ML20214R637

Technical Rept Supporting Cycle 15 Operation

ML20214R637
Person / Time
Site:	Haddam Neck File:Connecticut Yankee Atomic Power Co icon.png
Issue date:	06/30/1987
From:	CONNECTICUT YANKEE ATOMIC POWER CO.
To:
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ML20214R609	List: B12537, Proposed Tech Specs Supporting Cycle 15 Operation & Reflecting Upgraded Methods for Analysis of non-LOCA Events ML20214R608 ML20214R637
References
NUSCO-155, NUDOCS 8706080230
Download: ML20214R637 (40)

Text

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NUSCO-155 June 1987 i

CONNECTICUT YANKEE ATOMIC POWER COMPANY l

HADDAM NECK PLANT Technical Report Supporting Cycle 15 Operation 4

d Northeast Utilities P.O. Box 270 Hartford, Connecticut 6706080230 870601 PDR ADOCK 05000213' P PDR i

CONTENTS Page

1. I n t ro du c ti o n a nd S umma ry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1-1

2. Op e ra t i n g H i s t o ry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-1

3. General Description.................................. .... 3-1

4. Fuel System Design........................................ 4-1

5. Nuclear Design............................................ 5-1

6. Thermal-Hydraulic Design.................................. 6-1

7. Accident and Transient Analysis........................... 7-1

8. Startup Program - Physics Testing......................... 8-1

9. References................................................ 9-1 List of Tables Table 3-1 Cycle 14 Discharge Fuel................................... 3-3 4-1 Nominal Fuel Design Parameters............................ 4-5 5-1 Haddam Neck Plant Physics Parameters...................... 5-3 5-2 Haddam Neck Plant Cycle 15 Shutdown Margin Calculations With Maximum Stuck Rod....................... 5-4 6-1 Cycle 15 Thermal-Hydraulic Data........................... 6-2 List of Figures Figure 1

3-1 Haddam Neck Plant Cycle 15 Core Loading Pattern........... 3-4 I I

3-2 Haddam Neck Plant BOC15 Burnup Distribution I in MWD /MTU................................................ 3-5 )

3-3 Haddam Neck Plant Cycle 15 Control Rod Locations.......... 3-6 I l

5-1 Haddam Neck Plant Cycle 15 Relative Power Distribution at 150 MWD /MTU, HFP, AR0.................... 5-5 i

I

1. Introduction and Summary The objective of this report is to support the operation of the fifteenth cycle of the Haddam Neck Plant at its rated core power of 1825 MWt.

Included are the analyses outlined in the USNRC document, " Guidance for Proposed License Amendments Related to Refueling." Since it is the licensee's intention to replace expended fuel with fuel of similar design, references are made to previoucly supplied analyses wherever possible.

The nominal 11,850 MWD /MTU (419 effective full-power days) Cycle 15, based on an expected Cycle-14 burnup of 10,040 MWD /MTU (355 effective full power days), is scheduled to begin in September 1987. The end of cycle burnup window assumed for Cycle 14 was 9,500-11,040 MWD /MTU, which corresponds to an end of Cycle 15 burnup of 11,200-12,200 MWD /MTU. The reviews of the fuel mechanical performance.in Section 4, the thermal hydraulic performance in Section 6, and the accident and transient analysis in Section 7 were based on these burnup windows.

Based on the analyses performed and review of the proposed revisions to Technical Specifications, it is concluded that the Haddam Neck Plant can l l

be operated safely at the rated power level of 1825 MWt for Cycle 15.

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)

2. Operating History 1

Initial criticality for Cycle 14 occurred on May 6, 1986. The plant j phased on line May 10, 1986, and reached 100 percent power on May 22, 1986. The cold leg Resistance Temperature Detectors (RTDs) were relocated and replaced during the 1986 outage in order to_ eliminate previously observed cold leg temperature streaming effects. Startup test results.

showed that past fuel cycles have operated at a vessel average temperature approximately 5*F higher than the indicated value. Cycle 14, therefore, I

has operated at a vessel average temperature 5*F less than previous cycles. Analyses have shown that past operation at higher temperature has had a negligible effect on the fuel cycle design. The Cycle 15 100 percent power vessel average temperature will be increased from 557 r

to 562*F. The Cycle 15 design and analyses reflect this change. Cycle 14  !

1 operation is scheduled for completion in July 1987. No operating anomalies

, occurred during the fourteenth cycle that would adversely affect fuel performance.

E i

4 j 2-1

3. General Description The reactor core of the Haddam Neck Plant is described in detail in Section 4 of the Facility Description and Safety Analysis Report.(1) The Cycle 15 core consists of 157 fuel assemblies, each of which is a 15 by 15 array containing 204 fuel rods, 20 control rod guide tubes, and one incore instrument guide tube. The fuel pin cladding for 153 of the fuel assemblies is stainless steel (type 304) with an outside diameter (OD) of 0.422 inch and a wall thickness of 0.0165 inch. Four batch ISB Zircaloy

-clad lead test assemblies (LTAs) have fuel pin cladding with an OD of 0.422 inch and a wall thickness of 0.027 inch. All fuel consists of bevelled, dished-end cylindrical pellets of uranium dioxide. Batch 15A, 15C, 16, and 17 fuel pellets are 0.3825 inch in diameter and 0.458 inch in length. The 56 batch 17 fuel assemblies have an average nominal fuel loading of 411.5 kg of uranium and an undensified nominal active fuel length of 120.5 inches. The minimum batch theoretical fuel density is 94.9 percent for all Cycle 15 fuel batches.

