ML20205K355

From kanterella
Jump to navigation Jump to search
Technical Rept Supporting Cycle 14 Operation
ML20205K355
Person / Time
Site: Haddam Neck File:Connecticut Yankee Atomic Power Co icon.png
Issue date: 02/28/1986
From:
BABCOCK & WILCOX CO.
To:
Shared Package
ML20205K348 List:
References
BAW-1878, NUDOCS 8602270321
Download: ML20205K355 (52)
v • d • e


Text

t l' 4 '

~. BAW-1878 I February 1986 7

9 1

- CONNECTICUT YANKEE ATOMIC POWER COMPANY I Haddam Neck Plant Technical Report Supporting Cycle 14 Operation J

J -

~

] -

3 il Z

5 iI .

,U l

e Babcock &Wilcox peg 22;gg!Agg88alia a ucoermott company a p PDR M

t. s

~

BAW-1878 m

.t. February 1986 1

[ =

1-CONNECTICUT YANKEE ATOMIC POWER COMPANY Haddam Neck Plant Technical Report Supporting Cycle 14 Operation 9

l e

ij ru

.d i

.-l) .

W G

,Y L; ,

}

BABC0CK & WILCOX

}.

Nuclear Power Division F. O. Box 10935 Lynchburg, Virginia 24506-0935 a .. .

a ucoermoir como.ny C

m o

  • t
CONTENTS Page
1. INTRODUCTION AND

SUMMARY

..................... 1-1

2. OPERATING HISTORY . . . . . . . . . . . . . . . . . . . . . . . . . 2-1
3. GENERAL DESCRIPTION . . . . . . . . . . . . . . . . . . . . . . . . 3-1
4. FUEL SYSTEM DESIGN ........................ 4-1 -

~] 4.1. Fuel Assembly Mechanical Design . . . . . . . . . . . . . . . 4-1 6 4.2. Fuel Rod Design . . . . . . . . . . . . . . . . . . . . . . . 4-1

4. 2.,1. Cladding Collapse . . . . . . . . . . . . . . . . . . 4-2 q 4.2.2. Cladding Stress . . . . . . . . . . . . . . . . . . . 4-2 4.2.3. Cladding Strain . . . . . . . . . . . . . . . . . . . 4-2 4.2.4. Cladding Fatigue .................. 4-2

, 4.3. Thernal Design ....................... 4-2

\

I 5. NUCLEAR DESIGN .......................... 5-1 g 5.1. Physics Characteristics . . . . . . . . . . . . . . . . . . . 5-1 5.2. Nuclear Design Changes ................... 5-2 J

6. THERMAL-HYDRAULIC DESIGN ..................... 6-1 6.1. Design Analysis'and Safety Limits Curves .......... 6-1 6.2. Cycle 14 Thermal-Hydraulic Performance ........... 6-1
7. ACCIDENT AND TRANSIENT ANALYSIS . . . . . . . . . . . . . . . . . . 7-1 7.1. General Safety Analysis .................. 7-1 q 7.2. Control Rod Withdrawal Incident .............. 7-2 7.3. Isolated Loop Startup Incident . . . . . . . . . . . . . . . 7-3 7.4. Baron Dilution Incident .................. 7-4 7.5. Excess Feedwater Incident .................. 7-5 D; 7.6. Excessive Load Increase .................. 7-6 2 7.7. Dropped Rod Inci dent . . . . . . . . . . . . . . . . . . . . 7-6 7.8. Control Rod Ejection Incident ............... 7-7 3 7.9. Loss-of-Coolant Incident . . . . . . . . . . . . . . . . . . 7-8 3, 7.10. Loss-of-Flow Incident ................... 7-8 7.11. Steam Line Rupture Incident ................ 7-8

,, 7.12. Steam Generator Tube Rupture Incident ........... 7-9

'3 7.13. Loss-of-Load Incident ................... 7-9 "J 7.14. Loss-of-Feedwater Incident . . . . . . . . . . . . . . . . . 7-10 7.15. Fuel Handling Incident . . . . . . . . . . . . . . . . . . . 7-11
s 3 .

Embcock & Wilcox 1 - iii - -- --

., utoerman zampny

CONTENTS (Cont'd)

Page 7.16. Waste Gas Incident . . . . . . . . . . . . . . . . . . . . . 7-11 7.17. Hypothetical Accident ................... 7-11

8. TECHNICAL SPECIFICATIONS ..................... 8-1 8.1 Power Versus Offset Limits: Four-Loop Operation ...... 8-1 8.2 Power Versus Offset Limits: Three-Loop Operation ...... 8-3
9. STARTUP PROGRAM PHYSICS TESTING . . . . . . . . . . . . . . . . . . 9-1
10. REFERENCES ............................ 10-1 List of Tables Table 3-1. Cycle 13 Discharged Fuel .................... 3-2 >

3' 4-1. Nominal Fuel Design Parameters ................. 4-4 5-1. Physics Parameters -- Cycles 13 and 14 ............. 5-3

5. ? . Shutdown Ma~rgin Calculations-Haddam Neck Plant, Cycle 14 . . . . 5-5 6-1. Cycl e 14 The rmal -Hyd ra ul i c Da ta . . . . . . . . . . . . . . . . . 6-3 6-2. Cycle 14 Parameters -- Hot Conditions . . . . . . . . . . . . . . 6-6 7-1. Dropped-Rod Incident Comparison of Parameters . . . . . . . . . . 7-12 .

7-2. Steam Line Rupture Incident . . . . . . . . . . . . . . . . . . . 7-12 i b

List of Figures Figure  ;]

3-1. Haddam Neck Plant Cycle 14 Core Loading Diagram . . . . . . . . . 3-3 3-2. Haddam Neck Plant B0C 14 Burnup Distribution .......... 3-4 -,

Haddam Neck Plant Cycle 14 Control Rod Locations ........ 3-5 ^!

3-3. U 5-1. BOC 14 Two-Dimensional Relative Power Distribution (Full Power, No Control Rods Inserted, Equilibrium . . .

5-6 Xenon). .............................

Power Vs Offset Limits, Four-Loop Operation 0-125 EFPD . . . . . 8-5 a 8-1.

8-2. Power Vs Offset Limits, Four-Loop Operation, 125-250 EFPD . . . . 8-6 8-3. Power Ys Offset Limits, Four-Loop Operation, 250 EFPD-EOC . . . . 8-7 g 8-4. Power Vs Offset Limits, Three-Loop Operation 0-125 EFPD . . . . 8-8 74, '

8-5. Power Vs Offset Limits, Three-Loop Operation, 125-250 EPFD . . . 8-9 8-6. Power Vs Offset Limits, Three-Loop Operation, 250 EFPD-E0C . . . 8-10 7,.

I.N ta4 N &Mecens h'

- IV - a McDermott company W

l' *

1. INTRODUCTION AND

SUMMARY

The objective of this report is to support the operation of the fourteenth 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 previously supplied analyses wherever possible.

Based on the analyses performed and review of the Technical Specifications, it

.i s concluded that the Haddam Neck Plant can be operated safely at the

'~

rated power level of 1825 MWt for cycle 14.

e o

,B r

i

.a W

U E

, h __.,

. . - - . _ . . . . . . - - - - = = = - -

G- e +

.3 3

et D ESIDCOCk & MICOE 3 1-1 ,uc m , moi,co,n,.ny

7 '

l

.)

2. OPERATING HISTORY Initial criticality for cycle 13 occurred on November 3, 1984. The plant phased online November 9,1984, and reached 100% power on November 23, 1984.

Cycle 13 operation is scheduled for completion in January 1986. No operating anomalies occurred during the thirteenth cycle that would adversely affect fuel performance during cycle 14. .

The nominal 360 effective full-power day (EFPD) cycle 14 is scheduled to begin in March of 1986.

.a s

e

+

r Y

%e n'

f,b d

r ,

b 5 - ._

.m 5.:

t &

a sabcock awiscom 2-1

.9[1 -- -

. ._ . m,=,, my .- _

, l' v .

~I-i 1

3. GENERAL DESCRIPTION l

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 14 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 s, teel (type 304) with an outside diameter (0D) of 0.422 inch and a wall thickness of 0.0165 inch. Four batch ISB Zircaloy lead test assemblies (LTAs) have fuel pin cladding made of Zircaloy with an 00 of 0.422 inch and a wall thickness of 0.027 inch. The new fuel consists of bevelled, dished-end cylindrical pellets of uranium dioxide. Batch 16 fuel pellets are 0.3825 inch in diameter and 1

O.458. inch in length. The 52 batch 16 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 (see Table 6-2). The minimum batch theoretical fuel density is 94.9% for batches 9, 14, 15A, ISB, and 16 (cycle 14 batches).

i* Figure 3-1 is the core loading diagram for cycle 14 of the Haddam Neck Plant.

