIO R9 25.79 18 01 9.10 17 56 27.58 15 51 9 12 16.56 7 76 15 68 21 30 - 20 RIO R11 R9 R11 RIO R9 RIO R11 R9 RIO RB 17 76 9 10 25 52 9.33 16.95 26.14 18 45 8 46 21.85 14 97 25.98 ~18 R9 RIO R11 R9 R11 RIO R11 R10 R11 R8 26.45 17 65 9 34 25.61 9 30 18.02 8.77 16 74 6.52 25 27 16 R10 R8 RIO R11 R9 R11 R9 R11 R9 16.41 27.62 17.01 9.31 26.14 8.92 25.65 7.43 21 01 14 R11 R10 R9 RIO R11 R10 R11 R11 R9 9.37 15.56 26 05 18 13 8 93 17.88 7.89 6.73 20 22 12 R9 R11 R10 R11 R9 R11 R9 RIO R8 24.68 9 18 18.47 8 80 25.65 .7.89 22.00 l4 80 23.96 10 R11 RIO R11 RIO R11 R11 RIO R8 8 87 16.61 8 51 16 80 7.45 6.74 14 76 26 75 8 R9 R11 R9 R11 R9 R9 R8 22 02 7.82 21.95 6.55 21 00 20 25 24 09 6 R11 RIO RIO R8 R8 - P8DPB289, RELOAD 8 6 57 15 69 14.97 25 36 R9 - P8DPB289, RELOAD 9 4 I R9 19.96 R9 21.35 R8 25.97 BUNDLE 10 RTO - P8DPB289, RELOAD 10 R11 - P8DPB289, RELOAD 11 EXPOSURE [0WO/ST) 2 l l l l l l 23 25 27 29 31 33 35 37 39 41 43 I Figure 5.2.1 VY CYCLE 12 HALING DEPLETION, EOFPL BUNDLE AVERAGE EXPOSURES g -n-I

I I VERMONT YANKEE CYCLE 12 RODOED DEPLETlON I EOFPL BUNDLE NERAGE EXPOSURES I

R9 R9 RC R9 Rio R11 R9 R11 R9 R11 R9 2421 25.76 UA5 25.98 16.4 8 0.04 25.04 8.55 22.21 6.18 20.a3 22 I R9 25A5 R10 U.75 R11 8.38 R10 U.55 R8 27.75 R10 16.7 9 R11 835 R10 16.7 7 R11 734 RC ES3 R9 2142 20 I R10 U.50 R11 8.36 R9 25.28 R11 8.86 RIO U.12 R9 26.47 RC 18.7 2 R11 8.31 R9 2t99 RIO 15.19 R8 26.0 8 I R9 26.41 Rio U.72 R11 8.97 R9 25.72 Rt1 9.14 Rio 18.37 R11 8.66 RIO 16.91 R11 6.26 RS 25.38 #

I R10 16.4 3 R8 2735 R10 U.27 R11 9.16 R9 26.33 R11 8.79 R9 25.77 R11 7.08 RD 2t02 "

I R11 8.97 Rio 1521 R9 26.37 Rio 18.4 5 R11 8.78 RC E06 R11 737 R11 6.29 R9 2 0.19 U

~

I M 24.69 M

835 M

18.7 0 M

8.65 M

25.74 M

7.56 M

22.06 M

1435 M

24.01 D I R11 8.53 Rio 16.7 5 R11 8J1 Rio 16.9 3 R11 7.07 R11 6.29 RO 14.81 R8 2633 8 I

R9 R11 Ris R11 R9 R9 R8 22.04 734 22.04 8.25 20.98 20.22 24.0 6 M RM RC R8 R8 - P8DPB289, RELOAD 8 I 6.0 15.7 8 15.14 25.43 R9 - P8DPB289, RELOAD 9 4 RIO - P8DPB289, RELOAD 10 R9 R9 R8 B N ID R11 - P8DPB289, RELOAD 11 1936 2144 26.09 EXPOSt7tE (GWD/ST) 2 I -

2, 2. . 29 ., .3 - ,, M ., .3 FIGURE 5.2.2 l VY CYCLE 12 ROODED DEPLETION, EOFPL BUNDLE NERAGE EXPOSURES l

I l . --- .-- - --

M M M M M M M M M M M M M M M M VY CYCLE 12 COLD SHUTDOWN PERCENT DELTR K VS. CYCLE EXPOSURE 2.0 1.9 1.8

1. 7 =

1. 6 -

1. 5 x 24 x

,N /

e N

' G-s 1.1

[

a 1.0 /

Z 0.9 r

80.8 s

D 0.7 x

M 0. 6 0.5 0.4 l'EUHNIURL SI'EUII'ICHTION LIf111

~~---------- -----~~~~-- -----------"------------~~~-------"-~~~------- ---~~~~~---

0.3 0.2 0.1 O.0 .

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 CYCLE EXPOSURE (GWD/ST)

FIGURE 5.3.1 VY CYCI.E 12 COLD SliUTDOWN DEI.TA K IN PERCENT VERSUS CYCLE EXPOSfJRE

I 6.0 THERMAL-HYDRAULIC DESIGN The thermal-hydraulic evaluation of the Reload Cycle was performed I using the methods described in the following section.

