ML19284A448

Cycle 2 Reload Rept.

ML19284A448
Person / Time
Site:	Crystal River
Issue date:	02/28/1979
From:	FLORIDA POWER CORP.
To:
Shared Package
ML19284A445	List: ML19284A444 ML19284A446 ML19284A448
References
BAW-1521, NUDOCS 7903060434
Download: ML19284A448 (79)

Text

.

BAW-1521 February 1979 CRYSTAL RIVER UNIT 3

- Cycle 2 Reload Report -

J 790306043'[

BABCOCK & WILCOX Power Generation Group Nuclear Power Generation Division P. O. Box 1260 Lynchburg, Virginia 24505 Babcock & Wilcox

CONTENTS Page

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. Rtactor Coolant System Stress . . . . . . ..... 3-1 3.1.2. Reactor Coolant Purp Power Monitors . . . ..... 3-1 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 l 4.2.2. Cladding Stress . . . . . . . . . . . . . ..... 4-2 4.2.3. Cladding Strain . . . . . . . . . . . . . ..... 4-2 4.3. Fuel Thermal Design . . . . . . . . . . . . . . . ..... 4-2 4.4. Operating Experience . . . . . . . . . . . . . . . ..... 4-2

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 6.3. Flux / Flow Trip Setpoint Analysis . . . . . . . . . ..... 6-3 6.4. Loss-of-Coolant Flow Transients . . . . . . . . . ..... 6-3

7. ACCIDENT AND TRANSIENT ANALYSIS . . . . . . . . . . . . ..... 7-1 7.1. General Safety Analysis . . . . . . . . . . . . . ..... 7-1 7.2. Accident Evaluation . . . . . . . . . . . . . . . ..... 7-2 7.3. Rod Withdrawal Accidents . . . . . . . . . . . . . ..... 7-2 7.4. Moderator Dilution Accident . . . . . . . . . . . ..... 7-3 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  ;

I Babcock & Wilcox I l

l 1

CONTENTS (Cont'd)

Page 7.8. Less of Electric Power . . . . . . . . . . . . . . ..... 7-6 7.9. Steam Line Failure . . . . . . . . . . . . . . . . ..... 7-7 1 . 1 01 Steam Generator Tube Failure . . . . . . . . . . . ..... 7-7 7.11. Fuel Handling Accident . . . . . . . . . . . . . . ..... 7-8 7.12. Rcd Ejection Accident . . . . . . . . . . . . . . ..... 7-8 7.13. Maximum Hypothetical Accident . . . . . . . . . . ..... 7-9 7.14. Waste Gas Tank Rupture . . . . . . . . . . . . . . ..... 7-9 7.15. LOCA Analysis . . . . . . . . . . . . . . . . . . ..... 7-9 7.16. Failure of Small Lines Carrying Primary Coolant Outside Containment . . . . . . . . . . . . . . . . . . . ..... 7-10 7.16.1. Identification of Causes . . . . . . . . . ..... 7-10 7.16.2. Analysis of Effects and Consequences . . . ..... 7-10

8. PROPOSED MODIFICATIONS TO TECHNICAL SPECIFICATIONS . . . ..... 8-1

9. STARTUP PROGRAM -- PHYSICS TESTING . . . . . . . . . . . ..... 9-1 9.1. Precritical Tests - Control Rod Trip Test . . . . ..... 9-1 9.2. Zero Power Physics Teste . . . . . . . . . . . . . ..... 9-1 9.2.1. Critical Boron Concentration . . . . . . . ..... 9-1 9.2.2. Temperature Reactivity Coefficient . . . . ..... 9-1 9.2.3. Control Rod Group Reactivity Worth . . . . ..... 9-2 9.2.4. Ejected Control Rod Reactivity Worth . . . ..... 9-2 9.3. Power Escalation Tests . . . . . . . . . . . . . . ..... 9-3 9.3.1. Core Power Distribution Verification at %40, 75, and 100% FP With Nominal Control Rod Position ... 9-3 9.3.2. Incore Vs Excore Detector Imbalance Correlation

'erification at %40% FP . . . . . . . . . ..... 9-4 9.3.3. Temperature Reactivity Coefficient at %100% FP . .. 9-5 9.3.4. Power Doppler Reactivity Coefficient at N100% FP . . 9-5 9.4. Procedure for Failure to Meet Acceptance Criteria ..... 9-5 REFERENCES . . . . . . . . . . . . . . . . . . . . . . . ..... A-1 List of Tables Table 4-1. Fuel Design Parameters and Dimensions . . . . . . . . ..... 4-3 4-2. Fuel Thermal Analysis Parameters . . . . . . . . . . . ..... 4-4 5-1. Physics Parameters, Crystal River Three, Cycle 2 . . . ..... 5-3 5-2. Shutdown Margin Calculation for Crystal River 3, Cycle 2 . ... 5-5 6-1. Cycle 1 and 2 Thermal-Hydraulic Design Conditions . . .....

7-1. 6-4 Comparison of Key Parameters for Accident Analysis . . ..... 7-13 7-2. Bounding Values for Allowable LOCA Peak Linear Heat Rates ...

7-3. 7-13 Input Parameters to LOCF Transients . . . . . . . . . ..... 7-14

! 7-4. Summary of Minimum DNBR Results for Limiting LOCF Transients .. 7-14

- lii - Babcock 8.Wilcox

u Tables (Cont'd)

Table Page 7-5. Analysis Assumptions for MU&PS Letdown Line Rupture Accident . . 7-15 7-6. Activity Released to Environment Due to Rupture of MUSPS Letdown Line . . . . . . . . . . . . .. . .. . . . . ..... 7-16 7-7. Radiological Consequences of MUSPS Letdown Line Rupture Outside Containment . . . . . . . . . . . . . . .. . . . . . ..... 7-16 8-1. Technical Specification Changes . . .. . .. . . . . ..... 8-2 List of Figures Figure 3-1. Core Loading Diagram for Crystal River 3, Cycle 2 . . ..... 3-4 3-2. Enrichment and Burnup Distribution for Crystal River 3 Cycle 2 . . . . . . . . . . . . . . .. . . . . . . . ..... 3-5 5-1. BOC (4 EFPD), Cycle 2 Two-Dimensional Relative Power Distribution -- HFP, Equilibrium Xenon, Banks 7 and 8 Inserted . 5-6 7-1. Four-Pump Coastdown Power and Flow Transients, Crystal River 3 . . . . . . . . . . . . . . .. . .. . . . . ..... 7-17 7-2. Four-Pump Coastdown -- Hot Channel MDNBR Vs Time, Crystal River 3 . . . . . . . . . . . . . . .. . .. . . . . ..... 7-18 7-3. Locked-Rotor Power and Flow Transients, Crystal River 3 . . . . 7-19 7-4. Locked-Rotor, Crystal River 3 . . . .. . . .. . . . ..... 7-20 8-1. Reactor Core Safety Limits . . . . .. . .. . . . . ..... 8-11 8-2. Reactor Core Safety Limits . . . .. . . . . . . . . ..... 8-12 8-3. Trip Setpoints . . . . . . . . . . .. . .. . . . . ..... 8-13 8-4. Pressure-Temperature Limits . . . . ... ... . . . ..... 8-14 8-5. Regulating Rod Group Insertion Limits for Four-Pump Operation From 0 to 225 10 EFPD . . . . . . . .. ... . .. ..... 8-15 8-6. Regulating Rod Group Insertion Limits for Four-Pump Operation After 225 10 EFPD . . . . . . . . .. . .. . . . . ..... 8-16 8-7. Regulating Rod Group Insertion Limits for Three-Pump Operation From 0 to 225 i 10 EFPD . . . . . . ... .. . . . . ..... 8-17 8-8. Regulating Rod Group Insertion Limits for Three-Pump Operation After 225 10 EFPD . . . . . . . . ... .. . . . . ..... 8-18 8-9. Centrol Rod Locations . . . . . . . .. . ...... ..... 8-19 8-10. APSR Position Limits for 0 to 225 10 EFPD, Crystal River 3 . 8-20 8-11. APSR Position Limits After 225 10 EFPD, Crystal River 3 . . . 8-21 8-12. Axial Power Imbalance Envelope for Operation From 0 to 225 10 EFPD . . . . . . . . . . . . . . .. .. . . . . ..... 8-22 8-13. Axial Power Imbalance Envelope for Operation After 225 10 EFPD . . . . . . . . . . . . . . . . .... . . .. ..... 8-23 L

- iv - Babcock & \Vilcox

l. INTRODUCTION AND

SUMMARY

This report justifies tne operation of Crystal River Unit 3 (cycle 2) at a rated core power of 2544 MWt. Included are the required analyses to support cycle 2 operation; these analyses employ analytical techniques and design bases established in reports that have received technical approval by the USNRC (see references).

The design for cycle 2 raises the rated thermal power from 2452 to 2544 MWt; the latter corresponds to the ultimate core power level identified in the Crystal River Unit 3 FSAR.I Each accident analyzed in the FSAR has been re-viewed, and each review is summarized in this report. Some accidents were re-analyzed to include the reactor coolant pump power monitors 2, 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 Specifications have been reviewed, and the modifications re-quired for cycle 2 are justified in this report.

Based on the analyses performed, which take into account the postulated effects of fuel densification and the Final Acceptance Criteria for emergency core cooling systems (ECCS), it has been concluded that Crystal River 3, cycle 2, can be safely operated at the rated core power level of 2544 MWt.

1-1 Babcock & Wilcox

2. OPERATING HISTORY The Crystal River Unit 3 nuclear generating plant achieved initial criticality on January 14, 1977, and power escalation commenced on January 29, 1977. The 100% power level of 2452 MWt was reached on April 1, 1977. A control rod interchange was performed at 250 effective full-power days (EFPD).

At 268.8 EFPD, all burnable poison rod assemblies (BPRAs) and orifice rod as-semblies (ORAs) were removed as the result of mechanical difficulties experi-cnced with the BPRA/ ORA latching mechanisms. Two modified ORAs with retainers were located in the two fuel assemblies containing the primary neutron sources.

In addition, four fuel assemblies were replaced and four were relocated because of damage incurred during refueling operations.

Cycle 1 regained criticality on September 15, 1978, and regained 100% power operation on October 4, 1978. Cycle 1 is scheduled for completion in April 1979 after 450 EFPD. No operating anomalies occurred during cycle 1 that would adversely affect the fuel performance during cycle 2.

Cycle 2 is scheduled to start operation la June 1979 with an upgraded rated power level of 2544 MWt. The design cycle length is 275 EFPD.

2-1 Babcock & VVilcox

3. GENERAL DESCRIPTION 3.1. Plant Description 3.1.1. Reactor Coolant System Stress In support of the power upgrade, reactor coolant system (RCS) stresses were reviewed. Since the CR-3 RCS functional specification did not analyze power levels up to 2544 MWt, a new document was issued. This revised document was reviewed by the applicable engineering groups, and it was determined that no hardware changes were required; however, a revision was issued to the RCS Stress Report.