. The nominal initial enrichment for batches 9, 14, 15A, and 16 is 4.00 wt 4i  % uranium-235, and 3.41 wt % for batch ISB. ,

7 The 53 fuel assemblies that will be discharged at the end of cycle 13 are i

from batches 9 and 13 (see Table 3-1). ,

} The 52 batch 14, 48 batch 15A and 4 batch 158 assemblies will be shuffled to ld new locations at the beginning of cycle 14. Twice-burned assembly J36 (batch

'p 9), discharged at the end of cycle 8, will be reinserted as the center C assembly in cycle 14. The 52 fresh batch 16 assemblies will occupy the outer row of assemblies (see Figure 3-1). Figure 3-2 is an eighth-core map showing the assembly burnup distribution at the beginning of cycle 14.

)

a I

M 2

M W SabCOCk& MhCOE Ij, 3-1,_,_ _ ,, , ,, ,,,,,, , uc o,, ,, ,,,,,ny ,,,,_ ,

i

l Reactivity control is supplied by 45 full-length Ag-In-Cd control rods and by soluble boron shim. The cycle 14 locations of the 45 control rods and the group designations are indicated in Figure 3-3. The core locations of the total pattern (45 control rods) for cycle 14 are identical to those of the reference cycle indicated in the Haddam Neck Plant Cycle 13 Reload Report.2 Table 3-1. Cycle 13 Discharged Fuel No. of Batch assemblies Cycles burned 9 1 3 (cycle 13 center assembly) 13 52 3 Total discharged 53 I

a d

'l e.

Y IE LF p

vi S.:.

~~ ~

Babcock &WI8cos at 3-2 , uormate com,,ny [.]

1, . .

7 '

Figure 3-1. Haddam Neck Plant Cycle 14 Core Loading Diagram 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 16 16 16 P2 16 16 16 16 16 16 P33 P36 ZR51 P04 P42 "

16 16 16 16 14 14 ISB* 14 14 R25 P01 R24 R07 P44 R42 R40 P16 R08 M 16 16 15A 14 15A 15A 14 15A 15A 14 ISA P15 R03 R48 P26 R46 P28 R11 R31 P06 16 16 16 16 14 15A ISA 14 15A 14 15A 15A 14 P47 R38 R02 P19 RIO P02 R16 P10 R32 R34 P08 l 16 16 I 14 15A 15A 14 15A 14 15A 14 1hA 15A 14 P49 ROI P09 R39 P30 R18 P41 R43 P05 J 16 16 16 16 14 14 15A 14 15A 14 15A 14 1EA 14 1RA P12 ZR50 P34 R15 P11 R28 J36 R29 P50 R22 P40 ZR52 P43 H 16 16

.. 14 ISB* 14 15A 14 15A 9 15A 14 15A 14 ISB* 14 P03 R14 P35 R12 P27 R30 N21 R19 P45 R20 P31 G 16 16 16 16 14 15A 14 15A 14 15A 14 15A 14 15A 14 P17 R13 R47 N22 R35 P38 R21 P29 R23 R33 P07 16 15A 14 16 14 15A 15A 14 ISA 14 15A 14 15A P32 R37 R06 P25 R44 P39 R27 R04 P23 16 16 16 16 14 15A ISA 14 15A 14 15A 15A 14

+

R41 P46 R45 R26 P20 R17 R36 P14 R09 D 16 16 15A 14 15A 1EA 14 154 iga 14 154 P13 P24 2R49 P37 P18

-' 16 16 16 16 14 14 158* 14 14

'.} 16 16 16 14 16 16 16 o

XXX Fuel Assembly Serial No. 16 16 16 A

XXX Batch No.

16 Batch 16 Fuel (Fresh) Batch No._ _

9 4.00

*Zircaloy-clad LTA. 14 4.00 0

15A 4.00 15B 3.41 t, 16 4.00 u

A P 3-3 Babcock & WHcom

-() a u<onmost comoony

l Figure 3-2. Haddam Neck Plant B0C 14 Burnup Distribution r

, 8 7 6 5 4 3 2 1 1

H 21,577 12,914 26,339 8,272 26,609 10,148 23.010 0 G 24,394 8,233 21,727 8,620 24,933 0 0 F 26,327 11,125 12,537 21,946 0 i

E 12,914 22,749 0 0 t D 8,272 0

.l

. .I C

B ~U

)

i e'! ,

C l A 7t '

gi' Assembly burnup at 0 EFPD y XX,XXX of Cycle 14, mwd /mtU g;

, yt Babcock &Wilcox  ?'

3-4 , uco,,more comp.ny #

1. .

l i Figure 3-3. Haddam Neck Plant Cycle 14 Control Rod Locations

'i Called North 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 R

A P D A A 0 N B C B C B M D D L C D A D C K l

A A J-

~.

A B A A A A H B

A A G C D A D C F 0 .

D

' B D C B C B D A A D C 3 ,

i A C

}.s A X Rod group designation Group No. of rods Function J

B 8 antrol

'[ A 17 Control '

D 12 Safety

~~

, C 1 Safety

TOTAL 45 3

3

] ,

3-5 Babcocer &WHcos 4 MCDermott Company

\' .

i

4. FUEL SYSTEM DESIGN 4.1. Fuel Assembly Mechanical Design The cycle 14 core consists of the fuel assemblies of batches 14, 15A, and ISB, the fresh assemblies of batch 16, and one reinserted twice-burned assembly from batch 9. Batch 15B consists of four lead test 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 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 rod. The shift in grid position- does not adversely affect grid-to-grid matchup between adjacent fuel assemblies.

4.2. Fuel Rod Design The fuel rods for batches 9,14,15A, and 16 are 304 SS clad fuel rods of

.a nearly identical design. The stack length for batches 14, 15A and 16 is slightly longer (120.5 inches versus 120.3 inches) than the stack length for batch 9. All of the 304 SS clad fuel assemblies have the same uranium loading and enrichment. Batches 14,15A, and 16 have reduced diameter fuel pellets (by 1 mil), and a higher internal rod pressure (54.7 psia versus 14.7 psia) compared to batch 9. These changes result in. easier fuel loading, improved performance, and more creep resistance.

]

.a The Zircaloy-4 clad fuel rods of batch ISB have a shorter length to allow for

  • irradiation gr$wth of the cladding. The batch I5B fuel rods also have a O shorter fuel stack, thicker clad, smaller diameter fuel pellet, and a higher g prepressure in order to give equivalent performance to the 304 SS clad fuel "3 rods. The advantage of the batch ISB fuel rods is in the lower uranium loading and enrichment that is required, which results in less overall fuel 3 cost. The mechanical evaluation of the fuel rods is discussed below.

1

  • Babcock & WHcom 4-1 4

y __

- a uconmoir comp ny

4.2.1. Cladding Collapse Batch 14 fuel rods are the most limiting in terms of creep collapse because they have the highest previous incore exposure time. The batch 14 assembly '

power histories were analyzed to determine the worst case power history for creep collapse. This power history was used to analyze a fuel rod operating under conservative conditions for creep collapse. The creep collapse analysis predicts a collapse time longer than the maximum expected residence time of the fuel incore. The results are shown in Table 4-1.

4.2.2. Cladding Stress The Connecticut Yankee (Haddam Neck) fuel rods were analyzed in a conservative fuel rod stress analysis. For design evaluation, the primary membrane stress must be less than two-thirds of the minimum specified unirradiated yield strength, and all stresses must be less than this value. In all cases, the margin is in excess of 19.6%.

4.2.3. Cladding Strain The fuel design criteria specify a limit of 1.0% on cladding plastic tensile circumferential strain. The pellet is designed to ensure that cladding plastic strain is less than 1% at the design local pellet burnup and heat i

generation rate. The design burnup and heat generation rate are higher than the worst case values the Haddam Neck fuel is expected to see. The strain analysis is also based on the upper tolerance values for the fuel pellet diameter and density and the lower tolerance value for the cladding ID.

'}

s 4.2.4. Cladding Fatigue

  • A fatigue analysis was performed using conservative conditions to find the cumulative fatigue usage factor. Following the ASME Code, the fatigue usage f

. factor for the Haddam Neck fuel rods was calculated and compared to the maximum allowed factor of 0.9. The cumulative fatigue usage factor was found Il to be less than 0.2 for the 304 55 clad fuel, and less than 0.4 for the

~

Zircaloy-4 clad fuel. ],

,;. 3 4.3. Thennal Design All fuel in cycle 14 is thermally similar. The design of the batch 16 j '

assemblies is identical to that of batch 15A. The design of the batch ISB h'

E & MICOE Vf 8"2 A McDermott company

7

! Zircaloy-clad assemblies is such that the thermal performance of this fuel is less limiting than the designs used in the remainder of the core.