Figure 3-1 is the core loading diagram for Cycle 15 of the Haddam Neck Plant. The nominal initial enrichment for batches 15A, 15C, 16, and 17 is 4~.00 wt. percent uranium-235, and 3.41 wt. percent for batch ISB.

The 57 fuel assemblies that will be discharged at the end of Cycle 14 are from batches 9,11, and 14 (see Table 3-1).

3-1 j

The 44 batch ISA, 4 batch ISB and 52 batch 16 assemblies will be shuffled I

to new locations at the beginning of Cycle 15. Once-burned assembly R40 (batch ISC), discharged at the end of Cycle 13, will be reinserted as the center assembly in Cycle 15. The 56 fresh batch 17 assemblies will occupy the outer row of assemblies (see Figure 3-1). Figure 3-2 is a i

quarter-core map showing the assembly burnup distribution at the beginning of Cycle 15.

Reactivity control is supplied by 45 full-length Ag-In-Cd control' rods and by soluble boron shim. All 45 control rods are being replaced in Cycle 15 due to previously observed cracking and wear problems. The new control rods have extended life features including Inconel cladding for improved wear resistance, reduced poison diameter in the bottom twelve i

inches to offset pellet swelling and pressurized poison rods to enhance j

i creep resistance. The impact of the new control rods on the core physics

characteristics has been evaluated and determined to be negligible. The Cycle 15 locations of the 45 control rods and the group designations are indicated in Figure 3-3. The core locations of the 45 control rods for Cycle 15 are identical to those of Figure 4.2-2 of the Facility Description and Safety Analysis.

i 3-2

Table 3-1. Cycle 14 Discharged Fuel No. of Batch assemblies Cycles burned 9 1 3 (Cycle 14 center assembly) 11 4 4 14 52 3 Total discharged 57 r

t O

C l

l 1

3-3

i Figure 3-1. Haddam Neck Plant Cycle 15 Core Loading Pattern

~

15 14 13 12 11 10/^9 8 7 6 5 4 3 2 l l CALLED NCRTH l 17 17 17 R 17 17 17 15A 17 17 17 P 17 17 15A 15A 1G 15A ISA 17 17 N 17 17 15A 16 16 16 16 16 ISA 17 17 M l 17 17 ISA ISB 16 15A 16 15A 16 ISB 15A 17 17 L 17 15A 16 16 15A 16 ~16 16 15A 16 16 15A 17 K 17 17 15A 16 15A 15 16 15A 16 16 15A 16 15A 17 17 J 17 15A 16 16 16 16 15A 15C 15A 16 16 16 16 15A 17 H.

l 17 17 15A 16 15A 16 16 15A 16 16 15A 16 15A 17 17 G l

17 15A 16 16 15A 16 16 16 15A 16 16 15A 17 F 17 17 15A 15B 16 15A 16 15A 16 ISB 15A 17 17 E 17 17 15A 16 16 16 16 16 ISA 17 17 D 17 17 15A 15A 16 15A 15A 17 17 C 17 17 17 ISA 17 17 17 8

. 17 17 17 A initial Region # Assemblies w/o U235 44 4.00 15A(l) 15B 4 3.41 15C(2) 1 4.00 (1) Zr-clad Assemblies 16 52 4.00 (2) Reinsert from EOC13 discharge 17 56 4.00 l

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1 FIGURE 3-2.HRODAH NECK PLANT BOC15 BURNUP DISTRIBUTION IN MWD /MTU B 7 6 5 4 3 2 1 H 122s1 19080 10800 7sys 10800 7sys 22006 o G 2270s 10817 tesos 230s9 902s 21570 o o F to77s 10s12 suis ases ssou test 7 o '

i' E 7sss 230ss sese 2:232 issss o o D 1077s sota ssas teuse o o C 7sss 21sse tests o o B 2201o o o o I i

R o o 3-5

Figure 3-3. Haddam Neck Plant Cycle 15 Control Rod Locations i

15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 1

J R

A P 0 A A D N B C B C B H D D L C D A D C K A A J A A A H B A B A A A G C D A D C F D D E B C B C B D D A A D C A 0 A

X Rod group designation Group No. of rods Functions B 8 Control A 17 Control D 12 Safety C 8 Safety Total 45 3-6

4. Fuel System Design The Cycle 15 core consists of the fuel assemblies of batches 15A, ISB, and 16, the fresh assemblies of batch 17 and one reinserted once-burned assembly from batch 15C. Batch ISB consists of four lead tes4 assemblies (LTAs) with Zircaloy-4 clad fuel rods as opposed to the standard 304 SS clad design. The pertinent fuel parameters for all five batches are listed in Table 4-1. All fuel 4 assemblies are identical in concept and are mechanically interchangeable.

The upper and lower spacer grids were moved slightly on the Zircaloy-4 clad LTAs to accommodate the shorter Zircaloy-4 clad fuel rods. The shift in grid position does not adversely affect grid-to-grid matchup between adjacent fuel assemblies.