3.1.2. Reactor Coolant Pump Power Monitors In support of the power upgrade, reactor coolant pump power monitors (RCPPM) are being added to CR-3 during the EOC-1 refueling outage.2 The RCPPM anticipates a loss or reduction of the reactor coolant flow by moni-toring RC pump power and detecting abnormal power conditions indicative of an inoperable 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 less 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 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 3-1 Babcock & Wilcox

preclude the minimum DNBR from going below 1.30 for the four-pump coastdown transient and below 1.00 for the locked-rotor transient.

Hovtver, 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 preclude the minimum DNBR from going below the acceptance criteria.

Therefore, for power levels above 2500 MWt, nuclear overpower based on RCPPMs must 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 miniaum DNBR limits.

Each RCPPM string includes two current transformers and two potenti 11 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 bistable, which provides a contact output for selected overpower and under-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 monitoring 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 inside enclosures separated by barriers. Contact outputs from the RCFPM cabinets to the four RPS channels are arranged to provide ade-quate physical separation and electrical isolation between each channel. Ex-ternal 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 scismically qualified and located in a Class I structure. All supports for engineered safeguards cable trays and conduits are designed for OBE and SSE using the acceleration floor re-sponse spectra developed for applicable levels of the containment building, auxiliary building, intermediate building, and control complex.

3-2 Babcock & Wilcox

The current and potential transformers are not seismically qualified. However, separation of the cables carrying redundant transformer outputs to the RCPPM cabinets is provided in accordance with the separation criteria stated above.

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 system. Upon the receipt of two such signals, whatever the 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.1 The cycle 2 core consists of 177 fuel assem-blics (FAs), each of which is a 15-by-15 array containing 208 fuel rods, 16 control rod guide tubes, and one incore instrument guide tube. The fuel as-semblies in batches 2 and 3 have an average nominal fuel loading of 463.6 kg of uranium, whereas the batch 4 assemblies maintain an average nominal fuel loading 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).

Figure 3-1 is the core loading diagram for Crystal River 3. The initial en-richments of batches 2 and 3 were 2.54 and 2.83 wt % uranium-235, respectively.

All 56 batch 1 assemblies will be discharged at the end of cycle 1. The batch 4 design enrichment is 2.64 wt % uranium-235. Batch 2 and 3 assemblies will be shuffled to new locations. The batch 4 assemblies will occupy the periphery of the core. Figure 3-2 is an eighth-core map showing the burnup of each as-sembly at the beginning of cycle 2 and its initial enrichment.

Core reactivity will be controlled by 61 full length Ag-In-Cd control rod as-semblies (CRAs) and soluble boron shim. In addition to the full-length CRAs, eight axial power shaping rods (APSRs) are provided for additional control of the axial power distribution. The cycle 2 locations of the 69 control rods and the group designations are indicated in Figure 8-9. No control rod inter-changes or burnable poison rods are required for cycle 2. The nominal system pressure is 2200 psia, and the core average densified nominal linear heat rate (LHR) is 5.71 kW/ft at the rated core power of 2544 MWt.

3-3 Babcock & Wilcox

Figure 3-1. Core Loading Diagram for Crystal River 3, Cycle 2 A 17 Cycle 1 location 4 4 4 4 4 Y Batch Number 5 B7 W5 31 4 4 4 2 2 2 4 4 4 F7 C6 M2 C8 Mle C10 F9 C

4 4 2 2 3 2 3 2 2 4 4 H5 L2 M2 N3 R8 N13 N14 L14 E8 4 4 2 3 3 3 3 3 3 3 2 4 4 c6 510 013 E6 L1 D9 L15 E10 03 B6 C10 E 4 2 3 3 2 3 2 3 2 3 3 2 4 F3 812 F3 N11 K1 F8 K15 M4 Fil 84 F13 4 4 2 3 2 2 3 3 3 2 2 3 2 4 4 G2 Bil C12 A10 A9 D7 G8 C12 A7 A6 C4 BS C14 4 2 3 3 3 3 2 2 2 3 3 3 3 2 4 F12 H) H15 C4 H14 H7 H9 HS H2 K12 El H13 E4 4 2 2 3 2 3 2 2 2 3 2 3 2 2 4 K2 P11 012 RIO R9 K4 K8 N9 R7 R6 04 P5 K14 4 2 3 3 3 3 2 2 2 3 3 3 3 2 4 L3 P12 L5 E12 G1 B8 C15 05 Lil P4 L13 L

4 4 2 3 2 2 3 3 3 2 2 3 2 4 4 E6 P10 C13 M6 F1 N7 FIS M20 C3 F6 K10 4 2 3 3 2 3 2 3 2 3 3 2 4 MS F2 D2 D3 A8 D13 D14 F14 H11 4 4 2 3 3 3 3 3 3 3 2 4 4 L7 06 E2 08 E14 010 0 L9 4 4 2 2 3 2 3 2 2 4 4 P7 D11 P9 4 4 4 2 2 2 4 4 4 R

4 4 4 4 4 1 2 3 4 l

5 6 7 8 9 10 11 12 13 14 15 3-4 Babcock & Wilcox

Figure 3-2. Enrichment and Burnup Distribution for Crystal River 3, Cycle 2 8 9 10 11 12 13 14 15 2.54 2.54 2.83 2.54 2.83 2.54 2.54 2.64 H

12,361 16,481 14,632 16,450 10,718 16,191 14,617 0 2.54 2.83 2.83 2.83 2.83 2.54 2.64 K

16,449 10,125 7,565 11,734 10,774 13,714 0 2.54 2.54 2.83 2.54 2.64 2.64 b

14,614 16,592 7,535 14,503 0 0 M 8,190 11,662 1 , 59 2.54 2.64 2.64 17,494 0 0 2.64 0

0 P

i R

XXX Initial Enrichnent XXXXX BOC Burnup, mwd /atU 3-5 Babcock & Wilcox

4. FUEL SYSTEM DESIGN 4.1. Fuel Assembly Mechanical Design The types of fuel assemblies and pertinent fuel design parameters for Crystal River Unit 3, cycle 2 are listed in Table 4-1. New retainer assemblies will be used on four fuel assemblies, two containing primary neutron sources (PNS) and two containing regenerative neutron sources (RNS). The justification for the design and use of the retainers described in reference 5 is applicable to PNS and RNS retainers in the CR-3, cycle 2 fuel. All fuel assemblies are identical in concept and are mechanically interchangeable.

The new fuel assemblies have modified end fittings, primarily to reduce fuel assembly pressure drop and to increase holddown margin. The reload fuel assem-blies incorporate minor design modifications to the spacer grid corner cells which reduce spacer grid lateraction during handling. In addition, improved test methods (dynamic impact testing), show the spacer grids to have a higher seismic capability and therefore, an increased safety margin over the values reported in B&W topical report BAW-10035.21 All other results presented in the FSAR fuel assembly mechanical discussion are applicable to the reload fuel assemblies.

4.2. Fuel Rod Design The mechanical evaluation of the fuel rad is discussed below.

4.2.1. Cladding Collapse The batch 2 and batch 3 fuel are more limiting than batch 4 due to their longer previous incore exposure time. The batch 2 and 3 assembly power histories were analyzed and the most limiting assembly was determined.

The power history for the most limiting assembly was used to perform the creep collapse analysis using the CROV computer code and procedures described in topical report BAW-10084PA, Rev. 2.6 The actual power history for the first 2000 EFPH of cycle 1 was used, including the initial operation at 40 and 75%

4-1 Babcock & Wilcox

core power. The collapse time for the most limiting assembly was conservative-ly determined to be more than 30,000 EFPH which is greater than the maximum projected residence time of cycle 2 fuel (Table 4-1) .

4.2.2. Cladding Stress The batch 2 and 3 reinserted fuel assemblies are the limiting batches from a cladding stress point of view due to their lower density and longer previous exposure time. Batches 2 and 3 have been analyzed and documented in the Crys-tal River Unit 3 Fuel Densification Report.7 4.2.3. Cladding Strain The fuel design criteria specify a limit of 1.0% on cladding plastic circum-ferential strain. The pellet design is established for cladding plastic strain of less than 1% 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 The design minimum linear heat rate (LER) capability and the average fuel tem-perature for each batch in cycle 2 are shown in Table 4-2. LHR capabilities are based on centerline fuel melt and were established using the TAFY-3 code 8 with fuel densification to 96.5% of theoretical density. The batch 4 fuel has a higher initial density and a correspondingly higher linear rate capability (20.15 Vs 19.7 kW/ft) than the batch 2 and 3 fuel. Pellet resinter test data from the batch 4 will be evaluated to demonstrate that the fuel exceeds the design minimum LHR capability.

The power spike model used in cycle 2 analysis is identical to that presented in BAW-100559 except for modifications to F and F as described in reference 10 and accepted in reference 11. These same changes in the densification power spike model have been approved for Oconee 1, 2, 3, ANO-1. and TMI-l re-loads.

4.4. Operating Experience Babcock & Wilcox operating experience with the Mark-B, 15x15 fuel assembly has verified the adequacy of its design. As of November 30, 1978, the following 4-2 Babcock & Wilcox

experience has been accumulated for the nine operating B&W 177-fuel assembly plants using the Mark-B fuel assembly:

Maximum assembly Cumulative net

"#" P'

• Current -

electrical Reactor cycle Incore Discharged output, MWh Oconee 1 5 31,900 31,100 23,108,841 Oconee 2 3 19,800 33,700 19,165,343 Oconee 3 4 20,900 29,400 20,224,866 TMI-l 4 29,430 32,200 21,856,837 TMI-2 1 688 --

156,570 ANO-1 3 30,835 28,300 18,145,982 Rancho Seco 2 29,378 26,670 13,910,219 Crystal River 3 1 12,798 --

5,533,734 Davis-Besse 1 1 6,220 --

2,444,570 Table 4-1. Fuel Design Parameters and Dimensions Batch 2 Batch 3 Batch 4 Fuel assembly type Mark B3 Mark B3 Mark B4 No. of assemblies 61 60 56 Fuel rod OD, in. 0.430 0.430 0.430 Fuel rod ID, in. 0.377 0.377 0.377 Flexible spacers, type Corr'd Corr'd Spring Rigid spacers, type Ceramic Ceramic Zr-4 Undensified active fuel length, in. 144 144 143.6 Fuel pellet OD (mean specified), in. 0.370 0.370 0.370 Fuel pellet initial density (mean speci-fied), % TD 92.5 92.5 94.0 Initial fuel enrichment, wt % 235 U 2.54 2.83 2.64 Estimated residence time, EFPH 17,400 24,120 20,760 Cladding collapse time, EFPH >30,000 >30,000 >30,000 4-3 Babcock & Wilcox

Table 4-2. Fuel Thermal Analysis Parameters Densified fuel parameters (" Batch 2 Batch 3 Batch 4 Pellet diameter, in. 0.3641 0.3641 0.3648 Fuel stack height, in. 141.1 141.1 141.8 Nominal LHR at 2568 MWt, kW/ft 5.77 5.77 5.74 Avg fuel temp at nominal LHR, F 1330 1330 1280 LHR capability (CL fuel melt),

kW/ft 19.7 19.7 20.1 (a)Densification assumed to 96.5% TD. Densified parameters for batch-es 2 and 3 are from the Densification Report7 and are based on as-built data. Densified parameters for batch 4 fuel are based on fuel specification values.