Analyses for the incoming fuel and for batches 14, 15A, and ISB were performed 3

with the TAC 02 code, using the analysis methodology consistent with reference

4. Limits on heat generation rates for batches 9,14,15A, ISB, and 16 are documented in the Technical Specifications.5 The linear heat rates are based on a postulated loss-of-coolant accident (LOCA). The pin pressure for cycle 14 will remain below nominal system pressure.

i t

I L

  • mq

~,

=

. ;f

3

-J d

.1 m Babcock &WIIcom 4-3 _ . uc a n ,c.,,,,ny

(({

Table 4-1. Nominal fuel Design Parameters Batch 9 Batch 14 Batch 15A Batch ISB Batch 16 Manufacturer GAC/88W B&W B&W B&W B&W No. of assemblies 1 52 48 4 52 Previous irradiation, cycles 2 2' 1 1 0 Initial fuel enrichment -

4.0 4.0 4.0 3.41 4.0 Initial minimum fuel density, 94.9 94.9 94.9 94.9 94.9

% theor.

fuel pellet nominal diameter, in. 0.3835 0.3825 0.3825 0.36?0 0.3825 Active fuel stack nominal 120.3 120.5 120.5 119.0 120.5 p length, in.

u Batch burnup, 800, mwd /mtU 21,577 23,799 10,284 10,148 0 Initial gas pressure, psia 14.7 54.7 54.7 279.7 54.7 Gas composition, min % lie 80.0 95.0 95.0 95.0 95.0 Cladding material 304SS 30455 304SS Zirc-4 304SS Nominal cladding thickness, in. 0.0165 0.0165 0.0165 0.0270 0.0165 g Assembled fuel rod length, in. 126.68 126.68 126.68 126.125 126.68 f! Cladding collapse time, EfPH(a) 32,000 33,060 33,060 >31,200 33,060 D Design residence time EOC 14, EFPH(a) 29,568 29,952 18,912 18,912 9,600

  • f (*)EfPil denotes effective full-power hours.

n wa c.,

m.? ,: e > . , ,

a

,..,,em,.<-

w ,n n w _ . .

.O e m

i

5. NUCLEAR DESIGN 5.1. Physics Characteristics Table 5-1 compares the core physics parameters of cycles 13 and 14. The values for both cycles were generated using PD007. Since the fuel cycles differ in core average burnup, there are differences in core physics parameters.

t The number of fresh fuel assemblies loaded is the same for both cycle 14 and

~i the reference cycle. The design number of full-power days of operation and

( design cycle differential burnup differ between cycle 14 and the reference

, cycle. Figure 5-1 illustrates a representative relative power distribution for the beginning of the fourteenth cycle (B0C 14) at full power and normal c rod positions.

" Because cycle 14 core average burnup (B0C) is lower, the critical boron

., concentrations are higher than those for cycle 13. The ejected rod worths in a Table 5-1 are the maximum calculated values within the allowable rod insertion limits. Although they are similar, these values differ between cycles since

[- the isotopic and radial flux distributions are not identical. Calculated

. ejected rod worths and their adherence to criteria are considered at all times I in life and at all power levels in the verification of the existing rod

^

position limits (Technical Specification 3.10). The maximum stuck rod worths T for cycle 14 are used directly in the shutdown margin calculations, as shown a in Table 5-2. Ejected rod worths were analyzed 40 EFPD past EOC 14 to

,y accomodate possible coastdown past the design E0C.

O Shutdown reactivity is considered from hot full power (HFP) (4-loop operation) and 65% FP (3-loop operation) with sufficient margin for compatability with I _the. safety analyses.. A_IO% uncertainty. on_ net _ rod worth _ and_ the_ application ___ ___

of a flux redistribution penalty are additional conservatisms used in the g

3 3

sA ,

p - Babcocic & Wilcox g 5-1 ~ . uco,n,, ort conipany

~

shutdown analysis. Flux redistribution was accounted for since the shutdown analysis (power deficit) was calculated using a two-dimensional model.

The reference fuel cycle shutdown margin is presented in reference 2. To accommodate possible coastdown at E0C 14, the shutdown analysis was analyzed at 40 EFPD past design E0C 14. The remaining values for Doppler and moderator coefficients, boron and xenon worths, and effective delayed neutron fractions in Table 5-1 exhibit expected trends as a function of burnup.

5.2. Nuclear Design Changes The cycle 14 design meets all criteria including those applicable to radial power peaking, ejected rod worths, and shutdown margin. The same calcula-tional methods were used to obtain all of the important nuclear design parameters in cycle 14 as in the reference cycle. There are no significant

, operational procedural changes from the reference cycle with regard to control of axial or radial power shape, xenon, or tilt.  :

1

3 "I

.Q 0

3'

?

.. ?j

.D E SIDCOCN O W lICO E g, a McDermott company

{. .

l Table 5-1. Physics Parameters -- Cycles 13 and 14 I Cycle 13(a) Cycle 14(b)

Cycle design length, EFPD 338 360 Cycle design burnup, mwd /mtU 9,573 10,203 Average core burnup, E0C, mwd /mtU 21,711 21,625 Design initial cere loading, mtU 64.4 64.4 Critical boron -- B0C, no Xe, ppm HZP, all rods out(c) 1585 1608 HZP, groups B and A inserted 1210 1239 HFP, all rods out 1350 1433 4 .L Critical boron -- EOC, equil. Xe, ppm m HZP, all rods out 321 231 HFP, all rods'out 0 0 Control rod worths -- HFP, BOC, pcm Group A 1813 1924 Group B 748 893 Control rod worths -- HFP, E0C,(d) p ,

Group A 2244- 2245 3

Group B -

929 972 Maximum ejected rod worth -- HZP, pcm/F q y BOC (at rod insertion limit) 168/5.46 231/6.29 d EOC(d) (at rod insertion limit) 197/5.18 222/5.16 Maximum ejected rod worth - HFP, pcm/F q

BOC (at rod insertion limit) 26/2.42 23/2.44 5 EOC(d) (at rod insertion limit) 33/1.98 16/2.13 3 ~' ' ~

~ ~~~ Maximum stuck rod worth ~-- HZP, pcm ~~~

BOC 889 1196 EOC Id) 1354 1522 m

h

~ =~ a wucon _.

5-3 a nacoermor comp.ny

Table 5-1. (Cont'd)

Cycle 13(a) Cycle 14(b)

Power deficit, HZP to HFP, pcm BOC 1409 1361 E0C(d) 1686 1683 Doppler coefficient, pcm/"F BOC (100% power, no Xe) -1.39 -1.39 E0C (100% power, equil. Xe) -1.53 -1.53 E0C(d) (100% power, equil. Xe) -1.56 -1.56 Moderator coefficient -- HFP, pcm/*F BOC (no Xe, all rods out) -9.00 -8.30 EOC (equil. Xe, all rods out) -26.23 -27.00 EOC(d) (equil. Xe, all rods out) -26.52 -27.30 Boron Worth -- HFP, ppm /pcm B0C 0.148 0.149 E0C 0.136 0.136 t

'(

Xenon worth -- HFP, pcm BOC (4 days) 2293 - 2283 f, l EOC(equil.) 2355 2367 l ~.7 l

Eff. delayed neutron fraction -- HFP ':1 BOC 0.00621 0.00624 . _ ,

l EOC 0.00556 0.00555 -

(a) Based on cycle 12 length of 460 EFPD. ,)

(b) Based on cycle 13 length of 388 EFPD.

(c)HZP: hot zero power; HFP: hot full power. $

l (d)These values were calculated at 40 EFPD beyond the cycle design lengths to allow for possible coastdown at the end of the cycle. .4 i e) t.l b

ambcockawiscom 3 5-4 . ucoernione co,n,.ny .g

-\. .

]

1 Table 5-2. S5utdown Margin Calculations -- Haddam Neck Plant, Cycle 14, with Maximum Stuck Rod From HFP, 4 Loop, From 65% FP, 3 Loop, pcm pcm 0 EFPD 400 EFPD(a) 0 EFPD 400 EFPD(a)

Available Rod Worth Total rod worth, HZP 6302 6754 6302 6754 Max stuck rod, HZP -1196 -1522 -1196 -1522 Net worth 5106 5232 5106 5232 Less 10% uncertainty - 511 - 523 - 511 - 523 Total available worth 4595 4709 4595 4709 Required Rod Worth Pcwer deficit, HFP-HZP 1361 1683 757 696 Max allowable inserted 217 317 403 502 rod worth a

Flux redistribution 7. 187 6 163 q.

1

~'

Total required worth 1585 ,

2187 1166 1361 d'

Shutdown Margin 7

1 (Total available worth minus 3010 2522 3429 3348 v

total required worth) e I *

(*)These values were calculated at- 40 EFPD- beyond cycle' design' length to ~ ~

~

--- allow for;possible-coastdown-at.the end of- the cycle.. _ . .

