ML19210A518
| ML19210A518 | |
| Person / Time | |
|---|---|
| Site: | Three Mile Island  |
| Issue date: | 01/13/1976 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| Shared Package | |
| ML19210A510 | List: |
| References | |
| NUDOCS 7910300557 | |
| Download: ML19210A518 (54) | |
Text
. .
THREE MILE ISIRID UNIT 1 -
CYCLE 2 RELOAD REPORT (Based on a Cycle 1
'Burnup of kk0+ 10 EFFD) 4 EABC0CK & UILCOX Power Concration Group 1.uclear Po 'cr Generation Division
}4h[[ P.O. Box 1260 Lynchburg, Virginia 24505 7910800867 I
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 4.2 Fuel Rod Design . . . . .'. .... .. .... .... . 4-1 4.3 Ther=al Design . . . . . .. . ... ...... ... . 4-3 4.4 Material Design . . . . . .. .... . ...... ... 4-4 4.5 Operating Experiences . . ... ............. 4-4
- 5. NUCLEAR DESIGN . . . . . . . . .. . ............. 5-1 5.1 Physics Characteristics .
.. .......... .... 5-1 5.2 Analytical Input . . . . .. .... ......... . 5-2 5.? Changes in Nuclear Design . ............... 5-2
- 6. THERMAL-HYDRAULIC DESIGN . . . .. .... ..... . . , . . . 6- l 6.1 Thermal'-Hydraulic Design Calculations . .. . . . '.' . . '. ' F-1 6.2 DNBR Analysis . . . . . . ...... ....... .. 6-2
- 7. ACCIDENT AND TRANSIENT ANALYSIS ... .. . ......... 7-1 7.1 Ceneral Safety Analysis . .. . ......... .... 7-1 7.2 Rod Withdrawal Accidents . . .... . .. ... . ... 7-1 7.3 Moderator Dilution Accident . . ... ...... ... . 7-2 7.4 Cold Water (Pump Startup) Accident . . ......... 7-3 7.5 Loss of Coolant Flou . . .. . . ... ........ . 7-3 -
\
7.6 Stuck-Out, Stuck-In, or Dropped Control Rod Accident .. 7-4 7.7 Loss of Electric Power . .. .. . .. ..... .. .. 7-4 7.8 Steam Line Failure . . . .. .... ......... . 7-4 7.9 Steam Generator Tube Failure . .. .. .... ..... 7-5 1487 228 , I f i i t t
.I e
CONTENTS (Continued) Page 7.10 Fuel Handling Accident . . . . . . . . . . . . . . . . . 7-5 7.11 Rod Ejection Accident . . . . . . . . . . . . . . . . . 7-5 7.12 Maximum Hypothetical Accident . . . . . . . . . . . . . 7-6 7.13 Waste Gas Tank Rupture . . . . . . . .. . . . . . . . . 7-6 7.14 LOCA Analysis . . . . . . . . . . . . . . . . . . . . . 7-6
- 8. PROPOSED MODIFICATIONS TO TECHNICAL SPECIFICATIONS . . . . . . &-1
- 9. START-UP PROGRAM.. . . .. . . . . . . . . . . . . . . . . . . 9-1
- 10. REFERENCES . . . . .. ... . . . . . . . . . . . . . . . . . 10-1 List of Tables Table 4.1-1 Fuel Design Parameters . . . . . . . . . . . . . . . . . . 4-5 4.2-1 Fuel Rod Dimensions .. . . . .. . . . . . . . . . . . . 4-5 4.2-2 Input Su= mary for Cladding Creep Collapse Calculations *. . 4-6
- 4. 3-1 Fuel Temperature Analysis Parameters for Cycles 1 and 2 . 4-7 5.1-1 TMI-1, Cycle 2 Physics Parameters . . . . . . . . . . . . 5-3 5.1-2 Shutdown Margin Calculation TMI-1, Cycle 2 . . . . . . . . 5-5 6.2-1 Cycle 1 and Cycle 2 Design Conditions . . . . . . . . . . 6-3 7.1-1 Comparison of Key Parameters for Accident Analysis (440 EFPD) . . . . . . . . . . . . . . . . . . . 7-7 7.14-1 Allowable LOCA Peak Linear Heat Rate . . . . . . . . . . . 7-8 List of Figures Figure 3-1 TMI-1, Cycle 2 Core Loading Diagram . . . . . . . . . . . 3-3 3-2 TMI-1, Enrichment and Burnup Distribution for Cycle 2 . . 3-4 3-3 TMI-1, Cycle 2 control Rod Locations . . . . . . . . . . . 3-5 1487 229
- iii - .
List of Figures (Continued) Figure Page A. 3-1 Maximum Cap Size Vs Axial Position . .. . . . . . .. .. 4-8 4.3-2 Power Spike Factor Vs Axial Position . . . . . . . . . .. 4-9 5.1-1 30C (4 EFPD), Cycle 2 Two-Dimensional Relative Power Distribution . . . . . .. . . .. . .. . ... .. 5-6 6-1 Core Protection Safety Limit .. . . . . .. . . . . . .. 8-2 8-2 Core Protection Safety Limits . . . .. . . . .. . . .. 8-3 8-3 Core Protection Safety Bases . . . . . . . . . . . .. .. B-4 S-4 Protection System Maximu Allowable Set Points . . . . .. B-5 8-5 Protection System Maximum Allowable Set Points . . . . .. 8-6 8-6 Rod Position Limits for 4 Pump Operation Applicable During the Period From 0 to 152 1 10 EFPD; Cycle 2 . . .. 8-7 8-7 Rod Position Limits for 4 Pump Operation Applicable During the Period From 152 + 10 to 275 1 10 EFPD; Cycle 2
, 8-8 8-8 Rod Position Limits for 4 Pump Operation Applicable During the Period After 275 1 10 EFPD; Cycle 2 . . .. .. 8-9 8-9 Rod Position Limits for 2 and 3 Pump Operation Appli-cable During the Period From 0 to 152 +10 EFPD; Cycle 2 8-10 S-10 Rod Position Limits for 2 and 3 Pump Operation Appli-cable During the Period From 152 + 10 to 275 + 10 EFPD; Cycle 2 . . . . . . . .. . . . . . . . . . . . ..
B-ll 8-11 Rod Position Limits for 2 and 3 Pump Operation Appli-cable During the Period Af ter 275 1 10 EFPD; Cycle 2 . .. 8-12 6-12 Operational Power Imbalance Envelope Applicable to - Operation From 0 to 152110 EFPD; Cycle 2 . . . . . . .. 8-13 8-13 Operational Power Imbalance Envelope Applicable to Operation From 152110 to 275 i 10 EFPD; Cycle 2 . . .. 8-14 6-14 Operational Power Imbalance Envelope Applicable to Operation Af ter 275 i 10 EFPD; Cycle 2 . . . . . . . . .. 8-15 l 1487 230 !
. (
- iv -
f I i
r .
- 1. INTRODUCTION AND SU'O!ARY This report justifies the operation of the Three Mile Island-Unit 1
, Nuclear Station, Cycle 2 at the rated core power of 2535 Wt. Included are the required analyses, as outlined in the USNRC docu=ent "G 4%ce for Propo' sed License A=end=ents Relating to Refueling" June 1975.
To support Cycle 2 operacion of the Three Mile Islanc Unic 1 Nuclear Station, this report ecploys analytical techniques and design bases established in reports which were previously subczitted and accepted by the USNRC (see References). A brief su= mary of Cycle 1 and 2 reactor para =eters that are related to power capability is included in this report. All of the accidents analyzed in the FSAR have been reviewed for Cycle 2 operation. In those cases where Cycle 2 characteristics proved to be conservative 5.-ith respect to those analyzed for Cycle 1 operation, no new analysis was perfor=ed.
- The Technical Specifications have bum reviewed and the modifications required for Cycle 2 operation are justified in this report.
Based on the analyses perfor=ed, which take into account the postulated effects of fuel densification and the Final Acceptance Criteria, it has been concluded that Three Mile Island-Unit 1, Cycle 2 can be safely operated at the rated core power level of 2535 W t. 1487 231 e 1-1 o O
- 2. OPERATING HISTORY Three Mile Island-Unit 1 achieved initial criticality on June 5,1974, and power generation commenced on June 15, 1974. The 100% power level of 2535 MWt was reached on August 3,1974. A control rod interchange was per-formed at 256 effective full power days (EFPD). The fuel cycle is scheduled for completion February 14,1976 af ter 440 + 10 EFPD. No operating anomalies occurred during the first cycle which would adversely affect the fuel per-formance during the second cycle.