4-4 Babcock & Wilcox

5. NUCLEAR DESIGN 5.1. Physics Characteristics Table 5-1 compares the core physics parameters of cycles 1 and 2; these values were generated using PDQ07 14 for both cycles. Since the core has not yet reached an equilibrium cycle, differences in core physics parameters are ex-pected between cycles. The shorter cycle 2 will produce a smaller cycle dif-ferential burnup than cycle 1. The accumulated average core burnup will be higher in cycle 2 than in cycle 1 because of the presence of the once-burned batch 2 and 3 fuel. Figure 5-1 illustrates a representative relative power distribution for the beginning of the second cycle at full power with equilib-rium xenon and normal rod positions.

The critical boron concentrations for cycle 2 are given in Table 5-1. Control rod worths are sufficient to attain the required shutdown margin as indicated in Table 5-2. However, due to changes in isotopics and the radial flux dis-tribution, the hot, full-power control rod worths are somewhat less than those for cycle 1. The ejected rod worths for cycle 2 are lower than those in cycle 1 for the same number of regulating banks inserted. Values between cycles or between rod patterns are difficult to compare since neither the rod patterns from which the CRA is assumed to be ejected nor the isotopic distributions are identical. 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 insertion limits presented in section 8. The maximum stuck rod worths for cycle 2 are less than those for cycle 1. The adequacy of the shutdown margin with cycle 2 stuck rod worths is demonstrated in Table 5-2. The fol-lowing conservatisms were applied for the shutdown calculations:

1. Poison material depletion allowance.

2. 10% uncertainty on net rod worth.

3. Flux redistribution penalty.

5-1 Babcock & Wilcox

Flux redistribution was accounted for since the shutdown analysis was calcu-lated using a two-dimensional model. The shutdcen c2? c >h .lon at the end of cycle 2 was analyzed at 225 EFPD. This is the latest time ( 10 days) in core life at which the transient tank is nearly fully inserted. After 225 EFPD, the transient bank will be almost fully withdrawn, thus, the available shut-down margin will be increased.

The cycle 2 power deficits from hot zero power to hot full power are similar to but slightly more negative than those for cycle 1. The Doppler coefficients and xenon worths are similar for the two cycles. The differential boron worths at the start of cycle 2 are lower than those for cycle 1 because of depletion of the fuel and the associated buildup of fission products. The effective de-layed neutron fractions for both cycles show a decrease with burnup.

5.2. Changes in Nuclear Design The only major design change between the modified cycle 1 design and 3

the re-load cycle 2 design is the upgrading of the core power to 2544 MWt. The same calculational methods and design information were used to obtain the important nuclear design parameters. The operation without BPRAs/ORAs will be continued from cycle 1. No significant operational or procedural changes exist with regard to axial or radial power shape control, xenon control, or tilt control.

The operational limits and RPS limits (Technical Specification changes) for cycle 2 are presented in section 8.

5-2 Babcock & Wilcox

Table 5-1. Physics Parameters, Crystal River Three, Cycle 2( )

Cycle 1 Cycle 2 Design cycle length, EFPD 430 275 Design cycle burnup, mwd /mtU 12,849 8,500 Design average core burnup - EOC, mwd /mtU 12,849 17,364 Design initial core loading, mtU 82.1 82.3 Critical boron - BOC, ppm (no Xe)

HZP(b), group 8 (37.5% wd) 1415 1260 HZP, groups 7 and 8 inserted 1305 1185 HFP(b), groups 7 and 8 inserted 1210 991 Critical boron - EOC, ppm (eq Xe) group 8 (37.5% wd)

F 5 Control rod worths - HFP, BOC, %Ak/k Group 6 1.43 1.02 Group 7 1.13 0.85 Group 8 (37.5% wd) 0.42 0.49 Control rod worths - HFP, EOC, %Ak/k Group 7 g) 1.23 1.11 Group 8 (37.5% wd) 0.46 0.48(c)

Max ejected rod worth - HZP, %Ak/k P9C (N-12) 0.85 d)

EvC (N-12) 0.85 0.

0.5055 ((d) (c)

Max stuck rod worth - HZP, %Ak/k BOC (N-12) 4.00 EOC (L-14) 2.83 1.82(

1.88 )

Power deficit, HZP to HFP, %Ak/k BOC -0.86 -1.30 EOC -1.90 -2.06 Doppler coeff - BOC, 10-5 (Ak/k/ F) 100% power (0 Xe) -1.35 -1.50 Doppler coeff - EOC, 10-5 (Ak/k/ F) 100% power (eq Xe) -1.53 -1.58 Moderator coeff - HFP, 10-4 (Ak/k/ F)

BOC (0 Xe, Crit. ppm, group 8 ins) -0.11 -0.65 EOC (eq Xe,17 ppm, group 8 ins) -2.50 -2.52 Boron worth - HFP, ppm /%Ak/k BOC 100 106 EOC 103 94 Xenon worth - HFP, %Ak/k BOC (4 EFPD) 2.62 2.67 EOC (equil.) 2.70 2.74 5-3 Babcock & Wilcox

Table 5-1. (Cont'd)

Cycle 1 Cycle 2 Effective delayed neutron fraction - HFP BOC 0.00690 0.005843 EOC 0.00515 0.005164 The cycle 1 values given are for the original cycle 1 design.1 The modified cycle 13had a design cycle length of 510 EFPD, however, it is now planned to refuel after 450 EFPD. The 450 EFPD cycle 1 length has been input to the cycle 2 calculations as presented in this report.

HZP denotes hot zero power (532F T ); HFP denotes hot full power (579F Tavg}*

(c) Rod worths calculations for EOC-2 are done at 225 EFPD, the latest time in core life in which the transient bank is nearly full-in.

(d) Ejected rod worth for groups 5 through 8 inserted.

5-4 Babcock & Wilcox

Table %? . Shurdwn Margin Calculation for Crystal River 3, Cycle 2 BOC, % Ak/k EOC " , % Ak/k Available Rod Worth Total rod worth, HZP( ) 8.66 8.89 Worth reduction due to burnup of poison material -0.10 -0.18 Maximum stuck rod worth, HZP -1.82 -1.88 Net worth 6.74 6.83 Less 10% uncertainty -0.67 -0.68 Total available worth 6.07 6.15 Required Rod Worth Power deficit, HFP to HZP 1.30 2.06 Max allowable inserted rod worth 0.96 1.25 Flux redistribution 0.47 0.74 Total required worth 2.73 4.05 Shutdown Margin Total available minus total required 3.34 2.10 Note: Required shutdown margin is 1.00% Ak/k.

(" For shutdown margin calculations, this is defined as %225 EFPD, the latest time in core life in which the transient bank is nearly full-in.

(b)HZP denotes hot zero power; HFP denotes hot full power.

5-5 Babcock & Wilcox

Figure 5-1. B0C (4 EFPD), Cycle 2 Two-Dimensional Relative Power Distribution -- HFP, Equilibrium Xenon, Banks 7 and 8 Inserted 8 9 10 11 12 13 14 15 7

H 1.00 0.93 1.11 1.12 1.23 0.90 0.47 0.59 K 0.94 1.09 1.27 1.23 1.09 0.83 0.67 7 8 L 0.63 1.01 1.14 1.06 1.15 0.63 M 1.25 1.24 1.03 1.00 N 1.11 1.25 0.74 0 0.87 P

R Ng Inserted rod group No.

x.xx Relative pc;.er density 5-6 Babcock & kVilcox

6. THERMAL-HYDRAULIC DESIGN 6.1. DNBR Evaluations Crystal River 3 will be upgraded in power for cycle 2 operation from 2452 (cy-cle 1) to 2544 MWt rated core power. Thermal-hydraulic design calculations in support of cycle 2 operation assumed a rated power level of 2568 MWt for con-sistency with other B&W reactors and used the analytical methods documented in the Final Safety Analysis Report and updated in the Fuel Densification Report 7 l

The following changes in thermal-hydraulic conditions or assumptions were made for cycle 2 evaluations.

1. The B&W-2 CHF correlation 13 was used for cycle 2 instead of the W-3 cor-relation. The B&W-2 correlation, a realistic prediction of the burnout phenomenon, has been reviewed and approved for use with the Mark-B fuel assembly design. This correlation was used for the Crystal River 3, cycle 1, restart report and is currently used to license all operating B&W 3

plants with Mark-B fuel assembly cores.

2. The assumed system flow was changed from 105 (cycle 1) to 106.5% (cycle 2) of the design flow of 88,000 gpm/ pump primarily to make the thermal-hydrau-lic design basis for Crystal River 3 consistent with that assumed for other HSW plants of similar design and rated power level (e.g. , Oconee 1, 2, 3, ANO 1, and TMI-1). This assumption is fully justified by measured flow data from Crystal River 3, which indicates a system flow in excess of 109.5% of design flow, including allowance for measurement error.

3. The fresh incoming batch 4 fuel inserted for cycle 2 is the Mark-B4 assem-bly design. Batches 1, 2, and 3 used for cycle 1 were Mark-B3 assemblies.

The Mark-B4 fuel assemblies differ from the Mark-B3 assemblies primarily in the end fittings, which have been modified to reduce assembly pressure drop and increase holddown margin. The reduced assembly pressure drop causes a slight increase in flow through the B-4 assemblies relative to the B-3 design. For cycle 2 operation, the highest steady-state peaks 6-1 Babcock & Wilcox

will occur in the fresh batch 4 Mark-B4 fuel assemblies. No credit has been taken in thermal-hydraulic evaluations for any increase in B-4 assem-bly flow resulting from a mixed core that includes Mark-B3 assemblies.

Similar core configurations (Mark-B3 in combination with Mark-B4 assem-blies) have successfully operated in a number of B&W reactors, including Oconee 1, 2, 3, ANO 1, and TMI-1. Mark-B4 assemblies are currently in all B&W operating reactors.

4. The fuel rod bow model and associated DNBR penalty was updated for cycle 2 evaluations. B&W has submitted an interim rod bow penalty evaluation procedure 12 for use until a topical report is completed and reviewed. As shown in reference 12, when this interim procedure is used, there is no DNBR penalty due to fuel rod bow for fuel burnup to approximately 21,300 mwd /mtU. For Crystal River 3, cycle 2, the limiting (highest power) fuel assemblies are always in fuel burned less than 21,300 mwd /mtU, and no DNBR rod bow penalty is required.

5. No orifice rod assemblies (ORA) or burnable poison rod assemblies (BPRAs) will be in the cycle 2 core. All ORAs and BPRAs were actually removed from the core before completion of cycle 1 operation. A separate report was filed with the NRC to justify continued operation without the ORAs and BPRAs.3

6. A reference design radial x local power peaking factor (FAH) of 1.71 was used for cycle 2 evaluations. The cycle 1 FAH of 1.78 was reduced to 1.71 in conjunction with ORA and BPRA removal and was documented in reference

3. The continued use of an FAH of 1.71 is fully supported by nuclear cal-culations that predict a maximum cycle 2 FAH of 1.441.