T_ _

.3 Jr k

, - . ._ -ayyl8com 5-5 . . _ . _.. , m , , ,y 1

- ~- - - - - - - , . - --,s e , ,, , , , . . _ . _ _ _ - - - ,

Figure 5-1. BOC 14 Two-Dimensional Relative Power Distribution (Full Power, No Control Rods Inserted, Equilibrium Xenon) 8 7 6 5 4 3 2 1 H 0.87 0.98 0.91 1.14 0.94 1.01 0.93 0.80 G 0.98 0.90 1.12 1.02 1.16 0.95 1.17 0.69 F 0.91 1.12 0.97 1.18 1.15 0.97 1.00 E 1.14 1.02 1.18 1.16 1.'00 1.19 0.73 D 0.94 1.16 1.15 1.00 1.03 0.80 t

C -1.01 0.95 0.97 1.18 0.80 1

J B 0.93 1.17 1.00 0.73 1

A 0.80 0.69 ~)

la

{ .;

O X.XX Relative Power Density ,

. S

- T l'

NN b NE !e l

5-6 , ucoennort company fj i

~. ,

t. .

i

6. THERMAL-HYDRAULIC DESIGN 6.1. Desion Analysis and Safety Limit Curves The batch 16 fuel in cycle 14 was fabricated to the same tolerances as the batch 15A fuel which was evaluated in the cycle 13 analysis. The batch ISB Zircaloy-clad assembly design has been evaluated with TAC 02 3 and is no more limiting than the remainder of the core. Thus, the safety limit curves 6 and maximum design calculations ,10 applicable for operation in cycles 7 through 13 are also applicable for cycle 14.

6.2. Cycle 14 Thermal-Hydraulic Perfomance Results of the steady-state, thermal-hydraulic analysis for cycle 14 operation are given in Table 6-1. The corresponding design data are included for

' comparison and show that cycle 14 operation will result in less severe conditions than the maximum design conditions. The cycle 14 predicted

,. themal-hydraulic data are based on the design data listed in Table 6-2 and a the core loading shown in Figure 3-1. Additional inputs used to generate the predicted thermal-hydraulic data include the power distribution that gives the

] most limiting total peak during the cycle, maximum core inlet conditions, and a power level of 100% FP. .

The hot pin is located in assembly G-2 at 4 EFPD. The hot pin peaking was increased by the uncertainties and penalties listed in Table 6-1. The resulting minimum DNBR calculated for cycle 14 is equal to 2.48. The maximum heat flux and maximum linear heat rate are estimated to be 445,600 Btu /h/ft2 and 14.5 kW/ft, respectively. These data also include the uncertainties and 3 penalties listed in Table 6-1.

The limiting case- data reflect a revision in the core bypass flow from 9% to i

} 4.5%. The data also, reflect a proposed Technical-Specification change to the p RCS flowrate based' on steam generator tube plugging and measurement W uncertainty.

2 -

2  ;.

n -

Babcockawiscox 6-1

^

, ,ueo,,,,,,,,,,,,, -

The maximum fuel temperature was calculated with the TAC 02 computer code. The maximum steady-state fuel temperature during cycle 14 is expected to be less than 4102*F which is less than the maximum design value of 4780*F.

e* g I

e W .4

-s w

g-g, E

D

'?

A IE!.

e s --

== gg ,

E 6-2 . ucaron c% g

t. .

i Table 6-1. Cycle 14 Thermal-Hydraulic Data Design Cycle 14(a)

Hot Channel Factor Heat flux (Fq) 3.09 2.62 Enthalpy rise (FAH) 1.78 1.72 Coolant System Pressure Nominal, psig 2150 2000 Minimum steady-state, psig 2125 1970

, Coolant Flow Total flow rate, gpm 268,800 257,000

- Eff. flow rate for heat transfer, gpm 244,600(b) 245,400(c) 2

^ Total core flow area, ft 41.8 43.4 Average velocity along fuel rods, fps 13.67 14.25 Coolant Temperature, F

~

Nominal core inlet 549.5 537.0 a

l Max. core inlet, incl. instrumentation 553.3 544.6 error and deadband c\

.; Average rise in vessel 47.2 47.2

. Nominal vessel outlet 596.7 584.2 O Average rise in core 51.5 49.3

~

Average in vessel 573.1 560.6 Average in core 575.3 561.6 7

3 Nominal outlet, hot channel 634.3 596.3 Maximum outlet, hot channel 637.2 608.5(d)

_ Coolant Enthalpy . _ _ _ _ _ - _ _ - .

~

7 GJ Max. outlet, hot channel, Btu /lb 673.0 629.4(d)

Saturated enthalpy at min, pressure, Btu /lb 687.9 670.4 D

Il  !

Babcock &Wilcom

-y] _ 6-3_ _ , u c o ,,m ore c ,,y

Table 6-1. (Cont'd)

Design Cycle 14(a)

Heat Transfer Heat generated in fuel, % 97.4 97.4 Avg. film coeff, Btu /h-ft 2 'F 5205 5534 Avg. film AT, F 32.45 30.70 Max. U0 temp., F 4780 4102 2

Active heat tragsfer surface area of fuel rods, ft 35470 35540 Avg. heat flux, Btu /h-ft 2 170,980 169,900 2 445,600(d)

Max. heat flux, Btu /h-ft 528,330 Max. thermal output, kW/ft 16.9 14.5(d)

Max, cladding surface temp'. at nominal pressure, F 652.0 648.0 Minimum DNBR 1.84 2.48(d)

Power Distribution Uncertainties (Used for Cycle 14 Analysis)

Nuclear calculational uncertainty -- radial 1.05  ;

\

Quadrant power tilt penalty 1.06 Penalty for cycle 13, deviation "

from design E0C burnup 1.00 Multiplier to preserve design 3' pin peak (1.42) used in safety analysis 1.0032 a Heat balance error 1.02 7 Multiplier to obtain 14.5 kW/ft. 1.0453

.s Enthalpy rise engineering hot channel factor 1.099 2.'.

n Axial local grid factor 1.026 B

Heat flux engineering hot channel factor l'.02 S Densification spike factor 1.022 g C,;

bf 2

r amhencar & Wilcox -

N a McDermott company

~

{. .

  • Nuclear calculational uncertainty -- axial 1.024 Statistical densification factor 1.012 l

(a) Based on hot dimensions, 100% full power, and beginning-of-life fuel.

(b) Based on 9.0% bypass flow.

J (c) Based on 4.5% bypass flow.

(d) Includes engineering hot channel factors, maximum coolant inlet temperature, and minimum pressure.

,j r

i

.1

. .'l i

1 r

i j - _ _ _ - - - - - - - .- -- - --

7, t

4 s

b

's ,

.y y ~

+ 4* m

, 6-5 _ _ _ _

, ucm,,,,ott ceriy * **

Table 6-2. Cycle 14 Parameters -- Hot Conditions Total core heat output at 100% power, MWt 1825 Fuel rod OD, in. 0.4240 Fuel rod pitch, in. 0.5652 Active fuel length, in. 120.5(a)(b)

No. of fuel rods per assembly 204 No of assemblies in core 157 Flow area per assembly, in.2 39.78 Heated perimeter per assembly, in. 271.8 Wetted perimeter per assembly, in. 307.7 (a)Undensified. i (b)119.0 in, for Zircaloy-clad rods.

'M 'y S

3 J

E a'

i .?

k E FD 6-6 a uemm.oet eampany

7

7. ACCIDENT AND TRANSIENT ANALYSIS 7.1. General Safety Analysis In order to determine the effects of the cycle 14 reload and ensure that the cycle's thermal performance during hypothetical incidents is not degraded, I

each FDSA accident analysis was evaluated.

The initial conditions assumed in the original FDSA analyses were based on

" design" operating conditions (Table 6-1) and reflect operating procedures l

with uncertainties that were in effect during operation of the Haddam Neck Plant's first core. The cycle 14 operating temperature and pressure are lower than the values assumed in the original FDSA analyses. These reductions were implemented during a previous cycle. The lower temperature tends to improve thermal performance, while the reduced pressure degrades it. The basis for the magnitudes of these reductions was to maintain at least:ths same steady-state and accident thermal margin -exhibited in the FDSA. Therefore, the worst-case steady-state conditions that could exist consistent with cycle 14 operating procedures do not result in lower steady-state and accident thermal margins than those in the original FDSA analyses. Several accidents have been reanalyzed since cycle 1 due to changes in plant operations or regulatory cri-

, teria. The LOCA was re-analyzed during cycle 2 in compliance with the Interim Acceptance Criteria (IAC). Subsequent reanalyses addressed fuel densifica-

, tion, revised upper-head-fluid temperature, the gap increase plus prepressuri-

, zation modifications incorporated in batch 14 and steam generator tube plugging. The rod drop accident was re-analyzed during cycle 5 to support a modification of the control rod bank configuration. The steamline rupture

, ~

incident was reanalyzed during cycle 8 to support a change in the sequencing

? of the charging pumps during safety injection and in cycle 9 to support I

- automatic initiation of auxiliary feedwater. .l i _

{  ?

t, 1 Babcock & WHcom ~

-h 7-1 __ wm,, , . s.

l

O h For each FDSA Chapter 10 incident, the effects of the cycle 14 reload are com-pared to each of the current analyses to ensure acceptable thermal perform-ance.

