ML19322E510
| ML19322E510 | |
| Person / Time | |
|---|---|
| Site: | Crystal River  |
| Issue date: | 02/29/1980 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| Shared Package | |
| ML19322E508 | List: |
| References | |
| BAW-1607, NUDOCS 8003280276 | |
| Download: ML19322E510 (80) | |
Text
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BAW-1607 February 1989 I
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! l CRYSTAL RIVER UNIT 3
- Cycle 3 Reload Report -
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Babcock & Wilcox
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20IU;$ <32$$/3'
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L- BAW-1607 February 1980 I
L CRYSTAL RIVER UNIT 3 L
> - Cycle 3 Reload Report -
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- BABC0CK & WILCOX Power Gen'; ration Group E Nuclear Power Generation Division L P. O. Box 1260 Lynchburg, Virginia 24505 I
Babcock & Wilcox
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CONTENTS l
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- 1. INTRODUCTION AND
SUMMARY
. . . . . . . . . . . . . . . . ..... 1-1
- 2. OPERATING llISTORY . . . . . . . . . . . . . . . . . . . ..... 2-1
- 3. GENERAL DESCRIPTION . . . . . . . . . . . . . . . . . . ..... 3-1 3.1. Plant Description . . . . . . . . . . . . . . . . ..... 3-1 3.1.1. Reactor Coolant System Stress . . . . . . ..... 3-1 g 3.1.2. Reactor Coolant Pump Power Monitors . . . ..... 3-1 g 3.2. Core Description . . . . . . . . . . . . . . . . . ..... 3-3
- 4. FUEL SYSTEM DESIGN . . . . . . . . . . . . . . . . . . . ..... 4-1 4.1. Fuel Assembly Mechanical Design . . . . . . . . . ..... 4-1 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. Fuel Thermal Design . . . . . . . . . . . . . . . ..... 4-2 g i
4.4. Operating Experience . . . . . . . . . . . . . . . ..... 4-2 E
- 5. NUCLEAR DESIGN . . . . . . . . . . . . . . . . . . . . . ..... 5-1 5.1. Physics Characteristics . . . . . . . . . . . . . ..... 5-1 5.2. Changes in Nuclear Design . . . . . . . . . . . . ..... 5-2
- 6. TiiERMAL-IIYDRAULIC DESIGN . . . . . . . . . . . . . . . . ..... 6-1 6.1. DNBR Evaluations . . . . . . . . . . . . . . . . ..... 6-1 6.2. Pressure-Temperature Limit Analysis 6-2 Flux / Flow Trip Setpoint Analysis E
6.3. . . . . . . . . . ..... 6-2 3 6.4. Loss-of-Coolant Flow Transients . . . . . . . . . ..... 6-3 !
- 7. ACCIDENT AND TRANSIENT ANAL' ISIS . . . . . . . . . . . . ..... 7-1 7.1. General Safety Analysis . . . . . . . . . . . . . ..... 7-1 l 7.2.
7.3.
Accident Evaluation . . . . . . . . . . . . . . . .....
Rod Withdrawal Accidents . . . . . . . . . . . . . .....
7-2 7-2 El 7.4. Moderctor Dilution Accident . . . . . . . . . . . ..... 7-3 g ,l l 7.5. Cold Water Accident . . . . . . . . . . . . . . . ..... 7-4
! 7.6. Loss of Coolant Flow . . . . . . . . . . . . . . . ..... 7-4 7.6.1. Four-Pump Coastdown . . . . . . . . . . . ..... 7-5 7.6.2. Locked Rotor . . . . . . . . . . . . . . . ..... 7 -5 7.7. Stuck-Out, Stuck-In, or Dropped Control Rod Accident .... 7-6
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I CONTENTS (Cont'd)
Page 7.8. Loss of Electric Power . . . . . . . . . . . . . . ... .. 7-6 7.9. Steam Line Failure . . . . . . . . . . . . . . . . .. . . . 7-7 I 7.10.
7.11.
7.12.
Steam Generator Tube Failure Fuel Handling Accident . . . .
Rod Ejection Accident . . . .
7-7 7-8 7-8 7.13.
I 7.14.
7.15.
Maximum Hypothetical Accident Waste Gas Tank Rupture LOCA Analysis 7-9 7-9 7-9 7.16. Failure of Small Lines Carrying Primary Coolant Outside I Containment 7.16.1.
Identification of Causes 7.16.2. Analysis of Effects and Consequences 7-9 7-9 7-9
- 7.17. Main Feedwater Line Break . . . . . . . . . . . . .. . .. 7-11 7.18. Dose Consequences of Accidents . . . . . . . . . . ... .. 7-12
- 8. PROPOSED MODIFICATIONS TO TECilNICAL SPECIFICATIONS . . . ... . 8-1
- 9. STARTUP PROGRAM - PilYSICS TESTING . . . . . . . . . . . . .. . .. 9-1 i
- 9.1. Precritical Tests . . . . . . . . . . . . . . . . .. .. . 9-1 9.1.1. Control Rod Trip Test . . . . . . . . . .. . . . 9-1 9.1.2. RC Flow . . . . . . . . . . . . . . . . . .. . . . 9-1 9.1.3. RC Flow Coastdown . . . . . . . . . . . . .. ... 9-1 9.2. Zero Power Physics Tests . . . . . . . . . . . . . .. . .. 9-2 9.2.1. Critical Boron Concentration . . . . . . ... . . 9-2 9.2.2. Temperature Reactivity Coef ficient . . . ... .. 9-2 9.2.3. Cont rol Rod Group Reactivity Worth . . . . . ... 9-2 9.2.4. Ejected Control Rod Reactivity 'lorth . . . .. .. 9-3 9.3. Power Escalation Tests . . . . . . . . . . . . . . .. . .. 9-3 9.3.1. Core Power Distribution Verification at %40, 75, I and 1007. FP With Nominal Control Rod Position ... 9-3 9.3.2. Incore Vs Excore Detector Imbalance Correlation Verification at %40% FP . . . . . . . . . . .. . . 9-5 I 9.4.
9.3.3.
9.3.4.
Temperature Reactivity Coefficient at %100% FP ..
Power Doppler Reactivity Coefficient at %100% FP .
9-5 9-5 Procecure for Failure to Meet Acceptance Criteria . .. .. 9-6 REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . .. .. A-1 I
List of Tables Table 4-1. Fuel Design Parameters and Dimensions . . . . . . . . . .... 4-3 4-2. Fuel Thermal Analysis Parameters . . . . . . . . . . . . .. . . 4-4 5-1. Physi s Parameters, Crystal River Three, Cycle 3 . . . . . . . . 5-3 I
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1 Tables (Cont'dl Table Page 5-2. Shutdown Margin Calculation for Crystal River 3, Cycle 3. . . . 5-5 6-1. Cycle 1, 2, and 3 Thermal-llydraulic Design Conditions . . . . . 6-4 <
7-1. Comparison of Key Parameters for Accident Analysis . . . . . . . 7-13 7-2. Bounding Values for Allowable LOCA Peak Linear lleat Rates . . . 7-13 7-3. Input Parameters to 1.oss-of-Coolant-Flow Transients . . . . . . 7-14 7-4. Summary of Minimum DNBR Resulto for Limiting Loss-of-Coolant-Flow Transients . . . . . . . . . . . . . . . ..... 7-14 7-5. Analysis Assumptions for MU6PS Letdown Line Rupture Accident . . 7-15 7-6. Activity Releates to Environment Due to Rupture of MU6PS Letdown Line . . . . . . . . . . . . . . . . . . . . . ... . . 7-16 7-7. Radiological Consequences of MU6PS I.etdown Line Rupture Outs ide Containment . . . . . . . . . . . . . . . . . ..... 7-16 8-1. Technical Specification Changes . . . . . . . . . . . ..... 8-2 8-2. RPS Trip Setroints . . . . . . . . . . . . . . . . . . ..... 8-12 8-3. Quadrant Power Tilt Limits . . . . . . . . . . . . . . ..... 8-13 8-4. DNBR Limits . . . . . . . . . . . . . . . . . . . . . ..... 8-13 List of Figures Figure l 3-1. Core Loading Diagram for Cyrstal River 3, Cycle 3 . . ..... 3-4 3-2. Enrichment and Burnup Distribut ion for Crystal River 3, Cycle 3 . . . . . . . . . . . . . . . . . . . . . . . ..... 3-5 3-3. Control Rod Locations . . . . . . . . . . . . . . . . ..... 3-6 5-1. !!OC , Cycle 3 Two-Dimensional Relative Power Distribution -
!!FP, Equilibrium Xenon, Ba nks 7 and 8 Inserted . . . . ..... 5-6 7-1. Four-Pump Coastdown - llot Channel MDNBR Vs Time, Crystal g River 3 . . . . . . . . . . . . . . . . . . . . . . . ..... 7-17 7-2.
5 Locked-Rotor, Crystal River 3 . . . . . . . . . . . . ..... 7-18 8-1. Reactor Core Safety Limits . . . . . . . . . . . . . ..... 8-14 j
8-2. Reactor Core Safety Limits . . . . . . . . . . . . . . ..... 8-15 I 8-3. Reactor Trip Setpoints . . . . . . . . . . . . . . . . ..... 8-16 8-4. Pressure / Temperature Limits . . . . . . . . . . . . . ..... 8-17 8-5. Regulating Rod Group Insertlon Limits for Four-Pump Operation From 0 to 250 ? 10 EFPD . . . . . . . . . . . . . . . 8-18 8-6. Regulating Rod Group Insertion Limits for Four-Pump Operation After 250 t 10 EFPD l
8-7.
Regulating Rod Group Insertion Limits for Three-Pump 8-19 g g
Operation From 0 to 250 10 EFPD . . . . . . . . . . . . . . . 8-20 8-8. Regulating Rod Group insertion Limits for Three-Pump Operat ion Af ter 250 f 10 EFPD . . . . . . . . . . . . . . . . . 8-21 8-9. APSR Position Limits for 0 to 2501 10 EFPD, Crystal River 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8-22 8-10. APSR Position Limits Af ter .{501 10 EPPD, Crystal River 3 8-23 8-11. Axial Power Imbalance Envelopa for Operation From 0 to 250 l
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! 10 EFPD . . . . . . . . . . . . . . . . . . . . . . . . . . . 8-24 8-12. Axial Power Imbalance Envelope for Operation Af ter 250 i 10 EFPD g
. . . . . . . . . . . . . . . . . . . . . . . . . . . 8-25 g
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- 1. INTRODUCTION AND
SUMMARY
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This report justifies the operation of Crystal River Unit 3 (cycle 3) at a
[ rated core power of 2544 MWt. Included are the required analyses to support L.
cycle 3 operation; these analyses employ analytical techniques and design r
I bases established in reports that have received technical approval by tin.
USNRC (see references).
I f The design for cycle 3 raises the rated thermal power from 2452 to 2544 MWt; the latter corresponds to the ultimate core power 1cvel identified in the r Crystal River l' nit 3 FSAR.1 l
The upgraded power was analyzed f or cyc le 22 ,
L but the upgrade was not implemented and cycle 2 was operated at 2452 MWt; many of the vnalyses are again summarized in this report for completeness.
L Each accidert analyzed in the FSAR has been reviewed, and each review is sum-marized in this report. Some accidents were re-analyzed to include the re-3 L actor coolant pump power monitors , which are being installed during the refueling outage. It is worthy of note that several other Babcock & Wilcox cores of the same design are licensed for 2568 MWt. The Technical Spec if i-cations have been reviewed, and the modifications required f or cycle 3 are justified in this report, llased on the analyses performed, which take into account the postulated ef-fccts of fuel densification and the Final Acceptance Criteria for emergency 1 I
core cooling sistems (ECCS), it has been concluded that Crystal River 3, cycle 3, u,n be safely operated at the rated core power level of 2544 MWt.
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- 2. OPERATING !!ISTORY i l l
Cycle 1 of the Crystal River Unit 3 nuclear generat ing plant was completed on April 23, 1979, af ter 440 EFPD at 2452 MWt. Cycle 2, which achieved criti-cality on July 29, 1979, is scheduled for completion in April 1980 after ap-proximately 190 EFPD at the current rated power level of 2452 MWL. No oper-f ating anomalies have occurred during previous cycle operations that would W adversely affect fuel performance in Cycle 3.
Cycle 3 is scheduled to start operation in June 1980 with an upgraded rated W power level of 2544 MWL. The design cycle length is 320 EFPD.
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- 3. CENERAL DESCRIPTION 1
3.1. Plant Description L 3.1.1. Reactor Coolant System Stress In support of the power upgrade, reactor cool.nt system (RCS) stresses were L- reviewed. Since the Crystal River 3 (CR-3) f unctional specification did not
_ analyze power levels up to 2544 MWt, a new document was issued. The revised
_ document was reviewed by the applicable engineering groups, and it was deter-mined that no hardware changes were required; however, a revision was issued to the RCS Stress Report.
3.1.2. Reactor Coolant Pump Power Monicors In support of the power upgrade, reactor coolant pump power monitors (RCPPMs)
_ are being added to CR-3 during the EOC-2 refueling outage. 3 The RCPPM anticipates a loss or reduction of the reactor coolant flow by moni-torir.g RC pump power and detecting abnormal power conditions indicative of an L inopercht- pump. The status of each pump is transmitted by the RCPPM to each of four reactor protection system (RPS) channels. Two RCPPMs are supplied to provide redundant pump status information to each RPS channel. Logic in the RPS will act on the pump status information and take appropriate action as follows:
- 1. With three or four RC pumps operating, no action is taken by the RCPPM.
Reactor protection is provided by the nuclear overpower based on the RCS flow and axial power imbalance unit of the RPS.
- 2. With two or fewer RC pumps operating, the RCPPM trips the reactor.
