ML20154J525
| ML20154J525 | |
| Person / Time | |
|---|---|
| Site: | Summer  |
| Issue date: | 05/31/1988 |
| From: | SOUTH CAROLINA ELECTRIC & GAS CO. |
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| NUDOCS 8805270031 | |
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{{#Wiki_filter:m i t ! ATTACHHENT 1 SAFETY EVALUATION FOR THE VIRGIL C. SUMMER NUCLEAR STATION TRANSITION TO HESTINGHOUSE 17x17 VANTAGE 5 FUEL SOUTH CAROLINA ELECTRIC & GAS COMPANY VIRGIL C. SUMMER NUCLEAR STATION _. DOCKET No. 50-395 l i MAY, 1988 { G 00681:6/880504 G805270031 DR 880520 ADOCK 05000395 i DCD I
TABLE OF CONTENTS EAge Section 1
1.0 INTRODUCTION
5 2.0 SUHHARY AND CONCLUSIONS 7 3.0 MECHANICAL EVALUATION 15 4.0 N'JCLEAR EVALUATION 5.0 THERMAL AND HYDRAULIC EVALUATION 21 27 6.0 ACCIDENT EVALUATION 41 7.0
SUMMARY
OF TECHNICAL SPECIFICATION CHANGES 52
8.0 REFERENCES
LIST OF TABLES Table No. lLtle E121 Comparison of 17X17 LOPAR and 17X17 VANTAGE 5 Fuel 13 3.1 Assembly Design Parameters . 5.1 V. C. Sumer Thermal and Hydraulic Design Parameters 24 7.1 V. C. Sumer - Sumary Technical Specificat'on 49 Changes for VANTAGE 5 Fuel LIST OF FIGURES ILtle Eige F1aure N L 3.1 17x17 VANTAGE 5 LOPAR Fuel Assembly Comparison 14 i O 1 00681:6/880504
f] V
1.0 INTRODUCTION
The V. C. Summer Power Plant is currently operating in Cycle 4 with a Westinghouse 17x17 low-parasitic (LOPAR) fueled core For subsequent cycles, it is planned to refuel and operate the V. C. Summer Plant with the Hestinghouse VANTAGE 5 improved fuel design defined in Reference 1, except for replacing the VAkTAGE 5 Bottom Nozzle with a Debris Filter Bottom Nozzle (DFBN). As a result, future core loadings would range from approximately 50%-60% LOPAR and 40%-50% VANTAGE 5 transition core (Cycle 5) to eventually an all VANTAGE 5 fueled core. The VANTAGE 5 fuel assembly is designed as a modification to the current 17x17 LOPAR (standard fuel) and the optimized fuel assembly (0FA) designs, Reference 2. The VANTAGE 5 design features were conceptually packaged to be licensed as a
~
single entity. This was accomplished via the NRC review and approval of the "VANTAGE 5 Fuel Assembly Reference Core Report," HCAP-10444-P-A, Reference 1. The initial irradiation of u fuel region containing all the VANTAGE 5 design features occurred in the Callaway Plant during the last quarter of 1987. The N Callaway VANTAGE 5 licensing submittal was made to the NRC in March 31, 1987 (ULNRC-1470, Docket No. 50-483). Several of the VANTAGE 5 design features, such as axial blankets, reconstitutable top nozzles, extended burnup modified _ fuel assemblies and Integral Fuel Burnable Absorbers have been successfully licensed as individual design features and are currently in operating Hestinghouse plants. Also, four VANTAGE 5 demonstration assemblies are currently in a 3rd cycle of irradiation in the V. C. Summer core. Descriptions and evaluations of the DFBN and VANTAGE 5 design features are presented in Section 3.0 of this evaluation report. A brief summary of the VANTAGE 5 design features and its major advantages compared to the LOPAR fuel design are given below. Intearal Fuel Burnable Absorber (IFBA) - The IFBA features a zirconium diboride coating en the fuel pellet surface on the central portion of the enriched 00 2 fuel stack. In a typical reload core, approximately one fourth ( of the fuel rods in the feed region are expected to include IFBAs. IFBAs l provide power peaking and moderator temperature coefficient control. l 00681:6/880504 1
I Intermediate Flow Mixer (IFM) - Three IFM grids located between the four upper Increased h most Zircaloy grids (Figure 3.1) provide increased DNS margin. margin permits an increase in the design basis FAH and Fg . Reconstitutable Too Nozzle - A mechanical disconnect feature facilitates the top nozzle removal. Changes i, the design of both the top and bottom nozzles increase burnup margins by providing additional plenum space and room for fuel rod growth. ELtended Burnuo - The VANTAGE 5 fuel design will be capable of achieving extanded burnups. The basis for designing to extended burnup is contained in the approved Hestinghouse topical HCAP-10125-P-A, Reference 3. Blaatfth - The axial blanket consists of a nominal six inches of natural UO2 pelltets at each end of the fuel stack to reduce neutron leakage and to improve uranium utilization. This report is to serve as a reference safety evaluation / analysis report for the region-by-region reload transition from the present V. C. Summer LOPAR fueled core to an all VANTAGE 5 fueled core. This report examines the differences between the VANTAGE 5 and LOPAR fuel assembly designs and evaluates the effect of these differences on the cores during the transition to an all VANTAGE 5 core. The transition and VANTAGE 5 core evaluation / analyses were performed at a core thermal power. level of 2775 MHt with the following conservative assumptions made in the safety evaluations: a full power FAH of 1.62 for a reload transition core and 1.68 for the VANTAGE 5 fueled core (except for the large Break LOCA analysis which used 1.62 for both transition and full VANTAGE 5 cores), an increase in the maximum F to 2.45, 15% plant uniform steam generator tube plugging *, and a positive q moderator temperature coefficient (PMTC) of +7 pcm/*F from 0 to 70% power and then decreasing linearly to O pcm/*F between 70 to 1007. power.
- Assumes 15% of steam generator tubes in each generator are plugged and corresponds to the worst plugging level of any steam generator.
O 09681:6/880504 2 l l
The axial offset strategy will be the licensed RAOC with Base Loaded Technical - Specifications as an optional strategy during operation at or near steady-state equilibrium conditions. An FqSurveillance Technical Specification will be implemented. RAOC andgF surveillance were approved by the NRC, as shown in WCAP-10217-A (Reference 4). , This report utilizes the standard reload design methods described in Reference 5 and will be used as a basic reference document in support cf future V. C. Sumer Reload Safety Evaluations (RSE) for VANTAGE 5 fuel reloads. Sections 3.0 through 6.0 of the report sumarize the Mechanical, Nuclear, Thermal and Hydraulic, and Accident Evaluations, respectively. Section 7.0 gives a sumary of the Technical Specifications changes needed. Attachments 2 and 3 contain the Technical Specification change pages and non-LOCA safety analyses results, respectively. Attachment 4 contains the large and small break LOCA safety analyses. Attachment 5 contains the Thimble Plug Removal Evaluation. Attachment 6 contains the No Significant Hazards Consideration Justification, and Attachment 7 shows no adverse radiological consequence when using VANTAGE 5 fuel. Consistent with th'e Westinghouse standard reload methodology, Refere,cd 5, parameters are chosen to maximize the applicability of the safety evaluations , for future cycles. The objective of subsequent cycle specific RSEs will be to l verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this report, j In order to demonstrate early performance of the VANTAGE 5 design product features in a commercial reactor, foer VANTAGE 5 demonstration assemblies (17x17) were loaded into the V. C. Sumer Unit 1 Cycle 2 core and began power production in December of 1984. These assemblies completed one cycle of irradiation in October of 1985 with an average burnup of 11,357 MHD/MTV. Post-teradiation examinations showed all 4 demonstration assemblies were of good mechanical Integrity. No mechanical damage or wear was evident on any of the VANTAGE 5 components. Likewise, the IFH grids on the VANTAGE 5 demonstration assemblies had no effect on the adjacent fuel assemblies. All l four demonstration assemblies were reinserted into the V. C. Summer core for a 00681:6/880504 3
second cycle of irradiation. This cycle was completed in March of 1987, at which timo the demonstration assemblies achieved an average burnup of about 30,000 MWD /HTV. The observed behavior of the four demonstration assemblies at the end of 2 cycles of irradiation was as good as that observed at the end of the first cycle of irradiation. The four assemblies were reinserted for a third cycle of irradiation. In addition to V. C. Summer, individual VANTAGE 5 product features have been demonstrated at other nuclear plants. IFBA demonstration fuel rods have been irradiated in Turkey Point Units 3 and 4 for two reactor cycles. Unit 4 contains 112 fuel rods equally distributed in four demonstration assemblies. The IFBA coating performed well with no loss of coating integrity or adherence. The IFH grid feature has been demonstrated at McGuire Unit 1. The demonstration assembly at McGuire has been irradiated for two reactor cycles and is of good mechanical integrity. O O 00681:6/880504 4
2.0
SUMMARY
nND CONCLUSIONS Consistent with the Westinghouse standard reload methodology for analyzing cycle specific reloads, Reference 5, parameters were' selected to conservatively bound the values for each subsecuent reload cycle and to facilitate determination of the applicability of 10CFR50.59. The objective of subsequent cycle specific reload safety evaluations will be to verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this report. The mechanical, thermal and 1 hydraulic, nuclear, and accident evaluations considered the transition core effects described for a VANTAGE 5 mixed core in Reference 1. The summary of these evaluations for the V. C. Sunner core transitions to an all VANTAGE 5 core are given in the following sections of this submittal. The transition design and safety evaluations consider the following conditions: 2775 MHt core thermal power, 552.3'F core inlet temperature, 2250 psia system pressure and 277,800 gpm RCS thermal design flow. These conditions are used in core design and safety evaluations to justify safe operation with the conse'rvative assumptions noted in Section 1.0. The conditions summarized in the SER for the VANTAGE 5 reference core report, ~ HCAP-10444, have been considered in the V. C. Summer plant-specific safety _ evaluations. i The results of evaluation / analysis described herein lead to the following ;. 1 conclusions:
- 1. The Westinghouse VANTAGE 5 reload fuel assemblies for the V. C. Summer '
Nuclear Plant are mechanically compatible with the current LOPAR fuel assemblies, control rods, secondary source rods and reactor internals interfaces. The VANTAGE 5/LOPAR fuel assemblies satisfy the current design bases for the V. C. Summer reactor.
- 2. Evaluations / analyses have shown that all or any combination of thimble plugs may be removed from the Cycle 5 core and subsequent reload cores.
O . 00681:6/880504 5
Changes in the nuclear characteristics due to the transition from LOPAR to 3. VANTAGE 5 fuel will be within the range normally seen from cycle to cycle due to fuel management effects.
- 4. The reload VANTAGE 5 fuel assemblies are hydraulically compatible with previously irradiated LOPAR fuel assemblies.
- 5. The core design and safety analyses results documented in this report show the core's capability for operating safely for the rated V. C. Summer Plant design thermal power with an FAH of 1.62 for mixed fuel cores (eg. Cycle 5) and for all VANTAGE 5 fueled cores,* Fq - 2.45, uniform steam generator tube plugging levels up to 15%, and a positive MTC of +7 pcm/*F from 0 to 707. power and then decreasing linearly to O pcm/*F at 100% power.
- 6. Plant operating limitations given in the Technical Specifications will be satisfied with the proposed changes noted in Section 7.0 of.this report.
A reference is established upon which to base Hestinghouse reload safety evaluations for future reloads with VANTAGE 5 fuel.
- Allows for a number of low power LOPAR assemblies in essentially all VANTAGE 5 cores. The number and maximum power of the LOPAR assemblies will be defined in the cycle specific Reload Safety Evaluation (RSE).
00681:6/880504 6 1
3.0 MECHANICAL EVALUATION This section evaluates the mechanical design and the compatibility of the 17x17 VANTAGE 5 fuel assembly with the current LOPAR' fuel assemblies during the transition through mixed-fuel cores to an all VANTAGE.5 core. The VANTAGE 5 fuel assembly has been designed to be compatible with the LOPAR fuel assemblies, reactor internals interfaces, the fuel handling equipment, and the refueling equipment. The VANTAGE 5 design is intended to replace and be coupatible with fuel cores containing fuel of the LOPAR designs. The VANTAGE 5 design dimensions as shown on Figure 3.1 are essentially equivalant to these designs from an exterior assembly envelope and reactor internals interface standpoint. The design tasis and design limits are essentially the same as those for the LOPAR designs. As such, compliance with the "Acceptance Criteria" of the Standard Review Plan (SRP, NUREG 0800) Section 4.2 Fuel System Design was fully demonstrated. The sigritftcant new mechanical features of the VANTAGE 5 design relative to the current LOPAR fuel design in operation include the following: o Integral Fuel Burnable Absorber (IFBA) o Intermediate Flow Mixer (IFM) Grids o Reconstitutable Top Nozzle o Extended Burnup Capability l o Axial Blankets , o Replacement of six intermediate Inconel grids with Zircaloy grids o Reduction in fuel rod, guide thimble and instrumentation tube diameter Table 3.1 provides a comparison of the VANTAGE 5 and LOPAR fuel assembly design parameters. Another new mechanical design feature is the Debris Filter Bottom Nozzle (DFBN) which is used instead of the bottom nozzle described in the Reference 1 VANTAGE 5 report. O I 00681:6/880504 7
fuel Rod Performance Fuel rod performance for all fuel rod designs is shown to satisfy the SRP fuel rod design bases on a region by region basis. These 'same bases are applicable to all fuel rod designs, including the Westinghouse LOPAR and VANTAGE 5 fuel designs, with the only difference being that the VANTAGE 5 fuel is designed to achieve a higher burnup consistent eith WCAP-10125-P-A, Reference 3, and VANTAGE 5 fuel is designed to operate with a higher F3g limit. The design bases for Hestinghouse VANTAGE 5 fuel are discussed in Reference 1. Yhere is no effect from a fuel rod design standpoint due to having fuel with more than one type of geometry simultaneously residing in the core during the transition cycles. The mechanical fuel rod design evaluation for each region incorporates all appropriate des;gn features of the region, including any changes to the fuel rod or pellet geometry from that of previous fuel regions. The IFBA coated fuel pellets are identical to the enriched uranium dioxide pellets except for the addition of a thin coating on the pellet cylindrical . surface. Coated pellets occupy the central portion of the fuel column. The number and pattern of IFBA rods within an assembly may vary depending on specific application. The ends of the enriched coated pellets and enriched uncoated pellets are dished to allow for axial expansion at the pellet centeriine and vold volume for fission gas release. Analysis of IFBA rods includes any geometry changes necessary to model the presence of burnable absorber, and conservatively models the gas release from the coating. An evaluation and test program for the IFBA design features are given in Section 2.5 in Reference 1. Fuel performance evaluations are completed for each fuel region to demonstrate that the design criteria will be satisfied for all fuel rod types in the core under the planned operating conditions. Any changes from t'.e plant operating conditions originally considered in the mechanical design of a fuel region (for examole, a power uprating or an increase in the peaking factors) are addressed for all affected fuel regions. O l 00681:6/880504 8 l
I Fuel rod design evaluations are currently performed using the NRC approved models in References 6, 7, and 8 to demonstrate that the SRP fuel rod design criteria (including the rod internal pressure design basis in Reference 9) will be satisfied. Grid Assemblies The top and bottom Inconel (non-mixing vane) grids of the VANTAGE 5 fuel assemblies are nearly identical in design to the Inconel grids of the LOPAR fuel assemblies. The only differences are: 1) the grid spring and dimple heights have been modified to accommodate the reduced diameter fuel rod, and
- 2) the grid spring force has been reduced in the top grid. The six intermediate (mixing vane) structural grids are made of Zircaloy material rather than the Inconel used in the LOPAR design, the straps are thicker and the grid height is greater compared to the LOPAR design.
The Intermediate Flow Mixer (IFM) grids shown in Figure 3.1 are located in the three uppermost spans between the Zircaloy mixing vane structural grids and incorporate a similar mixing vane array. Their prime function is mid-span flow mixing in t'he hottest fuel assembly spans. Each IFH grid cell contains four dimp'les which are designed to prevent mid-span channel closure in the _ spans containing IFMs and fuel rod contact with the mixing vanes. This simplified cell arrangement allows short grid cells so that the IFM grid can accomplish its flow mixing objective with minimal pressure drop. The IFM grids are not intended to be structural members. The outer strap configuration was designed similar to current fuel designs to preclude grid hang-up and damage during fuel ha'ndling. Additionally, the grid envelope is smaller which further minimizes the potential for damage and reduces calculated forces during seismic /LOCA events. A coolable geometry is, therefore, assured at the IFM grid elevation, as well as at the structural grid elevation. O 00681:6/880504 9
Reconstitutable Too Nozzle . The reconstitutable top nozzle for the VANTAGE 5 fuel assembly differs from the LOPAR design in two ways: a groove is provided in each thimble thru-hole in the nozzle plate to facilitate attachment and removal; and the nozzle plate thickness is reduced to provide additional axial space for fuel rod growth. To remove the top nozzle, a tool is first inserted through a lock tube and expanded radially to engage the bottom edge of the tube. An axial force is then exerted on the tool which overrides local lock tube deformations and withdraws the lock tube from the insert. After the lock tubes have been withdrawn, the nozzle is removed by raising it off the upper slotted ends of the nozzle inserts which deflect inwardly under the axial lift load. With the top nozzle removed, direct access is provided for fuel rod examina-tions or replacement. Reconstitution is completed by the remounting of the nozzle and the insertion of lock tubes. Additional details of this design feature, the design bases and evaluation of the reconstitutable top nozzle are given in Section 2.3.2 in Reference 1. Oebris Filter Bottom Nozzle . It is planned to introduce the debris filter bottom nozzle (DFBN) into the V. C. Summer Region 7 fuel assemblies to reduce the possibility of fuel rod dtmage due to debris-induced fretting. The relatively large flow holes in a conventional bottom nozzle are replaced with a new pattern of smaller flow holes for the DFBN. The holes are sized to minimize passage of debris particles large enough to cause damage while providing sufficient flow area, comparable pressure crop, and continued structural integrity of the nozzle. Tests to measure pressure drop and demonstrate structural integrity have been performed to verify that the debris filter bottom nozzle is totally compatible with the current design. The 304 stainless steel DFBN is similar to the LOPAR design used for the V. C. Summer Region 6 fuel assemblies. Significant changes compared to the LOPAR l 00681:6/880504 10
design involve 1) a modified flow hole size and pattern as described above, and 2) a decreased nozzle height and thinner top plate (identical to the existing VANTAGE 5 bottom nozzle) to acccanodate the high burnup fuel rods. The DFBN retains the design reconstitution feature which facilitates easy } removal of the nozzle from the fuel assembly in t'he same manner as the bottom nozzle used for the Region 6 LOPAR fuel assemblies. Axial Blankets Although noted as a new mechanical feature of the VANTAGE 5 design and licensed in Reference 1, axial blankets have been and are currently operating in Westinghouse plants. A description and design application of this feature are contained in Reference 1, Section 3.0. g Mechanical Comoatibility of Fuel Assemblies Based on the evaluation of the VANTAGE 5/LOPAR design differences and hydraulic test results (Reference 1) and the evaluation of the DFBN, it is concluded that the two designs are mechanically compatible with each other. The VANTAGE 5 fuel rod mechanical design bases remain unchanged from that used l
- for the LOPAR Region 6 fuel assemblies in the V. C. Summer Cycle 4 core.