Operation of Cycle 2 is scheduled to begin in late April 1976. The design cycle length is 296 EFPD and no control rod interchanges are planned. 1487 232 2-1
- 3. GENERAL DESCRIPTION The TMI-l reactor core is described in detail in Section 3 of the Three Mile Island-Unit 1 Nuclear Station, Final Safety Analysis Report (Reference 1).
The Cycle 2 core consists of 177 fuel asse=blies, each of which is a 15 by 15 array containing 208 fuel rods,16 control rod guide tubes, and one incore instrument guide tube. The fuel rods in batches 2 and 3 of Core I have an undensified nominal active length of 144 inches while batch 4 has an undensified length of 142.6 inches. All fuel assemblies in Cycle 2 maintain a constant nominal fuel loading of 463.6 kg of uranium. The cladding is cold-worked Zircaloy-4 with an OD of 0.430 inch and a wall thickness of 0.0265 inch. The fuel consists of dished end, cylindrical pellets of uranium dioxide which are 0.700 inch in length and 0.370 inch in diameter. (See Table 4.1-1 for additional data.) Figure 3-1 is the core loading diagram for TMI-1, Cycle 2. The initial enrichments of batches 2 and 3 vere 2.75 and 3.05 vt 5 235U , respectively. The 56 batch 4 assemblies are enriched to 2.6h vt % 235U All of the batch 1 assemblies will be discharged at the end of Cycle 1. The batch 2 and 3 asse=blies will be shuffled to new locations. The batch h asse=blies vill occupy primarily the periphery of the core and eight locations interior to the core. Figure 3-2 is an eighth-core map showing the assembly burnup and enrich =ent distribution at
,the beginning of Cycle 2.
Reaccivity control is supplied by 61 full-length Ag-In-Cd control rods and soluble boron shim. In addition to the full-length control rods, eight axial power shaping ods are provided for additional control of axial power distribution. The Cycle 2 locations of the 69 control rods and the group designations are indicated in Figure 3-3. The core it :ations of the total pattern (69 control rods) for' Cycle 2 are identical to that of the reference cycle indicated in Section 3 of the FSAR. The group designations, however, dif fer between Cycle 2 and the reference cycle in order to minimize power peaking. No control rod interchange and no burnable poison rods are necessary during Cycle 2. 1487 233 3-1
The nominal system pressure is 2200 psia, and the densified nominal heat rate is 5.?8 kW/f t at the rated core power of 2535 MWt. The heat rate is slightly higher than.in Cycle 1 due to the shorter stack height of batch 4. 1487 234 3-2
Figure 3-1 TMI-1 Cycle 2 - Core Loading Diagram FUEL TRANSFER CANAL >-
~
A 4 4 4 4 4 B 4 4 4 2 3 2 4 4 4 F-7 A-8 F-9 4 3 3 2 2 3 2 2 3 3 4 C B-4 A-6 C-6 D-5 B-8 D-ll C-10 A-10 B-12 4 3 2 3 3 4 3 4' 3 3 2 3 4 D D-2 E-8 A-7 B-5 C-4 B-11 A'-9 H-ll D-14 E 4 3 3 3 2 - 2 2- 2 3 3 3 4 F-1 G-1 C-3 E-6 B-7 G-4 B-9 E-10 C-13 G-15 F-15 4 4 2 3 2 2 3 2 3 2 2 3 2 7 4 4 F-3 E-2 F-5 D-7 B-6 C-8 B-10 G-12 F-ll E-14 F-13 4 2 2 4 2 3 5 2 3 3 2 4 2 2 4 C G-6 E-4 G-2 F-2 D-3 G-8 C-12 F-14 G-14 E-12 G-10 4 3 3 3 2 2 2 2 2 2 2 3 3 3 4 H H-1 H-2 N-3 N-7 H-3 H-7 H-8 H-9 H-13 D-9 D-13 H-14 H-15 4 2 2 4 2 3 3 2 3 3 2 4 2 2 4 E K-6 M-4 K-2 L-2 0-4 K-8 N-13 1-14 K-14 - M-12 K-10 4 4 2 3 2 2 3 2 3 2 2 3 2 4 4 L L-3 M-2 L-5 K-4 P-6 0-8 P-10 N-9 L-11 M-14 L-13 4 3 3 3 2 2 2 2 2 3 3 3 4 M 0-3 L-1 K-1 M-6 P-7 K-12 P-9 M-10 0-13 K-15 L-15 4 3 2 3 3 4 3 4 3 3 2 3 4 N N-2 H-5 R-7 ,P-5 0-12 P-ll R-9 M-8 N-14 4 3 3 2 2 .3 2 2 3 3 4 0 0-6 N-5 P-8 N-11 0-10 R-10 P-12 4 4 4 2 3 2 4 4 4 P L-7 R-S L-9 R 4 4 4 4 4 1 2 3 ,
'4 5 6 7 8 9 10 11 12 13 14 15
'N- Batch 1487 235 Y Y- Previous Core Location
. '"T - ~ .
33 ,
Figure 3-2 THI-1 Enrichment and Burnup Distribution for Cycle 2 e 8 9 10 11 12 13 14 15 2.75 2.75 2.75 2.75 3.05 3.05 3.05 2.64 H 11748 16541 16106 16787 11508 13815 10552 0 3.05 3.05 2.75 2.64 2.75 2.75 2.64 X 11508 11825 13318 0 14689 17987 0 2.75 2.75 3.05 2.75 2.64 2.64
- 1. 16787 17056
- 11056 14515 0 0 M 3.05 3.05 3.05 2.64 8207 10132 7778 0 2.75 3.05 2.64 N 18057 7727 0 2.64 0 0 P
R xxx Initial Enrichment xxxx BOC Burnup (MWD /MTU)* 1487 236 3-4 s
I,. Figure 3-3' TMI-1 Cyc.'.e 2 - Control Rod Locations X s. 3 4 5 4 C 2 7 7 2 D . 6 8 3 8 6
~~
2 3 1 l' 3 2
~
4 8 5 6 5 8 4 G 7 1 3 3 1 7 5 3 6 4 6 3 5
-Y l1 l
. K 7 3 3 1 7-g L 4 8 5 6 5 8 4 M 2 3 1 1 3 2 iI
'6 .8 3' 8 6 0
2 7 7 2 P , 4 5 4 R
. l
.Z 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15
, Croup Nu=ber of Rods Functio'n-X Croup Nusber y 8 Safety 2 8 Safety 3 12 Safety 4 9 Safety 5 8 Control 6 8 Control 7 8 Control 8 8 APSR's Total 69 .
.jgg/ }}[
- 4. FUEL SYSTEM DESIGN 4.1. Fuel Assenhly Mechanical Design Pertinent fuel design parameters uce listed in Table 4.1-1. All fuel assemblies are identical in concept and are mechanically interchangeable. The new fuel assemblies incorporate minor modifications to the end fittings, primarily to reduce fuel assembly pressure drop and to increase holddown margin. All other results presented in the FSAR fuel assembly mechanical discussion are applicable to the reload fuel assemblies, i
4.2. Fuel Rod Design Pertinent fuel rod dimensions for resflual and new fuel are listed in Table 4. 2-1. The mechanical evaluation of the fuel rod is discussed below. Cladding Collapse: Creep collapse analyses were performed for three-cycie assembly power i histories for TMI-1. The batch 3 fuel is more limiting than batch 4 fuel due to the lower prepressurization and lower pellet density. A summary of the batch 3 and batch 4 fuel rod designs are contained in Table 4.2-2. The batch 3 assembly power histories were analyzed and the most limiting assen61y determined. Actual operating history was used where it was available. This included data through % 7000 EFPH operation and included the initial power operation at 40% and 80% core power. The predicted assembly power history for the most limiting assembly was used to determine the most limiting collapse time as described in BAW-10084P-A, (Reference 2). The following conditions were analyzed for the worst assembly power history. In all cases, the 2000 hr densification assumption described in Reference 2 was used since it was found to be the most severe case. 'th e coolant and cladding temperatures, and fast flux values were calculated at axial locaticas corresponding to the conditions listed below.