The fuel rods for batches 15A, 15C, 16, and 17 are 304 SS clad fuel rods of identical design. The stack length for batches 15A, ISC, 16, and 17 is 120.5 inches. All of the 304 SS clad fuel assemblies have the same uranium loading and enrichment.

The Zircaloy-4 clad fuel rods of batch ISB have a shorter length to allow for irradiation growth of the cladding. The batch ISB fuel rods also have a shorter fuel stack, thicker cladding, smaller diameter fuel pellet, and a higher prepressure in order to give equivalent performance to the 304 SS clad fuel rods. The advantage of the batch ISB fuel rods is in the lower uranium loading and 1

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enrichment that is required, which results in less overall fuel cost. The mechanical evaluation of the fuel rods is discussed below.

Cladding Collapse Batch 15A fuel rods are the most limiting in terms of creep collapse due to their having the highest previous incore exposure time and burnup. The power histories of the batch 15A fuel assemblies and the other 304 SS clad fuel assemblies were analyzed to determine a worst case enveloping power history for creep collapse. This worst case power history was used to analyze a fuel rod operating under conservative conditions for creep collapse. In addition, the other 304 SS clad fuel assemblies (batches 15C, 16, and 17) were analyzed under the same conditions for creep collapse. A similar analysis was done for the Zircaloy-4 clad LTAs, batch ISB, using a worst case power history and conservative conditions in the creep collapse analysis. For all five fuel batches the creep collapse -

analyses predict collapse times and burnups exceeding the maximum expected residence time and burnup of the fuel-I batches incore. The results are shown on Table 4-1.

4-2 l

l Cladding Stress The Haddam Neck fuel rods, both the 304 SS cladding and the Zircaloy-4 cladding designs, were analyzed by conservative stress analyses following ASME guidelines for pressure vessels. For design evaluation, the primary membrane stress intensity and any single stress must be less than two-thirds of the minimum specified unirradiated yield strength of the cladding. In all cases the margin is in excess of 13.8 percent.

Cladding Strain The fuel design criteria specify a limit of one percent o'n cladding plastic tensile circumferential strain. The pellet is designed to assure that cladding plastic strain is less than one percent (1%) at design local pellet burnup and heat generation rate. The design burnup and heat generation rate are higher than the worst case values any of the Haddam Neck fuel batches are expected to experience in Cycle 15. The strain analyses are based on the upper tolerance values for the fuel pellet diameter and density, and on the lower tolerance value for the cladding ID.

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1 Cladding Fatigue Fatigue analyses were performed using conservative conditions to find the cumulative fatigue usage factor. The fatigue usage factor for the Haddam Neck fuel rods was calculated following the ASME Pressure Vessel design code and compared to the maximum allowed factor of 0.9. The cumulative fatigue factor was found to be 0.2 for the 304 SS clad fuel, and 0.4 for the Zircaloy-4 clad fuel.

l All fuel in Cycle 15 is thermally similar. The mechanical design of the stainless steel clad batches 15A, 15C, 16, and 17 is identical.

l The design of the batch ISB Zircaloy-clad assemblies is such that l

l the thermal performance of this fuel is less limiting than the design used in the remainder of the core.

l Analyses for all fuel batches were performed with the TACO 2(2) code, using the analysis methodology consistent with Reference 3. The pin pressure for all fuel batches for Cycle 15 will remain below nominal system pressure.

The batch 17 fuel assemblies are identical in design to the other 304 SS clad fuel assemblies in the core during Cycle 15. This design has performed well in previous cycles with no adverse materials effects. Thus, all possible fuel-cladding-assembly-coolant material interactions have been proven to have no adverse effects on fuel performance when operating under the conditions expected in Cycle 15.

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_ _ _ _ _ _ _ _ _ _ _ . _ __a

Table 4-1. Nominal Fuel Design Parameters -

Batch ISA Batch ISB Batch ISC( ) Batch 16 Batch 17 Manufacturer B&W B&W B&W B&W B&W No. of assemblies 44 4 1 52 56 Previous irradiation, cycles 2 2 1 1 0 Initial fuel enrichment 4.0 3.41 4.0 4.0 4.0 1.'itial minimum fuel density,

% tueor. 94.9 94.9 94.9 94.9 94.9 Fuel pellet nominal diameter, in. 0.3825 0.3610 0.3825 0.3825 .3825 Active fuel stack nominal length, in. 120.5 119.0 120.5 120.5 120.5 20,800 21,200 12,300 8,600

{ Batch burnup, BOC, MWD /MTU 0 Initial gas pressure, psia 54.7 279.7 54.7 54.7 54.7 Gas composition, min % lie 95.0 95.0 95.0 95.0 95.0 Cladding material 304SS Zirc-4 304SS 304SS 304SS Nominal cladding thickness, in. 0.0165 0.0270 0.0165 0.0165 0.0165 Assembled fuel rod length, in. 126.68 126.125 126.68 126.68 126.68 Cladding collapse time, EFFH(* 31,500 >35,000 31,500 31,500 31,500 Design residence time EOC 15, EFPH(* 28,190 28,190 19,646 18,878 10,334

(* EFPH denotes ef fective full power hours.

(b) Reinsert from EOC13 discharge.