The effects of fuel densification on the results of the FDSA analysis have been evaluated in reference 8. Since the batch 16 fuel reload (new cycle 14 assemblies) contains fuel rods whose theoretical density is higher than those considered in WCAP-82139 , the conclusions set forth in reference 8 are still valid.

7.2. Control Rod Withdrawal Incident A control rod withdrawal incident is defined as an uncontrolled addition of ,

reactivity to the reactor core by withdrawal of control rods. Two classes of rod withdrawals were analyzed in the FDSA: uncontrolled withdrawal from a suberitical condition and uncontrolled withdrawal from power.

The key parameters for a subcritical condition are the Doppler coefficient of reactivity, the moderator temperature coefficient of reactivity, and the dif- i ferential worth of a control rod group. The assumption in the FDSA analysis that no protection trip function is initiated until the overpower trip set-point is reached dictates that the fastest rod withdrawal from subcritical is the case of most concern. Although not credible itself since a malfunction of t i

protective trips would have to be assumed, this case results in a transient that is more severe than any of the worst credible cases and thus provides an ,

exceedingly conservative basis for evaluation. )

Since the fastest insertion of reactivity results in the worst case, the least g negative design value of -1 pcm/*F for the Doppler coefficient of reactivity, the most positive expected value of +10 pcm/ F for the moderator temperature ,

coefficient of reactivity, and the maximum expected differential worth of a control rod group of 90 pcm/ inch was assumed in the FDSA analysis.

9 The corresponding hot zero power parameters (535F) predicted for cycle 14 6 are -1.72 pcm/*F for Doppler coefficient, -4.09 pcm/ F for the moderator temperature coefficient, and 40.08 pcm/ inch for the maximum expected f differential worth of a control' rod' group. Therefore, the key parameters assumed in the FDSA analysis adequately bound the cycle 14 predicted values so @

c h

a Babcock &WHcom- ,

7-2 a ucoermor company y

3

{+ ,

i .that the consequences of an uncontrolled rod withdrawal from a subcritt al condition would be no more' severe than those reported in the FDSA.

The uncontrolled rod withdrawal from power results in increases in heat flux and reactor coolant temperature. Unless terminated by manual or automatic action, the increase in heat flux and the reactor coolant temperature rise could eventually result in departure from nucleate boiling (DNB). Both slow and fast reactivity additions must be considered in the analysis because of the several ways (heat flux increase, temperature increase, and pressure de-crease) in which the DNB condition can be approached. .

The protective circuits that are effective against rod withdrawal at power are the overpower trip and the variable low-pressure trip. The combination of these two trip functions protects the core against a rod withdrawal incident 1 regardless of the rate of reactivity insertion. A very rapid reactivity in-sertion would cause power to increase significantly before the reactor coolant

, could respond, and the transient would normally be terminated by the overpower trip. The reference I analysis of T.hc red withdrawal incident from power in-cluded parametric studies on varying insertion rates, moderator coefficients, and Doppler coefficients, and identified the most severe transient as having the combination of- parameters- that results in simultaneous overpower 7 and variable low-pressure trips.

Because of the reduction in operating pressure and temperature, the combina-tion of parameters that would be necessary to reach the corner point of the protection diagram (Figure 10.2.1-5 of the FDSA) was reassessed in reference

} 10., The conclusion reached, as reviewed and approved by the NRC Staff, was 1

that the minimum margin to DNB for the most severe rod withdrawal from power i3 during cycle 7 was no less than the margin reported in the FDSA. This same

  • conclusion is applicable to cycle' 14.

T 7.3. Isolated Loop Startup Incident

,x If a reactor loop is isolated from the remainder of the reactor coolant system and subsequently brought back into operation without first matching its baron f

~

concentration and temperature to those of the system, an increase ~ in~ core're-a activity and power may occur. To prevent this and ensure the safe startup of ,

J an isolated loop, procedures (outlined in Section 10.2.2.1 of the FDSA) have q  :

w  ;

Batacock&WHcom b R( 7 acom.on ceny  !

.c

. _ . - - . , _ . - . - , , . . . . . , . - - . . . . _ _. ,.,,,.-..-y,r___. ___.,~.--.-..-r . . . . - . . . - . . - - - - - , , . . - - - - - - - - - - -

been established. In addition, a temperature-valve interlock prevents opening of the cold leg stop valve if the temperature difference between the hottest-operating loop cold leg and the isolated cold leg exceeds 20F. A valve-pump interlock prevents starting a reactor coolant pump unless the cold leg valve in that loop is closed.

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 leg temperatures. This case was identi-fied in the FDSA as the most severe transient from the standpoint of CNB.

To maximize reactivity addition, the FDSA analysis assumed the most negative moderator temperature coefficient expected (-35 pcm/*F). To minimize negative reactivity feedback, the least negative Doppler coefficient of -0.5 pcm/*F was assumed. The corresponding most negative moderator temperature coe#ficient +

and least negative Doppler coefficient predicted for cycle 14 are -27.30 pcm/ I

'F (at 400 EFPD cycle length) and -1.39 pcm/*F, respectively. Since the pre-dicted cycle 14 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 14 (

would be less severe than previously reported.

7.4. Boron Diluti. Incident Tj The boron dilution incident considers the inadvertent dilution of reactor g

! coolant boron concentration. The result would be a reactivity increase lead- [.

ing to either a power increase or a loss in shutdown margin. To cover all l q phases of plant operation, boron dilution during refueling, startup, and power {;

operation wa:: considered in the FDSA analysis. The key parameters in this incident are the initial boron concentration, the dilution flow, the concen- 0 2) tration at which the reactor would go critical or at which shutdown margin is lost, and the effective reactivity addition rate due to dilution. The cycle g 14 BOC critical boron concentration with all rods in at 70F, is 1122 ppm. The 5 initial concentration during refueling and startup that provides, with all 3.

9 rods in, a shutdown margin of 8000 pcm has been calculated to be 1952 ppm. 4 5

Bat > cock SiMcom a, 7-4 . uconmor company y

i With a maximum dilution flow of 180 gpm, the times required to dilute the reactor boron concentration from 1952 to 1122 ppm (the concentration at which the reactor would go critical) would be approximately 1.07 and 2.88 hours0.00102 days <br />0.0244 hours <br />1.455026e-4 weeks <br />3.3484e-5 months <br /> during refueling and startup, respectively. These are ample times for the operator to recognize the audible high count rate signal and to terminate the dilution flow before the reactor would go critical.

A boron dilution incident at power is concerned with a loss in shutdown margin

~

or a return to criticality after a reactor trip. Based on a maximum dilution flow rate of 180 gpm and a maximum concentration of 1608 ppra, the maximum cycle 14 effective reactivity addition rate for boron dilution at power is 0.90 pcm/sec. With a BOL cycle 14 shutdown margin of 3010 pcm, approximately I

55 minutes are available for the operator to terminate dilution. Therefore, even if an unintentional dilution of baron in the reactor coolant. did occur, numerous alarms and indications are available to alert the operator to these conditions. The maximum reactivity addition due to the dilution is slow enough to allow the operator to determine the cause of the addition and take corrective action before shutdown margin is lost.

7.5. Excess Feedwater Incident

,~ ~ ~

The ' excess feedwater incident 1s'a result of an abnormal ' sustained increase in j 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 steam generator, re-f sulting in a cold leg temperature reduction and a ' subsequent increase in reac-l2 tivity that would lead to a power excursion. The power excursion is maximized

!' with the most negative moderator temperature coefficient (ensures maximum re-in activity addition from the reduced coolant inlet temperature) and the least

  • negative Doppler coefficient (ensures minimum reactivity feedback). The val-l- ues assumed in the FDSA analysis are -35 pcm/ F for the moderator and -0.5 5 pcm/*F for the Doppler coefficients of reactivity. The corresponding values t

predicted for cycle 14 are -27.30 pcm/*F and -1.39 pcm/*F, respectively.

39 jg The values of moderator temperature and Doppler coefficient predicted for cycle 14 are within the bounds assumed in the previous _ safety analysis.

2

~ ~

Therefore,the]DSAanalys'iscoversthemostadverseexcessfeedwaterincident Q

h Ia 3 sabcockawucox 7-5

,g . _ ,,,_ ,uco,,,,,,co ,p,ny

that could be postulated to occur during cycle 14. The bounding analysis of the FDSA 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 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.