6.1 Steady-State Thermal Hydraulics Core steady-state thermal-hydraulic analyses were performed using the FIBWR [10,11] computer code. The FIBWR 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. FIBWR calculates the core pressure drop and total bypass flow for a given total core flow. The I power distribution, inlet enthalpy, and geometry are presumed known and are supplied to FIBWR. The power distribution is derived by the 3-D neutronic simulator SIMULATE [4]. Core pressure drop and total leakage flow predicted by the FIBWR code were used in setting the initial conditions for the system's transient analysis model.

I 6.2 Reactor Limits Determination I The objective for normal operation and anticipated transient events is to maintain nucleate boiling. Avoiding a" transition to film boiling protects I the fuel cladding integrity. ' Based on Reference 12, 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 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 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 bundle power. Thermal margin is stated in terms of the minimum value of the critical power ratio, MCPR, which corresponds to the most limiting fuel I assembly in the core. Both the transient (safety) and normal operating thermal limits in terms of MCPR are derived based on the GEIL correlation as described in Reference 12.

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I Vermont Yankee Technical Specifications [8] limit the operation of the Reload Cycle fuel to a Maximum Linear Heat Generation Rate (MLHGR) of 13.4 KW/ft. The basis for a MLHGR of 13.4 KW/ft can be found in Reference 3.

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I 7.0 ACCIDENT ANALYSIS I 7.1 Core Wide Transient Analysis Core wide transient simulations are performed to assess the impact of certain transients on the heat transfer characteristics of the fuel. The figure of merit used is the Critical Power Ratio (CPR). It is the purpose of the analysis to determine the minimum critical power ratio such that the safety limit is not violated for the transients considered.

I 7.1.1 Methodology The analysis requires two types of simulations. A system level I simulation is performed to determine the 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 fuel, referred to as " hot channel calculations". The hot channel simulations provide the bundle transient ACPR (the initial bundle CPR minus the minimum CPR experienced during the transient).

I The system level simulations are performed with the model documented in Reference 13. -

The hot channel calculations are performed with the RETRAN [14] and TCPYA01 [15] computer codes. The GEXL correlation [12] is used in TCPYA01 to evaluate critical power ratio. The calculational procedure is outlined below.

I The hot channel transient ACPR calculations employ a series of " inner" I and " outer" iterations, as illustrated by the flow chart in Figure 7.1.1.

outer loop iterates on the hot channel initial power level. This is necessary The because the ACPR for a given transient varies with Initial Critical Power Ratio (ICPR). However, only the ACPR corresponding to a transient MCPR equal to the safety limit (i.e., 1.07 + ACPR = ICPR) is appropriate. The approximate constancy of the CPR/ICPR ratio is useful in these iterations.

Each outer iteration requires a RETRAN hot channel run to calculate the transient enthalples, flows, pressure and saturation properties at each time step. These are required for input to the TCPYA01 code. TCYPA01 is then used I I  !

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.

I The inner loop iterates on the hot channel inlet flow. These iterations are necessary, because the RETRAN hot channel model calculates the I entrance loss coefficient when given the initial power level, flow, and pressure drop as input. The pressure drop is assumed equal to the core average pressure drop, and the flow is varied for a given power level until the calculated entrance loss coefficient is correct. FIBWR [10, 11] is utilized to estimate the correct inlet flow for a particular power ' level and pressure drop.

I 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 conditions are 104.5% core themal power and 100% core flow. The core axial power distribution for each of the exposure 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 t'he quality in the liquid region outside the separators is as close to zero as possible. For fast I

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.

Assumptions specific to a particular transient are discussed in the section describing the transient. In general, the following assumptions are made for all transients:

I 1. Scram setpoints are at Technical Specification [8] limits.

I 2. Protective system logic delays are at equipment specification limits.

I

I 3. Safoty/relicf valv3 cnd stfety v31va crpscitica tra b:sid en Technical Specification rated values.

4. Safety / relief valve and safety valve setpoints are modeled as being I at the Technical Specification upper limit. Valve responses are based on slowest specified response values.

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5. Control rod drive scram speed is based on the Technical Specification limits. The analysis addresses a dual set of scram speeds as given in the Technical Specifications. These are referred to as the " Measured" and the (slower) "67B" scram times.

7.1.3 Reactivity Functions I The methods used to generate the fuel temperature, moderator density, and scram reactivity functions are described in detail in Reference 16. The method is outlined below.

I 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, EOFPL-2000 MWD /ST, and BOC exposure statepoints.

These statepoints are characterized by exp"osure and void history I distributions, control rod patterns, and core thermal 'nydraulic conditions.

The latter are consistent with the assumed system transient conditions provided in Table 7.1.1.

The BOC base state is established by shuffling from the previously defined Current Cycle endpoint into the Reload Cycle loading pattern. A criticality search provides an estimate of the BOC critical rod pattern. The EOFPL and intermediate core exposure and void history distributions are calculated with a Haling depletion as described in Section 5.2. The EOFPL I state is unrodded. As such, it is defined sufficiently. However, EOFPL-1000 MWD /ST and EOFPL-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 i

I parameters are within operating limits, then this configuration becomes the i

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.

