ML19260A055
| ML19260A055 | |
| Person / Time | |
|---|---|
| Site: | Three Mile Island  |
| Issue date: | 11/30/1976 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| References | |
| BAW-1442, NUDOCS 7910290684 | |
| Download: ML19260A055 (56) | |
Text
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I BAW-1442 Nove=ber 1976 t $ ! ,, ( -. 5 I
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, L i r., THREE MILE ISLAND UNIT 1 CTCLE 3 RELOAD REPORT i auk i e
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I I 1 j L. 1 s 11' ( BABCOCK & WILCOX l Power Generation Group
,- Ricient Paver Generation Division i
P. O. Box 1260
- e. Lynchburg, Virginia 24505 l4 k [24 1
1 Babccck &Wilcox
.l-
b J o i COM I i Page g 1. INTRODUCTION AND
SUMMARY
. . . . . . . . . . . . . . . ...... 1-1 7' 2. OPERATING HISTORY ............. ........... 2-1 I 3. GENERAL DESCRIPTION' .............. ......... 3-1
- 4. FUEL SYSTEM DESIGR . . . . . . . . ................ 4-1
' 4.1. Fuel Assembly Hechanical Design .............. 4-1
.3 4.2. Fuel Rod Design ............. ......... -
4-1 '
^
4.2.1. Cladding Collapse .......,......... 4-1 4.2.2. Cladding Stress .. ........,,...... 4-2 4.2.3. Cladding Strain ....... .,......... 4-2 4.3. Thermal Desigu . . . . . . . . . . . . . .......... 4-3 e 4.3.1. Power Spike Model ................. 4-3
-\ 4.3.2. Fuel Temperature Analysis ...,,........ 4-3 4.4. Material Design ................. . . . . 4-3
. 4.5. Operating Experience . . . . . . . . . . . . . . . ..... 4-3
}:L, 5. NUCLEAR DESIGN . . . . . . . .............. ..... 5-1 5.1. Physics Characteristics ... ............... 5-1 5.2. Analytical Input . ..................... 5-2 5.3. Changes in Nuclear Design .............. ... 5-2 6.
){ THERMAL-HYDRAULIC DESIGN . . . ........ .......... 6-1 6.1. Thermal-Hydraulic Design Evaluation ............ 6-1 i
4 ,.* 6.1.1. Increased RCS Flow . . . . . . . . ......... 6-1 6.2. DNBR Analysis ................... .... 6-2 L 6.3. Pressure-Te=perature Limit Analysis ........... . 6-2 6.4. Flux-to-Flow Setpoint Evaluation . . . . . . . . . . ... . 6-2 s
- 7. ACCIDENT AND TRANSIENT ANALYSIS ................. 7-1 7.1. General Safety Analysis .. ................ 7-1 7.2. Rod Withdrawal Accidents . . ........... ..... 7-1 7.3. Moderator Dilution Accident ................ 7-2 7.4. Cold Water (Pu=p Startup) Accident . ............ 7-3
- 7. 5. Loss of Coolant Flow . .............. ..... 7-3 7.6. Stuck-out, Stuck-In, or Dropped Control Rod ........ 7-4
- 7. 7. Loss of Electric Power . . ................. 7-4 7.8. Steam Line Failure . . . . ................. 7-5 1
7.9. Steam Generator Tube Failure . . .............. 7-5
\479 \25 Babcock & \Vilcox L
CONTENTS (Cont'd)_ I'
.- Page
.. . 7-5 Fucl Handling Accident . . '. .' . . . . . . , , , . . .
' 7.10. . .. 7-5 .
7.11. Rod Ejection Accident ........,.............. ... ... ,. 7-6 7.12. Maximum Hypothetical Accident 7-6 ( 7.13. Waste Gas Tank Rupture . . ...... ... ...... 7-6 7.14. LOCA Analysis ... . . . . . .
.,.... ,. 8-1
- 8. PROPOSED MODIFICATIONS TO TECHNICAL SPECIFICATIONS 9-1
- 9. STARTUP PROCFJLM - PHYSICS TESTING . . . . . . . . . . . . . . . . .
A-1
'. REFERENCES ...
P#
- r. List of Tab _es_
I Table .,,, 4-5 I I- 4-1. Fuel Design Parameters ...........,.,,...., ,, 4-5
> 4-2. Fuel Rod Nominal Dimensions . . . . . . . . . , , . . .
6 4-3. Input Su:=narf for Cladding Creep Collapse Calculationa ~ 4-6 F Batch 3 . . . ... . . . . . . . . ........
. . .... . , 4-6 4-4. Fuel Thermal Analysis Parameters ..., 5-3 5-1. TMI 1 Cycle 3 Physics Parameters . ... .. ... 5-5
; :- 5-2. Shutdown Margin-Calculation for TMI 1, Cycle 3 .,,..... 7-7 7-1. Thermal-Hydraulic Design Conditions . . . . ... ....... 7-8 t 7-2. Comparison of Key Parameters for Accident Analysis
...... ....., ,, 7-8 1 7-3. Allevable LOCA Peak Linear Heat Rate 6.
e L List of Figures _ W Figure
. .... 3-3 I
3-1. TMI 1. Cycle 3 Loading Diagram . . .......... 3-4 3-2. TMI 1 Enrich =ent and Burnup Distribution for Cycle 3 . . . . . . 3-5 J, 3-3. THI 1. Cycle 3 Control Rod Locations........... ...........,.. 4-7 4-1. Mw4mm Gap Size Vs Axial Position . . . . 4-8 4-2. Power Spike Factor Vs Axial Position . . . . . , . . . . . . . . .
' ' 5-1. Two-Dimensional Relative Power Distribution - BOC ... .. 5-6 (4 EFPD), Cycle 3 .................. .... 8-2 8-1. Core Protection Safety Limits. Pressure Vs Temperature . . . . 8-3 8-2. Core Protection Safety Limits, Reactor Power Imbalance . ... 8-4 8-3. Core Protection Safety Bases . . . . . .... ......
4- 8-4. Protection System Maximum Allowable Setpoints, Pressure .. ... 8-5 Vs Temperature . . ......... ......... r 1479 126 _ ,;1 Babcock &Wilcox I
d L List of Figures (Cont'd) Page s 8-5. Protection System Maximum Allowable Setpoints, Reactor Power Imbalance . . ...................... 8-6 8-6. Red Position Limits for Four-Pu=p Operation, O to 100 i 1G EFFD, Cycle 3 . . . . . . . . .,. . . . . . . , . 8-7,, f 8-7. Rod Positiva Limits for Four-Pump Operation, j 100 i 10 to 246 i 10 EFPD, Cycle 3 ......... ..... 8-8
' 8-8. Rod Position ~ Limits for Four-Pump Operation After 246 t 10 ETPD, Cycle 3 ... ............ ..... 8-9 i
8-9. Rod Position Limits for Two- and Three-Pump Operation, s O to 100 t 10 EFPD, Cycle 3 . . . . . . . . . . . . . . . . . . '8-10 8-10. Ro.d Position Limits for Twi:)- and Three-Pump Operacion,
" 100 t.10 to 246 i 10 EFPD, Cycle 3 .............. 8-11 8-11. Rod Position Limits for Two- and Three-Pump Operation After 246 i 10 EFPD, Cycle'3 ........... ,,,. , 0-12
.- 8-12. Operational Power Imbalance Envelope for Operation From 0 to 100 i 10 EFPD, Cycle 3 ......... ...... 8-13 h 8-13. Operational Power Imbalance Envelope for Operation From 100 t 10 to 246 i 10 EFPD, Cycle 3 ........ ...... 8-14 8 8-14. Operational Power Imbalacce Envelope for Operation After
,, 246 i 10 EFPD, Cycle 3 .. .. ............... . 8-15 8-15. APSR Position Limits for Operation From 0 to 100 1 10 E.PD, Cycle 3 . . . . . . . . . . . . . . . . . . . . . , , . . . . . B-16 8-16. APSR Position' Limits for Operation From 100 i 10 to 246 3
i 10 LFPD, Cycle.3 ............... ,,,,,. 8-17 8-17. APSR Position Limits for Operation Af ter 246 i 10 . EFPD, Cycle 3 . . . ... . . . . . . . ... . . . . . . , , . . 8-18 L w bn t 1 i
\'
b ll r 1479 127 li a
- iv . E,tscock a.Wilcox I
- 1. INTRODUCTIO I AIID SUIO!ARY This report justifies operation of the Three Mile Island Nuclear Station Unit 1, cycle 3, at a rated core power of 2535 MWt. Included are the required analyses, as outlined in the US:lRC document, " Guidance fvr Propo- License Anendments Relating to Refueling," June 1975 To support cycle 3 operation of the TMI-1, this report employs analytical techniques and design bases established in reports that have been submitted and received technical approval of the USNRC (see refer-ences).
Cycle 3 reactor paraceters that are related to power capability are sunmarized in this report and referenced to cycle 2. All of the accidents analyzed in the FSAR have been reviewed for cycle 3 operation and in cases where cycle 3 charac-teristics proved to be conservative with respect to those analyzed for cycle 1 operation, no new analysis was performed.
! The Technical Specifications have been reviewed, and the modifications required I for cycle 3 operation are justified by this report. Based on the analyses per-g formed, which take into account the postulated effects of fuel densification and the ECCS Final Acceptance Criteria, it has teen concluded that TMI 1, cycle 3, can be safely operated at the rated core power level of 2535 MWt.