7.6. Excessive Load Increase An excessive load increase incident is defined as a rapid increase to steam generator steam flow resulting in a significant power mismatch between the reactor core power and the steam generator load demand. The excess load re-sults in a decrease in reactor coolant temperature. To maximize the resultant power excursion, the FDSA analysis considered the most negative moderator co-efficient of reactivity (-35 pcm/ F) and the least negative Doppler coeffi-cient of reactivity (-0.5 pcm/ F). The corresponding values predicted for cycle 14 are -27.30 and -1.39 pcm/ F for the most negative moderator coeffi-  :

cient and the least negative Doppler coefficient, respectively. This com-bination of moderator and Doppler coefficients would result"tra 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 14.

7.7. Dropped-Rod-Incident .

A dropped-rod incident is the unexpected release of a control rod, which would cause it to fall into the core. This incident was analyzed in reference 11 to determine the effect of the cycle 5 modification to control rod bank B;

]

sensitivity studies of the parameters pertinent to the rod drop incident were included. The ranges and variables included in these sensitivity studies and  ?

the corresponding cycle 14 predicted values are sumarized in Table 7-1.

Since the variables predicted for cycle 14 are bounded by the values' ge considererin the reference 11 -analysis, this analysis is used to determine ,

the cycle 14 thermal margin during the postulated dropped-rod incident. 3 S

y

~

- m4. l Baticock &WHcox g.

7-6 ,uco,,morte . ,ny

(* 4

! To use the reference 11 analysis for cycle 14, the maximum increase in local

., peaking as a result of a dropped rod is needed. This value is 1.083 for cycle l 14. With this value, Figures 3.2-6 and 3.2-7 of reference 11, and the maximum cycle 14 dropped-rod worth of 131 pcm, the minimum DNBR and maximum fuel centerline temperature can be determined. The minimum DNBR and maximum fuel centerline temperature resulting from the most severe rod drop incident postulated for cycle 14 are 1.77 and 4312F, respectively. It should be noted tha' these results are based on design hot channel factors (Fq = 3.09 and FaH

= 1.78) and uncertainties consistent with the FDSA.

7.8. Cor. trol Rod Ejection Incident 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. Table 5-1 provides the cycle 14

' values for the key parameters used in the FDSA control rod ejection analysis.

In all except the B0C hot zero power case, the cycle 14 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

~

i damage for the most severe rod ejection incidents that could be postulated.

For the cycle 14 BOC, zero power case the total F q predicted for cycle 14 m (6.29) is 26% greater than that assumed in the FDSA analysis (5.00), even U though the cycle 14 predicted ejected rod worth (231 pcm) is substantially c3 lower than that used in the FDSA analysis (830 pcm). The effect of the higher A

31 F q

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 26% change in linear heat generation rate from reference 9. The results of this calculation indicate 5 that the peak fuel centerline temperature will remain less than 3650F. Note that this calculation was perfomed using the 830 pcm FDSA rod worth value instead of the cycle 14 231 pcm value; thus, the actual peak fuel centerline

- temperature would be expected to_be considerably _less .than _3650F_These _ ._

temperatures are. considerably: lower than the 4700F UO2 melting temperature; p therefore, safety limits will be neither approached nor exceeded.

i' q

S .

ah&WHcom 7-7 _ . a ueon.a,,eamo,ny _

3..

7.9. Loss-of-Coolant Incident The loss-of-coolant incident was addressed and included the type of fuel that will be in the cycle 14 core. The effects of fuel densification were address-ed in reference 8 and the effects of a revised assumption concerning the upper head fluid temperature were addressed in reference 12. The resultant limiting linear heat generation rates are included in the basis for the axial offset limit curves (Figures 8-1 through 8-6).

7.10. Loss-of-Flow Incident A loss-of-coolant flow could result from a mechanical or electrical failure in one or more reactor coolant pumps or from a fault in the power supply to these pumps. If the reactor is at power at the time of the incident, the immediate effect of loss-of-coolant flow is a rapid increase in coolant temperature due to the reduction of heat removal capability. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly.

As indicated in the FDSA analysis, the most severe loss-of-flow incident  !

3 occurs from full power (1825 MWt) with the loss of all four reactor coolant pumps. Since the hydraulic design and total core pressure drop have not  ;

significantly changed from the original core, the flow coastdown curves used in the FDSA analysis still apply. The least negative values of the Doppler and moderator temperature coefficients (-0.5 and +10 pcm/*F) were -

assumed in the reference 1 analysis since they result in the maximum hot spot ,

heat flux during the initial part of the transient. -

The hot full-power least negative values of Doppler and moderator temperature q

coefficients predicted for cycle 14 are -1.39 and -8.30 pcm/*F, respectively. D Since.the cycle 14 predicted values are bounded by the FDSA study, the results of that study conservatively estimate the thermal margin for the most severe d

. loss-of-flow incident that could be postulated to occur during cycle 14.

"i 7.11. Steam Line Rupture Incident 9

The limiting steam line rupture incident identified in the FDSA is a i circumferential double-ended rupture of a 24-inch steam line from the steam generators upstream of the. steam line isolation valves. This rupture was i o,

analyzed in the FDSA from both full-and zero-power conditions. For both cases N':

~

2

~r Babcock & WHcox . 7 7-8 , %,,,, y g, e

9 .

! the FDSA analysis assumed a most negative moderator coefficient of -35 pcm/*F and a Doppler coefficient of .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 14 predicted values for moderator and Doppler coefficients are -27.30 and -1.39 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 14 is 2522 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 14 steam line rupture incident.

The maximum loss of shutdown margin was re-evaluated in reference .13 to i support a plant modification that implemented automatic initiation of

. auxiliary feedwater. The cycle 14 design was evaluated to assure that q sufficient shutdown margin is provided by the power dependent control rod insertion limit (Technical Specification 3.10) to prevent a return to criticality.

1 l The required shutdown margin based on cooldown deficits and available shutdown margin based on the cycle 14 ~ design are provided in Table 7-2 for the limiting hot full and hot zero power cases for four and three loop operation.

~

These results demonstrate that the cycle 14 design provides sufficient shutdown margin to prevent a return to criticality.

. 7.12. Steam Generator Tube Rupture Incident

.) 4 The integrity of the steam generator is significant from the point of view of radiological safety. The radiological consequences of this incident are inde-pendent of core loading and remain acceptable.

7.13. Loss-of-Load Incident

> 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. .

.j _ Nonnally, a .large loss will cause a turbine trip,_ either by a signal _from the _

generator-or switchgear 7or from the turbine ~ overspeed; trip signal.. -If the

^

t i

3 - --

Babcock &WWilcom - -

7 '

0 -

a uconmou m -r iy

I turbine 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 con-trol valves respond quickly enough to prevent the overspeed trip of the tur-bine, the load could be reduced to the station service level without a reactor trip. This is the largest credible 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 determin-ing 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 coupled with the least negative Doppler coefficient.

Values assumed in the FDSA were +10 and -0.5 pcm/ F for the most positive mod-erator coefficient and the least negative Doppler coefficient, respectively.

The corresponding hot full-power values predicted for cycle 14 are -8.30 and

-1.39 pcm/*F. The comparison indicates that the parameters assumed in the FDSA analysis bound the values predicted for the loss-of-load postulated to j' occur during cycle 14 without immediate reactor trip, and that the loss-of-load presents no hazard to the integrity of the core, the reactor coolant sys- ,

Pressure-relieving devices' incorporated in the tem, or the turbine cycle.

system are ample to limit the maximum pressure to acceptable values.

7.14. Loss-of-Feedwater Incident This incident is a reduction in feedwater flow to the steam generators when 2 1

operating at power without an equivalent reduction in steam flow, thus reduc-ing the water inventory in the affected steam generators. Since the complete y loss of feedwater flow requires the most rapid response from the reactor 23 protection system, it fonns the basis for the loss-of-feedwater flow incident 5.)

analysis reportcd in the FDSA. The maximum core power level before trip and d the minimum DNBR occur with the maximum positive value of the moderator tem- _

3 perature coefficient coupled with the minimum negative Doppler coefficient: y

+10 and -0.5 pcm/*F. The maximum positive moderator temperature coefficient and the minimum negative Doppler coefficient presented for cycle 14 at hot full power are -8.30 and -1.39 pcm/*F, respective,1y. The values assumed in the FDSA analysis adequately bound the values predicted for cycle 14; thus, g the results reported in section 10.3.6 of the FDSA are applicable to the cycle U 14 reload, a

,M A Micou p 7-10 ,uc u mor,com ,,ny g

9 P .

7.15. Fuel Handling Incident

-,. The fuel handling incident considers the possibility of dropping a fuel assem-t bly during fuel handling operations. The concerns over this incident are radiation exposure and accidental criticality. These concerns are independent of core loading.