5. Nuclear Design The physics methodology used to support the Haddam Neck Plant for the Cycle 15 reload design is documented in Reference 4. Fifty-six (56) fresh assemblies were required as feed for Cycle 15. Due to a change in the operating conditions (vessel average temperature) for Cycle 15, the 4

physics models were updated to incorporate the new, higher vessel average temperature. This increase had a very small impact upon the physics

! characteristics of the core. The effect of new rod cluster control assemblies on core neutronics and reactor safety has been evaluated and determined to be negligible.

Table 5-1 provides a summary of changes in the Cycle 15 kinetics character-istics compared with the current limits based on the reference safety analysis (5) . Since Cycle 15 is the first application of the non-LOCA transient design basis, the current and Cycle 15 limits are equivalent.

Table 5-2 shows the shutdown margin calculations for 4-loop and 3-loop operation at beginning and end of cycle conditions. As shown in Table 5-2, shutdown margins (revised Technical Specification limits) of 1800 pcm and 2600 pcm were used for 4-loop and 3-loop operation, respectively, in operational Modes 1 and 2. The impact of these proposed required shutdown margins were verified by the steam line break accident analysis documented in Reference 5.

5-1

The proposed Technical Specification values for the maximum nuclear enthalpyrisehotchannelfactors,Fh,of 1.60[1 + 0.3*(1-P)] for 4-loop operation with 0 $ P $ 1.0 of rated power, and 1.64[1 + 0.3*(0.65 - P)] for 3-loop operation with 0 $ P $ 0.65 of rated power were used in all design calculations in support of the proposed Technical Specification values. Westinghouse Relaxed Axial Offset Control (RAOC) methodology (0) was implemented for Cycle 15 for the Haddam Neck Plant.

The RAOC results support the proposed Technical Specification axial offset limits. For both 4-loop and 3-loop operation, the applicability of the proposed Technical Specification axial offset limits has been divided into two burnup ranges instead of the previous three burnup ranges, namely, 0-250 EFPD and 250 EFPD-E0C by eliminating the 0-125 EFPD burnup range and extending the 125-250 EFPD range to cover 0-250 EFPD.

The more limiting liner heat generation rate of 14.3 kW/ft (instead of the 14.5 kW/ft) was conservatively used for the entire 0-250 EFPD burnup range in the RAOC analysis. Figure 5-1 illustrates the Cycle 15 relative power distribution at 150 MWD /MTU, hot full power, all rods out atd equilibrium xenon conditions. Power distribution and burnup data presented are based upon an assumed end of Cycle 14 burnup of 9500 MWD /MTU. .

I l

1 5-2 I l

l Table 5-1. Haddam Neck Plant Physics Parameters Current Limit Cycle 15 Cycle Design Length, (EFPD) -

431 Cycle Design Burnup, (MWD /MTU) -

12200 Average Core Burnup, E0C, (MWD /MTU) -

21486 Design Initial Core Loading, (MTU) -

64.52 Most Positive Moderator Temperature +0.0 at HFP +0.0 at HFP Coefficient, (pcm/*F) +5.0 at 65% +5.0 at 65%

power and HZP power and HZP  ;

Most Negative Moderator Temperature Coefficient, (pcm/*F) -29.0 at or below HFP -29.0 at or below HFP Doppler Temperature Coefficient, at HFP, (pcm/*F) -1.07 to -1.64 -1.07 to -1.64 Delayed Neutron Fraction, ggf, (%) 0.47 to 0.67 0.47 to 0.67 Prompt Neutron Lifetime, (psec) 20 to 10 20 to 10 Maximum Differential Rod worth at Suberitical, (pera/in) 135 135  ;

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Table 5-2. Haddam Neck Plant Cycle 15 Shutdown Margin Calculations with Maximum Stuck Rod From HFP, From 65% Power, 4 loop, 3 Imop, pcm pcm Available Rod Worth BOL EOL BOL EOL Total rod worth less max. stuck rod, HZP 5700 5600 5700 5600 (1)Less 10% uncertainty 5130 5040 5130 5040 Required Rod Worth Doppler Defect 1010 960 670 640 Moderator Defect 260 770 10 20 Rod Insertion Allowance 640 530 660 990 Flux Redistribution 220 680 170 480 Void Effect 50 50 50 50 (2) Total Rod Worth Required 2180 2990 1560 2180 Shutdown Margin (1)-(2) 2950 2050 3570 2860 Required Shutdown Margin 1800 1800 2600 2600 5-4 l

l l

i i

i l

FIGURE 5-1.HRDDRM NECK PLRNT CYCLE 15 RELATIVE POWER DISTRIBUTION RT 15D MHD/MTU,HFP.RRO ,

I 8 7 6 5 LA 3 2 1 I

! H 1.oss 1.032 1.201 1.254 1.222 1.127 o. sus o.sut l

l G o.981 1.14o 1.172 1.os8 1.204 o.s70 1.01s o.sto 1

l F .nes t.tss 1.oss t.aas 1.ast o.ss7 o.egs l  :

E 1. ass 1.oss 1.234 o.sas 1.027 t.oss o. sos

'1 i

! D 1.aas t.ana 1.aqa t.oss 3.13o o.7sa i C 1.iss o.s77 o.s74 1.osa o.7sv i

l B o.847 1.022 o.852 o. sos i

i

! A o. sus o. sus i

l i

l i

5-5 4

. /

6. Thermal-Hydraulic Design All fuel assemblies in the Cycle 15 core are hydraulically similar.

The reactor core safety limit curves in the proposed Technical Specifications have been developed using the VIPRE methodology that has been previously reviewed and approved.(7) These safety limit curvesareconsistentwiththeenthalpyrisehotchannelfactorEfg and RCS flow rate provided in Tablo 6-1 and the proposed Technical Specifications.