I l-1 3abecek & Wilcox 1479 128
~. .. . . . . { . s - L L e,
- 2. OPERATING HISTORY 6.
i s TMI 1 achieved initial criticality on June 5,1974, and power generation com-menced on June 15, 1974.. The 100% power icvel of 2535 MWt was reached on August 3, 1974. A control rod interchange was performed at 256 effective full l' power days (EFPD). Cycle 1 was completed on Febrtiary 21, 1976, after 466 EFPD. 3, The reference second cycle achieved criticality on May 24, 1976, and attained r 100% power on June 1, 1976. No control rod interchange was planned for cycle 2, which is ccheduled for coupletion in =id March 1977 after 253 1 10 EFPD. No operating anomalies occurred during the first two cycles that would adverse-ly affect fuel performance during the third cycle, t.
- Operation of cycle 3 is ccheduled to begin in May 1977. The design cycle length is 270 1 10 EFPD, and no control rod interchanges are planned.
I, L 7, 1 b L 1
;. 1479 129 f 2-1 Babcock & Wilcox
n. F
-e
- 3. GENERAL DESCRIPTION p.
1 L The TMI 1 reactor core is described in detait in section 3 of the Final Safety y Analysis Report for that Unit.1 The cycle 3 core consists of 177 fuel as-( semblies, oach of which is a 15 by 15 array containing 208 fuel rods,16 con-trol rod guide tubes, and one incore instrument guide tube, The fuel rods in batches la and 3 have an undensified nominal active length of'144 inches, while batches 4 and 5 have undensified lengths of 142.6 sad 142.25 inches, respec- [ tively. All fuel assemblies in cycle 3 maintain a constant nominal fuel load-(t
- ing of 463.6 kg of uranium. The cladding is cold-worked Zircaloy-4 with an J
OD of 0.430 inch and a vall thicknece of 0.0265 inch. The fuel consists of . i dished-end, cylindrical pellets of uranium dioxide (see Table 4-2 for data). Figure 3-1 is the core loading diagram for THI 1, cycle 3. The initial en-p richments of batches la, 3, 4, and 5 were 2.06, 3.05, 2.64, and 2.85 wt % ura-nium-235, respectively. All of the batch 2 assemblies will be discharged at the end of cycle 2, and the batch 3 and 4 assemblies will be shuffled to new locations. The batch 5 assemblies will occupy the periphery of the core, and batch la assemblies will occupy 13 interior core locatione, Figure 3-2 is an eighth-core map showing each assenbly's burnup at the beginning of cycle 3 and its initial enrichment. Reactivity control is supplied by 61 full-length Ag-In-Cd control rod asse=blies (CRAs) and soluble boron shim. In addition to the full-length control rods, eight axial power shaping rods ( APSRs) are provided for additional control of axial power distribution. The cycle 3 locations of the 69 control rods and the group designations are indicated in Figure 3-3. Tne core locations of the 69 control rods for cycle 3 are identical to those of the reference cycle
- 2. The group designations, however, differ between cycle 3 and the reference cycle in order to minimise power peaking. No control rod interchanges and no 1479 130 3-1 3abcock & Wilcox
burnable poison rods are necessary for cycle 3. The nominal system pressure is 2200 psia, and the densified nominal linear heat rate is 5.71 kW/ft at the rated core power of 2535 IWt. The heat rate is slightly higher than in the reference cycle due to the shorter stack height of batch 5 1479 131 3-2 3abcock & Wilcox
._ _ ..-.. =
r-Figurc 3-1. TMI 1, Cycle 3 Loading Diagram 1 Fuel Transfer Canal
. . 5 5 5 5 5 ,,
I" 5 5 5 4 1A 4 5 5 5 B 6- P-11 H-11* P-5
~~
5 3 3 1A 4 4 4 1A 3 3 5
,- . C 0-12 0-11 0-9* N-9 P-6 N-7 O-7* 0-5 0-4
~~
- 5 3 3 4 4 4 3 4 4 4 3 3 5 0
N-13 K-9 P-12 R-9 R-10 P-8 R-6 R-7 P-4 K-7 N-3 5 3 4 4 3 3 4 3 3 4 4 3 5 i E
' M-13 N-14 0-13 N-11 L-9 P-10 L-7 N-5 0-3 N-2 M-3 5 5 1A 4 3 3 3 3 3 3 3 4 1A 5 5 r F K-13* K-15 M-12 M-11 N-10 N-8 N-6 M-5 M-4 K-1 K-3*
% 5 4 4 4 3 3 3 4 3 3 3 4 4 4 5 H-14 K-12 L-15 K-10 L-12 H-13 R-8 O-8 L-4 K-6 L-1 K-4 M-2 l
5 u 4 3 4 3 4 u 4 3 4 3 4 u 5 H If - .
-Y E-11* L-14 H-14' F-14 H-12 H-15 L-128 H-1 H-4 L-2 H-2 F-2 M-5*
I. 5 4 4 4 3 3 3 4 3 3 3 4 4 4 5 1 < K
} Z-14 C-12 F-15 C-10 T-12 C-8 A-8 H-3 F-4 G-6 P-1 C-4 E-2 o 5 5 1A 4 3 3 3 3 3 3 3 4 1A 5 5 L .
, C-13* C-15 E-12 ?,-11 D-10 D-8 D-6 E-5 E-4 C-1 C-3*
S 3 4 4 3 3 4 3 3 4 4 3 5 M
- E-13 D-14 C-13 D-11 F-9 B-6 T-7 D-5 C-3 D-2 E-3 5 3 3 4 4 4 3 4 4 4 3 3 5
<- g
~~
D-13 C-9 B-12 A-9 A-10 B-8 A-6 A-7 5-4 C-7 D-3 i 5 3 3 1A 4 4 4 1A 3 3 5 G-C-12 C-11 C-9* D-9 s-10 D-7 C-7* C-5 C-4 5 5 5 4 u 4 5 5 5 B-11 E-5* 3-5 - R
' I
. Z
~
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15
'w V
4 *-- Batch *' Indicates cycle 1 location. j B-11 +- Previous Core Location c 1479 132 3-3 Babcock & Wilcox
f t r Figure 3-2. TMI 1 Enrichment and Burnup Distribution for Cpcle 3
,- 8 9 10 11 12 13 14 15 2.06 2.64 3.05 2.64 3.05 2.64 2.06 2.85 ..
e 14,603 5,720 21,361 8,176 18.082 8,176 15,754 0 2.85 3.05 3.05 3.05 2.64 2.64 2.64 K 21,654 20,297 22,430 4,836 9,485 7,682 0 3.05 3.05 2.64 2.06 2.85 2.85 L 18,743 20,522 5.579 15,514 0 0 r 2.64 2.64 3.05 2.85 H 6,009 5,~520 17,941 0 g w. 3.05 3.05 2.85
'" N 22,137 16,809 0 ?
2.85 0 b 0 A P . R s 4 b f 2.06 + Initial Enrichment 14,603 e BOC Buruup, HWd/mtU f 1479 133 c 3-4 Babcock t.Wilcox {
qF 4 1' t Figure 3-3. TMI 1. Cycle 3 Control Rod Locations L X r i A 2 7 2 i B 5 5 3 C 3 q-6 8 7 1 7 8 0 P 4 1 1 4 3 E 3 6 8 2 2 8 6 3
- F lL. 1 4 4 1 5
. G 5 2 3 6 7 -y t H W- 7 4 3 4 4 1 5 5 1 I K q.
8 6 3 6 8 2 2 f L 4 1 1 4 3 I 'M 3 7 8 6 8 7
- l. '
3 5 5 3 0
. ~ .
2 3 -2 P R
' I Z
i
! l<' 1 2 3 4 5 4 y S. 9 10 11 12 13 14 l15 e.
l Croup No. of sQ function 1 8 Safety
! j' 2 9 Safety I' 3 12 Safety
+ Croup Number 4 8 Safety 7
5 8 Control . i! 6 8 Control 7 8 Control
8 8 APSRs i
Total 69 i ll
,d hli 1479 134 l 3-5 Babcock & Wilcox
'f
- k. FUEL SYSTFl! DESIGH 4.1. Fuel Assembly Mechanical Design Pertinent fuel design parameters are listed in Table b-1. All fuel assemblies are identical in concept and are mechanically interchangeable. Thc cycle 3 reload fuel (batch 5) relative t'o the reference cycle 2 reload fuel (batch h) incorporates minor design modification to the spacer grid corner cells, which reduces the potential for spacer grid hangup during handling. This modification does not alter the structural integrity of the spacer grid design. In addition, the results of dynamic impact :esting of the spacer grid design show that the grids have a higher seis=ic capability than that shown in the static tests previ-ously documented.2 All other results presented in the FSAR fuel assembly mechan-ical discussion are applicable to the cycle 3 reload fuel assemblies.
s. 4.2. Fuel Rod Design Pertinent fuel rod dimensions for cycle 3 fuel are listed in Table 4-2. The e rechanical evaluation of the fuel rod is discussed below. 4.2.1. Cladding Collapse e Creep collapse analyses were performed for three-cycle assembly power histo-s ries for THI 1. The batch 3 fuel is more limiting than batch 4 and 5 fuel due to the lower prepressurization, lower pellet density, and previous in-core exposure and power history. Batch 3 is also more limiting than the 13 batch la assemblies e that will be inserted in cycle 3 for their second burn due to the greater in-core exposure time for batch 3. Table 4-3 is a summary of the input to the analysis. The batch 3 assembly power histories were analyzed and the most
- limiting assembly determined. Actual operating history was used throughout cycle 1. This included tne initial power oceration at LO and 80" core pcwer.