7. The densification power spike was eliminated from DNBR evaluations based on the NRC approval of this change in reference 11.

The cycle 1 and 2 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 operation 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 for 6-2 Babcock & VVilcox

each operating condition represent a locus of points for which the calculated minimum DNBR is greater than the B&W-2 CHF correlation limit of 1.30.

6.3. Flux /Flev Trip Setpoint Analysis The flux / flow trip is designed to protect the plant during pump coastdowna from four-pump operation or to act as a high flux trip during partial-pump operation. Crystal River 3, cycle 2, will have redundant pump monitors on each pump, which will trip the reactor immediately upon the loss of power to two or more pumps. Therefore, the flux / flow trip setpoint need only protect the plant during a one-pump coastdown from four-pump operation.

The margin for any assumed flux / flow setpoint is determined with a transient analysis of a one-pump coastdown initiated from 102% indicated power (108%

real power). The 6% full power difference between real power and indicated power accounts for 4% FP neutron power measurerent error and a 2% FP heat balance error. Actual measured one-pump coastdown data are used in the analy-sis, and maximum additive trip delays are used between the time trip condi-tions are reached and actual control rod motion starts. Once a flux / flow trip limit is found to be adequate by thermal-hydraulic analysis, error ad-justments are made to account for flow measurement noise and instrument error before the actual trip setpoint is determined.

The recommended cycle 2 thermal-hydraulic flux / flow trip limit of 1.10 (actual in-plant setpoint of 1.07) resulted in a transient minimum DNBR of 1.75 (B&W-2) during the pump coastdown. This represents >34% DNBR margin to the correlation limit of 1.30.

6.4. Loss-of-Coolant Flow Transients The one-pump coastdown analysis was discussed in conjunction with the flux /

flow setpoint analysis in section 6.3. The four-pump coastdown and locked-rotor transients were also analyzed for cycle 2. The results of these analy-ses are discussed in section 7, " Accident and Transient Analysis."

6-3 Babcock & kVilcox

Table 6-1. Cycle 1 and 2 Thermal-Hydraulic Design Conditions Cycle 1 Cycle 1 Cycle 2,

<268.8 EFPD >268.8 EFPD 2544 MWt 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 power peaking factor, 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 (a) 431 x 103 452 x 103 CHF correlation W-3 B&W-2 B& W-2 Minimum DNBR, % power 1.61 (114) 2.14 (112) 1.98 (112) 1.92 (102) 2.27 (108) 2.12 (108) 2.49 (102) 2.33 (102)

\

'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 calculatior.2, as described in reference 9 and accepted in reference 11.

(

140.2 inches is a conservative (minimum) value used in cycle 2 analy-sis; it is the minimum densified length for any B&W fuel. Specific densified lengths for CR-3 fuel are given in Table 4-2.

6-4 Babcock & \Vilcox

7. ACCIDENT AND TRANSIENT ANALYSIS 7.1. r;eneral Safety Analysis Each FSARI accident analysis has been examined with respect to changes in cycle 2 parameters to determine the effect of upgrading the reactor power from 2452 to 2544 MWt. Because the FSAR accident analysis and dose calcula-tions, with the exception of the four-pump coastdown and locked-rotor acci-dents, were done at a higher power level than the requested upgrade (i.e.,

2568 versus 2544 MWt), it was only necessary to examine the cycle 2 parame-ters relative to the FSAR values to ensure that the thermal performance during hypothetical transients is not degraded. Although the FSARI states that all accidents were done at 2544 MWt, they were actually analyzed using the more conservative 2568 MWt.

The effects of fuel densification on the FSAR accident analysis results have been evaluated and are reported in reference 7. Since batch 4 reload fuel assemblies 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 MWt for consistency with other B&W reactors using the analytical methods documented in the FSAR I and updated in the Fuel Densification Report.7 The input param-eters used for these accidents are given in Table 6-1 and section 7.6. No new dose calculations were performed for the reload report except for the doses associated with the letdown line rupture (section 7.16), which is a requirement of reference 4. The other dose considerations in the FSAR were based on a 2568-MWt power level and maximum peaking and burnup for all core cycles; therefore, the dose considerations are independent of the reload batch or power upgrade to 2544 MWt.

The effects of BPRA and ORA removal from the cote on accident consequences were considered in reference 3. Basically, to gain DNBR margin to offset, the loss of core flow available for heat transfer, the design radial x local 7-1 Babcock & Wilcox

peaking factor (FAH) was reduced from 1.78 to 1.71. Additional details are included in section 6.1. Each accident is independently addressed in sections 7.3 through 7.16.

7.2. Accident Evaluation The key parameters that have the greatest effect 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 2 are given in Table 4-2.

Table 6-1 compares the cycle 2 thermal-hydraulic maximum design conditions to the previous cycle values, and a comparison of the key kinetics parameters from the FSAR and cycle 2 is provided in Table 7-1. Table 7-2 is a tabulation showing the bounding values for allowable LOCA peak linear heat rates for Crystal River 3, cycle 2 fuel.

It is concluded from the loss-of-flow analysis (section 7.6) and by examination of cycle 2 core thermal and kinetics properties with respect to acceptable FSAR values that this core reload will not adversely affect tha ability to safely operate the Crystal River 3 plant during cycle 2. Considering the previously accepted design basis used in the FSAR, the transient evaluation of cycle 2 is considered to be bounded by previously accepted analyses. The initial con-ditions of the transients in cycle 2 are bounded by the FSAR with the excep-tion of the four pump coastdown and locked rotor accidents, which were redone at a core power of 102% of 2568 MWt.

7.3. Rod Withdrawal Accidents This accident is defined as uncontrolled reactivity addition to the core due to withdrawal of control rods during startup conditions or from rated power conditions. Both types of incidents were analyzed in the FSAR.

The important parameters during a rod withdrawal accident are Doppler co-efficient, moderator temperature coefficient, and the rate at which reactiv-ity is added to the core. Only high-pressure and high-flux trips are accounted for in the FSAR analysis, ignoring multiple alarms, interlocks, and trips that normally preclude this type of incident.

7-2 Babcock & Wilcox

For positive reactivity addition indicative of these events, the most severe results occur for BOL conditions. The FSAR values of the key parameters for BOL conditions were -1.17 x 10-5 Ak/k/*F for the Doppler coefficient 0.0 Ak/k/*F for the moderator temperature coefficient and rod group worths up to and including a 12.9% A/k/k rod worth. Comparable cycle 2 parametric values are -1.5 x 10-5 Ak/k/*F for Doppler coefficient, -0.65 v 10-" ak/k/ F for moderator temperature coefficient, and maximum rod bank worth of 8.66% Ak/k.

The FSAR used an initial rated power level of 2568 MWt for these accidents.

This is more conservative than initializiang the accident at 102% of 2544 MWt and tripping the reactor at 110% of 2544 versus 112% of 2568 MWt since mcre enrrgy is added to the system for the FSAR analysis. For the accidents which trip on high pressure, the effect of a higher initial power level (i.e., 102%

of 2544 MWt) is to cause the pressure trip to occur slightly sooner. There-fore, cycle 2 parameters are bounded by design values assumed for the FSAR analysis. Thus, for the rod withdrawal transients, the consequences will no more severe than those presented in the FSAR.

7.4. Moderator Dilution Accident Boron in the form of boric acid is used to control excess reactivity. The boron content of the reactor coolant is periodically reduced to compensate for fuel burnup and transient xenon effects with dilution water supplied by the makeup and purification system. The moderator dilution transients con-sidered are the pumping of water with zero boron concentration from the make-up tank to the RCS under conditions of full power operation, hot shutdown, and refueling.

The key parameters in this analysis are the initial boron concentration, boron reactivity worth, and moderator temperature coefficient for power cases.

For positive reactivity addition of this type, the most severe results occur for BOL cor.ditions. 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 2 values are 1084 ppm for the initial boron concen-tration, 105 ppm /1% Ak/k boron reactivity worth and -0.65 x 10-4 ok/k/*F for the moderator temperature coefficient. The FSAR used an initial rated power 7-3 Babcock & \Vilcox

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 during this event. Sufficient time for operator action to terminate this transient is also shown in the FSAR even with maximum dilution and minimum shutdown margin.

The predicted cycle 2 parameter values of importance to moderator dilution transient are bounded by the FSAR design values, thus, the analysis in the FSAR is valid.

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, the increased flow rate will cause the average core temperature to decrease.

If the moderator temperature coefficient is negative, reactivity will be added to the core and a power increase will occur.

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 analyzed as the most severe transient. The initial power level of 50% of 2568 ffWt 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 10-4 Ak/k/ F and least negative Doppler coefficient of -1.17 x 10-4 Ak/k/ F. The corresponding most negative moderator temperature coefficient and least negative Doppler coefficient pre-dicted for cycle 2 are -2.52 x 10-4 Ak/k/ F and -1.5 x 10-5 Ak/k/*F, respec-tively. As the predicted cycle 2 moderator temperature coefficient is less negative and the Doppler coefficient is more negative than the values used in the FSAR, the transient results would be less severe than those reported in the FSAR.

7.6. Loss of Coolant Flow (LOCF)

A reduction in reactor coolant flow can be caused by mechanical failure or a loss of electrical power to the pumps. The LOCF transients were re-analyzed 7-4 Babcock & Wilcox

for cycle 2 operation and assumed an initial power level of 102% of 2568 MWt for consistency with other B&W reactors.

7.6.1. Four-Pump Coastdown (4PCD)

The 4PCD transient has been analyzed under conditions that represent the most conservative that can occur for cycle 2 operation. These conditions include such key parameters as initial flow rate, flow

• ate versus time for the tran-sient, initial power level, Doppler c- moderator temperature coeffi-cient, and reference design radial x local power peaking factor (FAH). Table 7-3 compares the key parameters used in the analysis with those predicted for cycle 2. For all parameters, the value used in the analysis is either equal to the cycle 2 parameter or is more conservative. Figure 7-1 shows the be-havior of the power and flow versus time for this transient.

The results of the analysis are shown on Figure 7-2. The minimum DNBR of 2.10 (BAW-2) obtained during the transient is well above the DNBR correlation limit of 1.30. The fuel and cladding temperatures are not shown since there was no increase in these parameters. It is therefore concluded that no fuel damage will occur.

Table 7-4 provides a comparison of MDNBRs between the FSAR, Fuel Densification Report, and cycle 2 for both one- and four pump coastdowns. Additional DNBR margin is shown for cycle 2 due to the use of the B&W-2 CHF correlation instead of the W-3 CHF correlation.

7.6.2. Locked Rotor (LR)

The locked-rotor accident has been analyzed under conditions that represent the most conservative that can occur for cycle 2 operation. These conditions are the same as those in section 7.6.1 (4 PCD). Table 7-3 compares the key parame-ters used in the analysis with those predicted for cycle 2. For all parame-ters, the value used in the analysis is either equal to the cycle 2 parameter or is more more conservative. Figure 7-3 shows the behavior of the flow and power versus time for this transient.