- 1. Assembly outlet conditions
- 2. Axial power peak of 1.0 in the upper part of core
- 3. >bximum axial peak in the u ner part of the core averaged over three cycles of operation.
1487 238 4-1 1 l
i The third condition above was found to be the cost limiting. In addition to the above, the worst batch 2 assecbly was analyzed u. sing . the same procedure described above. The conservatiscs in the analytical procedure are su=sarized below.
- 1. The CROV co=puter code was used to predict the tice to collapse. CROV conservatively predicts collapse times, as demonstrated in Reference 2. *
- 2. No credit is taken for fission gas release. Therefore, the net dif farential pressures used in the analysis are conservatively high.
- 3. The cladding thickness used was the LTL (lower tolerance limit) of the as-built ceasure=ents. The initial ovality
. of the cladding used was the UTL (upper tolerance 1Mt) of
- the as-built ceasurements. These values were taken from a
. statistical sa=pling of the cladding.
The cost limiting asse:bly, Batch 2, was found to 'have a collapse time greater than the naximus projected Cycle 2 life of 19,000' hours. This analysis was perfor=ed using the asst ptions on densification described in Reference 2. Cladding Stress: Since the Batch 3 fuel is the cost limiting from a cladding stress point - of view due to the low prepressurization and Icw denci c , the calculations performed in the TMI-l Fuel Densification Repo.-t, DAW-1389,. June 1973 (Reference 3), are the cost limiting. Fuel Pellet Irradiation $uelling: The fuel design criteria specify a licit of 1.0% on cladding circum-ferential plastic strain. The pellet design is set such that the plastic cladding strain is less than 1% at 55,000 SThEU. The conservatiscs in this analysis are listed below.
- 1. The caximu= specification value for the fuel pellet dia=eter uns used.
2. The raxicun specification value for the fuel pellet density was used. 1487 239 F 4-2
r s 3. The cladding ID used was the lowest permitced specification
, tolerance.
- 4. The maxi =us expected 3 cycle local pellet burnup is less '
, than 53,000 MND/NEU.
4.3. Thermal Design The core loading for Cycle 2 operation is shcun in Figure 3-1, There are 56 fresh (batch 4) fuel asse=blies and 121 once-burned (batches 2 and 3) fuel asse=blies. These asse=blies are thermally and geocetrically similar. E:vever, the batch 4 fuel has a hizhe- initial theoretical density ,(TDI), and a correspondingly higher linear heat rate capability (20.15 kW/ft vs 19.6 kW/ft) than does the batch 2 and 3 fuel. These linear heat rate limitations were . established utilizing the TAFT-3 (Raference 4) rode with full fuel deasifi-
. cation penalties.
Power Spike Model The power spike codel utilized in this analysis is identical to that presented in BAW-10055 (Reference 5) except for' two codifications. The modi-fications have been applied to Fg and Fkas described in Rsference 6. These probabilities have been changed to reflect additional data from operating - - reactors that support a sacewhat different approach and yield Iess severe . penalties due to power spikes. F was changed from 1.0 to 0.5. F was daanged _ from a Gaussian distribution to a linear distribution, which reflects a' decreasing frequency with increasing gap size. The maximum gap size versus azial position is shown in Figure 4.3-1, and the power spike factor versus axial position is shown in Figure 4.3-2. The calculated power spike and gap si;e were based upon 92.5% TDI and an enrichcent of 3 wt % 5 0. The corresponding values for batch 4 fuel would be s= aller because of the. increased density and lower enrich =ent of this fuel. Fuel Temoerature Analysis Thermal analysis of the fuel rods asse=ed in-reactor fuel densification to 96.5% theoretical density, TDF. The basis for the analysis utilized is c'.ven in References 4 and 5 with the following =cdifications:
. 1487 240 . 4-3
i 1. The option in the code for no restructuring of fuel has been used in the analysis presented here in accordance with the 1RC's interim evaluation of TAFY. ~
- 2. The calculated gap conductance vus reduced by 25% in ac ordance with the NRC's interim evaluation of TAFY.
During Cycle 2 operation the highest relative asse=bly povar levels occur is batch 3 fuel (See Figures 3-1 and 51-1). Fuel tecperature analysis for this fuel is doce ented in the TMI-l Fuel Densification Report (Reference 3) for 30C. The =azicus hot spot centerline fuel te=perature is predicted-on the basis of reference design peaking conditions, as shown in Table 4.3-1. Although batch 4 fuel has a, reduced active fuel length, and a correspondingly higher average linear heat race, t.ths caxicus predicted centerline tecperature of this fuel would be lower than that of batch 3 fuel, even with the same peaking factors applied. This is primarily due to the higher initial density of the batch 4 fuel. Therefore, the analysis perfor=ed for Cycle 1 is also applicable for Cycle 2. g 4.4. Materials Design ( '
~
The batch 4 fuel asscchlies are not new in concept and they do not utilize different component caterials. Therefore, the chemical compatibility of all possible fuel-cladding-coolant-assechly interactions for the batch 4 fuel asse=blies are identical to those of the present fuel. 4.5. Operating Experience -- - B&W's operating experiences with co= parable fuel asse=bly design have been de onstrated in the operation of six 177 fuel assecbly plants utilizing this fuel asse=bly design. The most recent application for fuel assecblies of similar design is described in the Oconee 1, Cycle 2 Reload Report (Reference 7) . 1487 2'41 S e 4-4
v Table 4.1-1. Furl Design Parameters Residual New Fuel Assembly Fuel Assembly Batch 2 Batch 3 Batch 4
- 1. Fuel Assembly Type W-B3 &-B3 E-B4
- 2. Number 61 50 56
- 3. Initial Fuel Enrichment 2.75 3.05 2.64
- 4. Initial Fuel Density
% Theoretical 92.5 92.5 93.5
- 5. Initial Fill Gas Pressure, psia proprietary data to be supplied under separate cover.
. 6. Batch Burnup, BOL, MMD/>frU 15,892 10,180 0
- 7. Clad Collapse Time '
Effective Full Power Hours >19,000 >19,000 >19,000 Table 4.2-1. Fuel Rod Dimensions Residual Fuel New Fuel Component Batches 2 & 3 Batch 4
- 1. Fuel Rods 0.D. inches .430 .430 I.D. inches .377 .377
- 2. Fuel Pellet 0.D. inches .370 .370 Density, % Theoretical 92.5 93.5
- 3. Undensified Active Fuel Length, inches 144 142.6
- 4. Flexibic Spacers , Type Corrur,ated Spring Spacer
- 5. Solid Spacers, Material Zr0 Z#~'
2 t
.t*
1487 242 4-5
Table k.2-2. Incut Surnary for Cladding Creen Collacce Calculations Batches 2 & 3 Batch h Pellet OD (=ean specified), in. 3700 3700 Pellet Density (=ean Specified),
% TDI 92 5 93 5 Densified Pellet OD, in. 3650 3663 Cladding ID (mean specified), in. 377 377 Reactor System Pressure, psia 2200 2200 Stack Height (undensified), in. Ikk 1h2.6 1487 243 h-6
Table 4.3-1. Fuel Temperature Analysis Parameters for Cycle 2** Reactor Core Power Level, MWt 2568 System Pressure, psia 2200 Reactor Vessel Average Coolant Temperature, F 579 Fraction of Heat Generated in Fuel and Cladding 0.973 F 1*78 AH F"z 1.70 Fq (Nuclear) 3.03 Fq (Nuclear and Mechanical) 3.12 Average Thermal Output kW/ft - Batches 2, 3 5.77 Batch 4 5.80* Average Fuel Temperature, F 1335 Maximum Fuel Centerline Temperature at Hot Spot, F 4780 Densified Active Fuel Length, In. - Batches 2, 3 141.2 Batch 4 140.4* Linear Heat Rate to Central Fuel Melt, kW/ft - Batches 2, 3 19.6 Batch 4 20.15* Initial Theoretical Density (TDI) - Batches 2, 3 92.5 Batch 4 93.5*
- Batch 4 fuel not used for analysis because batch 2,3 fuel is more limiting.
** Reference 3, BAW-1389.
1487 244 4-7
}
ll i i Figure 4.3-1. Maximum Cap Size Vs Axial Position i i 4.0 1 - TDF = 96.5% 4 J TDI = 92.5%. 1 3.0 - E t
?