The predicted assembly power history for the mest limiting assembly was used to determine the most limiting collapse time as described in BAW-10084P, Rei. 1.3 The following conditions were analyzed for the worst assembly power history. In all cases, the 2000-hour densification assumption described in reference { 1479 135 4-1 Babcock t.Wilcox
( r i . 6 3 was used since it was found to be the most severe. The coolant an'd cladding c temperatures and fast flux values were calculated at axial locations corre-t ,, sponding to the conditions listed below. I' l 1. Assembly outlet conditions
- 2. Axial power peak of 1.0 in upper part of core.
I_
- 3. Maximum axial peak in upper part of the core averaged over three cycles of operation. ,
m e The third condition above was found to be the most limiting. The conserva-f tisms in the analytical procedure'are summarized below. N 4"
- 1. The CROV computer code was used to predict tha time to collapse. .CROV conservatively predicts collapse times, as demonstrated in reference 3.
- 2. No credit is taken for fission gas release. Therefore, the net differ-r ential pressures used in the analysis are conservatively high.
- 3. The cladding thickness used was the.LTL (lower tolerance 31mit) of the as-built measurements. The initial ovality cf the cladding used was the UTL J,. (upper toicrance limit) of the no-built measurements. These values Jere I
taken from a statistical sampling of the cladding.
! The most limiting assembly in batch 3 was found to have a collapse time longer s
than the nav4="= projected three-cycle core exposure of 2h,288 EFFH. This analysis utilized the assumptions on densification described in reference 3. 4.2.2. cladding Stress
~ ', Since the batch 3 fuel is the most limiting from a cladding stress point of view because of its low prepressurization and low density, the calculations e
performed in the TMI 1 Fuel Dewification Report are the most limiting. 4.2.3. Cladding Strain The fuel design criteria specify a limit of 1.0% on cladding cirewnferential plastic strain. The pellet design for each batch is set so that the plastic j cladding strcin is less than 1% at 55,000 KJd/mtU. The conservatists in this analysis are as follows:
- 1. The maximum specification value for fuel pellet diameter was used. '
- 2. The maximum specification value for fuel pellet density uas used.
- 3. The cladding ID used was the lowest permitted specification tolerance.
- 4. The maximum expected three-cycle local pellet burnup is less than 55,000 HWd/mtU.
L 1479 136 4-2 Babcock a.Wilcox
k ( 4.3. Thermal Design , All fuel assemblies in this core are thermally and geometrically similar. The variation in linear heat rate capability between fuel batches results primarily from the different specified fuel densities (Table 4-4). Linear heat rate ca-pabilities are based on centerline fuel melt and were established using the t TAFY-3 code 5 with full fuel densification penalties. r 4.3.1. Power Spike Model
,. The power spike model used for the cycle 3 analysis is the same as that used in the reference cycle 2.6 The ==v4=n= gap size yersus axial position is chown
~
in Figure 4-1, and the power spika factor versus axial position is nhown in Figure 4-2. The calculated power spike and gap size were based on 92.5% fuel density (TD) and an enrichment of 3 wt % uranium-235. The corresponding values for batch 4 and 5 fuel would be smaller because of the increased density and lower enrichment of these batches. 4.3.2. Fuel Temperature Analysis Thermal analysis of the fuel rods assumed in-reactor deneification to 96.5% TD. The analysis used is the same -as that documented in references 4 and 6, cycles 1 and 2 recpectively. The average fuel temperatures shown in Table 4-4 for ths various fuel batches sre taken from the analyses which define the linear heat rate (LHR) capability for each fuel type. The analysis of batches e _ la and 3, which incorporated "as-built" fuel data, is presented in reference
, _ 4. Batches 4 and 5 were analyzed with the same methods but these analyses
, were based on the lower tolerance limit (LTL) of the fuel density _ specifica-tion. All fuel meeting specification requirements would thus be expected to exhibit a LHR capability (LF.R to centerline, fuel melt) greater tnan or equal to that shown in Table h-4.
4.4. Material Design The chemical compatibility of all possible fuel-cladding-coolant-assembly in-
, teractions for the batch 5 fuel assemblies is identical to that of the present
, 4.5. Operating Experience 1479 137 B&W's operating experience with the Mark B 15 by 15 fuel assembly design has verified the adequacy of the fuel assembly design. As of September 1, 1976, the following operating experience has been amassed for the six B&W 177-fuel
, assembly plants using the Mark B fuel assembly:
4-3 Babcock & Wilcox
e .
. Max assembly Cu=ulative Current burnup,
. net electrical
;. 3 Reactor cycle Wd/mtU output, ifIn I Oconee 1 3 21,304 13,826,959 ,,
oconee 2 2 19,020 8,687,947
' 19,163 8,745,066 Oconee 3 1
, n!I 1 2 20,777 10,105,588 i Arkansas one 1 15,450 6,922,082 I Rancho Seco 1 7,553 2,042,758 l.ai- .
I 1 r-l _ f b i 1 1479135 l - 4_4 Babcock &Wilcox
Table k-1. Fuel Design Parameters Residual fuel asse=blies New fuel assemblies Eatch la 3atch 3 Batch k bat:h 5 Fuel assembly type Mark 3-2 Mark E-3 Mark Bh Mark Eh No. of assemblies 13 60 56 48 Initial fuel enrich, 5 23s U 2.06 x 05 2.6h 2.85 Initial fuel density, 5 TD 92.5 92.5 93.5 9h.0 3atch burnup, BOL, mwd /=tU 15,51S 19,869 6734 0 Cladding collapse time, EFPH >30,000 >30,000 N30,000(a) N30,000(a)
- Initial fill gas pressure (min spec) psia Table h-2. Fuel Red Nominal Dimensions Residual fuel New fuel Batch la Batch 3 Batch h batch 5 Fuel rods OD, in. 0.k30 0.h30 0.h30 0.430 ID, in. 0.377 0.377 0.377 0 377 Fuel pellet OD, in. 0.370 0.370 0.370 0.3695 Density, 5 TD 92.5 92.5 93.5 9h.0 Length, in. 0.T 0.7 0.7 0.6 Undensified active fue, lhh ihh 142.6 1h2.25 length, in.
Flexible spacers, type Corrugated Corrugated Spring Spring spacer spacer Solid spacers, material Zr0 2 Zr0 Z#~h Z#~h 2
" A detailed three-cycle collapse analysis will be performed fer subsequent reload reperts. A cladding collapse time of N30,000 hours is a preliminary estimate based on a ec=parisen of batch k and 5 design parameters with octch 3.
- Proprietary lkfO 5 4-5 Batecek & Wil:ox
Table h-3. Input Su=.ary for Cladding Creep Collapse Calculations - Batch 3 Pellet OD (mean specified), in. 0.3700 Pellet density (mean specified), % TD 92.5 Densified pellet dic=eter, in. 0.3650 Cladding ID (mean specified), in. 0.377 l Reactor system pressure, psia 2200 l Stack height (undensified), in. ik4.0
- Cladding Ovality, (UTL), in.
- Cladding Thickness, (LTL), in.
- Prepressure (tin. specified), psia
- Post Densification Prepressure (cold), psia l
i Table 4-4. Fuel Thermal Analysis Parameters Batch la Batch 3 Datch 4 Batch 5 lf ( Initial density, % TD 92.5 92.5 93.5 94.0 l Pellet diameter, in. 0.370 0.370 0.370 0.3695 ~. Noninal stack height, in. 144.0 144.0 142.6 142.25
, Densified Fuel Parameters (Densification to 96.5 % TD Assumed)
- Pellet diameter, in. 0.3640 0.3640 0.3645 0,3646
, Fuel stack height, in. 141.12 141.12 140.46 140.47 q Nominal LER at 2568 MWe, kW/ft 5.77 5.77 S.80 5.80 gr Avg fuel temp at nominal LER, P 1335 1335 3320 1315 LHR to ( fuel melt, kW/ft 19.6 19.6 20.15 20.15
- Proprietar/
1479 140 h-6 Babecek & Wilecx
- " - - & M n n r"--"I m M n n .9 m. M M 9 9 Figure 4-1. Maximu:n Cap Size Vs Axial Position 4.0 TDF = 96.5% TDI = 92.5% e = 3.00%
, 't . 0 -
e is w
" 2.0 -
c N s E
- o. 2 L 2 -
1.0 i e i t t t 4
- 0 20- 40 60 80 100 120 140 h N Axial Position, inches txt
- su O
g a - x - 9* w e
~~
% . _ ._ . ~ .a gf6h@d*%'Q'%#R% '" , ~~ ' Figure 4-2. Power Spike Factor Vs Axial Position 1.10 TDF = 96 5% , TDI = 92.5% e = 3.00% 1.08 O u C
. g 1.06 -
c. i bc. cn a k
- 1.04 -
o. I O 1.02
- m su er 0 -
Q- 4 1,00 140 N O 20 40 60 80 100 120 c= E Axial Position, inches g - n , bd . . e
~
b.
