ML20134N542

Yankee Nuclear Power Station Core Xviii Performance Analysis

ML20134N542
Person / Time
Site:	Yankee Rowe
Issue date:	08/23/1985
From:	Paul Bergeron, Cacciapouti R, Stephen Schultz YANKEE ATOMIC ELECTRIC CO.
To:
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ML20134N533	List: ML20134N530 ML20134N538 ML20134N542
References
YAEC-1496, NUDOCS 8509050184
Download: ML20134N542 (101)

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4 Yankee Nuclear Power Station Core XVIII Performance Analysis August 1985 Major Contributors: J. Abdelghany K. J. Morrissey D. Beller D. A. Rice F. Carpenito R. Paulson V. Chandola K. E. St. John Y. Fujita W. J. Szymezak J. Kendall Approved By: '

P. A. Bergeron/flanager (Ddte)

Transient AnalMis Group Approved By: # 3/Ff R[/J. Cacciafouti, Manager /(Dat'e)

Reactor Physics Group Approved By: 3 85 S. V. ' Schultz', Mana}hr (Date)

Nuclear Evaluation and Support Group Approved By: 1 S A. Hus'ain, Mahager I (Date)

LOCA Group Approved By: '

B. C'. (Dat.e)

Nuclear Slifer, Manag utpar Engineering g ~ tment Yankee Atomic Electric Company Nuclear Services Division 1671 Worcester Road Framingham, Massachusetts 01701 8509050184Bj0 29 PDR ADOCK O pg P

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DISCLAIMER OF RESPONSIBILITY This document was prepared by Yankee Atomic Electric Company and is completely true and accurate to the best of our knowledge, information and belief. It is authorized for use specifically by Yankee Atomic Electric Company, and the appropriate subdivisions within the Nuclear Regulatory Commission only.

With regard to any unauthorized use whatsoever, Yankee Atomic Electric Company, and its 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 This report describes the mechanical, thermal-hydraulic, physics and j cafety analyses necessary for Yankee Reload Core XVIII.

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TABLE OF CONTENTS Section Page DISCLAIMER OF RESPONSIBILITY..................................... 11 ABSTRACT......................................................... iii LIST OF FIGURES.................................................. vi LIST OF TABLES................................................... viii ACEIKELEDG EMENTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . x

1.0 INTRODUCTION

..................................................... 1 2.0 OPERATING HISTORY OF CURRENT CYCLE............................... 2 3.0 GENERAL DESCRIPTION.............................................. 4 4.0 FUEL MECHANICAL AND THERMAL DESIGN............................... 8 4.1 Mechanical Design.......................................... 8 4.2 Thermal Design............................................. 10 4.3 Operating Experience....................................... 11 5.0 NUCLEAR DESIGN................................................... 24

5.1 Physics

Characteristics.................................... 24 5.2 Reactor Physics Analytical Computer Codes.................. 25 5.3 Changes in Analytical Methods.............................. 26 6.0 THERMAL-HYDRAULIC DESIGN......................................... 40 7.0 ACCIDENT ANALYSIS................................................ 50 1 7.1 Introduction............................................... 50 l

l 7.1.1 Initial Operating Conditions....................... 50 7.1.2 Reactor Trip Setpoints and Instrumentation Delays............................................. 51 7.1.3 Reactivity Coefficients............................ 51 l

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7.2 Control Rod Withdrawal Incident............................ 51 t 1

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i TABLE OF CONTENTS (continued) i Section Egge 7.3 Boron Dilution Incident.................................... 52 7.3.1 Introduction....................................... 52 7.3.2 Analysis and Results............................... 52 l 7.3.2.1 Boron Dilution During Modes 1 and 2...... 52 7.3.2.2 Boron Dilution During Mode 3............. 53 l

7.3.2.3 Boron Dilution During Modes 4 and 5...... 54 j 7.3.2.4 Boron Dilution During Mode 6............. 54 7.3.2.5 Failure to Borate Prior to Cooldown...... 55 7.3.3 Conclusions........................................ 56 7.4 Control Rod Drop Incident.................................. 56 7.5 Isolated Loop Startup Incident............................. 57 7.6 Loss of Load Incident...................................... 57 7.7 Loss of Feedwater Flow Incident............................ 58 7.8 Loss of Coolant Flow Incident.............................. 58 7.9 Control Rod Ejection Accident.............................. 59 7.10 Steam Line Break Accident.................................. 60 7.11 Steam Generator Tube Rupture Incident...................... 60 7.12 Other Accidents and Transients............................. 61 7.13 Transient Analysis Summary................................. 61 8.0 STARTUP PR0 GRAM.................................................. 74 9.0 LOSS-OF-COOLANT ACCIDENT......................................... 77 9.1 Introduction............................................... 77 9.2 Small Break L0CA........................................... 77 9.3 Large Break L0CA................................. ......... 79 i 9.3.1 Core Inlet Temperature Increase.................... 79 9.3.2 SIAS Satpoint Decrease............................. 80 9.3.3 Steam-Emergency Core Cooling (ECC) Water Injection Mode 1.......................... ......... 80 9.3.4 Axial Power shape sensitivity Study................ 80 9.3.5 . Core IVIII Burnup Sensitivity Study................ 82 9.4 Conclusion................................................. 83

10.0 REFERENCES

....................................................... 89 i

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LIST OF FIGURES number Ittit f.ast 2-1 Comparison of Measured and Calculated Radial Power Distribution (MOL) 3 3-1 Core XVII Loading Pattern 5 3-2 Core XVIII Loading Fattom Showing Dumup on Recycled Fuel 6 t

3-3 Fosition of Core XVII Assemblies in Core XVIII 7 4-1 Centerline Temperature vs. LHGR st BOC for Fresh Fuel 14 4-2 Volume Averaged Temperature vs. LMGR at BOC for Fresh Fuel 15 4-3 Centerline Temperature vs. LHGR at BOC for Exposed Fuel 16 4-4 Volume Average Temperature vs. LNGR at BOC for Exposed Fuel 17 4-5 Centerline Temperature vs. LHGR at EOC for Fresh Fuel 18 4-6 Volume Average Temperature vs. LHCR at EOC for Fresh Fuel 19 4-7 Centerline Temperature vs. LHCR at ROC for Exposed Fuel 20 4-8 Volume Average Temperature vs. LHGR at IOC for Exposed Fuel 21 4-9 Yankee Core XVIII Core Locations of Modified Assemblies 22 4-10 Yankee Core XVIII Lattice Locations of Solid Rods 23 5-1 Yankee Core XVIII Relative Radial Power Distribution 500 mwd /NTU All-Rods-Out 31 5-2 Yankee Core XVIII Relative Radial Power Distribution 8,000 mwd /MTU All-Rods-Out 32 5-3 Yankee Core XVIII Relative Radial Power Distribution 14,000 leid/MTU All-Rods-Out 33 5-4 Yankee Core XVIII Relative Radial Power Distribution ,

500 mwd /NTU Bank C Fully Inserted 34 5-5 Yankee Core XVIII Relative Axial Power in Assembly as Control Rod Bank C is Inserted 35 5-6 Core XVIII control Rod Bank Identification 36 i  ;

I 5-7 Rod Insertion Limit vs. Power Level 37 t

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gg.T. OF FIGURES (continued)

Number Title p.gge 5-8 Core XVIII Itultiplier for Reduend Power as a Function of Rxposure 38 5-9 Core XVIII Ilultiplier for Ienon Redistribution as a Function of Rxposure 49 6-1 Reactor Core Safety Limit - All Loops in operation 51 9-1 Yankee Core XVIII LOCA Axial Shape Study 85 Break Spectrum Axial Power Shapes 9-2 Yankee Core ZVIII LOCA Axial Shape Study 86 Rominal Vs. Xenon Axial Power Shapes Exposed Fuel HFP 14 GW4/NTU l

9-3 Ysnkee Core XVIII LOCA Axial Shape Study 87 Cosine Vs. I:non Axial Power Shapes BOC 18 Exposed Fuel 9-4 Core XVIII t.llowable Peak Rod LHGR vs. Cycle Burnup 88 l

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LIST OF Tejk(S Title Egge EhB).1E Nominal Mechanical Design Parameters 12 4-1 Summary of Nuclear Characteristics 27 5-1 Shutdown Requirements, % M 29 l 5-2 5-3 Comparison of Control Rod Worths 30 6-1 Thermal-Hydraulic Data Sheet for Yankee Core XVIII During 4-Loop Operation 42 6-2 Thermal-Hydraulic Data Sheet for Yankee Reference Analysis '

During 4-Loop Oparation 44 6-3 Summary of Hot Spot and Hot Chant.el Factors for 46 Yankee Core XVIII 6-4 Summary of Hot Spot and Hot Channel Factors for Yankee Reference Analysis 47 6-5 Wominal Hot Channel DNBR, FAH and Fg as Functions of Group C Position for Core XVIII 48 Initial Operating Conditions 62 7-1 ,

Reactor Trip Setpoints snd Instrumentation Delays 63 7-2 Modes 1 and 2 Boron Dilution 64 7-3 7-4 Mode 3 Boron Dilution 65 Modes 4 and 5 Beron Dilution 66 7-5 7-6 Mode 6 Boron Dilution - All Loops Isolsted Minimum Water 67 Level 68 7-7 Mode 6 Boron Dilution - Realistic Water Volume Control Rod Drop Incident Parameters 69 7-C HIP Rod Ejection Accident Parameters 70 7-9 NFP Rod Ejection Accident Parameters 71 7-10

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r LIST OF TABLES (continued)

Number Title Enge_

7-11 Yankee Core XVIII Safety Analyris Sunsaary of Results 72 0-1 Yankee Core XVIII Startup Test Acceptance Criteria 76 9-1 Core XVIII Burnup Sensitivity Study Results 80 f

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1.0 INTRODUCTION

This report is submitted to the Nuclear Regulatory Commission to support operation of the Yankee Nuclear Power Station through the forthcoming Core XVIII (reload cycle). This reload cycle will contain forty (40) fresh fuel assemblies fabricated by combustion Engineering (C-E) and thirty-six (36) essemblies from Core XVII fabricated by Exxon Nuclear Corporation (ENC). The introduction of the new fuel is necessary in order to maintain sufficient reactivity for continued operation at full rated power.

This report contains sections dealing with the mechanical, thermal-hydraulic, physics and safety analysis aspects of this reload.

The approach to licensing the reload core presented in this report is similar to that used in the licensing of Core II through Core XVII.

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2.0 OPERATING HISTORY OF CURRENT CYCLE The current operating cycle is Core XVII. Core XVII started producing power on June 8, 1984, and is scheduled to shutdown on October 19, 1985.

During the period of operation through approximately 10,000 mwd /MTU, the core operated normally in an essentially all-rods-out condition. For the remainder of full power life (12,500 mwd /MTU), and until 80% power is achieved in ,

coastdown, the plant will operate with control rod Group C restricted to 80 to 83 inches withdrawn. After this time, Group C will be fully withdrawn and the coastdown will continue in the all-rods-out condition for the remainder of Core XVII. Both plant measured data and Reactor Physics calculations have confirmed that the gross power distribution changes only slightly as a function of time during core life. The middle of life power distribution shown in Figure 2-1 is, therefore, representative of the entire cycle. No abnormalities which would adversely affect the power distribution have been detected thus far during Core XVII operation.