0' 5 ! E d 2.0 - 1 & . u a 8 N E 1.0 - O _ i i i I I I 0
.s 0 20 40 60 80 100 120 140 00 N
Axial Position, inches N 4 (.JI
0 4 1 0 I 2 1 n o i t i s 0 o I 0 P 1 l a i x A s s e V h c r I 0 i n o 8 t , c n a o F i t e i k s i o p S P 0 l r I 6 a e i w x o A P 2 3 4 I 0 4 e r u g i F %% 5 5 6 2 9 9 I 0
= - 2 F I D D T T O
0 9 8 7 6 5 4 3 2 1 0 1 0 0 0 0 0 0 0 0 0 0 1 1 1 1 1 1 1 1 1 1 1 u3Nm $#uy2 o Aa NA&
!: ii
- T*
- 5. HUCLEAR DESIGN 5.1 Physics Characteristics Table 5.1-1 compares the core physics parameters of cycles 1 and 2.
The values for both cycles were generated using PDQ07. Since the core has not yet reached an equilibrius cycle, differences in core physics paraseters are to be expected between the cycles. The shorter cycle 2 will produce a s= aller cycle differential burnup than that for cycle 1. The accumulated average core burnup will be higher in cycle 2 than in cycle 1 because of the presence of the once-burned batch 2 and 3 fuel. Figure 5.1-1 illustrates a representative relative power distribution for the beginning of the second cycle at full power with equilibrium xenon and nor=al rod positions. The critical boron concentrations for cycle 2 are lower in all cases than for the cycle 1. The control rod worths for hot full power (due to changes in radial flux distribution and isotopics) are somewhat less than those : for cycle 1 (at BOL), although they are sufficient to maintain the required shutdown cargin, as indicated in Table 5.1-2. Reference fuel cycle shutdown cargin is presented in the TMI-I FSAR Table 3-5. The stuck rod and ejected rod worths are considerably lower for cycle 2 than for cycle 1-except for the stuck rod worth at the end of cycle. Cycle 2 is only slightly higher for this case than cycle 1 and no adverse safety impli-cations are associated with this higher worth. The adequacy of the shut-down margin with cycle 2 stuck rod worths is deconstrated in Table 5.1-2. For the shutdown calculations the following conservatis=s were applied:
- 1) Poison naterial depletion allowance
- 2) 10% uncertainty on nec rod worth
- 3) Flux redistirbution penalty Flux redistribution was accounted for since the shutdown analysis was calculateu using a two-dicensional codel. The shutdown calculation at the end of cycle 2 is analy:cd at approximately 275 EFPD's. This is '.hc latest 1487 247 5-1
time (i 10 days) in core life in which the transient bank is nearly fully inserted. After 275 EFPD's the transient benk will be almost fully with-drawn thus increasing the potential shutdown margin. The cycle 2 power deficits from hot zero power to hot full power are higher than those for cycle 1 due to a more negative moderator co-efficient in cycle 2. The differential boron worths and total xenon worths for cycle 2 are' lower than for cycle 1 due to depletion of the fuel and the associated buildup of fission products. Effective delayed neutron fractions show a similar trend with burnup. 5.2 Analytical Input The cycle 2 incore measurement calculation constants used for com-puting core power distributions were prepared in the same manner as the reference cycle. 5.3 Changes in Nuclear Design There were no relevant changes in core design between the reference and reload cycles. The same calculational methods and design information were used to obtain the important nuclear design parameters. In addition, no significant operational procedure changes exist from dbe reference cycle with regard to axial or radial power shape control, xenon control, or tilt control. The operational limits (Technical Specifications changes) for the reload cycle are shown in Section 8. - 1487 248 0 5-2
Table 5.1-1 TMI-1, Cycle 2 Physics Parameters Cv .le 2** Cycle 1* Cycle length, EFPD 296 466 Cycle burnup, mwd /mtU 9144 14,396 Average core burnup - E0C, mwd /mtU 18,072 14,396 Initial core loading, mtU 82.1 82.1 Critical boron - BOC, ppm HZP - all rods out 1415 1634 HZP - groups 7 and 8 inserted 1252 1494 - HFP - groups 7 and 8 inserted 1066 1382 Critical boron - EOC, ppm HZP - all rods out 440 480 HFP - group 8 (37.5% withdrawn, equil. Xe) 96 180 Control rod worths - HFP, BOC, %Ak/k Group 6 1.20 1.58 Group 7 0.96 0.99 Group 8 (37.5% ud) 0.54 0.44 Control rod worths - HFP, EOC, %Ak/k Group 7 1.33 1.37 Group 8 (37.5% wd) 0.51 0.26 Ejected rod worth - HZP, %Ak/k BOC 0.59+ 0.48++ EOC 0.58+ 0.72++ Stuck rod worth - HZP, %Ak/k BOC 2.16 4.27 E0C 2.22 2.69 Power deficit, HZP to HFP, %Ak/k BOC -1.65 -1.32 EOC -2.49 -2.10 Doppler coeff - BOC, 10' (Ak/k/ F) 100% power (0 Xe) -1.51 -1.51 Doppler coeff - EOC, 10-5 (ak/k/ F) 100% power (equil Xe) -1.55 -1.67
!bderator coe f f - HEP, 10" (ak/k/ F)
BOC (0 Xe, 1000 ppm. groups 7 and 8 inserted) -1.03 -0.23 EOC (equil Xe, 17 ppm, group 8 inserted) -2.60 -2.70 1487 2u49 5-3
Table 5.1-1 (Continued) Cycle 2** Cycle 1* Boron worth - HFP, pps/%Ak/k BOC (1000 ppm) 109 98 EOC (17 ppm) 101 95 Xenon worth - HFP, %Ak/k BOC (4 days) 2.60 2.71 EOC (equilibrium) 2.66 2.65 Effective delayed neutron fraction (HFP) d BOC .00577 . 00690 EOC .00516 . 00514
*For Cycle 1 length of 466 EFPD
** Based on Cycle 1 length of 440 EFPD
+ Ejected rod value for group 5, 6, 7 and 8 inserted 4+ Ejected rod value for Group 6, 7, and 8 inserted 1487 250 4
5-4 -
Table 5.1-2 Shutde.n M1rgin Calculation TMI-1, Cycle 2 I. Available Rod Worth BOC, % Ak/k EOC*, .C Ak/h
- a. Total rod worth, HZP** 9.77 9.80
- b. Worth reduction due to burnup of poison material -0.19 -0.30
- c. Maximum stuck rod, HZP -2.16 -2.22
- d. Net worth 7.42 7.28
- e. Less 10% uncertainty -0.74 -0.73
- f. Total available worth 6.68 6.55 II. Required Rod Worth
- a. Power deficit, HFP to HZP 1.65 2.49
- b. Max. allowable inserted rod worth 1.07 1.27
- c. Flux redistribution 0.40 1.00
- d. Tctal required worth 3.12 4.76 III. Shutdown Margin (I.f. minus II.d.) 3.56 1.79 Note: Required shutdown margin is 1.00% Ak/k
*For shutdcun cargin calculations this is defined as ~ 275 EFPD, the latest time in core life in which the transient bank is nearly full in.