- 5. NUCLEAR deb 1GN k--
5.1. Physics Characteristics i, Table 5-1 compares the core physics parameters of cycles 2 and 3; these values for both cycles vere generated using PDQ07. Since the core has not yet reached an equilibrium cycle, differences in core physics parameters are to be expected between the cycles. 8 The shorter design life of cycle 3 will produce a corresponding smaller cycle differential burnup than designed for the reference cycle 2.
- - The lower ac-
- j". cumulated everage core burnup at end of cyde 3 is due to the shorter cycle, partially offset by the precence of once-burr.cd batch la and 4 fuel and twice-t f burned batch 3 fuel. Figure 5-1 illustrates a representative relative power
}- distribution for the beginning of the third cycle at full power with equ111b-rium xenon and group 7 & 8 inserted.
The critical boron concentrations for cycle 3 ar6 approximately the same as those
. of the design reference cycle 2. The hot full-pever control rod worths are some-8-
what less in cycle 3 than in cycle 2;.nevertheless, they are sufficient to
, c:aintain the required shutdown margin, as indicated in Table 5-2. The dif-l ferences in the parameters between cycles 2 and 3 are due to changes in radial
- flux distribution, isotopics, and the dif ference in cycle lengths. The ejected
}
rod worths in Tabic 5-1 are the relmum calculated values. It is difficult to compare valuee between cycles or between rod patterns since neither the rod patterns from which the CRA is ejected nor the isotopic distributions are iden- [ tical. Calculated ejected rod worths and their adherence to criteria are con-h sidered at all times in life and at all power levels in the development of the t rod position limits presented in ection 8. The maximum stuck rod worth at the end of cycle for cycle 3 is similar to that for the design cycle 2 but higher at the beginning of the cycle. No adverse safety implications are associated with this higher verth since the adequacy of the shutdown e l cycle 3 stuck rod worths is demonstrated in Table 5-2. iz' I n '1 'h43 i 5-1 Babcock s.Wilcox
j The following conservatisms were applied for the shutdown calculations:
- 1. ' Poison material depletion allowance.
- 2. 10% uncertainty on net rod worth. ,
- 3. Flux redistribution penalty. ~~
Flux redistribution was accounted for since the shutdown analysis was calculated { using a two-dimensional model. The shutdown calculation at the end of cycle 3 is analyzed at approximately 246 EFPD. This is the latest time ( 10 EFPD) In
)
1' core life in which the transient bank is nearly fully inserted. After 246 EFPD the transient bank will be almost fully withdrawn, increasing the potential shutdovn margin'.
- t. The cycle 3 power deficits fron hot zero power to hot full power are slightly lower than those for cycle 2 due to the less negative moderator coefficients l# in cycle 3.
The differential boron and xenon worths are similar for cycles 2 j p. and 3. The effective delayed neutron fractions for cycle 3 show a decrease
- 'l with burnup, as also reported in the reference cycle 2. -
j 5.2. Analy_tical input 1 i The cycle 3 incore measurement calculation constants used for computing core powel distributions were prepared in the same manner as for the reference cy.cle. 5.3. Changes in Nuclear Design - There were no relevant changes in core design between cycle 3 and the reference cycle 2. The same calculational methods and design information were used to ob-tain the i=portant nuclear design parameters. The only significant operational I procedure change from the reference cycle 2 is the specification of APSR position limits, in addition to the usual regulating CRA position and power imbalance limits, based upon LOCA analyses. The APSR position limits will provide additional control of power peaking through an i= proved definition of the core power distri- { bution. The operational limits (Technical Specification changes) for cycle 3 are shown in Section 8. The FLAE computer code 7 '8 was used in setting the Technical Specification limits. 1479 144 8 i I 5-2 Babcock & Wilcox 4 i
(' , Table 5-1. TMI 1 Cycle 3 Physics Parameters
- Cycle e 3(a) Cycle 2(b)
, Cycle length, EFPD 270 296 - Cycle burnup, Wd/mtU 8,341 9,144 Average core burnup - EOC, Wd/mtU 18,352 J8,612 Initial core loading, mtU 82.1 82.1 Critical boron - BOC, ppm (no Xe)
}' HZP, groups 1-8 out (e) 1,317 1,350 HZP, groups 7 and 8 inserted 1,155 1,187
, HFP, groups 7 and 8 inserted (e) 998 J.004
, Critical boron - EOC, ppm (eq Xe)
, HZP, groups 1-8 out 380 390 h HFP, group 8 (37.5% ud, eq Xe) 84 46 Control rod worths - HFP, BOC, % Ak/k Group 6 1. J 8 1.17 6
Group 7 0.84 0.97 - f Group 8 (37.5% wd) 0.54 0.54 Control rod worths - HFP, EOC, % Ak/k J Group 7 J.32 2 Group 8 (31.5% vd) 1.13 0.50 c)(I ) 0.50 Hax ejected rod worth - HZP, % Ak/k( } - BOC 0.34 0.57 EOC 0.77 0.54 Hax stuck rod worth - HZP, % Ak/k
] BOC 2.k2 2.15 EOC 2.06 2.21 Power deficit, HZP to HFP, % Ak/k BOC -1.58 -1.64 EOC -2.15 -2.48 Doppler coeff - BOC, 10-5 (Ak/k/'F) j 100% power (0 Xe) -1.47 -J.49 J Doppler coeff - EOC, 10-5 (ak/k/*F)
- L 100% power (eq Xe) -1.51 -1.53 Moderator coeff - HFP, 10
, (Ak/k/*F)
BOC (0 Xe, 1000 ppm, groups 7, 8 ins) -0.91 1.06 L EOC (eq Xe, 17 ppm, Group 8 ins) -2.54 -2.63 Boron worth - HFP, ppm /% Ak/k BOC (1000 ppm) 107 108 EOC (17 ppm) 97 100 Zenon worth - HFP, % Ak/k BOC (4 EFPD) 2.59 2.61 EOC (equilibrium) 2.64 2.67 1479 145 5-3 Babcock & Wilcox
i Table 5-1. (Cont'd) - Cycle 3(# Cycle 2(b)
. Eff delayed neutron fraction - HFP -
l BOC 0.00584 0.00577 EOC 0.00524 0.00516 , (*} Based on 253 EFPD at 2535 We, cycle 2. ( Based on 466 EFPD at 2535 We, cycle 1. I")246 EFPD (d) Ejected rod worth for groups 5-8 inserted.
~' '
(*)HZP denotes hot zero power (532 F Tave); HFP denotes hot full power (579 F Tave). ji . I I ( l. b I j 1479 146 5-4 Babcock & Wilcox
Table 5-2. Shutdown Margin Calculation for TMI 1. Cvele 3 BOC,5 Ak/k EOC(" , 5 Ak/k Available Rod 'Jorth Total rod worth, HZP(b) g,59 g,gg k' orth reduction due to burnup of poison material -0.26 -0.32 Maximum stuck rod, HZP -2.k2 -2.06 Net worth 5.91 6.11 Less 10% uncertainty -0.59 -0.61 Total available vorth 5.32 5.50 Recuired Rod 'Jorg Power deficit, EFF to HZP 1.58 2.15 Max allovable inserted rod vorth 1.1h 1.22 Flux redistribution 0.h0 0.73 Total required worth 3.12 k.10 Shutdown Marcin Total available - total required 2.20 1.h0 Note: Required shutdown =argin is 1.00% Ak/k
" For shutdown =argin calculations this is defined as N2h6 F.FPD, the latest ti=e in core life at which the transient bank is nearly full-in.
(b)HZP denotes hot zero power (532o F Tave); HFP denotes hot full power (579 F Tave). 1479 147 I i 5-5 Babcock & Wilcox
J l Figure 5-1. Two-Dimensional Relative Power Distribution - BOC (4 EFPD), Cycle 'a (Full Power, Equilibrium Xenon, Groups 7 and 8 Inserted) 8 9 10 11 12 13 14 15 7 ! H 1.09 1.29 1.10 1.26 1.28 1.20 0.55 0.78 i
. K 1.12 1.05 1.11 1.35 1. 24 1.10 0.87 8
L 1.05 1.08 1.17 0.96 1.29 0.77 w I M 1.21 1.07 0.92 0.95 6 e.
?:
l' , 7 i N i, 0.51 - 0.63 0.55
!l ,'/ 0 0.48 l
l
; P
__ t y . R L I 7 + Inserted Rod Group Nu:nber 6 0.51 + Relative Power Density 1479 143 5-6 Babcock & Wilcox l 4
f l jr . r i r' . I
- 6. THERMAL-HYDRAULIC DESIGN 6.1. Thermal-Hydraulic Design Evaluation The thermal-hydraulic design evaluation for support of cycle 3 operation uti-lized the same methods and models described'in references 1, 4, and 6. The
' core configuration for cycle 3 includes 13 once-burned batch la fuel assemb-lies (Mark B-2), 60 twice-burned batch 3 assemblies (Mark B-3), 56 once-burneC batch 4 assemblies (Mark B-4), and 48 fresh batch 5 assemblies (Mark B-h). Dur-E ing cycle 3 operation the highest assembly power always occurs in either the m.