As stated in the accident analyses of the CR-3 FSAR, in the event of a loss r~
L of reactor coolant flow due to failure of one or more of the RC pumps at the present licensed power level of 2452 MWt, the transient is terminated by the present RPS flux-flow trip. The present RPS action is quick enough to E
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I prevent the minimum DNBR going below 1.30 for the four-pump coastdown transient and below 1.00 for the locked-rotor transient.
However, at thermal power levels above 2500 MWt, RPS action by the flux-flow comparator is not fast enough - in the event of loss of more than one RC pump - to keep 'he minimun DNBR from going below the acceptance criterion.
Therefore, for power levels above 2500 MWt, nuclear overpower based on RCPPMs cust be added to the RPS trip functions to reduce the response time of the RPS and thereby terminate the transient quickly enough to ensure compliance with the minimum DNBR limits.
Each RCPPM string includes two current transformers and two potential trans-formers to measure the current and voltage on the RCP power feed lines. The transformers provide input to an electronic watt transducer, which produces an output signal proportional to real power. This power signal is fed into a g
bistable, which provides a contact output for selected overpower and undm - 5 power setpoints. The bistable output contact actuates four separate relays.
A contact from each relay is wired to its respective RPS channel. Thus, one pump monitor string provides status information for one pump to each of four RPS channels. An identical redundant string using separate transformers and monitcring equipment again provides status information for the same pump to the four RPS channels. In the event of failure of one string, all four RPS channels would still have the necessary pump status information via the re-dundant string.
The complete RCPPM system is constructed so that equipment belonging to re-dundant strings is placed inside enclosures separated by barriers. Contact outputs from the RCPPM cabinets to the four RPS channels are arranged to pro-vide adequate physical separation and electrical isolation of each channel.
External signal cable and equipment separation for this installation complies with IEEE 384-1977 and Regulatory Guide 1.75. Where separation cannot be maintained, physical barriers are included.
RCPPM cabinets and equipment specified are seismically qualified and located in a Class I structure. All supports for engineered safeguards cable trays 3 and conduits are designed for OBE and SSE using the acceleration floor re-sponse spectra develcoed for applicable levels of the containment building, auxiliary building, intermediate building, and control complex.
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separation of the cables carrying redundant t ransformer outputs to the RCPPM cabinets is provided in accordance with the separation criteria stated above.
j The current and potential transformers are not seismically qualified because they are not required to safely shutdown the reactor. The loss of the current or potential transformers would result in a " pump inoperable" signal to the
[ RPS.
Upon receipt of t wo such signals, wha tever t he cause, the RPS trips the reactor.
3.2. Core Description The CR-3 reactor core is described in detail in Chapter 3 of the Final Safety Analysis Report for the unit.I The cycle 3 core consists of 177 fuel assem-blles (FAs), each of which is a 15-by-15 array containing 208 fuel rods; 16
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I cont ro l rod guide tubes; and one incore instrument guide tube. The fuel as-semblies in batches 2, 3, and 5 have an average nominal fuel loading of 463.6 I
kg of uranium, whereas the batch 4 assemblles maintain an average nominal fuel loaning of 468.6 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 (see Table 4-2 for data).
Iigure 3-1 is the core loading diagram for cycle 3 of Crystal River 3. The inittal enrichments of batches 2, 3, and 4 were 2.54, 2.83, and 2.64 wt %
oranlum-235, respectIvely. Fifty-two of the batch 2 assemblles wilI be dis-charged at the end of cycle 2. The batch 5 design enrichment is 2.62 wt 7, uranium-235. Batches 3 and 4 and the remaining batch 2 assemblies will he shuf fled to new locations. The batch 5 assemblies will occupy the periphery l of the core. Figure 3-2 is an eighth-core map showing the burnup of each as-sembly at the beginning of cycle 3 and its initial enrichment.
Core reac t is ity will be cont rolled by 61 full-length Ag-In-Cd cont rol rod as-semblics (CRAs) and soluble boron st'm. In addition to the full-length CRAs, eight axial power shaping rods (APSRs) are provided for addit ional control of the axial power distribution. The cycle 3 locations of the 69 control rods and the group designations are unchanged f rom cycle 2 and are shown in Figure 3-3. Control rod group 7 wi.11 he withdrawn at 250 ' 10 EFPD of operatlon.
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Figure 3-1. Core Loading Diagram for Crystal River 3 Cycle 3 I
A 5 5 5 5 5 F7 C9 F9 .u Cycle 2 Location B 5 5 5 5 5 5 3 3 3 Y Batch Number u C 013 D7 N3 L1 P8 L15 N13 D9 03 4 4 5 g 3 4 2 4 4 3 C7 N2 M2 D5 R8 Dil M14 N14 G13 D 5 5 3 4 4 3 4 3 4 4 3 G4 B12 F6 K5 K1 L8 l K15 K11 F10 B4 G12 5 3 4 2 3 4 3 l_
4 3 2 4 3 W C12 Bil E9 E5 D6 BIO D10 Ell E7 B5 C4 4 4 3 3 3 4 3 3 3 4 4 C C6 A10 E4 A9 F4 B6 D8 F14 F12 A7 E12 A6 G10 5 5 3 4 3 4 3 4 3 4 3 4 3 4 3 G3 1l14 111 5 H10 F2 H4 11 8 H12 L14 11 6 III !!2 K13 H 5 3 2 4 3 4 3 2 3 4 3 4 2 3 K6 RIO M4 R9 L4 L2 N8 P10 L12 R7 M12 R6 K10 K 5 3 4 3 4 3 4 3 4 4 5
3 3 4 3 012 P11 M9 M5 N6 P6 N10 Mll M7 PS 04
' 5 5 5 4 4 3 3 3 4 3 3 3 4 4 K4 P12 L6 G5 G1 F8 G15 L10 M 5 3 4 2 3 4 3 4 Gil 3
P4 4
K12
$ l 2 3 5
K3 D2 E2 N5 A8 N11 E14 D14 09 N 5 5 5 5 3 4 4 3 4 3 4 4 3 0
I N7 D3 F1 B8 F15 D13 N9 C3 5 $
4 3 4 4 2 4 4 3 4 L7 07 L9 P 5 5 5 5 5 5 3 3 3 un l R 5 5 5 5 5 l
l 1 2 3 4 5 6 l7 8 9 10 11 12 13 14 15 I
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Figure 3-2. Enrichment and Burnup Distribution for
, Crystal River 3, Cycle 3
-- 8 9 10 11 12 13 14 15 3
2.54 2.83 2.64 2.83 2.64 2.54 2.83 2.62
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17,741 17,349 6,503 20,533 3,574 17,114 16,719 0 2.64 2.83 2.64 2.83 2.64 2.83 2.62 m K 6,502 13,721 ' 987
, 18,191 3,693 16,069 0 m
2.83 2.83 2.64 2.64 2.62 2.62 14,929 14,444 5,624 6,883 0 0 m
E 2.54 2.64 2.83 2.62 17,914 4,163 18,295 0 k
2.83 2.62 2.62 L 16,720 0 0
{ 2.64 4,827 l
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x.xx Initial Enrichment l xx xxx BOC Burnup, mwd /mtU l
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I Figure 3-3. Control Rod Locations I
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GROUP NUMBER OF RODS FUNCT10N I 8 SAFETY 2
E 8 SAFETY 3 3 12 SAFETY 4 9 SAFETY E
5 8 CONTROL 5
. 6 8 CONTROL 7 8 CONTROL 8 8 APSRs TOTAL 69 I
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- 4. Ft1EL SYSTEM DESIGN l
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{ 4.1. Fuel Assembly Mechanical Design The types of fuel assemblies and pertinent fuel design parameters for Crystal River l! nit 3, cycle 3 are listed in Table 4-1. Two assemblies will conta in f primary neut ron sources (PNS), and two assemblies will contain regeneralIve l neutron sources (RNS) in cycle 3. The justification for the design and use of the retainers described in reference 4 is applicable to RNS and PNS :e-I tainers in the CR-3, cycle 3 fuel.
cept and are mechanically Interchangeable.
All fuel assemblies are identical in con-All other results presented in the Crystal River 3 Cycle 2 Reload Report are applicable to the reload fuel 2
assemblles.
I 4.2. Fuel Rod Design The f uel pellet end configuration has changed f rom a spherical dish for batch-es 1 through 4 to a truncated cone dish for batch 5; this minor change facil1-tates manufacturing.
The mechanical evaluation of the fuel rod is d iscussed below.
I 4.2.1. Cladding Collapse Creep collapse analyses were performed for three-cycle assembly power histories for Crystal River 3.
Batches 2 and 3 are more limiting than batches 4 and 5 due to their previous incore exposure time. A batch 3 fuel assembly was de-termined to have the most limiting power history and was, therefore, analyzed for creep collapse.
The limiting power history was used to calculate the fast neutron flux level for the energy range >l McV. The collapse time for the most limiting assembly was conservatively determined - 4 greater than the three-cycle design life.
The collapse times reported fr Tbble 4-1 are based on the procedures set forth in references 5 and 6.
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- 4.2.2. Ciadding Stress i
The batch 2 and 3 reinserted fuel assemblies are the limiting batches from a cladding stress point of view because of their lower density and longer prc- W vlous exposure time. Batches 2 and 3 have been analyzed and documented in the Crystal River Unit 3 Fuel Densification Report.7 i
I 4.2.3. Cladding Strain The fuel design criteria specify a limit of 1.0% on cladding plastic circum- I ferential strain. The pellet design is established for cladding plastic strain of less than 17. at values of maximum design pellet burnup and heat =
generation rate, which are considerably higher than the values the CR-3 fuel is expected to be. The strain analysis is also based on the maximum specifI- '
cation tolerance for the cladding ID.
4.3. Fuel Thermal Design All fuel assemblies in this core are thermally similar. The fresh batch 5 fuel inserted for cycle 3 operation introduces no significant dif ferences in fuel thermal performance relative to the fuel remaining in the core. The de-sign minimum linear heat rate (LilR) capability and the average fuel temperature i for each batch in cycle 3 are shown In Tabl e 4-2. LIIR capabilities are based on centerline fuel melt and were established using the TAFY-3 code with fuel dens i f icat ion to 96. 5% of theoret ical density.18 4.4. Operating Experience Babcock & Wilcox operating experience with the Mark B 15 x 15 fuel assembly has verifled the adequacy of its design. As of December 31, 1979, the fol-lowing experience has been accumulated for the eight operating B6W 1//-fuel assembly plants using the Mark-B fuel assembly:
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Maximum assembly }
hurnup, mwd /mtU Cumulative net Current electrical output, Reactor cycle Incore Discharged MWh Oconee 1 6 18,610 40,000 28,291,915 Oconee 2 4 27,900 33,700 23,829,671 Oconee 3 5 23,100 29,400 23,492,274 I TMI-l ANO-1 4
4 32,400 23,600 32,200 33,222 23,857,504 21,462,382 !
Rancho Seco 3 37,462 29,378 18,400,06.
I Crystal River 3 2 20,656 15,264 8.718,554 '
Davis Besse 1 1 11,600 --
5,755,467 l l
(a) As of December 31, 1979. i i
As of September 30, 1979 - latest data available.
I Table 4-1. Fuel Design Parameters and Dimensions Batch 2 3 lI 4 5 j Fuel assembly type Mark B-3 Mark B-3 Mark B-4 Mark B-4 Number of assemblies 9 60 56 52
( Fuel rod OD, in. 0.430
, 0.430 0.430 0.430 Fuel rod ID, in. 0.377 0.377 0.377 0.377 W Flexible spacers, type Corrugated Corrugated Spring Spring Rigid spacers, type Ceramic Ceramic Zirc-4 Z i rc-4 Undensifled active fuel length, Jn. 144 14., 143.6 141.79 i Fuel pellet (mean specifled), in. O.370 0.3697 0.370 0.36S6
_ Fuel pellet init ial density (mean speciffed), % TD 92.5 92.5 94.0
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95.0 Initial fuel enrichment, wt % 235 U 2.54 2.83 2.64 2.62 Estimated residence time, EFPH 22,800 22,800 18,720 20,640 Cladding collapse time, EFPH >25,000 >25,000 >30,000 >30,000 l
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I Table 4-2. Fuel Thermal Analysis Parameters Batches 2/3 Batch 4 15atch 5 No. of assemblies 9/60 56 52 Nominal pellet density, % TD 92.5 94.0 95.0 Pellet diameter, in. 0.370 0.3697 0.3686 Stack height, In. 144.0 143.62 141.8 Densified Fuel Parameters "
Pellet diameter, in. 0.3641 0.3648 0.3649 ,
Fuel stack height, in. 141.1 141.8 140.74 Nominal LilR at 2568 MWt, kW/ft 5.77 5.74 5.79 Avg fuel temperature at nominal 1330 1280 1310 LilR , F LilR to centerline fuel melt, kW/ft 19.7 20.1 20.1 Core average densified LIIR at 2544 MWt is 5.71 kW/ft DensifIcation to 96.5% TD assumed.
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- 5. NUCLEAR DESIGN l
5.1. Physics Char ac t er i s t ic si l l
Table 5-1 compares the core physics parameters of cycle:, 2 and 1; these vaines were generated using PDQ07 8 for both cycles.
Since the core ha:. no t yet '
reached an equilibrium cycle, d if f erence:, in core physics parameters are ex-pected between cycles. The longer cycle 3 will produce a larger cycle dif-j ferential burnup than cycle 2. The accumulat ed average core burnup uill be higher in cycle 3 than in cycle 2 because of the presence of the once-burned t
batch 4 and twice-burned batch 2 and 3 f uel. Figure 5-1 illustrates a rep:e-sentative relative power distribution for the beginning of the third cycle at f ul1 power with equilibrium xenon and normal rod positions.
The critical boron concentrations for cycle 3 are given in Table 5-1. Control {
rod worths are suf f icient to at tain the requir( d shutdown margin as Indicated l In Table 5-2. The hot full power control rod worths vary l ittle bet ween cycles <
2 and 3. The ejected rod worths for cycle 3 are higher than those in cycle 2 l f or the san;e number of regulat ing banks insert.ed ; however, values between cy-cles are difficult to compare since the isotopic distributions are different.