Results of the steady state, thermal-hydraulic analyses for Cycle 15 operation are given in Table 6-1. Since Cycle 15 is the first application of the approved VIPRE methodology, the Design and Cycle 15 values are equivalent. Two VIPRE analyses were performed to evaluate the minimum steady state DNBR and maximum UO2 temperature.

The case to determine the minimum DNBR assumes a limiting top peak

axial power shape consistent with the positive axial offset limits N

i and the limiting F AH all wed by the proposed Technical Specifications.

The case to determine the maximum UO2 temperature assumes a limiting i

total peak consistent with the axial offset limits and the limiting i

Linear Heat Generation Rate allowed by Technical Specifications.

Both cases assume the limiting RCS flow rate, pressurizer pressure and core inlet coolant temperature.

4 6-1 i

l

)

l Table 6-1. Cycle 15 Thermal-Hydraulic Data Design Cycle 15 Hot Channel Factor Nuclear Enthalpy Rise (FAH) 1.60 1.60 Coolant System Pressure 1 Nominal, psig 2,000 2,000  ;

Minimum steady-state, psig 1,960 1,960 (includes instrument error and j dead band)

Coolant Flow Rate Core flow rate *, gpm 233,870 233,870 Total core flow area, ft2 42.24 42.24 Average velocity along fuel rods, fps 13.01 13.01 Coolant Temperature, 'F 4 Nominal core inlet 538 538 Maximum core inlet 544.6 544.6 (includes instrument error and dead band)

Maximum rise in vesset 51.89 51.89 Maximum rise in core 54.16 54.16 Maximum average in vessel 570.55 570.55 Maximum average in core 571.68 571.68 Maximum vessel outlet 596.49 596.49 Coolant Enthalpy Hot channel outlet, BTU /lbm 668.53 668.53 Saturated enthalpy at minimum 667.41 667.41 pressure, BTU /lbm Heat Transfer Heat generated in fuel, % 97.4 97.4 Maximum UO2 temperature, 'F 4224.6 4224.6 Active heat transfer surface area, ft 2 35,540 35,540 Average heat flux, BTU /h-ft 2 174,066 174,066 I

Maximum heat flux, BTU /h-ft 2 476,430 476,430 Maximum thermal output, kW/ft 15.5 15.5 Maximum cladding surface temperature, *F 702.9 702.9 Minimum DNBR 2.04 2.04

• Includes measurement uncertainty and assumes 4.5% bypass flow fraction.

6-2

7. Accident and Transient Analysis The non-LOCA design basis transients for the Haddam Neck Plant have been reanalyzed (5) by Northeast Utilities Service Company (NUSCO) on behalf of Connecticut Yankee Atomic Power Company (CYAPCO). This revised design basis will replace the original FDSA analyses and subsequent analyses that have been performed by Westinghouse, Yankee Atomic Electric Company, and NUSCO during previous operating cycles.

The reanalysis has been submitted to consolidate the non-LOCA design basis and resolve issues related to the original design basis identified during the course of the Systematic Evaluation Program and by the Connecticut Yankee Plant Design Change Task Group.

Several of the accidents from this reanalysis are required to support the proposed revision to Technical Specifications for the Cycle 15 Reload. The reanalysis of these accidents will replace the current design basis for the boron dilution, rod withdrawal, dropped rod, and loss of flow accidents. NRC approval of these accidents is expected prior to the startup of Cycle 15 in mid-September 1987.

The results of the steam line break accident reanalysis support the proposed Technical Specification requirement for shutdown margin in operational Modes 1, 2, and 3. However, this accident is limiting at the end of cycle and approval of this accident may conservatively be deferred until after startup. The existing design basis for the steam line break conservatively shows acceptable results through mid-cycle operation.

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The remaining accidents in the non-LOCA transient reanalysis will replace the current design basis when final NRC approval is received.

This approval is expected after startup.

l The effects of the reload design on the design basis non-LOCA transients (5) were examined. In all cases, it was found that the effects were accommodated within the conservatism of the initial

! assumptions used in the non-LOCA transient reanalysis. Therefore, the conclusions presented in the non-LOCA transient reanalysis remain valid.

A core reload can typically affect accident analysis input parameters in the following cases: core kinetics characteristics, shutdown margin, control rod worths, and core peaking factors. Cycle 15 parameters in each of these areas were examined as discussed below to ascertain whether new accident analyses were required.

Kinetics Parameters A comparison of Cycle 15 kinetics parameters with current limits established by the non-LOCA transient reanalysis is presented in Table 5-1. All parameters are within current limits, therefore, no reanalysis is required, i

7-2 l l t

I._________________________

. - i Shutdown Margin Changes in minimum shutdown margin requirements may impact i the safety analyses, particularly the steam line break and boron dilution accidents. The Cycle 15 Technical Specifications proposed for shutdown margin have been revised to be consistent with the revised design basis for the non-LOCA transients.

i Control Rod Worths i

Changes in control rod worths mey affect shutdown margin.

Table 5-2 shows that the Cycle 15 shutdown margin require-ments are met.