The results of the analysis are shown on Figure 7-4. The maximum fuel temper-ature does hor. exceed the initial centerline fuel temperature of 4400F.

This temperature starts to decrease around 2 seconds into the accident. The analysis for the maximum transient cladding and fuel temperatures conservatively 7-5 Babcock & Wilcox

assumed film boiling at a DNBR of 1.43 instead of the correlation limit of 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 ll20F at 5.5 seconds after initiation of the accident. Less than 0.5% of the fuel pins in the core will experience a DNBR of less than 1.30, 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 If a control rod is dropped into the core while operating, a rapid decrease in neutron power would occur, accompanied by a decrease in core average coolant temperature. In addition, the power distribution may be distorted due to a new control rod pattern. Therefore, under these conditions, a return to rated power may lead to localized power densities and heat fluxes in excess of design limitations.

The key parameters for this transient are moderator temperature coefficient, worth of dropped rod, and local peaking factors. The FSAR analysis was based on 0.40% Ak/k rod worth with a moderator temperature coefficient of -3.0 x 10-4 Ak/k/ F. For cycle 2, the maximum worth rod at power is 0.20% Ak/k and the moderator temperature coefficient is -2.52 x 10-4 Ak/k/*F. Since the predicted rod worth is less and the moderator temperature coefficient more positive, the consequences of this transient are less severe than the results presented in the FSAR.

The effect of initializing these accidents at 2568 MWt as done 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 concern is low system pressure. Starting the accident at a higher power level (i.e., 102% of 2544 MWt) would yield slightly higher system pressures.

7.8. Loss of Electric Power Two types of power losses were considered in the FSAR: 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.

7-6 Babcock & VVilcox

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 largest offsite dose occurs for the second case, i.e., loss of all electrical power except unit batteries, and assuming operation with failed fuel and steam generator tube leakage. These results are independent of core loading; therefore, the results of the FSAR are applicable for any reload.

The dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. An initial power level of 102% of 2544 (2595) would yield doses approximately 1% higher than those reported in the FSAR. These doses remain well below the limits of the 10 CFR 100 guideline.

7.9. Steam Line Failure A steam line failure is defined as a rupture of any of the steam lines from the steam generators. Upon initiation of the rupture, both steam generators start to blow down, causing a sudden decrease in primary system temperature, pressure, and pressurizer level. The temperature reduction leads to positive reactivity insertion and the reactor trips on high flux or low RC pressure.

The FSAR has identified a double-ended rupture of the steam line between the steam generator and steam stop valve as the worst-case situation at end-of-life conditions.

The key parameter for the core response is the moderator temperature coeffi-cient which in the FSAR was assumed to be -3.0 x 10-4 Ak/k/*F. The cycle 2 predicted value of moderator temperature coefficient is -2.52 x 10-4 Ak/k/*F.

This value is bounded by that used in the FSAR analysis; hence, the results in the FSAR represent the worst situation.

The FSAR used an initial power level of 2568 MWt for these accidents. This is more conservative than running 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.

7.10. Steam Generator Tube 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 for this 7-7 Babcock & Wilcox

incident is the potential radiological release, which is independent of core loading. Hence, the FSAR results are applicable to this reload.

The dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. An initial power level of 102% of 2544 (2595) MWt would yield doses approximately 1% higher than those reported in the FSAR. These doses remain well below the limits of the 10 CFR 100 guideline.

7.11. Fuel Handling Accident The mechanical damage type of accident is considered the maximum potential source of activity release during fuel handling activity. The primary concern is over radiological releases which are independent of core loading, and, therefore, the results of the FSAR are applicable to all reloads. The dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. An initial power level of 102% of 2544 (2595) would yield doses approxi-mately 1% higher than those reported in the FSAR. These doses remain well below the limits of the 10 CFR 100 guideline.

7.12. Rod Ejection Accident For reactivity to be added to the core at a more rapid rate than by uncontrolled rod withdrawal, physical failure of a pressure barrier component in the 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. The values used in the FSAR and densification report at BOL conditions of -1.17 x 10-5 Ak/k/"F Doppler coefficient, 0.0 bk/k/ F moderator temperature coefficient, and ejected rod worth of 0.65% Ak/k represented the maximum possible transient.

The use of a 0.65% Ak/k maximum rod worth is conservative in comparison to the cycle 2 predicted value of 0.52% Ak/k. Furthermore, the cycle 2 predicted values of -1.5 x 10-5 Ak/k/*F Doppler and -0.65 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 7-8 Babcock & Wilcox

input bound the cycle 2 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 consequences of this accident are independent of core loading. Therefore, the results reported in the FSAR are applicable for all reloads.

The dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. An initial power level of 102% of 2544 (2595) MWt would yield doses approximately 1% higher than those reported in the FSAR. These doses remain well below the limits of the 10 CFR 100 guideline.

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 consequences of this incident are independent of core loading and, therefore, the results reported in the FSAR are applicable to any reload.

The dose calculations reported in the FSAR were based on a rated power level of 2568 MWt. An initial power level of 102% of 2544 (2595) MWt would yield doses approximately 17. higher than those reported in the FSAR. These doses remain well below the limits of the 10 CFR 100 guideline.

7.15. LOCA Analysis Generic LOCA analyses for B&W 177-FA lowered-loop NSSs have been performed using the Final Acceptence Criteria ECCS Evaluation Model. The large-break analysis is presented in a topical report 15, and is further substantiated in a letter report 20 The small break analysis is presented in a letter reportI8 These analyses used the limiting values of key parameters for all plants in the category. Furthermore, the average fuel temperaturc as a fuuetion of linear heat rate and lifetime pin pressure data used in the LOCA limits analysis 15 are conservative compared to those calculated for this reload. Thus, these 7-9 Babcock & \Vilcox

analyses and LOCA limits provide conservative results for the operation of Crystal River Unit 3, cycle 2.

Crystal River Unit 3's proposed long-term ECCS modification for small break LOCA is presented in reference 19.

The LOCA analyses used a power level of 2772 MWt, which is conservative rela-tive to the 2544 MWt rating. Table 7-2 shows the bounding values for allowa-ble LOCA peak linear heat rates for Crystal River Unit 3, cycle 2.

7.16. Failure of Small Lines Carrying Primary Coolant Outside Containment 7.16.1. Identification of Causes A break in fluid-bearing lines that penetrate the containment could result in the release of radioactivity to the environment. There are no instrument lines connected to the RCS that penetrate the containment. However, other piping lines from the RCS to the 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 active component will result 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 safety evaluation criterion for this accident is that resultant doses shall not exceed 10 CFR 100 limits.

7-10 Babcock & Wilcox

7.16.2.2. Methods of Analysis The CRAFT computer code 16 was used to determine the loss-of-coolant character-istics of this letdown line 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 oper-ating 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 as-sumption extends the time to reactor trip /ESFAS actuation and increases the mass and energy releases to the auxiliary building. Termination of the acci-dent was assumed following ESFAS actuation on low RC pressure (1500 psig) and closure of the letdown isolation valves inside the containment. An instrument error of 6% of full range was assumed for the ESFAS actuation pressure, and the letdown isolation valve was assumed closed 7.4 seconds after the ESFAS pressure setpoint was reached. The 7.4-second time period for complete valve closure considers both the instrumentation response time and the actual valve closure time. Credit was not taken for a reduction in break flow during the time the isolation valves were closing.

7.16.2.3. Environmental Consequences The time required for the RCr to reach the actuation pressure of 1350 psig (1500 psig minus 6% of 2500 psia) for the ESFAS 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 raunds of reactor coolant was assumed to volatilize and become air-borne in the auxiliary 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 filtered through HEPA and charcoal filters in the building's ventilation system before being exhausted to the environment. The analysis is based on a conservatively estimated charcoal filter iodine removal efficiency of 90%. The assumptions used in the evaluation of the offsite doses are summarized in Table 7-5. The atmospheric dispersion factors (X/Q) used to calculate the two-hour doses at 7-11 Babcock & Wilcox

the exclusion area boundary and the low population zone boundary are also listed in Table 7-5. The fission product activities released to the environ-ment during the accident are listed in Table 7-6.

7.16.2.4. Results of the Analysis The dose consequences of the letdown line rupture accident are presented in 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 population zone for the first two hours after the acci-dent. The resulting doses are small fractions of the 10 CFR 100 limits.

7-12 Babcock & \Vilcox

Table 7-1. Comparison of Key Parameters for Accident Analysis FSARI , Predicted densif'n7 cycle 2 Parameter value Cycle 1 3 value BOL Doppler coeff, 10-5 Ak/k/ F -1.17 -1.47 -1.5 (268 EFFD)

EOL Doppler coeff, 10-5 Ak/k/*F -1.30 -1.66 -1.58 (510 EFPD)

BOL moderator coeff, 10-+ Ak/k/*F 0" -0.75 -0.65 (268 EFPD)

EOL moderator coeff, 10-4 Ak/k/*F -4.0( -2.42 -2.52 (510 EFPD)

All-rod bank worth at BOL, HZP, 12.9 9.12 8.66

% Ak/k (268 EFPD)

Boron reactivity worth (HFP), 100 101 105 ppm /1% Ak/k Max ejected rod worth (HFP), % Ak/k 0.65 0.55 0.52 Dropped rod worth (HFP), % Ak/k 0.65 0.20 W. 20 Initial boron cone'n (HFP), ppm 1150 795 1084

( )+0.50 x 10-4 Ak/k/*F was used for the moderator dilution accident.

(b) 3.0 x 10-4 Ak/k/ F was used for the steam line failure analysis and dropped rod accident analysis.

Table 7-2. Bounding Values for Allowable LOCA Peak Linear Heat Rates Core Allowable elevation, peak LHR, ft kW/ft 2 15.5 4 16.6 6 18.0 8 17.0 10 16.0 7-13 Babcock & Wilcox

Table 7-3. Input Parameters to LOCF Transients Cycle 2 value Value used in analysis Initial flow rate, % >109.5 106.5 of 352,000 gpm Flow rate Vs time > Fig. 14-17, FSAR Fig. 14-17, FSAR (4PCD) r.;g.14-19a, FSAR Fig.14-19a, FSAR (LR)

Initial powe: level, 102% of 2544 102% of 2568 MW Dopple coeff, Ak/k/*F -1.5 x 10-5 -1.27 x 10-5 Moderator temp coeff, -0.65 x 10-4 0 Ak/k/*F FAH 1.44 1.71 Table 7-4. Summary of Minimum DNBR Results for Limiting LOCF Transients Cycle 1 Densif'n FSARI report 2 Cycle 2 Transien t (W-3) (W-3) (B&W-2)

One-pump coastdown (flux / flow #

NR NR 1.75 trip)

Four-pump coastdown (flux / flow 1.45 1.39 2.10 trip, cycle 1; pump monitor trip, cycle 2)

(#

NR: not reported.