I 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 I

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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 terperature (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 generated by quasi-statically varying the inlet subcooling only. The moderator enthalpy source distribution is in equilibrium with the calculated nuclear power. The moderator density reactivity functions for the pressurization transients are generated by quasi-statically varying the core pressure. A series of calculations are performed for various inlet moderator temperatures. The moderator enthalpy source distribution is that of the base state case. .

~

I In order to qualitatively compare the core reactivity characteristics between different base configurations, core average reactivity coefficients at I selected conditions are provided in Table 7.1.2.

parameters for RETRAN are also provided.

Calculated point kinetics The reactivities versus scram insertion are calculated at constant, pre-transient moderator conditions. These are fitted to yield highly detailed scram reactivity curves. The curves are combined with the appropriate rod I 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 for the " Measured" scram time analysis. Figures 7.1.5 through 7.1.7 display similar curves for the "67B" scram time analysis.

I

I 7.1.4 Tr nsi"nts Analyzid Past licensing analysis has shown that the core wide transients which I result in the minimum core thermal margins are:

1. Generator load rejection with complete failure of the turbine bypass system.

2. Turbine trip with complete failure of the turbine bypass system.

3. Loss of feedwater heating.

The "feedwater controller failure" (maximum demand) transient is not a I limiting transient for Vermont Yankee, because.of the plant's 110% steam flow 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 core wide transients analyzed are provided in the following section.

7.2 Core Wide Transient Analysis Results The transients selected for conside~ ration were analyzed at exposure I points of EOFPL, EOFPL-1000 MWD /ST, and EOFPL-2000 M'.JD/ST; the loss of feedwater heating transient was also evaluated at BOC conditions. A summary of the results of the analyses is provided in Table 7.2.1.

7.2.1 Turbine Trip Without Bypass Transient (TTWOBP)

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 normally open to relieve pressure, remain closed. A reactor protection system signal is generated by the turbine stop valve closure switches. Control rod 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 I valva ch;st le lump:d into tha centrol volume upstream of ths turbina otcp valves. Predictions of the salient system parameters at the three exposure points are show in Figures 7.2.1 through 7.2.3 for the " Measured" scram time analysis.

7.2.2 Generator Load Rejection 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 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 c 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.

7.2.3 Loss of Feedwater Heating Transient (LOFWH)

A feedwater heater can be lost in such a way that the steam extraction line to the heater is shut off or the feedwater flow bypasses one of the heaters. In either case, the reactor will receive cooler feedwater, which will produce an increase in the core inlet subcooling, resulting in a reactor 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 assumption for the maximum expected single heater or group of heaters that can be tripped or bypassed by a single event.

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Vermont h nkte h = a ceram setpoint of 120% of ratcd p:wsr es part of the Reactor Protection System (RPS) on high neutron flux. In this analysis, no credit was taken for scram on high neutron flux, thereby allowing the I reactor power to reach its peak without scram. This approach was selected to provide a bounding and conservative analysis.

The transient response of the system was evaluated at several exposures during the cycle. The transient evaluation at EOFPL-1000 MWD /ST was found to be the limiting case between BOC to EOFPL. 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 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 is the same as that used for the core wide transient analysis (Section 7.1.1).

The initial conditions and modeling assumptions discussed in Section 7.1.2 are applicable to this simulation.

The transient is initiated by a siniultaneous closure of all four MSIV's. 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 of 10% above vessel design pressure for upset conditions, or 1375 psig.

7.4 Local Rod Withdrawal Error Transient Results l l

The rod withdrawal error is a local core transient caused by an operator erroneously withdrawing a control rod in the continuous withdrawal I mods. If th3 cora is cptrsting at its cptrating linits for MCPR cnd LHGR st 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 System to stop the 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 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 experierree the spatial power shape transient as well as the overall core power increase.

The core conditions and control rod pattern for the bounding case are specified using the following set of concurrent worst case assumptions:

1. The rod sheuld have high reactivity worth. This is provided for by analysis cf tne core at peak reactivity exposure for each test pattern with xenon free conditions superimposed. The xenon free conditions and the additional control rod inventory needed to maintain criticality exaggerates the worth of control rods

'substantially when compared to normal operation with normal xenon levels. A fully inserted high worth rod is selected as the error rod.  ;

2. The core is initially at 104.5% power and 100% flow.

I I 3. The core power distributicn is cdjustsd with th~s Evsiltble ccntrol l

rods to place the locations within the four by four array of bundles around the error rod as nearly on the operating limits as practical. .

l 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 channels to remain operable if at least half of the 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 from the A and C levels; the I 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 LPRM strings. The RBM channel responses are evaluated separately at these eleven input failure conditions. Then, for each channel taken separately, the lowest response as a function of error rod pcsition is chosen for comparison to the RBM setpoint.

The event is analyzed separate 1Iy in each of the four quadrants of I 2.

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.