**HZP denotes Hot Zero Power /HFP denotes Hot Full Power.
1487 251 5-5
Figi_ ' .-l BOC (4 EFPD), Cycle 2 Two- mensional s - Relative Power Distribution - Full Power, Equilibrium Xenon, Normal Rod
- Positions (Groups 7 and 8 Inserted) e 8 9 10 11 12 13 14 15 H
1.40 1.27 1.22 1.17 1.21 .85 .77 .60 7 K 1.27 1.41 1.37 1.23 1.23 .57 .69 .59 8' L 1.22 1.37 1.18 1.13 - 1.12 .95 .95 .52 M 1.17 1.23 1.13 1.35 1.29 1.24 .92 . 8 N 1.21 1.23 1.12 1.29 1.10 1.09 .66 7 - 0 .85 .57 .95 1.24 1.69 .72 9 P .77 .69 .95 .92 ,
.66 R .60 .59 .52 '
x Inserted Rod Group No. 1487 252 x.xx Relative Power Density . 5-6
j
- c. ERMAL-EYDRAULIC DESIGN l 6.1 Thermal-Hyd raulle Des ten Calculations -
t Thermal-hydraulic design ' calculations for support of cycle 2 operation utilized predominantly the ss=e analytical methods previously docu=ented in . references 1 and 3. Adjust =ents to these calculations were made to account I for the introduction of the Mk-B4 assemblies in batch 4, to consider the ! minimum actual Reactor Coolant system flow rate as censured during first ; cycle operation, and to incorporate the B&U-2 critical heat flux correlation in place of the previously used W-3 correlation. , l Introduction of Mark B-4 Assemblies - As discussed in section 4.1, th.e Ek-B4 assemblies differ . fren the Mk-B3 assemblies pri=arily in the end fittings. This difference causes a slight reduction in the flow resistance of the B4 assemblies relative to the B3's. Since the B4 asse= biles are loaded pri=arily on the periphery of the core, the hottest (highest power) asse=bly is a B3 asse=bly (sca Figures 3-1 and 5.1-1) . In order to conservatively account for the introduction of the B-4 assechlies, the ther=al-hydraulic model utilizad ~ the cycle 3 configuration (117 34, 60 B3 asse=blies), but retained the 33 assemblics in the hottest core locations. This assumption increases the conservatisc of the cycle 2 design by reducing the calculated hot asse=bly flow rate. . Increased RC System Flow . RC flow data obtained during Cycle 1 operation verified. that the system flow was greater than the design flow. The ninicu= =easured value was 108 percent of design. For the Cycle 2 thertal-hydraulic design analysis the increase in system flev was conservatively chosen to be 106.5 percent of Cycle 1 design flow. Therefore Cycle 2 100% flow is 139.8 x 10 lb/hr. B&W-2 DNS Correlation The B&W-2 DNB correlation is a core realistic prediction of the burnout pheno =cna (as described in references 8 and 9) and has been re-viewed and approved for use with the 15:k 3 fuel assc=bly design. In the application of the b5W-2 DN3 corrcletion to the TMI-1 Cycle 2 core, two =odifications in the use of the correlation have been instituted. 1487 253 6-1
- 1. The limiting design DNBR of 1.30 was used representing a 95 percent confidence level for 95 percent population pro-tection. A limiting DNBR of 1.32 which has been used for this correlation in previous design analyses, represented a 99 percent confidence level for a 95 percent protection.
This change is consistent with industry practice and the statistical standards associated with limiting design DNBR values accepted by the NRC Staff and ACRS.
- 2. The pressure range applicable to the correlation has been extended downward from 2000 psia to 1750 psia. This re-vision is based on a review of rod bundle CHF data taken at pressures below 2000 psia which shows that the B&W-2 correlation conservatively predicts the data in this range.
The use of this correlation in conjunction with increased system flow for the Cycle 2 analysis indicates that the calculated margin to DNB' is higher than had been predicted for first core operation as chown in Table 6.2-1.
~
6.2- DNBR Analysis In addition to the items discussed above, the maximum design conditions considered in the FSAR and generic fuel assembly geometry based on total Mark B as-built data were taken into account. This resulted in a minimus DNBR of 1.919 at 112 percent power for undensified fuel. The effects due to densification can be divided into two categories: (1) the result of reduced stack height and (2) the power spike resulting from densification induced gaps in the fuel column. The active length was calculated to be 141.2 inches, a reduction from the undensified length of 144.0 inches. These de:Mified and undensified lengths are based on fuel Batches 2 and 3 as discussed in Section 4.3.1. The axial flux shape which resulted in the maximum change in DNBR from the original design value was an outlet peak with a core offset of
+11.8%. The spike magnitude and the maximum gap size are disc in Section 4.3 and the values used in the analysis are 1.037 and 3.1 . aches, respectively. The results of the two effects are -1.88 and -1.06 percent change in minimum hot channel DNBR and peaking margin, respectively. The changes in these margins are summarized in Table 6.2-1 which includes com-parisons of other pertinent Cycle 1 and Cycic 2 data. ~
6-2 1487 254' O
-7 Table 6.2-1 Cycle 1 and Cycle 2 Design Condis.ons Cycle 1* Cycle 2 -
Power Level, M'Je 2568 2568 2200 2200 Systes Pressure, psia Reactor Coolant Flow, % Design Flov 100.0 106.5 Vessel Inlet Coolaat Tenperature - 100% Power, OF 554.0 555.6 Vessel Outlet Coolant Temperature-100% Power, OF 603.8 a02.4 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. (undensified) 1hh ihh+
Average Heat Flux (100% Power), 171,470 174,870*** Btu /h-ft2 S
' Maxinus Heat Flux (100% Power),
466,903*** Btu /h-ft2 (for DNBR Calc. ) 457,825 W-3 B&W-2 CHF Correlation
- Pdnicus DNBR (Max. Design Conditions, 1.55 1.919 '
At overpower indicated)no densification (11h5 pover) (112% power) pen alty. - Hot Channel Factors Enthalpy Rise 1.011 1.011
' 1.014 1.014 Heat Flux Flow Area 0.93 0.98 Densification Effects Change in DNBR Margin, % -6.0S** -1.88 Change in Power Peaking Margin, % -2.82** -1.06
*Re fe rence 1. Three Mile Island, Unit 1 FSAR, Docket No. 50;289
~
** Reference 3, "Three Mile Island, Unit 1 Fu21 Densifiction Report,"
*. Eabcock 6 Wilcox Report BAU-1389 (Proprietary), June 1973.
k'* F.or densified fuel
+ Batches 2 & 3 li=iting assenbly. Fcr batch h; 1h2.6 inches.
~
;~3
_ 1487~255
- 7. ACCIDENT AND TRANSIENT ANALYSIS 7.1 Ceneral Safety Analysis Each TSAR (Reference 1) accident analysis has been exa=ined with respect to changes in Cycle 2 para =eters to determine the effects of the Cycle 2 reload and assures that ther=al perfor=ance during hypothetical, transienta is not degraded.
Core ther=al parameters used in the FSAR accident analysis were design operating values based on calculational values plus uncertainties. Co=- parison of first core values (FSAR values) of core ther=al paraceters with para =sters used in Cycle 2 analysis are given in Table 6.2-1. These are parz=eters coc=en to all of the accident analyses presented herein. For each' accident of the FSAR, a discussion of the accident and the key para =eters are provided. A comparison of the key paraceters (See Table 7.1-1) from the FSAR and Cycle 2 is provided with the accident discussions tc show that the ini tial conditions for the transients are bounded by the FSAR analysis { The effects of fuel densification on the FSAR accident results nave been evaluated and are reported in Reference 3. Since Batch 4 reload fuel asse:blies contain fuel rods whose theoretical density is higher than those considered in Reference 3. the conclusions in that reference are still valid. Calculational techniques and cethods rt=ain consistent for Cycle r
. analysis with those used for the ESAR. Additional DKER cargin is shown for -
Cycle 2 by the use of the P&W-2 CHF correlation rather than the W-3 CEF correlation. No new dose calculations war perfor=ed for the reload report. The dose considerations lu the FSAR were basec on maxicus peaking and burnup for all core cycles and therefore the dose considerations are independent.of the reload batch. 7.2 Rod Uithdrawal Accidents This accident is defined as uncontrolled reactivity addition to the core i fro: withdrawal of control rods during startup conditions or from rated power
, conditions. Both types of incidents were analyzed in the FS AR.
The important para =eters during a red withd:cual accident are Doppler c: efficient, modcrator tenperature coef ficient and the rate at which reaccivity is added to the core. Only high pressure and high flux trips are accounted for
~
- 7-'
1487 256
in the FSAR analysis, ignoring multiple alarms, intetlocks and trips that normally precludes this type of incident. For positive reactivity addition indicative of these events, the most severe results occur for BOL conditions. The FSAR values of the key parameters for BOL conditions were -1.17x10-5 (ak/k/F) for the Doppler coefficient, 0.Olb.k/k/F ) for the mocerator temperature coefficient and rod group worths up to and including a 10% ak/k rod group worth. Comparable Cycl
- 2 parametric values are -1.51x10-5 (ak/k/F) for the Doppler coefficient,-1.03xlC (ak/k/F) for the moderator temperature coefficient, and maximum rod group worth of 9.8% ok Therefore, 7
Cycle 2 parameters are bounded by design values assumed for the FSAR analysis. Thus, for the rod withdrawal transients, the consequences will be no more severc than those presented in Reference 1 or 3. _7. 3 Moderator Dilution Accident Boron in the form of boric acid is utilized to control excess reactivity. The boron content of the reactor coolant is periodically reduced to ecmpensate for fuel burnup and transient xenon effects with dilution water supplied by the makeup anii purification system. The moderator dilution transients considered are the pumping of water with zero boron concentration from the makeup tank to , the reactor coolant system under conditions of full power operation, hot shutdown and during refueling. The key parameters in this analysis are the initial boron concentration, boron reactivity worth, and the moderator temperature coefficient for power cases. 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 ppe/1% ak/k baron reactivity worth and 0.5x10 ' ak/k/F for the moderator te=perature coefficient. Comparable Cycle 2 valuco are 1066 ppm for the initial boron concentration, 84 ppm /1% Ak/k boron reactivity worth and -1.03x10" (ak/k)/F for the moderator temperature coefficient. The FSAR shows that the core and RCS are adequately protected during this event. Sufficient time for operator action to terminate this transient is also shown in the FSAR even with maximum dilution and minimum shutdown margin. The predicted Cycle 2 para =cters of importance to moderator dilution transient are bounded by the FSAR design values, thus, the analysis in the FSAR is valid.