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9 FIGURE 2-1 YANKEE CORE XVil COMPARISON OF MEASURED AND CALCULATED RELATIVE RADIAL POWER DISTRIBUTION 0.619 0.789 0.603 0.611 0.634 0.804 0.820 0.631 2.4 1.9 2.1 3.3 0.761 t133 t106 t066 t132 0.753 0.760 1.12 6 t110 t076 t131 0.756

-0.1 -0.6 0.4 0.9 -0.1 0.4 0.763 1.17 0 1.082 1.104 1.177 1.067 1.173 0.754 0.759 1.151 1.073 t097 1.17 3 1.0 61 1.15 8 0.757

-0.5 -1.7 -0.8 -0.6 -0.3 -0.6 -1.3 0.4 0.631 1.14 4 1.078 1.072 1.19 2 1.222 t078 1073 1.135 0.614 0.634 1.12 4 1.065 t067 t181 1.209 1.076 1071 1.134 0.629 0.5 -1.8 -1.2 -0.5 -0.9 -1.1 0.0 -0.2 -0.1 2.4 0.818 1.099 1.187 1.229 1.111 1.116 t183 1.181 1.099 0.805 0.815 1.083 1.173 1.219 t109 1.1M 1.180 1.17 9 1.100 0.816

-0.4 -t5 -1.2 -0.8 -0.2 -0.2 -0.3 -0.2 0.1 1.4 0.808 1.12 4 1.185 t184 1.1M 1.109 1.226 1.114 1.093 0.795 0.805 1.106 1.177 1.17 8 t1M 1.109 t217 1.105 1.085 0.806

-0.4 -1.6 -0.7 -0.5 0.0 0.0 ,- 0.7 -0.8 -0.7 1.4 0.597 1.106 1.046 1.070 1.220 1.194 1.070 1.076 1.147 0.645 0.619 1.122 1.067 1.070 t211 1.186 1.066 1.0 61 1.12 3 0.646 3.7 1.4 2.0 0.0 -0.7 -0.7 -0.4 -1.4 -2.1 0.1 0.718 1.129 1.039 1.17 3 1.185 1.084 1.17 4 0.780 0.737 1.M4 1.053 1.17 2 1.T77 1.076 1.153 0.772 2.6 1.3 1.3 -0.1 -0.7 -0.7 -1.8 -1.0 0.695 1.105 1.072 1.107 1.142 0.772 0.713 t1M 1.077 1.107 1.130 0.773 2.6 0.8 0.5 0.0 -1.1 0.1 0.599 0.792 0.801 0.634 WEASUREMENT (7000 MWD /WTU) 0.616 0.805 0.813 0.642 cal.CULATION (7000 WWD/MTU) 2.8 1.6 1.5 L3 PERCENT DIFTERENCE

i 3.0 CENERAL DESCRIPTION Figure 3-1 is a schematic of the Yankee core showing the Core XVII loading pattern. In this scheme, the inner region consists of 40 recycled essemblies and the outer region consists of 36 fresh assemblies. Core XVIII will utilize 40 fresh assemblies, and 36 recycled assemblies. The fresh essemblies fabricated by combustion Engineering and the recycled assemblies fabricated by Exxon Nuclear all have an initial enrichment of 3.7 w/o U-235.

The core average exposure for beginning-of-life Core XVIII is 6148 Wd/MTU compared to 6562 mwd /MTU for Core XVII. The full power lifetime of Core XVIII is estimated to be 14,000 mwd /MTU compared to 12,500 mwd /MTU for Core XVII.

Figure 3-2 shows the Core XVIII loading pattern giving the location of both fresh and recycled fuel, as well as the recycled fuel's exposure.

Figure 3-3 shows the position that each recycled assembly occupied in Core XVII.

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FIGURE 3-1 CORE XVII LOADING PATTERN 1 1 3 4 F F F F s e 1 e a 10 F F F F 13 14 15 1s 17 18 11 11 F F F F 12 13 24 ts as 17 to to to tt F F F F 32 33 34 35 36 37 3s 19 30 31 y

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' F F F F s0 et et es se sE ss E9 F F F F s1 se s9 70 71 72 F F F F 73 74 75 1s F F F F F= FRESH ASSDiBLY l

I FIGURE 3-2 YANKEE CORE XVIII BOL ASSEMBLY AVERAGE EXPOSURE (MWD /MTU) 3 1 2 3

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, 12122 11640; 17024 16730 to to 11 22 13 24 u u 11 ts 16739 9358 9582! 12127 9554 11986 32 33 34 35 38 37 38 29 30 31 i

17222 12077 16884' 16717 9707 11506- 11263 og 43 44 es as 47 es 3s 40 41 11641 11521 9386 16793, 16770 11949 17171 52 s3 se ss ss s7 se as so 51 12146 9567 ~11949! 9592 9772 !16584 se so at sz 83 e4 ss as 16777 17161! 11283 11950 si es e s, 70 7t' 72 11489 vs 76 1s 7s

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FIGURE 3-3 POSITION OF CORE XVII ASSEMBLIES IN CORE XVIII t t 3 4 5 s 1 e a 10 11 11 13 to 15 to 11 to 11 3 59 60 57 13 24 ts 24 11 ts to to at et 9 28 58 75 76 48 33 34 35 36 31 3e Es 30 31 32 65 38 71 50 73 5 10 43 44 45 44 41 es 39 40 41 42 67 72 4 27 6 39 12 54 55 E6 51 Es 4e 50 51 52 53 29 1 2 19 49 68 et es se ss s6 se so et 20 17 18 74 si es es 70 11 72 66 CORE XVII POSITION 13 74 75 18 l

4.0 FUEL MECHANICAL AND THERMAL DESIGN 4.1 Mechanical Design Forty fresh assemblies manufactured by Combustion Engineering (C-E) will be inserted into the Yankee reactor for Core XVIIIoperation. The mechanical design parameters are described in detail in Reference 4.1. Table 4-1 lists the design parameters for the recycled (Batch IN-6) and fresh fuel ,

(Batch CE-1) assemblies.

The fuel design for the initial C-E batch is similar to the current vendor (Exxon Nuclear Corporation) design, except for the following:

4 a) The cold, beginning-of-life shoulder gap (space between the tops of the fuel rod end caps and the underside of the upper end fitting u flow plate) has been increased in the C-E fuel to 1.215 inches. To provide space for the increase in shoulder gap, the lower end fitting flow plate has a reduced thickness compared to that in the current vendor's fuel design.

b) The lowest spacer grid in the C-E design is fabricated from Inconel-625 material and is attached to the lower end fitting by a skirt fabricated from the same material. The design and material are consistent with those in all operating C-E fuel assemblies, c) The remaining spacer grids in the C-E fuel are fabricated entirely from Zircaloy-4 material. The all-Zircaloy grid design is also consistent with the design of grids in other C-E fuel. The grids in the current vendor's recent design have Zirealoy strips with Inconel springs captured within the strips.

d) The C-E spacer grids are the same height as those in other C-E fuel. Their height is somewhat less than that in the current vendor's design. The trailing edge of the lowest grid and the ,

leading edges of the remaining grids have been placed at the same elevations for the two designs in order to minimize flow perturbations caused by height differences.

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e) The upper end fitting in the C-E design includes a hold-down arrangement that differs mechanically from that in the current vendor's design. Hold-down forces are very similar. In the C-E design, r hold-down ring is engaged axially by the upper core plate and its downward travel creates the required spring compression.

In the current vendor's design the entire upper nozzle is engaged and moves relative to the remainder of the fuel assembly. The C-E design has the same basic arrangement as the hold-down features in other C-E fuel and provides improved stability at the upper end of the assembly.

f) The hold-down springs in the C-E design are fully shrouded by the hold-down ring from the effects of the cross-flow present in the upper end fitting. The current. vendor's design includes a shroud which extends part way around the circumference of each spring.

g) The insertion of solid Zircaloy rods in ten lattice locations in each of four B-type assemblies and three lattice locations for each of four A-type assemblies. The addition of special guide bars to two corner locations in the four B-type assemblies and one location j in the four A-type assemblies. The use of a fixed spacing device to provide extra rigidity for the solid Zircaloy rods. These changes will be for the assemblies in the core locations shown in Figure 4-9. Figure 4-10 shows the assembly lattice locations for the solid Zircaloy rods and additional guide bars.

Further descriptions of the recycled fuel rod mechanical design and cualyses are provided in Reference 4.2. These design analyses, including the cladding collapse time verification analysis, remain valid with respect to the aforementioned fuel design parameter modifications. Mechanical and chemical compatibility of the fuel assemblies with the inservice reactor environment is also addressed in References 4.1 and 4.2.

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4.2 Thermal Desimn The fuel thermal effects calculations were performed using the GAPEXX digital computer code (Reference 4.3). The methodology of calculation for Core XVIII is unchanged from previous reload analyses.

The GAPKIX code calculates pellet-to-clad gap conductance from a combination of theoretical and empirical models which predict fuel and cladding thermal expansion, fission gas release, pellet awelling, pellet densifiestion, pellet cracking, and fuel and cladding thermal conductivity.

I The thermal effects analysis encompassed a study of fuel rod temperature response as a function of the detailed cycle burnup and power. -

The fuel rod types and power histories examined in detail include:

Maximum Fresh Pin: fuel rod with the maximum average power for the fresh (reload) fuel, Hi-Power Exposed Pin: fuel rod with the maximum average power for

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the exposed (recycle) fuel.

Figures 4-1 through 4-4 illustrate the effect of linear heat generation rate (LHGR) on calculated temperatures at the beginning-of-cycle. Figures 4-5 through 4-8 illustrate the effect of LEGR on calculated temperatures for end-of-cycle conditions. These resultant temperatures are similar to those which have been reported in previous Yankee reload analyses (References 4.4 and 4.5).

These results demonstrate that the beginning-of-life (BOL) conditions yield the maximum predicted fuel temperatures. This is due to a prediction of the maximum diametral gap at BOL. The calculated internal fuel rod 1

I pressures are less than operating coolant system pressure throughout Core XVIII operation.

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4.3 Operatinz Experience The fifth batch of Exxon nuclear fuel loaded into the reactor will be discharged during this refueling outage. The batch average burnup is cpproximately 28,850 W d/MTU with a peak assembly burnup of approximately 32,850 W d/MTU. The batch of Exxon nuclear fuel discharged from the reactor during the last refueling outage had a batch average burnup of approximately 27,450 Ed/MTU with a peak assembly burnup of approximately 31,400 W d/MTU.

I The first six batches of Exxon nuclear fuel (including those remaining from Core XVII operation) have performed as expected.

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TABLE 4-1 NOMINAL MECHANICAL DESIGN PARAMETERS l

Recycled Puel Fresh Fuel (36 assemblies) (40 assemblies) l

! Fuel Pellets Fuel Material (sintered UO2 UO2 pellets)

Initial Enrichment, 3.7 3.7 w/o U-235 Pellet Density, 94.0 94.75

% theoretical Pellet Diameter, inches 0.3105 0.3105 Fuel Rod Active Length, inches 91.0 91.0 Overall Rod Length, 95.34 95.34 inchec Upper Plenum Length, 2.10 1.54 inches Fuel Rod Pitch, inches 0.472 0.472 Diametral Gap (cold), 0.0065 0.0065 inches Fill Gas Helium Helium Fill Gas Pressure, psig 250.0 250.0 Cladding Material Zr-4 Zr-4 l Outside Diameter, inches 0.365 0.365 Thickness, inches 0.024 0.024 Inside Diameter. 0.317 0.317 inches Guide Bars Material Zr-4 Zr-4 Number per Assembly 8 8 Length, inches 96.32 96.52

TABLE 4-1 (Continue 6)

NOMINAL MECHANICAL DESIGN PARAMETERS Recycled Fuel Fresh Fuel (36 assemblies) (40 assemblies)

Fuel Assembly Number of Assemblies 36 40 Fuel Rod Array 16x16 16x16 Fuel Rods per Assembly Type A 231 231 Type B 230 230 Fuel Rod Axial Clearance, 0.98 1.215 inches Outside Dimensions Assembly Cross 7.614x7.156 7.614x7.156 Section, inches overall Length, 111.775 111.790 inches Spacer Crids Material Zr-4 Zr-4 Number per Assembly 6 6 Weight of Contained

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r Uranium Type A Kg U 234 231.3 Type B, Kg U 233 230.3

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FIGURE 4-9 YANKEE CORE XVill CORE LOCATIONS OF MODIFIED ASSEMBUES 1 2 3 4 CE (4) 5 6 7 8 9 10 CE DO(ON (4) (4) 11 12 13 14 15 16 17 18 CE DO(ON (4) (12) 19 20 21 22 23 24 25 26 27 28 CE (4) ,

29 30 31 32 33 34 35 36 37 36

,. 00(ON EXXON (12) (4) 39 40 41 42 43 44 45 46 47 48 DO(ON (19) 49 50 51 52 53 54 55 56 57 58 CE D0(ON 00(ON DO(ON (12) (4) (4) (12) 59 60 61 62 63 64 65 66 CE (12) 07 68 69 70 71 72

'CE (12) 73 74 75 76 ASSEMBLY NUMBER CE FUEL TYPE (12) $ OF INERT R0r,S s

w._.-. _ _ _ _ __- _ - - _- - --__ - - - - _ .