Rod Bow It is predicted that the 17x17 Vantage 5 rod bow magnitudes, like those of the j Westinghouse OFA fuel, will be within the bounds of existing 17x17 LOPAR assembly rod bow data. The current NRC approved methodology for compar.ing rod bow for two different fuel assembly designs is given in Reference 10. Rod bow in fuel rods containing IFBAs is not expected to differ in magnitude or frequency from that currently observed in Westinghouse LOPAR fuel . rods - under similar operating conditions. No indications of abnormal rod bow have j been observed on visual or dimensional inspections performed on the test IFBA rods. Rod growth measurements were also within predicted bounds. l O 00681:6/880504 11
Fuel Rod Heae Fuel rod wear is dependent on both the support conditions and the flow environment to which the fuel rod is subject 6d. Due'to the LOPAR and VANTAGE 5 fuel assembly designs employing different grids. there is an unequal axla1 pressure distribution between the assemblies. Crossflow resulting from this unequal pressure distribution was evaluated to determine the induced rod vibration and subsequent wear. Hydraulic tests, (Reference 1, Appendix A.l.4) were performed to verify that the crossflows were negligible and also to check hydraulic compatibility of the LOPAR and VANTAGE 5 designs. The VANTAGE 5 ft'el assembly was flow tested adjacent to a 17x17 0FA, since vibration test results indicated that the crossflow effects produced by this fuel assembly combination would have the most detrimental effect on fuel rod wear. Results of the wear inspection and analysis discussed in Reference 1, Appendix A.I.4, revealed that the VANTAGE 5 fuel assembly wear characteristic was similar to that of the 17x17 0FA when both sets of data were normalized to the test duration time. It was concluded that the VANTAGE 5 fuel rod wear would g be less than the maximum wear depth established, Reference 11. for the 17x17 0FA at EOL. O 00681:6/880504 12
_ _ . . _ _ . _ m l l TABLE 3.1 . j COMPARISON OF 17x17 LOPAR '!j and . 17x17 VANTAGE 5 FUEL ASSEMBLY DESIGN PARAMETERS , i< , 17x17 17x17 I l' PARAMETER LOPAR DESIGN VANTAGF 5 DESIGN j' t Fue ' w Length, in 1F9.765 159.975 Fue i 6: Length, in 151.56 152.255 ll Assembly Envelope, in 8.426 8.426 Compatible with Core Internals Yes Yes )* Fuel Rod Pitch, in .496 .496 Number of Fuel Rods /Assy 264 264 ' Number / Guide Thimble Tubes /Asg 24 24 Number / Instrumentation Tube /Assy 1, 1
; Fuel Tube Material Zircaloy 4 Zircaloy 4 d fuel Rod Clad 00, in 0.374 0.360 Fuel Rod Clad Thickness, in .0225 .0225 Fuel / Clad Gap, mil 6.5 6.2 Fuel Pellet 01ameter, in .3225 .3088 -
l Fuel Pellet Langth, in .530 .507 Guide Thimble Material Zircaloy 4 7.trealoy 4 j Guide Thimble 00, in. .482 .474 Instrumentation Tube Material 71rcaloy 4 Zircaloy 4 ! Instrumentation Tube 00, in. .482 .474 lO 00681:6/880504 13 l
~ _ . ~ _ - - _ . _ _ _ _ . - - . - - _ _ _ _ _ _ . _ - _ _ _ _ _ _ - . - - - -
4 109.975 1 2.383 + *- h 3.475 152.25 m ( ),
=
8.426 0 8.426 o 8.372 0 8.418 . u r / i E Mi M6 E 2T *- 7L _ _ -__ ~ ~. 112 92 92 37 71 62 51 27 30 72 5 84 133 47 I7Xl7 VANTAGE 5 FUEL ASSEMBLY 159.765 = 4 2.738 + *- 4 = 3.670 151.56 =
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112 82 92 27 71 72 Si 17 30 62 6 190 153 96 133 37 l 17XI7 RECONSTITUTABLE LOPAR FUEL ASSEMBLY l l V.C. SUMMER UNIT I FIGURE 3.1 l 14 17XI7 VANTAGE 5 / LOPAR l FUEL ASSEMBLY COMPARISON
(m ( ,) 4.0 NUCLEAR DESIGN The evaluation of the transition and equilibrium cycle VANTAGE 5 cores presented in Reference 1, as well as the V. C. Sumer specific transition and equilibrium core evaluations, demonstrate that the impact of implementing { VANTAGE 5 does not cause a significant change to the physics characteristics , of the V. C. Sumer cores beyond the normal range of variations seen from cycle to cycle. The methods and core models used in the V. C. Summer reload transition core analysis are described in References 1, 5, 12, and 13. These licensed methods and models have been used for V. C. Summer and other previous Westinghouse reload designs using the LOPAR and VANTAGE 5 fuel. No changes to the nuclear design philosophy, methods, or models are necessary because of the transition to VANTAGE 5 fuel. Increased emphasis will be placed on the use of three-dimensional nuclear models because of the axially heterogeneous nature of the VANTAGE 5 fuel design when axia1 blankets and reduced length HABA/IFBAs are used, i From the nuclear design area, the following V. C. Summer Technical Speclfi-cation changes are proposed: !
- 1) Increased F AH limits. These higher limits will allow loading pattern designs with lower leakage which in turn will allow longer j cycles. l l
- 2) Increased F limit. The increased F limit will provide greater 9
9 flexibility with regard to accommodating the axially heterogeneous cores (blankets and short burnable absorbers).
- 3) F gSurveillance. This revision to surveillance requirements on the heat flux hot channel factor, F (z), has been proposed to increase 9
plant operating flexibility while more directly monitoring Fg . Rather than performing surveillance on Fxy(2), the radial component p of the total peaking factor, surveillance is pe,-formed directly on U F (z). 9 Th'e steady-state F (z) is measured and increased by 9 00681:6/880504 15
applicable uncertainties. This qinntity is further increased by an analytica.1 factor called H(z) which accounts for possible increases in the steady-state Fg(z) resulting from operation within the allowed axial flux difference limits. The r'esulting Fg (z) is compared to tne F (z) limit to demonstrate operation below the heat 9 flux hot channel factor limit.
- 4) RAOC/ Base Load implementation. The RAOC strategy allows greater operator flexibility with regard to core operation. The margin created by the increased Fg limit is being partly converted into operational flexibility.
During operation at or near steady state equilibrium conditions, core peaking factors are significantly reduced. Through the use of a Base Load Tech Spec, this reduction in core peaking factors are recognized. As illustrated in Attachment A to this section, the AFD operating spaces may be presented as a function of cycle burnup to further enhance operational flexibility during portions of the cycle. This is accomplished by performing an analysis consistent with the methodology described in Reference 4, which' takes credit for _
"burndown" characteristics of both radial and axial power shapes.
It is proposed that the resulting AFD operating spaces for RAOC and Base Load operations be deleted from the Tech Spec and instead, be incorporated in the Peaking Factor Limit Report (PFLR). This elircinates the potential necessity of Tech Spec amendments or AFD limits for future reload cycles, while providing adequate assurance ~ that the correct AFD operating spaces will be followed. A sample PFLR containing AFD operating limits and W(Z) values can be found in the Attachment A. O 00681:6/880504 16
)
l
- 5) Positive MTC. Due to increased boron concentration associated with a higher enriched fuel and longer fuel residence times, it is necessary to increase the MTC limit. The increased limit value is +7.0 pcm/*F from HZP to 70% rated power, then a negative' linear ramp from 70%
power to 0.0 pcm/'F at HFP. _ Power distributions and peaking factors show slight changes as a result of the incorporation of axial blankets, reduced length HABA/IFBAs, and increased ^ peaking factor limits, in addition to the normal variations experienced with different loading patterns. The usual methods of enrichment variation and burnable absorber usage can be employed in the transition and full VANTAGE 5 cores to ensure compliance with the peaking factor Techniul Specifications. The key safety parameters evaluated for V. C. Summer reactor as it transitions to an all VANTAGE 5 core show little change relative to the range o) parameters experienced for the all LOPAR core. The changes in values of the key safety parameters are typical of the normal cycle-to-cycle variations experienced as loading patterns change. As is current pract!ce, each reload core design will be eva10ated to assure that design and safety limits are j satisfied according to the reload methodology. The design and safety limits , will be documented in each cycle specific reload safety evaluation (RSE) _; report which serves as a basis fcr any significant changes which may require a { future NRC review. I O 00681:6/880504 17
l l l Attachment A to Section 4.0 SAMPLE PEAKING FACTOR REPORT FOR V. C. SUMMER PLANT _ l l l l l O. i O 00681:6/880504 18
i
/ PEAKING FACTOR LIMIT REPORT FOR V. C. SUMMER CYCLE 5 RA0C AND BASE LOAD OPERATION
[ This Peaking Factor Limit Report is provided in accor' dance with Paragraph 6.9.1.11 of the V. C. Sumer Technical Specifications. _ The V. C. Sumer Cycle 5 allowed Axial Flux Difference (AFD) operating space for RAOC operation for beginning-of-life (BOL), middle-of-life (MOL), and end-of-life (E0L) portions are shown in Figures 1 through 3, respectively. The cycle burnup ranges applicable to each specified operating space are indicated in each of the figures. The specified allowed AFD operational spaces for RAOC operation were determined using the method described in Reference 1. The V. C. Sumer Cycle 5 elevation dependent N(z) values for RAOC operation at BOL, MOL, and near E0L are shown in Figures 4 through 6, respectively. This information is sufficient to determine H(z) versus core height for Cycle 5 burnups in the range of 0 MHD/MTU to EOL burnup through the use of three point interpolation. The V. C. Sumer Cycle 5 allowed AFD target bands during base load operation _ for BOL, MOL, and EOL were determined to be as follows: t
- + or - x % about a measured target value I' BOL (0 - xxxx MHD/HTU)
HOL (xxxx - xxxx MHD/HTV) : + or - x % about a measured target value EOL (xxxx - xxxxx MHD/MTU) : + or - x % about a measured target value The V. C. Sumer Cycle 5 elevation dependent H(z) values for base load operation between x% and 100% of rated thermal power with the above specified target band about a measured target value at BOL, MOL, and near EOL are shown in Figures 7 through 9, respectively. This information is sufficient to determine H(2) versus core height for Cycle 5 burnups in the range of 0 MHD/MTu to EOL burnup through the use of three point interpolation. H(z) values for RAOC and base load operation were calculated using the method described in Part B of Reference 1. 00681:6/880504 19
NO ' The minimum allowable power level for base load operation, APL , for V. C. Summer Cycle 5 is xx percent of rated thermal power. The appropriate W(z) function is used to confirm that the heat flux hot channel factor, FQ(z), will be limited to the Technical Specification values Fq(z) 1 2.45 / P (k(z)) for P > 0.50 and Fq(z) 1 4.90 (k(2)) for P 1 0.50 The appropriate elevation dependent W(z) values, when applied to a power distribution measured under equilibrium conditons, demonstrate that the initial conditions assumed in the LOCA are met, along with the ECCS acceptance criteria of 10CFR50.46. (1) WCAP-10215-P-A, Relaxation of Constant Axial Control - Fq Surveillance Technical Specification O I l l l 9 00681:6/880504 20
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FIGJRE 2 l AXIAL FIl3X DIFTERDCE LD1ITS AS A FUCTICN & I pAUD DIEWAL PWER FOR CYCII BUmmP 4000 -10000 MO/)fIU 1 1
1 i . SAMPLE n v 120 110 il (-1 ;,190) ( + 7,100) : ifi .
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i AXIAL FLUX DIFFERDCE LIMITS AS A FUCTION OF e ragaga, pcwER FOR cycle Bum;UP 10000 Mo/MIV - IDL l
SAMPLE l-go WIeff (FT.) DOL W(2) 0.0 1.0000
- 9.3 1.0000 .
u e 0.4 1.0000
- 9.5 1.0C00
- 0.8 1.0000
. e 1.0 1.000o
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- 1.8 1.0000 2.0 1. H01
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" 5.8 1.1880
# d 5.8 1.17S2 0.0 1.1750 O2 1.1773
, 0.4 1.1836 LN 8.0 1.1884
= e.. '*
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e 10,8
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V. C. SOFIR BAOC W(Z) AT BOL (150 MWD /MTU) . ice aw eor 7w in, ocwoeo as rea 7ecs spec 4222 o h j
1 SAMPLE (FT.) W(2) I o 0.0 1.0000
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- 11.s 1.0000 t
e 11.8 1.0000 1.0000 l e 12.0 i V. C. Su2GR R/4C W(Z) AT M3L __ (8000 MWD /MTV) l TOP AND BOTTOM 15*. E.xCLUDED AS PER TECH EPEC 4122 G 4 l l l l- _ ., . --
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SAMPLE . e m0.0 Os 1.0000 , o 0.2 1.0000 12 e 0.4 1.0000 e 0.9 1.0000
*" e C.8 1.0000 e
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, g E.8 1.2182 1
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t.> 3.8 1.2130 e *** "
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8.8 1.1204 I I f
, __.__. 4.0 1.1270 1 1.2318
- 4.2 l 4.4 1.2847 .
- a. 4.5 1.2285 I
- _ 4.8 1.2225 1
^ 5.0 1.2312 l N 3.2 1,2331
- 8.4 1.2So1
" 5.8 1.2961 tu 5.8 1.2850
-= 6.0 1.9031 S.2 1.3183 e.4 1.2300 i 5.5 1.3282 0.8 1,8425 to 7.0 1.S427
.. . . .._ l 7.2 1.2348 1
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-- 9.8 1.2232 10.0 1.2410 tu, , , e a e a 10.2 1.2488 BOTTOM CORE EIGHT (FEET) 70p l llj
= 10.8 iC 1.oooo e 11.o 1.00o0 e 11.2 1.0000 e 11.4 1.0000
= 11.8 1.oooo FIGUPI 6 e 11.8 1.oooo a 12.0 1.oooo V. C. SLMKER RACC W(Z) hTM EOL (14000 MWD /MTV)
TOP AND BOTTOM 15% EXCLUDED AS PER TECH SPtC 4212 G O 1
SAMPLE ma e 0.0 1.000o O t.n e 0.3 e e 0.4 0.8 1.0000 1.0000 1.000o e C.8 1.0000 e 1.0 1.0000 e 1.3 1.00c0 e 1.4 1.0000
, e 1.8 1.000o e 1,a 1.0000 3.0 1.0064 S.S 1.00lJ a 3.4 1.0848 i'
3.5 1.0038 t.s e 2.8 1.0821
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- S.8 1.0854 e 0.8 1.0838
'
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4 1 e go,4 g,cooc COPI K lGHT FEE 1 TOP e 10.8 1. Coco BOTTOM e 1o.8 1.cooo e 11.0 1.000o i e 11.2 1. coco e 11.4 1.o000 11.8 1.000o e
- 11.8 1.0000 l e 12.0 1.0000 7
V. C. SuthER B7.SEUDAD W(Z) FDR Foss BEIkm 80% AND 100% OF ,, pm im7d, IOG2 WITHDJ X % AFD OF THE MEASURED TAPGET AT BOL (150 MWD /MTU) j Top AND BOTTOM 15** EXCLUDED AS PER TECH SPEC 4.2.2 4 G
SAMPLE
~
o a ma0.0 1.0000 1.0000 0.3 ' km
- 0.4 1.0000
. e 0.8 1.0000 e 0.8 1.000o e 1.0 1.0000 e 1.2 1.0000 to e 1.4 1,g000 e 1.0 1.0000 e 1.8 1.0000 1.112S 3.0 3.3 tu g,4 1.10,88 g,q 43 S.S 1.0898 S.8 1.0643 3.0 1.0888 S.2 1.0836
*" 8.4 1.0785 S.S 1.0774 3.8 1.07S4 4.0 1.0754 4.2 1.0742
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^ -
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; S.2 1.0060
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- t
- S.8 1.0798 I , 7.0 1. 04 M tm I
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@v c* sucER BASamD W(Z) FOR pm EE%'Em 00% AND 10 (8000 MWD /MTU)
O lop AND bot 10W 15% ErCluoED As PER TECH $P(C 4 2 2AG
SAMPLE
~
- a
. (FT ) .If.( 2_)
(s~3 - . ....., . .. e 0.0 1.0000
- 9.8 1.0000
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, g,4 g,ggon
- 1.9 1.9000 e 1.8 1.0000 3.0 1.1854 S.3 1.1273
** 3.4 1.1178 S.5 1.1000 S.8 1.0975 S.0 1.0870
,,, 8.3 1.9811 gg 8.4 1.0828
* ^
S.S 1.0844
*
- 8.8 1.0851 4.0 1.0057 to 4.2 1.0050
, 4.4 1.0859 4
4.0 1.0857 4.8 1.0034
^ O.0 1.0843 N
*g,,
S.2 1.0013
- S.4 1.1006
" , S.S 1.1064 S.S 1.1172 1.1237 t.m
^/*""*r S.0 S.E 1.1200 S.4 1.1327 S.S 1.1350 S.S 1.1357 O t.m 7.0 7.2 7.4 7.S 1.1950 1.1327 1.1282 1.1243 7.8 1.1104 S.0 1.1100 t.m S.2 1.1086 _;
8.4 1.1114 S.S 1.1134 ' S.8 1.1143 !. S.0 1.1145 i tc i ~~ S.2 1.1142 i i S.4 1.1136 1 f S.S 1.1130 ' l S.8 1.1130
, i ' 10.C 1.1141 t#
4 e 4 e g 10.2 1.1147 8 8 e 10.4 1.0000 CORE KlGHT (FEET) Top e 20.5 i.0000 BOTTOM e 10.8 1.0000 e 11.0 1,0000 e 11.2 1. _ 0 e 11.4 1.0000 e 11.8 1.0000 e 11.8 1,0000 FIQ'RE 9 e 12.0 1.0000 l V. C. SUfER BASELCAD W(Z) FOR FOCRS BEIkIIN 80% RD 100% OF l RAn:o Ucan rocR WImm X % Am OF DE MEASURID GPET hTAR DOL l (14000 MWD /MTU) O l 10P AND DOTTOW 15% DCtVDED AS PER TECN SPfC 422,4 G t i l i
5.0 THERMAL AND HYDRAULIC DESIGN The analysis of the LOPAR and VANTAGE 5 fuel is based on the Improved Thermal Design Procedure (ITDP) described in Reference 14. T'he LOPAR fuel analysis uses the HRB-1 DNB correlation in Reference 15 while the VANTAGE 5 fuel utilizes the WRB-2 DNB correlation in Reference 1. These DNB correlations take credit for the significant improvement in the accuracy of the critical heat flux predictions over previous DNB correlations. The HRB-2 DNB correlation also takes credit for the VANTAGE 5 fuel assembly mixing vane design. A DNBR limit of 1.17 is applicable for both the HRB-1 and HRB-2 correlations. In addition, The H-3 DNBR correlation is used where appropriate (e.g., accidents analyzed in Sections 15.2.1 and 15.4.2 of Attachment 3). Table 5.1 summarizes the pertinent thermal and hydraulic design parameters. The design method employed to meet the DNB design basis is the Improved Thermal Design Procedures which has been approved by the NRC, Reference 16. Uncertainties in plant operating parameters, nuclear, and thermal parameters, and fuel fabrication parameters are considered statistically such that there is at least 95 percent probability at a 95 percent confidence level that the minimum DNBR will be gre.ater than or equal to 1.17 for the limiting power rod. Plant parameter uncertainties are used to determine the plant DNBR _ uncertainties. Those DNBR uncertainties, combined with the DNBR limit, establish a DNBR value which must be met in plant safety analyses. Since.the parameter uncertainties are considered in determining the design DNBR value, the plant safety analyses are performed using values of input parameters without uncertainties. For this application, the minimum required DNBR values ' for tile LOPAR fuel analysis are a 1.35 for thimble cold wall cells (three fuel rods and a thimble tube) and 1.36 for typical cell (four fuel rods). The design DNBR values for the VANTAGE 5 fuel are a 1.32 and a 1.33 for thimble and typical cells, respectively. In addition to the above considerations, a plant-specific DNBR margin has been considered in the analyses. In particular, safety analysis DNBR limits of ! 1.44 for thimble and 1.48 for typical cells for LOPAR fuel, and 1.60 and 1.68 for thimble and typical cells respectively for the VANTAGE 5 fuel, were employed in the safety analyses. The DNBR margin between the DNBRs used in l 00681:6/880504 21 i
l. i l O the safety analyses and the design DNBR values is broken down as follows. A fraction of the margin is utilized to accommodate the transition core penalty (12.5% for VANTAGE 5 fuel and none for LOPAR fuel) and the appropriate fuel rod bow DNBR penalty, Reference 10, which is less.than 1.3%. The existing 6.3% margin in the LOPAR fuel and 17.5% margin in the VANTAGE 5 fuel between the design and safety analysis DNBR limits also includes a greater than 5% DNBR margin in the LOPAR fuel and a greater than 3.7% DNBR margin in the VANTAGE 5 fuel reserved for flexibility in the design. , The LOPAR and VANTAGE 5 designs have been shown to be hydraulically compatible in Reference 1. The major impact of thimble plug removal on the thermal-hydraulic analysis is the increase in bypass flow which is reflected in Table 5.1. The phenomena of fuel rod bowing, as described in Reference 10, must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Internal to the fuel rod, the IFBA and fuel pellet designs are not expected to increase the propensity for fuel rods to bow. Ext?rnal to the VANTAGE 5 fuel rod, the Inconel non-mixing vane and Zircaloy mixing vane grids provide fuel rod support. Additional restraint is _, I' provided with the Intermediate Flow Hixer (IFH) grids. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNBR are used to offset the effect of rod bow. The safety analysis for the V. C. Summer Plants maintain sufficient margin between the safety analysis limit DNBRs and the design limit DNBRs to accommodate full-flow and low-flow DNBR penalties. The Hestinghouse transition core DNB methodology is given in References 2 and 17 and has been approved by the NRC via Ref'erence 18. Using this methodology, transition cores are analyzed as if they were full cores of one assembly type (full LOPAR or full VANTAGE 5), applying the applicable transition core penalties. This penalty is included in the safety analysis limit DNBRs such that sufficient margin over the design limit DNBR eY'sts to accommodate the transition core penalty and the appropriate rod bow DNBR penalty. 00681:6/880504 22 r
The fuel temperatures for use in safety analysis calculations for the VANTAGE 5 fuel are evaluated using the same methods as those used to evaluate the LOPAR fuel. Hestinghouse uses the PAD perfor. nance code described in Reference 6 to perform both design and licensing calc'ulations. When the code is used to calculate fuel temperatures to be used as initi.al conditions in safety analyses, a conservative thermal safety model, Reference 7, is used. O I l O 00681:6/880504 23 l l I
i 1 f 'j ~ V TABLE 5.1 i: V. C. SUMMER THERHAL AND HYDRAULIC DESIGN PARAMETERS l-Design }: Thermal and Hydraulic Desian Parameters Parameters l>- (Using ITDP) , Reactor Core Heat Output, HHt 2775 Reactor Core Heat Output, 106BTV/Hr 9469 Heat Generated in Fuel, % 97.4 Core Pressure, Nominal, psia 2280 [ Radial Power Distribution (LU;AR) 1.56 (1+0.3(1-P)P
^
(V-5) 1.62 (1+0.3(1-P)]* Limit DNBR for Design Transients Typical Flow Channel (LOPAR) 1,48 (V-5) 1.68 Thimble (Cold Hall) Flow Channel (LOPAR) 1.44 (V-5) 1.60 DNB Correlation (LOPAR) HRB-1 (V-5) HRB-2
- The 4% radial power uncertainty has been removed for statistical combination with other uncertainties in the ITDP analysis.