Calculated ejected rod worths and their adherence to criteria are considered at all times in life and at all power levels in the development of the rod in-sertion limits presented in rection 8. The maxirvun stuck rod worths for cycle l
3 are greater than those for cycle 2. The adequacy of the shutdown margin l with cycle 3 stuck rod worths is demonstrated in Table 5-2. The following conservatisms were applied f or the snutdown calculations:
- 1. . poison material depletion allowance.
- 2. 10% uncertainty on net rod worth.
- 3. ; lux redistribution penalty.
Flux redistribution was accounted for since the shutdown analysis was calcu-lated using a two-dimensional model. The shutdown calculation at the end of H cycle 3 was analyzed at 250 EFPD. This is the latest time ( 10 EFPD) in core F
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. . , - . - - e. - _ - - . .-,%.. --,,.g -.,,.w_ , ,,wy,,yp-em, ,.,--_9%w.p ..-,y,,, g ww y ,.9-.-%
I life at which the transient bank is nearly fully inserted. After 250 EFPD, the transient bank will be almost fully withdrawn, thus, the available shut-down margin will be increased.
The cycle 3 power deficit from hot zero power to hot full power is identi- ,
i cal to the cycle 2 deficit at BOC, but slightly more negative than the cycle 2 deficit at EOC. I The Doppler coef ficients and xenon worths are similar for the two cycles. The dif ferential boron worths are similar for cycles 2 and 3.
The effective delayed neutron fractions for both cycles show a decrease with burnup.
- 5. 2. Changes in Nuclear Design There is no major change between the designs of cycle 2 and cycle 3; tlie up-
>;rading of the core power level to 2544 MWt was cons # 4 red in cycle 2 oesign and will be implemented in cycle 3. The same calculat onal methods and de-sign information were used to obtain the important nuclear design parameters.
No significant operational or procedural changes exist with regard to exial or radial power shape control, xenon control, or tilt control. The opera-tional limits and RPS limits (Technical Specification changes) for cycle 3 are presented in section 8.
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Table 5-1. Physics Parameters, Crystal River Three, Cycle 3 L
Cycle 2 Cvele 1(")
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Design cycle length, EFPD 275 320 Design cycle burnup, mwd /mtU 8,500 9,881 Design average core burnup - EOC, mwd /mtU 17,364 17,955
~
l Design initial core loading, mtU 82.3 82.3 Crit ical boron - 110C, ppm (no Xe)
IIZl'(b ) , group 8 (37.5% wd) 1,260 I
1,375 IlZ P , groups 7 and 8 inserted 1,185 1,299 IIFP(b> , groups 7 and 8 insert ed 991 1,129 Critical boron - EOC, ppm (eq Xe)
IlZ P) 305 l 290
, gyp 7 group 8(37.5%wd) 51 30 Control rod worths - IIFP, BOC, %Ak/k Group 6 1.02 1.07 Group 7 0.85 0.82 Group 8 (37.5% wd) 0.49 0.48 Control rod worths - IIFP, EOC, %Ak/k Group 7 1.11 W 1. 07 (d )
Group 8 (37.5% wd) g 0.48 0.48 Max ejected rod worth ") - HZP, %Ak/k n0C (N-12) 0.55 0.63 EOC (N-12) 0. 50 (c ) 0. 61 (d ) l i
Max stuck rod worth - IlZP, %Ak/k BOC (N-12) 1.82 EOC (L-14) 1.88(c) 1.
L.92M (d)
Power deficit, IIZP to IIFP, %Ak/k Il0C
-1.30 -1.30 EOC -2.06
-2.I1 Doppler coeff - BOC, 10-5 (Ak/k/"F) 100% power (0 Xe) -1.50 -1.50 Doppler coeff - EOC, 10-5 (Ak/k/ F) 100% power (eq Xe) -1.58 -1.62 Moderator coeff - IIFP, 10 (Ak/k/ F)
Il0C (O Xe, critical ppm, group 8 inserted) -0.65 -0.41 EOC (eq Xe, 17 ppm, group 8 inserted) -2.52 -2.61 Boron worth - IIFP, ppm /%Ak/k BOC 106 107 EOC 94 94 Xenon worth - IIFP, %Ak/k BOC (4 EFPD) 2.67 2.64 EOC (equilibrium) 2.74 2.74 5-3 Babcock 8. Wilcox
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Table 5-1. (Cont'd)
Cycle 2 Cycle 3 Effective delayed neutron fraction - IIFP '
BOC 0.00584 0.00594 EOC 0.00516 0.00519 l (a) Cycle 3 data are for the conditions stated in this report; the cycle 2 values given are at the core conditions identified in reference 2.
II7,P denotes hot zero power (532F T #8); IIFP denotes hot full power (579F T"VE).
(c) Rod worths for EOC-2 are calculated at 225 EFPD, the latest time in core lif e in which the transient bank is nearly full-in.
(d) Rod worrbs for EOC-3 are calculated at 250 EFPD, the latest time in core life in which transient bank is nearly full-in.
(e) Ejected rod worth for groups 5 thorugh 8 inserted.
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if'
\NNk
- @%' \
kby gn*+/ _ EE < EAT,em
- %'i? * -
TEST TARGET (MT-3)
L 'l 1
1 1.0 gm!al l EE 1
- ru Ek'1 m
= .
l,l g D bb \'
l l.8 i
1.25 IA 1.6 1
k MICROCOPY RESOLUTION TEST CHART
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i Table 5-2. Shutdown Margin Calculation for Crystal River 3, Cycle 3 l
L HOC, %Ak/k EOC " , %Ak/k Ava ilable Rod Worth Total rod worth, lizP( )
I 9.26 9.26 L Worth reduction due to burnup of poison material
~0.37 -0.42 m-Maximum stuck rod worth, HZP -1.85
% -1.92 Net worth 7.04
~
6.92 Less 10% uncertainty -0.70 -0.69 Total available worth 6.34 6.23 Required Rod Worth Power deficit, liFP to !!ZP 1.30 2.08 Max allowable inserted rod worth 1.06 1.36 Flux redistribution 0.55 1.05 p .
L Total required worth 2.91 4.49 Shutdown Margin Total available minus total required 3.43 1.74 Note: Required shutdown margin is 1.00% Ak/k.
" For shutdown margin calculations, this is defined as s250 EFPD, the latest in.
time in core life in which the transient bank is nearly full-(b)liZP: hot zero power, lil'P : hot full power.
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Figure 5-1. BOC (4 EFPD), Cycle 3 Two-Dimensional Relative Power Distribution - IIFP, Equilibrium Xenon, Banks 7 and 8 Inserted 8 9 10 11 12 13 14 15 11 1.03 1.09 1.26 1.14 1.33 0.95 0.47 0.53 i
1 1.22 1.10 1.28 1.14 1.23 0.83 0.60 I-I I,
N 0.65 N 1.02 1.14 8
1.25 1.12 0.57 M 1.01 1.28 1.06 0.92 l
l N 1.12 1.12 0.64 I
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0.68 P
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l Inserted rod group No.
Nx x.xx Relative power density I I
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- 6. TllERMAL-ilYDRAULIC i)ESIGN I 6.1. DNBR Evaluations Crystal River 3 will be upgraded in power for cycle 3 operation from 2452 to I 2544 MWt rated core power. Thermal-hydraulic design calculat ions in suppt>rt of cycle 3 operation assumed a rated power level of 2568 MWt for consistency with other B&W reactors and used the analytical methods documented in the l Final Safety Analysis Report l and updated in the Fuel Densification Report.7 The following changes in thermal-hydraulic conditions or assumptions were
- made for cycle 2 and 3 evaluations.
- 1. The B6W-2 CHF correlation 3 was used instead of the W-3 correlation. The l B6W-2 correlation, a realistic prediction of the burnout phenomenon, has been reviewed and approved for t3e with the Mark-B fuel assembly design.
. This correlation was used for the Crystal River 3, cycle 2 reload report ?
and is currently used to license all operatin; B&W plants with Mark-B fuel assembly cores.
- 2. The assumed system flow was changed from 105% (cycle 1) to 106. 5 (c yc l e a l 2 and 3) of the design flow of 88,000 gpm/ pump primarily to nuke the ther-mal-hydraulic design basis for Crystal River 3 consistent with that as-sumed for other B&W plants of similar design and ratt! power level (e.g.,
l L Oconee 1, 2, 3, ANO-1, and TML-1). This assumption is fully just if;ed by measured flow data from Crystal River 3, which Indicates a syst em flow in r~
L excess of 109.sz of design flow, including aitomance ror measurement error.
- 3. The fresh incoming batcl; 5 fuel inserted for cycle 3 is the Mark 3-4 as-sembly design. Batches 2 and 3 are Mark B-3 assemblies, shile batch +
is Mark B-4. The Mark B-4 fuel assemblies differ from the Mark B-3 assem-blies primarily in the end fittings, which have been modified to reduce assembly pressure drop and increase holddown margin. The reduced assem-bly pressure drop causes a s1ight increase in flow through the B-4 u
6-1 Babcock & Wilcox
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assemblies relative to the B-3 design. No credit has been taken in thermal-hydraulic evaluations for any increase in B-4 assembly flow re-sulting from a mixed core that includes Mark B-3 assemblies. Sin 11ar core "
configurations (Mark B-3 in combination with Mark B-4 assemblies) have succes.sfully operated in a number of B&W reactors, including Oconce 1, 2, 3, ANO-1, and TMI-1. Mark B-4 assemblies are currently in all B&W operating reactors.
- 4. A rod bow penalty has been calculated according to the p ocedure approved in reference 10. The burnup used is the maximum fuel assembly burnup of the batch that contains the limiting (maximum radial x local peak) fuel assembly. For cycle 3, this burnup is 31,358 mwd /mtU in a batch 3 assem-bly. The resultant net rod bow penalty after inclusion of the 1% f. low area reduction factor credit is 2.8% reduction in DNBR. The rod bow pen-alty is more than offset by the 10.2% DNBR margin included in trip set-points and operating limits.
5.
A reference design radial > local power peaking factor (Fg) of 1.71 was used for cycle 2 and 3 evaluations. The cycle 1F of 1.78 was reduced g AH to 1.71 in conjunction with ORA and BPRA removal. Il E 6 The densification power spike was eliminated from DNBR evaluations based on the NRC approval of this change in reference 12.
The cycle 1, 2, and 3 maximum design conditions and significant parameters are shown in Table 6-1.
6.2. Pressure-Temperature Limit Analysis The pressure-temperature limit curves for four- and three-pump operations are shown in Figure 8-4. The most limiting of these curves (four pump) provides '
the basis for the RPS variable-low pressure trip function. The curves are based on a minimum DNBR of 1.433, which provides 10.2% margin to the CHF cor-relation limit. The margin is incorporated to provide flexibility for future cycle designs to avoid the potential need for revising setpoints on a cycle- '
by-cycle basis. W 6.3. Flux / Flow Trip Setpoint Analysis The flux / flow trip is designed to protect the plant during pump coastdowns from four-pump operation or to act as a high flux trip during partial-pump 6-2 Babcock & Wilcox l ,
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B operation. Crystal River 3, cycle 3, will have redundant pump monitors on each pump, which will trip the reactor inenediately upon the loss of power to two or more pumps. Therefore, the flux / flow trip setpoint need only protect i
the plant during a one-pump coastdown f rom four-pump operat ion.
The margin for any assumed flux / flow setpoint is determined with a transient i
analysis of a cne pump coastdown initiated f rom 102% indicated power (108%
real power). The 6% full power difference between real power and ind ica t ed power account s for 4% FP neutron power measurement error and a 2% Fp heat balance error. Actual measured one-pump coastdown data are used in the anal-l ysis, and m4ximum additive trip delays are used betweeen the time trip condi-tions are reached and actual control rod motion starts. Once a flux / flow j trip limit is found to be adequate by thermal-hydraulic analysis, error ad-3 justments are made to account for flow measurement noise and instrument error I
. before the actual trip setpoint is determined.
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The recommended cycle 3 thermal-hydraulic flux / flow trip limit of L.10 (actual in-plant setpoint of 1.07) resulted in a transient minimum DNBR of 1.75 (BW-2) during the pump ocastdown. This represents > M% DNBR margin to the correla-t.on limit of l.30.
6.4. Loss-of-Coolant-Flow Transients The one pump coastdown analysis was discussed in conjunct.fon with the flux / flow I setpoint analysi, in section 6.3.
transients were also analyzed for 2568 MWt.
The four-pump coastdown and locked-rotor The results of these anal,vse e discussed in acetion 7, "Acc ident and Transient Analysis."
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I Table 6-1. Cycle 1, 2, and 3 Thermal-Hydraulic Design Conditions Cyc19 1
<268.8 EFP2 Cycle 1
>268.8 EFpD Cycles 2&J 2544 MWt Il Design power level, MWt 2452 2452 2568 System pressure, psia 2200 2200 2200 Reactor coolant flow,
% design 105 105 106.5 Ref design radial x local g power peaking factor, g FAH 1.78 1.71 1.71 Ref design axial flux shape 1.5 cosine 1.5 cosine 1.5 cosine Hot channel factors Enthalpy rise 1.011 1.011 1.011 Heat flux 1.014 1.014 1.014 Flow area 0.98 0.98 0.98 Densified active length, in. 141.12 140. 2 (b) 140. 2 (b )
Avg heat flux at 100%
power, Btu /h-ft2 167 x 103 168 x 103 176 x 103 Max heat flux at 100%
power, Btu /h-ft2 446 x 103 (" 431 x 10 3 452 x 103 CHF correlation W-3 B&W-2 B&W-2 =
Minimum DNBR (J. power) 1.61 (114) 2.14 (112) 1.98 (112) m.