FIGJRE 4-10 YANKEE CDRE XVIII IATTIG IDCATICNS CF SOLID R@S (G FUEL ASSEMBLIES)

YANKEE ASSEMBLY TYPE A X X X

_X D

+f LX' D

o -

O X X YANKEE ASSEMBLY TYPE B X X X O

h h - INST. 'I1[IMBLE O

O Z - mIw m f

O M X H -rew mIm m X D-neream 4M M 6 6 MM M X X MooooOZ 5.0 WUCLEAR Dg3IGN 5.1 Physics Character 13t hg Table 5-1 presents a sumary of pertinent physics data for Core XVIII.

The data is comparable to that of the current operating cycle, Core XVII.

Core XVIII radial power distributions at hot-full-power, all-rods-out conditions are presented for cycle exposures of 500, 8,000, and 14,000 ledd/NTU. In Figures 5-1 through 5-3, the maxisua unrodded radial pin power peak, F , is 1.621 at 500 BAdd/NTU,1.549 at 8000 andd/NTU, and 1.463 at 14,000 Infd/NTU. The radial power distribution at 500 Itid/NTU with control rod Croup C fully inserted is given in Figure 5-4. In this case, the Peak F is 1.794. Figure 5-5 shows the relative axial power distribution as control rod Group C is inserted for the assembly containing the peak F , at Location 42 (D-6).

Table 5-1 presents various reactivity coefficients for anticipated BOL and 30L (14,000 ledd/NTU) reactor conditions. The moderator temperatura coefficient at BOL Core XVIII is less negative than the comparable Core XVII value primarily due.to a larger boron concentration. The change in total control rod worth is due to the change in the power distribution between the two cores. The effective delayed neutron fractions for core XVIII are similar to those for Core XVII. All of the values used in the accident analysis are, chosen in a conservative manner for each analys)s.

Core XVIII, like Cores XIV through IVII, uses control Bank C as its controlling rod group. Table 5-2 presents a summary of calculated control rod group worths for Core XVIII and comparable Core XVII data. The reactivity C11owances are listed for Cores XVIII and XVII with the resulting excess shuttown margin also tabulated. The control rod group configuration remains c.onsistent with previous cores and is presented in Figure 5-6. The calculated teactivity worth for each control bank at hot-full-power conditions is given in Table 5-3 for both BOL and 30L reactor conditions. The control rod insertion limit curve for use during Core XVIII operation is unchanged from Core XVII operation and is given in Figure 5-7.

l Benon redistribution effects are accounted for by the use of two factors the Ionon Redistribution pactor (IRF) and the Reduced Load IIultiplier (Rut). The IRF accounts for changes in axial peaking due to s e trol rod Group C motion in the full power operating band (80 to 90 inches withdrawn). The multiplier is defined as the ratio of the maximum value of Fg in W analyucany derW %peaW unonhed axial sWe 6 h [

• "" "* #* "I ** * #** " '"

• value of F2*

factor is unity. This is consistent with the methodology used to derive the UIOR limits which were generated based on the worst top-peaked axial shapes.

The top-peaked axial shapes bound both nominal and botton-yeakad shapes in tosias of UIGR limits. The Rut designates a power level at which the plant l sist remain for 24 hoters if control rod Group C is inserted below 80 inches tithdrawn. This time period allows for dampening of the xenon transient and, therefore, the power redistribution. The factors as a function of Core IVIII cycle average burnup are presented in pisures 5-0 and 5-9. t Another parameter, the gF factor, was forinerly used to account for (

the change in power distribution due to control rod insertion in the full ,

power operating band. This action would temporarily shif t power to the bottosa  ;

cf the core. The current LHCR methodology incorporates limits which ,

conservatively bound nominal and botton-peaked shapes, and since the xenon redistribution factor already accounts for the maximum conservative perturbation that could occur to measured power distributions, this factor is no lonsor warranted.

I 5.2 Reactor Physics Analytical Computer Codes L

i I Yankee core depletion calculations since Core FIII have been made using o pDg/HARet0NY (References 5.1, 5.2) model with each fuel rod represented explicitly. Thus, tlhen the Core XVII assemblies were reshuffled into Core EVIII, the recycled fuel had the burnup of each fuel rod represented E explicitly. Few group cross sections, for use in PDQ, were obtained with the L50 PARD (Reference 5.3) program. l The SINULATE (Reference 5.4) program has been used for the calculation sf reactivity parameters such as moderator temperature coefficients, fuel temperature coefficients, boron worth and critical boron concentrations.

Comparisons with Cores XIV through IVII startup measurements were found to be in excellent agreement with the calculated SINULATE reactivity parameters.

For analysis of flux measurements. Yankee uses the INCORE (Reference 5.5) program. As input. INCORE requires theoretical data (PDQ) in the form of radial assembly and fuel rod powers, plus the fast and thermal fluxes in the instrumentation thimbles. The edits in the PDQ/ HARMONY are creenged to give this dets as a function of lifetime.

5.3 Channes in Analytical Methods There have not been any analytical methodology changes initiated for Core XVIII. However, the PDQ-7 model of the Yankee core is now full-core instead of the half-core model previously used. This change was necessary to evaluate the effects of changes made to some individual assemblies as described in section 4.1.

TABLE 5-1 SUltlARY OF WUCLEAR CHARACTERISTICS Core IVIII Core XVII Total Control Rod Worth, % M Hot Full Power BOL 11.55 11.26 Mot Full Power EOL 11.94 11.71 Dissolved Boron Dissolved Baron content for criticality BOL, All-Rods-Out, ppm 480 F, No Io, Peak sa 2065 1935 Hot Zero Pwwer, No Ie, Peak Sm 2050 1847 Not Full Power Eq. Io and Sm 1575 1385 Dissolved Boron Content for Refuelins, ppm 2031 1924 i

BOL, All Rods In, K = .93 1 verse Baron Worth, pps/% AP 680F, BOL, No Xe, Peak Sm 105 105 BOL, Hot Zero Power, No Ie, Peak sm 135 136 BOL, Hot Full Power Eq. Ie 136 137 Reactivity Coefficients (All-Rods-Out)

Moderator Temperature Coefficients, Af /0F Hot Full Power, BOL -1.05x10-4 -1.14x10-4 Hot Full Power, ROL, 0 ppm -3.13x10-4 -3.05x10'4 Fuel Temperature Coefficients, AP/0F Hot Full Power, BOL -1.46x10-5 -1.48x10-5 Hot Full Power, 20L, 0 ppm -1.50x10-5 -1.55x10-5 l

Moderator Void Coefficients, M /% void Hot Full P'ower, BOL -0.83x10-3 -0.87x10-3 Hot Full Power, 20L, 0 ppm -2.48x10-3 -2.32x10-3 l

c TABLE 5-1 (Continued)

I N Y OF WVCLEAR_ CHARACTERISTICS Core IVIII Core XVII Moderator Pressure Coefficients, AP/ psi Hot Full Power, DOL 1.10x10-6 1,19xio-6 Mot Full Power, IOL, 0 ppm 3.27x10-6 3.18x10-6 Effective Deleyed Neutron Fraction BOL, HZP .006492 .006422 Effective Delayed Neutron Fraction, ROL, HZP .005424 .005494 Prompt Neutron Lifetime.p sec, BOL, HZP 20.25 20.35 Prompt Neutron Lifetime.p sec, ROL, NZP 23.14 22.55 l

d e

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I l ,

l f

l TABLE 5-2 SHUTDOWN REQUIREMENTS. %6P (HFP)

Core XVIII Core XVII E E BOL EOL

1. Total Control Rod Worth 11.55 11.94 11.26 11.71

2. Worth of Stuck Rod 3.03 3.24 3.00 3.16

3. Total Worth Less Stuck Rod (1-2) 8.52 8.70 8.26 8.55

4. Total Worth Less 7.5% Uncertainty (3 x .925) 7.88 8.05 7.64 7.91 Allowances 5.1 Fuel Temperature Variation .74 .77 .76 .80 l

5.2 Moderator Temperature Variation .24 .72 .26 .70 5.3 Moderator Voids .02 .07 .03 .06 5.4 operational Maneuvering Band .06 .12 .06 .13 5.5 Shutdown Margin 5.50 5.50 5.50 5.50

5. Total Allowances 6.56 7.18 6.61 7.19

6. Excess Shutdown Margin (4-5) 1.32 .87 1.03 .72 i

i

TABLE 5-3 COMPARISON OF CONTROL ROD WORTHS (HFP)

Core XVIII Core XVII BOL EOL BOL EOL

%AP %M %AP %A9 Group C 1.73 1.82 1.74 1.82 Group A 1.38 1.49 1.36 1.47 Group B 2.56 2.51 2.42 2.43 Group D 5.88 6.12 5.74 5.99 Total 11.55 11.94 11.26 11.71 Stuck Rod 3.03 3.24 3.00 3.16 Total Less Stuck Rod 8.52 8.70 8.26 8.55 t

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,a

FIGURE 5-1 YANKEE CORE XVill RELATIVE RADIAL POWER DISTRIBUTION 500 MWD /MTU, ALL-RODS-0UT 0.564 0.740 0.767 0.568 .

0.715 1.104 1.073 1254 1.105 0.685 0.711 t151 1.161 1.1G1 1.062 1.054 1.141 0.694 0.595 t116 1.055 1.199 L210 1.146 1.19 3 1.166 1.105 0.554 0.791 1.252 t060 1.154 1.076 1.074 1.214 t170 1.087 0.745 0.759 1074 1.15 6 1.213 1.079 1.076 1.146 1.073 1.269 0.778 0.563 1.10 8 t152 1.19 6 1155 1.215 1.201 t062 1.134 0.594 0.732 1.152 1.052 1.071 1.17 9 L173 1.166 0.721 l

0.733 1.127 1.267 1.093 1.13 2 0.730 0.611 0.802 0.775 0.587

?