7.16. Waste Gas Incident The waste gas incident is defined as an unexpected and uncontrolled release to the atmosphere of the radioactive xenon and krypton fission gases stored in the waste gas decay tank. The consequences of this incident are independent of core loading; therefore, the results reported in reference 12 are applic-g able for any reload.

7.17. Hypothetical Accident Regardless of the ability of the safety injection and core deluge systems to a prevent major fission product releases to the containment, the safety of reac-I tar 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. _ .

.l lt

! L 7

3 i

.J 3 __ _ . . _ __ _ __._ ._ _ __ . _ _ _ _ . _ . .-

b

h J rt - Babcock &Wilcom i 7-11 _a ucht cwriy _.

g _

o

  • Table 7-1. Dropped-Rod Incident -- Comparison of Parameters Assumed in Cycle 14 predicted, Parameter reference 11 hot full-power Moderator temp coeff, -38.50 to -1.50 -8.30 (0 EFPD) pcm/*F -27.30 (400 EFPD)

Maximum dropped rod 500 to 30 125 (0 EFPD) worth, pcm 131 (400 EFPD)

Doppler coeff. pcm/*F -1.11 -1.39 (0 EFPD)

-1.56 (400 EFPD)

Maximum local peaking Figure 3.2-8, 1.083 (0 EFPD) increase with dropped rod reference 11 1.074 (400 EFPD)

Table 7-2. Steam Line Rupture Incident Required Available shutdown shutdown Subcritical Conditions margin, ocm maroin, pcm margin, pcm Four Loop Hot Full Power 1900 2522 622 Hot Zero Power 3460 4241 781 _I Three Loop  :

1 Hot Full Power 2870 3348 478 Hot Zero Power 3910 4287 377 ,

a

  • J El U

Cl d

Ee b

2 L

~ "

Babcock & Wilcox - '

r-7-12 , ua, mon co.ony g

\'

  • k 1

i

, 8. TECHNICAL SPECIFICATIONS 8.1 Power Versus Offset Limits: Four-Loop Operation Axial power offset limits are prescribed to prevent core total power peaks above a specified limit. Technical Specification Figures 3.18-la , 3.18-lb, and 3.18-1c (Figures 8-1 through 8-3) show the specific cycle 14 four-loop power versus offset limits in relation to new four-loop generic Technical -

Specification limits. The applicability of the generic limits is demonstrated by the comparison with cycle 14 specific limits. The limits in Figures 8-1,

{. 8-2, and 8-3 preserve the LOCA linear heat rate limit and a 1.42 design radial

, pin peak throughout the fuel cycle. They also incorporate the measurement uncertainty corresponding to monitoring axial offset.

The limits presented here were generated from a three-dimensional power distribution analysis, including the effects of thermal-hydraulic feedback, Durnup, transient xenon, and control rods. Margins to the linear heat rate limit are calculated from the power distribution data base simulated for cycle 14. The total peak in the hottest fuel pellet is used to establish

~

the operating limits, and is found by applying the following multipliers to the calculated cycle 14 peak:

} F (x,y,z,t)=F (x,y,z,t).RL(x,y,t).F (t).F q p D

. AL.S(z).F'.F R

.QT.HBE.F u c 1

F c

2

$ where a F = highest pellet relative power density (RPD),

4 i F = FLAME 3 calculated peak, p

' RL = radial-local factor as a function of assembly position and burnup, 1.00 to 1.60, FD = burnup-dependent multiplier to preserve a .1.42 design pin peak, 1.00-to 1.08, cp 1

~

.~b 5 . Eh&Wilcom 8-1 _ _ _ _ .,_. wo,,,,on -y _

FR = maximum allowable increase in radial peak with decreasing power assumed for safety analyses, 1.00 to 1.20, AL = axial-local grid factor,1.026, S = spiking factor as a function of axial elevation in the active core, 1.00 to 1.06, Fe = hot channel factor, 1.020, 4

F = nuclear uncertainty, 1.075, u

QT = quadrant tilt peaking factor,1.060, HBE = heat balance error, 1.020, F = penalty for cycle 13 deviation from nominal by -25/+35 EFPD,1.00 to c

1 1.03, F = effect on peak caused by depletion with bank B partially inserted, c

2 1.00 to 1.04.

The radial-local factor is an indication of the increase in relative power \

density of the hottest pin relative to the average of the assembly in which it resides. These pin peak-to-assembly average ratios are calculated at each time step for each assembly during the PDQ07 cycle depletion.

The peaking margin for each node of each assembly is calculated by comparing g the calculated, augmented linear heat rate at each node (using Fq , above) with the burnup-dependent allowable LOCA linear heat rate at each node (using Fq ,

above) and detennining a percentage difference:

F (x,y,z,t) - AVGLHR - F0P ,l a 100 3

% MARGIN = (1- )

LIMIT (t) where % MARGIN = margin to LOCA linear heat rate limit, %,

~

F = defined above,  !

9 .J l AVGLHR = core-average densified linear heat rate, 5.656 kW/ft, FOP = fraction of power, i

LIMIT = burnup-dependent LOCA linear heat, rate limit, kW/ft. ,

l

)

y

- NabCoCk &MfCox ~

l 8-2 , ucc,rmott company .;g b

\' '*

I The densified linear heat rate is obtained from the core-average undensified linear heat rate by applying the following augmentation factors:

F = stack shortening factor, 1.007 ss Fsd = statistical density factor, 1.012 The results of the margin calculations are then plotted, and offset limits are determined which preserve the LOCA linear heat rate limits given above in each of the Technical Specification windows.

For four-loop operation, the value of c y, is equal to 1.00 because the effect of the previous cycle length's deviation from nominal is implicit in the FD factor. The function represented by FR provides a maximum allowable increase in radial peak with power equal to 1.42 x (1. + 0.2 x (1-F0P)), where F0P =

, fraction of power. Values of the burnup-dependent LOCA linear heat rate limit 1

i used in the analysis for four-loop operation are as follows:

7 0-125 EFPD: 14.30 kW/ft

{

125-250 EFPD:. 14.50 kW/ft 250 EFPD to End of Design Life: 15.50 kW/ft

]

- 8.2 Power Versus Offset Limits: Three-Loop Operation 1 Technical Specification Figures 3.18-1d, 3.18-le, and 3.18-1f (Figures 8-4 through 8-6) show the specific cycle 14 three-loop power versus offset limits R in relation to new three-loop generic Technical Specification liinits. The

~

applicability of the generic limits is demonstrated by the comparison with e

cycle 14 specific limits. The limits in Figures 8-4, 8-5, and 8-6 preserve-

~

the LOCA linear heat rate limit, and incorporate the measurement uncertainty corresponding to monitoring axial offset.

The three-loop offset limits were generated from three-dimensional power distributions which simulated operation at three-loop conditions. The multipliers described in Section 8.1 were applied to the calculated ~ total peaks, margins to the linear heat rate limit were calculated, and three-loop offset limits. were. determined. . _ _ _ _ . _ _. . _ _._

m .

w q .

il m - meswucox

8-3 - acoennon compny

F For three-loop operation, the value of cy is greater than 1.00 because F D becomes smaller than the value required to preserve the cycle length flexibility of -25/+35 EPFD. The function represented by F pr vides a R

maximum a 'llowable increase in radici peak with power equal to 1.42 x (1.0236 +

0.2 x (0.65-F0P)); assuming that the radial peak increases in this manner allows a 1.454 design radial pin peak tio be maintained for three-loop operation at 65% FP, and bounds the actual increase in peak with the -25/+35

.EFPD cycle length flexibility included. Values of the burnup-dependent LOCA linear heat rate limit used in the analysis for three-loop operation are as follows:

0-125 EFPD: 9.295 kW/ft 125-250 EFPD: 9.425 kW/ft 250 EFPD to End of Design Life: 10.075 kW/ft ,

i

  • g ,

<.E lI er

!.D

$I "D

ib 3

L1 E

a 9

a s ,

g Babcock &Wilcox g a McDermott company g:

I l -

J 0

0 1

D P

b 0

, 0 9

F E

5 N 4, 6

9 4

1 e

, 0 8

2 1

\ (

l e

c de n

, 0 7

0 s

)

\ C y

G e

0 6

t i

m) 0 7, \ -

0 i a Ll (

1 4 6 -

0 , 5 0

e1 8

) )

- , 4 s . 0 0 f3 f

0 1,

\ 4, 3 0

M Oe r 4

\ 0 2

su (

1 (

0 Vg J -

rF i

)

0

\ O ,

2 e n 0 0 w o 1, , 1 t

oi 0 0 0 20 0 0 0 0 0 0 e P t 1 9 8 6 5 4 3 2 s nc a ( 1 f_ _

1 0 f f

oi ) O if ti 0

0 )Eo 0 1  %

) ,

ac 1, 0 re ep 1 1 7, 0

/

pS 2 O

- 0 , -

>- ( 3 pa l

) / (

0 oc 0

! 3

- oi 0 O ,

Ln h 1, / )