1.92 (102) 2.27 (108) 2.12 (108) g 2.49 (102) 2.33 (102)
(a)The maximum heat fluxes shown are based on reference peaking and average flux. For cycle 1, thermal hydraulic calculations also includ-ed the densification spike factor in the DNBR calculations. B&W no longer considers this spike factor in DNBR calculations, as described in reference 7 and accepted in reference 12.
j 140.2 inches is a conservative (minimum) value used in cycle 2 and 3 analyses; it is the minimum densified length for any B&W fuel. Spe-cific densified lengths for CR-3 fuel are given in Table 4-2.
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- 7. ACCIDENT AND TRANSIENT ANALYSIS 7.1. General Safety Analysis
.- Fach FSARI accident analysis has been examined with respect to changes in cy- l
- cle 3 parameters to determine the effect of upgrading the reactor power from 2452 to 2544 MWt. llecause the FSAR accident analysis, with the exception of the four-pump coastdown and locked-rotor accidents, was done at a higher power
{ 1evel than the requested upgrade (i.e., 2568 versus 2544 MWt), it was only necessary to examine the cycle 3 parameters relative to the FSAR values to en-
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sure that the thermal performance during hypothetical transients is not degrad-ed. Alt hough the FSARI states that all accident.s were done at 2544 MWt, they were actually analyzed using the more conservative 2568 MWt.
The ef f ects of f uel densification on the FSAR accident ant,1ysis results have been evaluated and are reported in reference 7. Since batch 5 reload fuel as-m semblies do not contain fuel rods whose theoretical density is lower than those
- considered in reference 7, the conclusions (with the exception of the four-
- pump coastdown and locked-rotor accidents) in reference 7 are still valid.
These two accidents have been re-evaluated at 102% of 2568 MWL for consistency with other ll&W reactors using the analytical methods documented in the FSAR I and updated in the Fuel Densification Report.7 The input parameters used for these accidents are given in Table 6-1 and section 7.6. The letdown line rup-ture is analyzed in sectic. 7.16, the environmental dose assessnent for all
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accidents is summarized in 1-ection 7.18.
7 2.
,_._ Accident Evaluation The key parameters that have the greatest effeet on determining the outcome of a transient can typically be classified in three major areas. core thermal parameters, thermal-hydraulic parameters, and kinetics parameters, including the reactivity feedback coefficients and control rod worths. Fuel thermal analysis parameters for each batch in cycle 3 are given in Table 4-2.
7-1 Babcock & Wilcox
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I, Table 6-1 compares the cycle 3 thermal-hydrau'.ic maximum design conditions t
, to the previous cycle values, and a comparison of the key kinetics parameters from the FSAR and cycle J is provided in Table 7-1. Table 7-2 is a tabulation (
showing the bounding values for allowable 1.0CA peak linear heat rates for Crystal River 3, cycle 3 fuel.
It is concluded from the loss-of-flow analysis (seetIon 7.6) and by examination of cycle 3 core thermal and kinetics propert ies wit h respect to acceptable FSAR values that this core reload will not adversely affect the abilit y to safely operate the Crystal River 1 plant during cycle 3. ConsiderInn t he previously accept'ed design basis used in the FSAR, the t ransient evaluation of cycle I is considered to be hounded by previously accept ed analyse: . The init !al con-l ditions of the transients in cycle 3 are bounded by the FNdt wi th t he excep- E tion of the t our pump coast dewn and locked rotor accidents, which were redone at a core power of 1022 of 2568 t!Wt.
- 7. 3. Rod Withdrawa1 Acc ident s This accident is det ined as uncontrolled reactivity addit ion to the core due to withdrawal ot control rods during 4tartup conditions or from rated power conditions. Both types of incident s were analyzed in the FSAR.
The import ant parameters during a rod withdrawal acc ident are Doppler co-efficient, moderator temperature coetfielent, and the rate at which reactiv- E it y is added to the core. Only high-pressure and high-flux trips are accounted for in the FSAR analysis, innoring multiple alarms, interlocks, and trips that normally preclude this type of incident.
l For positive reac t ivity addit ion indicative of these events, the most severe 1
l resul ts occur for 110L conditions. The FSAR values of the key parameters for BOL conditions were -1.17 x 10-5 Ak/k/aF for the Doppler coef fic ient , 0.0 Ak/k/ F for the moderator temperature coef f ic ient and rod group worths up to ani including a 12.97: A/k/k rod worth. Comparable cycle 3 parametric values are -1.5 5 Ak/k/oF for Doppler coefficient, -0.41 10-4 Ak/k/"F for moderator temperature coef t le ient , and maximum rod bank worth of 9.26% Ak/k.
The FSAR analyses used an init ial rated power l evel of 2568 51Wt with a reactor trip at 112% of 2568 StWt. For the acc iden t s that trip on high flux, this is more conservative than initializing the accident at 10.'Z o f 2 54 4 31Wt and trip-ping the reactor at 114% oi 2544 since more energy is added to the system in l
7-2 Babcock & Wilcox l 1
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_ _ _ _ _ - . _ - _ _ - . __ - .._._ _ ___._ .__.__-. . . ~ _ _ _ _ _ - - _ _ _ - - - - _ _ _ _ _ _ _ _ _ _ _ --
the FSAR analysis. For the accidents that trip on high pressure, the pressure trip would occur a little sooner with the higher initial power level (2594 MWt
= 102% of 2544 MWt) than with the lower initial power used in the FSAR (2568
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MWt). Therefore, cycle 3 parameters are bounded by design values assumed for the FSAR analysis. Thus, for the rod withdrawal transients, the consequences will be no more severe than those presented in the FSAR and the Fuel Densifi-cation Report.
7.4. Moderator Dilution Accident A
Boron in the form of boric acid is used to control excess reactivity. The
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boron content of the reactor coolant is periodically reduced to compensate l
for fuel burnup and transient xenon effects with dilution water supplied by the makeup and purification system. The moderator dilution transients con- l sidered are the pumping of water with zero boron concentration from the make-l__
up tank to the RCS under conditions of full power operation, hot shutdown, and l L, refueling.
The key parameters in this analysis are the initial boron concentration, boron reactivity worth, and moderator temperature coefficient for power cases. l For positive reactivity addition of this type, the most severe results occur
, for BOL conditions. The FSAR values of the key parameters for BOL conditions were 1150 ppm for the initial boron concentration, 100 ppm /1% Ak/k boron re-activity worth and +0.5 x 10-4 Ak/k/*F for the moderator temperature coeffi-cient. Comparable cycle 3 values are 1129 ppm for the initial boron concen-tration, 107 ppm /1% ok/k boron reactivity worth and -0.41 x 10-4 ok/k/ F for the moderator temperature coefficient. The FSAR used an initial rated power level of 2568 MWt for these accidents. The effect of a higher initial power (i.e., 102% of 2544 MWt) is to cause the pressure trip to occur sooner.
The FSAR shows that the core and RCS are adequately protected dur hg *b's event. Sufficient time for operator action to terminate this transient is also shown in the FSAR even with maximum dilution and ninimum shutdown margin.
The predicted cycle 3 parameter values of importance to moderator dilet ion transient are bounded by the FSAR design values, thus, the analysis in the FSAR is valid.
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I 7.5. Cold Water (Pump Startup) Accident The NSS contains no check. or isolation valves in the RCS piping; therefore, the classical cold water accident is not possible. However, when the reactor is operated with one or more pumps not running, and the pumps are then started, g
the increased flow rate will cause the average core temperature to decrease. W if the moderator temperature coefficient is negative, reactivity will be added to the core and a power increase will occur.
Protective interlocks and administrative procedures exist to prevent the starting of idle pumps if reactor power is above 22%. However, these restric-tions were not assumed, and two-pump startup from 50% of 2568 MWt power was l analyzed as the most severe transient. The initial power level of 50% of 2568 MWL is slightly more conservative than initializing the transient at 50%
of 2544 MWt.
To maximize reactivity addition, the FSAR analysis assumed the most negative moderator temperature coefficient of -4.0 x 10- Ak/k/ F and least negative Doppler coefficient of -1.17 x 10 '* Ak/k/*". The corresponding most negative i moderator temperature coefficient and least negative Doppler coefficient pre-dicted for cycle 3 are -2.63 x 10-4 Ak/k/ F and -1.5 x 10% Ak/k/ F, respec-tively. As the predicted cycle 3 moderator temperature coefficient is less l
negative and the Doppler coefficient is more negative than the values used in El 5
l l the FSAR, the transient results would be less severe than those reported in ,
! the FSAR.
7.6. Loss of Coolant Flow (LOCF)
A reduction in reactor coolant flow can be caused by mechanical failure or a I i loss of electrical power to the pumps. The LOCF transients were re-analyzed for cycle 3 operation and assumed an initial power level of 102% of 2568 MWt 3 for consistency with other B&W reactors.
7.6.1. Four-Pump Coastdown (4PCD) ll l l
The 4PCD transient has been analyzed under conditions that represent the most conservative that can occur for cycle 3 operation. These conditions include gj
- l such key parameters as initial flow rate, flow rate versus time for the tran-sient, initial power level, Doppler coefficient, moderator temperature coef fi-cient, and reference design radial x local power peaking factor ( Pt.H ) . Table 7-3 compares the key parameters used in the analysis with those predicted for 7-4 Babcock & Wilcox
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cycle 3. For all parameters, the value used in the analysis is either equal j to the cycle 3 parameter or is more conservative.
The results of the analysis are shown on Figure 7-1. The minimum DNBR of l 2.10 (BAW-2) obta ined dur ing the transient is well above the DNBR correlation limit of 1.30. The luel and cladding t emperatures are not shown since there
- was no increase in these parameters, it is therefore concluded that no fuel i
damage wilI occur.
l Table 7-4 provides a comparison of MDNBRs between the FSAR, Fuel Densitiration Report, Cycle 2, and cycle 3 for both one- and four-pump coastdowns Add it ion-al DNBR mare,in is shown for cycles 2 and 3 due t.o the use of t. he BhW-2 CHF co r-relation instead of the b-3 CHF correlat lon, j 7.6.2. Lockei Rotor (LR)
The locked-rotor accident has been analyzed under cond it ions that represent the j most conservative that can occur for cycle 3 operation. These conditions are the same as those in section 7.6.1 (4 PCD). Table 7-3 comparet the key parame-ters used in the analvsis with those predicted for cycle 3. For all parameters, the value used in the analysis is either equal to the cycle 3 parameter or is
!I more conservative.
I The results of the analysis are shown on Figure 7-2. The maxiraum fuel temper-ature does not exceed the initial centerline fuel temperature of 440JF.
)
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1 This temperature start s to decrease around 2 seconds into the acc iden t . The
{
analysis for the maximum transient cladding and fuel temperatures conservativelv assumed film boiling at a DNBR of 1.43 instead of the c o r re l a t. i o n 1imit of I 1.30 (refer to section 6). The DNBR reached the 1.43 value at approximately 1.2 seconds, after which the cladding temperature increased to a maximum of Il20F at S.5 seconds after initiation of the accident. Less than 0.5* of the f fuel pins in the core will experience a DNBR of less than 1.41. and no pins will experience a DNBR less than 1.00. For those pins that experience DNB, the cladding temperature will not exceed ll20F.
7.7. Stuck-Out, Stuck-In, or Dropped
}
Control Rod Accident L.
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 averace coolant temperature. In addillon, the power distribution may he distorted due to the 7-5 Babcock & Wilcox
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new control rod insertions. Therefore, under these condit:.ons, a return to ;
rated power may lead to localized power densities and heat fluxes in excess of design limitations.
The key parameters for this transient are moderator temper ture coefficient, worth of dropped rod, and local peaking factors. The FSAR analysis was based on 0.40% ok/k rod worth with a moderator temperature coefficient of -3.0 x 10 '*
Ak/k/*F. For cycle 3, the maximum worth dropped rod at pover is 0.20% Ak/k and the moderator temperature coefficient is -2.63 x Ak/k/' F. Since the pre-dicted rod worth is less and the moderator temperature coef ficient more posi-tive, the consequences of this transient are less severe than the results presented in the FSAR.
The effect of initializing these accidents at 2568 MWt as cone in the FSAR versus using 102% of 2544 MWt is judged insignificant or slightly beneficial since as shown in Figures 14-20 and -21 of the FSAR, the parameter of primary Starting the act.ident at a higher power level concern is low system pressure.
g (i.e., 102% of 2544 MWt) would yield slightly higher systen pressures. E 7.8. Loss of Electric Power Two types of power losses were considered in the t'SAR: a loss of load condi-tion, caused by separation of the unit from the transmission system, and 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 with and without turbine runback.
When there is no runback, a reactor trip occurs on high RC pressure or tempera-ture. This case resulted in a non-limiting accident. The limiting accident for offsite dose considerations thus becomes the loss of all electrical power except unit batteries, and assuming operation with failed fuel and steam gen- E erator tube leakage. The environmental dose assessment is presented in sec-tion 7.18.
7.9. Steam Line Failure j
l A steam line failure is defined as a rupture of any of the steam lines f rom I t
the steam generators. Upon initiation of the rupture, both steam generators start to blow down, causing a sudden decrease in primary system temperature, g pressure, and pressurizer level. The temperature reduction leads to positive reactivity insertion and the reactor trips on high flux or low RC pressure.
7-6 Babcock & Wilcox
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The FSAR has identified a double-ended rupture of the steam line between the l
steam generator and steam stop valve as the worst-case situation at end-of-life conditions.
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The key parameter for the core response is the moderator temperature coeffi-cient which in the FSAR was assumed t o be -3.0 x 10 4 Ak/k/ F. The cycle 3 i
predicted value of moderator temperature coef ficient is -2.63 x 10-4 Ak/k/ F.
This value is bounded by that used in the FSAR arial sis; hence, the results in th2 FSAR represent the worst situation.
i The FSAR used an Initial pt wer level of 2568 MWL for these accidents This is l more conservative than rutuing the accident at 102% of 2544 Et and tripping the reactor at 110% versue the current 112% setpoint since more energy is added to the system for the FSAE analysis i
7.10. Steam Generator Tub Failure A rupture or leak in a steam generator tube allows reactor coolant and associ-ated activity to pass to the secondary system. The FSAR analysis is based on complete severance of a steam generator tube. The primary concern f or t his incident is the potential radiological release. The environmental dose assess-7 ment is presented in section 7.18.