L

~~

[

RGURE 5-2 YANKEE CORE XVill RELATNE RADIAL POWER DISTRIB1JT10N

8000 MWD /MTU, ALL-RODS-OUT 0.614 0.780 0.795 0.612 .

0.744 1.005 1.069 1.215 1.091 0.720 0.7AO 1.12 9 1.138 1.151 1068 1.054 1.119 0.725 0.634 1.10 0 1.056 1.184 1.196 1146 1.17 6 1.137 1.000 0.602 0.815 1.214 1.068 1.15 3 1.093 t196 1.091 1.151 t073 0.779 0.796 1.071 1.146 1.198 1.095 1091 1.144 1.070 t216 0.798 0.630 1.098 1.13 0 1.17 9 1.151 1.194 t178 1.052 1.102 0.627 0.756 1.12 8 1.050 1.070 1.15 4 1.135 1.12 9 0.7 41 0.755 t100 1.215 1074 1.10 3 0.747 0.642 0.816 0.800 0.627 RGURE 5-3 YANKEE CORE XVill RELATIVE RADIAL POWER DISTRIBUTION 14,000 MWD /MTU, ALL-RODS-0UT 0.657 0.8 14 0.825 0.658 -

0.773 1.088 1.067 1.191 1.086 0.759 l

0.770 1.11 3 1.118 1.133 1.064 1.050 1.10 0 0.763 O.672 1.090 1.050 1.15 9 1.16 9 1.12 9 1.155 1.119 1.086 0.651 0.835 1.188 1.062 1.134 1.082 1.081 1.17 0 1.134 1.070 0.813 0.822 1.066 1.12 8 1.#0 1.084 1.081 1.12 7 1.064 1.191 0.826 0.667 1.088 1.110 1.156 1.132 1.167 1.156 1.048 1.002 0.666 0.778 1.110 1.044 1.063 1.135 1.116 1.114 0.771 4.777 1.087 1.187 1.067 1.093 0.775 0.674 0.834 0.826 0.667 j

FIGURE 5-4 YANKEE CORE XVill RELATIVE RADIAL POWER DISTRIBUTION  ;

500 MWD /MTU, BANK C FULLY INSERTED 0.570 0.679 0.696 0.560 0.800 1.13 2 0.832 0.957 1.104 0.751 0.791 1.294 1.240 0.992 0.881 1.107 1.267 0.767 0.596 1.13 2 1.121 1.330 1.327 1.253 1.312 1.233 1.12 2 0.554 0.730 0.969 0.890 1.270 1.258 1.252 1.324 0.992 0.840 0.681 0.706 0.836 0.993 1.335 1.261 1.254 1.253 0.889 0.970 0.708 0.594 1.14 3 1.234 1.326 1.267 1.329 1.324 1.12 0 1.14 0 0.590 i

l 0.8 21 1.130 1.116 0.893 1.007 1.248 1.304 0.797

'O.816 1.141 0.977 0.855 1.16 2 0.815 2

0.612 0.739 0.719 0.596 i

YANKEE R0WE CORE XVIII RELRTIVE RXIRL POWER IN RSSEMBLY 42 VS GROUP C INSERTION 2.0 a - ALL RODS OUT o - 0 IN 25 PERCENT a - 0 IN 50 PERCENT

+ - C IN 75 PERCENT A

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0 10 20 30 40 50 60 70 80 90 i RXIRL LOCRTION (INCHES)

FIGURE 5-6 CORE XVIII J

i.

CONTROL R0D CROUP 1

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IDENTIFICATION i

I I 1 ,2 3 4 D

5 5 7 8 9 JO i ~

. D C . --,

11 12 13 14 15 16 17 18 ,

. B B D J

J' 19 20 21 22 23 24 25 26 - 27 , ,28 D B A B .

l 29 30 31 32 33 34 35 36 37 38 l

C- A A C 39 40 41 42 43 44 45 46 47 ' 48

! B A B' D I

i 49 50 51 52 53 54 55 56 , 57 58 i D B B . .

59 60 61 62 63 64 65 66 C D 67 68 69 70 71 72  ;

5 -

D  !

73 74 75 76 ASSEMBLY POSITION # e l

l I

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Figure 5-7 YANKEE CORE XVIII ROD INSERTION LIMIT VERSUS POWER LEVEL 1

o - CORE XVIII i

F l

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2 M OPERATION j O l j CL ,

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0' 0 10 20 30 40 50 SO 70 80 90 l CONTROL GROUP C POSITION (INCHES WITHDRAWN)

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YANKEE CORE XVIII XENON REDISTRIBUTION FACTOR 1.10 O

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0 5 0 5 0 0 9 9 8 0 1 0 0 0 0 b 3 o s. N o S B B B e h

6.0 THERMAL-HYDRAULIC DESIGN The thermal-hydraulic evaluation of the reload cycle has been performed utilizing basically the same methodology as the previous seven reload analyses (Core II through Core XVII).

The reference thermal-hydraulic analysis was performed in support of the Final Safety Analysis Report (FSAR, Reference 6.1). The thermal-hydraulic cnalysis of the reload cycle was performed by adjusting code input to reflect the reload cycle power distributions and thermal-hydraulic characteristics.

Table 6-1 contains the pertinent thermal-hydraulic parameters for the reload cycle and Table 6-2 contains the same information for the reference analysis.

tote that the reference analysis performed in support of the FSAR was not cpecific to any core cycle. Therefore, all values calculated in the reference cualysis were design values. Table 6-3 rummarizes both the predicted and' design hot spot and hot channel factors for the reload cycle. Table 6-4 provides the same information for the reference analysis (design values cnly). Table 6-5 indicates the behavior of the hot channel DNBR, F H ""d F at full power conditions versus Rod Group C position for Core XVIII.

Predicted hot channel factors are based on power distributions at 500 mwd /MTU l

cycle exposure (the limiting exposure for power peaking) when Rod Group C is inserted up to 25 percent, even though rod restrictions do not permit operation at full power in this mode.

As evidenced by the information provided in Tables 6-1 through 6-5, the reload core has significant margin to DNB, coolant quality and fuel centerline melt limits. The design DNBR for the reload core is slightly lower than the l reference analysis DNBR, 2.93 versus 3.07 at full power 4-loop operation.

This difference is due primarily to an allowance for flow variations associated with the' slight variation in hydraulic resistance between the C-E and Exxon fuel assemblies obtained from flow test results (Reference 4.1) in conjunction with core inlet flow variations.

Safety limit curves for the reload cycle are provided in Figure 6-1.

These curves were developed in the Cycle 15 analysis to conservatively bound

l future reload cores. No modification of these curves is required for Care XVIII operation.

The effect of rod bow has been considered for Core XVIII operation. As r quired in Reference 6.2 for the Exxon recycled fuel, a 34% DNBR credit is n:eded to offset the very conservatively applied full-closure rod bow 1

\

penalty. Generic credit of 13.2% DNBR margin was accepted in Reference 6.2.

The most limiting anticipated transient is the 1 out of 4 pump loss of flow, i Based on design conditions, this event results in a minimum DNBR in excess of 1.8. Thus, 27.8% margin to a DNBR of 1.3 exists for this limiting event, which can be applied to the remaining 20.8% requirement for rod bow. For the C-E fuel, a rod bowing evaluation was performed which demonstrated that no bowing penalty is required for Core XVIII.

l .

l

/

l

TABLE 6-1 THERMAL-HYDRAULIC DATA SHEET FOR YANKEE CYCLE 18 DURING 4-LOOP OPERATION General Characteristics Core XVIII Desian T tal Core Power, 600 618 Wt Fraction of Heat, .973 .973 Generated in Fuel 2,000 1,925 Main Coolant Pressure, psig 520 524 Main Coolant Inlet Temperature, OF 563 568 Reactor Vessel Outlet Temperature, OF Average Core Outlet 567 572 Temperature OF Average Core 58.6 60.3 Rnthalpy Rise, Btu /lb Total Coolant Flow 38.3x10 6 38.3x10 6 Rate, Ib/hr Heat Transfer Flow 35.0x106 35.0x106 Rate, ib/hr Average Mass a 2.29x10 6 2.29x106 Velocity, Ib/hr-ft 2 Average Coolant 13.5 13.6 Velocity in Core, ft/see Core Pressure Drop, 15 15 psi Reactor Vessel 30 30 Pressure Drop, psi Average Rod Heat 158,454 163,207 Flux, Btu /hr-f t 2 Average Film 5,646 5,687 l

I coefficient, Btu /hr-ft2 _oy Average Film i 28.1 28.7 Temperature Difference OF l

Average Linear 4.44 4.57 Rod Power, kW/ft l 33.4 34.4 Average Specific

! Power, kW/kgU 4 Power Density, 90.1 92.8 kW/ liter Hydraulic Diameter, in 0.412 0.412 Assembly Heat 165 165 Transfer Area, ft2 l

l l

TABLE 6-1 (continued)

Hot Channel and Hot Spot Parameters Core XVIII Desian Maximum Heat Flux. 389,365 450,451 Btu /hr-ft2 Maximum Linear Rod Power, 10.91 12.6 hM/ft Maximum Clad Surface 634 637 Temperature, OF Maximum Centerline Pellet 2,901 3.271 Temperature, OF Hot Channel Outlet 591 608 Temperature. OF Minimum W-3 DNBR 4.06 2.93 TABLE 6-2 THERHAL-HYDRAULIC DATA SHEET FOR YANKEE REFERENCE ANALYSIS DURING 4-LOOP OPERATION General Characteristics Desian Total Core Power, 618 MWt Fraction of Heat .973 Generated in Fuel Main Coolant 1,925 Pressure, psig Main Coolant Inlet 524 Temperature OF Reactor Vessel 568 outlet Temperature, OF Average Core outlet 572 Temperature, OF Average Core 60.3 Enthalpy Rise, Btu /lb Total coolant Flow 38.3x10 6 Rate, lb/hr Heat Transfer Flow 35.0x106 Rate, lb/hr Cominal Channel Hydraulic .412 Diameter, in Average Hass 2.29x106 Velocity, ib/hr-ft 2 Average Coolant 13.6 Velocity in Core, ft/see Core Pressure Drop, 13.7 psi Reactor Vessel 32 Pressure Drop, psi Average Rod Heat '

161,624 Flux, Btu /hr-ft 2 Assembly Heat 167 Transfer Area, ft2 i Average Film 5,638 Coefficient, Btu /hr-ft2 _oy Average Film 28.7 Temperature Difference, OF Average Linear 4.53 Rod Power, kW/ft Average Specific 34.7 Power, kW/kgU Power Density, 92.8 kW/ liter TABLE 6-2 (continued)

Hot Channel and Hot Spot Parameters Design Maximum Heat Flux e 446,405 Btu /hr-ft2 Maximum Linear Rod Power e 12.5 W/ft Maximum Clad surface 637 Temperature OF Maximum Centerline Pellet 3,315 Temperature, OF Hot Channel Outlet 605 Temperature, OF Minimum W-3 DNBR 3.07 l

l s'

! TABLE 6-3

SUMMARY

OF HOT SPOT AND HOT CHANNEL FACTORS FOR YANKEE CYCLE 18 Cycle 18 Desian Heat Flux Factors Buclear Heat Flux Factor 2.30 2.59 Fuel Densification Factor 1.03 1.03 Engineering Heat Flux Factor 1.04 1.04 T;tal Heat Flux Factor 2.45 2.76 Enthaley Rise Factors Statistical Enthalpy Rise Factor 1.08 1.08 Lower Plenum Factor 1.05 1.05 Nuclear Enthalpy Rise Factor 1.56 1.80

?

r l

i 1

l I

l f

l

i TABLE 6-4

SUMMARY

OF HOT SPOT AND HOT CHANNEL FACTORS FOR YANKEE REFERENCE ANALYSIS Heat Flux Factor Buclear Heat Flux Factor 2.59 Fuel Densification Factor 1.03 Engineering Heat Flux Factor 1.04 Total Heat Flux Factor 2.76 Enthalpy Rise Factors Statistical Enthalpy Rise Factor 1.08 (Fuel Density, Pellet Diameter and Enrichment, Rod Diameter, Pitch and Bowing)

Lower Plenum Factor 1.05 Nuclear Enthalpy Rise Factor 1.80 I,

l TABLE 6-5 NOMINAL HOT CHANNEL DNBR, F AND F AS FUNCTIONS 33 OF GROUP C POSITION FOR CORE IVIII Croup C .

Fyg: FQ DNBR Position *(inches inserted) 0 1.575 2.20 4.40 7.5 1.579 2.22 4.33 15.0 1.585 2.29 4.23 22.5 1.592 2.42 4.06 CPower Dependent Insertion Limits Restrict Group C Insertion to 10 inches at Full Power.