,_, 1487 257
7.4 Cold Water (Puno Startup) Accident The NSS does not contain any check o- isolation valves in the reactor coolant system piping, therefore, the cla"sical cold water accident is not po s s ib le. However, when the reactor is operated with one or more pumps not running, and these pumps are then started, the increased flos rate will cause the average core temperature to decrease. If the coderator temperature coefficient is negative, reactivity will be added to the core and a power increase will occur. Protective interlocks exist and administrative procedures are imposed to prevent the starting of idle pu=ps if the reactor power is above 30%. However, these restrictions were not assumed and two pump startup from 50% power was analy:ed as the cost severe transient. To maxi =ite reactivity addition, the FSAR analysis used the most negative
~
=oderator temperature coef ficient of -3.0x10 ' ak/k/F and least negative Doppler
' -5 Ak/k/F. The corresponding most negative moderator temperature coef ficier.t of-1.17x 10 _4 coefficient and least negative Doppler coef ficient predicted by Cycle 2 are -2.60x10
-5 ok/k/F, respectively. As the predicted Cycle 2 moderator ak/k/F and -1 lx10 te=perature coefficient is less negative and the Doppler coefficient is core 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 A reduction in the reactor coolant flow rate can occur from mechanical f ailures or f rom a loss of electrical power to the pumps. With four independent pu=ps available, a techanical failure in one pucp will not affect cperation of the others. With the reactor at power, the effect of loss of coolant flow is a rapid increase in coolant temperature due to reduction of heat receval capability. This temperature increase could result in DNB if corrective action were not taken im=ediately. The key para =eters for 4 pu=p coastdown or locked rotor incident are the flow rate, flow coastdown characteristics, Doppler coefficient, coderator temperature coefficient, and hot channel DNB peaking factors. The conservative initial conditions assumed for the densification report (Reference 3) were: FSAR values
-5 ak/k/F Doppler coefficient, +0.5x10~ ak/k/F of flow and coastdown, -1.1Tx10 oderator terperature coefficient, with densified fuel power spike and peaking.
The results showed the DN3R renained above 1.3 (W-3) for the 4-pump coastdewn and the fuel cladding temperature remained below criteria limits for the locked rotor transient. 7-3 1487 258
I The predicted values for Cycle 2 are -151X10 -5 ak/k/F Doppler coefficient,
~ -1.03x10 ' ak/k/F moderator temperature coefficient and peaking factors as shown in Table 6.2-1. Since the B&W-2 CHF correlation was used for Cycle 2 and the pre-dicted Cycle 2 values are bounded by those used in the Cycle 1 densification repo>t, the results of the Cycle 1 analysis represents the most severe consequences from a loss of flev incident.
7.6 Stuck-Out, Stuck-In, or Drooned Control Rod Accident If a control rod is dropped into the core while operating,a rapid decrease in neutron power would occur, accompanied by a decrease in core average coolant temperature. In addition, the power distributien =ay be distorted due to a new control red pattern. Therefore, under these conditions a return to rated power may lead to localized power drnsities and heat fluxes in excess of design limitations. The key paraceters for this transient are coderator temperature coefficient, worth of dropped rod, and local peaking factors. The FSAR analysis was based on 0.46% ak/k and 0.36% ak/k rod worths with a coderator temperature coefficient
-4 of -3.0x10 ok/k/F. For Cycle 2, the maxi =us worth rod at power is 0.20% ak/k and
~
a moderator temperature cocfficient of -2.60x10 ak/k/F. Since the predicted rod worth is less and the coderator temperature coefficient more positive, the consequences of this transient are less severe than the results presented in the FSAR. 7.7 Loss of Electric Power - Two types of power losses were considered'id the FSAR: 1) a loss of load - condition, caused by separation of the unit from the transmission system, and
- 11) a hypothetical condition which results in a complete loss of all system and unit power except the unit batteries.
The FSAR analysis evaluated the loss of load both with and without turbine runb ack. 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 power except unit batteries, and assuming egeration with f ailed fuel and steam generator tube leakage. These results are independent of core loading, and therefore, the results of the FSAR are applicable for any reload. 7.8 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 blowdown, causing a sudden decrease in primary system temperature, pressure and pressurizer level. The temperature reduction leads to positive reactivity 7-' I 1487 259
insertion (at EOL, the coderator temperatures 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 stop valve as the worst case situation, at end-of-life conditions. The key parameter for the core response is the moderator temperature coefficient which was in the FSAR assumed to be -3.0x10" ak/k/F. The Cycle 2 predicted value cf moderator temperature coefficient is -2.60x10 f.k/k/F. This value is bounded by those used in the FSAR analysis and hence, the results in the FSAR represent the worst situation. 7.9 Steam Generator Tube Failure. A rupture or leak in a steam generator tube allows ;cactor coolant and associated activity to pass to the secondary system. The FSAR analysis is based on complete severence of a steam generator tube. The primary ancern for this incident is the potential radiological release, which is independent of core loading. Hence, the FSAR results are applicable to this reload. 7.10 Fuel Handling Accident The mechanical damage type of accident is considered the maximum potential source.of activity release during fuel handling activity. The primary concern - is radiological releases which are independent of core loading and, therefore, the results of the FSAR are applicable to all reloads. 7.11 Rod Eiection Accident For reactivity to be added to the core at a more rapid rate than by uncontrolled rod withdrawal, physical failure of a pressure barrier component in the control rod drive assembly must occur. Such a failure could cause a pressure differential to act.on a control rod assembly and rapidly eject the assembly from the core. This incident represents the most rapid reactivity insertion that can be reasonably postulated. The values used in the FSAR and
-5 densification report at BOL conditions of -1.17x10 Ak/k/F Doppler coefficient,
+0.5x10 ak/k/F moderator temperature coefficient, and ejected rod worth of 0.65% ak/k represent the maximum possible transient. The corresponding Cycle 2
-5 parametric values of -1.51x10 ak/k/F Doppler, -1.03x10- Ak/k/F moderator temperature coefficient, are both more negative than used in Reference 3 and a maximum predicted ejected rod worth of 0.18% ak/k assure that the results will be less severe than those presented in the FSAR and densification report (References 1 and 3) 7-5 1487 260
7.12 Maximum Hypothetical Accident There is no postulated mechanism whereby this accident can occur since this would require a multitude of failures in the engineered safeguards. The hypothetical accident is based solely on a gross release of radioactivity to the j reactor building ar.d is independent of core loading. Therefore, the results reported in the FSAR are applicable for all reloeds. 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 followirg operation with 1% defective fuel. Rupture of the tank would result in the release of its radioactive contents to the plant ventilation system and to the atmosphere through the unit vent. The consequences of this incident are independent of core loading, and therefore, the results reported the FSAR are applicable to any reload. 7.14 LOCA Analysis A generic 10CA analysis for.B&W 177 FA lowered-loop NSS has been performed using the Final Acceptance Criteria ECCS Evaluation Model. This study is reported in BAW-10103, Rev. 1 (Reference 10). The analysis in Reference 10 is generic in nature since the limiting values of key parameters for all plants in this category were used. Furthermore, the average fuel temperature as a function of the linear heat rate and the lifetime pin pressure data used in the BAW-10103, Rev. 1 LOCA limits analysis are conservative and/or identical compared to.those calculated for this reload. Thus, the analysis and the LOCA limits reported in Reference 10, provide conservative results for the operation of TMI-L, Cycle 2. Table 7.14-1 shows the bounding values for allewable LOCA peak linear heat rates for TMI-l Cycle 2 fuel. I487 26i h 7-6 I, e
Table 7.1-1 COMPARISON OF !Y PARAMETERS FOR ACCIDENT ANA. .IS (440 EFPD) PARAMETER FSAR & DENSIFICATION VALUE PREDICTED CYCLE 2 VALUE Doppler Coefficient, BOL -1.17x10-5 (ak/k)/F -1.51x10-5 (ak/k)/F
, EOL -1.33x10-5 (ak/k)/F -1.55x10-5 @ /k)/F Moderator Coefficient, BOL +0.5x10 (ak/k)/F -1.03x10 (ak/k)/F
,EOL -3.0x10'0 (ak/k)/F
~
-2.60x10 ' (ak/k)/F All Rod Group Worth 10%4k/k 9.8%Ak/k Initial Boron Concentration 1200 ppm 1066 ppm Boron Reactivity Worth, cold 75 ppm /1% ak/k 84 pps/1% ak/k Maximum Ejected Rod Worth 0.65% ak/k O.18% ak/k Dropped Rod Worth, HFP 0.46% ak/k O.20% Ak/k 1487~262 7-7
Table 7.14-1 ALLOWABLE LOCA PEAK LINEAR HEAT RATE Allowable Peak Linear Core Elevation, ft. Heat Rate, kW/ft 2 15.5
- 4 16.6 6 18.0 8 17.0 10 16.0 1487 263 4
e 7-8
I s
- 8. PROPOSED MODIFICATIO:IS TO TECE:IICAL SPECIFICATI0 IS .