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7_. l_L . __Fue l Ha nd ling Acc iden t The mechanical damage type of accident is considered the maximum potent tal source of activity release during fuel handling activity. The primary con-F cern is over radiological releases.
L The environmental dose assessment is pre-sented in section 7.18.
e j 7.12. Rod Ejection Accident For reactivity to be added to the core at a more rapid rate than by uncont.iolled rod withdrawal, physical failure of a pressure barrier component in the CRDA must occur. Such a failure could cause a pressure differential to act on a CRA and rapidly eject the assembly from the core. This incident represents the most rapid reactivity insertion that can be reasonably postulated. Th-r-
L values used in the FSAR and densification report at BOL conditions of -1.17 L
10-5 Ak/k/*F Doppler coefficient., 0.0 Ak/k/*F moderator temperature coefficieut, and ejected rod worth of 0.65% Ak/k represented the maximum possible transient.
L 7-7 Babcock & Wilcox
-..-vm,--,-----,e,--w-r--,--e.v---w--eve-
I The use of a 0.65% Ak/k maximum rod worth is conservative in comparison to the cycle 3 predicted value of 0.63% Ak/k. Furthermore, the cycle 3 predicted values of -1.5 x 10-5 Ak/k/ F Doppler and -0.41 x 10-5 Ak/k/ F moderator tem-perature coefficient are both more negative than used in the FSAR analysis.
The FSAR used an initial rated power level of 2568 MWt for this accident. This is more conservative than initializing the accident at 102% of 2544 MWt and tripping the reactor at 110% versus the current 112% setpoint since more energy is added to the system for the FSAR analysis. For the accident which trip on high pressure, the effect of higher initial power level (i.e., 102% of 2544 MWt) is to cause the pressure trip to occur slightly sooner. Since the FSAR input bound the cycle 3 predicted values, the results in the FSAR and densifi-cation report are applicable to this reload 7.13. 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 hypo-thetical accident is based solely on a gross release of radioactivity to the reactor building. The environmental dose assessment is presented in section 7.18.
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- 7. 14. Waste Gas Tank Rupture The waste gas tank was assumed to contain the gaseous activity evolved from degassing all the reactor coolant following 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 environmental dose assessment is presented in section 7.18. =
7.15. LOCA Analysis Generic LOCA analyses for B&W 177-FA lowered-loop NSSs have been performed 1
using the Final Acceptance Criteria ECCS Evaluation Model. The large-break l analysis is presented in a topical report l 3, and is further substantiated in a l 1etter report l4 The staall break analysis is presented in a letter report 15, These analyses used the limiting values of key parameters for all olants in the l
l category. Furthermore, the average fuel temperature as a functio.. >f linear heat rate and lifetime pin pressure data used in the LOCA limits analysis l3 atu conservative compared to those calculated for this reload. Thus, these 7-8 Babcock & Wilcox
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analyses and LOCA limits provide conservative results for the operation of Crystal River Unit 3 at 2544 MWt.
Crystal River Unit 3's proposed long-term ECCS modification for small break
- LOCA is presented in reference I ti .
The LOCA analyses used a power level of 2772 IfWt, which is conservat Ive rela-
[ tive to the 2544 MWt rating. Table 7-2 shows the bounding values for allow-able LOCA peak linear heat rates for Crystal River Unit 3, cycle 3.
- 7.16. Failure of Small Lines Carrying Primary Coolant Outside Containment j 7.16.1. Ident i ficat ion of Causes A break in fluid-bearing lines that penetrate the containment could result in l the release of radioactivity to the environment. There are no instrument lines connected to the RCS that penetrate the containment. However, other piping i
t lines from the RCS to t.he makeup and purification system and the decay heat removal system do penetrate the containment. Leakage through fluid penetra-tions not serving accident-consequence-limiting systems is minimized by a double-barrier design so that no single credible failure or malfunction of an I active cor ionent will resul t in loss of isolation or intolerable leakage. The installed double barriers take the form of closed piping, both inside and out-side the containment, and various types of isolation valves.
The most severe pipe rupture relative to radioactivity release during normal plant operation occurs in the makeup and purification system. This would be a rupture of the letdown line just outside the containment but upstream of the letdown control valves. A rupture at this point would result in a loss of reactor coolant until the RCS pressure dropped below its low pressure setpoint at 1500 psig;. When this pressure is reached, the emergency injection signal initiates closure of the letdown isolation valve inside the containment, thus terminating the accident.
7.16.2. Analysis of Effects and Consequences 7.16.2.1. Safety Evaluation Criteria The sa Nty evaluation criterion for this accident is that resultant doses shall l not exceed 10 CFR 100 limits. i l
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- 7. l fi. 2. 2. Methods of Analysis The CRAFT 2 computet code f' was used to determine the loss-of-coolant charac-teristics of this letdown line rupture accident. The multinode model included a detailed model of the RCS and additional noding simulating the letdown line piping, valves, and coolers. Before the accident, the reactor was assumed to be operating at 2603 MWt with a letdown flow of 140 gpm. A complete severance of the 2.5-inch letdown line between valves MU-V40 or MU-V41 and MU-V49 was assumed. Coincident with this accident, the makeup control valve was assumed to go to a full-open position so that the maximum makeup flow is available.
This assumption extends the time to reactor trip /ESFAS actuation and increases the mass and energy releases to the auxiliary building. Termination of the accident was assumed following ESFAS actuation on low RC pressure (1500 psig) and closure of the letdown isolation valves inside the containment. An instru-ment error of 6% of full range was assumed for the ESFAS actuation pressure, and the letdown iselation valve was assumed closed 7.4 seconds after the ESFAS pressure setpoint was reached. The 7.4-second time period for the complete valve closure considers both the instrumentation response time and the actual valve closure t ime . Credit was not taken for a reduction in break flow during the t ime the isolation valves were closing.
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7,.16. 2 . 3. Environmental Consequences The time required for the RCS to reach the actuation pressure of 1350 psig '
(1500 psig minus 6 of 2500 psia) for the ESF'S to initiate isolation is con-servatively calculated to be 752 seconds, including valve closure time. For the 2.5-inch letdown line, a total reactor coolant mass of 45,760 pounds is released into the auxiliary building. Ten percent of the iodine contained in the 45,760 pounds of reactor coolant was assumed to volatilize and become air-borne in the auxiliar building. The remaining 90% was assumed to remain in the liquid which drains into the auxiliary building sump tank.
The airborne radioactive nuclides in the auxiliary building are f il t ered through HEPA and charcoal filters in the building's ventilation system before being exhausted to the environment. The ana!ysis is based on a conservatively estimated charcoal filter lodine removal efficiency of 90%. The assumptions used in the evaluation of the offsite doses are summarized in Table 7-5. The atmospheric dispersion f act ors (X/Q) used to calculate the two-hour doses at Il 7-10 Babcock & Wilcox
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the exclusion area boundary and the low population zone boundary are also l
u listed in Tabic 7-5. The fission product activities released to the environ-ment during the accident are listed in Table 7-6.
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7.16.2.4. Results of the Analysis The dose consequences of the letdown line rupture accident are presented in L Table 7-7. The table presents (1) the thyroid dose due to inhalation of iodine activity, (2) the whole body doses from gamma radiation due to immersion in the gas cloud, and (3) the skin doses from beta radiation due to immersion in the cloud for individuals located at the outer boundaries of either the exclu-sion area or the low populatien zone for the first two houra after the acci-dent. The resulting doses are small fractions of the 10 CFR 100 limits.
7.17. Ma in Feedwater Line Break A feedwater line f a ilu re is defined as a rupture of the feedwater line to the
{ steam generator. The rupture results in a reduction in the heat removal from the primary coolant system. With this reduction the reactor coolant system pressure and temperature will increase until the reactor trips on high reac-I tor coolant pressure at 11.8 seconds after the break. The FSAR analyzed the rupture c! the main feedwater header at the steam generator inlet nozzles as the worse case, since this case results in the most rapid steam generator I blowdown.
Because the feedwater accident is an overheating event, BOL values of Doppler I and moderator coefficients represent the most positive reactivity addition to the core.
Table 7-1 shows that the FSAR value for these parameters are more positive than the cycle 3 value, (i.e., FSAR used - 1.17 x 10-5 Ak/k/F and 0 Ak/k/F for the Doppler and moderator coefficients respectively, while cycle 3 pred icts
-1.5 4 10-5 Ak/k/F and -0.41 x 10~4 Ak/k/F for these two parameters). There-fore, the cycle 3 value is bounded by the FSAR analysis and the FSAR represents the worst s i tua t ion .
The effect of a higher initial mwer level on this accident (i.e., IUM of 2544 MWt) is to cause the pressure ocip to occur 0.14 seconds sooner and the peak system pressure to be 17 psi greater, still within the allowable code pressure limit. The reactor coolant system design will accommodate 14'3 minutes of safe shutdown operation at the higher power level. Thereafter, the operator 7-11 Babcock & Wilcox I
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can provide a controlled cooldown of the plant utilizing the auxiliary feed- ll water system and steam relief through the atmospheric or condenser dump valves.
Mnce core coverage can be maintained and reactor coolant system pressure re-main within code allowable limits, the safety evaluation criteria are met.
7.18. Dose Consequences of Accidents The only new detailed dose calculations performed for cycle 3 were for the letdown line rupture accident. The dose consequences of that accident were calculated using the same source terms as the other FSAR analyses and are pre-sented in Table 7-7. The table presents (1) the thyroid dose due to inhala-tion of iodine activity, (2) the whole body doses from gamma radiation due to immersion in the gas cloud, and (3) the skin doses f rom beta radiation due to immersion in the cloud for individuals located at the outer boundaries of ei-ther the exclusion area or the low population zone for the first two hours af ter the accident. The resulting doses are small fractions of the 10 CFR 100 limits.
All dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. A power level of 102% of 2544 MWt (or 2595 MWt) would yield doses approximately 1% higher than those reported in the FSAR. =
Although the dose evaluations in the FSAR were based on conservative values for fuel burnup and power peaking, improved fuel utilization and improved fuel burnup data have resulted in higher plutonium-to-uranium fission rat ios. Since plutonium has a higher iodine fission yield than uranium, more iodine activity W is produced, and thus the thyroid doses are expected to be slightly higher than reported in the FSAR. Generally, the plutonium fission yield for noble gases is lower than for uranium, which would result in lower noble gas inventories that would tend to lower the whole body doses below those reported in the FSAR unless the iodine release is large enough to result in an overall dose increase.
Experience with dose evaluations for reload cores in similar nuclear units indicate that the thyroid doses associated with the FSAR accidents typically increase by less than 20% although some accidents with very small doses have -
on occasion caused increases of a facter of 2. Even with similar increases, all thyroid and whole body doses resulting f rom accidents during cycle 3 are well below the 10 CFR 100 limits. W I
7-12 Babcock & Wilcox
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Table 7-1. Comps ison of Key Parameters for Accident Analysis I
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FSARl ,
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~ densif'n Cycle 3 '
l Parameter value7 Cycle 1 11
/alue BOL Doppler coeff, 10-5 Ak/k/"F -1.17 -1.47 -1.5
[- (268 EFFD)
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EOL Doppler coeff, 10-5 Ak/k/*F -1.30 -1.66 -1.62 (510 EFPD)
I BOL moderator coeff, 10-4 Ak/k/ F 0( ) -0.75 -0.41 (268 EFPD) r b EOL moderator coeff, 10-4 Ak/k/'F -4.0 0) -2.42 -2.63
- (510 EFPD)
All-rod bank worth at BOL, HZP, 12.9 9.12 9.26
% Ak/k (268 EFPD)
Boron reactivity worth (HFP), 100 101 107 ppm /1% Ak/k l
[ Max ejected rod worth (HFP), % Ak/k 0.65 0.55 0.63 Dropped rod worth (HFP), % Ak/k 0.65 0.20 0.20 Initial boron cone'n (HFP), ppm 1150 795 1129
(")+0.50 x 1; 4 Ak/k/ F was used for the moderator dilut"lon accident.
Ih) 3.0 x 10-4 Ak/k/ F was used for the steam line failure analysis and dropped rod accident analysis, a
f 4 Table 7-2. Bounding Values for Allowable LOCA Peak Linear Heat Rates l Core Allowable i elevation, peak LilR, i ft kW/ft l L
2 15.5
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4 16.6 1
g 6 18.0 a
8 17.0 10 16.0 1
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7-13 Babcock & Wilcox
-. , , - . , - , - - - - - - _ , - - - - , - , , . --y - - . , _ - , . _ _ _e_ --.-----,_,-.,,w,,w,--yy
l Il Table 7-3. Input Parameters to Loss-of-Coolant-Flow Transients Cycle 3 value Value used in analysis Initial flow rate, % >109.5 106.5 of 352,000 gpm Flow ra' Vs time > Fig. 14-17, FSAR Fig. 14-17, FSAR (4PCD)
Fig.14-19a, FSAR Fig.14-19a, FSAR (LR)
Initial power 1cvel, 2544 102% of 2568 MW Doppler coeff, Ak/k/ F -1.5 x 10-5 -1.27 x 10-5 Moderator temp coeff. -0.41 x 10 '+ 0 Ak/k/ F =
F/.ll 1.47 1.71 Table 7-4. Summary of Minimum DNBR Results for Limiting -
Loss-of-Coolant-Flow Transients '
Cycle 1 Densif'n u FSARI report Cycle 2 Cycle 3 Transient (W-3) (W-3) (B&W-2) (B&W-2)
One-pump coastdown (flux / flow NR NR 1.75 1.75 trip) l Four-pump coastdown (flux / flow trip, cycle 1; pump monitor 1.45 1.39 2.10 2.10 I j trip, cycles 2 and 3) 1 t
(")NR: not reported.