! Core Peaking Factors All core peaking factors for Cycle 15 were within the new design basis limits.

The effects of the reload design on the current design basis non-LOCA transients, except for the boron dilution, rod withdrawal, dropped 1  :

rod, and loss of flow accidents, were also evaluated. Each of these )

I accidents is discussed below:

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Isolated Loop Startup If a reactor loop is isolated from the remainder of the reactor coolant system and subsequently brought back into operation without first matching its boron concentration and temperature to those of the system, an increase in core reactivity and power may occur. To prevent this and ensure the safe startup of an isolated loop, procedures (outlined in Section 10.2.2.1 of the FDSA) have been established. In addition, a temperature-valve interlock prevents opening of the cold leg valve if the temperature difference between the hottest operatirg loop cold leg and the isolated cold leg exceeds 20*F.

Although it is improbable that the operator would neglect to follow any one of the established procedures in starting up an isolated loop, the FDSA analysis assumed violation of two such procedures. The maximum reactivity insertion produced by such a violation results from starting the isolated loop pump at power after neglecting to match cold l

leg temperatures. This case was identified in the FDSA as the most severe transient from the standpoint of DNB.

To maximize reactivity addition, the FDSA analysis assumed the most negative moderator temperature coefficient expected (-35 pcm/*F). To minimize negative reactivity 7-4

?

l feedback, the least negative Doppler coefficient of -0.5 pcm/*F was assumed. The limiting most negative moderator j j temperature coefficient and least negative Dcppler coefficient i

i for Cycle 15 are -29.0 pcm/*F and -1.07 pcm/*F, respectively.

! Since the limiting Cycle 15 moderator temperature coefficient is less negative and the Doppler coefficient more negative than the corresponding values assumed in the FDSA analysis, the consequences of an isolated loop startup during

! Cycle 15 would be less severe than previously reported.

.l i Excess Feedwater i

i The excess feedwater incident is a result of an abnormal sustained increase in feedwater flow to one or more steam generators in excess of that needed to maintain the steam

generator water level. The excess feedwater would absorb

extra heat from the reactor coolant loop of the affected i

steam generator, resulting in a cold leg temperature

{ reduction and a subsequent increase in reactivity that i

i would lead to a power excursion. The power excursion is maximized with the most negative moderator temperature coefficient (ensures maximum reactivity addition from the reduced coolant inlet temperature) and the least negative i Doppler coefficient (ensures minimum reactivity feedback).

i The values assumed in the FDSA analysis are -35 pcm/*F for l the moderator and -0.5 pcm/*F for the Doppler coefficients 7-5 i

of reactivity. The corresponding values predicted for I

Cycle 15 are -29.0 pcm/*F and -1.07 pcm/*F, respectively.

The limiting values of the. moderator temperature and Doppler coefficient for Cycle 15 are within the bounds  ;

assumed in the previous safety analysis. Therefore, the l t

FDSA analysis covers the mo~sp adverse excess feedwater incident that could be postulated to occur during Cycle 15.

The bounding analysis of the FSDA indicates .that one or-more of the high steam generator water level alarms will alert the operator to trip the reactor (and turbine),

thereby preventing damage to the turbine due to steam generator overflow. If the operator takes no action, the t

core inlet temperature will decrease and heat flux will increase up to the point of reactor trip. The'two effects tend to compensate each other with respect to,DNB, and there is no fuel damage.

Excessive Load Increase 1

An excessive load increase incident is defined as a rapid increase to steam generator steam flow, resulting in a l significant power mismatch between the reactor core power and the steam generator load demand. The excess load results in a decrease in reactor coolant temperature. To .

I maximize the resultant power excursion, the FDSA analysis considered the most-negative moderator coefficient of- I reactivity (-35 pcm/*F) and the least negative Doppler coefficient of reactivity (-0.5 pcm/*F). The limiting -j 7-6 l

, i**

values for Cycle 15 are -29.0 pcm/*F and -1.07 pcm/*F for the most negative moderator coefficient and the least negative Doppler coefficient, respectively. This combination of moderator and Doppler coefficients would result in a milder power excursion in the event of an excessive load increase and a less severe core transient than the case analyzed in the FDSA. Therefore, the case analyzed in the FDSA bounds the most severe excessive load increase incident that could be postulated to occur for Cycle 15.

Rod Cluster Control Assembly Ejection The control rod ejection incident is postulated to occur by the failure of a control rod drive mechanism housing, permitting a control rod to be rapidly ejected from the core. This incident represents the most rapid reactivity insertion that can be reasonably postulated.

In all except the BOC hot zero power case, the Cycle 15 parameters are bounded by the values in the FDSA analysis.

The results reported in Section 10.2.7 of the FDSA for the bounded cases conservatively predict the margins to fuel damage for the most severe ro'd ejection incidents that could be postulated.

For the Cycle 15 BOC, zero power case, the total F q

predicted for Cycle 15 (7.8) is 56% greater than that 7-7 J

assumed in the FDSA analysis (5.00), even though the limiting Cycle 15 ejected rod worth (500 pcm) is sub-stantially lower than that used in the FDSA analysis (830 pcm). The effect of the higher F was conservatively evaluated by increasing the peak fuel centerline temperature quoted in the FDSA (2930F) for this case by an amount predicted by the change in fuel centerline temperature for a 56% change in linear heat generation rate from Reference 8.