The Technical Specifications have been revised for Cycle 2 operation. The changes cade are as a esult of:
. (1) Using the B&W-2 CEF correlation rather than U-3 as discussed in Section 6.1, (2) The use of a 95/95 confidence level rather than 99/95 as discussed in Section 6.1, (3) The use of 106.5% of design flow rather than 100% as discussed in Section 6.1, (4) The use of the Final Acceptance Criteria LOCA analyses for restricting peaks during operation as discussed in Section 7.14.
Based upon the Technical Specifications derived from the analyses presented In this report, the Final Acceptance Criteria ECCS limits will not be exceeded and the ther=al design criteria vill not be violated.
- [-
e 1487 264 e e t 8-1
y 2600 2400 ACCEPTABLE .
._ OPERATION 2200 - - -
a / E C S a. 5 2000 f 3 e
/ UNACCEPTABLE j OPERATION 1800 7
1600 l 560 580 600 640 660 620 Reactor Outlet Temperature. F. A COREPROTECTIONSAFETYLYidlT Figurel 8 -1 8- 2 1487 265
grmalPowerLevel,5 UNACC, .ABLE OPERATION 120 - (-19.112) (112) (+ a.i 2) ACCEPTABLE Kw/ft Limit 4 PUMP Kw/ft Limit 100 -- OPERATION O (.gs,si) M S 83) (86.7) (-us.ss.7) ACCEPTABLE 80 -- 3 & 4 PUMP @(+45.ss.7) OPERAT10N (-us.5: i) 60 __ (59.1) ( q$,s9, ) ACCEPTABLE 2,3 & 4 PUMP @ OPERAT10N 40 - - i
.r . . . .-- - -
t 20 -- . s . i t
! I I I I 1
. -60 -40 -20 0 +20 +40 +60 - l Reactor Power imbalance, 5 l CURVE
- REACTOR COOLANT FLOW (Ib/hr) ,
1 139.8 x 106 2 104.5 x 106 , 3 68.8 x 106 CORE PROTECTION SAFETY LIMITS F i g u r e ' 8-2 8-3
- O -
2600 2400 i2@ c* Yf E. 2200 l 2 / 5 0
; 2000 '/
g -- Y
/\
E 1800 w . 1600 ' i 560 580 600 620 640 660 Reactor Outlet Temperature. F REACTOR COOLANT FLOW . CURVE (LBS/HR) POWER PUMPS OPERATING (TYPE OF LIMIT) i 139.8 x 106 .(100%)* 112% Four Pumps (DNBR Limit) - 2 104.5 x 106 (7h.7%) 86.7% Three Pumps (DNBR Limit) 3 68.8 x 10 6 (h 9.2%) 59.1% One Pump in Each Loop (Quality Limit)
- 106.5% of Cycid.1 Design Flav.
I CORE PROTECTION SAFETY ' BASES F i gu r e 8-3 ' 8-4 i h 1487 267 ~ - -- --
2500 2300 P = 2355 PSIG ACCEPTABLE E OPERATION 2100 S m
- 2 J/
E 0 $/ 1900
=
8 + u % UNACCEPTA8!.E P = 1800 PSIG OPERATION 3 o .
~ ~ 170d -
1500 540 560 580 600 620 640 Reactor Outlet Temperature, i 4 0 PROTECT 10N SYSTEM MAX l MUM ALLOWABLE SET POINTS i gu r e 8-4 8-5 , 1487 268 . i
. Power Level, 5 UNACCEPTABLE 120-OPERATION (108.0) 4% 100 -
4,' u\' .,
+
+N .##/
- I ACCEPTABLE BIT'(80.7) 4 PUMP OPERATION 60- - ACCEPTABLE (53.1) 3 & 4 PUMP OPERATION 40- -
ACCEPTABLE i
~
2,3 & 4 PUMP OPERATION
, 20-- ,
o , W . 6 E ' T T + + . ll Il il 11 m a E i i i I lm i i
-60 -40 -20 0 +20 +40 +60 Power Imbalance, ", A-PROTECTION SYSTEM MAXIMUM ALLOWABLE SET POINTS F i gu r c 8-5 8-6 1487 269
100 iss.s.lo2 iss.v.lo2 POWER LEVEL 19s.4.92 CUT 0FF 90 - i s 8. s,8 5 207,85 80 RESTRICTED REGION RESTRICTED REGION
. 70 S
m 12s.7,s7 60 - o e 50 - 300.44.5 2 40 PERMISSIBLE 222.3,44.5 OPERATING 30 - REGION 20 - 10 - t t l t 8 l l 1 l 1 1 1 1 0 20 40 60 80 100 120 140 160 180 200 220 260 280 300 Rod index, 5 Withdrawn y 0 25 50 75 100 0 2 5 ~^ 50 75 100 i i f f f I f I t t Group 5 Group 7 0 25 50 75 100 f I f Group 6 ; e ROD POSITICN LIMITS FOR 4 PUMP OPERATICN APPLICABLE CURING THE PERIOD FROM O TO 152 1 10 EFPD; CYCLE 2 .I Figure 8-6 g 1487 270 . J
; . ?-
OPERAil0N IN THIS REGION 875.7.102 202.1.102
~
IS NOT ALLOWED POWER LEVEL 90 - 202.i.92 CUTOFF
. RESTRICTED SH - 70 -
DOWN MARGIN s2s.s.73 U
= LIMIT .n y 60 - + 36.50 50 -
g RESTRICTED PERMISSIBLE ' "*"' REGION
= 40 - OPERATING 222.3.us REGION 30 -
20 - - o)l5 10 - ,
- 0 ' ' ' ' ' ' ' ' ' ' ' ' ' i 0 20 40 60 80 100 120 140 160 180 200 220 240 260 280 300
- .' Rod inder. 5 Withdrawn O 25 50 75 100 f
0 25 50 75 100 i i f i f f f f f Group 5 Croup 7 0 25 50 75 100 i i . , Group 6 R00 POSITION LIMITS FOR 4 PUMP OPERATICN APPLICABLE CURING THE PERICO FROM 152 1 10 TO 275 1 10 EFP0; CYCLE 2 Figure 8-7 8-8 - 1487 271
II
-s e ; ? f '. .