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Table 7-5. Analysis Assumptions for MUSPS Letdown
- Line Rupture Accident Data and Assumptions Used to Estimate Radioactive Souret*
- Power level, MWt 2568 Percent of fuel rods leaking, % l.0 Table for escape rate coef f 11-11 Reactor coolant activities Nuclide Activityt_mC1/mL 85Kr" 1.6
-~
85 Kr 11 87 Kr 0.87 88 Kr 2.8 131Xe* 2.5
_ 133Xe* 2.9 133 Xe 260 135Xe" 0.97
! 135 Xe L 6.2 138 Xe 0.53 131 1 3.3 I
l 1321 133I 5.0 3.9 134 1 0.52 135I 2.0 i
L-Data and Assumptions Used to Estimate Radioactivity Released Total mass of reactor coolant released 45,760 to auxiliary building, Ib
- Charcoal filter ef ficiency for
- Iodine, % 90 Noble gas, % 0 Fraction of airborne iodine 0.1 E Dispersion Data Exclusion area boundary, meters 1340 F Low population zone boundary, meters 8047 i Atmospheric dispersion percentile, % 5 0-2 h atmospheric dispersion factors, s/m3
- at exclusion area boundary 1.6 x 10-4 at low population zone boundary 1.4 x 10-5 m
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Table 7-6. Activity Released t( Envfronment Due to Rupture of MU6PS LetJNn Line ll Nuclide Activity, Ci 85Kr* 46.1 85Kr 317 l
87Kr 25.1 88 Kr 80.6 131Xe* 72.0 133Xe* 83.5 133 Xe 7490 135Xe 27.9 135xe 179 13e Xe 15.3 1311 0.95 132 I 1.44 l
l 133t 1,12 1 31*I 0.15 1351 0.58 i
I Table 7-7. Radiological Consequences of MU&PS Letdown Line Rupture Outside Containment I
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0-2 b dose at exclusion area boundary, Rem Thyroid (inhalation) 0.111 Whole body (gamma) 0.033 SL.n (beta) 0.049 l)
=
0-2 h dose at low population boundary, Rem ,
Thyroid (inhalation) 9.75 x 10-3 Yhole body (gamma) 2.91 x 10-3 3 kin (beta) 4.27 x 10-3 I
7-16 Babcock & Wilcox
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I Figure 7-1. Four-Pump Coastdown - Ilot Channel MDNBR Vs Time, Crystal River 3 i
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- 8. PROPOSED MODIFICATIONS TO TECHNICAL SPECIFICATIONS I
1 1
All technical specifications have been reviewed by Florida Power Corporation and H&W and some were revised for cycle 3 operation. The Technical Specifi-cation sections to which modifications have been made are listed in Table 8-1 l and are shown on the following pages. The reanalysis of Technical Specifica-tions for cycle 3 operation used the same analytical techniques as the cycle 2 design.2 The review of tant Technical Specifications based on the analyses presented in this report, and the proposed modifications contained in this section, ensure that the Final Acceptance Criteria ECCS limits will not be exceeded nor will the thermal design criteria be violated.
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Table 8-1. Technical Specification Changes Tt a Spec Report page j No. (figure, Nos. (figure tabl,Nos.) Nos.) Reason for change _
- 1. 3 8-3 Rated thermal power increased to 2544 MWt.
2.0 8-4 thru 8-11 Because of the large number of changes to Bases for 2.0 (8-1 thru -4) section 2.0 in cycles 1, 2, and 3, the en-l s Table 8-2 tire section is presented here to avoid E confusion. The flux /Aflux envelopes changed due to the power upgrade. Flux / g flow trips changed witia the addition of the g)
RC pump monitors; the trips are now based on a one pump versus fot -pump coastdown.
3.1.l.6 (3.1- Figures Specs 3.1.3.6, 3.1.3.9, tad 3.2.1 reflect 1, -3, -3. -4) 8-5 thru 8-8 revised nuclear parameters as a result of the cycle 3 reload, including the power up- g grade.
B 3.1.3.9 (3.1- Figures 8-9, 9, -10 8-10 i 3.2.1 (3.2-1, Figures 8-11, l -2) 8-12 3.2.4 (Table Table 8-3 Tilt limits were reduced to reflect in- l 3.2-2) creased detector depletfon. l 3.2.5 (Table Table 8-4 Flow rates were recalculated based on 2544 3.2-1) MWt.
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L 1 1.0 DEFINITIONS I
DEFINED TERMS I
1.1 The DEFINED TERMS of this section appear in capitalized type and are l applicable throughout these Technical Specifications. '
TH_ERMAL POWER I 1.2 THERMAL POWER shall be the +otal reactor core heat transf er rate to the reactor coolant.
RATED THERMAL POWER
- 1. 3 RATED THERMAL POWER shall be a total reactor core heat transfer rate to the reactor coolant of 2544 MWt.
OPERATIONAL MODE
- 1. 4 An OPERATIONAL MODE shall corrcspond to any one inclusive combination l of core reactivity condition, power level, and average reactor coolant temperature specified in Table 1.1.
ACTION 1.5 ACTION shall be those additional requirements specified as corollary statements to each principal specification and shall be part of the speci-I fications.
- 1. 6 A system, subsystem, train, component or device shall be OPERABLE or I have OPERABILITY when it is capable of performing its specified function (s).
Implicit in this definition shall be the assumption that all necessary at-tendant instrumentation, controls, normal and emergency electrical power I sources, ccoling or seal water, lubrication or other auxiliary equipment, that are required for the system, subsystem, train, component or device to perform its function (s), are also capable of performing their related sup-port function (s).
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I 2.0 SAFETY LIMITS AND LIMITING SAFETY SYSTEM SETTINGS I
2.1. SAFETY LIMITS REACTOR CORE 2.1.1. The combination of the reactor coolant core outlet pressure and out-let temperature shall not exceed the safety limit shown in Figure 2.1-1.
APPLICABILITY: MODES 1 and 2.
ACTION:
When the point defined by the combination of reactor coolant core outlet pressure and outlet temperature has exceeded the safety imit, be in HOT STANDBY within one hour.
REACTOR CORE 2.1.2 The coubination of reactor THERMAL POWER and AXIAL POWER IMBALANCE 3 shall not exceed the safety limit shown in Figure 2.1-2 for the various com- 5 binations of three and four reactor coolant pump operation.
APPLICABILITY: M0 E 1.
ACTION:
4henever the point defined by the combination of Reactor Coolant System flow, AXIAL POWER IMBALANCE and THERMAL POWER has exceeded the appropriate safety 1t.it be in HOT STANDBY within one hour.
REACTOR COOLANT SYSTEM PRESSURE 2.1.3 The Reactor Coolant System pressure shall not exceed 2750 psig.
APPLICABILITY: MODES 1, 2, 3, 4 and 5.
ACTION:
MODES 1 and 2 Whenever the Reactor Coolant System pressure has exceeded 2750 g psig, be in HOT STANDBY with the Reactor Coolant System pres- g sure within its limit within one hour.
MODES 3 and 4 Whenever the Reactor Coolant System pressure has exceeded 2/50 g psig, reduce the Reactor Coolant System pressure to within its g limit within 5 minutes.
2.2. LIMITING SAFETY SYSTEM SETTINGS l REACTOR PROTECTION SYSTEM STEPOINTS 2.2.1 The Reactor Protection System instrumentation setpoints shall be set consistent with the Trip Setpoint values shown in Table 2.2-1.
I 8-4 Babcock & Wilcox
SAFETY LIMITS AND LIMITING SAFETY SYSTEM SETTINGS APPLICABILITY: As shown for each channel in Table 3.3-1.
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ACTION:
With a Reactor Protection System instrumentation setpoint less conservative than the value shown in the Allowable Values column of Table 2.2-1, declare the channel inoperable and apply the applicable ACTION statement requirement I of Specification 3.3.1.1 until the channel is restored to OPERABLE status with its trip setpoint adjusted consistent with the Trip Setpoint value.
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2.1 SAFETY LIMITS BASES 2.1.I and 2.1.2 REACTOR CORE The restrictions of this safety limit prevent overheating of the fuel cladding and possible cladding perforation which would result in the re-lease of fission products to the reactor coolant. Overheating of the fuel ,
cladding is prevented by restricting fuel operation to within the nucleate l boiling regime where the heat transfer coefficient is large and the cladding surface temperature is slightly above the coolant saturation temperature. '
Operation above the upper boundary of the nucleate boiling regime would i
result in excessive cladding temperatures because of the onset of departure i from nucleate boiling (DNB) and the resultant sharp reduction in heat trans- (
fer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POWER and reactor coolant temperature and pressure have been related to DNB through the BAW-2 DNB correlation. The DNB correlation has been developed to predict the DNB flux and the location of DNB for axi-ally uniform and nonuniform heat flux distributions. The local DNB heat flux ratio, DNBR, defined as the ratio of the heat flux that would cause DNB at a particular core location to the local heat flux, is indicative of the margin to DNB.
The minimum value of the DNBR during steady-state operation, normal op-erational transients, and anticipated transients is limited to 1.30. This value corresponds to a 95? probability at a 95' confidence level that DNB will not occur and is chosen as an appropriate margin to DNB for all operat-ing conditions.
The curve presented in Figure 2.1-1 represents the conditions at which a minimum DNBR - 1.30 is predicted for the maximum possible thennal power, 112 when the reactor coolant flow is 139.7 x 106 lb/h, which is 106.55 of l the design flow rate for four operating reactor coolant pumps. This curve 3 is based on the f*'nwing nuclear power peaking factors with potential fuel 5 densification effects:
F = 2.57; F = 1.71; F = 1.50.
g l
The design limit power peaking factors are the most restrictive calculated at f ull power for the range from all control rods fully withdrawn to minimum E'
5,
- allowable control rod withdrawal, and form the core DNBR design basis.
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SAFETY LIMITS F
BASES
- The reactor trip envelope appears to approach the safety limit more l closely than it actually does because the reactor trip pressures are measur ed at a location where the indicated pressure is about 30 psi less than core out-let pressure, providing a more conservative margin to the safety limit.
The curves of Figure 2.1-2 are based on the more restrictive of two ther-mal limits and account for the effects of potential fuel densification and
- potential fuel rod bow:
- 1. The 1.30 DNBR limit produced by a nuclear power peaking factor of F" = 2.57 or the combination of the radial peak, axial peak andhositionoftheaxialpeakthatyieldsnolessthana1.30 DNBR.
- 2. The combination of radial and axial peak that causes central fuel melting at the hot spot. The limit is 19.7 kW/ft.
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Power peaking is not a directly observable quantity and therefore limits
__ have been established on the basis of the reactor power imbalance produced by the power peaking.
r u The specified flow rates for curves 1. and 2 of Figure 2.1-2 correspond to the expected minimum flow rates with four pumps and three pumps respective-ly.
The curve of Figure 2.1-1 is the most restrictive of all possible reactor coolant pump-maximum thermal power combinations shown in BASES Figure 2.1.
p The curves of BASES Figure 2.1 represent the conditions at which a minimum L DNBR of 1.30 is predicted at the maximum possible thermal power for the num-ber of reactor coolant pumps in operation.
These curves include the potential effects of fuel rod bow and fuel densification.
{ The DNBR as calculated by the BAW-2 DNB correlation continually increases from point of minimum DNBR, so that the exit DNBR is always higher. Extrapo-lation of the correlation beyond its published quality range of 22% is justi-fied on the basis of erperimental data.
BASES For each curve of BASES Figure 2.1, a pressure-temperature point above and to the left of the curve would result in a DNBR greater than 1.30 or a local quality at the point of minimum DN3R less than 22% for that particular reactor coolant pump situation. The 1.30 DNBR curve for four pump operation is more restrictive than any other reactor coolant pump combination because any pressure / temperature point above and to the left of the four pump curve will be above and to the left of the other curves.
8-7 Babcock & \Milcox
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! SAFETY LIMITS I'
2.1.3 REACTOR COOLANT SYSTEM PRESSURE The restriction of this Safety Limit protects the integrity of the Reactor Coolant System from overpressurization and thereby prevents the re-i lease of radionuclides contained in the reactor coolant from reaching the E containment at.nosphere. E l The reactor pressure vessel and pressurizer are designed to Section III of the ASME Boiler and Pressure Vessel Code which permits a maximum transient l pressure of 1104, 2750 psig, of design pressure. The Reactor Coolant System i piping, valves and fittings, are designed to USAS B 31.7, February,1968 Draft Edition, which permits a maximum transient pressure of 110r>, 2750 psig, of component design pressure. The Safety Limit of 2750 psig is therefore consistent with the design criteria and associated code requirements.
The entire Reactor Coolant System is hydrotested at 3125 psig,125% of l
j design pressure, to demonstrate integrity prior to initial operation.
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L 2.2 LIMITING SAFETY SYSTEM SETTINGS p BASES L
2.2.1 REACTOR PROTECTION SYSTEM INSTRUMENTATION SETPOINTS The Reactor Protection System Instrumentation Trip Setpoint specified in Table 2.2-1 are the values at which the Reactor Trips are set for each param-eter. The Trip Setpoints have been selected to ensure that the reactor core and reactor coolant system are prevented from exceeding their safety limits.
Operation with a trip setpoint less conservative than its Trip Setpoint but within its specified Allowable Value is acceptable on the basis that the L_ difference between each Trip Setpoint and the Allowable Value is equal to or less than the drif t allowance assumed for each trip in the safety analyses.
The Shutdown Bypass provides for bypassing certain functions of the Reactor Protection System in order to permit control rod drive tests, zero power PHYSICS TESTS and certain startup and shutdown procedures. The purpose of the Shutdown Bypass RCS Pressure-High trip is to prevent normal operation with Shutdown Bypass activated. This high pressure trip setpoint is lower than the normal low pressure trip setpoint so that the reactor must be tripped before the bypass is initiated. The Nuclear Overpower Trip Setpoint of s 5.0%
_ prevents any si 5nificant reactor power from being produced. Sufficient natu-ral circulation would be available to remove 5.0% of RATED THERMAL POWER if none of the reactor coolant pumps were operating.