7-14 Babcock & Wilcox

Table 7-5. Analysis Assumptions for MU6PS Letdown Line Rupture Accident Data and Assumptions Used to Estimate Radioactive Source Power level, MWt 2568 Percent of fuel rods leaking, % 1.0 Table for escape rate coeff 11-11 Reactor coolant activlties Nuclide Activity, rCi/mi 85Kr* 1.6 85Kr 11 87Kr 0.87 88Kr 2.8 131Xe' 2,5 133Xe* 2.9 133 Xe 260 135Xe* 0.97 135Xe 6.2 138 Xe 0.53 1311 3.3 1 321 5.0 I?31 3.9 1341 0.52 135I 2.0 Data and Assumptions Used to Estimate Radioactivity Released Total mass of reactor coolant released 45,760 to auxiliary building, lb Charcoal filter efficiency for Iodine, % 90 Noble gas, % 0 Fraction of airborne iodine 0.1 Dispersica Data Exclusion area boundary, meters 1340 Low population zone boundary, meters 8047 Atmospheric dispersion percentile, % 5 0-2 h atmospheric dispersion factors, s/m 3 at exclusion area boundary 1.6 x 10-4 at low population zene boundary 1.4 x 10-5 7-15 Babcock & Wilcox

Table 7-6. Activity Released to Environment Due to Rupture of MU6PS Letdown Line Nuclide Activity, Ci 85Kr* 46.1 85Kr 317 87 Kr 25.1 88 Kr 80.6 131Xe* 72.0 133Xe" 83.5 133 Xe 7490 m

135xe 27.9 135xe 179 138 Xe 15.3 1311 0.95 1321 1.44 1331 1.12 134 1 0,15 1351 0.58 Table 7-7. Radiological Consequences of MU&PS Letdown Line Rupture Outside Containment 0-2 h dose at exclusion area boundary, Rem Thyroid (inhalation) 0.111 Whole body (gamma) 0.033 Skin (beta) 0.049 0-2 h dose at low population boundary, Rem Thyroid (inhalation) 9.75 x 10-3 Whole body (gamma) 2.91 x 10-3 Skin (beta) 4.27 x 10-3 7-16 Babcock & Wilcox

Figure 7-1. Four-Pump Coastdown Power and Flow Transients. Crystal River 3 1.00 -

Flow 0.90 -

0.80 -

g Power o

o 0 0.70 -

u.

8 E

] 0.60 -

b 5

m 0.50 -

0.40 -

0.30 _

, i i 0 0.5 1.0 1.5 2.0 Time, Seconds 7-17 Babcock & Wilcox

Figure 7-2. Four-Pump Coastdown - llot Channel MDNBR Vs Time, Crystal River 3 2.34 2.30 _

y 2,26 -

5 m

o g 2.22 -

a:

E o

E E 2.18 -

c E

2.14 -

2.10 i , i 0 0.5 1.0 1.5 2.0 Time, Seconds 7-18 Babcock & Mcox

Figure 7-3. Locked-Rotor Power and Flow Transients, Crystal River 3 1.00 0.90 -

0.80 -

Flow

/

8 0.70 -

8 i

0.60 g

u- Power E

[ 0.50 -

E a.

0.40 -

0.30 -

0.20 -

0.10 -

t t 1 t f  !

0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 Time, Seconds 7-19 Babcock & Wilcox

E-Mf8 '0!ied 8NO mnw!u!H a u o a w a n d A A -

~

~ '

' o s'

o*

w m

W e

+

m o

- m e

a m

h 4 a o e o o.

h -

9 2 e$

O - V U e C O 3 0 x w g 4a a e

a a s -

o. .

U W M o

g O -

.-2 i s G

'Is o 4

8 o

6C

• m "z o g -

J o

o o o 8

m -o 8 o o

o o

o o

o o

e s e j, 'a;nge;adwal acepng pelo unm!xeH i t i , , , , ,

o o o o o o o

2 o

N o

O o

D o

w o

2 8

N a a a m n n m 3, 'o;nte;adwai Jagua3 [anj wnt' J!xeH 7-20 Babcock & Wilcox

8. PROPOSED MODIFICATIONS TO TECHNICAL SPECIFICATIONS All technical specifications have been revieweu by Florida Power Corporation and B&W and some were revised for cycle 2 operation. The Technical Specifi-cation sections to which modifications have been made are listed in Table 8-1 and are shown on the following pages. The reanalysis of Technical Specifica-tions for cycle 2 operation used the same analytical techniques as the modi-fled cycle 1 design.3 The review of the 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.

8-1 Babcock & \Milcox

Table 8-1. Technical Specification Changes Tech Spec Report page No. (figure, Nos. (figure table Nos.) Nos.) Reason for change 1.3 8-3 Rated thermal power increased to 2344 MWt.

2.1 (2.1-1, 8-4 (8-1, Pressure / temperature limits changed due to

-2, bases 8-2, 8-4) higher assumed flow, use of BAW-2 CHF corre-for 2.1, lation, and power upgrade (section 6). The Figure 2.1) flux /Aflux envelopes changed due to power upgrade.

2. 2 (2. 2-1, 8-5, 8-6, Flux / flow trips changed due to cddition of RC 2.2-1, bases Table 8-2, pump power monitors, setpoint is new based on for 2.2) Figure 8-3 one-pump versus four pump coastdown.

3.1. 3. 6 (3.1- Figures Specs 3.1.3.6, 3.1.3.7, 3.1.3.9, and 3.2.1 1, -2, -3, -4 ) 8 8-8 reflect revised nuclear parameters as a re-sult of the cycle 2 reload, including the power upgrade.

3.1. 3. 7 (3.1-7, -8,

-9) 3.1. 3. 9 (3.1- Figures 8-10, 9, -10 8-11 3.2.1 (3.2-1, Figures 8-12,

-2) 8-13 3.2.2 8-7, 8-8 Reflects increased power.

3.2.4 (Table Table 8-3 Error-adjusted limits changed to reflect in-3.2-2) creased age of detectors.

3.2.5 (Table Table 8-4 Reflects increased power and assumed flow 3.2-1) (see section 6.1).

8-2 Babcock & Wilcox

1. 0 DEFINITIONS DEFINED TERMS 1.1 The DEFINED TERMS of this section appear in capitalized type and are applicable thrJughout these Technical Specifications.

THERMAL POWER 1.2 THERMAL POWER shall be the total reactor core heat transfer 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 correspond to any one inclusive combination 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-fications.

OPERABLE - OPERABILITY

1. 6 A system, subsystem, train, component or device shall be OPERABLE or 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 sources, cooling 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).

CRYSTAL RIVER - UNIT 3 8-3 Babcock & VVilcox

2.1 SAFE Y LIMITS BASES 2.1.1 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 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 result in excessive cladding temperatures because of the onset of departure 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 of 1.30 is predicted for the maximum possible thermal power, 112% when the reactor coolant flow is 139.6 x 106 lb/h, which is 106.5% of l the design flow rate for four operating reactor coolant pumps. This curve is based on the following nuclear power peaking factors with potential fuel densification effects:

F"7= 1.50.

= 2.57; F = 1.71; H

The dcsign limit power peaking factors are the most restrictive calculated at full power for the range from all control rods fully withdrawn to minimum allowable control rod withdrawal, and form the core DNBR design basis.

CRYSTAL RIVER - UNIT 3 8-4 Babcock & Wilcox

l LIMITING SAFETY SYSTEM SETTINGS BASES RCS Outlet Temperature - High The RCS outlet temperature high trip 5 619*F prevents the reactor outlet temperat're J from exceeding the design limits and acts as a backup trip for all power excursion transients.

Nuclear Overpower Based on RCS Flow and AXIAL POWER IMBALANCE 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 &: creasing 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 power level increases or the reactor coolant flow rate decreases. The power level setpoint produced by the power-to-flow ratio provides overpower DNB protection for all modes of pump operation. For every flow rate there is a maximum permissible power level, and for every power level there is a minimum 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 79.9% and reactor flow rate is 74.7%, or flow rate is 5 70.1% and power is 75%.

For safety calculations the maximum calibration and instrumentation errors for the power level were used.

CRYSTAL RIVER - UNIT 3 8-5 Babcock & Wilcox

LIMITING SAFETY SYSTEM SETTINGS BASES The AXIAL POWER IMBALANCE boundaries are established in order to pre-vent reactor thermal limits from being exceeded. These thermal limits are either power peaking kW/ft limits or DNBR limits. The AXIAL POWER IMBALANCE 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 power level trip and associated reactor power-reactor power-imbalance bound-aries by 1.07% for a 1% flow reduction.

RCS Pressure - Low, High and Variable Low 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 from low power or a slow reactivity insertion from high power, the RCS pressure-high setpoint is reached before the nuclear overpower trip setpoint. The trip setpoint for RCS pressure-high, 2355 psig, has been established to maintain the system 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 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 a pressure reduction. It also prevents reactor operation at pressures below the valid range of DNB correlation limits, protecting against DNB.

Due to the calibration and instrumentation errors, the safety analysis used an RCS pressure-variable low trip setpoint of (11.80 Tout F-5245.2) psig.

Reactor Containment Vessel Pressure - High The reactor containment vessel pressure-high trip setpoint, c 4 psig, provides positive assurance that a reactor trip will occur in the unlikely event of i steam line failure in the containment vessel or a loss-of-coolant accident, even in the absence of an RCS pressure-low trip.

l CRYSTAL RIVER - UNIT 3 8-6 Babcock & Wilcox

POWER DISTRIBUTION LIMITS NUCLEAR HEAT FLUX HOT CHANNEL FACTOR - Fg LIMITING CONDITION FOR OPERATION 3.2.2 F shall be limited by the following relationships:

q 3.08 Fq s p THERMAL POWER where P = RATED THERMAL POWER and P s 1.0.

APPLICABILITY: MODE 1.

ACTION:

With Fg exceeding its limit:

a. Reduce THERMAL POWER at least 1% for each 1% Fo exceeds the limit within 15 minutes and similarly reduce the NucTear Overpower Trip Setpoint and Nuclear Overpower based on RCS Flow and AXIAL POWER IMBALANCE Trip Setpoint within 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.

b. Demonstrate through in-core mapping that Fg is within its limit within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after exceeding the limit or reduce THERMAL POWER to less than 5% of RATED THERMAL POWER within the next 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />.

c. Identify and correct the cause of the cut of limit condition prior to increasing THERMAL POWER above the reduced limit re-quired by a or b, above; subsequent POWER OPERATION may proceed provided that FQ is demonstrated through in-core mapping to be within its limit at a nominal 50% of RATED THERMAL POWER prior to exceeding this THERMAL POWER and within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after attaining 95% or greater RATED THERMAL POWER.

SURVEILLANCE REQUIREMENTS 4.2.2.1 FQ shall be determined to be within its limit by using the incore detectors to obtain a power distribution map:

CRYSTAL RIVER - UNIT 3 8-7 Babcock & Wilcox

POWER DISTRIBUTION LIMITS BASES F

g Nuclear Enthalpy Rise Hot Channel Factor, is defined as the ratio of the integral of linear power along the rod on which minimum DNBR occurs to the average rod power.