6. once-burned batch 4 or the fresh batch 5 fuel. Thus, the hot assembly is always a Mark B-4 fuel assembly. 8 Both the Mark B-2 and B-3 fuel assemblies have slightly greater resistance to flow than do the Mark B-4 assemblies. The cycle 2 analysis6 , used also for the cycle 3 evaluation, considered a core configuration of 117 B-4 and 60 B-3 assemblies, with the hot assembly assumed to be a Ebrk B-3 type. This analysis is conservative for cycle 3 because the predicted hot assembly coolant flow rate is less than would be predicted for a corresponding Mark B-4 hot assembly. s 6.1.1. Increased RCS Flow Reactor coolant flow data obtained during the first two cycles of operation verified that the system flow was at locat 108% of the (first core) design flow
$ rate.
For the cycle 2 and 3 thermal-hydraulic design analyses this increase in
- 6. system flow was conservatively chosen to be 106.5% of the cycle 1 design flow.
The incorporation of this increased flow in the thermal-hydraulic calculations was accompanied by a corresponding increase in the reactor coolant inlet tem-perature, from 554 to 555.6F. This increase in core inlet temperature is both necessary and realistic since the integrated control system maintains a con-stant reactor vessel average coolant temperature. Both the increased reactor coolant flow and inlet temperature are changes in calculational parameters only (relative to first core calculations) and do not represer.c changes in plant operation. 1479 149
,_1 .Bascocks. win x
P h e
, 6.2. DNBR Analysis
- The BAW-2 CHF correlation has been used for termal-hydraulic analysis of cycles 2 and 3. This correlation has been reviewed and approved by the NRC for use with the Mark B fuel assembly design'. The cycle 2 (and 3) DNBR analysis resulted in 8
a predicted mini =um DNBR of 1 919 at the 112% overpower maximum design condition 1 for undensified fuel. This analysis was based on a nominal active fuel length of lh inches, a _ The effect of fuel densification on minimum DNBR is primarily a result of the reduction in active fuel length, which increases the average heat flux. For
" this evaluation, the active fuel length of batch 4 fuel (140.46 inches densi-fied) was used, resulting in a reduction in the calculated minimum DNBR of 2.18% (from 1.919 to 1.877 at design overpower).
The potential effect of fuel rod bow on minimum DNBR is considered by incorpo-rating cuitable margins into DNB-limited RPS setpoints. 6.3. Pressure-Temperature Limit Analysis Pressure-temperature limits for cycles 1 and 2 were based on the assunption
' that one core barrel vent valve failed open, thus reducing by 4.6% the effec-tive coolant flow for heat transfer. An NRC staff evaluation 10 of operating g data from B&W plantsll relieved B&W of the vent valve penalty.
For cycle 3 operation the cycle 2 pressure-temperature limits have been retained as the basis for the variable low-pressure trip function. Additional analyses have s. been conducted without the vent valve penalty to demonstrate that there is sufficient margin to offset potential rod bow effects'. 6.4. Flux-to-Flow Setpoint Evaluation No change has bean made to the flux / flow trip setpoint for cycle 3 operation. The transient thermal-hydraulic analysis on which this setpoint was based is conservative wit.h respect to cycle 3 operation. This analysis assumed that all vent valves remain closed and included densification effects. The results of the analysis demonstrate sufficient DNBR margin to offset the maximum p e-
, dicted fuel rod bow penalty.
1479 150 9 6-2 ' Babcock & Wilcox
L f . d. f
's . ACCIDENT AND TRANSIENT ANALYSIS
.{
, 7.1. General Safety Analysis 1l*
Each FSARI accident analysis has been examined with respect to changes in cycle 3 parameters to determine the effects of the cycle 3 reload and to ensure that thermal performance during hypothetical transients is not degraded. i' l- Core thermal parameters used in the FSAR accident analysis were design operat-4 ing values based on calculational values plus uncertainties. First core values
., (FSAR values) of core thermal parameters are compared with the parameters used in cycle 3 analysis in Table 7-1. These are parameters common to all of the accident analyses presented herein. For each accident of the FSAR, a discus-
.f. sion of the accident and the key parameters are provided. A comparison of the key parameters (see Table 7-2) between the FSAR and cycle 3 is provided with the
[{ i accident discussions to show that the initial conditions of the transients are
; bounded by the FSAR analysis.
The effects of fuel densfiication on the FSAR accident results have been evalu-i ated and are reported in reference 4. Since batch 5 reload fuel assemblies con-I tain fuel rods whose theoretical density is higher than those considered in
. reference 4, the conclusions in that reference are still valid.
Calculational techniones and methods remain consistent for cycle 3 aralysis with those used for the FSAR. Additional DNBR margin is shown for cycle 3 as it was for cycle 2 by the use of the B&W-2 CHF correlation rather than the W-3 e . CHF correlation. No new dose calculations were performed for the reload report. The dose con-g siderations in the FSAR were based on maximum peaking and burnup for all core
- l. cycles; therefore, the dose considerations are independent of the reload batch.
7.2. Rod Withdrawal Accidents This accident is defined as uncontrolled reactivity addition to the core caused by withdrawal of control rods during startup or rated power cond%%. g L types of incidents were analyzed in the FSAR. 7-1 Babcock & Wilcox c
e. The important parameters during a rod withdrawal accident are Doppler coeffi-cient, moderator temperature coefficient, and the rate at which reactivity is p added to the core. Only high-pressure and high-flux trips are accounted for .. in.the FSAR analysis; multiple alarms, interlocks, and trips that normally pre-clude this type of incident are ignored. I - s For positive reactivity addition indicative of these events, the most severe
~1 results occur for BOL conditions. The FSAR values of the key parameters for
- BOL were -1.17 x 10-5 (ak/k/*F) 'or the Doppler coefficient, 0.0 (ak/k/*F)
[ for the moderator temperature o efficient, and rod group worths up to and in-cluding a 10% ak/k rod group worth. Comparable cycle 3 parametric values are (' -1.47 x 10-5 (ak/k/*F) for the Doppler coefficient, -0.91 x 10~4 (ak/k/*F) for , fF the moderator temperature coefficient, and maximum rod grcup worth of 8.597
.h ok/k. Therefore, cycle 3 parameters are bounded by design values assumed for I[,
the FSAR analysis. Thus, for the rod withdrawal transients, the consequences l_'k will be no more severe than those presented in reference 1 or 4. I f g ,, ) 7.3. Moderator Dilution Accident
'y Boron (in the form of boric acid) is used to control excess reactivity. The
, [l .
boron content of the reactor coolant is periodically reduced to compensate for
!. fuel burnup and transient xenon effects, with dilution water supplied by the g makeup and purification system. The moderator dilution transients considered
'o are the pumping of water with zero boron concentration from the makeup tank to the reactor coolant system (RCS) under conditions of full-power operation, hot shutdown, and during refueling. The key parameters in this analysis are the f
)
initial boron concentration, boron reactivity worth, and the moderator tempera-ture coefficient for power cases. e i For positive reactivity addition of this type,'the most severe results occur at BOC conditions. The FSAR values of the key parameters for BOC conditions were { ( 1200 ppm for the initial boron concentration, 75 pps/1% ak/k ecid boron reactivity
' worth, and 0.5 x 10" ak/k/*F for the moderator temperature coefficient.
l{* Comparable cycle 3 values are 998 ppm for the initial borou concentration, 82 pps/1% ak/k cold boron reactivity worth, and -0.91 x 10 (ak/k/*F) for the moderatcr temperature coefficient. The FSAR shows that the core and RCS are j , adequately protected during this event. Sufficient time for operator action
, to termintte thi.s transient is also shown in the FSAR, even with maximum 1479 152 i l- 7-2 Babcock &Wilcox
7 dilution and minimum shutdown margin. The predicted cycle 3 parameters of im-portance to moderator dilution transient are bounded by the FSAR design values; thus, the analysis in the FSAR is valid.
- 7. r. . Cold Water (Pump Startup) Accident The NSS contains no, check or isolation valves in the RCS piping; therefore,
) the classic cold water accident is r.ot possible. However, when the reactor is operated with one or more pumps not runaing, and these pumps are then started, the increased flow rate will cause the . average core temperature to decrease.
- If the moderator temperature coefficient is negative, reactivity will be added to the core and a power increase will occur.
H Protective interlocks exist, and administrative procedures are imposed to pre-I vent starting idle pumps if the reactor power is above 30%. However, tbese I
\' ' restrictions were not assumed, and two-pump startup from 50% power was analyzed as the most severe transient.
e, To maximize reactivity addition, the FSAR analysis used the most negative mod-I erator temperature coefficient of -3.0 x 10 ok/k/'F and least negative Doppler coefficient of -1.2 x 10-8 Ak/k/ F. The corresponding most negative moderator l-temperature er efficient and least negative Doppler coefficient predicted in - g, cycle 3 are -2.54 x 10 . Ak/k/'F and -1.47 x 10-5 Ak/k/*F, respectively. Since the predictei r e cle 3 moderator temperature coefficient is less negative and the Doppler coefficient is more negative than the values used in the FSAR, the transient results would be less severe than those reported in the FSAR. 7.5. Loss of Coolant Flow i 1 A reduction in the reactor coolant flow rate can be caused by mechanical fail-3 ures or a loss of clectrical powet to the pumps. With four independent pumps available, a mechanical failure in one pump will not affect operation of the others. With the reactor at power, the effect of loss of coolant ficw is a ( rapid increase in coolant temperature due to reduction of heat removal capabil-icy. This temperature increase could result in DNB if corrective action were ] not taken immediately. The key parameters for four-pu=p coastdown or a locked-f rotor incident are the flow rate, flow coastdown characteristics, Doppler coef-ficient, moderator temperature coefficient, and hot channel DNB peaking factors. The conservative initial conditions assumed for the densification report" were FSAR values of flow and coastdown, -1.2 x 10-5 Ak/k/*F Doppler coefficient, 14719 153 7_3 Babeeck a.Wilcox
(
+0.5 x 10'" Ak/k/*F moderator temperature coefficient, with densified fuel pow-er spike and peaking. The results showed that the DNBR remained above 1.3 (W-3) i for the four-pump coastdown, and the fuel cladding temperature remained below crJteria limits for the locked-rotor transient. ..