A n

M 1

FIGURE 6-1 l ,

j .

I

,i I

k '

l 4 -

! 6M l

0 640 i

s 1

! an.

! 620 5

i s MAIN COOLANT

{ g SYSTEM j g PRESSURE a

a- 600

! I 1

j es 2600 psia 3

j 3 580 0 2400 psia l

i 5

x 2200 psia

• 560 l ,

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! M 2000 psia l [c 540 l

w I $

1800 psia

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i 520 J

1600 psia f

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500 80 90 100 110 120 130 70 Indicated Reactor Power, Percent REACTOR CORE SAFETY LIMIT - ALL LOOPS IN OPERATION 1

7.0 ACCIDENT ANALYSIS 7.1 Introduction The safety evaluation of the reload cycle is presented in this cection. Each transient in the following subsections is compared with the most recent reference analysis. For most of the transients, the reference analysis was presented in the Final Safety Analysis Report (FSAR, Reference 7.1). Additional analyses have been performed, both within and ceparate from reload reviews. Where appropriate, these investigations serve j cs the reference analysis for the Core XVIII reload cycle review. j To facilitate this comparison, tables of the parameters applicable to cil of the transients are presented in the following three subsections.

7.1.1 Initial Operatinz Conditions Table 7-1 provides the initial operating conditions for most of the transients. Any deviations from the values indicated are noted in the discussion of the specific transient. Table 7-1 also shows the design pellet centerline temperature and design W-3 DNB ratios.

Only minor differences in the basic plant parameters exist between the reload core and the reference analysis. These minor differences are the following:

l

1. The maximum Linear Heat Generation Rate (LHGR) is slightly hisher for Co.e XVIII than for the reference analysis. This difference resulted from an increase in the core average LHGR associated with the addition ot dummy fuel pins to selected assemblies, and was i I

accounted for in the thermal-hydraulic analysis of Core XVIII.

l 1

2. The minimum DNB ratio at design conditions for Core XVIII is marginally lower than for the reference analysis. The impact of this minor reduction in design DNBR will be addressed in the review of each appropriate transient.

l 4

7.1.2 Rosetor Trio Setooints and Instr'smentation Delars l Table 7-2 presents the reactor trip setpoints and instrumentation  ;

h delays applied in the transient analysis. With the exception of the rain coolant high pressure trip, these values remain the same as for the previous core. For the reference analysis, credit has been taken for the high main 1

coolant pressure trip in the analysis of the loss of load transient, which is discussed in Section 7.6.  !

7.1.3 Reactivity Coefficients i

l The moderator sad fuel temperature coefficients, where they are 1

important to the analysis, will be discussed on an event-by-event basis.

l 7.2 Control Rod Withdrawal Incident i

For the reference analysis (FSAR, Reference 7.1), a bounding analysis was performed with the following assumed conditions:

1. Design peaking factors were used even though lower peaking factors exist during the incident;

2. Core power was assumed to be at the overpower trip setpoint;

3. Coolant pressure was assumed to be at the lower end of the operating band;

4. Coolant temperatures were assumed to be at the values consistent with steady-state operation at the overpower trip setpoint.

For the reload cycle, these conservatively assumed conditions remain bounding. However, the design DNBR is marginally lower for Core XVIII than 1 was calculated for the reference analysis. The reference analysis shows that j the minimum DNBR for this event remains significantly above 1.3 (greater than f i 2.0). Accounting for the slightly lower design DNBR for Core XVIII, the consequences of this event are still within fuel design limits for the reload i i

core.  ;

7.3 Boron Dilution Incident 7.3.1 Introduction l

The most recent review of the time to loss of shutdown margin portion cf the boron dilution analysis was performed for the Core XVII reload submitted to the NRC in Reference 7.3. This analysis reviewed every mode of operation allowable in the Technical specifications. The Core XVII time to 1:ss of shutdown margin analysis was also presented in the FSAR (Reference 7.1), along with an additional analysis which investigated the minimma DNBR which could occur for a boron dilution event in Mode 1. It was  ;

concluded that the minimum DNBR was bounded by the results of the control rod withdrawal transient discussed in section 7.2. This conclusion remains valid fer the current reload cycle.

For the reload cycle, each mode of operation has been analyzed with the current cycle values for initial and critical boron concentrations. The  ;

results from these studies will be described in the detailed description of cach operational mode. These results are for the limiting beginning-of-life conditions. As the fuel eg osure increases during the cycle, the time for loss of shutdown margin increases.

7.3.2 Analysis and Results 7.3.2.1 Boron Dilution Durina Modes 1 and 2 1 l The significant parameters for the reload cycle during Modes 1 and 2 at the most limiting condition are shown in Table 7-3. As shown, the reload l

cycle in the limiting case remains well above the minimum allowable time criterionof15minNtes. The maximum reactivity insertion rate during a Mode l l 1 or 2 boron dilution was reanalyzed in support of the FSAR, which serves as the reference analysis for this portion of the boron dilution analysis. The reactivity insertion rate for Core XVIII is bounded by the reference analysis

because of the conservatively large initial boron concentration (2527 ppm) cssumed in the analysis for the FSAR. It was shown for the FSAR analysis that cven with this high value of boron concentration the reactivity insertion rate during boron dilution is bounded by the results of the rod withdrawal incident.

l

[

I l

l For the Mode 1 or 2 boron dilution event, the Low Pressure Surge Tank l

I (LPST) level, temperature and pressure indication, as well as the high alarms 3

on these three parameters, would provide information to aid ths operator in diagnosing the condition. Should the reactor be in the automatic control mode during a dilution at power, the control rod group currently selected would insert to offset any temperature increase resulting from the core power / steam flow mismatch. Audible indication of this control rod motion would be cvailable from the Containment Sound Monitoring System. Even if the control rods insert at the maximum rate of 6 inches / minute, it would take Cyproximately 15 minutes for the control rod group to be fully inserted.

During this time, the operator would be alerted by the high average temperature and high neutron flux alarus. Because of the available alarms and indications, there is ample time'to allow the operator to diagnose the situation, terminate the dilution and restore adequate shutdown margin.

7.3.2.2 Boron Dilution Durinz Mode 3 The analysis for Core XVIII boron dilution during Mode 3 used the same cyproach as the previous reload analysis. The significant parameters for the limiting cases are presented in Table 7-4. As shown, the reload cycle in the limiting case remains well above the minimum allowable time criterion of 15 cinutes.

i Close surveillance would be required for a feed and bleed operation .

performed at 100 spm, and the possibility of inadvertently feeding unborated water at this magnitude is extremely improbable. However, assuming this unlikely situation did occur, the operator would have a minimum of four alarms to alert him to a possible inadvertent dilution. These alarms include the

! high neutron flux alarm and the LPST high level, temperature and pressure clarms. Additional alarms would also be expected from the bleed line radiation monitor and possibly the main coolant pump low cooling flow alarm.

The latter alarm could be expected due to increased component cooling flow to l cool the LPST in response to its high temperature, resulting in a decrease in the cooling flow to the main coolant pumps. This event can be terminated quickly and easily from the Control Room by isolating the charging line, chutting off the charging pumps, or isolating the source of domineralized water.

l

7.3.2.3 Boron Dilution Durina Modes 4 and 5 The most limiting boron dilution for these two modes of operation is from Mode 5 conditions as demonstrated in the previous cycles. The significant parameters for the 10aiting case from these conditions are shown in Table 7-5. The reload cycle remains well above the minimum allowable time criterion of 15 minutes.

Table 7-5 shows that ample time exists for the operator to acknowledge the high neutron flux alarm. In addition, high level, temperature and pressure alarms on the LPST would key the operator to the event. The operator would then take corrective action to terminate the dilution from the control Room by isolating the charging line, shutting off the charging pumps, or isolating the source of domineralised water.

7.3.2.4 Boron Dilution Durina Mode 6 (Refuelina)

The current reference analysis for the boron dilution incident from Mode 6 conditions was performed in support of the FSAR (Reference 7.1). The tnalysis assumed the minimum volume of water to be diluted, four isolated reactor coolant loops with the upper head drained, concurrent with the largest possible dilution rate of 100 gal / min unborated water. Additionally, a shutdown margin of 7.5% A/ was assumed. The significant parameters for the tinimum water volume case of the Mode 6 boron dilution incident are given in Table 7-6. As shown, the reference analysis is more limiting than the reload cycle.

It is important to note that this combination of conditions could only cecur When the reactor vessel head is about to be unbolted. The nominal conditionsduringtNobulkoftimeatMode6 includes 32feetofwaterabove the top of the irradiated fuel assemblies and the tagging out of service of equipment that would make possible inadvertent reactivity increases, as required by Technical Specifications. A more realistic way, therefore, to examine the Mode 6 boron dilution event was used as an additional case in the FSAR analysis, and is repeated for Core XVIII. In this additional case, the active dilution volume assumes that all 4 loops are isolated, and the shield tank cavity is filled to 32 feet above the top of the fuel assemblies with f

i-one-half of the shield tank volume contributing to the active dilution l v:1ume. The significant parameters for this case are shown in Table 7-7.

Under all conditions, greater than 30 minutes is available for operator action.

In Mode 6, the operator is provided with indication of a possible boron dilution via the audible count rate signal, Which would increase, and the high neutron flux alarm. The operator can then take corrective action from the I

control Room by isolating the charging line, shutting off the charging Pumps er isolating the source of domineralised water.

7.3.2.5 Epilure to Borate Prior to Cooldown Because of the large negative temperature coefficient of reactivity at and of cycle, any decrease in main coolant system temperature increases the core reactivity state. Consequently, during the process of cooldown of the main coolant system, it is necessary to ensure adequate shutdown margin by ensuring that adequate boron concentration and/or control rod worth is cvailable.

The failure to ensure adequate shutdown margin prior to cooldown was es-evaluated for Core XVIII with the following basic assumptions:

6 m) The moderator defect vs. temperature curve is used in assessing the reactivity addition, since the moderator temperature coefficient is, a function of temperature.

/

b) The reactor is initially 1% suberitical at an average temperature of $20 F, the maximum allowed temperature at hot standby (Mode 3) condition,s.

0 c) Theshutdownmarginat520Fis5.5%A/,theminimumrequiredby the Technical specifications, d) The main coolant system temperature is reduced at the rate of 100 F/hr, the maximum cooling rate permitted. >

l In order to make the reactor critical from these initial conditions, the average coolant temperature must be reduced to approximately 495 F.

This temperature reduction requires approximately 15 minutes to accomplish.

This is ample time for the operator to diagnose the condition and take necessary corrective action. For a complete loss of shutdown margin, a cooldown to approximately 325 F would be required and would require t

typroximately 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. l

\

l J T 7.3.3 Conclusions The probability of erroneous dilution is considered very small because cf the equipment, ' controls, and administrative procedures provided for boron dilution activities. However, in the unlikely event that an unintentional dilutio'n of boron in the main coolant system occurs, numerous alarms and indications are available to alert the operator of the condition. These clates include high level, temperature and pressure on the Low Pressure Surge Tank, and the high neutron flux alarm, as well as the audible count rate oignal in place during Mode 6 conditions. If the reactor is initially critical at the time dilution begins, the automatic safety features of the reactor protection system would ensure acceptable plant performance. For boron dilutions initiated during any operational mode, adequate time exists for the operator to determine the cause of the dilution and take corrective cetion before a complete loss of shutdown margin occurs.