4 OPERATION IN THIS REGION 132.102
~ 253.s.ior IS NOT ALLOWED POWER LEVEL 90 -
CUT 0FF ''* 80 - M- SHUTOOWN MARGIN 3 LIMIT REGION o 60 - at J 50 - s,,,,,7 2 222.3.47 PERMISSIBLE . 30 - OPERATING REGION 20 - 15.g.15 .' I l I f f I f I I I , , ; , 0 20 40 60 80 100 120 140 160 180 200 220 240 260 280 300 Rod index. f. Witndrawn O. 25 50 75 100 0 25 50 75 I I I I I 100 I I g g , Group 5 Group 7 0 25 50, 7,5 100 . Group 6 RCD POSITION LIMITS FOR 4 PUMP OPERATION APPLICABLE CURING THE PERICO AFTER 275 1 10 EFPD; CYCLE 2 8-9 " 1487 272
sos, tot 15s,102 100 - 2cs.so2 2 a,102 1 RESTRICTED REGION RESTRICTED RESTRICTED REGION 90 - FOR 2 ANO 3 PUMP REGION FM2W3 OPERATION R3 p p p 32s.s ss
- c 80 .
OPERATIOP' 200. a 2. s
; , 222.2.s2.5 a
70 j RESTRICiED REGION a FOR 3 PUMP 60 -
$ OPERATION PEREISSI6LE 50 OPERATING 222.3.5s.5 300,5s.5
- REGION 3 40 -
o E 3 30 -
== $ 20 -
g 10 .. # - m 0.o e I t t t , I , t 0 t t t t 20 40 60 80 t t 100 120 140 160 180 200 220 240 260 280 300 O 25 50 Rod Index. i Withdr. inn 75 100 t i 0 25 1 I 50 75 100 i t Group 5 I f ' Group 7 0 25
' 50 75
' ' 100 i Grcup 6 s, R00 POSITICN LIMITS FOR 2 AND 3 PLNP OPERATICN APPLICABLE CURING THE PERIOC FRCM 0 TO 152 1 10 EFP0; CYCLE 2 8-10 Figure 8-9 1487 273
OPERATION IN THIS REG.10N 178,102 100 - 207.102 218.802 IS NOT ALLONED I RESTRICTED REGION FOR 2 AND 3 PUMP
, '90 -
= OPERATION
- 3,,,,g
. 1
% 80 222,as E ~
o SHUT 00NN MARGIN RESTRICTED REGION 70 -
,7 FOR 3 PUMP !! PERMISSIBLE OPERATION $ 60 -
OPERATING - REGION 222,s9 300,s9 d' 3 50 - 2 a 3 40 -
% 30 -
49 , d
=
20 - E" O. e s , 10 I 1 1 I 1 1 f I f f I f I t 0 20 40 60 80 100 120 140 160 180 200 220 240 260 280 300 Rod Index. 5 Witndrawn 0 25 50 75 100 0 25 50 75 100 t 1 1 i f 9 f f f I
. Group 5 3roup 7 0 25 50 75 100 .
I t I t i Group 6 ROD POSITION LIMITS FOR 2 AND 3 PUMP OPERATION APPLICABLE CURING THE PERIOD FRCM 152 1 10 TO 275 i 10 EFPD; CYCLE 2 Figuro 8-10 - 8-11 . 1487 274 .
s
~
OPERATION IN THIS REGION 882.102
~ 224,802 233, 02 IS NOT ALLOWED I 4---
- -RESTRICTE0 E 90 -
159.sr REGION FOR 3 22:.3.er i. jn 80 2 AND 3 PUMP e RESTRICTED REGION E 70 - FOR 3 PUMP d' o SHUTOOWN MARGIN OPERAil0N [ 60 y LIMIT ss.so 222.3.60 3 50 - g" ss.s,47 PERMISSIBLE 5 40 - OPERATING E a 30 - REGION I ' . j 20 - ase s.es , 10 I f f ' I I I I I I f I 1 1 0 20 40 60 80 100 120 140 160 180 200 220 240 - 250 280 300 Rod inder, 5 Withdrawn 0 25 50 f 75 100 t f 0 25 f f t 50 75 100 t g g g Group 5 Group 7 0 25 50 , i 75 100
- i , ,
s, Group 6 . 1 ROD POSITION LIMITS FOR 2 AND 3 PLHP - (PERATION APPLICABLE CURING THE PERICO AFTER 275 i 10 EFP0; CYCLE 2 ' Figure 8-11 ! 8-12 1487 275
3 1 Power, ",of 2535 MWt i o t
'l RESTRICTED REGION
-l1.22,102 11.22,102
-100 '
'l
-11.04,92 I 10.12,92
-14.45,85 13.6,85 .
. - 80
-18.75,75 18.75,75
. 70
- 60
- 50 .
PERMISSIBLE OPERATING
-20.03,44,5 REGION. 40 -
. . 30
- 20
- - 10
. I I I I I I I I 1 50 40 -30 20 -10 0 10 20 30 40 50 Axial Power imbalance, ", i OPERATIONAL POWER IMBALANCE ENVELOPE APPLICABLE TO OPERATION FROM 0 TO 152 i 10 EFPD- CYCLE 2 F i gu r e 8-12 8-13 , 1487 276
4 1
. a Power, 5 of 2535 MWt l
I RESTRICTED REGION ! 16.32,102 ,,
,11.22,102 $
J
-15.64,92 ,, gg 10.12,92 -
-15.3,85 .
-- 80 70
'7*75'7' 60
-- 50
-I '7
PERMISSIBLE OPERATING REGION
-. 40 30
-- 20
-- 10 I I I I I f f I I l
-50 -40 -30 10 0 10 20 30 40 50 Axial Power Imbalance, 5 *
.c s, .
OPERATIONAL POWER IMBALANCE ENVELOPE : APPLICABLE TO OPERATION FROM 152 + 10 TO 275 1 10 EFP0; CYCLE 2 , i' F i gu r e 8-13 8-14 : 1 1487 277 :
Power, ",of 2535 MWt . RESTRICTED REGION
-- 100 -
22.54,92 13.8.92
-- 80
-- 70 15.75,63 60 50
-22.56,47 PERMISSIBLE . ,
OPERATING c- 40 REGION - 30 *
~
-- 20 10 i i i i , , , , ,
-50 -40 -30 20 -
10 0 10 20 30 40 50 Axial Power trabalance, ", z s, OPERATIONAL POWER IMBALANCE ENVELOPE APPLICABLE TO OPERATION AFTER 275 1 10 EFP0; CYCLE 2 F i g u r e 8-14 ! 8-15 1487 278 e N ==w- .ww-..
- 9. STARTUP i'it0GRArt ihc planned startup testing asauciated with core performance are provided below. These tests verify that core performance is within the assumptions of the safety analysis and provide the necessary data for con-tinued safe plant operation.
i Zero Power Tests
- 1. Critical Boron Concentration
, 2. Temperature Reactivity Coefficient
- 3. Control Rod Group Worth 4 Ejected Rod Worth .
Power Tests
- 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 Verifi-cation at Approximately 75% FP
- 4. Power Doppler Reactivity Coefficient at Approximately 10C% FP
- 5. Temperature Reactivity Coefficient at Approximately 100% FP 1487 279
10 REFERENCES
- 1. Three Mile Island-Unit 1 Nuclear Station, Final Safety Analysis Report, Docket No. 50-289.
- 2. A.F.J. Eckert, H. W. Wilson, and K. E. Yoon, "P.ogram to Determine In-Reactor Performance of B&W Fuels - Cladding Creep Collapse," B AW-10084P-A, Babcock & Wilcox, January 1975.
- 3. Three Mile Island Unit 1 Fuel Densification deport, BAW-1389, Babcock &
Wilcox, June 19 73.
- 4. C. D. Morgan and H. S. Kao, "TAFY - Fuel Pin Temperature and Gas Pressure Analysis," BAW-10044, Babcock & Wilcox, May 1972.
- 5. Fuel Densification Report, BAW-10055. Rev.1, Babcock & Wilcox, June 1973.
- 6. "Densification Kinetics and Power Spike Model," Meeting with USAIC, July 3, 1974; J. F. Harrison (B&W) to R. Lobel (USAEC) Telecon, " Power Spike Factor,"
July 18,1974. *
- 7. Oconee 1 Cycle 2 - Reload Report, BAW-1409 (Rev.1), Babcock & Wilcox, October 10, 1974.
- 8. ' Correlation of Critical Heat Flux in a Bundle Cooled by , Pressurized Water,"
BAV-10000, Babcock & Wilcox, March 1970.
- 9. " Correlation of Critical Heat Flux in a Bundle Cooled by Pressurized Water,"
BAW-10036, Babcock & Wilcox, February 1972.
- 10. "ECCS Analysis of B&W's 177 FA Lowered-Loop NSS," BAW-10103, Rev.1, September 1975. -
1487 2130 p 6 10-1}}