It has been determined by extensive analysis of possible operating power shapes that the design limits on nuclear power peaking and on minimum DNBR at full power are met, provided:

Fq s 3.08 F 6g 5 1.71 Power Peaking is not a directly observable quantity and therefore limits have been established on the bases of the AXIAL POWER IMBALANCE produced by the power peaking. It has been determined that the above hot channel factor limits will be met provided the following conditions are maintained.

1. Control rods in a single group move together with no individual rod insertion differing by more than 6.5% (indicated position) from the group average height.

2. Regulating rod groups are sequenced with overlapping groups as re-quired in Specification 3.1.3.6.

3. The regulating rod insertion limits of Specification 3.1.3.6 and the axial power shaping rod insertion limits of Specification 3.1.3.9 are maintained.

4. AXIAL POWER IMBALANCE limits are maintained. The AXIAL POWER IMBAL-ANCE is a measure of the difference in power between the top and bottom halves of the core. Calculations of core average axial peak-ing factors fro many plants and measurements from operating plants ur. der a variety of operating conditions have been correlated with AXIAL POWER IMBALANCE. The correlation shows that the design power shape is not exceeded if the AXIAL POWER IMBALANCE is maintained within the limits of Figures 3.2-1 and 3.2-2.

The design limit power peaking factors are the most restrictive calcu-lated at full power for the range from all control rods fully withdrawn to minimum allowable control rod insertion and are +,e core DNBR design bash.

Therefore, for operation a* a fracticn of RATfD THERMAL POWER, the design limits are met. When using incore detectors uo make power distribution maps to determine FQ and FaH:

a. Themeasurementoftotalpeakingfactor,Fheas , shall be increased by 1.4 percent to account for manufacturing tolerances and further increased by 7.5 percent to account for measurement error.

CRYSTAL RIVER - UNIT 3 8-8 Babcock & Wilcox

Table 8-2. RPS Trip Setpoints Table 2.2-1. Reactor Protection System Instrumentation Trip Setpoints__

Functional unit Trip setpoint Allowable values

1. Manual reactor trip Not applicable Not applicable

2. Nuclear overpower 5 105.5% of RATED THERMAL POWER 5 105.5% of RATED THERMAL POWER with four pumps operating with four pumps operating 5 79.9% of RATED THERMAL POWER s 79.9% 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 POWER IMBALANCEa a

5. RCS pressure-low 1800 psig 2 1800 psig

= 6. RCS pressure-high 5 2355 psig 5 2355 psig

7. RCS pressure-variable- 2 (11.80 Tout F-5209.2) psig 2 (11.80 Tout F-5209.2) psig lowa

8. Nuclear overpower 5 125% of RATED THERMAL POWER 5 125% of RATED THERMAL POWER based on RCPPMsa with three pumps operating with three pumps operating 5 0% of RATED THERMAL POWER 5 0% of RATED THERMAL POWER with less than three pumps op- with less than three pumps op-erating erating

9. Reactor containment 5 4 psig 5 4 psig vessel 5 a g Trip may be manually bypassed when RCS pressure s 1720 psig by actuating the shutdown bypass, o provided that (1) the nuclear overpower trip setpoint is 5 5% of RATED THERMAL POWER, (2) the

x- shutdown bypass RCS pressure-high trip setpoint of 5 1720 psig is imposed, and (3) the shutdown e* bypass is removed when RCS pressure > 1800 psig.

w

Table 8-3. Quadrant Power Tilt Limits Table 3.2-2. Quadrant Power Tilt Limits Steady-state Transient Maximum limit limit limit Measurement independent 4.92 11.07 20.0 QUADRANT POWER TILT QUADRANT POWER TILT as measured by:

Symmetrical incore 3.46 8.96 20.0 detector system Power range channels 1.96 6.96 20.0 Minimum incore de- 1.90 4.40 20.0 tector system Table 8-4. DNBR Limits Table 3.2-1. DNB Margin Four RC pumps Three RC pumps Parameter operating operating a

Reactor coolant hot leg 5 604.6 5 604.6 temperature, T g, F a

2 2,061.6 2 2,057.2 Reactor coglant pres-sure, psig Reactor coolant flow 2 374,880 2 280,035 rate, gpm

^ 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 prr minute or a THERMAL POWER step increase of greater than 10% of RATED THERMAL POWER.

8-10 Babcock. 4. Wilcox

Figure 8-1. Reactor Core Safety Limits 2400 RCS Pressure-High Trip 2300 -

RC Outlet Temp High Trip 2200 -

.?

E

[ 2:0g _ ACCEPTABLE y OPERATION e

ct 5

5 2000 -

+

J' N Safety Limit O $s' d

1900 - #

S

/ UNACCEPTABLE OPERATION 1800 RCS Pressure Low Trip I t I t i 580 590 600 610 620 630 Reactor Outlet Temperature, *F a-11 Babcock & Wilcox

Figure 8-2. Reactor Core Safety Limits

(-43.68,112) (26.88,112)

(-50,105) / l10

._100 ACCEPTABLE (5000) 4 PUNP OPERATION

(-43.68,86.4)

~ ~

(26.88,86.4)

(-50.79.4)

ACCEPTABLE U'7"*'I

-- 70 3&4 PUMP OPERATION h

__ 20

_ _ 10 i i i i i I

-60 -40 -20 0 +20 +40 +60 Reactor Power Imbalance, %

8-12 Babcock & Wilcox

Figure 8-3. Trip Setpoints

-- 120 00 - - 110 (ioy) s- 4#*

4s / ACCEPTABLE g

/ I 4 PUMP -

- 100 g *#4g

(-37.85, 99) l OPERATION l (37.41,43)

- 90 I

I 80 I(79.9) l ACCEPTABLE I

(-37.85,71.9) l 3&4 PUMP l

- - 70 l OPERATION l (37. l.65.9 )

l I

_ _ 60 g i I

__ 50 l I l _ _ 40 l l 1 l -

- 30 I l l

$ ol Io -

4 dl

- 20 lg ,

7 7I I T S I  !

1:  : _

_ l0 si

l  !:ii  :

i i l i t

-40 -20 0 +20 +40 Reactor Power lmbalance, %

8-13 Babcock & Wilcox

Figure 8-4. Pressure-Temperature Limits 2400 r I

,cm 2200 -

E i 2 E

g 2000 -

5 2

8 1800 -

1600 i i i 550 600 620 640 660 Reactor Outlet Temp, F POWER PUMPS OPERATING CURVE  % OF 2568 (TYPEOFLIMIT) l ll2 4 PUMPS (DNBR) 2 86.4 3 PUMPS (DNBR) 8-14 Babcock & Wilcox

Figure 8-5. Regulating Rod Group Insertion Limits for Four-Pump Operation From 0 to 225 10 EFPD

!!0 (180,102) (225, i t,. )

100

- Power Level Cutoff (90f. UNACCEPTABLE ofRatedThermalPower) OPERATION g _ (180,90) iN i (225,90)

O g g0 - (165,80) (240,80)

UNACCEPTABLE

% OPERATION 2 70 -

2 ACCEPTABLE OPERATION 2 60 -

50 -

(300'50) >

0 (100,50) u

. t40 -

5 5

c. 30 -

20 -

10 -

0 ' i  ! . ' ' ' ' ' ' '

0 50 100 150 200 250 300 Rod index, f. Withdrawn 0 2,5 50 75 O 50 IQO 2,5 7,5 10,0 Group 5 Group 7 0 25 50 75 100 l t  !  ! I Group 6 8-15 Babcock & Wilcox

Figure 8-6. Regulating Rod Group Insertion Limits for Four-Pump Operation After 225 10 EFPD llo (270,102) 100 "

90 -

Power Level Cutoff (00%

of Rated Thermal Power) 37 80 -

u (255,80) i O

70 -

Ts E

b UNACCEPTABLE g 60 -

OPERATION

?

= 50 -

ACCEPTABLE (175.50) OPERATION o

" L40 -

C e 30 -

20 -

10 -

(0,0) 0 i i e i t ' '

0 50 !00 150 200 250 300 Rod index, % Withdr3wn 0, 25, 50 75

, , , 10,0 0, , 25, 50 75 1g0

- i Group 5 Group 7 0 25 50 75 100 1 t i f i '  !

Group 6 8-16 Babcock & Wilcox

Figure 8-7. Regulating Rod Group Insertion Limits for Three-Pump Operation From 0 to 225 10 EFPD 100 90 --

(247 0.76.5) 80 -

(158.3,76.5) u 70 -

y UNACCEPTABLE 2 OPERATION

.. 60 -

E b

1. SO -

(100,50) (300.50 0) '

E '

4; ACCEPTABLE 40 -

[ OPERATION c

i 30 -

2 20 20 -

(0.0) 0 ' ' ' ' ' ' ' ' ' ' '

O 25 50 75 100 125 150 175 200 225 250 275 300 Rod index, % Withdrawn 0 25 50 75 100 0 25 s

'0 75 i ) } t 100 t I f f I Group 5 Group 7 0 25 50 75 100 l t l l l Group 6 8-17 Babcock & Wilcox

Figure 8-8. Regulating Rod Group Insercion Limits for Three-Pump Operation After 225 10 EFPD 100 90 -

80 - UNACCEPTABLE OPERATION (246,76.5) u g 70 -

n.

[e 60 -

N u 50 -

(175,50) a

% 40 _

o W'

u

• 30 -

ACCEPTABLE f OPERATION 20 -

10 -

(0,0) 0 , , , , , i t i t i i  !