00681:6/880504 24
TABLE 5.1 (Continued) V. C. SUMMER THERMAL AND HYDRAULIC DESIGil PARAMETERS Design. HFP Nominal Coolant Conditions Parameters Vessel Minimum Measured Flow + Rate (including Bypass), 100 lbm/hr 106.2 GPM 283,500 g Vessel Thermal Design Flow + Rate (including Bypa.:e), 100 lbm/hr 104.1 GPM 277,800 Core Flow Rate (excluding Bypass, basec on TDF) 106 lbm/hr 94.8 GPM 253,080 Fuel Assembly Flow Area ++ 2 for Heat Transfer, ft (LOPAR) 41.55 (V-5) 44.04 Core Inlet Mass Velocity, 100 lbm/hr-ft (Based on TOF) (LOPAR) 2.28 (V-5) 2.15
+ Includes 157. steam generator tube plugging
++ Assumes all LOPAR or VANTAGE 5 core 0
00681:6/880504 25
-k/ TABLE 5.1 (Continued) V. C. SUMMER THERHAL AND HYDRAULIC DESIGN PARAMETERS Design ', Thermal and Hydraulic Desian Parameters Parameters t (Based on Thermal Design Flow) Nominal Vessel / Core Inlet Temperature, 'F 552.3 Vessel Average Temperature, 'F 585.5 Core Average Temperature, 'F 590.5 Vessel Outlet Temperature, 'F 618.7 Average Temperature Rise in Vessel, 'F 66.4 Average Temperature Rise in Core, 'F 72.0 f O ! Heat Transfer j l Active Heat Transfer Surface Area," (LOPAR) 43,598 [ 2 46,779 2 ft (V-5) 2 - 189,820 Average Heat Flux, BTU /hr-ft (LOPAR) (V-5) 197,200 Average Linear Power, kw/ft 5.45 ! i
+++
Peak Linear Power for Normal Operation, kw/ft 13.30 o Assumes all LOPAR or VANTAGE 5 core m Based on 2.45 F qpeaking factor 00681:6/880504 26
6.0 ACCIDENT EVALUATION g, 6.1 Non-LOCA Accidents This section addresses the impact of the VANTAGE 5 design features and modified safety analysis assumptions for the V. C. Summer Plant non-LOCA accident analyses. 6.1.1 VANTAGE 5 Design Features The design features of VANTAGE 5 fuel, considered in the non-LOCA analysis are: f
- Fuel Rod Dimensions
- Axial Blankets
- Integral Fuel Burnable Absorbers (IFBAs) and Wet Annular Burnable Absorbers (WABA)
Intermediate Flow Hixer Grids (IFHs)
- Reconstitutable Top Nozzle
- Fuel Enrichment
- Extended Burnup Fuel Assembly Design l f
A brief description of each of these and its consideration in the safety _j analyses follows. l Fuel Rod Dimensions I The VANTAGE 5 fuel rod dimensions which determine the safety analysis temperature versus linear power density relationship include' rod diameter, j pellet diameter, initial pellet-to-clad gap size, and stack height. The ' non-LOCA safety analysis fuel temperature and rod geometry assumptions consider this geometry change and bound both LOPAR (Standard) and VANTAGE 5 fuel. O 1406v:10/880517
3 Axial Blankets and IFBAs p) Axial blankets reduce power at the ends of the rod which increases axial peaking at the interior of the rod. Used alone, axial blankets reduce DNB margin, but the effect may be offset by the presence'of part length Integral Fuel Burnable Absorbers (IFBAs) which flatten the power distribution. The not effect on the axial shape is a function of the number and configuration of IFBAs in the core and time in life. The effects of axial blankets and IFBAs on the reload safety analysis parameters are taken into account in the reload design process. The axial power distribution assumption in the safety analyses kinetics calculations have been determined to be applicable for evaluating the introduction of axial blankets in the V. C. Summer plant. IFH Grids and Reconstitutable Too Nozzle The IFM grid feature of the VANTAGE 5 fuel desica increases ONB margin. The fuel safety analysis limit DNBR values contair, significant DNB margin (see Section 4.0). This ONB margin was' set to en;ure that the core thermal safety limits for the VANTAGE 5 fuel with an F gg cf l'.68 are acceptable. However, for the transition cycles the LOPAR fuel core limits with an FHH of 1,62 are more restrictive than the VANTAGE 5 fuel core limits. Thus the most restrictive core limits correspond to the LOPAR fuel design. Any _ transition core penalty is accounted for with the available DNBR margin. The IFM grid feature of the VANTAGE 5 fuel design increases the core pressure drop. The control rod scram time to the dashpot is increased from 2.3 to 2.7 seconds. The increased drop time primarily affects the fast reactivity transients. These accidents we're reanalyzed for this report. The revised safety anal,. tis assumption was incorporated in all the reanalyzed events requiring this parameter and the remaining transients have been evaluated. Core flow areas and loss coefficients were preserved in the design of the reconstitutable top nozzle. As such, no parameters important to non-LOCA safety analyses are impacted. O 28 1406v:lD/880517
Fuel Enrichment The VANTAGE 5 fuel design increased fuel enric;-ant is conservatively bounded by the maximum safety analysis assumptions. Extended Burnuo Fuel Assembly Desian _ HCAP-10125-P-A, "Extended Burnup Evaluation of Westinghouse Fuel," evaluates the impact of extended burnup co the design and operation of West;nghouse fuel. The major effect of the extended burnup rod design is on power shaping between fresh and burned assemblies. 6.1.2 Modified Safety Analysis Assumptions Listed below are the analysis assumptions which represent a departure from that currently used for V. C. Summer.
- Changes in Moderator Temperature Coefficient (Most Positive, Most Negative)
Increased Design Enthalpy Rise Hot Channel Factor (Fgq) and Fg
- Increase in Allowable Steam Generator Tube Plugging level
- Reactor Coolant System Flow Raduction _
- Thimble Plug Deletion Oebris Filter Bottom Nozzle
- Increased Overpower /0vertemperature HT Reactor Trip Response Time
- Improved Thermal Design Procedure l
A brief description'of each of tnese assumptions follows. Qtances in Moderator Temoerature Coefficient (Most Positivo. Most Negative) A positive moderator temperature coefficient (M'TC) of +7 pcm/ degree F from 07. } to 707. power and decreasing linearly to O pcm/ degree F at 1007. power was incorporated into the safety analyses performed for this report. O 1406v:lD/880517 29
In general, the analyses presented are based on a +7 pcm/ degree F moderator , b temperature coefficient, which is assumed to remain' constant for variations in temperature. Exceptions are rod ejection and rod withdrawal from subcritical which are based on a MTC of +7 pcm/ degree F at zero power nominal average temperature and which, due to moderator temperature f'eedback modeled in the TWINKLE diffusion-theory code, becomes less positive for Mgher temperatures. Incorporation of the described lev 11 of PMTC into the safety analyses is, in all cases, a conservalve assumption for this report. In order to accommodate longer fuel cycles and extended fuel burnup, a negative moderator temperature coefficient of -50 pcm/ degree F corresponding to end of life, full power conditions was conservatively incorporated into the safety analyses performed for this report. Increased Design Enthalpy Rise Peaking Factor (FAH) and Fg The F AH for the LOPAR and Vantage 5 fuel during the transition cycles is 1.62. The non-LOCA calculations applicable for the VANTAGE 5 core have , assumed a full power Fah of 1.68. This is a conservative safety analysts assumption for this report. . The design core limits for this report incorporate the increased FAH for both the LOPAR and VANTAGE 5 fuel. The inc < -e in the Technical Specification maximum LOCA F from 2.25 to 9 ' 2.45 for both LOPAR and VANTAGE 5 fuel is conservatively accounted for in the non-LOCA transients. l-Increased Steam Generator Tube Pluaaina All non-LOCA safety analyses reenalyzed for this report have incorporated up to a maximum of 15% plant total steam generator tube plugging. It is assumed t, sat no one steam generator exceeds 15% tube plugging. O 1406v:lD/880517 30
Reactor Coolant System Flow Reduction All non-LOCA safety analyses reanalyzed for this report have incorporated a reduction in the reactor coolant system flow. The reduced flow corresponds to a thermal design flow of 92600 gpm/ loop, and a minimu'm measured flow of 94500 gpm/ loop. _ Thimble Plua Deletion The non-LOCA analyses performed incorporated the impact of thimble plug deletion. Thimble plug deletion affects core pressure drops and bypass flow. These effects have been conservatively incorporated into the non-LOCA safety analyses. Debris Filter Bottom Nozzle The VANTAGE 5 fuel design will also include the Debris Filter Bottom Nozzle (DF8N). In the DFBN, the relatively large flow holes in the conventional bottom nozzle are replaced with'a new pattern of smaller flow holes. These
~
holes are sized to minimize the passage of debris particles large enough to cause damage while still providing sufficient flow area, comparable pressure drop, and continued structural integrity of the nozzle'. As such, no _ parameters important to the non-LOCA safety analyses are impacted. Increased Overoower/0vertemoerature AT Reactor Trio Resoonse Time The total time delay of the overtemperature AT and overpower AT trips (including RTD time response, trip circuitry and channel electronics delay) assumed in the non-LOCA analyses is 8.5 seconds. The 8.5 second delay includes a 7 second first order lag incorporated into the determination of the time at which the overiemperature AT and overpower AT trip setpoints are reached. The remaining 1.5 seconds is the delay from the time at which the trip signal is initiated until the rod sluster control assemblies are free to drop into the core. O 31 1406v:lD/880517
Imoroved Thermal Desian Proceduret /m V a The calculational method utilized to meet the ONB design basis is the ITOP, discussed in Reference.14. Uncertainties in plant operating parameters are statistically treated such that there is at least.a 95 percent probability at a 95 percent confidence level that the minimum DNBR will be greater than 1.17. Since the parameter uncertainties are considered in determining the design DNBR value, the plant safety analyses are performed using nominal input j parameters without uncertainties. ,' The LOPAR fuel DNB analyses use the HRB-1 correlation, while the VANTAGE 5 fuel analyse! use the HRB-2 correlation. ., , 6.1.3 Non-LOCA Safety Evaluation Methodology The non-LOCA safety evaluation process is described in References 1 and 2. The process determines if a core configuration is bounded by existing safety , analyses in order to confirm that applicable safety criteria are satisfied. i The methodology systematically identifies parameter changes on a cycle-by-cycle basis which may invalidate existing safety analysis assumption's and identifies the transients which require reevaluation. This methodology is applicable to the evaluation of VANTAGE 5 transition and full cores. - Any required reevaluation identified by the reload methodology is one of two types. .If the identified parameter is only slightly out of bounds, or the transient is relatively insensitive to that parameter, a simple evaluation may be made which conservatively evaluates the magnitude of the effect and explains why the actual analysis of the event does not have to be repeated. Alternatively, should the deviation be large and/or expected to have a significantly or not easily quantifiable effect on the transients, reanalyses ' are required. The reanalysis approach will typically utilize the analytical methods which have been used in previous submittels to the NRC. These methods are t'ose h which have been presented in F5ARs. 3cbsequent submittals to the NRC for a specific plant, reference SARs, or report 'submittals for NRC approval. O 1406v:10/880517 32
The ke'y safety parameters are documented in Reference 5. Values of these safety parameters which bound both fuel types (LOPAR and VANTAGE 5) were assumed in the safety analyses. For subsequent fuel reloads, the key safety parameters will be evaluated to determine if vloiations of these bounding values exist. Reevaluation of the affected transients would take place and would be documented for the cycle-specific reload design, as per Reference 5, 6.1.4 Conclusions Descriptions of the transients reanalyzed for this report, method of analysis, results, and conclusions are contained in Attachment 3. The analytical procedures and computer codes used are identified in Section 15.1. Attachment 3 has been prepared conforming to the format of the V. C. Summer FSAR. For each cf the accidents reanalyzed, it was found that the appropriate safety criteria are met. In addition, evaluations have been performed regarding the impact of VANTAGE 5 fuel on the steam line break mass and energy release analyses, both inside and outside containment. The results of th'is evaluation verify that the mass and energy releases previously calculated, are not adversely impacted by the transition to VANTAGE 5 fuel. 6.2 LOCA Accidents _ 6.2.1 Large Break LOCA 6.2.1.1 Description of Analysis / Assumptions for 17x17 VANTAGE 5 Fuel l Consistent with the methodology developed in the VANTAGE 5 Reference Core Report (Reference 1), a large break loss-of-coolant accident (LOCA) analysis based on a full VANTAGE 5 core has been performed to define peaking factor limits for use during and subsequent to the transition to VANTAGE 5 fuel at V. C. Summer. The Westinghouse 1981 Evaluation Model with BASH (Reference 19) was utill:ed for a spectrum of c.dd leg bieH - Ley assumptions incluce: 1 0 1406v:lD/880517 33
o Core thermal power of 2775 MHt. o 157. uniform steam generator tube plugging. o AFAH of 1.62. o Fuel data based on the Revised Thermal Hodel '(Reference 7). o A limiting chopped cosine power shape (Reference 26). During the transition period, a PCT penalty is applied to the full VANTAGE 5 core for the purpose of demonstrating conformance with the 10CFR50.46 PCT limit of 2200'F. (See Section 5.2.3) Reference 25 states three restrictions related to the use of the 1981 Evaluation Model + BASH calculational model. The application of these restrictions to the plant specific large break LOCA analysis was addressed with the following conclusions: V. C. Summer is neither an Upper Head Injection (UHI) nor an Upper Plenum Injection (UPI) plant, so Restriction 1 does not apply. V. C. Summe- Plant specific LOCA analysis analyzed both minimum and maximum ECCS cases to address Restriction 2. The Cd - 0.4 Double Ended Cold Leg Guillotine (DECLG) with minimum ECCS flows was found to result in most limiting consequences. Generic sensitivity studies have been performed by Hestinghouse which justify the continued use of the chopped cosine power shape as limiting for 3-loop plants which addresses restriction 3. 4 1406v:10/880517 34 . - - - - . _ _ ~ . . . . _ . - - - . - - - _ . . . _ ~ - -
6.2.1.2 Method of Analysis The methods used in analyzing the V. C. Summer Power Plant for VANTAGE 5 fuel, including computer codes used and assumptions, are described in detail in , Section 15.4.1.1.2. 6.2.1.3 Results The Double Ended Cold Leg Guillotine (DECLG, CD-0.4) with minimum ECCS was found to result in the most limiting consequences. The peak clad temperature was 2141'F at a total peaking factor of 2.45. The maximum local metal-water reaction was 10.13%, and the total core wide metal-water reaction was less than 0.37. for all cases analyzed. The clad temperature ' tansients turned around at a time when the core geometry was still amenable to cooling. The impact of the transition core cycles was conservatively assumed to be a 50*F increase in calculated peak cladding temperature which would yield a transition core PCT of 2191.0'F. The transition core penalty can be accommodated by the mir'in to the 10CFR50.46 limit of 2200*F. h The results of this analysis, including tabular and plotted results of the break spectrum analyzed, are provided in Appendix C which has been prepared , using the NRC Standard Format and Content Guide, Regulatory Guide 1.70, Revision 1 for accidents applicable to the V. C. Summer plant. 6.2.1.4 Conclusions The large break LOCA analysis performed for the V. C. Summer Power Plant has demonstrated that for breaks up to a double-ended severance of the reactor coolant piping, the Emergency Core Cociing System (ECCS) will meet the acceptance criteria cf Title 10 CFR Part 50 Section 46, that is:
- 1. The calculated peak cladding temperatm o will remain below the required 2200'F.
O 1406v:lD/880517 35
^ -
- 2. The 4 9 ? fuel cladding that reacts chemically with the water or steam does aed one percent of the total fuel rod cladding.
3, The localized cladding oxidation limit of 17 percent is not exceeded during or after quenching.
- 4. The core remains amenable to cooling during and after the LOCA.