I The analysis is performed using the three dimensional steady state SIMULATE core model [4]. Necessary properties cf that model for use in this analysis are:

1. Accurate bundle power calculation as shown by the PDQ and gamma scan comparisons.

l Il f 2. Accurste LPRM signti calculiticn es chown by ths d;tcilsd TIP trcco I comparisons.

3. Accurate control rod worths and core power coefficient as shown by the consistent core eigenvalues.

l-Two separate cases are presented from numerous explicit SIMULATE l 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 the abnormal xenon condition at the most reactive point in the cycle for the given rod pattern configuration. The initial conditions for Case 2 approximate the 104.5% power conditions with an expected control rod pattern and equilibrium xenon. The transient results, the ACPR and 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 are not initially near the MCPR operating limit. 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 MCPR margin. ,

The Case 1 RBM channel responses are shown in Figures 7.4.3 and 7.4.4.

They also note 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 Rod Block Monitor setpoint is taken from Figure 7.4.1. The same figure demonstrates margin to the 1% plastic strain limit. The MLHGR values include the 2.2% power spiking penalty.

7.5 Misloaded Bundle Error Analysis Results l

7.5.1 Rotated Bundle Error l

l The primary result of an assembly rotation is a large increase in local 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 i

- }]

8 racctivity, d:pinding en th] cyp rura Cnd void fr:cti n stctas. Th3 R-fcetor increase results in a CPR reduction, while the local pin peaking factor increase results in a higher pin linear heat generation rate. The objective of the analysis is to insure that in the worst possible rotation, the safety limit linear heat generation rate and CPR are not violated with the most limiting monitored bundles on their operating limits.

To analyze the CPR 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 [12). 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. These MCPR values are, in addition, modified slightly to account for the change in reactivity resulting from the rotation. For each sequence, the MCPR for the properly oriented assemblies 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.

To determine the Maximum Linear Heat Generation Rate (MLHGR) resulting from a rotation, the ratios of the maximum rotated bundl,e local peaking factor to the maximum properly oriented bundle lo' cal peaking are determined for the l' expected range of exposure and void conditions. The maximum of this ratio is applied to the operating limit LHGR of 13.4 kw/ft. This maximum rotated bundle LHGR 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.

The results of the rotated bundle analysis are given in Table 7.5.1.

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 used for the misloaded bundle are the standard values for the fuel type.

l -,-

I

I The analysis consists of an iterative procedure which examines the effects of explicit misloadings and/or successively eliminates potential misloading locations from any MCPR safety limit violations. The first step is to use SIMULATE to determine the largest possible ACPR which could result, at any location, as the result of misloading a high reactivity assembly into the location. This maximum ACPR is then subtracted from all the other bundle I CPR's in the core. This is done at the various cycle exposures. Even with this maximum ACPR applied, some locations will never exceed the MCPR safety limit of 1.07. These locations are eliminated from further investigation.

The next iteration consists of applying the same procedure to the locations which appeared to violate the safety limit when the maximum ACPR from the first iteration was applied. Since these locations are of higher reactivity than those eliminated in the first iteration, they will result in a smaller ACPR when misloaded. Using this smaller ACPR, some of the remaining I locations will be eliminated from potential CPR safety limit violations. This procedure is continued until all locations are shown to be above the MCPR safety listit due to a misloading, or until a limiting location is identified.

Using the above procedure, it has been demonstrated that for the Reload Cycle all possible mislocations result in calculated MCPRs well above the

~

1.07 safety limit, assuming the initial operating CPR limit set by the rotated bundle analysis. This CPR limit is shown in Table 7.5.1. This makes the mislocated bundle analysis less limiting than the rotated bundle analysis.

I 7.6 Control Rod Drop Accident Results The control rod sequences are a series of rod withdrawal and banked 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 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.

I

I Some cut-of-c qu*.ncs centrol r:dm could eccrum pst2ntially high I 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.

The sequence entered into 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 gram [17, 18]. Below 20% power, however, the. sequences must be examined for incremental rod worth. This is done throughout the cycle using the full core, xenon free SIMULATE model.

Both the A and B sequences were examined. It was found that the highest worth occurred in the first red pull of the second group. Any of the first four rod arrays shown in Figures 7.6.1 and 7.6.2 may be designated as the first group pulled. But, then 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 followed at the plant, the worths will always be I 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 [19] apply which severely reduce the rod incremental worths. This makes the xenon free, hot standby worths much less than the cold, xenon free worths as demonstrated in Reference 9.