0 25 50 75 100 125 150 175 200 225 250 275 300 Rod index, 7. Withdrawn 0 2,5 50 0 25 g i 7,5 10,0 i i 5,0 7, 5 100 s

Group 5 Group 7 0 25 50 75 100 i ,

, i , i Group 6 8-18 Babcock & Wilcox

Figure 8-9. Control Rod Locations I @ @ @

G @ @ @

@ @ @ @ G @ @

@ Q @ @ @ @ @

O @ @ @ @ @

@ @ O O

, @ @ @ l GROUP NUMBER OF RODS FUNCTION I 8 SAFETY 2 8 SAFETY 3 12 SAFETY 4 9 SAFETY 5 8 CONTROL 6 8 CONTROL 7 8 CONTROL 8 8 APSRs TOTAL 69 8-19 Babcock & Wilcox

Figure 8-10. APSR Position Limits for 0 to 225 10 EFPD, Crystal River 3 110 (7,102) (31,102) 100 -

a (31,90) 90 - '( 7,9 0 )

u UNACCEPTABLE

,I g39,gn) OPERATION a- 80 , (0,80)

~

n

?

jE 70 -

s 18 j 60 -

'E

"- 50 -

(100,50) '

d x ACCEPTABLE e 40 -

OPERATION 30 -

20 -

10 -

0 i , , , i i i 0 10 20 30 40 50 60 70 80 90 100 Rod Position ,7, Withdrawn 8-20 Babcock & Wilcox 9

Figure 8-11. APSR Position Limits After 225 10 EFPD, Crystal River 3 110 (8,102) (34,102) 100 -

90 -

,(s,90) (34,30) UhACCEPTABLE OPERATION 80 ' (o,80) N'M E

x E 70 -

e b 60 -

50 -

ACCEPTABLE 'U g OPERATION e 40 -

5

[ 30 -

20 10 0 i i i 1 ' ' ' ' '

0 10 20 30 40 50 60 70 80 90 100 Pod Position, f. Withdrawn 8-21 Babcock & Wilcox

Figure 8-12. Axial Power Imbalance Envelope for Operation From 0 to 225 10 EFPD Pcwer, % of Rated Thermal Power

(-10.2,102) (10.2,102)

-- 100

(-15.3,90) (10.8,90)

-- 90 4

(-25.4,80), - 80 I,(12,80)

- 70 ACCEPTABLE UNACCEPTABLE OPERATION UNACCEPTABLE OPERATION - -. 60 OPERATION

_ _ 50 E_

__ 40 4 5

5

_ 50 $

t

__ 20 0 4

_ _ 10 t f f f f 1  ! t

-40 -30 -20 10 0 10 20 30 40 Axial Power Imbalance, %

8-22 Babcock & Wilcox

Figure 8-13. Axial Power 1mbalance Envelope for Operation After 225 10 EFPD Power, % of Rated Thermal Power

(-19.3.102) (18.3,102)

-100

(-20.7,90) - -

- 90 4 (18,92)

(-29,80) 1

- - 80 > (19.6,80)

ACCEPTABLE -

- 70 OPERATION

~ ~

UNACCEPTABLE OPERATION 50

. 40

$5 5p

_ _30 UNACCEPTABLE Iy OPERATION

. _20

- 10 t 1 1 l 1 f f i t I

-50 -40 -30 -20 -10 0 10 20 30 40 50 Axial Power imbalance, %

8-23 Babcock & Wilcox

9. STARTUP PROGRAM - PHYSICS 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 the unit.

9.1. Precritical Tests -- 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. Acceptance 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 criterion of 1.40 seconds corrected to a 75%-inserted position (by rod insertion versus time correlation) is 1.66 seconds.

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 rod withdrawal required in achieving equilibrium boron. The acceptance criterion placed on critical boron concentration is that the actual boron concentration must be within t100 ppm boron of the predicted value.

9.2.2. Temperature Reactivity Coefficient The isothermal temperature coefficient is measured at approximately the all-rods-out configuration and at the hot zero power rod insertica limit. The 9-1 Babcock & Wilcox

average coolant temperature is varied by first decreasing then increasing tem-perature by 5'F. During the change in temperature, reactivity feedback is compensated by discrete change ir. rod motion, the change in reactivity is then calculated by the summation of reactivity (obtained from reactivity calculation 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 0.4 x 10-4 (Ak/k)/*F.

The moderator coefficient of reactivity is calculated by adjusting the mea-sured temperature coefficient by the predicted value of the fuel Doppler co-efficient. This value must not be in excess of the acceptance criteria of 0.9 x 10-4 Ak/k/*F.

9.2.3. Control Rod Group Reactivity Worth Control bank group reactivity worths (groups 5, 6, and 7) are measured at hot zero power conditions using the boron / rod swap method. The boron / rod swap method consists of establishing 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 in incremental steps. The reactivity changes that 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 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, 7 worth:

predicted value - measured value measured value x 100 < 15

2. Sum of groups 5, 6, and 7:

predicted value - measured valu measured value x 100 s 10 9.2.4. Ejected Control Rod Reactivity Worth After the CRA groups have been positioned near the minimum rod insertion limit, the ejected rod is borated to 100% withdrawn and the worth obtained by adding the incremental changes in reactivity by boration.

9-2 Babcock & \Vilcox

After the ejected rod has been borated to 100% withdrawn and equilibrium boron established, the ejected rod is then swapped in versus the controlling rod group and the worth determined by the change in the previously calibrated con-trolling rod group position. The boron swap and rod swap values are averaged and error-adjusted to determine ejected rod worth. Acceptance criteria for the ej ected rod worth test are as follows:

predicted value - measured value 1, x 100 -< 20 measured value

2. Measured value (ecror-adjusted) < l.0% Ak/k 9.3. Power Escalation Tests 9.3.1. Core Power Distribution Verification at 440, 75, 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 pla-teau. Rod index is established at a nominal full-power rod configuration at which 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.

The following acceptance criteria are placed on the 40% FP test:

1. The worst-case maximum linear heat rate must be less than the LOCA limit.

2. The minimum DNBR must be greater than 1.30.

3. The value obtained from the extrapolation of the minimum DNBR to the next power plateau overpower trip setpoint must he 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 setpoint must be less than the fuel melt limit or the extrapalated value of imbalance must fall outside the RPS power / imbalance / flow trip envelope.

5. The quadrant power tilt shall not exceed the limits specified in the Tech-nical Specifications.

9-3 Babcock & Wilcox

6. The highest measured and predicted radial peaks shall be within the fol-lowing limits:

predicted measured measured x 100 <8

7. The highest measured and predicted total peaks shall be within the fol-lowing limits:

predicted measured measured x 100 < 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 plateau may be accomplished without exceeding the safety limits specified by the safety analysis with regard to DNBR and linear heat rate.

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 accept-ance criteria are as follows:

1. The highest measured and predicted radial peaks shall be within the follow-ing limits:

predicted measured x measured 100l<5

2. The highest measured and predicted total peaks shall be within the follow-ing limite:

predicted measured measured x 100 < 7.5 9.3.2. Incore Vs Excore Detector Imbalance Correlation Verification at %40% FP Imbalances are set up in the core by control rod positioning. Imbalances are read simultaneously on the incore detectors ad excore power range detectors for various imbalances. The imbalances from the excore detectors must exceed 9-4 Babcock 8.Wilcox

those on the incore detectors by a factor of 1.25. It the ratio of excore to incore detector imbalance is less than 1.25, gain amplifiers in the excore de-tector signal processing equipment are adjusted to provide the required gain.

9.3.3. Temperature Reactivity Coefficient at %100% FP The average reactor coolant temperature is decreased and then increased by about 5*F at constant reactor power. 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 reactivity coefficient 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 N100% FP Reactor power is decreased and then increased by about 5% FP. The reactivity change is obtained from the change in controlling rod group positicn. Control 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 coefficient is cal-culated f rom the measured reactivity change, adjusted as stated above, and the measured power' change. Acceptance criteria state that the measured value shall be more negative than -0.55 x 10-4 (Ak/k)/%FP.

No core power distribution test will be performed at the previous fu!.1 power level of 2452 MWt. Since 2452 MWt is 96.4% of the new thermal power level (2544 MWt) and the final power distribution test will be performed at 98-100%

of 2544 MWt, the difference is too small to necessitate an additional test.

This is further supported by the several B&W cores of this design which are operating at 2568 MWt.

9.4. Procedure for Failure to Meet Acceptance Criteria FPC reviews the results of all startup tests to ensure that all acceptance criteria are met. If the review of the test indicates that the results are well within the acceptance criteria, no further evaluation is conducted. If the review indicates that the results are approaching or close to the accep-tance criteria limits, further evaluation of that particular test or other 9-5 Babcock & Wilcox

supporting tests is performed to look for trends. This evaluation will deter-mine whether additional support data are required to discover any abnormal conditions. If acceptance criteria for any test are not met, an evaluation is performed before the test program is continued. This evaluation is performed by site test personnel with participation by Babcock & Wilcox technical person-nel as required. Further specific actions depend on c. valuation results. These actions can include repeating the tests with more detailed attention to test prerequisites, added tests to search for anomalies, or design personnel per-forming detailed analyses of potential safety problems because of parameter deviation. Fower is not escalated until evaluation shows that plant safety will not be compromised by such escalation.

9-6 Babcock & Wilcox

REFERENCES 1

Crystal River Unit 3, Final Safety Analysis Report, Docket 50-302, Florida Power Corp.

2 Crystal River Unit 3, Technical Specification Change Request No. 27, Dock-et 50-302, License DPR-72, November 29, 1978.

3 Crystal River Unit 3, Licensing Considerations for Continued Cycle 1 Oper-ation Without Burnable Poison Rod Assemblies and Orifice Rod Assemblies, BAW-1490, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, July 1978.

" Standard Format and Content of Safety Analysis Reports for Nuclear Power Plants," USNRC Regulatory Guide 1.70, Revision 3, November 1978.

5 BPRA Retainer Design Report, BAW-1496, Babcock & Wilcox, Lynchburg, Vir-ginia, May 1978.

6 A. F. J. Eckert, H. W. Wilson, and K. E. Yoon, Program to Determine Per-formance of B&W Fuels - Cladding Creep Collapse, BAW-10084P-A, Rev. 2, Babcock & Wilcox, Lynchburg, Virginia, January 1979.

7 Crystal River Unit 3, Fuel Densification Report, BAW-1397, Babcock & Wil-cox, Lynchburg, Virginia, August 1973.

8 C. D. Morgan and H. S. Kao, TAFY - Fuel Pin Temperature and Gas Pressure Analysis, BAW-10044, Babcock & Wilcox, Lynchburg, Virginia, May 1972.

9 Fuel Densification Report, BAW-10055, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, July 1973.

10 K. E. Suhrke (B&W) to S. A. Varga (NRC), Letter, "Densification Power Spike," December 6, 1976.

11 S. A. Varga (NRC) to J. H. Taylor (B&W), Letter, " Update of BAW-10055, Fuel Densification Report," December 5, 1977.

12 J. H. Taylor (B&W) to D. B. Vassallo (NRC), Letter, " Determination of the Fuel Rod Bow DNB Penalty," December 13, 1978.

A-1 Babcock & Wilcox

13 Correlation of Critical Heat Flux in Bundle Cooled by Pressurized Water, BAW-10000A, Babcock & Wilcox, Lynchburg, Virginia, May 1976.

14 Babcock & Wilcox Version of PDQ User's Manual, BAW-10ll7P-A, Babcock &

Wilcox, Lynchburg, Virginia, January 1977.

15 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.

16 CRAFT 2 - FORTRAN Program for Digital Simulation of a Multinode Reactor Plant During Loss of Coolant, BAW-10092, Babcock & Wilcox, Lynchburg, Virginia, April 1975.

17 S. A. Varga (NRC) to J. H. Taylor (B&W), Letter, " Comments on B&W's Sub-mittal on 'ombination of Peaking Factors," May 13, 1977.

18 J. H. Taylor (B&W) to S. A. Varga (NRC), Letter, "ECCS Small Break Analy-sis," July 18, 1978.

19 W. P. Stewart (FPC) to R. W. Reid (NRC), Letter, d Crystal River Unit 3, Docket No. 50-302, Operating License No. DPR-72, ECCS Small Break Analy-sis," January 12, 1979.

20 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 a Revised System Pressure Distribution," July 8, 1977.

21 R. V. DeMars and R. R. Steinke, Fuel Assembly Stress and Deflection Analy-sis for LOCA and Seismic Excitation, BAW-10035, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, December 1972.

A-2 Babcock & Wilcox