The results of this calculation indicate that the peak fuel centerline temperature will remain less than 4200*F.

These temperatures are considerably lower than the 4700 F UO melting temperature; therefore, safety limits will be 2

neither approached nor exceeded.

Steam Line Break The limiting steam line break incident identified in the FDSA is a circumferential double-ended break of a 24-inch steam line from the steam generators upstream of the steam line isolation valves. This break was analyzed in the FDSA from both full- and zero power conditions.

For both cases the FDSA analysis assumed a most negative moderator coefficient of -35 pcm/*F and a Doppler coefficient l

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of -0.5 pcm/*F. This combination of moderator and Doppler coefficients results in the maximum peak power before reactor trip and, thus, the minimum DNBR. The corresponding Cycle 15 limiting values for moderator and Doppler coefficients are -29.0 pcm/*F and -1.07 pcm/*F, respectively.

These values are bounded by the FDSA analysis. The shutdown margin assumed in the FDSA was 3400 pcm. The minimum shutdown margin for Cycle 15 is 2230 pcm.

However, since the concern is DNB at the time of reactor trip, the lower shutdown margin does not have an impact on the minimum DNBR. Therefore, the DNB margin reported in the FDSA represents a conservative estimate of the minimum DNBR for the postulated Cycle 15 steam line break incident. .

The maximum loss of shutdown margin was reevaluated in Reference 9 to support a plant modification that imple-mented automatic initiation of auxiliary feedwater. The Cycle 15 design was evaluated to assure that sufficient shutdown margin is provided by the power dependeni control rod insertion limit to prevent a return to criticality.

s The required shutdown margin based on cooldown deficits and available shutdown margin based on the Cycle 15 design was conservatively evaluated for the limiting hot full and hot zero power cases for four and three loop operation.

The end of the Cycle 15 shutdown margin and a conservative 7-9 a

8 estimate of the middle of cycle cooldown deficit demonstrate that the Cycle 15 design provides sufficient margin to prevent a return to criticality through mid-cycle operation.

Steam Generator Tube Rupture The integrity of the steam generator is significant from the point of view of radiological safety. The radiological consequences of this incident are independent of core loading and remain acceptable.

Loss of Load A loss-of-load incident is a large, rapid reduction in generator load causing a similar reduction in the heat extracted from the reactor coolant system. Normally,.a large loss will cause a turbine trip, either by a signal from the generator or switchgear, or from the turbine overspeed trip signal. If the 'urbine t load is above the permissible low-power level, the turbine trip will cause an immediate reactor trip, thus preventing any significant pressure or temperature rise in the reactor coolant system. However, if the turbine control valves respond i l

quickly enough to prevent the overspeed trip of the turbine, the load could be reduced to the station service level without a reactor trip. This is the largest credible 1

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load loss and is the basis for the loss-of-load analysis presented in the FDSA. The FDSA analysis considered a range of values for both moderator temperature and Doppler coefficients in determining DNB margin and reactor vessel integrity. As a result, it was concluded that the most adverse loss-of-load transient is one having the most positive moderator temperature coefficient coupled with the least negative Doppler coefficient. Values assumed in the FDSA were +10 pcm/*F and -0.5 pcm/*F for the most positive moderator coefficient and the least negative l

Doppler coefficient, respectively. The limiting hot full power values for Cycle 15 are 0.0 pcm/*F and -1.07 pcm/ F.

l The comparison indicates that the parameters assumed in the FDSA analysis bound the limiting values for the loss of load presents no hazard to the integrity of the core, the reactor coolant system, or the turbine cycle.

Pressure-relieving devices incorporated in the system are ample to limit the maximum pressure to acceptable values.

Loss of Normal Feedwater Flow This incident is a reduction in feedwater flow to the steam generators when operating at power without an equivalent reduction in steam flow, thus reducing the water inventory in the affected steam generators. Since the complete loss of feedwater flow requires the most rapid response from the reactor protection system, it 7-11

forms the basis for the loss of feedwater flow incident analysis reported in the FDSA. The maximum core power level before trip and the minimum DNBR occur with the most positive value of the moderator temperature coefficient, coupled with the least negative Doppler coefficient: +10 pcm/*F and -0.5 pcm/*F. The limiting positive moderator temperature coefficient and the limiting negative Doppler coefficient for Cycle 15 at hot full power are 0.0 pcm/ F and -1.07 pcm/*F, respectively. The values assumed in the FDSA analysis adequately bound the values predicted for Cycle 15; thus, the results reported in Section 10.3.6 of the FDSA are applicable to the Cycle 15 reload.

Fuel Handling Incident The fuel handling incident considers the possibility of dropping a fuel assembly during fuel handling operations.

The concern over this incident are radiation exposure and accidental criticality. These concerns are independent of core loading.

Waste Gas Incident The waste gas incident is defined as an unexpected and uncontrolled release to the atmosphere of the radioactive I l

xenon and krypton fission gases stored in the waste gas 7-12 w____

,.o decay tank. The consequences of this incident are independent of core loading; therefore, the results reported in Reference 10 are applicable for any reload.

Hypothetical Accident Regardless of the ability of the safety injection and core deluge systems to prevent major fission product releases to the containment, the safety of nuclear plants has historically been evaluated based on a " hypothetical accident." The hypothetical accident involves a gross release of fission products from the fuel to the containment.