The predicted values for cycle 3 are -1.47 x 10 Ak/k/*F Doppler coefficient, ~
-0.91 x 10
- Ak/k/*F moderator temperature coefficient, and peshing factors as lI ,
shown in Table 7-1. Since the B&W-2 CHF correlation was used for cycle 3, and l' the predicted cycle 3 values are bounded by those used in the cycle 1 densifi-( cation report, the results of the cycle 1 analysis represent the most severe i consequences from a loss o'f-flow incident. I
'* 7.6. Stuck-Out, Stuck-In, or Dropped Control Rod
[ If a control rod is dropped into the core while operating, a rapid decrease in f neutron power would occur, accompanied by a decrease in core average coolant temperature. In addition, the power distribution may be distorted due to a new r control rod pattern. Therefore, under these conditions a return to rated power ! ,, may lead to localized power densities and heat fluxes in excess of design limi-tations. f, , i The key parameters for this transient are moderator temperature coefficient, the worth of the dropped rod, and local peaking factors. The FSAR analysis was based on 0.36% Ak/k and 0.h6". Ak/k rod worths with and eithout Xenon, respectively, and with a moderator temperature coefficient of -3.0 x 10 " Ak/k/0F, For cycle 3,
~
t the maximum-worth rod at power is 0.20", Ak/k and the moderator temp-
~
erature coefficient is -2.54 x 10 ' Ak/k/ F. Since the predicted rod worth is less and the moderator temperature coefficient more positive, the consequences of this transient are less severe than the FFAR results. i' 7.7. Loss cf Electric Power L Two types of power losses were considered in the FSAR: a loss cf load condi-tion, caused by separation of the unit from the transmission system; and a hypothetical condition that results in the complete loss of all tystem and unit J power except the unit batteries. L The FSAR analysis evaluated the loss of load both with and without turbine run-back. When there is no runback, a reactor trip occurs on high reactor coolant pressure or temperature. This case resulted in a non-limiting accident. The largest offsite dose occurs for the second case, i.e., loss of all electrical 7-4 Bab o & ilco
( l N powe execpt unit batteries and assuming operation with failed fuel and steam g generator tube leakage. These results are independent of core loading; there-fore, the results of the FSAR are applicable for any reload. 7.B. Steam Line Failure A steam line failure is defined as a rupture of any of the steam lines from the steam generators. Upon initiation of the rupture, both steam generators start to blow down, causing a sudden decrease in primary system temperature,
,. preasure, and pressurizer level. The temperature reduction leads to positive reactivity insertion (at EOL, the moderator temperature coefficient is negative), ,, and the reactor trips on high flux or low RC pressure. The FSAR has identified a double-ended rupture of the steam line between the steam generator and steam s
cop valve as the worst-case situation EOL conditions. g ( The key parameter for the core response is the moderator temperature coeffi-cient, which in the FSAR was assumed to be -3.0 x 10* Ak/k/*F. The cycle 3 i predicted value of moderator temperature coefficient is -2.54 x 10 " Ak/k/'F.
~
This value is bounded by those used in the FSAR analysis; hence- the results ( in the FSAR reprecent the worst situation. 7.9. Steam Generator Tube Failure P A rupture or leak in a steam gene:ator tube allows reactor coolant and associ-ated activity to pass to the secondary syste The FSAR analysis is based on r complete severence of a steam generator tube. The primary concern for this incident is the potential radiological release, w '-h is independent of core loading. Hence, the FSAR results are applicable to this reload. 7.10. Fuel Handling Accident The mechanical damage type f accident is considered the. maximum potential t[- source of activity release during fuel handling activity. The primary concern
?
is radiological releases, which are independent of core loading; therefore, the ruults of the FSAR are applicable to all reloads. 7.11. Rod Rjection Accident For reactivity to be added to the core core rapidly than by uncontrolled rod
, withdrawal, physical failure of a pressure barrier component ir. the control rod drive assembly must occur. Such a failure could cause a pressure differ-ential to act on a CRA and rapidly ej ec. the assembly from the core. This 1479 155 7-5 Babcot.'
x .Wilcox
1 4 I ( incident represents the most rapid reactivity insertion that can be reasonably postulated. The values used in the FSAR and the densification report at BOL
-5 ok/k/*F Doppler coefficient, +0.5 x 10" Ak/k/*F conditions of -1.17 x 10 moderator temperature coefficient, and ejected rod worth of 0.65% Ak/k repre .
sent the maximum possible transient. The corresponding cycle 3 parametric
~
values of -1.47 x 10 Ak/k/*F Doppler, -0.91 x 10 " Ak/k/'F moderator tem-perature coefficient, are both more negative than those used in reference 4,
,. and a maximum predicted ejected rod worth of 0.55% Ak/k ensures that the re-7 suits will be less severe than those presented in the FSARI and the Jensifi-I cation report".
,. 7.12. Maximum Hypothetical Accident There is no postulated mechanism whereby this accident can occur since it r-i would require a multitude of failures in the engineered safeguards. The hypo-(. . _' - thetical accident is based solely on a gross release of radioactivity to the I
,, reactor building and is independent of core loading. Therefore, the results i, reported in the FSAR are applicable for all reloads.
l ., 7.13. Waste Gas Tank Rupture The waste gas tank was assumed to contain the gaseous activity evolved from degassing all of the reactor coolant following operation with 1% defective
, fuel. Rupture of the tank would result in the release of its radioactive con-cents to the plant ventilation system and to the atmosphere through the unit vent. The consequences of this incident are independent of core loading; therefore, the FSAR results are applicable to any reload.
/
7.14. LOCA Ar.,alysis I
, A generic LOCA analysis for a B&W 177-fuel assembly lowered-loop NSS has been
, performed using the Final Acceptance Criteria ECCS evaluation model.12 This analysis is generic since the limiting values of key paraceters for all plants in this category were used. Furthermore, the average fuel te=perature as a i
function of the linear heat rate and the lifetime pin pressure data used in i the reference 12 LOCA limits analysis are conservative and/or identical compared to those calculated f9r this reload. Thus, the analysis and the LOCA limits [ reported in reference 12 provide conservative results for the operation of TMI 1, cycle 3. L' Table 7-3 shows the bounding values for allevable LOCA peak linear heat rates 1479 156 i for TE 1, cycir 3 fuel. 7-6 Babcock t.Wilcox
Table 7-1. Thermal-Hydraulic Design Conditions Cycle 1 1 Cycles 2 and 3 Power level, HWe 2568 2568 2200 2200 i System pressure, psia Reactor coolant flow, % design 100.0 106.5 ! Vessel inlet coolant temp -- 100% power, F 554.0 555.6 Vessel outlet coolant temp -- 100% power, F 603.8 602.4 Ref design radial-local power peaking factor 1.78 1.78 Ref design axial flux shape 1.5 cosine 1.5 cosine
' Active fuel length, in. Table 4 -4 Table 4-4 Avg heat flux (100% power), Btu /h-ft2 171,470 174,870 6 g
Maximum heat flux (100% power), 457,825 466,903 6 I Btu /h-ft2 (for DNBR calculation) CHF correlation W-3 B&W-2 Minimum DNBR (max. design conditions, 1.55 1.919 no densification penalties) (114% power) (112% power) { t.
, Hot channel factors i
1.011 1.011
/ Enthalpy rise 1.014 Heat flux 1.014 Flow area 0.98 0.98 Minimum DNBR with densif'n penalty 1.46 4 1.877("}.
(114% power) (112% power) g
- Cycle 3; 1.805 for cycle 2.6
;< s
,v -
i 1 i 1479 157
' Babcock & VVilcox 7_7
(
' l'-
Table 7-2. Comparison of Key Parameters for Accident Analysis F FSAR and ref Predicted cycle Parameter 4 value 3 value
-5 I
~~
Doppler coeff, 10 Ak/k/*F L BOL -1.17 -1.47
. EOL -1.33 -1.51
~
Moderator coeff, 10 " ak/k/*F
} BOL 40.5 -0.91 p EOL -3.0 -2.54 k All rod group worth, % Ak/k, HZP 10 8.59 I
I* Initial boron co'ne, ppm , EFP n., 1200 998
-{ Boron reactivity worth, cold, 75 82
,q- ppa /% Ak/k I1 Hax ojected rod worth, FJP, % Ak/k 0.65 0.55 fI Dropped rod worth, HFP, % Ak/k 0.46 - 0.20 2 r d's a Table 7-3. Allowable LOCA Peak Linear Ucat Rate 5 Core Allowable peak elev, ft _t.HR, kW/ft 2 15.5 !# 4 1,6. 6 6 18.0
/ 8 17.0 10 16.0 1
r k i 1479 158 l' 7-8 Babcock 3.Wilcox l . . _ . --
- 8. PROPOSED MODIFICA.'IO:iS TO TECHNICAL SPECIFICATI0:S The Technical Specifica; ions have been revised for cycle 3 operation. The changes were made as a esult of the follouing:
- 1. An analysis incorporating the effects of fuel rod bow on core parameters.