7.7 Stability Analysis Results I An analysis of reactor stability has not been specifically performed for this Reload Cycle. Vermont Yankee intends to implement the stability monitoring as described in Reference 21.  !

i l

l 1 I I 1

I TABLE 7.1.1 VY CYCLE 12

SUMMARY

OF SYSTEM TRANSIENT MODEL INITIAL CONDITIONS FOR CORE WIDE TRANSIENT ANALYSES Core Thermal Power (MWth) 1664.0 Turbine Steam Flow (% NBR) 105 Total Core Flow (10 61bm/hr) 48.0 Core Bypass Flow (10 6 1bm/hr) 5.3 Core Inlet Enthalpy (BTU /lbm) 520.9 Steam Dome Pressure (psia) 1034.7 Turbine Inlet Pressure (psia) 986.0 6

Total Recirculation Flow (10 1bm/hr) 23.4 Core Plate Differential Pressure (psi) 18.5 Narrow Range Water Level (in.) 35 Average Fuel Gap Conductance (See Section 4.2)

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M M M M M M M M M M M M M M M M M TABLE 7.1.2 VY CYCLE 12 TRANSIENT ANALYSIS REACTIVITY COEFFICIENTS AT SELECTED CONDITIONS Cycle Exposure Point (MWD /ST)

Calculated Parameter EOFPL EOFPL-1000 EOFPL-2000 BOC Axial Shape Index(l) -0.0874 -0.1726 -0.1003 -0.1402 Moderator Density Coefficient 20.87 23.95 25.13 18.26 (Subcooling', (/au(2)

PresFure = 1050 psia Subcoo18ag = 30 BTU /lbm Moderator Density Coefficient 23.11 26.05 27.92 (3)

(Pressurization), g/Au ~

Pressure = 1050 psia Inlet Enthalpy = 520 BTU /lbm Fuel Temperature Coefficient -0.278 -0.284 -0.292 -0.263 at 11300F, //0F I

C Effective Delayed 0,005437 0.005517 0.005579 0.006013 Neutron Fraction Prompt Neutron Generation 42.52 42.35 41.16 39.73 Time in Microseconds P -P

• T B 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

E TABLE 7.2.1 1

VY CYCLE 12 CORE WIDE TRANSIENT ANALYSIS RESULTS I Peak Prompt Power (Fraction of Peak Avg.

Heat Flux (Fraction of ACPR j

Exposure P8X8R I

Transient Initial Value) Initial Value)

Turbine Trip EOFPL 2.990 1.213 .21 I Without Bypass,

" Measured" Scram Time EOFPL-1000 2.011 1.091 .08 I EOFPL-2000 1.000 1.000 .00 Turbine Trip EOFPL 3.412 1.258 .27 I Without Bypass, "67B" Scram Time EOFPL-1000 2.518 1.151 .14 EOFPL-2000 1.191 1.000 .00 2.942 .20 I Generator Load Rejection Without Bypass, EOFPL EOFPL-1000 1.899 1.202 1.071 .06

" Measured" Scram Time EOFPL-2000 1.000 1.000 .00

~

I Generator Load Rej ection Without Bypass, EOFPL EOFPL-1000 3.471 2.486 1.253 1.140

.27

.14 "67B" I Scram Time EOFPL-2000 1.082 1.000 .00 I Loss of 100 0 F Feedwater Heating EOFPL EOFPL-1000 1.205 1.218 1.197 1.210

.16

.17 1

]

EOFPL-2000 1.215 1.207 .17 BOC 1.201 1.193 .15 I

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l I TABLE 7.3.1 VY CYCLE 12 OVERPRESSURIZATION ANALYSIS RESULTS Maximum Pressure at Reactor Conditions Vessel Bottom (psix)

" Measured" Scram Time 1272 "67B" Scram Time 1294 l

l I TABLE 7.5.1 VY CYCLE 12 ROTATED BUNDLE ANALYSIS RESULTS Resulting Initial MCPR Resulting MCPR LHGR (kw/ft) 1.24 1.07 17.66 I

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TABLE 7.6.1 CONTROL ROD DROP ANALYSIS - ROD ARRAY PULL ORDER I The order in which rod arrays are pulled is specific once the choice of first group is made.

First Group Second Group Successive Group Pulled is: Pulled Must Be: Is Banked Out Array 1 Array 2 Array 3 or 4 I Array 2 Array 1 Array 4 Array 3 or 4 Array 1 or 2 Array 3 Array 4 Array 3 Array 1 or 2 I

TABLE 7.6.2 VY CYCLE 12 CONTROL ROD DROP ANALYSIS RESULTS Maximum Incremental Rod Worth .87% AK Calculated Cold, Xenon Free Bounding Analysis Worth for Enthalpy 1.30% AK Less than 280 Calories per Gram ,

(References 17, 18 and 19) .

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is Loss Coefficient Corr No Revise Flow Yes If RETRAN/TCPYA01 Hot Channel Run  ;

I V Has ACPR No Converged? New ICPR I Yes STOP FIGURE 7.1.1 FLOV CHART FOR THE CALCULATION OF ACPR USING Tile RETRAN/TCPYA01 CODES 1

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" MEASURED" SCRAM TIME

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" MEASURED" SCRAM TIME I I

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FIGURE 7.1.5 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL12, "67B" SCRAM TIME I I

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I FIGURE 7.1.7 INSERTED ROD WORTH AND ROD POSITION VERSUS TIME FROM INITIAL ROD MOVEMENT AT EOFPL12-2000 MWD /ST, "67B" SCRAM TIME

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-2.0 O.0 .5 1.0 1.5 2.0 2.5 3.0 T!hECEC)

TTWOBP EOFPL-2 MST i

i FIGURE 7.2.3-3

, TURBINE TRIP WITHOUT BYPASS, EOPPL12-2000 MWD /ST TRANSIENT RESPONSE VERSUS TIME, " MEASURED" SCRAM TIME i

I

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o.