- 5. The core temperature is reduced and decay heat is removed for an extended ,
period of time. This is required to remove the heat produced by the long-lived radioactivity remaining in the core. Thus, the ECCS analysis for the V. C. Summer Power Plant is in compliance with the requirements of 10CFR50.46 including consideration for trans'ition core configurations. 6.2.2 Small Break LOCA 6.2.2.1 Description of Analysis / Assumptions for 17x17 VANTAGE 5 Consistent with the methodology developed in the VANTAGE 5 Reference Core Report (Reference 1), a small break loss-of-coolant accident (LOCA) analysis - was performed assuming a full core of VANTAGE 5 fuel to determine the peak l' clad temperature. The currently approved Small Break ECCS Evaluation Model, using NOTRUMP, Reference 21, was utilized for a spectrum of cold leg breaks. Attachment 4, Section 15.3.1, includes a full description of the analysis and assumptions utilized. Key assumptions include an FAH of 1.68,'a total j peaking factor corresponding to 2.5 at the core mid-plane,157. uniform steam I generator tube plugging, and a core' thermal power level of 2775 MHt. Sensitivity studies performed using the NOTRUMP small break evaluation model have demonstrated that VANTAGE 5 fuel is more limiting than 0FA fuel in calculated ECCS performance. Similar studies using the HFLASH evaluation model, Reference 22, have previously shown that 0FA fuel is more limiting than LOPAR fuel. For the small break LOCA, the effect of the fuel difference is 1406v:lD/880517 36
more pronounced during core uncovery periods and, therefore, shows up predominantly in the LOCTA-IV calculation in the evaluation model analysis. Consequently, the previous conclusion drawn from the HFLASH studies, regarding the fuel difference, may be extended to this NOTRUMP , analysis. Thus, only VANTAGE 5 fuel was analyzed, since it is the more. limiting of the two types of fuel residing in the core. 6.2.2.2 Method of Analysis The methods of analysis, including codes used and assumptions, are described in detail in Attachment 4, Section 15.3.1. 6.2.2.3 Results The small break VANTAGE 5 LOCA analysis for the V. C. Summer Power Plant, utilizing the currently approved NOTRUMP Evaluation Model, resulted in a peak clad temperature of 2095'F for the 3.0 inen diameter cold leg break. The analysis assumea the limiting small break power shape consiste.it with a LOCA Fg envelope of 2.50 at core midplane elevation and 2.26 at the top of the core. .The maximum local metal-water reaction is 5.69 percent, and the total core metal-water reaction is less than 0.3 percent for all cases analyzed. The clad temperature transients turn around at a time when the core geometry is still amendable to ecoling. These results are applicable for V. C. Summer with a full core of VANTAGE 5 fuel and with transition cores. 6.2.2.4 Conclusions Analyses presented in Attachment 4. Section 15.3 show that one centrifugal charging punp, together with the accumulators, provide sufficient core flooding to keep the calculated peak clad temperature well below the required limits of 10CFR50.46. It can also be seen that the ECCS analysis remains in compliance with all other reaniremer.ts of 10CFR50.46. Adequate orotection is therefore afforded by the ECCS in the avant of a small breal LOCA. O 1406v:10/800517 37
i ! 7 6.2.3 Transition Core Effects on LOCA (% i .
/
The V. C. Summer large and small break analysis have been performed in accordance with the transition core LOCA methodology defined in Section 5.2.3 of VANTAGE 5 Reference Core Report. These analyses n're based on a full core of VANTAGE 5. To cover a transition core, the maximum PCT penalty of 50*F has been applied to the full VANTAGE 5 core results for large breaks to ensure conformance with the 10CFR50.46 PCT limit of 2200*F. Application of this maximum penalty, conservatively accounts for the potential increases in PCT due to the effects of mixed core hydraulic resistance mismatch as described in Reference 2. No PCT penalty has been applied to the small break results since mixed core hydraulic resistance mismatch is not a significant factor for the analysis. Following the transition period, the large break LOCA analysis will apply without the 50*F PCT penalty. 6.2.4 Containment Integrity Mass and Energy Releases The extent to which fuel changes can impact containment mass and energy O) v releases, used to determine containment peak prossure, is dependent upon changes to: ;
- 1) Core fluid volume. _;.
- 2) Core stored energy.
- 3) Core hydraulic resistance.
The VANTAGE 5 fuel design utilizes a fuel rod of smaller diameter than the ! 17xl? LOPAR (standard) fuel presently installed in the V. C. Summer Power Plant. This smaller fuel rod diameter leads to a reduction in core stored energy which is beneficial in reducing the mass and energy releases calculated i for a hypothetical LOCA. The small VANTAGE 5 fuel rod will also result in a slight increase in core fluid volume; and, the use of Intermediate Flow Hixing grids will increase hydraulic resistance. These changes are off set by-the reduction in core stored energy. Based on this offset, a reanalysis of containment integrity mass and energy releases is not necessary for the implementation of VANTAGE 5 fuel at the V. C. Summer Power Plant. Thus, the 1406v:lD/880517 38 l 1
)
implementation of VANTAGE 5 fuel at the V. C. Summer Power Plant will not g result in an increase in the containment peak pressure reported in the V. C. W Summer FSAR or increase the offsite radiological consequences associated with high containment pressures resulting from a hypothetical LOCA. 6.2.5 Steam Generator Tube Rupture _ The consequences of a Steam Generator Tube Rupture (SGTR), as analyzed in the V. C. Summer Power Plant FSAR, are dependent upon the initial reactor and steam generator conditions of power, pressure, and temperature. Changes in initial operating conditions as a result of implementation of VANTAGE 5 fuel at V. C. Summer Power Plant have been evaluated and concluded that the consequences of a SGTR will not be increased by the implementation of VANTAGE 5 fuel. Thus, a reanalysis of the FSAR Steam Generator Tube Rupture was determined to be unnecessary for the implementation of VANTAGE 5 fuel and the current FSAR SGTR analysis remains applicable. 6.2.6 Blowdown Reactor Vessel and Lo6p Forces O The forces created by a hypothetical break in the RCS piping are principally caused by the motion of the decompression wave through the RCS. The strength of the decompression wave is primarily a result of the assumed break opening time, break area, and RCS operating conditions of power, temperature, and pressure. These parameters will not be affected by a change in fuel at the V. C. Summer Power Plant from 17x17 Standard to VANTAGE 5. The forces in the vicinity of the core are affected by the core flow area / volume. An increase in core flow area / volume will tend to more effectively dissipate the decompression wave resulting in a reduction of the forces acting on the reactor vessel internals. VANTAGE 5 fuel, having a smaller rod diameter than 11x17 standard fuel, increases the core fiow area and volume which is beneficial in reducing forces associated with a hypothesized LOCA. Forces acting on the RCS loop piping as a result of a hypothesi:'ed LOCA are not influenced by changes in fuel assembly design. Thus, the implementation of VANTAGE 5 fuel at the V. C. Summer Power Plant will not result in an increase O 1406v:lD/880517 39
. . . . . . _ . __m .. .
of the calculated consequences of a hypothesized LOCA on the reactor vessel s internals or RCS loop piping. The current FSAR analysis for forces on the reactor internals and RCS piping resulting from a hypothesized LOCA remains applicable to the application of VANTAGE 5 fuel at the V. C. Summer Power Plant. 6.2.7 Post-LOCA Long-Term Core Cooling (ECCS Flows, Core Subcriticality, and Switchover of the ECCS to Hot leg Recirculation) The implementation of VANTAGE 5 fuel at the V. C. Summer Power Plant does not impact the assumptions for decay heat, core reactivity, or boron concentration l for sources of water residing in the containment sump Post-LOCA. Thus, these licensing requirements associated with LOCA are not significantly affected by the implementation of VANTAGE 5 fuel. Additionally, Westinghouse performs an independent check on core suberiticality for each fuel cycle operated at the V. C. Summer Plant. O 3 i i. l 1 i l lO F 40 1406v:lD/880517 i-I
7.0
SUMMARY
OF TECHNICAL SPECIFICATION CHANGES The proposed changes to the Virgil C. Summer Nuclear Station (VCSNS) Technical Specifications are summarized in Table 7.1. These changes reflect the impact of the design, analytical methodology, and safety ana' lysis assumptions outlined in the SCE!,G amendment request and are given in the proposed Technical Specification page changes (see Attachment 2 of this report). A brief overview of the significant changes follows. 7.1 Core Safety Limits Core safety limits and associated bases for 3-loop operation during modes 1 and 2 are revised to reflect the impact of the transition to VANTAGE 5 with:
- 1. The use of ITDP and the WRB-1 and WRB-2 DNS Correlation.
- 2. An F 3g of 1.62 (see Section 7.11).
- 3. Reduced RCS flow to accommodate the increased resistance of the VANTAGE 5 fuel assembly and 2% flow margin over and above that required to support SG tube plugging up to 15% (see Section 7.2).
The proposed limits corresponds to those for the LOPAR fuel which are limiting during the transition period. Less limiting values will be possible with a full core of VANTAGE 5. 7.2 Thermal Desien Flow The VCSNS thermal design flow is being decreased frem 96,200 gpm per loop to 92,600 gpm per loop. This flow reduction accommodates:
- a. The increased resistance of the VANTAGE 5 fuel assembly.
- b. Up to 15% SG tube plugging in all three SG's,
- c. 2% additional flow margin.
140sv:10/0s1688 41
i The revision proposed for Table 2.2-1 corresponds to the minimum measured flow s value used as input to the ITDP DNBR analyses for the loss of flow event. 1 The reduced thermal design flow has also been factored into the limiting j conditions of operation defined by RCS flow and th.e Nuclear Enthalpy Rise Hot l' Channel Factor (see Section 7.11) within Technical Specification 3.2.3. Indicated RCS flow is derived from the reduced thermal design flow based on l VCSNS's currently approved flow measurement uncertainty of 2.1%. l f l 7.3 OPAT/0 tot Setooints Revisions to the limiting safety-system settings for the thermal overpower AT and overtemperature AT trip functions are proposed to maintain consistency with the non-LOCA Accident Analyses provided in the Transition Safety Evaluation. .These trip functions provide primary protection against departure from nucleate boiling and fuel centerline melting (excessive kw/ft) during postulated transients. The proposed settings have been derived consistent with WCAP-8745, "Design Bases.for the Thermal Overpower aT and Thermal Overtemperature AT Trip Functions", based on the core safety limits i (see Section 7.1) for the VANTAGE 5 transition and with instrument uncertainties accounted for. , 7.4 Shutdown Marcin for Modes 3, 4, and 5 Figure 3.1-3 of the VCSN.', Technical Specifications defines shutdown margin requirements as a function of average RCS boron concentration during Modes 3, 4, and 5. The proposed revisions are based on the reanalyses of the boron dilution event with VANTAGE 5 fuel (see the Transition Safety Evaluation) and are required to maintain the current bases of the Technical Specification. No new technology was employed in the VANTAGE 5 Boron Dilution analysis. The revised limits reflect solely the impact of VANTAGE 5; the more negative boron worths and higher initial boron concentrations (both initial and critical) are the primary factors leading to the modified shutdown margin requirements. O 1405v:1o/051688 42 _ __ __. _ ____-____-______N
7.5 Moderator Temoerature Coefficient , Revisions to the VCSNS Technical Specifications for Moderator Temperature coefficient are proposed. The BOL limits are increased from 0 AK/K/*F for the all rods withdrawn (BOL, H2P) to +7.0 pcm/*F from if2P to 70% rated power with a negative linear ramp from +7.0 pcm/'F at 70% power to 0.0 pcm/'F at HFP. This change is required due to the increased RCS boron concentrations for VANTAGE 5 and the positive shift in moderator coefficient caused by the larger H/V ratio and small MTU icading for the smaller tracared to LOPAR) VANTAGE 5 fuel rod. The E0L limits are also in:reased (more negative) to accommodate longer fuel cycles and extended fuel burnup. 7.6 Berated Water Sources The technical specification requirements for borated water sources during all operating modes were evaluated to determine if current limits remain applicable for 'het transition to VANTAGE 5 fuel. The bases for the technical specification were preserved by current limits except for the ' minimum contained borated water volume in the boric acid storag'e system during Modes 1-4. An increase in the minimum water volume of 100 gallons (from 13200 gallons to 13300 gallons) is requested tc assure minimum shutdown margin from full power equilibrium xenon conditions, in conjunction with the abov.e, the examples of maximum expected boration capabilities in the bases section are deleted. These examples were based on operation with LOPAR fuel and are not applicable for the transition to VANTAGE ' 5 fuel. Margin to the technical specification limits will be confirmed in the future on a cycle specific bases. 7.7 Red Droo Time The VANTAGE 5 guide thimbles are identical to those in the LOPAR design except for a reduction in the guide thimble diameter and length above the dashpot. 1405v:10/Os1688 43
j Q The reduction to the guide tube diameters is required due to the thicker ' zircaloy grid straps and reduced cell size; whereas; the VANTAGE 5 thimble tube is shorter due to the reconstitutable top nozzle feature. To accommodate these changes the scram time to the dashpot for accident analyses is increased frem 2.3 seconds to 2.7 seconds for the transitiert to VANTAGE 5. The increased rod drop time has been used in the safety analyses provided in the Transition Safety Evaluation. 7.8 Axial Flux Difference Axial power distribution control at VCSNS is currently achieved by following the Constant Axial Offset Control (CAOC) procedure. In conjunction with the transition to VANTAGE 5, SCE&G proposes to replace CAOC with a modified version of the NRC approved Relaxed Axial Offset Control (RAOC) procedure described in WCAP-10216-PA. The proposed implementation at VCSNS includes:
- a. Operation within the standard RAOC AFD limits as a function of power,
- b. Option to operate in a base load mode above a minimum allowable power level with AFD maintained within a specified target band about a target flux difference.
- c. Use of burnup dependent AFD limits and baseload target band widths on a cycle specific basis. These valves would be supplied in the peaking
. factor limit report.
The Modified RAOC Technical Specification is a logical extension of the Westinghouse Stand'rd a RAOC Technical Specification. Basically, this concept involves the removal of the "Axial Flux Difference Limits as a function of Rated Thermal Power" figure from the Technical Specifications. The AFD Technical Specification is then modified to reference a cycle specific Peaking Factor Limit Report (PFLR) as the reference document for the AFD limit figt.e(s). O 1405v;1o/es1688 44
The base lead option provides the additional capability to make use of the ' reduction in core peaking factors that result due to the limited amount of xenon skewing that occurs during operation at er near equilibrium conditions. 7.9 Heat Flux Hot Channel factor - Fg (z) lt is proposed to increase the VCSNS Fg limit from 2.25 to 2.45 for greater flexibility and to accomodate the axially heterogeneous aspects (tlankets and short burnable absorbers) of the VANTAGE 5 core. Furthermore, the K(z) curve, which defines the axial dependancy of gF , is modified to remove the third line segment applicable to the tcp of the core. The full power F li: nit value of 2.45 was selected to support steam 0 generator tube plugging level of up to 15% while still limiting large break LOCA peak clad temperature values to less than 2200'F, with transition core penalties included. The axial' power profile used to perform the VCSNS small-break LOCA' analysis was derived using the recently improved Westinghouse power shape methodology. Among the most notable aspects of this methodology are the use of a comprehensive data base and the elimination of the third line segment from the K(z) Technical Specification curve. - 7.10 QSurveillance Another reque:ted Technical Specification change revises the surveillance technique for the heat flux hot channel factor frem F,y (z) to the NRC , approved Fg (z) Surveillance described in WCAP-10216-PA. This revised surveillance procedure for the total peaking factor, F0 , will increase plant operating flexibility while more directly monitoring the parameter of interest. It is further proposed to supply the height dependent analytical factor W (z) on a burnup dependent, cycle specific bases within the Peaking Factor Limit Report to take advantage of changes in the axial flux and xenon concentration distributions which occur with burnup and from cycle to cycle. O 1405v:1o/051688 45
i .~. . . 7.11 Nuclear Enthalpy Rise Hot ' Channel Factor The following F AH values (includes uncertainties) are proposed for the VANTAGE 5 transition. F3g = 1.56 [1 + 0.3 (1-P)] _ where P is the fraction of full power. These higher values allow increased fuel cycle design flexibility and lower leakage core loading patterns. 7.12 DNB Parameters The proposed limits on DNS related parameters (T,yg and Pressurizer Pressure) assure that each are maintained within the normal steady state envelope of operation assumed in the transient and accident analyses. The proposed revisions are consistent with new accident analyses supplied in the Transient Safety Evaluation which utilizes the ITOP (see Section 5.0) for DNS evaluations. O LJ The T,yg reflects the nominal baseline T,yg of 585.5'F assumed in the VANTAGE 5 analysis in order to support full power operation with:
- 1. 15% uniform SG tube plugging. i
- 2. A thermai design flow conservatively reduced 2% below that required to support 15% SG tube plugging and the use of VANTAGE 5 fuel.
- 3. No increase in the V. C. Summer original THOT f 618.7'F.