- 2. The quadrant tilt limit was revi._ from a maximum actual core tilt li=it of h.925 to a maximum actual core tilt limit of 3.kl%.
7,s used in setting the Technical Specification
- 3. The i.J}E computer code Limits.
- h. Specifying APSR position limits in addition to the usual regulating control rod position and power intalance limits. The AFSR position limits vill pr-vide additional control of power peaking through an improved definition of the core power distribution.
1479 159 3-1 3abcock & Wilcox
f,. L . J l - Figure 8-1. Core Protection Safety Limits, Prescure Vs, Te:::perature
, 2600 f.
i{' d I 2400 l(
- i. .
ACCEPTABLE
/ E OPERATION fy 2200
(
.: a
/~
t t - t . ., 5 ie 2000 ---
.u j , UNACCEPTABLE
! OPERATION e 1600 6 I r i 560 560 600 620 640 660 Reactor Outlet Temperature, f i r 1 i I i
, 1479 160 8-2 Babcock r.Wilcox
f r- , t
, Figure 8-2. Core Protection Safety Limits, I
Reactor Power Imbalance
' instmal Power Level, 5 i
l - 120 , (.33,1I2) -- h 10 ACCEPTABLE 4 PUMP Ks/ft Limit . 100 Kv/f t Limit .'[ OPIRAfl0N f" (.'58,80) ( 33.88.7) . . 80 (38.7) (+'8'8 } (+35,88.7) 1 3 h ~~ 80 ACCEPTABLE 3 & 4 PUMP OPERAfl0N
- s -- 70 1
(-58,84.7) (+68,84,1) 7- (.33,59.1) I " *
.. 65
- h ACCEPTABLE
}. '
-- 50 2,3 t. 4 PUk' (<-47,48)
(48,47) L- OPERAfl0X
.. 40 T*
(_ .- - 30 I 20 1
.. If J
l I f t t I I 80 40 20 0 20 40 80 Reactor Poser la.nalancs, 5 CURVE REACTOR COOLANT FLOT (in/nt) i 138.8 108 2 104.5 108 3 88.8 a 108 r ' ~ 1479 161 8-3 Babcock s.Wilcox
L Figure 8-3. Core Protection Safety Bases J
'1 l
l' 2400 I 1 j \
'; 2 2200 --
p, i E. (
, t 2 c U /
1 $~ 2000 / 2 w'3 a I l c { f800 y , ((', i 1800 -- [ 550 580 800 $20 840 880 [ Resotor Outlet Temperatura, 'F Power, Fumps Operatir.8 (Type 6 d' curn Emactor coolant now.10 lb/h t_ of Limit) [ 1 139.8 (1002)*' 112 Four (ItG1 Limit)
'q 2 104.5 (74.7%) 86 7 Three (Ist3R Limic) .
L 3 48.8 (49.23) 39.1 One/ Loop (Quality Limit) h *106.5% of Cycle 1 Design now.
\
i 1479 162 64 Babcock &Wilcox
. - _ - - . .. .~. _-
Figure 8-4. Protection System Maximum Allowable j Setpoints. Pressure Vs Temperature f ' ( 2500 1 j P = 2405 psig ( 2300 h E. ACCEPTABLE p T = 610 F ( OPERATION
~
2l00
, e.
g. I' h
+,
rh S o 1900 5' I [i g UMtCCEPTABl.E T-
- P = 1800 psig OPEkJ10H 1700 i
f 1500 i 540 560 ~580 600 620 640 Reactor Outlet Temperature, F 4 I I t 1479 163 8-5 Babcock a,Wilcox l
. . . . . - - . _ . = ._ _ __
6 Figure 8-5. Protection System Maximum Allowable Setpoints, Reactor Power Imbalance Power Levtl [ ,1, il - 120
?
(-20.108)
- 110 (108)(+21.108) 1, .
{ g% - - 100 g ACCEPTABLE O,3
+ 4 PUNP - -
90 * {-
+' OPERATION (80.7)
(-46.80) . '(48,80) 80
- c. I .- 70 ACCEPTABLE -
! l' ' 3 & 4 PUMP IL OPrairloh -- 60
'1 -
1 l (53.I) (-46.52.7) ,, (+48,S2.7)
) l --
l 40 l [ i
, ACCEPTABLE F; 2,3 8 4 PUMP - -
30 (-46.25.1) OPERAT10ll (+4 8,25.1 ) f '
< -- to i e a' -
10 ly 7 i 18 18
- l. g gi
.; 7 i i #1 i i _I r, i i e 1
50 -40 -30 -20 -10 0 10 20 30 40 50 i, ' Power labalance 5 1479 164 8-6 Babcock &Wilcox
s Figure 8-6. Rod Position Limits for Four-Pump Operation,
- k. ' 0 to 100 i 10 EFPD, Cycle 3 l
112.se in,se2 2n.s. 02 100 - OPERATION IN THl3 , REGICH 15 NOT ALLOIt0 to - In.2,so 225.s 70tER LEVEL CUT 0FF i 80 - RE ON 880. 80 * \ .so 233.s so
. RESTRIOTED
$HUTOCIN 39s.t.70 REGION 10 EARGIN
' tas.s.io aoo,n LIMIT .
* - . 12s.o,so.
f CO E ' N 50 - es.u . FERulS$18LE '1
% . OP(RATING i
- 40 ', / REEL 0N f
a 3g e
' s 20 -
N e,as , g ' 10 - 1 ' t t t t t i . t t : i, 9 25 50 75 800 125 150 175 200 225 250 275 200 jj Rod indez, 5 Ilthdrawn
' 25 9 i,u
[4 5,a Group 7 8 25 50 75 100
- i t i i i Group 6
[ ' 0 25 50 15 100
- t. I t t t troup 5 a
f N I I 1479 165 d e .[ 8-7 Babcock & Wilcox j
.i .
G Figure 8-7. Rod Position Limits for Four-Pump Operation, 100
. + 10 to 246 _+ 10 EFPD, Cycle 3 100 . OPERATION IN THl3 REGION f ,
i- 15 NOT ALLO 1E0 POWER LEVEL E CUT 0ff
;. 50 -
i n. .. .o , ,. . . . o 8"88' .(' I) 80 -
, iso.2,so RESTRICTED I -
REGION 1 ns.2,70 l ,,,,,, g 33,,,,,,
- e SHUT 001N MARGIN *
, g 60 (y its.o so 50 . es,so @ PERMIS$1BLE
,- OPERATING
; 40 - BEGION s
~
/
30 -
@ 20 -
68.15 7,,gg , 20 - o,0
- l I I f I f I t I i I 0 25 50 15 800 125 150 175 200 225 25C 215 300 Rod indet, 5 Ilthdrawn '
O' 25 50 15 500 g i f f f I
- Group 1 1 g 2,5 5,0 7) 10,0 itt Group 6 A
100 0, 25, 5,0 7,5 . Grotip 5 di k i - B L p f 1479 166
] 8-8 Babcock & Wilcox
i Figure 8-8. Rod Position Limits for Four-Pump Operation After 246 + 10 EFPD, Cycle 3 rso. lot 158.s.102 100 - OPERATION lH THl3 REGION IS NOT ALLOIED 00TO F ag e e lo - 80 . RESTRIOTED . 237.s.so h REGION
' ~
ise.70
~ SHUT 00TN BARGIN
$ ,,, 80 -
p = LIMIT
. 50 -
. 110.50 PERWl55lBLE ,,i i e e . OPERATlHG
; g 40 REGION
' !; 30 - -
? 20 -
I " so.ls 10 o.o l l l 1 *I I I I l I l' 1 8 25 50 75 800 125 150 175 200 225 250 775 300 Red Inges. 5 l.itharasal O 25 50 15
, I I i 10,0 Eroup 7 r 8 2,5 la 7) 10,0 sr..,s k g 5,0 15 100 2,5 6 Stoup $ -
p s i
~
i 1479 167 1 1 g_9 Babcock t.Wilcox
i J i Figure 8-9. Rod Position Limits for Two- and Three-Pump Oper-
; ation, O to 100 + 10 EFPD, Cycle 3 f5 .
812,802 lis let I t.let 250.102 !' i, 100 - OPERAilGM IN THl3 RESTRIGTED f(Ell 0N REGION 13 NOT ALLGIED FOR 3 Pulf CPERAllGN . 90 - sie,se { j tic,se. soo,es- , i' 3gg , SHU1001N E ILARGIN 2 g LIRli 81b78 k $70 -. p b ,h, i 3 60 - @y
, I, ; frRElsslBLE OPfRATING i
IS0 - ts.so . REGIO,N A 5
. i $- 340 a - -
~
4 l
'e P.estricted Perion for 2 & 3 w.p operation 20 0.15
' ~ ( 10 - 8' 8 l 'l I I I I I I ( t t 9 25 50 75 100 125 150 175 200 225 250 215 300 t Red lades, 5 tithdrasal 8 25 50 15 100
, t I t t t Group 7 i
, 0 25 50 75 100 t t t t Ersup 6 6 25 58 75 108 .