7.4 Control Rod Drop Incident For the reference analysis, a bounding analysis was performed assuming steady-state operation at a conservatively determined limiting end point for the transient. The following assumptions were made in the analysis:

1. The highest calculated radial peaking factors with uncertainties j

for any dropped rod were used;

2. The design core power of 618 MWt was assumed as the power level; r

3. Main coolant pressure was assumed to be at the low pressure trip setpoint to minimize the DNB ratio;

4. Core inlet coolant temperature was assumed to be at the design value consistent with the design core power.

Table 7-8 summarizes the significant parameters for this bounding cvaluation. The information provided in Table 7-8 demonstrates that the reload cycle is conservatively assumed to be at essentially the same bounding l operating conditions as the reference analysis, but with lower power peaking.

The slight increase in the maximum linear heat generation rate for the reload cycle results in a small decrease in the minimum DNBR. However, the reload cycle evaluation results provided in Table 7-8 show that fuel performance is otill well within the acceptable limits for this event.

7.5 Isolated Loop Startup Incident current restrictions preclude power operation with a loop out of s rvice, thus this transient was not addressed specifically for Core XVIII.

The analysis of this event for the reference core (Core XI, Reference 7.4) demonstrated that the minimum DNB ratia was greater than 2.97 during the transient, while the peak fuel cienterline temperature was less than 3485 F.

The most significant parameter for this incident is the most negative expected value of the moderator temperature coefficient of reactivity (MTC). Since this coefficient is less negative for the reload cycle (-4.19x10 /F va. -4.57x10 4[/Fforthereferencecycle), the results would be bounded by those shown for the reference cycle.

l 7.6 Loss of Load Incident l For the reference analysis (FSAR, Reference 7.1), a bounding analysis was performed using the design value of moderator temperature coefficient at BOC,HZPconditions(Og/F). Note that the FSAR analysis included a change in the upper limit on the pressurizer safety valve setpoint tolerance fror. +0%

to +3%. The main coolant system high pressure trip was also credited in the cnnlysis. In general, this incident is not sensitive to minor changes in core i psrameters. In addition to the reference analysis, numerous parametric l

analyses were perforined in support of the Core XIV reload submittal (Reference 7.5). These sensitivity studies on moderator temperature coefficient and Doppler coefficient demonstrated the minor impact of core physics parameters on the loss of load transient. Thus, the core XVIII plant response to a loss of load would be within system design limits, and therefore caceptable.

7.7 Loss of Feedtrater Flow Incident The most recent review of the loss of feedwater flow incident was perforined in support of the FSAR (Reference 7.1). Rach of the assumptions made in this analysis bounds the Core XVIII system characteristics. For this event, the most significant requirement is maintenance of a steam generator h:st sink. The analysis provided in Reference 7.1 concluded that plant p rformance for this event was acceptable. It was concluded that,the l combination of the reactor protection system and emergency feedwater system ccoured the integrity of the core and primary and secondary system pressure l boundaries by 1) reactor trip on low steam generator water level, and 2) cuxiliary feedwater flow sufficient to assure adequate steam generator liquid inventory for primary system cooldown, decay heat removal, and main coolant pump heat removal for the entire course of the event.

Since the Core XVIII operating conditions are bounded by this analysis, it is concluded that Core XVIII response to a loss of feedwater is acceptable.

7.8 Loss-of-Coolant Flow Incident As demonstrated in the reference analysis (FSAR, Reference 7.1), the Icss-of-coolantflodtransientissensitivetocoreparameters, reactor protection system setpoints and steady-state thermal margin. The moderator temperature coefficient for the reload core is more negative than for the

~

reference analysis (-0.55x10

/Fvs.0.04[/Fforthereference cnalysis). In addition, the minimum scram rod worth for the reload core l

(7.78%Af) is much greater than that assumed in the reference analysis j (6.0% A/) . Therefore, with the exception of the steady-state thermal margin, Core XVIII is bounden by the reference analysis.

At steady-state design conditions, the DNBR for the reload core is cpproximately 4.6% less than the reference analysis. For the case of a complete loss-of-coolant flow (the only case resulting in DNB), the censitivity study performed for the reference analysis was utilized to determine the impact of the change in design DNBR on the number of fuel pin failures. The total amount of pin failures predicted for the reload core remains below 1.15%, and thus satisfies the acceptance criteria for this event.

It is important to note that the 4 of 4 pump loss of flow transient is j i

considered to be a very unlikely event due to the main coolant pump power I cource diversification. Two main coolant pumps are powered by the generator While the two other pumps are powered by the two separate off-site AC lines.

Even with a complete loss of off-site power, two main coolant pumps would produce near full flow while the generator is coasting down, and thus reduce (

l I

the severity of this transient.

The results of the 2 of 4 loss-of-coolant flow event will be more favorable for the reload core for the following reasons:

1. Initial fuel temperatures are more favorable.

2. Overall impact of core reactivity parameters is favorable.

The 1 of 4 loss of flow analyzed in the Cycle 18 thermal-hydraulic cnalysis is the limiting anticipated loss of flow transient. The value generated in the thermal-hydraulic analysis, a minimum DNBR of 1.82, shows the Isrge DNBR margin available.

7.9 Control Rod Eiection Accident The most recent control rod ejection accident analysis was performed in support of the Final Safety Analysis Report (Reference 7.1). A comparison of pertinent parameters, including calculational uncertainties, affecting the cvent for the reference analysis and the reload core is provided in Tables 7-9 cnd 7-10.

Both the full power and zero power rod ejection parameters are more fcvorable for the reload core, with the exception of the BOL post-ejection peaking and the EOL delayed neutron fractions for both cases. The minor difference in the EOL delayed neutron fractiono is more than compensated, however, by the reduced ejected rod worth of 0.25%4)0f or full power and 0 for zero power versus 0.5%djd j 0.83% Aj and 0.93%d4for the reference analysis values. The real impact of the delayed neutron fraction is best shown through 0

thecomparisonofthe///valuesinTables7-9and7-10. The large values ured in the reference analysis significantly bound the reload cycle.

Due to the larger post-ejection peaking at BOL conditions for the reload cycle, both the zero and full power rod ejection transients were conservatively reanalyzed. The results for both cases indicate that the cverage enthalpy of the hottest fuel pin is less than 200 cal /gm, and that the centerline enthalpy is below 250 cal /gm. Therefore, no fuel cladding damage will occur for this event.

7.10 Steam Line Braak Accident The reference analysis for the main steam line break transient was performed in support of the Final Safety Analysis Report (Reference 7.1). The moderator reactivity feedback parameters and control rod worths for Core XVIII were found to be more limiting for the steam line break than the values generated for the reference analysis. As a result, a reanalysis of the main steam line break was performed for the reload cycle. This analysis included a lower safety injection actuation setpoint than was assumed for the reference cnalysis (1600 psig vs. 1700 psig, including uncertainties). The results for thelimitingcaseshowaminimumsuberiticalityof-0.22%4/. Therefore, there is no return to power following the steam line break. This precludes DNB and j

fuel cladding damage.

7.11 Steam Generator Tube Rupture Incident The analysis of this event for the reference Core II cycle demonstrated that the results are not sensitive to the core design. Thus, the results of the analysis presented in Reference 7.4 will also apply to the reload cycle.

7.12 Other Accidents and Transients Previous analyses (Reference 7.6) have demonstrated that the Waste Gas Incident, Fuel Handling Incident, Reactor Containment Pressure Analysis, Hypothetical Accident, and Post-Accident Hydrogen Considerations are not censitive to core design changes and, therefore, the results presented in i l

Reference 7.7 still apply to the reload cycle.

7.13 Transient Analysis Sununary The results of the transient analysis review of Core XVIII are shown in T:.ble 7-11. Provided are the criteria for each incident, and the results for the reference analysis as well as the Core XVIII analysis. These results provide assurance of continued safe operation of the Yankee plant.

4

. - - _ __. . - - _ - _ _ _ _ _ - _ - _ . . _ - . . _ _ _ _ - ~ . _ . _ _ _ . . . _ _ - . _ _ _ _ - _ _ . . _- - . . . .

TABLE 7-1 INITIAL OPERATING CONDITIONS Reference Analysis Reload Core FSAR Parameter (4-Loop) (4-Loop)

Re ctor Power, NWt 600 + 18 600 + 18 Cire Inlet Temperature, OF 520 + 4 520 + 4 Main Coolant Pressure, psia 2015 i 50 2015 i 50 Cinimum Reactor Coolant Flow, 35.0 35.0 106 lb/hr Maximum Linear Heat Rate, 12.6 12.5 kW/ft Axial Heat Flux Profile Cosine Cocine Total Heat Flux Factor 2.76 2.76 T;tal Enthalpy Rise 1.78 1.78 Factor Maxinum Pellet Centerline 3271 3315 Temperature, OF *

(Design Conditions)

Minimum W-3 DNB Ratio 2.93 3.07 (Design Conditions) i l

l l

TABLE 7-2 REACTOR TRIP SETPOINTS AND INSTRUMENTATION DELAYS Trio Functions Setpoint* Delay Time (sec*)

High Startup Rate 5.2 decades / min. 0.3 High Neutron Flux 112 percent 0.4 Pump Current Deviations high/ low current 0.6 on two pumps High Pressurizer Water Level 209 inches 0.65 Low Main Coolant Pressure 1735 psig 0.6 Low Steam Generator Water Level -13.0 inches 2.0 Main Steam Line Isolation Trip 200 psig 0.16 High Main Coolant Pressure 2350 psig 2.0 d

8 Identical values for both Core XVII and Core XVIII with the exception of the main coolant high pressure trip credited in the reference loss of load analysis.

TABLE 7-3

[l095S_1 .aWD 2 BORON DILUTION

• FSAR Parameter Reload Core (Core XVII)

Minimas Coolant Volume, ft 3 ** 2400 2400 Limiting Initial Boron 1951 1858 Concentration, ppm Time to Loss of Shutdown 56.9 58.0 Margin, Minutes

. o*

o These operating modes are defined as follows:

Mode 1 2 Description Power Operation Startup Reactivity, K gg e >.99 >.99 Power Level, % Rated 12 <2 Coolant Temperature,0F 2330 >330 o* Assumes one main coolant loop isolated.

TABLE 7-4 MODE 3 BORON DILUTION

• FSAR Parameter Reload Core (Core XVII)

Minimum Coolant Volume, ft 3aa 1700 1700 Limiting Initial Boron 1893 1770 Concentration with Uncertainties, ppm Time to Loss of Shutdown 37.3 36.3 Margin, Minutes L

• Mode 3 is defined as follows:

Description Hot Standby Reactivity, K,gg < .99 Power Level, % Rated 0 Coolant Temperature,0F >330,

• • Assumes three main coolant loops isolated.

TABLE 7-5 MODES 4 AND 5 BORON DILUTION

• FSAR Parameter Reload Core (Core XVII)

C1nimum Coolant Volume, ft 3 ** 1276 1276 Limiting Initial Boron 2007 1932 Concentration, ppm Time to Loss of Shutdown 25.5 26.2 Margin, Minutes O These operating modes are defined as follows:

Mode 4 5 Description Hot Shutdown Cold Shutdown Reactivity, K,gg <.96 f.96 Power Level, % Rated 0 0 Coolant Temperature,0F 200 < T < 330 $200

• • Assumes four main coolant loops isolated, with provisions for upper reactor vessel head draining.

TABLE 7-6 MODE 6 BORON DILUTION

• ALL LOOPS ISOLATED.

MINIMUM WATER LEVEL FSAR Parameter Reload Core (Core XVII)

Minimum Coolant Volume, ft 3 1276 1276 Initial Boron Concentration, ppm 2269 2137 Time to Loss of Shutdown 35.6 34.5 Margin, Minutes

• Mode 6 is defined as follows:

Description Refueling Reactivity, K gr $.95 Power Level, % Rated 0 Coolant Temperature,0F 1140 l

TABLE 7-7 MODE 6 BORON DILUTION

• WATER VOLUME INCLUDING SHIELD TANK l l

l Reference Analysis Parameter Reload Core (FSAR) i Minimum Coolant Volume, ft 3** 8650 8650 Initial Boron Concentration, Ppm 2007 1932 Time to Loss of Shutdown 173.0 177.0 Margin, Minutes l

l A Mode 6 is defined as followc:

Description Refueling Reactivity, K,gg <.95 Power Level, % Rated 0 Coolant Temperature,0F 1140 C* Assumes 4 loops isolated and shield tank cavity filled to 32 feet above fuel assemblies, with 1/2 of the shield tank volume contributing to the active dilution volume.