The analyses provided herein either reflect the use of or are conservative relative to the proposed T,yg of 585.5'F. O 140$v:1o/051688 46 l I i
7.13 OTaT,/0 pat Trio Resconse Times To reflect the assumptions used in the VCSNS analyses performed to support the transition to VANTAGE 5 fuel, a response time of 8.5 seconds for the overtemperature AT and overpower AT is requested. It corresponds to the total time delay (including RTD and thermowell tir.e response, trip circuit and channel electronic delay) from the time the temperature difference in the loop exceeds the trip setpoint until the rods are free to fall and was assumed within the non-LOCA analyses. Its use is intended to allow the potential future removal of the RTD bypass manifold without the need to perform additional safety analyses. 7.14 RC Pumo Underfrecuency Trio Resoonse Time A decrease in the RC Pump Underfrequency trip response time for 0.9 to 0.6 seconds is requested. This change reflects the new analysis value for the Complete Loss of Forced Reactor Coolant Flow accident in the Transition Safety Evaluation. O 7.15 ECCS Accumulators A change to the contained borated water volume of the ECCS accumulators is _ requested. The proposed change would require maintenance of accumulater water volume between 7489 and 7685 gallons during Modes 1, 2, and 3 (pressurized pressure above 1000 psig). The net benefit is a total gain of 2 inches of usable inventory which improves core reflooding during a large break LOCA. This impact is included in the large break LOCA analyses provided in Attachment 4. An evaluation of the potential effects of this increased accumulator water volume in areas of safety other than the Large Break LOCA has been performed. Areas considered include small break LOCA, non-LOCA analyses, LOCA forces, mass and energy releases, and the steam generator tube rupture event. In all cases the increased accumulator volume had ei'.her a beneficial or inconsequential impact on the analysis results. O 1405v:10/o.1688 -
1 7.16 Charcing Pumo Surveillance ( . Minimum SI flow for the VANTAGE 5 transition LOCA and non-LOCA analyses assumed that the charging pump mini-flow path (recirculation) remained open for the duration of the accident. Consequently, from a safety Analysis standpoint, isolation of mini-flow to maximize SI is an optional action (i.e., not required for accident mitigation) during.the injection phase. To maintain consistency with the new analysis assumptions, it is proposed that the charging pump flow balance test limits be changed to reflect the mini-flow path open. i-O P 4 l-1 6 I i l 1405v:10/Cs1688 48 r,-=- --ow..-,,--- , - - -,-,~,,-.r.,, -
TABLE 7.1 VIRGIL SLMiER . TECHNICAL SPECIFICATION CHANGES FOR CYCLE 5 RELOAD FAGE SECTION DESCRIPTION OF CHANGE JUSTIFICATICN 2-1 2.1.1 Delete reference to two Two loop operation is not loop operation currently licensed 2-2 Figure 2.1-1 Core limits are revised Core limits are revised for Vantage 5 fuel due to IEP, new peaking factors and reduced RCS flow 2-5 Table 2.2-1 Setpoints and Themal Setpoints and TDF are 2-8 Design Flow (TDF) are consistent with the new 2-9 changed safety limits, instrtnent 2-10 uncertainty and reduced flow B 2-1 Bases 2.1.1 Discussion of Themal The analysis is based on aM Hydraulic analysis ITDP methodology and the WRB-1 and -2 correlations B24 Bases 2.2.1 DNBR 1 hit replaced Future changes in analyses by "safety analysis will net require a charge DNER limit" in Bases 3/4 1-3a Figure 3 1-3 Char.ged shutdown margin This is based on for Modes 3, 4 aM 5 reanalysis of boren dilution with Vantage 5 fuel 3/4 1 4 3 1.1 3 Charged the MTC 11 nits This is based on Vantage 5 - fuel 3/4 1-5 4.1.1 3 Changed the MTC Limits m is is based on Vantage 5 fuel 3/4 1-12 3 1.2.6.a Changed mini =tn borated This is based en vantage 5 water voltne for boric fuel and exteMed fuel acid systec cycles 3/4 1-19 3 1 3.4 Charged rod drop t1=e This is based on vantage 5 fue! 3/4 2-1, 3 2.1 Changed AFD requirements This is based on Vantage 5 3/4 2-2, fuel, RAOC and base load 3/4 2-3 operation O 49
TABLE 7.1 VIRGIL SUMMER O TECHNICAL SPECIFICATION CHANGES FOR CYCLE 5 RELOAD PAGE SECTION DESCRIPTION OF CHANGE JUSTIFICATION 3/4 2-4 3 2.2 manged F(Q) - This is based on the optimized selection of - _ parameters 3/4 2-5, 4.2.2.1 Changed F(0) Surveillance This is based an 3/4 2-6 Requirements Vantage 5 fuel, RAOC and base load operation 3/4 2-7 Figure 3 2-2 manged K(z) curve The K(z) curve reflects the and figure ntaber to limits in accondance with 3 2-1 LOCA analysis. 3/4 2-8 323 manged P to reflect This is based on the 3/4 2-9 F-Delta-H of 1.62 and optimized selection of changed figure ntnber parameters 3 2-3 3/4 2-10 Figure 3 2-3 Changed total RCS flow This is based on the rate requirenent and analyzed RCS flow limit figure ntnber to 3 2-2 3/4 2-16 Table 3 2-1 Changed DNB parameters m anges reflect new O analysis values , 3/4 3-9 Table 3 3-2 Changed OTDT, OPDT This is based on the " response times optimized selection of parameters j 3/4 3-10 Table 3 3-2 Changed underfrequency The change reflects new response time analysis v.Gue 3/4 5-1 351 Chan e xi acetnulator This is baset on the water voltre revised LOCA .vxi non- , LOCA analysis [ ] i O 4 50 b p
TABLE 7.1 VIRGIL SUMMER - TECHNICAL SPECIFICATION CHANGES FOR CYCLE 5 RELOAD PAGE SECTION DESCRIPTION OF CHANGE JUSTIFICATION 3/45-6 4.5.2 Changed charging pump - his is based on the flow balance limit.to revised safety analyses, reflect testing with where the charging punp recirculation recirculation was not isolated during the accidents 3/4 10-2 4.10.2.2 Changed surveillance Reflects correct section nurbers to specification nunbers reflect use of FQ as for FQ surveillance opposed to Fxy surveillance B 3/4 1-2 3/4 1.1 3 Charged !ffC limits This is based en Vantage 5 fuel B 3/4 1-3 3/4.1.2 Changed boration volunes h is removes the cycle specific values. B 3/4 2-1 3/4.2, 3/4.2.1, osanged F(Q), DNB limit, his is based on optimized B 3/4 2-2 3/4.2.2 and deleted'Fxy, and revised selection of parameters, B 3/4 2-3 3/4.2.3 discussion of AFD Vantage 5 fuel, ITDP, B 3/4 2-4 RAOC and baseload B 3/4 2-5 operation B 3/4 5-1 3/4.5.1 Added ccrnment for borated The accunulators are water and secondary pipe actuated in the steam line break analysis ruptures 6-18 6.9 1.11 Change discussion of This is based on Vantage 5
^ peaking factor report fuel, RAOC and base load operation. The approved methodology is referenced O
51
1 ZZZ.;Ji._ Z Z. ~~ .. .. L . ._
8.0 REFERENCES
. 1. Davidson, S.L. and Kramer; H.R.; (Ed.) "Reference Core Report VANTAGE 5
~ Fuel Assembly," HCAP-10444-P-A, September 1985.
- 2. Davidson, S.L.; Iorti, J. A., "Reference Core Report - 17x17 Optimized .
Fuel Assembly," HCAP-9500-A, May 1982. f
- 3. Davidson, S. L. (Ed.), et al., "Extended Burnup Evaluation of Westinghouse Fuel" HCAP-10125-P- A, December 1985.
- 4. Miller, R. W. et al.. "Relaxation of Constant Axial Offset Control-F 9
Surveillance Technical Specification". HCAP-10217-A, June 1983.
- 5. Davidson, S. L. (Ed.), et al., "Westinghouse Reload Safety Evaluation Methodology, "HCAP-9272-P-A, July 1985.
- 6. Hiller, J. V., "Improved Analy~tical Models ijsed in West',nghouse fuel Rod Design Computations," HCAP-8720 (Proprietary), October 1976. l l
- 7. Leech, H. J., et al., "Revised PAD Code Thermal Safety Model" _
HCAP-8720-A2 (Proprietary). October, 1982.
- 8. George, R. A., (et al.), "Revised Clad Flattening Model," HCAP-8377 (Proprietary) and NCAP-8381 (Non-Proprietary), July 1974.
- 9. Risher, D. H., (et al.), "Safety Analysis for the Revised Fuel Rod
! Internal Pressure Design Basis," HCAP-8963-P-A (Proprietary), August 1978. ! 10. Skaritka, J., (Ed.), "Fuel Rod Bow Evaluation, HCAP-8691, Revision 1 (Proprietary), July 1979. l
- 11. Davidson, S. L., Iorti, J. A. (Eds.), "Verification Testing and Analyses I cf the 17x17 Optimized Fuel Assembly," HCAP-9401-P-A, August 1981.
i i 52 00681:6/880504
I i
- 12. Camden, T. M., et al., "PALADON-Hestinghouse Nodal Computer Code,"
HCAP-9485-P-A, December 1979.and Supplement 1, September 1981.
- 13. Davidson, S. L. (Ed.), et al., "ANC: Westinghouse Advanced Nodal Computer Code," HCAP-10965-P-A, September 1986. --
- 14. Chelemer, H., Boman, L. H., Sharp, D. R., "Improved Thermal Design Procedure," HCAP-8567, July 1975.
- 15. Motley, F. E., et al., "New Westinghouse Correlation HRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane grids," HCAP-8762-P-A and WCAP-8763-A, July 1984.
- 16. Letter from NRC to Hestinghouse from Stolz to Eiche1dinger, SER on HCAP-7956, 8054, 8567 and 8762 dated April 1978.
- 17. Letter from E. P. Rahe (B) to Miller (NRC) dated March 19, 1982, NS-EPR-2573, HCAP-9500 and NCAPS 9401/9402 NRC SER Mixed Core Compatibility Items.
- 18. Letter from C. O. Thomas (NRC) to Rahe (8) - Supplemental Acceptance No.
2 for Referencing Topical Report WCAP-9500, January 1983.
- 19. Kabadi, J. N. et al., "The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code," HCAP-10266-P-A, March 1987, (Westinghouse Proprietary).
- 20. Letter H. Johnson (H) to J. Lyons (NRC), "Submittal of HCAP-10266 Addendum 1, BASH Power Shape Sensitivity Studies," January 26, 1987, Revised June 22, 1957.
- 21. Lee, N. Rupprecht, S. D., Schwart, H. R., Tauche, H. D., "Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," HCAP-10054-P-A (Proprietary) and HCAP-10081-A (Non-Proprietary) August 1985.
O 53 00681:6/880504
i ] i l i i
- 22. Esposito V. J., Kesavan, K., and Maul, B. J.; "H-FLASH-A Fortran-IV l l- Computer Program for Simulation of Transients in a Multi-Loop PHR " l i HCAP-8200 (Proprietary), July 1973. , l 4
4: i
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- 6 l@
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4 4 c 00681:6/880504 , i
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~.
l t ATTACHMENT 2 TECNNICAL SPECIFICATIONS CHANGE PAGES i FOR THE _ V. C. SUMMER PLANT TRANSITION TO 17x17 VANTAGE 5 FUEL O 5 0 00681:6/880504
- i t' - .
2.0 SAFETY LIMITS MD LIMITIN5 $AFETY $YSTEM SETTINGS 2.1 SAFETY LIMITS t -
- RE ACTOR CORE 2.1.1 The combination of THERPML POVER, pnseurizer pressure, and the highest shall not exceed the limits shown in
^
operating loop coolant tasperature (TFigures 2.1-1 anM L b4-for 3 M185$) operation 1 APPLICA3fL77Y: MDES 1 and 2. ACTION: Wenever the point defined by the combination of the highest operating loop ( average tee.perature and THERXAL POVER has exceeded the appropriate pressurizer pressure line, be in NT STMD5Y within I hour, and comply with the require-sents of Specification 6.7.1. REACTOR COOLANT SYSTEN PRES $URE 2.1.2 The Reactor Coolant Systes pressure shall not exceed 2735 psig. j APPLICA3 f LITY: MDES 1, 2, 3, 4 and 5. . ACTION: M DES 1 and 2 Wenever the Reactor Coolant Systes pressure has exceeded 2735 psig, be in MT STMD5Y with the Reactor Coolant Systen pressure within its lirit _. within 1 hour, and comply with the requirements of Specification 6.7.1. .. , MDES 3, 4 and 5 Wenever the Reactor Coolant Systes pressure has axceeded 2735 psig, reduce the Reactor Coolant systes pressure to within its limit within 5 minutes, and comply with the requirements of Specification 6.7.1. s SimER - UNIT 1 2-1 O
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*W en operating in the reduced RTP region of Technical Specification 3.2. l igure 3.2-3), the restricted power level must be considered 100% RTP fo this figure.
p Figure 2.1-1 Reactor Core Safety Limit - Three Loops in Operation \
\
Amendment No. 45 SUMMER - UNIT 1 2-2
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O ~ O - O~ . E TABLE 2.2-1 3 9 REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS e c
- _ Total Functional Unit ,
Allowance (TA) Z_ S Trip Setpoint Allowable Value l ~ 1. Manual Reactor Trip Not Applicable NA NA NA NA .
! 2. Power Range, Neutron Flux 7.5 4.56 0 1109% of RTP 1111.2% of RTP I High Setpoint Low Setpoint 8.3 4.56 0 125% of RTP 127.2% of RTP
- 3. Power Range Heutron F1ux 1. 6 0.5 0 15% of RTP with $6.3% of RTP witle High Positive Rate a time constant a time constant j 12 seconds 12 seconds
- 4. Power Range, Neutron Flux 1.6 0.5 0 15% of RTP with $6.3% of RTP with I m High Negative Rate a time constant a time constant 12 seconds 12 seconds
- 5. Intermediate Range, 17.0 8.4 0 -<25% of RTP -<31% of RTP Neutron Flux .
- 6. Source Range Neutron Flux 17.0 10.0 0 <105 cps -<1.4 x 105 cps 9.8 7. ?9 1.9 4 /.fdd <
l
- 7. Overtemperature AT 7d- .2M4- .J.4t- See note 1 See note 2 s'.z 2.2 L 1.9 '
- 8. Overpower AT M X X See note 3 See note 4
- 9. Pressurizer Pressure-Low 3.1 0.71 1.5 11870 psig 11859 psig
- 10. Pressurizer Pressure-H!gh 3.1 0.71 1.5 12380 psig $2391 psig R 11. Pressurizer Water Level-High 5.0 2.18 1.5 192% of instrument 193.8% of instrument 3 span span 5
z 12. Loss of Flow 2. 5 1.0 1.5 190% of .4eep 189.2% of 4eep P .
& ! g. ' L
- e ri p '!--? -
O #73 f all"vt b at -?vt a aJ # set es mg L :m p t : p ' i - = M , ?nn nn._ mea 5aatED gg4 5,,ggg RTP = RATED THERMAL POWER EL0 4d d p[g g c wmmam me.asae.eo vt.o.s = 9*1500 ym/ loop 4$ l 9 Yo SPA ^2 foR. DGt.TA -T (KTDs) n.0 c /.2 Y. Foe *PRESsu212EK PRESSutE
o O' U. ,
, TABLE 2.2-1(Continuedl s g
REACTOR TRIP $YSTEM INSTR;;HENTATION TRIP SETPOINTS j -
' HOTATION l.
NOTE 1: DVERTEMPERATURE AT 4 AT1AT,[Kg-Kafy } [T.- T'] + K3 (P - P') - f (al)) ifhere: AT = Measured AT by RTD Manifold Instrumentation
- i. .
AT, 6 K Indicated AT at RATED THERMAL POWER
~
Kg iW W },103 2p p o.o3006 ! K. I = The function generated by the lead-lag controller for T g
't e=
} dynamic compensation , ,
tg. & tg = Time constants utilfred in the lead-1de controller for Tg, tg)aJ8 secs. , . In # 4 SeC5-I T = Average temperature *F. ~
- 5Ff.i ,
l T' $ .5sht*F Reference T,q at RATED THERMAL POWER Ks 1M M e0.00l47 , l I P = Pressurizer pressure, psig 4 P' 1 Jr 2235 pstg, Nominal RCS operating pressure 1
$ = laplace transform operator, sec 1
)1 i I 1 : i D *
.R D D
m . O e ._ 0 b i, TAntE 2.2-1 (Continised) er. J k REACTL 1 RIP SYSTEM INSTRIMENTATION TRIP SETPOINTS 9 NOTATION (Continued)
- e i
- 5 NOTE 1: (Continued)
[
- and f,(at) is a function of the Indicated difference between top and bottom detectors of the j power-range nuclear len chambers; with gains to be selected based on measured instrtament -
l F response during plant startup tests such that: 14 . 1. j (1) for g +g g between 'y(percent and + percent f (at) = 0 where qg W g am pecut ,
- RATED THERMAL POWER in the top and bottom halves of the core respectively, and gg + g is b
total THERMAL. POWER in percent of RATED THERMAL POWER. ! -2t (11) for each percent that the magnitude of qt -9b exceeds -p percent, the aT trip setpoint - l :f shall be aatomatically reduced by d percent of its value at RATED THERMAL POWER. l 2.17 4 i ~ 4 (Ill) for each percent that the magnitude of qt - q, exceeds */ percent, the AT trip setpoint , sha11 be automatically reduced by A41 percent of its value at RATED THERMAL POWER. 2.l4 i NOTE 2: The channel's sauteum trip setpoint sha11 not exceed its computed trip point by more than
; 2.,o Rpercent AT span.
I NOTE 3: OVERPOWER AT l ai5at,IR-R(3Y,,3) 1 - R. ti - T=] 1 , . i Where: of = as defined in Wate 1 I j AT, = as derined.in Note 1 i R. ix c=4 1.o 27r I Rs 1F 0.02/*F for increasing average temperature and 0 for decreasing average temperature y'* = The function generated by the rate-lag controller for T,,, dynamic cmnsation
, 5 l
l-
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.o..
l 1 l 1ABLE 2.2-1 (Continced) REACTOR TRIP SYSTEM INSTRUNENTATION TRIP SETPOINTS
' ' NOTATION (Continued) l $'
- c-
- f. NOTE 3: (Continued) g ts = Time con tant utilf red in the rate-lag controller for T,,g. ts >(10 secs.
0 00lf
- K. 21)f 0.paliSO/*F for T > 1" and K. = 0 for T 1 T" T = as de ined in Note 1
- 5.
1" $ . "F Reference T,,g at RATED TitERMAL POWER S = as defined in Note 1 NOTE 4: The channel's maximum trip setpoint shall not exceed its computed trip point by more than f<f' percent ai span.
,, 2. 6) a c2 i
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- l 1
,m 2.1 SAFETY LIMITS BASES 2.1.1 REACTOR CORE el and The restrict'ons of this Safet Limit prevent ov rheating of the ossible cladding erforation which ould result in t release 1' cladding of fiss on is pr vented
'p oducts to the re tor coo? ant. Ove heating the of the f nucleate bo ling regime where the gb restricting fuel peration to with he transfer coeffi ient is large and the cladding surf ce temperature is sli tly above the co lant saturation t mperature.
q the nucleate bo ling regime could g eration above t upper boundary o e bq result in excessive cla ing temperatures ecause of the on et of departure D W from nu leate bniling (0 ) and the result t sharp reductio in heat transfe y coeffici nt. ONB is not directly measura le parameter duri g operation and { therefor THERMAL POWER an Reactor Coolant emperature and P ssure have been 4 ' h related t DNB through the -3 correlation. he W-3 DNB corre tion has been
- developed o predict the DN flux and the loc ion calofDNB DNB for a heat. fluially ratio,uniform ONBR,
- t The b nd non-uni orm heat. flux di ributions.
t flux that woul cause DNB at a p rticular fined as t e ratio of the h c re locat' ion to the incal hea flux, is indicat' ve of the margin o ONB. , v
$g The minim value of the DN R during steady s ate operation, n rmal This ope ational tra ients, and antic' pated transients is percent correspond to a 95 percent robability at a limited to 1.3 .
confide ce 13 vel f g vai that NB will not occur and is cho n as an appropri te margin to ONB or a'l
~
h opera ing conditio s. of points of f T e curves of 'gures2.1-1andk.1-2showtnelo t THERMAL POWC_R, React Coolant System hressure and aver e temperisture fo b{ l' which t minimum DNB isnolessthan130,ortheaver e enthalpy at th 4 l vessel e it is equal t theenthalpyofgaturatedliquid. - h N of >8 nd These curves are based on an enthalpy hot channel factor, FAkH'llow)ance a is \ N areferencecosinewithapeagof1,55foraxialpowershape.at reduced power based on included for an increase in.Fg 0.3 Fh='/,% I:>&S [l+ 3NC(1-P)] where P is the fraction of RATED THERMAL POWER These limiting heat flux conditions are higher than those calculated for the range of all control rods fully withdrawn to the maximum allowable control rod insertion assuming the axial power imbalance isWhen within thethe limits axial powerof the f t (delta I) function of the Overtemperature trip. O imbalarne is not within the tolerance, the axial power imbalance effect on the i \ Overtemperature delta T trips will reduce the setpoints to provide protection consistent with core safety limits. B 2-1
, SUMMER - UNIT 1 I
INSERT 1 The resteictions of this safetty limit prevent overheating of the fuel and possible cladding perforation which would result in the release of fission products to the reactor coolant. Overheating of the fuel cladding is prevented by restricting f .el 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 could result in excessive cladding temperatures because of the onset of departure from nucleate boiling (DIB) and the resultant sharp reduction.in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POWER and Reactor Coolant Temperature and Pressure have been related to DiB. Bis relation has been developed to predict the DB flux and the location of DIB for axially uniform and non-unifom heat flux distributions. The local DG heat flux ratio (DiBR) defined as the ratio of the heat flux that would cause WB at a particular core location to the local neat flux, is indicative of the margin to DG. The DiB design basis is as follows: there must be at least a 95 percent probability that the minimum UGR of the limiting rod during Condition I and II events is greater than or equal to the DER limit of the DNB correlation being used (the WRB-1 or KRB-2 correlation in this application). The correlation DGR limit is established based on the entire applicable experimental data set such that there is a 95 percent probability,with 95 percent confidence that DiB will not occur when the minimum DGR is at the DER limit (1.17 for the WRB-1 or WRB-2 Correlation) . In meeting this design basis, uncertainties in plant operating parameters, nuclear and thermal parameters, and fuel fabrication parameters are considered statistically such that there is at least a 95% probability with 95% confidence level that the minimum DGR for the limiting rod is greater than or equal to the DiBR limit. The uncertainties in the above plant parameters are used to deter.nine the plant DGR - uncertainty. Bis NGR uncertainty, combined with the correlation DIER limit, establishes a design DNER value which must be met in plant ::afety analyses using values of input parameters without uncertainties. In addition, margin has been maintained in the design by meeting safety analysis DGR limits in performing safety analyses. The curves of figure 2.1-1 show the loci of points of THERMAL POWER, Reactor Coolant System pressure and average temperature below which the calculated DGR is no less , than the design DGR value or the average enthalpy at the vessel exit is less than the enthalpy of saturated liquid. O
w =
. LIMITING 5AFETY $YSTEM SETTINGS
' 'I BASES l
REACTOR TRIP SYSTEM INSTRtMENTATION SETPOINTS (Continued)* The various reactor trip circuits automatically gpen the reactor trip breakars deaever a condition monitored by the Reactor Protection System ' reaches a preset er esiculated level. In addition to redundant channels and trains, the design approach pmvf des a Reactor Protection System which monitors numerous system variables, therefore, providing protection system functional ! diversity. The Reactor Protection Systen initiates a turbine trip signal whenever reactor trip is initiated. This prevents the reactivity insertion " that would otherwise result from excessive reactor systes cooldown and thus avoids unnecessary actuation of the Engineered Safety Features Actuation System. ,, Manual Reacter Trip The Reactor Protection Systee includes manual reactor trip capability, Power Rance Neutron Flux In each of the Power Range Neutron Flux channels there are two independent bistables, each with its own trip setting used for a high and low range trip setting. The low setpoint trip provides protection during suberitical and low power operations to mitigate the consequences of a power excu-ston beginning from low power, and the high setpoint trip provides protection during power i operations to sitigate the consequences of a reactivity excursion from all power levels. f s The low setpoint trip say be annually blocked above P-10 (a power level !! of tpproximately 10 percent of RATED THERMAL POVER) and is automatically I reinstated below the P-10 setpoint. Power Rance. Neutron Flux Mich Rates The Power' Range Positive Rata trip provides protection against rapid flux increases which are characteristic of a rupture of a control rod drive housing. Specifically, this trip complements the Power Range Neutron Flux High and Low
- trips to ensure that the criteria are met for' rod ejection fnas mid power. .