1 i I 9 j f p treup 5 t e b 1479 168 8-10 Babcock 8.Wilcox L
( k i Figure 8-10. Rod Position Limits for Two- and Three-Pump , Operation, 100 f 10 to 246 i 10 EFPD, Cycle 3 17s,102 140.102 100 - RESTRICTED REGl0F FOR 3 FUMP . CPERATION IN THl3 REGICH
- OPERATION, i
90 y, is NOT ALLDIED ,,,,,, ,,,,
, g 80 -
1, =. . . J 5 70 - . u
= 60 - FERulSSIBLE OPERATING 3- j SH'JT001N WARGiN . REGION = i u LIMli . N == 50 - e: .50 5 ,g *; 2 40 -
C j, t
., ; 30 -
in 20 -
- sc.15
.p f y10 -
- 0. 0 '
g I t t I t t t t t t t I 8 25 50 75 100 125 150 175 200 225 250 275 300 Red Irulsz, 5 titMrassi
~
8 .25 50 75 100 t t t t t Stoup 1 1 O 25 53 J5 800 t I t 1 g
," Gren; 3
' 100 S, 2,5 50 15 i 1
' steup 5 I
l. 1479 169 8-11 Banc7<:k & Wilcox
f 2 Figure 8-11. Rod Position Limits for Two- and Three-Pump Operation After 246 i 10 EFF0, Cycle 3 {, - 134 102 238,101 100 - OPERATION IN THIS REEL 0N
- 15 NOT ALLOYED RE31R10TED REGION FOR l
- to -
3 PUNP OPERATION E e so - . k $
- t E ~ SHUT 001N
[> [ EARGIN ;- ., E LIMIT FIREISSIBLE DPfRATING U.= 00 REGION
,\ u 2 50 - 188.58 .
3 h., 40 - 2 V 6 g 30 - [ 's e i20 -
) g
- ss,es
)
10 (. - I oo I 6 I I I t I I e I i t i I 25 50 15 100 125 150 175 200 225 250 275 300 I Red index, 5 Titttdrawn . 9 25 50 15 100
, i f I t I g Group 7 l
e e 2I 5e i ni 100 i Group 5 1 o I 25 t 50 I 75 t lost
)
Eroup 5 l L b I 1479 170. 8-12 Babcock s Wilcox
1' . I k Figure 8-12. Operational Power Imbalance Envelope for Operation From 0 to 100 + 10 EFPD, Cycle 3 Power, 5 of 2535 lat s' RESTRICTED REGION
~
-17.54,102 "
12.04,102
-100
-18.00,90 P
- 90 12.os,so l , , 41}0
- 70 '
PERMISSI'Bl.E 8 07ERATING-- 80 REGION
- 50
- 40 f
s
- - 30 P
- 20 .
- 10
, a i e . . . i e i 40 -30 -20 -10 0 10 20 30 40 50 Axial Power imbalance, 5
\' L 1479 171 8-13 Babcock & Wilcox
6
> Figure 8-13. Operational Power Imbalance Envelope for Operation From 100 ~+ 10 to 246 + "
10 EFPD, Cycle 3 ,
- 5 Power, 5 pf 2535 Mit RESTRICTED REGION .
/ -24.38,102 9.49,802 'r 1 - -100 l, -22.50.90 .. 90 9.49,90 s - 80 s
r '
. 70 ? '
PERMISSIBLE
, OPERATlHG 1
' - - 00 REGION
. -- 50
!' x: i ) ,
-- 40 ,
- 4. . .
-- 30 20
- - 10 i
, 4 i e' s e e 4 e r i
, 50 -40 -30 10 0 10 20 30 40 50 Axlal Power Imualance, 5 h
1479 172 8-14 Babcock & Wilcox
b 6 Figure 8-14. Operational Power, Imbalance Envelope for . Operation After 246 + 10 EFPD, Cycle 3 Power, 5 of 2535 Ut , , RESTRICThu REGION
-31.21,10 19.2s,102 i
- -100
-28.8,90 - 90 14.2s,90
, i
{
- 80 .
. PERMISSIBLE
'., OPERATING ~ ~ 70 ' REGION {
; . . 80 .
- 50
- 1
- - 40 .
l
- - 30 L .
P
- - 20 10 ,
, i e i I t a g g g g
-50 -40 -30 -20 -10 0 10 20 30 40 50 6 .
. Axial Power Imualince, 5 .
i479 175 p i 8-15 Babcock & Wilcox
I i Pigure 8-15. APSR Position Limits for Operation From I, O to 100 1 10 EFFD, cycle 3 . l
'; 18.'2,102 34.7,102 100 -
90 - 15.s,90 I 3 E.s,90 RESTRICTED
- REGION 80 -
s.s,s0 37.9,a0 i 70 to,70 s.s,70 )vs.2,70 g - t ' 60 - h N 45.2,60 100,60 g 50 - PERNISSIBLE { 40 - OPERATING E REGION
- 30. ,- -
r . 20 - 4 10 -
, , e e , , , , g j 0 10 20 30 40 50 60 70 80 90 100 APSR 5 fithdrawn 1479 174 8-16 Babcock s.Wilcox
Figure 8-16. APSR Position Limits for Operation From 100 i 10 to 246
- 10 EFPD, Cycle 3
' 20.6,102 38.8,102 100 -
r
, 90 -
e,,,go. is. s. n
', RESTRICTED 6.5,So u4.s,eo REGION 80- ,
I g 70 _ o,70 s.s,7o 45.2,70 i} - E a 60 _ 45.2,60 100,6'O i g i
- 50 _
t E f *- 4 0 - . PERillSSIBLE 30 - OPERATlHG l REGION L
; 20 _
10 _
> f t t t t t I t 1 0 10 20 30 40 50 60 70 80 90 100
?.
L APSR 5 Iltadrawn b L 1479 175 8-17 Babcock a.Wilcox
Figure 8-17. APSR Position Limits for Operation After '
' 246
- 10 EFPD, Cycle 3 l,
s.5,102 4s.0,102 100 - L l 2.5,90 RESTRICTED 90 REGION
- 80 ..a2.s,s0 0,80 .
L . 70 e
-E i: 80 -
f00,60 E 49.0,60 t U
; 50 - PERMISSIBLE a f OPERATING j 3 J 40 - REGION E' .
,- u 30 - .
~
.i .
20 - i . 10 - L I t t t I t f s t
' O 10 20 30 40 50 60 70 80 90 100
. APSR 5 WithM ann i ,
L h 1479 176 8-18 Babcock & Wilcox
P
- 9. STARTUP PROGRAM - PHYSICS TESTING The planned startup testing progra:n associated with core performance is out-1hed below. These ' tests verify that core performance is within the assump-
,c: -
of the safety analycia and provide the neccesary data for continued nafo, plant operation. Zero Power Testa
- 1. Critical boron concentration.
- 2. Temperature reactivity coefficient.
- 3. Control rod ~ group worth.
- 4. Ejected rod worth.
Power Testa
- 1. Core power distribution verification at approximately 40, 75, and 100% FP, normal control rod group configuration.
- 2. Core power distribution verification at approximately 40% FP with worst-case dropped rod fully inserted.
- 3. Incore/out-of-core detector imbalance correlation verification at approxi-mately 75% FP.
- 4. Power Doppler reactivity coefficient at approximately 100% FP.
- 5. Temperature reactivity coefficient at approximately 100% FP.
r-e i 147o ~7 91 Babcock s.Wilcox
. . . . ~ . _ . _ _ . . - . - . . -. . . . . . _ _ - ., _ _ - - _ . . . . ~ _
a f REFERENCES . p 1 Three Mile Island Unit 1 Nuclear Station, Final Safety Analysis Report, , USNRC Docket No. 50-289. 2 Fuel Assembly Stress and Deflection Analysis for Loss-of-Coolant Accident and Seismic Excitation, BAW10008, Part 2, Rev.1, Babcock & Wilcox, June 1970. 3 Program to Determine In-Reactor Performance of B&W Fuels - Cladding Creep
, Collapse, BAW-10084P, Rev. 1. Babcock & Wilcox, October 1976. ;" . s + :..
3 7tfI 1 Fuel Densificatiora Report, BAW1389, Babcock & Wilcox, June 197.3. J t.'#' g.m 5 3 n C. D., Morgan and H. S. Kao, TAFY - Fuel Pin Temperature and Gas Pressure
,. Analysis, BAW-10044, Babcock & Wilcox, May 1972.
6 Three Mile Island Unit 1 - Cycle 2 Reload Report, Babcock & Wilcox, July
; .e
[, 1976. ' ( ' 7 FLAME 3 - Three-Dimensional Nodal Code for Calculating Core Reactivity and
~
Power Distributions, BAW10124A, Babcock & Wilcox, August 1976. # 8 Verification of Three-Dimensional FLAME Code, BAW-10125A, Babcock & Wilcox, August 1976. 9 Correlation of Critical Ecat Flux in a Bundle Cooled. by Pressurized Water, g n BAW10000A, Babcock & Wilcox, June 1976. t ' 10 { anc Letter, A Schwencer to K. E. Suhrke, November 1975. - l 11 "B&W Operating Experience of Reactor Internals Vent Valves," Letter, K. E. Suhrke to A. Schwencer, August 4, 1975. 12 ECCS Analysis of B&W's 177-FA Lowered-Loop NSS, BAW-10103, Rev. 1, Babcock
& Wilcox, September 1975.
147,9 178 1 A-1 Babcock & Wilcox 1}}