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O.0 1.0 2.0 3.0 4.0 5.0 6.0 0.0 t.0 2.0 3.0 4.0 5.0 0.0 TIE (EC) T!hECEC)

(1RWOBP EOFPL MST Q RWOBP EOFPL h67 FIGURE 7.2.4-2 GENERATOR LOAD REJECTION WITilOUT BYPASS, EOFPL12 TRANSIENT RESPONSE VERSUS TIME, " MEASURED" SCRAM TIME

' M M M M M M m m g g g g g 2.0 ,

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G_RWOOP EOFPL M3T I

FIGURE 7.2.4-3 GENERATOR LOAD REJECTION WITl!OUT BYPASS, EOFPL12 TRANSIENT RESPONSE VERSUS TIME, " MEASURED" SCRAM TIME

ll ,

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Q_RWOBP EOFR.-1 MST FIGURE 7.2.5-3 1

GENERATOR LOAD REJECTION WITIIOUT tiYPASS, EOPPL12-1000 MWD /ST TRANSIENT RESPONSE VERSUS TIME, " MEASURED" SCRAf1 TIME

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1 i

FIGURE 7.2.6-3 GENERATOR LOAD REJECTION WITilOUT BYPASS, EOPP',12-2000 MWD /ST IMANSILNI HLSPUNbt VERSUS IIME, " MEASURED" SCRAM TIME l

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'I REACTOR INITIAL CONDITIONS AND TRANSIENT SUMMRRY FOR THE VY CYCLE 12 ROD WITHDRAWRL ERROR LICENSING CASE g INITIAL CONDITIONS 43-39 10 14 10 35 14 18 18 14 31 10 to I 27- 18 0 0 18 23- 14 14 19- 18 0 0 18 15 -- 10 10 I 11 -

07 14 10 18 14 18 10 14 03 --

g 02 06 10 14 18 52 56 30 34 38 42 CORE THERMAL POWER - 1864 MWT CORE AVE. PRESSURE - 1041.9 PSIP.

COPE PLOW - 48 MLB/HR INITIRL MCPR - 1.2621 l CYCLE EXPOSURE ZERO XENON

- 4350 MWD /ST INITIAL LHGR - 13.4 KW/PT I RWECdNTROLR00ATC00RD'S26-27 TRANSIENT

SUMMARY

I RBN ROD SETPOINT POSITION ' DELTA CPR MLHGR (KW/PT) 104 10 .10 i

15.1 105 10 .10 15.1 106 '2 .13 18.2 107 14 .16 16.8 108 18 .20 17.4 I "'""""'-

REACTOR INITIAL CONDITIONS AND TRANSIENT

SUMMARY

FOR THE VY CYCLE 12 ROD WITHDRAWAL ERROR CASE 1 I

I REACTOR INITIAL CONDITIONS AND TRANSIENT

SUMMARY

g FOR THE VY CYCLE 12 ROD WITHDRAWAL ERROR NORMAL CASE I INITI6L CONDITIONS I

39 34 35 31 34 8 34 27-23- 34 10 0 10 34 19-IS -- 34 8 34 11 --

07 34 02 06 10 14 18 22 26 30 34 38 42 CORE THERM 8L POWER - 1664 MWT CORE AVE. PRESSURE - 1041.9 PSIA CORE FLOW - 48 MLB/HR INITIAL MCPR - 1.4141 CYCLE EXPOSURE - 4350 MWD /ST INITI6L LHGR - 12.2 KW/PT EQUILIBRIUM XENON -

RWE CONTROL ROD BT C00RD'S 22-31 g TRANSIENT

SUMMARY

RBM ROD SETPOINT POSITION DELTA CPR NLHGR (KW/FT)

I 104 18 .06 12.3 105 20 .08 12.3 106 22 .09 12.3 ll 107 108 26 34

.11 12.5

.15 13.3

'I FIGURE 7.4.2 REACTOR INITIAL CONDITIONS AND TRANSIENT SUMMAR FOR THE VY CYCLE 12 ROD WITHDRAWAL ERROR CASE 2 I

M M M M M M M M M M M M M M M M M M M e .

9

~

RBM RESPONSE TO RWE. A+C _5

~

VERSUS RWE CONTROL R00 POSITION 5  %

U U d 2,, NOTES: C+

-;E b

2

}. ALL INTERCEPTS ARE DETERMINED BY THE B+D CHANNEL. 3 S 2.