The consequences of this incident are independent of core loading.

The small break LOCA design basis has also been reanalyzed by NUSCO on behalf of CYAPC0(II) . This revised design basis will replace the original Westinghouse design basis when NRC final approval is-received. This approval is expected prior to the startup of Cycle 15.

The large break LOCA design basis for Cycle 15 continues to be the reviewed and approved Westinghouse methodology that demonstrates compliance with the Interim Acceptance Criteria.(12) 7-13

.a

f.

The LOCA design basis is affected by the proposed reduction in the RCS flow rate provided in Technical Specifications. The current and proposed small break LOCA results are not significantly affected by the reduction in the flow rate. No fuel cycle design limits are established by the small break LOCA. The large break LOCA for three and four loop operation establishes the limiting Linear Heat Generation Rates (LHGRs) for Technical Specifications. The impact of the RCS flow rate on LHGR limits was reviewed and approved for the Cycle 14 reload.(13) The four loop LHGR limits for the current RCS flow rate Technical Specification limit of 257,000 gpm are 14.45, 14.75, and 17.0 kW/ft for the 0-125, 125-250 and 250 EFPD-E0C burnup windows respectively. The proposed RCS flow rate reduction to 246,000 gpm yields LHGR limits of 14.35, 14.6, and 16.9 kW/ft, respectively.

The current LHGR Technical Specification limits of 14.3, 14.5, and 15.5 kW/ft, respectively, remain bounding. The current LHGR limits for three loop operation also remain bounding. The four Zircaloy clad Lead Test Assemblies (LTAs) will undergo their third and final cycle of irradiation during Cycle 15. Reference 14 provides an Exemption from the requirements of 10 CFR 50.46 and Appendix K on the condition that the calculated steady-state peak LHGR for the LTAs remains less than 11 kW/ft. The Cycle 15 calculated steady-state LHGRs are less than 10 kW/ft.

l l

l 7-14

- ___ ____-__ - _ ____-_______ _____-__-_____-__ _ ____-_-____ _ - _ _ _ _ a

7 s

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8. Startup Program -- Physics Testing The planned startup tests associated with core performance are outlined below. These tests verify that core performance is within the assumptions of the safety analysis and provide the necessary data for continued safe plant operation.

Pre-Critical Tests

1. Hot control rod drop-time testing.

2. Control rod coupling verification.

Zero-Power Tests

1. Critical boron concentration.

2. Temperature reactivity coefficient.

a. All rods out.

b. Banks B, A, and D in.

3. Control rod group worth for banks B, A, and D.

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.. i-Power Tests

1. Core power distribution mapping <,80 and 100 percent. full power, normal bank configuration.

2. Excore/incore correlation verification.

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9. Reft ences

1. Facility Description and Safety Analysis (FDSA), Haddam Neck Plant, NY0-3250-5, Connecticut Yankee Atomic Power Company.

2. Y. H. Hsii, et al, TAC 02 -- Fuel Pin Performance Analysis, BAW-10141 PA, Babcock & Wilcox, Lynchburg, Virginia, June 1983.

3. J. H. Taylor (B&W) to J. S. Berggren (NRC), Letter, "B&W's Responses to TACO 2 Questions," April 8, 1982.

4. J. F. Opeka to C. I. Grimes, Haddam Neck Plant, Millstone Nuclear

, Power Station, Unit Nos. 2 and 3 - Physics Methodology for PWR Reload Design, September 12, 1986, a

5. Haddam Neck Plant Reanalysis of Non-LOCA Design Basis Accidents, i

Docket No. 50-213, Northeast Utilities Service Company, June 30, i

i i 1986, and revisions dated March 10, 1987 and May 7, 1987.

I 6. R. W. Miller, et al, " Relaxation of Constant Axial Of fset Control,"

WCAP-10216-P-A, June 1983.

)

7. F. M. Akstulewicz to J. F. Opeka, "NUSCO Thermal Hydraulic Model t Qualification, Volume II (VIPRE)," Topical Report NUSCO 140-2, October 16, 1986.

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8. Effects of Fuel Densification on the Connecticut Yankee Reactor, WCAP-8213, Westinghouse Electric Company, Pittburgh, Pennsylvania, October 1973.

9. W. G. Council to Director of Nuclear Reactor Regulations, Automatic Initiation of Auxiliary Feedwater, January 30, 1980.

10.

D. C. Switzer, CYAPCO, to A. Giambusso, USAEC, Letter December 5, 1972.

11. W. G. Council to J. A. Zwolinski, lladdam Neck Plant - Small Break LOCA Topical Report Tt!I Action Plan Items II.K.3.5, II.K.3.30, II.K.3.31, December 20, 1984.

12. R. A. Purple to D. C. Switzer, Supplement to the Safety Evaluation by the Directorate of Licensing, U.S. Atomic Energy Commission, Docket No. 50-213, Connecticut Yankee Atomic Power Company, lladdam Neck Plant, December 27, 1974.

13. F. ?!. Akstulewicz to J. F. Opeka, ifaddam Neck Plant Cycle 14 Reload Technical Specifications, April 14, 1986.

14. W. A. Paulson to W. G. Council, lladdam Neck Plant - Exemption from 10 CFR 50.46 and Appendix K, October 2, 1984.

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