The Power Range Negative Rate trip provides protection for control red drop accidents. At high power, a rod drop accident of a single or sultiple l rods could cause local flux peaking whte.h could cause an unconservative local : DM8R to exist. The Power Range Negative Rate trip will prevent this free occurring by tripping the reactor. No credit is taken for operation of the < Power Range Negative Rate trip fcr those control rod drop accidents for which DNBR's will be greater than h% rat /.W T v4 /u s.. l Intermediate and Source Rance, Nuclear Flux
~
The Intermediate and Source Range, Nuclear Flux trint eirovide reactor core protaction during reactor startup to sitigate the consequences of an SUP94ER - UNIT 1 B 2-4
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r 4 REACTIVITY CONTROL SYSTEMS - MODERATOR TEMPERATURE COEFFICIENT LIMITING CONDITION FOR OPERATION
- 3.1.1.3 The moderator temperature coefficient (MTC) shall be:
h lim;.h: slow h1 F,$vre 3, i - 0,. at
- a. Less positive than & d:lte L/k/*r for the :11 -ed ui+'.dre"n, 5:gfa iac a' cy:12 life '00L), het ::r: T"ER.".f.L PO' DER -;endi t4en.
-50 4
- b. Less negative than -4t t x 10 delta k/k/*F for the all rods l withdrawn, end of cycle life (EOL), RATED THERMAL POWER condition.
APPLICABILITY: Specification 3.1.1.3.a - MODES 1 and 2" only# Specification 3.1.1.3.b - MODES 1, 2 and 3 only# ACTION:
- a. With the MTC more positive than the limit of 3.1.1.3.a above operation in MODES 1 and 2 may proceed provided:
- 1. Control rod withdrawal limits are established and maintained suf ficient to restore the MTC to less positive than 9-htta de l *' fr 5AcaN
(~ s 'M Fg 3,f_ o (j k'k/*f within 24 hours or be in HOT STANOBY within' the next 6 hours. These withdrawal limits shall be in addition to the insertion limits of Specification 3.1.3.6.
- 2. The control rods are maintained within the withdrawal limits established above until a subsequent calculation verifies that the MTC has been restored to within its limit for the all rods ,
withdrawn condition.
- 3. In lieu of any other report required by Specification 6.9.1, a Special Report is prepared and submitted to the Commission pursuant to Specification 6.9.2 within 10 days, describing the value of the measured MTC, the interim control rod withdrawal limits and the predicted average core burnup necessary for restoring the positive MTC to within its limit for the all rods withdrawn condition.
- b. With the MTC more negative than the limit of 3.1.1.3.b above, be in HOT SHUTDOWN within 12 hours.
"With K,ff greater than or equal to 1.0
#See Special Test Exception 3.10.3 O
SUMMER - UNIT 1 3/4 1-4 6
-. - - - , . . + _ - . ,,_,,,___,.-,_.,,,._---,..,_._,--,____v. _ _ , . _ _ , - , . 1- - ...-- . , . - _
REACTIVITY CONTROL SYSTEMS SURVEILLANCE REQUIREMENTS
-- ~ -
4.1.1.3 The MTC shall be determined to be within its limits during each fuel
, cycle as follows:
- a. The MTC shall be measured and compared to the BOL limit of Specifi-cation 3.1.1.3.a. above, prior to initial operation above 5% of RATED THERMAL POWER, after each fuel loading.
(\ b. TheMTCshg11bemeasuredatanyTHERMALPOWERandcomparedto Gr3 x 10 delta k/k/ F (all rods withdrawn, RATED THERMAL POWER gf condition) within 7 EFPD after reaching an equilibrium boron concen-tration of 300 ppm. IntheevengthiscomparisonindicatestheMTC is more negative than 4-iPx 10 delta k/k/ F, the MTC shall be remeasured, and compared to the EOL MTC limit of specification 3.1.1.3.b, at least once per 14 EFPD during the remainder of the fuel cycle. d-) , O I . l 1 ( l O SUMMER - UNIT 1 3/4 1-5 1
(-.
\ j 1 F .9
< .8 o
C UNACCEPTABLE g .7 OPERATION W W ACCEPTABLE g .6 OPERATION 8 w .s 3 i2
$ .4 k a
#,.3
\(
i i w .2 8 2 ,) 0 0 10 20 30 40 50 60 70 80 90 100
% OF RATED THERMAL POWER _..
HGURE 3.10 MODERATOR TEMPERATURE COEFFICIENT VS POWER LEVEL i l I l l I Summer UNIT 1 3/4 1.Sa
W e 1 REACTIVITYCONTROLSYSTIHS BORATE 0 VATER SOURCES - OPERATING O' t1MITING CONDITION FOR OPERATION 3.1. 2. 6 As a sinimum, the following borated water source (s) shall be OPERABLE ) as required by Specification 3.1.2.2: ) l A boric acid storage system with: 33'300 l a.
- 1. A minf aun contained borated water volume of gallons,
- 2. Between 7000 and 7700 ppe of boren, and
- 3. A minimum solution temperature of 65'F.
- b. The' refueling water storage tank with:
- 1. A minimum contained borated water volume of 453,800 gallons,
- 2. A minimum boron concentration of 2300 ppe, and )
- 3. A minimum solution temperature of 40'F.
. APPLICABILITY: NODES'1, 2, 3 and 4. h ACTICN:
- a. With the boric acid storage system inoperable and being used as one ~
of the above required borated water sources, restore the storage system to OPERABLE status within 72 hours or be in at least HOT STANOBY within the next 6 hours and borated to a SHUTDOWN MARGIN equivalent to at least 2 percent delta k/k at 200'F; restore thu boric acid storage systee to CPERABLE status within the next 7 days or be in COLD SHUTDOWN within the next 30 hours,
- b. With' the refueling water storage tank inoperable, restore the tank ,i to OPERABLE status within one hour or be in at least HOT STANDBY within the next 6 hours and in COLL SHUT 00W within the following l
l 30 hours. T O SUM ER # UNIT 1 3/4 1-12 l
l l : \ . t _ REACTIVITY CONTROL SYSTEMS ROD DROP TIME .
. LIMITING CONDITION FOR OPERATION 2.7
- 3.1.3.4 The individual full length (sinutdwn and control rop time from
- the fully withdrawn position shall be less than or equal _ yeconds from beginning of decay of stationiry gripper,. coil voltage to ashpot entry with:
- a. T,yg greater thah or e ua1 I t'o 5$1*F, and
(,
, b. All reactor toolant pumps operating.
APPLICABILT TY: M DES 1 ,an'd 2. ACTION: - . With the drop time of any full length rod determined to exceed the above lir.it, restore the red drop time to within the above limit prior to proceeding to E DE 1 or 2.
." SURVEILLANCE REQUIREMENTS l 4.1.3.4 The rod drop time of full length rods shall be demonstrated through
- measuruent prior to reactor criticality
- a. For all rods follwing each ruoval of the reactor vessel head. -
- b. For specifically affected individual rods following any maintenance I on or modification to the control rod drive system which could ,
affect the drop time of those specific rods, and
- c. At least once per 18 months.
( ! C SUPHER - WIT 1 3/4 1-19 l I
l REPu>c6 Tex 7 on Pace l 3/4.2 POWER DISTRIBUTION LIMITS 3/4.2.1 AXIAL FLUX DIFFERENCE (AFD) LIMITING CONDITION FOR OPERATION
.2.1 The indicated AXIAL FLUX DIFFERENCE (AFD) shall be maintained within a i target band (flux difference units) about the target flux difference.
APPLI ILITY: H0DE 1 above 50% of RATED THERMAL POWER
- ACTION:
- a. With the indicated AXIAL FLUX DIFFERENCE outride of the i target band a ut the target flux difference and with THERMAL WER.
- 1. Abov 90% of RATED THERMAL POWER, within 15 mi es either:
a) Re ore the indicated AFD to within the arget band N limi , or p b) Reduce ERMAL POWER to less than of RATED THL.tMAL y POWER. k 2. Between 50% and .of RATED THERMA POWER: a) ay continue provided: h POWER OPERATIO The indicate FDtdnotbeenoutsideofthe15% 1) g target band for, )re than 1 hour penalty deviation s cumulative du ng he previous 24 hours, and 3 2) The indicate AFD is ithin the limits shown on Figure 3.gl. Otherwi , reduce THERMAL POWER to W 1ess t n 50% of RATED T RMAL POWER within 30 minutes and r,e uce the Power Range eutron Flux-High Trip { Setpoints to less than or eq 1 to 55% of RATED THERMAL POWER within the next hours. b) Sur,/eillancetestingofthePower' Rang Neutron Flux C(annels may be performed pursuant to Sp ification 4.3.1.1 provided the indicated AFD is main ined within the limits of Figure 3.2-1. A total 16 hours operation may be accumulated with the AFD outs e of the target band auring this testing without penalty viation.
- b. ERMAL POWER shall not be increased above 90% of RATED THE POWER unless the indicated AFD is within the 15% target band an ACTION a.2.a) 1), above has been satisfied.
l /SeeSpecialTestException3.10.2 SUMMER - UNIT 1 3/4 2-1
. . ~ . - ~ -
POWER DISTRIBUTION LIMITS O \CTION(Continued) i
- c. THERMAL POWER shall not be increased above 50% of RATED THERMAL T POWER unless the indicated AFD has not been outside of the 15% !
target band for more than 1 hour penalty deviation cumulative dur g j the previous 24 hours. Power increases above 50% of RATED THE 1 OWER do not require being within the target band provided the : a umulative penalty. deviation is not violated. l + SURVEILLANCE QUIREMENTS , s - N 4.2.1.1 The indica d AXIAL FLUX OIFFERENCE shall be determin to be within its limits during POW OPERATION above 15% of RATED THERMAL OWER by:
- a. Monitoring the ndicated AF0 for each OPERABLE core channel:
- 1. At least onc per 7 days when the AF0 M itor Alarm is OPERABLE,
> and
- 2. At least once pe hour for the firs 24 hours after restoring b the AFD Monitor Al m to OPERABLE atus,
- b. Monitoring and logging the ndicate AXIAL FLUX OIFFERENCE for each
% OPERABLE excore channel at i st o ce per hour for the first 24 hours and at least once per 3 utes thereafter, when the AXIAL O3N FLUX OIFFERENCE Monitor Alarm the indicated AXIAL' FLUX DIFF EN inoperable. The logged values of shall be assumed to exist during the interval preceding each oggin .
4.2.1.2 The indicated AFD shall considered o side of its 15% target band ~ 3 when 2 or more OPERABLE excore e nnels are indica ing the AFD to be outside the target band. Penalty devi ion outside of the % target band shall be accumulated on a time basis : 4 One minute pena y deviation for each one minu of POWER OPERATION ( a. outside of th target band at THEMAL POWER leve 50% of RATED HERMAL POWER, and equal to or above
- b. One-half nute penalty deviation for each one minut of POWER OPERATI outside of the target band at THERMAL POWER vels between 15% a 50% of RATED THERMAL POWER.
4.2.1.3 The arget flux difference of each OPERABLE excore channel s 11 be determined measurement at least once per 92 Effective Full Power Day The provision of Specification 4.0.4 are not applicable. 4.2.1. The target flux difference shall be updated at least once per 31 Effe ive Full Power Days by either determining the target flux difference pur uant to 4.2.1.3 above or by linear interpolati6n between the most recently asured value and 0 percent at the end of the cycle life. The provisions of pecification 4.0.4 are not applicable. N O , SUMER - UNIT 1 3/4 2-2
INSERT 2 3 2.1 The indicated AXIAL FLUX DIFFERENCE ( AFD) shall be maintained within: 0
- a. the allowed operational space defined in the Peaking Factor Lilnit Report (PFLR) for Relaxed Axial Offset Control (RAOC) operation,'
or
- b. within the target band specified in the PFLR about'the target flux difference during base load operation.
APPLICABILITY: M0EE 1 above 50% of RATED THERMAL POWER". ACTION:
- a. For RAOC operation with the indicated AFD outside of the applicable limits specified in the PFLR,
- 1. Either restore the indicated AFD to within the PFLR specified limits within 15 minutes, or
- 2. Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within 30 minutes and reduce the Power Range Neutron Flux - High Trip setpoints to less than or equal to 55% of RATED THERMAL POWER within the next 4 hours,
~
- b. For Base Load operation above APL" with the indicated AFD outside of the applicable target band about the target flux difference:
- 1. Either restore the indicated AFD to within the PFLR specified target band within 15 minutes, or
- 2. Reduce THERMAL POWER to less than APL" of RATED THERMAL POWER and -
discontinue Base Load operation within 30 minates.
- c. THERMAL POWER shall not be increased sbove 50% of RATED THERMAL POWER unless the indicated AFD is within the applicabl; RAO limits.
- See Special Text Exception 3 10.2. .
ND is the minimum allowable power level for base load operation and will be
##APL provided in the Peaking Factor Limit Report per Specification 6.9.1.11.
O
INSERT 2 (continued) v POWER DISTRIBUTION LIMITS SURVEILLANCE REQUIREMENTS , 4.2.1.1 The indicated AFD shall be determined to be within its limits during POWER OPERATION above 50% of RATED THERMAL POWER by:
- a. Monitoring the indicated AFD for each OPERABLE excore channel at least once per 7 days when the AFD Monitor Alarm is OPERABLE.
- b. Monitoring and logging the ~ indicated AFD for each OPERABLE excore channel at least once per hour for the first 24 hours and at least once per 30 minutes thereafter, when the AFD Monitor Alarm is inoperable. The logged values of the indicated AFD shall be assumed to exist during the interval preceding each loggf.ng.
4.2.1.2 The indicated AFD shall be connidered outside of its limits when two or more OPERABLE excore channels are indicating the AFD to be outside the limits. 4.2.1 3 When in Base Load operation, the target axial flux difference of each OPERABLE excore channel shall be determlined by measurement at least once per 92 Effective Full Power Days. The provisions of Specification 4.0.4 are not applicable:. p 4.2.1.4 hhen in Base Load operation, the target flux difference shall be updated at least once per 31' Effective Full Power Days by either determining the target flux difference in conjunction with the surveillance requirements of Specification 4.2.13 above or by linear interpolation between the most recently measured vilue and the calculated value at the end of cycle life. The provisions of Specification 4.0.4 are not applicable. 4 0
h-TL pp iAbg b k. O [ l S u m mER- D.o rr l 3/4 2-3 1 1 -
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m+... - 3 - ADDITIONAL AMRM '5ELOW Po E
- ~ "~ZON VIOLATION TARGET E 10 _-
BAND FOR ONE H R CUMULATIVE E , IN ANY 24. HOUR PE OD ::
.e -
r - ..'-- O 60 -40 30 20 10 0 10 20 3 40 IN CATED FLUX DIFFERENCE (All(PERCENT) Figure 3. Axial Flux Difference Limits as a Function of Rated Thermal Power REflACG w TA qyoscrE fAc&
~
SilWER IINIT 1 1/4 ?-3
e.2 w s? 10:53 m. ees 001
! l
@ERDISTRIBUTIONLIMITS 3/4.2.2 HEATFLUXHOTCHANNElFACTOR-Fg l :
LIMITING CONDITION FOR OPERATION , 3.2.2 F A(Z) shall be limited by the following' relationships: ' 2.4f Fq (Z) $ (K(Z)] for P > 0.5 Fg (Z) $ E+rt$ [K(Z)] for P $ 0.5 . where P = THEW 4AL POWER RArau TMERMAL POWER g and K(Z) is the function obtained from Figure 9:t-t for a given core height location.
.. APPLICA8ILITY: M00E 1. ,, ,
ACTION. , With Fg (Z) exceeding its limit:
- a. Reduce THERNAL POWER at least 3 for each 3 Fg (Z) exceeds the !
limit within 15 minutes and similarly reduce the Power Range Neutrcn Flux-Migh Trip Setpoints within the next 4 hours; POWER _. OPERATION may proceed for up to a total of 72 hours; subsequent POWER OPERATION may proceed provided the Overpower delta T Trip Setpoints have been reduced at least N for each n F (Z)
"exceeds' the limit'.
9
- b. Identify and correct the cause of the out of limit condition prior to increasing ~ THERMAL POWER aoove'tM reduced limit required by a, above; THERMAL POWER may then be increased provided F (Z) is , demon-strated through incore mapping to be within its limitq O i r
H
!I S - UNIT 1 3/4 2-4 Menmen No. g
gy4 9 5 90 Wit 015TRIBLTT10N LIMITS g[Dl6LL f(SN 9=L u,+ fu~ m rem SURVElu.AMCE R10VIREMENTS
- 4. L 2.1
- p n ylsfens of Specification 4.0.4 are not applicable.
4.2.2.2 F, hall he evaluated to setsreina if Fq (Z) is within i Itait by: ,
- s. bing ne novable 'incere detecters to obtain a power. istribution ERML POWER.
snap at TMERML Poster gnater than 5% of RATED
- b. Increasin the measured F component of the er distribution map .