,E O

g;- THE Box (D) SHOWS THE RESPONSE O WITH NO INSTRUMENT FAILURES. ~

I tij w

E U2 g CN a.*:

~5

• l g, us - us LAJ E

w E

1. e E.9_~

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$~ -

• 8 8 T OO 6'.00 l'O.00 15.00 2'O.00 25.00 3'O.00 3h.00 4'O.00 45.00 6030 RWE CONTROL ROD POSITION j FIGURE 7.4.3 l

VY CYCLE 12 RWE CASE 1-SETPOINT INTERCEPTS DETERMINED BY Tile A+C CHANNEL

W W M M M M M M M M M M M M M M M M M

~_

RBN RESPONSE TO RWE, B+D 3~

VERSUS RWE CONTROL ROD POSITION Z Z J- NOTES:

}. R M $ TP NT INTERCEPT l$ NARKED d- 2. ROD IS STOPPED AT NOTCH FOLLOWING ~$

= TWO INCHES OF FREE ROD NOTION.

• tAJ IAJ z 3. THE 80x (D) SHOWS THE RESPONSE 2 E go WITH NO INSTRUMENT FAILURES. to "

c" 50 H- -H 11J LAJ

W u Z m

W CN z No e

O- T Q-m- _ m

%3_ OE

- IflM -w 107 n6 . -

105

~~ t04 -d O O O, '

10 12 14, lh o l

I DD 5'.00 10.00 l'5.00 2'O.00 2'5.00 3'O.00 3b.00 4'O.00 4'5 00 6030 l RWE CONTROL ROD POSITION FIGURE 7.4.4 VY CYCLE 12 RWE CASE 1-SETPOINT INTERCEPTS DETERMINED BY THE B+D CilANNEI,

m-immum I 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 19 - 3 4 3 4 3 15 1 2 2 1 11 4 3 4 3 4 07 2 1 1 2 1

03 3 l l l l 02 06 10 14 18 22 26 30 34 38 42 RGURE 7.6.1 FIRST FOUR ROD ARRAYS PUl f ED IN THE A SEQUENCES l

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 19 - 3 4 I 15 2 1 3

2 3

1 4

2 3

11 3 4 4 3 07 2 1 2 03 3 3 I I I 02 08 10 14 18 22 26 30 34 38 42 FlGURE 7.6.2 FIRST FOUR ROD ARRAYS PULLED IN THE 8 SEQUENCES I

I

I 8.0 STARTUP PROGRAM I Following refueling and prior to vessel reassembly, fuel assembly position and orientation will be verified and videotaped by underwater television.

I The Vermont Yankee Startup Program will include process computer data checks, shutdown margin demonstration, in-sequence critical measurement, rod -

scram tests, power distribution comparisons. TIP reproducibility, and TIP symmetry chects. The content of the Startup Test Report will be similar to that sent to the Offict of Inspection and Enforcement in the past [22].

I I

I E I .

I 9

I I

I I I

l l

l 9.0 LOSS-OF-COOLANT ACCIDENT ANALYSIS l

l The results of the complete evaluation of the loss-of-coolant accident for Vermont Yankee as documented in Reference 23 provide required support for the operation of the Reload Cycle. No new fuel types have been introduced in this reload, therefore, the MAPLHGR limits as a function of average planar I exposure remain the same as in the Current Cycle [8).

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REFERENCES

1. M. A. Sironen and P. A. McGahan, Vermont Yankee Cycle 9 Summary Report, YAEC-1367, June 1983.

2. M. A. Sironen and R. C. Potter, Vermont Yankee Cycle 10 Summary Report, YAEC-1438, September 1984.

I 3. General Electric Standard Application for Reactor Fuel (CESTARII),

NEDE-24011-P-A-5, GE Company Proprietary, August 1982, as amended.

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 @ alysis of Boiling Water Reactors 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 Therinal 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 Effects (FROSSTEY) Code Qualification and Application, YAEC-1265P, June 1981.

8. 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 Cycle 9 Core Performance Analysis, YAEC-1275, August 1981.

10. A. A. F. Ansari, Methods for the Analysis of Boilink Water Reactors:

Steady-State Core Flow Distribution Code (FIBWR), YAEC-1234, December 1980.

11. 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 Repcrt EPRI NP-1923, Project 1754-1 Final Report, July 1981.

12. General Electric Company, GEXL Correlation Application to BWR 2-6 Reactors, NEDE-25422, GE Company Proprietary, June 1981.

13. 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.

14. EPRI, RETRAN - A Program for One-Dimensional Transient Thermal-Hydraulic Analysis of Complex Fluid Flow Systems, CCM-5, December 1978.

15. 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.

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16. J. M. Holzer, Methods for the Analysis of Boiling Water Reactors Transient Core Physics YAEC-1239P, Austst 1981.

17. C. J. Paone, et al., Rod Drop Accident Analysis for Large Boiling Water Reactors, NEDO-10527, March 1972.

18. 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.

19. D. Radcliffe and R. E. Bates, " Reduced Notch Worth Procedure", SIL-316, November 1979.

I 20. R. C. Stirn, et al., Rod Drop Accident Analysis for Large Boiling Water Reactor Addendum No. 2 Exposed Cores NEDO-10527, Supplement 2, January 1973.

21. Letter, dated April 24, 1985, Cecil O. Thomas, Chief of Standardization g and Special Projects Branch, to H. C. Pfefferlen, GE Manager of BWR g Licensing Programs, " Acceptance for Referencing of Licensing Topical Report NEDE-240ll, Revision 7, Amendment 8, ' Thermal Hydraulic Stability Amendment to GESTAR II.'"