[ Manual Reactor Trip The Manual Reactor Trip is a reduadant channel to the automatic Reactor Protection System instrumentation channels and provides manual reactor trip capability.
Nuclear Overpower
. A Nuclear Overpower trip at high power level (neutron flux) provides re-
^
actor core protection against reactivity excursions which are too rapid to be protected by temperature and pressure protective circuitry.
[~ l During normal station operation, reactor trip is initiated when the re-L actor power level reaches 105.5% of rated power. Due to calibration and in-strument errors, the maximum actual power at which a trip would be actuated l could be 112%, which was used in the safety analysis.
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l 8-9 Babcock & Wilcox
l LIMITING SAFETY SYSTEM SETTINGS Il j 1
BASES RCS Outlet Temperature - High_
The RCS outlet temperature high trip 5 619 F prevents the reactor outlet temperature from exceeding the design limits and acts as a backup trip for E all power excursion transients. E
_ Nuclear Overpower Based on RCS Flow and AXIAL POWER IMBALANCE g gt The power level trip setpoint produced by the reactor coolant system flow is based on a flux-to-flow ratio which has been established to accommo-date flow decreasing transients from high power.
The power level trip setpoint produced by the power-to-flow ratio pro-vides both high power level and low flow protection in the event the reactor 3 power level increases or the reactor coolant flow rate decreases. The power 3 level setpoint produced by the power-to-flow ratio provides overpower DNB protection for all inodes of pump operation. For every flow rate there is a g maximum permissible power level, and for every power level there is a minimum g permissible low flow rate. Typical power level and low flow rate combinations for the pump situations of Table 2.2-1 are as follows:
- 1. Trip would occur when four reactor coolant pumps are operating if power is 2 107.0% and reactor flow rate is 100%, or flow rate is 5 93.5% and power level is 100%.
- 2. Trip would occur when three reactor coolant pumps are operating if power is 2 80.2% and reactor flow rate is 74.7%, or flow rate is 5 69.9% ano power is 75%.
For safety calculations the maximum calibration and instrumentation errors for the power level were used.
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a p LIMITING SAFETV SYSTEM SETTINGS BASES r-L The AXIAL POWER IMBALANCE boundaries are established in order to pre-vent reactor thermal limits from being exceeded. These thermal limits are c either power peaking kW/f t limits or DNBR limits. The AXIAL POWER IMBALANCE L reduces the power level trip produced by the flux-to-flow ratio so that the boundaries of Figure 2.2-1 are produced. The flux-to-flow ratio reduces the s power level trip and associated reactor power-reactor power-imbalance bound-l aries by 1.07% for a 1% flow reduction.
l
'sCS Pressure - Low, High and Variable Low l
L The high and low trips are provided to limit the pressure range in
{ which reactor operation is permitted.
During a slow reactivity insertion startup accident fr~n low power or a I
l slow reactivity insertion from high power, the RCS pressure- .gh setpoint is reached before the nuclear overpower trip setpoint. The trip setpoint for RCS pressure-high, 2300 psig, has been established to maintain the system I
l pressure below the safety limit, 2750 psig, for any design transient. The RCS pressure-high trip is backed up by the pressurizer code safety valves for RCS overpressure protection, and is therefore set lower than the set pressure L_ for these valves, 2500 psig. The RCS pressure-high trip also backs up the nuclear overpower trip.
The RCS pressure-low,1800 psig, and RCS pressure-variable low (11.80 Tout F-5209.2) psig, trip setpoints have been established to maintain the DNB ratio greater than or equal to 1.30 for those design accidents that result in I a pressure reduction. It also prevents reactor operation at pressures below I
the valid range of DNB correlation limits, protecting against DNB.
I Due to the calibration and instrumentation errors, the safety analysis used an RCS pressure-variable low trip setpoint of (11.80 Tout F-5249.2) psig.
Reactor Containment Vessel Pressure - High The reactor containment vessel pressure-high trip setpoint, c 4 psig, I provides positive assurance that a reactor trip will occur in the unlikely event of a steam line failure in the containinent vessel or a loss-of-coolant accident, even in the absence of an RCS pressure-low trip.
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8-11 Babcock & Wilcox
Table 8-2. RPS Trip Setpoints Table 2.2-1. Reactor Protection System Instrumentation Trip Setpoints i Functional unit Trip setpoint Allowable values ;
- 1. Manual reactor trip Not applicable Not applicable r
- 2. Nuclear overpower 5 105.5% of RATED THERMAL POWER s 105.5% of RATED THERMAL POWER with four pumps operating with four pumps operating 5 80.2% of RATED THERMAL POWER 5 80.2% of RATED THERMAL POWER with three pumps operating with three pumps operating
- 3. RCS outlet temp-high 5 619 F 5 619*F
- 4. Nuclear overpower Trip'setpoint not to exceed the Allowable values not to exceed based on RCS flow and limit line of Figure 2.2-1 the limit line of Figure 2.2-1 AXIAL POWEd IMBALANCEa
- 5. RCS pressure-low # 2 1800 psig 2 1800 psig
{" 6. RCS pressure-high 5 2300 psig 5 2300 psig
- 7. RCS pressure-variable- 2 (11.80 Tout F-5209.2) psig s (11.80 Tout F-5209.2) psig l lowa
- 8. Nuclear overpower More than one pump inoperable. More than one pump inoperable.
based on RCPPMsa
- 9. Reactor containment 5 4 psig s 4 psig vessel i I
" Trip may be manual'y bypassed when RCS pressure s 1720 psig by actuating the shutdown bypass, co provided that (1) the nuclear overpower trip setpoint is s 5% of RATED THERMAL POWER, (2) the
! gp shutdown bypass RCS pressure-high trip setpoint of s 1720 psig is imposed, and (3) the shut-
) g down bypass is removed when RCS pressure > 1800 psig.
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Table 8-3. Quadrant Power Tilt Limits Table 3.2-2. Quadrant Power Tilt Limits l- Steady-state Transient Maximum limit limit limit 7 Measurement independent 4.92 11.07 20.0 I
QUADRANT POWER TILT QUADRANT POWER TILT as measured by:
i L Symmetrical incore 3.31 8.81 20.0 detector system
{ Power range channels Minimun incore de-1.96 1.90 6.96 4.40 20.0 20.0 tector system Table 8-4. DNBR Limits Ta bl e 3. 2-1. DNB Margin Four RC pumps Three RC pumps Parameter operating operating
[
Reactor coolant hot leg 5 604.6 5 604.6 temperature, T '
H Reactor coglant pres- 2 2,061.6 2 2,057.2 m
sure, psig Reactor coolant flow 2 139.7 x 106 2 104.4 x 100 rate, lbm/hr d
Applicable to the loop with two RC pumps operating.
b
[ limit not applicable during either a THERMAL POWER ramp in-crease in excess of 5% of RATED THERMAL POWER per minute or a THERMAL POWER step increase of greater than 101, of RATED THERMAL POWER.
[ 8-13 Babcock & Wilcox
I Figure 8-1. Reactor Core Safety Limits 2400 ,
I 2300 RCS PRESSURE-HIGH TRIP j RC OUTLET TEMP I
/ E HIGH TRIP
. 2200 a
l
/
mi 8 0 5 5 2100 ACCEPTABLE OPERATION f
a 2000 -
SAFETY LIMIT
'IV E 1900 H
Y c, UNACCEPTABLE 1800 RCS PRESSURE /
LOW TRIP 580 590 600 610 620 630 640 I
Reactor Outlet Temperature, F I
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l Figure 8-2. Reactor Core Safety Limits I
L
- 120
{ (-31.4,112)
~ ~
( +16. 8.112 )
(-40,107) CEPTABLE 4 F PUMP OPERATION L -
- 100 (+40,100)
(-31.4,84.2)
- 90
(+16.8,84.2)
(-40,79)( -
- 80
{ t-E
- 70 ( +40, 7 2. 3) a.
ACCEPTABLE 3 & 4 ; - - 60 PUMP OPERATION $
%- - 50
[ E a: -- 40 o
- 30 S- - 20 o
- 10
( i , , , i i > i I i
-60 -50, -40 -30 -20 - 10 0 + 10 +20 +30 +40 +50 +60 Reactor Power ImDalance, s l
g_t5 Babcock & Wilcox
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Figure 8-3. Reactor Trip Setpoints Ii I
(-15,107) - - 110
(+5.0,107.0)
- 100
(-30,96)
ACCEPTABLE (28,93.5) 4 PUMP -
- 90 OPERATION
(-15,80.2) (+5.0,80.2) 80
(-30,70) < -
- 70 ACCEP T ABLE > (28,67) 3 & 4 PUMP g OPERATION g- - 60 s1.
- 50 ,
5 g-cr
- 40 l
=
- 30 l
- 20 O
E- - 10 I e ! I 3
I t t f l I 40 -30 -20 -10 0 +10 +20 +30 +40 +50 g
Reactor Power Imualance, %
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Figure 8-4. Pressure / Temperature Limits 2400 1
oo 2200 I .
. L 2 l
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0 2 2000 - -
I s
- n ._ _ - _ _ _
l -
1800 l ,
I 1600 1 580 600 620 640 660 Reac to r Outlet Temp, F I CURVE FLOW
(*, DESIGN)
POWER
($ OF 2568 MWt)
PUMPS OPERATING (TYPE OF LIMIT) 1 106.5 112 4 PUMPS (DNBR) 2 74.7 86.4 3 PUMPS (DNBR)
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I' Figure 8-5. Regulating Rod Group Insertion Limits for Four-Pump Operation From 0 to 250 ! 10 EFPD 110 POWER LEVEL CUT 0FF (905 RATED THERMAL (175.102) (230.102) 10 0 POWER) 90 _
N (175,90) (230,90) 5 80 -
UNACCEPTABLE (160.80) (250.80)
E OPERATION 3 70 E
o 60 -
i
._ 50 ;- (100,50) i a (300.50) 40 e
ACCEPTABLE E
30 b
(60,30) '
( 33, 15) 10
/ (0,0) , , , , ,
0 50 100 150 200 250 300 Rau Index, '4 WI tild rawn 0 25 50 75 100 0 25 50 75 100 t f f f I I I I I I Group 5 Group 7 9 29 59 15 l00 g
Group 6 E
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Figure 8-6. Regulating Rod Group Insertion Limits for Four-Pur:p l Operrition After 250 1 10 EFPD 110 100 _ (261,102)
POWER LEVEL CUT 0FF 90 -
(90', OF RATED THERMAL POWER) r (261,90) 80 -
(240 80) i
~
j UNACCEPTABLE I 3
- 60 -
OPERATION E
I -
o 50 -
(175.50) 40 -
d= ACCEPTABLE E 30 -
OPERATION 20 -
10 -
0 ' '
0 50 100 150 200 250 300 Rao Index, t Witnarawn 0 25 50 75 100 0 25 50 75 100 t f f f f ! '
i f I Group 5 Group 7 0 25 50 75 100 t i ! f I Group 6 8-19 Babcock & Wilcox
I Figure 8-7. Regulating Rod Group Insertion Limit s for Three-Pump Operation From 0 to 250 10 EFPD 110 100 I
l UN ACCEPT ABL E l 90 -
OPERATION 80 -
(250,76 5)
{
(170.76.5) 70 -
E j
2 60 -
l 50 -
I (300.48) 5
- 40 -
a
% (95.37.5) d' I 30 ACCEPTABLE 20 -
OPERATION I
10 _ (33.11.25) 0 0 50 100 150 200 250 300 Rau Iridex, *, Wi tnorann 0 25 50 75 100 0 25 50 75 100 1 ? I t t i I ! I i Group 5 Group 7 0 25 50 75 100 t I f f I Group 6 I
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Figure 8-8. Regulating Rod Group Insertion Limits for Three-Pump Operation After 250 1 10 EFPD 110 s
l H 100
[ 00 -
UN ACCEP T ABL E g OPERATION 80 -
h (240.76.5)
I k
5 70 -
(300.16.5) 3 60 -
1 =
50 -
(175.48) y 40 -
J' ACCEPTABLE 30 -
OPERATION I 20 -
I 10 55.15)
(45,11.25) 0 (0 0) , , , , ,
0 50 100 150 200 250 300 Rau Index, ", Wiinurann I 0 t
25 t
50 t
75 f
100 I
0 t
25 t
50 t l 75 100 g
Group 5 Graup 7 0 25 % 75 100 0 I l t I cup 6 I
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I Figure 8-9. APSR Position Limits for o to 250 + 10 EFPD, Crystal River 3 110 l
100 -
( 7. 5, 10 2) 90 .-
(7.5 90) (36,90) UNACCEPTABLE OPERATION
$ 80 - (0,80) (45,80)
E I'1 70 -
l c
e 3 60 -
E 50 -
- ( 100,50)
ACCEPTABLE g 40 ._
OPERATION 5 30 -
20 -
10 -
0 0 10 20 30 40 50 60 70 80 90 100 5 Rod Posi ti on, fs Wi ttiarawn I
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Fir,ure 8-10.
APSR Posit ion Limits Af ter 250 + 10 EFPD, Crystal River 3 j 110 100 (7.5,102)
~
b 90 -
(40,90)
(7.5,90) UNACCEPTABLE
" OPERATION f g 80 - (0,80) 50,80)
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a.
E E 70 -
5
= 60 ACCEPTABLE a OPERATION E 50 -
, (100,50) 40 h
O 30 -
l 20 -
10 -
0 t i i l
e i , i i i 0 10 20 30 40 50 60 70 80 90 100 Rod Position, % Witnarawn I
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I Figure 8-11. Axial Power Imbalance Envelope for Operation From 0 to 250 t 10 EFPD
-110 I
(-20,102) c p (+10.2,102)
-100
(-20,90) -
- 90 (+10.8.90)
(-25.80) ; -- 80 ( + 12, 80 )
a.