1 by 35 to a ountformanufleturingtolerances nd further increasing the value by to account for esasurement artainties,
- c. Comparing the F computed (F C) obtaine 'in b, above to:
- 1. R(FATP) for the appropriate The F , limits or RATED TMERML measured core pla es given in , and f. be ov, and
- 2. The relationship:
F =F I [14.2( )) where F is the Itait or fr tional THERML POWER operation I expressed as a functi of F [ d P is the fraction of RATED ; THERML POWER at whi h F, was meas e d. l ule:
- d. Rameasuring F , accord ng to the following se I
- 1. When F C
is g repriate
. terthantheF[Iimitforthe relations ip, additional esasured core plane but less than the F 1 C
power distr ution maps shall be taken and F compa d to F l WF 9: a) ther within 24 hours after exceeding by 20% of TED C THERML POWER er greater, the THERML POWER at whic F was last estemined, er b) At least once per 31 EFPD, whichever occurs first. d SLMER - UNIT 1 3/4 2-5 l l
.. POWERCISTRIsuY10NLIMITS Sutyt1LLAMCE REQUIREMENTS (Continued) ,/
- 2. tAs F C g,'less than er equal to the F P limi or the app riata measund core pfane, addittenal powe distribution amps 11 he taken and F C campare'd te F F at least esca per 11 EIPO. 8
- e. r SATED THERP%L POWER (F hall be provided for The F,11alta all core planes e taining bank "O' contro rods and all unrodded care planes in a Ra tal Peaking Factor it Raport per
- e Specification 6.9.1. .
, f. The.Fg limits of e. , ve, are no applicable in the following core planes regions as asas in perc4nt of core height free the bottoe sf the fuel:
- 1. Lower core n gton from 15%, inclusive.
- 2. Upper core region fros 5 100%, inclusive.
- 3. Grid plane m gfons e 17.8 1 , 32.1 2 3 , 46.4 2 2%,
60.6 1 2% and 74.9 2%, inclus e. (17 x 17 fuel elements). ( 4. Core plane regi s within i 2% of re height (* 2.88 inches) about the bank nd position of th bank "D" control rods.
- g. With F C exceed gF the effects of F on (2) shall be evaluated m
l to detaruine i Fq (2) is within its limits. , 4.2.2.3 When Fq '(I) i asasured for other than F, detaruination an everall asasured Fq (I) shall be obtained from a power distribution map and ressed by 35 to account f annufacturing tolerances and further increased b 5% to account for asasu nt uncertainty. I l 4 l l SL30ER - UNIT 1 1/4 2-6 Ame.*nt No. 35 s
. . ~ - .
I 893 748 1275 SCtta 92 l POW (R 0151RIBU110N LIMITS h*htf 'l 1 SURVEltt.ANCE REQUIR(M[NTS 4.2.2.1 The provisions of specification 4.0.4 are not applicable. 4.2.2.2 For RAOC operation Fg (t) shall be evaluated to determine if Fq(z) is within its limit by:
- a. Using the novable incore detectors to obtain a power distribution '
map at any in!RML POWER greater than 5% of RATED THERML POVER. l
- b. Increasing the seasured Fg (t) coesponent of the power distribution map by 3% to account for manufacturing tolerances and further increasing the value by 5% to account for sessurement uncertainties.
Verify the requirements of Specification 3.,2.2 are satisfied,
- c. Satisfying theggwing relationship:
M F q(z)^< M x K( r_) f o r P > 0. 5 WPN Fq "( t ) ~c " "I*) for P c o.5
~
l W(t) x 0.5 where F (2) is the measuredgF (z) increased by the alignjes for manuf acturing tolerances and-measurement uncertainty, M is the F g limit, K(t) is given in Figure g, P is the relative THERML _ POWER, and W(t) is the cycle dependent function that accounts for power distribution transients encountered during normal operation. This function is given in the Peaking Factor Limit Report as per Speci-fication 6.9.1.5 tt. l
- d. Measuring qF "(t) according to the'following schedule:
- 1. Upon. achieving equilibrium conditions af ter exceeding by 10% or more of RATED THERML POWER, the THERML POWER at which Fg (t) was last determined,a or
- 2. At least once per 31 Ef fective Full Power Days, whichever occurs first.
"During power escalation at the beginning of each cycle, power level say be increased until a power level for extended operation has been achieved and a power distribution sep obtained.
O fu re, e.~ - u n . t J 3h 2 f
E 803 748 3275 SCEDG 03 ,, POWIR_0151RIBtfT10N L1 HITS 1#
$URVE!LLM E REQUIREMENTS (Contf rued)
- e. T'? ::_ .. n::t: S d ' i, M ' :.; Wih ne.ma.ximm vd e cif-esf4spr if F" I2) oAner hn Nru;2d[since k StTfevakiaskath+(x)ine the previous determination the following actions shall be taken: q (3)o~f F ,reely either of N
- 1) Fg (z) shall be increased by 2% over that specified in $pecifi-cation 4.2.2.2c. or N
- 2) Fg (z) shall be seasured at least once per 7 Effective full ;
Power Days untti two succes:1ve aaps indicate that de muimm va.Lc d
*eirtime YFf(2) g g,) y ys t not uee increasing.
mau f+(a) m
- f. With the relationships specified in Specification a.2.2.2c. above not being satisfled: ove& c ge h,g4(a)St-
- 1) Calculate the4 percenffq(2) exceeds its liedt by the following expression: myimum
~
. 3. -
R
- - ' - - - M Fn (t) x V(z) -1 x 300 for P 10.5
#' T T(z)
O _ .-
=
M
= l ,
;}7 {' Fn (z) x W(r) f,3,; x 300 fo r P < 0. 5
~ y se r.v g g x g(z)J' ,
l 2) One of the following actions shall be taken: I L a) Within 15 minutes, centrol the AF ( gnewAFDlimits which are determined by reducing WTts M t I 1 by 1% AFD for each percent Fg (z) exces its lialts as deter- , mined in Specification 4.2.2.2f.1). Within 8 hours, reset the AFD alars setpoints to these modified limits, or ; b) Comply with the requirements of Specification 3.2.2 for Fq (z) ~ exceeding its limit by the percent calculated above, or ' c) Verify that the requirements of Specification 4.2.2.3 for Base Load operation are satisfied and enter Base Load operation. O s- a,.a vna I
SCEtG 84 2 883 748 3275 a POWER 0151R18UT10N LIMITS SURVEILLANCE Rf0VIREMfi'_5 (Continued) 9 The lialts specified in Specificadons 4.2.2.2c, 4.2.2.2e. , and 4.2.2.2f. above are not applicable in the following core plane regions: 1.' Lower c'cre region from 0 to 15%, inclusive.
- 2. Upper core region from 85 to 103, inclusive.
U ff the 4.2.2.3 Base Load ope' ration is permitted at powers above APL following conditions are satisfied:
- a. Prior to entering 8sse Load operation, maintain THERML POWER above ,
APL O and Ie'ss than or equal to that allowed by Specification 4.2.2.2 i h for at least the orevious 24 hours. Maintain Base load operation s9 g ecMe gg surveillance (AfD within*.tMt:W target flux difference) during this time period.- Base Lead operation is then persitted providing THERML
- NU POWER ks enintained between APLand APL OL or between APL HO and 100% (whichever is most limiting) and FQ surveillance is maintained O!
pursuanttol$pecification4.2g4g APL isdefinedashm;nimum Va.k e d' g gBL , Mz % x M Z) 1),.x 103
, R,where: de kW) 'h2)*W2)BL
. g F (f) is the measured F (z) increased by the allevances for g
manufacturt59 tolerances and seasurement uncertainty. The F limit' is K t) is given in Figure . V(r)gg is the cyc1 dependent function that accounts for limited power distribution transients en-countered during base load operation. The function is given in tne Peating Factor Lleit Report as per Specification '.T.'.r.
- b. During Base' Load operation, if the THERML POVER In d'thbsed belo HO APL then the conditions of 4.2.2.3.a shall be $stisfied before reentering lBaseLeadoperation.
4.2.2.4 During Base load OperationqF (2) shall be evaluated to detertine if Fq (Z) is within its 1imit by:, f
- a. Using the movable incere detectors to obtain a power distribution sap at any fHER.ML PWER above APLND,
- b. Increasing t.he seasured Fg (Z) component of the power distribution map by 3% t'o account for manufacturing tolerances and further increasing the value by 5% to account for sensurement uncertainties.
Verify the , requirements of Specification 3.2.2 are s.atisfied. O h ew - lin . s .LI 3h 1-4s l
^
l z ses 74s s27s seus es POWER O!5TR18U710N LIMITS
\
- f. $ 6( ( .I
$URVEILLANCEREQUIREMENTA(Coritinued) !
- c. Satisfyl e(followingrelationship:
F((Z)i,,y 2 for P > APL ND where: 2.k - FltZ) is the measured q FThe F (Z).g limit is 8 4 3. i P. t-i K(2 is given in Figure 3,6=t. P is the relative THERMAL POVER. W(Z))gt i s the epcle cependent function that accounts for limite power distribution transients encountered during normal operation. This function 6't.1.E Specification ik given 11in the Penking Factor Lialt Report as per
- d. Measuring F((Z),in conjunction with target flux difference deter-mination accord lng to the following schedule:
- 1. Prior to estering BASE LOAD operation after satisfying Section -
4.2.2.3 unless a full core flux map has been taken in the previous 31 EFPD with the relative thersal power having been NO maintained above APL for the 24 hours prior to mapping, and
- 2. At least or}ce per 31 effective full pcwer days.
- e. "'"=w { ',,: 4 G:.;, with A d aKImknVa N**
_.m . c. ' verYa.CollNQI hu '- =;2[since the previous determination F (Z) either of the following actioqs shall at taken: 4
- 1. F (Z) shall be increased by 2 percent over that specified in 4.2.2.4.c,or
- 2. F (Z) shall be sessured at least once per 7 EFPD ur.til successive baps indicate that deMMimm VdC.
- p b oveM e co/a b er (2-)
sauises'g over-t. h(z) gis not increasing. i
- f. With the relationship specified in 4.2.2.4.c above not being satisfied, eithe,r of the following actions shall be taken:
- 1. Place the chre in an equilbrium condition where the limit in 4.1.2.2.c 15 satisfied, and remessure (2), or i
O %, ,m j m.a 4
2 803 748 3275 SCE*G 06 I P. F e r C ' POWER 015TRIBUT10N L1 HITS I SURVEILLANCE REQUIREMENTS (Continued)
- 2. Congly Eith the requirements of 5 ecification 3.2.2 for F (Z) 3 exceeding expres5jon: its limit by A thI[e'rNnt calculated with the fo l overtie care.he.7/d
.M[ Fl(Z), x W(Z)8l j 'j', -1 ) x 100 for P l NO
(@ea ==. 1 APL I
- g. The limits specified in 4.2.2.4.e, 4.2.2.4.e. and 4.2.2.4.f above are not appitcab1: in the following core plan regions:
- 1. Lower core region 0 to 15 percent, inclusive.
- 2. Upper c re region 85 to 100 percent, inclusive.
4.2.2.$ When Fg(Z) is! measured for reasons other than testing the requirements of specification 4.2.2;2 an overall measured F g(r) shall be obtained from a power distribution map and increased by 3% to account for manufacturing tolerances and further increased by 5% to account for atasurement uncertainty. O [ l i i I 9 Ju m me - tie. t 1 j 2/414c l l
1 I Y lj [ \ 1 i I
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a i i ! , ! l. i};i if-{j . y _ p ,. . q. y. ::p. .1;. ip.;1 : h. l . .ig i l. . _ . : jf :i 7 -l. -iry .l . .
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- :i I: I- i i -l ij, p 7. .
'. l, i !
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- p. 1, i .
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- 1. . . _ i\ l l -
l l f. O.0
/ I lYl l 1 l ;
4 6 10 12 0 2 OP BOTTOM OF CORE HEIGHT, FT. : OF FUEL : FUEL I FIGURE 3,2 2 K(Z) NORMAllZED F (Z) Q AS A FUNCTION OF CORE HEIGHT h 4.,,AJ l T& && O F & AE 5.2-( Summu. UNIT 1 3/427 ,
.,-..-.-n.,-.-_-- . - . . , - _ _ . _ - . - . ._ _
O V - 1.2 1.1 ' 1 9 -
.a
.8
, .7 3
c .6 -
- t g.5 C
i .4 3
.3 Core Heicht (ft) K_12] ,
0.0 1.0
'2 Q
(/ 60 12.0 1.9 0 925
.1 0
0 1 2 3 4 5 6 7 8 9 10 11 12 l OROM TOP ! CORE HEIGHT, FT. 3 OF OF FUEL FUEL l l FIGURE 3.2-1 K(z)- NORMAll2ED Fn(z)'AS A FUNCTION OF CORE HEIGHT l l r s Summer UNIT 1 M 2*7
POWER DISTRIBUTION LIMITS 3/4.2.3 RCS FLOW RATE AND NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR LIMITING CONDITION FOR OPERATION 3.2.3 The combination of indicated Reactor Coolant System (RCS) total flow rate and R shall be maintained within the region of allowable operation shown on Figure 9:4 3, for 3 loop operation. 3.2- 2 Where:
- a. R Fb '
??44,[1.0 + N (1.0 - P)]
b.
#*9HERMALP$dR ,
P = BTED THERXAL POWER
- c. Fh=MeasuredvaluesofFhobtainedbyusingthemovableincore detectors to obtain a power distribution map. The measured values of Fh shall be used to calculate R since Figure P.*3,J.2-2 includes measurement uncertainties of 2.1% for flow and 4% for inceremeasurementofFh,and APPLICABILITY: MODE 1. g ACTION:
With the combination of RCS total flow rate and R outside the region of accept-able operation shown on Figure M 3.2-2a _
- a. Within 2 hours either:
- 1. Restore the combination of RCS total flow rate and R to within the above limits, or
- 2. Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER and reduce the Power Range Neutron Flux - High trip setpoint to less than.or equal to 55% of RATED THERMAL POWER within the next 4 hours,
- b. Within 24 hours of initially being outside the above limits, verify through incore flux sapping and RCS total flow rate comparison that the combination of R and RCS total flow rate are restored to within the abovo limits, or reduce THERMAL POWER to less than 5% of RATED THERMAL POWER within the next 2 hours.
- c. Identify and correct the cause of the out-of-limit condition prior to increasing THERMAL POWER above the reduced THERMAL POWER limit required by ACTION items a.2. and/or b. above; subsequent POWER OPERATION may proceed provided that the combination of R and indicated RCS total flow rate are demonstrated, through incere flux mapping and RCS total flow rate comparison, to be within the realen
, g of acceptable operation shown on Figure T.fk.L, prior to exceeding the following THERMAL POWER levels: 3. 2 -2 SIM4FR - llWTT 1 1/A 7-A
) !-
l POWER DISTRIBUTION LIMITS () ACTION: (Continued) A nominal 50% of RATED THERMAL POWER, 1.
- 2. A nominal 75% of RATED THERMAL POWER, and
- 3. Within 24 hours of attaining greater than or equal to 95% of RATED THERMAL POWER.
t SURVEILLANCE REQUIREMENTS 4.2.3.1 The provisions of Specification 4.0.4 are not applicable. 4.2.3.2 The combination of indicated RCS total flow rate and R shall be l determined to be within the region of acceptable operation of Figure 3:2rG: 3242
- a. Prior to operation above 75% of RATED THERMAL POWER after each fuel loading, and
- b. At least once per 31 Effective Full Power Days.
4.2.3.3 'The indicated RCS total flow ra sh 11 be verified to be within the region of acceptable operation of Figure
^
at least once per 12 hours when g the most recently obtained,value of R obtained per Specification 4.2.3.2, is (-~j y_, assumed to exist. 4.2.3.4 The RCS total flow rate indicators shall be subjected to a CHANN.EL CALIBRATION at least once per 18 months. 4.2.3.5 The RCS total flow rate shall be determined by measurement at leest
-~
once per 18 months, i l l l . SUMMER - UNIT 1 3/4 2 9 l l 1
MEASUREMENT UNCERTAINTIES OF 2.1% FOR FLOW AND 4.0% FOR INCORE MEASUREMENT OF FN.iH ARE - INCLUDED IN THIS FIGURE 38 ACCEPTABLE UNACCEPTABLE OPERATION REGION OPERATl_ON REGION 36 34 2 2 k ~ I e 32 C 6 i2 30 9 5 E a:
' N' t "* (1 00,28 36) m 28
*~.m (1 00,28 08) 9+- '
- u. u. - 5 NOH i
(1 00[27 (100 2751) 79)i l 'Q
,, ((, (1.00.27 23)
(1.00,26 94) a 26 i 24
.9 .95 1 1.05 1.1 R = FN3s/1,56 [1.0 + 0.3(1.0 P)]
FIGURE 3.2 2 RCS TOTAL FLOW RATE VS. R THREE LOOP OPERATION NOTE. When operationg in this region. the restricted power lewels shall be cos4ered to be 1005. of f ated thermal power (RTP)f or hgwre 21 1 SUMMER UNIT 1 34210
DELE TG MEASUREMENT UNCERTAINTIES OF 2.II NFOR Fl.OW l O AND 4.0% FOR INCORE MEASUREMENT OF F ARE INCLUDEO IN THIS FIGURE
/
e 34 A EPTABLE UNAC TA 51.E l OPERA ON REGION OPERA N REGION
. y \
- t:
E E a n '
., j 5 S
- E /
i a I 3 . o -
,. N o 30
- / (1. 29.47)
$ , u ., , . ,
(1.0 29.17) m -
, __, ~
.00, 8.8 fe , SEE NOTE 1, ,; ,5 (1.00,28. 9) 28
[l[ (( (1.00,27. 26 4 24 , 1.10
.90 0.95 1.00 1.05 R E PLM Gppg R = Fj/1.49 (1.0 + 0.2(1.0 Pi]
O n. FIGURE 3.2 3 RCS TOTAL FLOW RATE VS R THREE LOOP OPEfl ATION g ppg gg ! uors. m . <. m., w e . ,i...in. . m .i.4 i.,S. en. n .
.., intri t., n o w,. 2.t 1.
pgg,g-
..a . .. 4 i. se teos .: estem ine,m.i m er o - nut? 1 1/a 2-10 Amendment No. 33. 60
O O o ; j ,. TABLE 3.2-1 ,
^
m DNB PARAMETERS h 9 LIMITS z U 3 Loops In 2 Loops in
- PARAMETER Operation Operation 5*89.8'F ,
Reactor Coolant System T,yg <SWT Pressurizer Pressure > JZ W inia *. 2206 ysart R. m h t a "Limit not applicable during either a THERMAL POWER ramp in excess of 5% of RATED THERMAL ; POWER per minute or a THERMAL POWER step in excess of 10% of RATED THERMAL POWER.
**These values left blank pending NRC approval of two-loop operation.
\ l I ' '
i ', ( O ,j. ,, T
- TABLE 3.3-2 -
h 9 REACTOR TRIP SYSTEM INSTRUNENTATION RESPONSE TIMES RESPONSE TIME 5 FUNCTIONAL UNIT _ ' Manual Reactor Trip Not Appilcable 1. 1
- 2. Power Range, Keutron Flux $ 0.5 seconds *
- 3. Power Range, Neutron Flux, 3 Not Applicable High Positive Rate
- 4. Power Range, Neutron Flux, High Negative Rate $ 0.5 seconds *
- 5. Intermediate Range, Neutron Flux NotAppilcable; ,
M*
- 6. Source Range, Neutron Flux Not'Epilcable.
g. Y
- 7. Overtemperature AT $ M secorw' e*
Overpower AT
"^^--
m
- U S 285 **C"'d5 8.