1

l l

l TABLE 7-8 1

f CONTROL ROD DROP INCIDENT PARAMETERS Reference Analysis Parameter Reload Core (FSAR)

Core Power,IEft 618 618 Core Inlet Temperature, OF 524 524 Main Coolant Pressure, psia 1750 1750 Maximum Linear Heat Rate 14.7 14.4 (with uncertainties),

itW/ft Maximum Fuel Centerline 3693 3704 Temperature, OF Minimum W-3 DNB Ratio 2.12 2.17 I

o

,e s

TABLE 7-9 .

H_ZP ROD EJECTION ACCIDENT PARAMETERS Parameter Reload Core Reference Analysis E E (FSAR)

Moderator Temperature Coefficlent (10-4 A//*F) +0.12 -1.99 +0.27 Doppler coefficient (10-5 4pfoF) -1.13 -1.17 -1.10 Ejected Rod Worth

(%of) .79 .83 .93 Delayed Neutron

.006492 .005424 .005490 Fraction (/)

[/p 1.218 1.525 1.694 Fq Following Rod Ejection 4.51 4.04 4.22 4

l l TABLE 7-10 i

l NFP ROD EJECTION ACCIDENT PARANETERS Parameter Reload Core Reference Analysis EQk EQL (FSAR)

Moderator Temperature coefficient (10-4 A//oF) -0.55 -2.63 0.0 Doppler coefficient (10-5affoF) -1.00 -1.04 .766 Ejected Rod Worth

(% A/) .2109 .2500 .5 Delayed Neutron Fraction (/) .006492 .005424 .005743

/3/ .325 .460 .871 Fg Following Rod Ejection 3.37 2.58 3.17 i

?

~ _ - . _ _ _ - _ _ - - - _ _ _ _ - _ _ _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ .

TABLE 7-11 YABEEE CYCLE 18 SAFETY ANALYSIS ,

SUBGEARY OF RESULTS 1

Incident I Section Criteria Reference Analysis Core XVIII i Control Rod 7.2 MDWBR > 1.30 MDNBR > 2.00 MDNBR greater than 1.90 Withdrawal RCS pressure RCS pressure RCS pressure l

l _ < 2750 psia -

= 2503 psia less than 2600 psia Boron Dilution 7.3 Suberitical: Suberitical: Suberitical:

Sufficient time for Creater than 25 minutes - Greater than 25 min. -

operator action Medes 3, 4 and 5 Modes 3, 4 and 5 Greater than 30 min. - Greater than 30 min. -

l Mode 6 Mode 6 critical: Critical: Critical:

Reactivity addition Bounded by control cod Bounded by control rod y rate withdrawal withdrawal Control Rod Drop 7.4 MDNBR 2 1.30 MDNBR = 2.10 MDNBR greater than 2.00 No fuel centerline Fuel centerline Fuel centerline melt temperature = 37500F temperature = 36930F Isolated Loop 7.5 MDNBR 2 1.30 MDNBR greater than 2.97 power operation with loop Startup No fuel centerline Fuel centerline out of service prohibited melt temperature = 34850F by Technical Specifications Loss of Load 7.6 RCS pressure Maximum RCS pressure = Maximum RCS pressure =

less than 2750 psia 2663 psia 2663 psia

TABLE 7-11 (continued)

YANEEE CYCLE 18 SAFETY ANALYSIS StNSEARY OF RESULTS Incident 1 Section Criteria Reference Analysis Core XVIII Loss of Feedwater 7.7' Sufficient time for , Emergency feedwater Emergency feedwater initiation of required 15 minutes required 15 minutes emergency feedwater following event following event Loss of Coolant 7.8 10CFR100 Less than 1.25% fuel Less than 1.15% fuel Flow failure failure Control Rod 7.9 10CFR100 No clad damage No clad damage Ejection w

Y Steam Line 7.10 Maintain fuel rod No return to critical No return to critical Rupture integrity Steam Generator 7.11 10CFR100 Radiological doses well Radiological doses well Tube Rupture within 10CFR100 within 10CFR100 l

8.0 STARTUP PROGRAN l

Following refueling and prior to vessel reassembly, fuel assembly prsition will be verified by underwater television and videotaped.

The Core XVIII Startup Program will include the following tests:

1. Control rod operability tests will be performed by moving each rod group in turn from 0" to 90" to 0" and verifying control rod movement by the rod position indicators.

2. Control rod drop time measurements will be conducted by withdrawing one rod group at a time to 90 inches, dropping it and measuring individual rod drop times with a recording oscillograph.

3. Just critical boron concentration is determined by placing the reactor just critical, allowing for system equilibrium and taking a series of main coolant boron samples. This will be done as close as possible to the conditions of all rods out, Group C inserted and Groups A.and C inserted.

4. Control rod group worths of Group C, Group A, and Group B will be determined. This is done by establishing a boron change, balancing the reactivity change with a control rod position change and measuring the reactivity worth of the rod steps with the reactivity computer.

5. Isothermal temperature coefficient measurement is completed by changing main coolant temperature and measuring the reactivity change wl'th the reactivity computer. Several such measurements are taken for each of a number of equilibrium boron conditions and control rod group insertions.

6. Power and xenon defects are interred using a reactivity balance before and af ter power ascension.

7. Power distributions will be measured as soon as the reactor is at steady-state power (50%1 Power Level:E 75%). This is done with the incore instrumentation system. A power distribution map will also be taken at a low power level to check for gross quadrant tilt.

8. A startup test report on the above will be submitted to the NRC 90 days after startup.

The acceptance criteria for the prediction of key core parameters is defined in Table 8-1. The permissible deviation from predicted values are colected to insure the adequacy of the safety analysis. In these tests, the nominal measured value is compared to the nominal calculated value.

Corrections are made for any differences between the measurement and calculational conditions.

If the criteria in Table 8-1 are not met, the deviations are evaluated relative to the assumptions in the safety analysis for the given core parameters. The Plant Operations Review Committee reviews the evaluation prior to power operation.

I I

I e

(- .

TABLE 8-1 YANKEE CORE XVIII STARTUP TEST ACCEPTANCE CRITERIA Measurement Conditions Criteria

1. Control Rod Drop Time Operating temperature Drop times no greater than 2.5 seconds 20 Critical Boron Hot zero power, near Measurement within Concentration all rods out i10% of predicted value

3. Control Rod Group Hot zero power, Groups Worth of each group Worths C, A and B within 17.5% of the predicted value

4. Control Rod Group Hot zero power, Groups If the criteria in Worths C, A and B Measurement (3) is not met, the total worth of all Groups measured must be within 17.5% of the predicted value

5. Isothermal Temperature Hot zero power, near Measurement within Coefficient all rods out 10.5x10-4 Aj0 /0F

- of predicted value

6. Radial Power Above 50% power with all The measured reaction Distribution rod groups greater than rates within 15% of 80 inches withdrawn the predicted value in the high power assemblies 9.0 LOSS-OF-COOLANT ACCIDENT 9.1 Introduction The most recent large break spectrum loss-of-coolant accident analysis was perforined for Core XVI (Reference 9.1) while the most recent small break spectrum analysis was performed for Core XIII (Reference 9.2). Additional 1crge break spectrum analyses described below have been performed in support cf Core XVIII. These analyses form the basis for Cycle XVIII operation. The C:re XVIII analysis takes into account the following changes from previous cC1culations:

a) A slightly modified reload fuel design, b) Accommodation of noncosine axial power shapes, c) A higher allowable core inlet temperature, and d) A model change reducing the so-called injection delta-P penalty during pumped ECC flow.

The effects of these changes on ECCS performance are discussed in Sections 9.2 cnd 9.3 for the two categories of LOCAs (small break and large break).

9.2 Small Break LOCA The assumptions made in performing the Core XIII small break analysis (Reference 9.2) conservatively encompass any minor changes in physics p;rameters that occu,e from cycle to cycle. For Core XVIII operation, the Safety Injection Actuation Signal (SIAS) low pressure actuation setpoint is lowered from 1700 to 1650 psis while the allowable core inlet temperature is increased from 515 to 520 F. For purposes of analysis, the revised values cf these p.rame.cers with uncertainties are 1600 psig and 524 F, respectively.

The change in SIAS low pressure actuation setpoint does not affect the small break LOCAs considered in Reference 9.2 for two reasons. First, the RCS pressure decay following these small break LOCAs is fairly rapid such that the I

SIAS is generated very quickly. The lowering of the SIAS setpoint will not citer this behavior. Secondly, the ICCS pump start time is unaffected since i ECCS pump start follows emergency diesel generator start upon loss of off-site power at time zero. Similarly, increasing T-inlet to 524 F will not alter small break results since saturation conditions within the RCS are achieved i v:ry rapidly following the small break. The increase in T-inlet from 519 F to 524 F will not measurably alter the time to reach saturation within the RCS.

The core XIII small break LOCA analysis was performed with the limiting fuel stored energy which was calculated to occur for the fresh fuel at BOC conditions. A cosine axial power distribution was used for this analysis.

The 10-inch, 7 1/2-inch, 5-inch, 4-inch and 2 1/4-inch diameter breaks were cnalyzed with a PLHCR of 12.85 kW/ft. The 4-inch break with a PCT of 1793 F was found to be the limiting small break.

The following assessment was made to evaluate the impact of the top-skewed power shapes on the small break LOCA results:

a) Cosine axial power shapes are expected up to 4 GWd/MTU; top-skewed power shapes are expected beyond 4 GWd/MTU. However, the allowable Core XVIII PLHGR is maintained below 11 kW/ft beyond 4 GWd/MTU which is significantly lower than the 12.85 kW/ft value used in the Core IIII small break analysis.

b) More than 400 F margin in PCT is available for the most limiting small break.

c) Core mixture level behavior for the worst small break was examined'. The total uncovery time at various core elevations were estimated. It was concluded that the uncovery time for a top-skewed shape power peak was very close to the uncovery time for a cosine shape power peak. Therefore, for a given PLHCR, the PCT calculated for a top-skew power shape will be very close to that obtained with a cosine power shape.

Based on these reasons, it was concluded that Core XIII small break LOCA analysis conservatively bounds Core XVIII operation and that a large break LOCA with a top-skewed power shape would still bound a small break LOCA.

9.3 Larne Break LOCA The large break LOCA analysis for Core XVIII has been perforised in two phases. In the first phase, similar to past reload analyses, the applicability of the most current break spectrum analysis (Core XVI) to the reload core (Core XVIII) was assessed and a burnup sensitivity study was performed to estrablish linear heat generation limits. These analyses were based on cosine axial power distributions. In the second phase, in response t9 Reference 9.3 for Core XVIII, additional break spectrum and'burnup s nsitivity analyses were performed to address the issue of core axial power

'sistributions and conforinance to Section I. A of Appendix K to 10CFR Part 50, which requires that a "... range of power distribution shapes and peaking fcctors representing power distributions that may occur over the core lifetime" be considered.

In the following sections, the following items are addressed for Core XVIII large break LOCA: 1) core inlet temperature increase to 524 F; 2) lowering of the SIAS low pressure actuation setpoint to 1600 psig; 3)

Steam-Emergency Core Cooling (ECC) interaction model; 4) axial power shape s:nsitivity study; and 5) Core XVIII burnup sensitivity analysis.