0 0 r' rc 4 4, , 4 ue 1

- ) 0 oT ( 0 3 0 F(

7, -

5

( -

N 1

1 4

(

/ 0 6

[-

8 e .

f ,

0 61 r

u g

/ )

, 7 I .

i F / 0

'4, ,

0 8

M

/9 6 0

9 j

ti d -

,1 0

0 d

h is b t bh t

4 1 [ . , g 8Y i

n s

3  ;

. = i}- i i .I'

figure 8-2. Four-Loop Operation Power Vs Offset Limits, 125-250 EFPD (Technical Specification Figure 3.18-lb)

(-18,100) (- 5.100) (10 1 0) (16,100) ,

100

/ . . 90

/

\

/ \

- 80

/ \

(20,70)

(-42,70) 1 -30,70) -

- yo (42,70)

/ \

, / -

- 60 \

E / \

/ g- - 50 \

/ 5 \

[(-96,40) (-30,40) O E- - 40 O (20,40) (96,40) Y

- 30

- - 20 I

jp t -

- 10

- - Cycl e 14

- Generic l ,

$D e i i  ; e i e e a e e i e i i I  : I

$ -100 80 60 40 20 -10 0 10 20 30 40 50 60 70 80 90 100 2$M  % Offset TEi1 '?h D E10 ?LWf) UW L i1 w) u i: ._

j a

]

j e

, f i

L g

n 0

i i 0 s 1 f

D e

h i 9

0 d

o \ 0

)4, 6

4 1 c e 0

\

i 8

n (

9 e r E l e

c y e n

t o C G i

7 0

~

D P

F

)

0 7,

\ -

0 6

E 0

5 2

4 x 6-O

0 5

2

) 0 0 \ ) - i 4

s 0 \ 0 t 1 i) 4, 0 mc 4 0 i 3 i1 L -

8 2

(  % . (

2 t1

)

0 O i 0 2

_ e . 0 s3 1 f .

i t

fe 0 0 0 0 '0 0 0 0 0 0 0 0 e Or 1 0 9 8 6 5 4 3 2 1 s u ( 1 1 f

_ sg Vi )

- 0 f

O

- F 0 r 0 0  %

en 1 )

h 1

wo 0 0 i -

oi Pt 2 7, 0

- 0 2 a ( 3 e -

nc -

oi ) (

if 0 0 ti 0 3 0

ac 1 i

re ep 4

) 0 0

O pS 2

/ 4, e 4

pa l

( / 0 3 0 oc (

i 5 oi -

Ln rc ue h 0 7, -- e 0

6 .

F(

oT 9

4

/ 0 T

(

/ ')

i 7

3

/ 0 0 8 / 4, 6

i 8

/

e 9 0 r -

e 9 I

u d

g

. i 0 F i 0

1 4

7 .

T" h , *8M ',

,r$3gg

4 Figure 8-4. Three-loop Operation Power Vs Offset Limits, 0-125 EFPD (Technical Specification Figure 3.18-2a)

(-8,65) (-7,65) - 70 (10,65) (14,65)

/ -

- 60 g

/ \

/ \ -

/ -

- 50 \

(-41,45.5)I (-30,45.5) (20,45.5 1,45.5)

/ -

-40 N

/ N

/ N

/ N

/ -

-30

[ Y (-96,26) (-30,26) O g 2- -20 O(20.26) (96,26)

==

-10 - + - Cycle 14 m Genric l i i e i i n' i i i a a i a i a a i i i 100 80 60 40 20 -10 0 10 20 30 40 50 60 70 80 90 100 m  % Offset et E

h i

go O

1 E 48M

[i'd! ['6!T bif Eh3 fi.;) Qi;) ' : ~i L, _' _._ ___

' f 9

,na $U M h d Ab da ._ _ ,

._ a _ ,

j _.J .__ J i

3 l

Figure 8-5. Three-Loop Operation Power Vs Offset Limits, 125-250 EFPD (Technical Specification Figure 3.18-2b) t

(-16,65) (-12,65)- - 70(10,65) (16,65)

- 60 \

/

/ , - - 50 \

(-42,45.5)g (-30,4 5. 5) *

(20,45.5

/ N

/ -

- 40 N s' p u N

\

/ N cn / E- - 30 b (-30,26)O E O(20,26) (96,26) f(96,26)

I, - u

- 20

- 10 .

- - Cycle 14 O Generic I I I I I I E I I I i i i I i I I I I E

[ -100 80 60 40 20 -10 0 10 20 30 40 50 60 70 80 90 100 i

6  % Offset 7

1 5o O

em 48M 1 9 - - - - - -

Figure 8-6. Three-loop Operation Vs Offset Limits, 250 EFPD to End of Design Life (Technical Specification Figure 3.18-2c)

(-24,65 (-20,65)' '(7010,6 (20,65)

/ -

- 60 g

/ \

/ -

- 50 \

(-49,4 5. (-30,4 5. 5) 42,45.5)

(20,45.5 )

\

/ -

- 40 g

. / N

/ b- - 30 \

" g

. AT-96,26) (-30,26)O j O(20,26) (96,26) %

e. . go

- 10 - - Cycle 14 O Generic i t i i i i t i t I i t i I I I i f i 1

-100 80 60 40 20 -10 0 10 20 30 40 50 60 70 80 90 100

,[  % Offiset E

t i D l

5 i

M

'7i5il 9!!'E E {l I;2fB (: ;I Ctif, zuj 1,2_? 4 _ , _ . , _ _ _

~ -

J.

9. 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.

Prs-Critical Tests

1. Hot control rod drop-time testing.

Zero-Power Tests

_ 1. Critical boron concentration.

2. . Temperature reactivity coefficien..
a. All rods out.

i b. Banks B, A, and D in.

3. Control rod group worth for banks B, A, and D.
4. Differential baron worth.
5. Ejected rod worth.

Power Tests

1. Core power mapping < 80 and 100% full power, norinal bank configuration.
2. Excore/incore correl.ation, verification.

v .

I s

, -] W i ,L l -

5 f.i e

S~l '

n=w & WHcom r a A4cDermott company

] ,. ,, -

1 1

i

10. REFERENCES
1. Facility Description and Safety Analysis (FDSA), Haddam Neck Plant, NY0-3250-5, Connecticut Atomic Power Company.
2. Technical Report Supporting Cycle 13 Operation, Haddam Neck Plant, Connecticut Yankee Atomic Power Company, BAW-1822, Rev. 1, Babcock &

Wilcox, Lynchburg, Virginia, June 1984.

i

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

^

4. J. H. Taylor (B&W) to J. S. Berggren (NRC), Letter, "B&W's Responses to TACO 2 Questions,' April 8, 1982.
5. Ha,ddam Neck Plant, Technical Specification 3.17.

1 J 6. Haddam Neck Plant, Technical Specification 2.2.

s 7. Technical Report Supporting Cycle. VI Operation and Proposed License Amendments, Haddam Neck Plant, Section 1, Docket No. 50-213, Connecticut Yankee Atomic Power Company, May 1975.

8. Description and Safety Analysis, Including the Effects of Fuel Densifica-tion on the Connecticut Yankee Reactor, Cycle V, Docket No. 50-213, t Connecticut Yankee Atomic Power Company, November 1973.

~~

-0. Effects of Fuel Densification on the Connecticut Yankee Reactor,

]- WCAP-8213, Westinghouse Electric Company, Pittsburgh, Pennsylvania,

' ~

October 1973.

h 10. Technical Report Supporting Cycle VII Operation Ug. Proposed License Amendments, Haddam Neck Plant, Docket No. 50-213, Connecticut Yankee

Atomic Power Company, May 3,1976, and Supplement, June 22, 1976.
11. Connecticut Yankee Cycle 5 Modification of Control Rod Bank B, YAEC-1080, 3

i Connecticut Yankee Atomic Power Company, December 10, 1974.

h4 P Bat >COCit & MICOE f 10-1 a ucmrmon wny

.c s. ,,

REFERENCES (cont'd)

12. D. C. Switzer to A. Schwencer (NRC), Letter, May 2, 1976.
13. W. G. Counsil to Director of Nuclear Reactor Regulation, Automatic Ini-tiation of Auxiliary Feedwater, January 30, 1980.

a 4

4 i

.g i

.D

! .fi

'. t u

'"x o

d

- -s C5

A -

O f

, e,i

  • ai-i $

i '

10-2 m, ,, ,y ,f

' , . . _ . . . . .._. ,,_ _ __ ___ _ . _ . , , ~ . _ . . . _ _ . . _ , , - . _ . . . _ . _ , _ _ , _ . . _ . . . _ , . _ . _ _ . . , _ _ . , _ _ . -

Retrieved from "https://kanterella.com/w/index.php?title=ML20205K355&oldid=3020988"