- 70 2as 5 -- 60 UN ACCEPT ABL E ACCEPTABLE g OPERATION OPERATION j -- 50 I
-- 40 i
f -
- 30 I
-- 20 I
- 10
-30
-20
-10 0
+10 +20
+30 I
Axial Po-er imbalance, ',
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p Figure 8-12. Axial Power Imbalance Envelope for L Operation After 250 10 EFPD i
110 .
(+18 3.102)
I (-25.102)c
- 100
(-25.90) -- 90 (+18 0.90)
(-30.80) -
- 80 (+19.6.80)
=
- 70 E
..- 60 UNACCEPTABLE ACCEPTABLE y -- 50 OPERATION E
I OPERATION
~
o
- 40
-- 30 E
- 20
-- 10 t t I i !
-30 -20 -10 0 +10 +20 +30 Axial Power Imtalance, 5 I
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^
- 9. STARTUP PROGRAM - PliYSICS TESTING The planned startup test program associated with core performance is outlined below. These tests verify that core performance is within the assumptions of the safety analysis and provide confirmation for continued safe operation of p the unit.
m 9.1. Precritical Tests 9.1.1. Control Rod Trip Test Precritical control rod drop times are recorded for all control rods at hot full-flow conditions before zero power physics testing begins. Acceptable criteria state that the rod drop time from fully withdrawn to 75% inserted shall be less than 1.66 seconds at the conditions above.
It should be noted that safety analysis calculations are based on a rod drop time of 1.40 seconds from fully withdrawn to two-thirds inserted. Since the most accurate position indication is obtained from the zone reference switch at the 75%-inserted position, this position is used instead of the two-thirds inserted position for data gathering. The acceptance cirterion of 1.40 seconds corrected to a 75%-inserted position (by rod insertion versus time correlation) is 1.66 seconds.
RC Flow
~
9.1.2.
RC Flow with four RC pumps running will be measured at hot zero power, steady-state conditions. Acceptance criteria require that the measured flow be with-in allowable limits.
9.1.3. RC Flow Coastdown
{
The "oastdown of RC flow from the tripping of the RC pump with highest flow from four RC pumps running will be measured at hot zero power conditions. The coastdown of RC flow versus time will then be compared to the required RC flow versus time. Acceptance criteria require that the measured flow rate exceed the minimum.
9-1 Babcock 8. Wilcox
I 9.2. Zero Power Physics Tests 9.2.1. Critical Boron Concentration Criticality is obtained by deboration at a constant dilution rate. Once criti-cality is achieved, equilibrium boron is obtained and the critical boron con-centration determined. The critical boron concentration is calculated by cor-recting for any tod withdrawal required in achieving equilibrium boron. The g
acceptance criterion placed on critical boron concentration is that the actual W boron concentration must be within t100 ppm boron of the predicted value.
9.2.2. Temperature Reactivity Coefficient I The isothermal temperature coefficient is measured at approximately the all-rods-out conf igurat ion and at the hot zero power rod insertion limit. The average coolant temperature is varied by first decreasing then increasing tem-I perature by 5 F. During the change in temperature, reactivity feedback is com- 3 pensated by discrete change in rod motion, the change in reactivity is then calculated by the summation of reactivity (obtained from reactivity calculation i on a strip chart recorder) associated with the temperature change. Acceptance criteria state that the measured value shall not differ from the predicted value by more than 10.4 x 10 ~ '+ (Ak/k)/ F (predicted value obtained from Physics Test Manual curves).
The moderator coefficient of reactivity is calculated in conjut.: tion :rith the temperature coefficlent measurement. After the temperature coefficlent has been measured, a predicted value of fuel Doppler coef ficient of reactivity is added to obtain moderator coef ficient. This value must not be in excess of the g acccotance criteria limit of +0.9 x 10 - '+ (Ak/k)/*F. .
W 9.2.3. Control Rod Group Reactivity Worth l
Control bank group reactivity worths (groups 5, 6, and 7) are measured at hot zero power conditions usir.3 the boron / rod swap method. The boron / rod swap method consists of establi.hing a deboration rate in the reactor coolant sys-tem and compensating for the reactivity changes of this deboration by inserting control rod groups 7, 6, and 5 incremental steps. The reactivity changes that 5 occur during these measurements are calculated based on Reactimeter data, and differential rod worths are obtained from the measured reactivity worth versus the change in rod group position. The differential rod worths of each of the l
I 9-2 Babcock & Wilcox
-,,moe. ,
A controlling groups are then summed to obtain integral rod group worths. The
- acceptance criteria for the control bank group worths are as follows:
~ 1. Individual bank 5, 6, ' worth:
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predicted value - measured value measured value x 100 s 15
- 2. Sum of groups 5, 6, and 7:
predictejl_value - measured value x 100
- measured value s 10 i
9.2.4. Ejected Control Rod Reactivity Worth After CRA groups 7, 6. and 5 have been positioned near the minimum rod inser-tion l im it , the e geted rod is horated to 100% withdrawn and the worth ob-tained by adding the incremental changu in reactivity by boration.
Af ter the ejected rod has been borated to 100% withdrawn and equilibrium boron established, the ejected rod is then swapped in versus the controlling I rod group and the worth determined by the change in the previously calibrated controlling rod group position. The boron swap and rod swap values are aver-aged and error-adjusted to determine ejected rod worth. Acceptance criteria for the ejected rod worth test are as follows:
predicted value - measured value measured value x 100 s 20
- 2. Measured value (error-adjusted) c 1.0% tsk/k The predicted ejected rod worth is given in the Physics Test Manual.
9.3. Power Escalation Tests I 9.
- 1. Core Power Distribution Verification at s40, 7 5, and 100% FP With Nominal Control Rod Position Core power distribution tests are performed at 40, 75, and 100% full power (FP). The test at 40% FP is essentially a check on power distribution in the core to identify any abnormalities before escalating to the 75% FP plateau.
Rod index is established at a nominal full power rod configuration at which I the core power distribution was calculated. APSR position is established to provide a core power imbalance corresponding to the imbalance at which the core power distribution calculations were performed.
9-3 Babcock s.Wilcox
The following acceptance criteria are placed on the 40% FP test:
- 1. The worst-case maximum Iincar heat rate must be less than the LOCA 1.. nit.
- 2. The minimum DNilR must be greater than 1.30. '
- 3. The value obtained from the extrapolation of the minimum DNIlR to the next power plateau overpower trip setpoint must be greater than 1.30 or the extrapolated value of imbalance must fall outside the RPS power / imbalance /
flow trip envelope.
- 4. The value obtained from the extrapolation of the worst-case maximum linear heat rate to the next power plateau overpower trip setpoi~ must be less l
l than the fuel melt limit or the extrapolated value of imbalance must fall ou t s id e the RPS power / imbalance / flow trip envelope.
- 5. The quadrant power t ilt shall not exceed the limi.ts specified in the Tech-nical Specifications.
- 6. The highest measured and predicted radial peaks shall be within the follow-i ing limits:
I predicted value - measured value measured value x 100 s8
- 7. The highest measured and predicted total peaks shall be within the follow-Ing limits:
predicted value - myisured value x 100 measured vtlue c 12 Items 1, 2, 5, 6, and 7 above are established to verify core nuclear and ther-mal calculational models, thereby verifying the acceptability of data from =
these models for input to safety evaluations.
Items 3 and 4 establish the criteria whereby escalation to the next power pla- I teau may be accomplished without exceeding the safety limits specified by the j safety analysis with regard to DNBR and linear heat rate. W l
The power distribution tests performed at 75 and 100% FP are identical to the 40% FP test except that core equilibrium xenon is established prior to the 75
- and 100% FP tests. Accordingly, the 75 and 100% FP measured peak acceptance criteria are as follows:
l 9-4 Babcock & Wilcox g
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- 1. The highest measured and predicted radial peaks shall be within the follow-l L ing limits:
predicted value - measured valu measured value x 100 s5
- 2. The highest measured and predicted total peaks shall be within the follow-I g Ing limits:
predicted value - measured value x 100 s 7.5 measured value 9.3.2. Incore Vs Excore Detector Imbalance j Correlation Verification at %40% FP Imbalances are set up in the core by control rod positioning. Imbalances are read simultaneously on the incore detectors and excore power range detectors for various imbalances. The excore detector offset versus incore detector off-set slope must be at least 1.15. If the excore detector offset versus incore detector offset slope criterion is not met, gain amplifiers on the excore de-tector signal processing equipment are adjusted to provide th. required gain.
9.3.3. Temperature React ivity Coef fic ien t a t N100% FP The average reactor coolant temperature is decreased and then increased by about 5 F at constant reactor power.
I The reactivity associated with each tem-perature change is obtained from the change in the controlling rod group posi-tion. Controlling rod group worth is measured by the fast insert / withdraw method. The temperature react ivity coef ficient is calculated from the mea-sured changes in reactivity and temperature.
Acceptance criteria state that the moderator temperature coefficient shall be negative.
9.3.4. Power Doppler Reactivity Coefficient at %100% FP I Reactor power is decreased and then increased by about 5% FP.
change is obtained from the change in controlling rod group position. Control The reactivity rod group worth is measured using the fast insert / withdraw method. Reactivity corrections are made for changes in xenon and reactor coolant temperature that occur during the measurement. The power doppler reactivity coef ficient is calculated f rom the measured reactivity change, adjusted as stated above, and the measured power change.
9-5 Babcock & Wilcox l
i i I The predicted value of the power doppler reactivity coefficient is given in the Fhysics Test Manual. Acceptance criteria state that the measured value shall be more negative than -0.55 x 10 (Ak/k)/% FP.
9.4. Procedure for Failure to Meet Acceptance Criteria i
Florida Power reviews the results of all startup tests to ensure that all ac-ceptance criteria are met. If the review of the test indicates that the re-sults are well within the acceptance criteria, no further evaluation is con-ducted. If the review indicates that the results are approaching or close to the acceptance criteria limits, further evaluation of that particular test or other supporting tests is performed to look for trends. This evaluation will determine whether additional support data are required to discover any abnor-l ma l cond i t ions. If acceptance criteria for any test are not met, an evalua-tion is performed before the test program is continued. This evaluation is t
performed by site test personnel with participation by Babcock & Wilcox tech-nical personnel as required. Further specific actions depend on evaluation I results. These actions can include repeating the tests with more detailed
- attention to test prerequisites, added tests to search for anomalies, or de-sign personnel performing detailed analyses of potential safety problems be-cause of parameter deviation. Power is not escalated until evaluation shows that plant safety will not be compromised by such escalation.
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1 REFERENCES 1
1 Crystal River Unit 3, Final Safety Analysis Report, Docket 3-302, Florida l Power Corporation.
2 Crystal River Unit 3, Cycle 2 Reload i:eport, BAW-]S21 Babcock & Wileox, I Lynchburg, Virginia, February 1979.
3 Crystal River Unit 3, Technical Specit ication Change Request No. 27, Dot.k-et 50-302, License DPR-72, November 20, 1978.
4 BPRA Retainer Design Report, BAW-1496. Babcock & Wilcox, Lynchburg, Vir-ginia, May 1978.
I ,
A. F. J. Ec ke r t , H. W. Wilson, and K. E. Yoon, Program to Determine Per-formance of B&W Fuels - Cladding Creep Collapse, BAW-10084P-A, Rev. 2, I Babcock & Wilcox, Lynchburg, Virginia, January 1979.
I TACO - Fuel Pin Performance Analysis, BAW-10087P-A, Rev 2, Babcock & Wilcox.
t' Ly nc hburg, Virginia, August 1977.
7 Crystal River Unit 3, Fuel Densification Report, BAji- 13 9'/ , Babcock & Wilcoy.
Lynchburg, Virginia, August 1973.
8 Babcock & Wilcox Version of PDQ User's Manual, BAW-10117P-A, Babcock &
Wilcox, l.ynchburg, Virginia, January 1977.
'3 Correlat ion of Critical Heat Flux in Bundle Cooled by Pressurized Water, 1
BAW-10000A, Babcock & Wilcox, Lynchburg, Virginia, May 1976.
10 L. S. Rubenstein (NRC) to J. H. Taylor (B&W), Letter, " Evaluation of In-l terim Procedure for Calculating DNBR Reductions due to Rod Bow," October 18, 1979.
11 l Crystal River Unit 3, Licensing Considerations for Continued Cycle 1 Oper-ation Without Burnable Poison Rod Assenblies and Orifice Rod Assemblies, BAW-1490, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, July 1978.
A-1 Babcock & Wilcox
f 12 S. A. Varga (NRC) to J. H. Taylor (B&W), Letter, " Update of BAW-10055, Fuel Densification Report," December 5, 1977.
13
- R. C. Jones, J. R. Biller, and B. M. Dunn, ECCS Analysis of B&W's 177-FA Lowered-Loop NSS, BAW-10103A, Rev. 3, Babcock & Wilcox, Lynchburg, Virginia, July 1977.
14 J. H. Taylor (B&W) to R. L. Baer (NRC), Letter, "LOCA Analysis for B&W's
- 177-FA Plants With Lowered-Loop Arrangement (Category 1 Plants) Utilizing l a Revised System Pressure Distribution," July 8, 1977.
15 J. H. Taylor (B&W) to S. A. Varga (NRC), Letter, "ECCS Small Break Analy-
, sis," July 18, 1978.
16 W. P. Stewart (FPC) to R. W. Reid (NRC), Letter, " Crystal River Unit 3, i
Docket No. 50-302, Operating License No. DPR-72, ECCS Small Break Analy-sis," January .2, 1979.
17 CRAFT 2 - FORTRAN Program for Digital Simulation of a Multinode Reactor Plant During Loss of Coolant, LAW-10092, Babcock & Wilcox, Lynchburg, Virginia, April 1975.
18 C. D. Morgan and H. S. Kao, TAFY - Fuel Pin Temperature and Gas Pressure j Analysis, BAW-10044, Babcock & Wilcox, Lynchburg, Virginia, May 1972.
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