22. Letter, FVY 84-132, dated November 6, 1984, W. P. Murphy to T. E. Murley, Regional Administrator, " Cycle XI Startup Test Report".

23. Loss-of-Coolant Accident Analysis for Vermont Yankee Nucicar Power Station, NEDO-21697, August 1977, as amended.

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APPENDIX A CALCULATED CYCLE DEPENDENT LIMITS The MCPR limits appropriate for the Reload Cycle are calculated by I 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.

With regard ~to MAPLHGR, no new fuel types have been introduced. The MAPLHGR limits given in Reference 8 apply to the Reload Cycle. The MCPR limits in Reference 8 are also bounding for the Reload Cycle. These are found in Reference 8 as Table 3.11-2 and are reproduced here as Table A.2. On Table A.2, as in the Technical Specifications, End of Cycle (EOC) is understood to mean End of Full Power Life (EOFPL).

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M M M M M TABLE A.1 VERMONT YANKEE NUCLEAR POWER STATION LIMITING CYCLE 12 MCPR RESULTS Value of "N" in RBM Average Control Rod Cycle MCTR for Equation (1) Scram Time Exposure Range P8X8R Fuel 42% BOC to EOFPL-2 CWD/T 1.27

" MEASURED" EOFPL-2 GWD/T to EOFPL-1 CWD/T 1.27 EOFPL-1 CWD/T to EOFPL 1.28 BOC to EOFPL-2 GWD/T 1.27 "67B" EOFPL-2 GWD/T to EOFPL-1 GWD/T 1.27 EOFPL-1 GWD/T to EOFPL 1.34 41% BOC to EOFPL-2 GWD/T 1.24

" MEASURED" EOFPL-2 GWD/T to EOFPL-1 GWD/T 1.24 EOFPL-1 GWD/T to EOFPL 1.28 BOC to EOFPL-2 GWD/T 1.24 "67B" EOFPL-2 GWD/T to EOFPL-1 GWD/T 1.24 EOFPL-1 CWD/T to EOFPL 1.34 4 40% BOC to EOFPL-2 CWD/T 1.24

" MEASURED" EOFPL-2 CWD/T to EOFPL-1 GWD/ 1.24 EOFPL-1 CWb!T to EOFPL 1.20 BOC to EOFPL-2 GWD/T 1.24 "67B" EOFPL-2 GWD/T to EOFPL-1 GWD/T 1.24 EOFPL-1 CWD/T to EOFPL 1.34 NOTES:

(1) The Rod Block Monitor (RBM) trip setpoints are determined by the equation shown in Table 3.2.5 of the Technical Specifications [ Reference 8].

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M M TABLE A.2 VERMONT YANKEE NUCLEAR POWER STATION TECHNICAL SPECIFICATION MCPR OPERATING LIMITS MCPR Operating Limit for Value of "N" in RBM Average Control Rod Cycle Fuel Type (2)

Equation (1) Scram Time Exposure Range 8X8 8X8R P8X8R 42% Equal or better BOC to EOC-2 GWD/T 1.29 1.29 1.29 than L.C.O. EOC-2 GWD/T to EOC-1 GWD/T 1.29 1.29 1.29 3.3 C.1.1 EOC-1 GWD/T to EOC 1.30 1.30 1.30 Equal or better BOC to EOC-2 GWD/T 1.29 1.29 1.29 than L.C.O. EOC-2 GWD/T to EOC-1 GWD/T 1.33 1.31 1.31 3.3 C.1.2 EOcel GWD/T to EOC 1.36 1.35 1.35 41% Equal or better BOC to EOC-2 GWD/T 1.25 1.25 1.25 than L.C.O. EOC-2 GWD/T to EOC-1 GWD/T 1.26 1.25 1.25 3.3 C.1.1 EOC-1 GWD/T to EOC 1.30 1.30 1.30 Equal or better BOC to EOC-2 GWD/T 1.25 1.25 1.25 y than L.C.O. EOC-2 GWD/T to EOC-1 GWD/T 1.33 1.31 1.31 6 3.3 C.1.2 EOC-1 GWD/T to EOC 1.36 1.35 1.35

_,40% Equal or better ,

,BOC to EOC-2 GWD/T 1.25 1.25 1.25 than L.C.O. EOC-2 GWD/T to EOC-1 CWD/T 1.26 1.25 1.25 3.3 C.1.1 EOC-1 GWD/T to EOC 1.30 1.30 1.30 Equal or better BOC to EOC-2 GWD/T 1.25 1.25 1.25 than L.C.O. EOC-2 CWD/T to EOC-1 CWD/T 1.33 1.31 1.31 3.3. C.1.2 EOC-1 GWD/T to EOC 1.36 1.35 1.35 NOTES: ,

(1) The Rod Block Monitor (RBM) trip setpoints are determined by the equation shown in Table 3.2.5 of the Technical Specifications.

(2) The current analyses for MCPR Operating Limits do not include 7X7 fuel. On this basis, further evaluation of MCPR operating limits is required before 7X7 fuel can be used in Reactor Power Operation.