9.3.1 Core Inlet Temperature Increase The Technical Specification setting the maximum allowable core inlet temperature has been, increased from 515 F to 520 F. Including uncertainties, the s'esumed values for this parameter for analysis purposes is 519 F and 524 F, respectively. The effect of raising the core inlet temperature from 519 F to 524 F on the large break LOCA calculations was cddressed by performing sensitivity analyses. These sensitivity analyses consisted of large break calculations for Core XVIII assuming the 524 F core inlet temperature and cosine axial power distributions. The results were compared to the Core XVI break spectrum analysis (cosine-based), which had b:en performed with a core inlet temperature of $19 F. It was found that a

(

l ccre inlet temperature of 519'F produced more conservative results than 524'F, so it was decided that the Core XVIII analysis would be done assuming o core inlet temperature of 519 F.

9.3.2 SIAS Setsoint Decrease For Core IVIII operation, the SIAS low pressure actuation setpoint has been lowered to 1650 pois from 1700 peig. For the purpose of analysis, the value of this parameter with uncertainties are 1600 pois and 1650 psig respectively. The change in SIAS low pressure actuation setpoint does not affect large break LOCA calculations for two reasons. First, the RCS pressure decay following a large break LOCA is very rapid such that the SIAS is generated sinost instantaneously. The lowering of the SIAS setpoint will not citer this behavior. Secondly, the ICCS pump start time is unaffected since it follows emergency diesel generator start upon loss of off-site power at time sero. Thus, the time at which ICC water is injected into the reactor vessel is unaffected by the change.

9.3.3 Steam-Emeraency Core Coolina (ECC) Water Interaction Model l

l The steam-ECC water interaction model described in Reference 9.4 has i been '.ncorporated into the Yankee RELAP4-EM flood model. In this model, a 0.15 paid frictional pressure loss penalty is applied during reflood to account for the thermal-hydraulic interaction between steam and pumped SCC

' flow. The penalty applied during accumulator water injection (1.8 psid) remains unchanged.

l The axial power shape sensitivity analyses described below and the Core IVIII burnup study (except for cosine-based burnup points) ut111 e the above pumped ECC injection pressure loss penalty model.

1 l

9.3.4 Axial Power Shane sensitivity Study Since 1975, Yankee reload LOCA analyses have been based on cosine axial power distributions. This assumption closely represents actual axial power shapes during the first several months of operation. However, after cpproximately 4 GWd/MTU, the actual core axial power shapes become flatter and

botton-skewed. These flatter shapes are susceptible to xenon transients )

induced by allowable control rod motion at full power. The resulting skewed axial power shape becomes more peaked in the upper half of the core. This cffect is least-pronounced during early to mid-cycle operation when the axial shapes are more cosine in nature. By the end-of-cycle, top-skewed axial power distributions can occur during a xenon transient.

In order to address this shape behavior in the LOCA analysis, a partial break spectrum analysis was performed to determine the sensitivity to the assumed axial power shape. This analysis consisted of performing, at BOC 18 conditions, blowdown and hot channel calculations (to end-of-bypass) for the break spectrum points considered in the Core XVI analyses. The axial power shape used was the 1.259 top-skewed shape shown in Figure 9-1. (Also shown on F15ure 9-1 is the 1.2 cosine axial power shape used in the Core KVI break spectrum analysis.) At end-of-bypass, two break sizes, the 0.8 DECLG and the 1.00 DECLS, were candidates for the limiting break. Since the top-skewed shape shown in Figure 9-1 is more appropriate for end-of-cycle conditions, the C:sessment of the above two break sizes was continued at a core average burnup cf 14 GWd/NTu. This assessment began with hot channel calculations for high power exposed fuel at 9.4 kW/ft, utilizing the boundary conditions from the 1.0 DECLS and 0.8 DECLG top-skewed break spectrum blowdown results, respectively. The results showed that the 1.0 DECLS break size was the more limiting of the two breaks. Thus, the 1.0 DECLS top-skewed break size was chosen as the limiting break for noncosine type axial power shapes.

With the question of limiting break size aside, an assessment of nominal power shape versus xenon-induced top-skewed power shape was made. .

Shown on Figure 9-2 are the axial power shapes employed in this assessment on o linear power basis. The nominal shape was evaluated for the recycled fuel at14GWd/NTuatafeakLHGRof9.4kW/ftutilizingthe1.0DECLSblowdown results as boundary conditions. The results for this case yielded a peak clad temperature of 1645 F, whereas the top-skewed case conducted at 9.4 kW/ft cxceeded 2200 F. These results show that the xenon-induced transient axial power shape is clearly more limiting than the nominal power shape. Therefore, the " worst case" burnup-dependent xenon-induced power shapes were employed in the Core XVIII burnup sensitivity analysis.

9.3.5 core XVIII turnue Sersitivity study Based upon the above findings, burnup sensitivity analyses were perfomed utilizing the " worst case" xenon transient axial power shapes predicted to occur for a given fuel type at the particular burnup being analysed. Additionally, hot channel boundary conditions were taken from the blowdown calculation for the limiting break (1.0 DECLS) identified above for the top-skewed axial power profiles. Except for the use of the " worst case" menon power shape end the modified pumped ICC injection delta-P model, the come analytical techniques employed in previous burnup sensitivity analyses were again used for Core IVIII.

The burnup sensitivity study addressed both fresh fuel (C-E) and recycled fuel (INC) performance in Core XVIII. The fresh fuel (C-E) was cvaluated at six exposures: SOC 18, 0.25 GWd/MTU, 1 GWd/NTU, 4 GWd/MTU, 10 GWd/NTU and 14 GWd/NTU. The effect of burnup on the recycled fuel (ENC) was determined at five exposures SOC 18, 4 GWd/NTU, 10 GWd/MTU, 14 GWd/MTU and Roc 18. Since the exposed fuel is more limiting, the Roc 18 recycled limit bounds the fresh fuel at ROC 18.

Of the above burnup points, those points with less than 4 GWd/MTU (SOC 18 to 1 GWd/NTU) have been analyzed using cosine axial power distributions, and employ boundary conditions from the blowdown calculation for the limiting break size (0.8 DECLS) identified for cosine shapes in the core XVI break spectrum study. During this period, the cosine axial power shape closely approximates the " worst case" tenon shapes predicted to occur.

Shown on Figure 9-3 is a comparison of the cosine axial power shape used in the burnup evaluation for the recycled fuel at BOC 18 to the " worst case" xenon shape at that time. This shape behavior is very typical during the carly part of core operation. Thus, utilizing cosine-based limits at these burnup points is appropriate. The results of the Core XVIII burnup sensitivity study are given in Table 9-1. The Technical specification limit for monitoring purposes is defined on the basis of power generation in the fuel rod, thus the total LHCR in Table 9-1 is reduced by the direct moderator heating fraction, 2.7%, and the resulting limit values are plotted in Figure 9.4.

9.4 Conclusions Based on the analysis presented in Sections 9.0 through 9.3, operation within the limits specified in Figure 9-4 yields LOCA results within the specifications of 10CFR50.46.

a 4

I 1

l

TABLE 9-1 CORE IVIII BURNUP SENSITIVITY STUDY RESULTS CAB Fuel PLHGk PCT GWd/MTU Iggg, (W/ft) (E) 0.00 Fresh 10.20 1994 0.25 Fresh 11.25 2145

+

1.00 Fresh 11.80 2157 4.00. Fresh 11.00 1971 10.00 Fresh 9.60 1973 14.00 Fresh 9.40 2146 0.00 Recycled 11.45 2160 4.00 ,

Recycled 10.50 1924 10.00 Recycled 9.30 2148 14.00 Recycled 9.10 2093 17.50 Recycled 8.00 1565 i

l (1) Cycle 18 Average Burnup (2) Peak Linear Heat Generation Rate (3) Peak Clad Temperature

-s4-l l

1

YANKEE CORE 18 LOCA - RXIAL SHAPE STUDY 8REAK SPECTRUM AXIAL POWER SHAPES C16-(1.2 COSINE) C18-(1.2S9 TOP SKEN) 1.5 '

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0 10 20 30 40 50 60 70 80 90 100 PERCENT OF CORE HEIGHT 14' N

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Figure 9-2 Yankee Core XVIII LOCA Axial Shape Study Nominal vs. Xenon Axial Power Shapes Exposed Fuel HFP 14 GWD/MTU 8-w Ew r a.

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Figure 9-3 Yankee Core XVIII LOCA Axial Shape Study Cosine vs. Xenon Axial Power Shape BOC18 Exposed Fuel 8

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RLLOWABLE PEAK RCD LHGR VERSUS CYCLE BURNUP 12 o - FRESH FUEL o - EXPOSED FUEL i '

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0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 AVERAGE BURNUP (GWD/MTU)

10.0 REFERENCES

4.1 CE Report, "First C-E Supplied Rowe Fuel Batch Design and Development Report", YR-CE-R-027, June 3, 1985.

4.2 Proposed Change No. 125, " Core III Refueling", rubmitted to USNRC on July 14, 1975, and Supplement No. 3 to Proposed Change No. 125, submitted November 7, 1975.

4.3 K. P. Galbraith, "GAPEII: A Computer Code for Predicting Pellet-to Cladding Heat Transfer Coefficients", IN-73-25, August 13, 1973.

4.4 YAEC-1325. " Yankee Nuclear Power Station Core XVI Perforinance Analysis,"

September 1982.

4.5 YAEC-1397 " Yankee Nuclear Power Station Core XVII Performance Analysis,"

January 1984.

5.1 W. R. Cadwell, "PDQ-7 Reference Manual" WAPD-TM-678, January 1967.

5.2 E. J. Breen, O. J. Marlowe and C. J. Pfeifer, " HARMONY: System for Nuclear Reactor Depletion Computation", WAPD-TM-478. January 1965.

3.3 R. F. Barry " LEOPARD - A Spectrum Dependent Nonspatial Depletion Program", WCAP-2795, March 1965.

5.4 D. M. VerPlanck, " SIMULATE - A Reactor Simulator Code", August 1973.

5.5 W. D. Leggett and L. D. Eisenhart, "The INCORE Code", WCAP-7149, December 1967.

6.1 Final Safety Analysis Report, Yankee Nuclear Power Station, July 1985.

6.2 USNRC Letter, D. Crutchfield to J. A. Kay, dated July 22, 1981.

7.1 Final Safety Analysis Report, Yankee Nuclear Power Station, July 1985.

7.2 USNRC Letter D. Crutchfield to J. A. Kay, dated July 22, 1981.

7.3 YAEC-1397, " Yankee Nuclear Power Station, Core XVII Performance Analysis," January 1984.

)

7.4 Proposed Change #115. " Core XI Refueling," submitted DOL /AEC on March 29, 1974..

7.5 Letter WYR 78-99, dated November 21, 1978, D. E. Vandenburgh to USNRC,

) " Additional Information - Core XIV Refueling."

7.6 Change No. 97 to License DPR-3 (Docket No. 50-29).

7.7 Amendment No. 9. Letter from K. R. Goller, DOL /AEC to YAEC, Attn:

G. C. Andognini, July 30, 1974.

l

9.1 YAEC-1325, " Yankee Nuclear Power Station Core XVI Performance Analysis,"

September 1982.

9.2 Proposed Change No. 145, Supplement No. 7. WYR 77/90, " Additional Yankee Rowe Core XIII Small Break Analysis," September 21, 1977.

9.3 USNRC Letter, J. A. Zwolinski to J. A. Kay, dated May 22, 1985,

Subject:

"Confirination of ECCS Codes."

l 9.4 Letter, FYR 85-88, dated August 16, 1985 G. Papanic, Jr., YAEC to T. A. Zwolinski, USNRC,

Subject:

"LOCA Injection Delta-P Penalty."

l l

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_ _ _ _ _ .