- 9. Pressurizer Pressure--Low $ 2.0 seconds
- 10. Pressurizer Pressure--High $ 2.0 seconds ,
Pressurtzer Water Level--High Not Appilcable. - ., 11. Neutron detectors arc exempt from response time testing. Response time of the. neutron flux signal portion m of the channel shall be measured from detector output or input of first electronic component in channel,
i .. . O - O . O, L TABLE 3.3-2 (Continued) - i X, REACTOR TRIP SYSTEM INSTRUNENTATION RESPONSE TIMES
~
s FUNCTIONRL tm1T RESPONSE TIME
. 12. A. Less of Flow - Single loop -
~
(Atmve P-8) ,
$,1.0 seconds
- 9. Less of Flow - Two Loops *
(Abeve P-7 and below P-8) i 1.'0 seconds - 3
- 13. Steam Generator 1Anter Level--Low-tow < 2.0 seconds
- 14. Steam /Feedwater Flow Miseetch and ',
- Low Steam Generator Water Level Not Appitcabid
{ 15. Underveltage-Reactor Coelant Pumps i 1.5 sec .. ! Y 16. Underfrequency-Reacter Coelant Pumps i jkg secondsj. g . , .
- 17. Turbine Trfp .
A. Low Fluid 011 Pressure Not Applicable B. Turbine Step Valve Closure Not Appitcable r
- 18. Safety injectfen input from ESF , Not Applicable
- 19. Reacter Trip System Interlocks Not Applicable i
- 20. Reacter Trfp Breakers Not Appilcable -
- 21. Automatic Trip Logfc Not App 11 cable .
b
~F
[s .^. , Q
l 1 l
; 3/4.5 EMERGENCY CORE COOLING SYSTEMS l
3 /4. 5.1 ACCLMULATORS
\
LIMITING COCTTION T0X OPERATION , l l l ( 3.5.1 Each reactor t:oolant systes accoulater shall be CPE1 TABLE with:
- a. The isof ation valve open, gf
- b. A contained borated water volse of between 7p and 7)iG gallons, i c. A boren concentration of between 1200 and 2500 ppe, and
- d. A nitrogen cover pressure af between 600 and 656 psig.
APPtit u ftTTY: EDES 2, I and 3.8 ACTION: l
- a. With one acessulater truyerable, except as a result of a closed isolation valve, restore .the inoperable accumulator to OPERABLE status within ena Scur ur te in at least ICT STANDBY within the next 6 hours and in H:TT .5 HUT 00mH within the following 6 hours.
.. b. With one accumulator ineperable due to the isolation velve being closed, either imediately open the isolation valve or be in at laast MT STACSY within one hour and in HOT SHUTDOW within the following 12 hours.
SURVEILLANCE REQUIkEMENTS 4.5.1.1 Each accoulator shall be seasonstrated CPERABLE:
- a. At taast once per 12 hours by:
- 1. Verifying the contained borated watar volume and nitrogen sever presure in the tank.s. and
- 2. Verifying that each accoulator isolation valve is open.
"Pressurizer pressure aboys 1000 psig.
- O SumIR -IMIT 1 3/4 5 1
I SPECIAL TEST EXCEPTIONS 3/4.10.2 GROUP HEIGHT, INSERTION AND POWER DISTRIBUTION LIMITS LIMITING CONDITION FOR OPERATION 3.10.2 The group height, insertion and power distribution limits of Specifi-cations 3.1.3.1, 3.1.3.5, 3.1.3.6, 3.2.1 and 3.2.4 may be suspended during the performance of PHYSICS TESTS provided:
- a. The THERMAL POWER is maintained less than or equal to 85% of RATED THERMAL POWER, and
- b. The limits of Specifications 3.2.2 and 3.2.3 are maintained and determined at the frequencies specified in Specification 4.10.2.2 below.
APPLICABILITY: MODE 1 ACTION: With any of the limits of Specifications 3.2.2 or 3.2.3 being exceeded while the requirements of Specifications 3.1.3.1, 3.1.3.5,.1.1.3.6, 3.2.1 and 3.2.4 are suspended, either: J
- a. Reduce THERMAL POWER sufficient to satisfy the ACTION require- 4 ments of Specifications 3.2.2 and 3.2.3, or
- b. Be in HOT STANDBY within 6 hours. ;
i. , SURVEILLANCE REQUIREMENTS I 4.10.2.1 The THERMAL POWER shall be determined to be less than or equal to 85% of RATED THERMAL POWER at least once per hour during PHYSICS TESTS. l . l. 4.10.2.2 The Surveillance Requirements of the below listed Specifications (0. M b-) l shall be erformed at least once per 12 hours during PHYSICS TESTS: C eL o9. +.2.1.Q M n, f. 2. 2 . 5.
- a. A Specifications 4.2.2.2 and,N hd SF
- b. Specification 4.2.3.2. ,
O SUMMER - UNIT 1 3/4 10-2
EMERGENCY CORE COOLING SYSTEMS SURVEILLANCE REQUIREMENTS (Continued) l
. i.
- h. By performing a flow balance test, during shutdown, following completion of modifications to the ECCS subsystems that alter the subsystem flow characteristics and verifying that: ,
- 1) For centrifugal charging pump lines, with a single pump running)( omd. M acM,im :
a) Thesumoftheinjectionlineflowrates,excludingthe highest flow rate, is greater than or equal to 3M.gpm, and 337 b) The total pump flow rate is less than or equal to 680 gpm.
- i. By performing a flow ~ test, during shutdown, following completion of modifications to the ECCS subsystems that alter the subsystem flow characteristics and verifying that:
- 1) For residual heat removal pump lines, with a single pump running thesumoftheinjectionlineflowrates.isgreaterthanor equal to 3663 gpm. -
t 4 1 SUMMER - UNIT 1 3/4 5-6 l l
REACTIVITY CONTROL SYSTEMS gl 1 W BASES i MODERATOR TEMPERATURE C0 EFFICIENT (Continued) involved subtracting the incremental change in the EC associated with a core ) ! condition of all rods inserted (most positive MDC) to an all rods withdrawn l condition and, a conversion for the rate of change of moderator density with I temperature at RATED THERML POWER conditions.
~
This value of the M C was then transformed into the Itaiting MTC value -[8x 10~4 delta k/k/*F. The MTC value of N x 10~4 delta k/k/*F represents a conservative value (with corree- 1 i tions for burnup and soluble boron) at a core condition of 300 ppm equilibrium boron concentra ion and is obtained by making these corrections to the limiting MTC value of x 10 ~4 k/k/*F. The surveillance requirements for measurement of the MTC at the beginning and near the end of the fuel cycle are adequate to confirm that the MTC resains within its limits since this coefficient changes slowly due principally to the reduction in RCS boron concentration associated with fuel burnup. 3/4.1.1.4 MINIMUM TEMPERATURE FOR CRITICALITY This specification ensures that the reactor will not be made critical with the Reactor Coolant Systen average temperature less than 551*F. This limitation is required to ensure 1) the moderator temperature coefficient is within its analyzed temperature range, 2) the protective instrumentation is within its normal operating range, 3) the pressurizer is capable of being in an OPERABLE status with a steam bubble, and 4) the reactor pressure vessel is above its minimus RTNDT temperature. 3/4.1.2 80 RATION SYSTEMS The boren injection system ensures that negative reactivity control is available during each mode of facility operation. The components required to ) perform this function include 1) borated water sources, 2) charging pumps, -
- 3) separate flow paths, 4) boric acid transfer pumps, and 5) an energency power supply free OPERABLE diesel generators.
With the RCS average tempe ature above 200'F, a minimum of two boron in-jection flow paths are required to ensure single functional capability in the 1 event an assumed failure renders one of the flow paths inoperable. The borativn capability of either flow path is sufficient to provide the required SHUTDOW T O SWNER - Uh!T 1 8 3/4 1-2
4 l REACTIVITY CONTROL SYSTEMS O BASES l BORAT!0N SYSTEMS (Continued) bW J fQ$@ MARGIN from expected op g conditions 1.77% delta k/k or as required by Figure 3.1-3 after apnon decay and Sod 1down to 200'F. The maximum expected boration capabilf tv/ requirement 3efurs from full power equilibrium xenon condi-tions and is C ? b d " ' i gallons of 7000 ppa borated water from the beric acid storage a r > ',^Mallons of 2300 ppe borated water from the refuel-ing water storage tank
,973,gg g With the RCS temperature below 200*F, one injection systed is acceptable without single failure consideration on the basis of the stable reactivity condition of the reactor and the additional restrictions prohibiting CORE ALTERATIONS and positive reactivity changes in the event the single injection systas becomes inoperable.
The limitation for a maxism of one centrifugal charging pump to be OPERABLE and the Surveillance Requirement to verify all charging pumps except the required OPERABLE ' pump to be inoperable below 275'T provides assurance that a mass addition pressure transient can be relieved by the operation of a single PORV. A n,v/c) The boron capability required below 200'F is sufficient to provide the i avir e ex r ia r 3.1-3 / O a#4r e swutoow" a^act" r t a rc #t < it */* r af ter xenort_ decay and cooldown from 200'F to 140'F. This condition is s gallons of 7000 ppe borated water from the boric acid sto ag d> i l _eithe l anks or allons of 2300 ppe borated water from the refueling water storage tank. 3'7 900 2,*700 The contained, water volume limits include allowance for water not available because of discharge line location and other physical characteristics. The OPERABILITY of one boron injection systas during REFUELING ensures that this system is available for reactivity control while in MODE 6. 3/4.1.3 W)VABLE C0KTROL ASSDeLIES The specifications of this section ensure that (1) acceptable power distribution Itaits are maintained (2) the minimus SHUTDOWN MARGIN is main-tained, and (3) limit the potential effects of rod misalignment on associated accident analyses. OPERA 81LITY of the control red position indicators is required to determine control rod positions and thereby ensure compliance with the control rod alignment and insertion limits. O SUP94ER - UNIT 1 8 3/4 1-3
eaes? 10 s to.006 oc2 3/4.2 POVER DISTRIBUTION LIMIVS BASES The specificatto s of this section provib assurance of fuel ntegrity uring Condition I (No al Operation) and II (L cidents of Moderat Frequency) ents by: (a) saintai ing the minimum DN8R in he core greater th or equal to 1.30 during normal o ration and in short to transients, and (b limiting the fission gas release, uel pellet temperature d cladding mechant 1 prop rties to within ass d design criteria. In ddition, limiting t peak SEPL AC E linea power density durin condition I events prov das assurance that a gg initia conditions assianed r the LOCA analyses are met and the ECCS ac eptance criteri limit of 2200*F is t exceeded. L s E R.T The finitions of certai het channel and peaki factors as used in 3 these spec fications are as fol : F9 (Z) He t Flux Hot char.nel F ctor, is defined as the maximus local hea flux on the surface of a fuel red at core e ovation Z divided by t average fuel rod at flux, allowing for nufacturing toleran on f 1 pellets and rods. Fh Nuclear nthalpy Rise Hot Ch nnel Factor, is define as the ratie of the int al of itnear power long the rod with the ighest integrated ; power to average rod powe ! F,Y(Z) Radia'i. Peak Factor, is defin as the ratto of peak owtr densit to average r density in the rizontalplaneatcoreelevationf. 3/4.2.1 AXIAL FLUX O!FFERENCE , The limits on AL FLUX O!FFE E (AFD) ssure that t Fg(Z) upper bound envelope of 2.25 t s the normali ed axial king factor s not exceeded [ during either no 1 operation o in the eve of xenon re istribution folloding REPt. A c E power changes. , Target f,ux difference J determined at equilibri xenon conditions The TcSEET full length cods may be po ccordance with their l respective / insertion 11s1%sitioned andshouldpeinserted with% the core r theirin normal position 9 for steady state operati n at high poyar levels. T: value of the target flux > differspeobtained r these con 4ftions divided y the fraction of RATED j THERMAL' POWER is the rget flux di ference at RA THERMAL POWER for the assoc (ated core bu conditions Target flux ifferences for othd THERMAL POWyAlevelsare ined by mul plying the ED THERMAL POWER value by the appropriate fyactional THE L POWER level. The periodic ting of the prgetfluxdifprencevaluei necessary to eflect core burnup onsiderations. Co wf i u u.E.t o u N e.x r Pu E k 9 StMtfR - UNIT 1 B 3/4 2-1
POWER DISTRIBUTION LIMIT BASES _, AXIAL FLUX DIFFERENCE (Continued) Although i is inten d that t plant w I be oper ed with he AFD within the + target b nd about e target lux diffe ence, du ng rapid plant THERMX POWER re etions, ntrol rod motion wil cause t AFD to 4 deviate out de of th target b d at red ed THERMA POWER le els. Thi deviation ill not a ect the enon redi ribution fficient to chan e the p envelope f peaking factors ich may b reached o a subse ent returp to RATED T RMAL POW (with t . AFD with n the targ t band) p ovided the time hw durati of the viation limited. Accordin y, a 1 ho r penait eviation limit umulativ during e previous 24 hours i provide 4 for oper ion outside q of t e target and but w. thin the mits of F ure3.2-ywhileat ERMAL POW R levels etween 5 RATED TH L POWE . For THE L POWER 1 elsbetween15%and)$and90% 50% of RATED THERMAL OWER, de ations o the AFD tside of/the targe,t' band are /ess signif cant. Th penalty o 2 ht;urs 4 ctual t e reflects this reduded signifi ance. N i / Q PyvisionsformonitorngtheAFD n an auto atic basi are de"ived from the plant process computer through th AFD Moni r Alarm, he computer deter . mined the one Minute aver ge of each of the OP ABLE exco detecto* tputs WV and'provides sage imme ately if e AFD for t least 2 more OpERABLEexco,analarm re channels are outsi e the tar t band an the THERMAiPOWER i Jreater than 90% of ED THERMA POWER. D ing opera on at THE L POWER flevelsbet, ween 50% d 90% and tween 15% nd 50% RA" 0 THERMAL WER, th , computer / outputs a alarm mess ge when th penalty d iation act ulates ; beyond the limits f I hour a d 2 hours, espective . Figure B 3 4 2-1 show a typical onthly tar t band.
~'
I' i 3/4.2.2 and 3/4.2.3 HEAT FLUX HOT CHANNEL FACTOR and RCS FLOWRATE AND NUCLEAR ENTHALPY RI5E h0T CHANNEL FACTOR The limits on heat flux hot channel factor, RCS flowrate, and nuclear , enthalpy rise hot channel factor ensure that 1) the design limits on peak , local power density and minimum DNBR are not exceeded and 2) in the event of a LOCA the peak fuel clad temperature will not exceed the 2200'F ECCS acceptance criteria limit. Each of these is measurable but will normally only be determined periodically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient to insure that the limits'are maintained provided:
- a. Control rods in a single group move togetner with no individual rod insertion differing by more than + 13 steps, indicated, from the
~
group demand position. O b. Control rod groups are sequenced with overlapping groups as described in Specification 3.1.3.6. SU!HER - UNIT 1 B 3/4 2-2 l
INSERT 3 g The specifications of this section provide assurance of fuel integrity during Condition I (Nomal Operation) and II (Incidents of Moderate Frequency) events by: (1) maintaining the calculated DtGR in the core at or above the design limit during nomal operation and in short-tem transients, and (2) limiting the fission gas release, fuel pellet temperature, and cladding acch'anical properties to within assumed design criteria. In addition, limiting the peak linear power density during Condition I events provides assurance that the initial conditions asgumed for the LOCA analyses are met and the ECCS acceptance criteria limit of 2200 F is not exceeded. The definitions of certain hot channel and peaking factors as used in these specifications are as follows: F9(Z) Heat Flux Hot Giannel Factor, is defined as the maximum local heat flux on the surface of a fuel rod at core elevation Z divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods; Ffg Nuclear Enthalpy Rise Hot Channel Factor, is defined as the ratio of the
'ntegral of linear power along the red with the highest integrated power to the average rod power.
O 6 0
l p
, C INSERT 4 Be limits on AXIAL' FLUX DIFFERECE ( AFD) assure that the Fn(Z) upper bound envelope of 2.45 times the nomalized axial peaking factor Is not exceeded during either nomal operation or in the event of xenon redistribution following power changes. _
he limits on AFD will be provided in the Peaking Factor Limit Report (PFLR) per Technical Specification 6.9.1.11. Target flux difference is detemined at equilibrium xenon conditions. The full-length rods may be positioned within the core in accordance with their respective insertion limits and should be inserted near their nomal position for steady-state operation at high power levels. The value of the target flux difference obtained under these conditions divided by the fraction of RATED THEMAL POWER is the target flux difference at RATED THEFFAL POWER for the associated core burnup conditions. Target flux differences for other THEMAL POWER levels are obtained by cultiplying the RATED THEFF.AL POWER value by the appropriate fractional THEFF.AL POWER level. Be periodic updating of the target flux difference value is necessary to reflect core burnup consideraticns. ND At power levels below APL , the limits on AFD are defined in the PFLR consistent with the Relaxed Axial Offset Control (RAOC) operating procedure and limits. These limits were calculated in a manner such that expected operational transients, e.g. A load follow operations, would not result in the AFD deviating outside of those (^,) limits. However, in the event such a deviation occurs, the short period of time allowed outside of the limits at reduced power levels will not result in significant i xenon redistribution such that the envelope of peaking factog would change ' sufficiently to prevent operation in the vicinity of the APL power level. ND - At power levels greater than APL , two modes of operation are pemissible; 1) ' RAOC, the AFD limit of which are defined in the PFLR and 2) Base Load operation, I which is defined as the maintenance of the AFD within P{fDR specifi ations band about a target value. Re RAOC operagg procedure above APL is the same as that defined for operation below AFL . However, it is possible when following extended lead following maneuvers that the AFD limits may result in restrictions in the i maximum allowed power or AFD in order to guarantee operation with F (z) less than its limiting value. To allow operation at the maximum permissible wer level the Base Load operating procedure restricts the indicategDAFD to relativgy small target , band (as specified in the PFLR) and power swings (APL i power 5 APL or 100% Rated Remal Power, whichever is lower) . For Base Load operation, it is expected that the plant will operate within the target band. Operation outside of the target band for the short time period allowed will not result in significant xenon redistributi an such that the envelope of peaking factors would change sufficiently to prohibit continued operation in the power region defined above. To assure there is no residual xenon redistribution impact from past operation onge Base Load operation, a 24-hour waiting period at a power level above APL and allowed by RAOC is necessary. D2 ring this time period load changes and rod motion are restricted to that allowed by the Base trad procedure. After the waiting pericd extended Base Load operation is pemissible. l
The computer determines the one minute average of each of the OPERABLE excore . detector outputs and provides an alarm message immediately if the AFD for at least 2 of 4 or 2 of 3 OPERABLE excore channels are: 1) outside the allowed delta-I power operating space (for RAOC operation), or 2) outside the allowed delta-I target band (for Base Lead operation) . These alarms are active when g'ger is greater than: 1) 50% of RATED THERMAL POWER (for RAOC operation), or 2) APL " (for Base Load operation). Penalty deviation minutes for Base Load operation are not accumulated based on the short period of time during which operation outside of the target band is allowed. O O
REPL A CE (dlTH Aew OI M TIO M S. B 3 y .2 - 3 .g U- Ow v v 7%s . i l1 (Nb l C,bA Mje. M ffc3 tN s Figuax FRem m h spec 3 l i-i
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20 2
- -20 -10 0 10 INDICATED AXIAL FLUX DIFFERENCE (%)
FIGURE B 3/4 21 INDICATED AXIAL FLUX OlFFERENCE s VERSUS THERMAL POWER O
- B 3/4 2-3 EUMMER - UNIT 1
i O . o T ,) 6 py v4 M bW h . . O O uur 1 su.mm ez 3 w 2-3
POWER DISTRIBUTION LIMIT 1 BASES HEAT FLUX HOT CHANNEL FACTOR and RCS FLOWRATE and NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR (Continued) ,
- c. The control rod insertion limits of Specifications 3.1.3.5 and 3.1.3.6 are maintained. _
- d. The axial power distribution, expressed in terms of AXIAL FLUX DIFFERENCE, is maintained within the limits.
N F 3H willbemaintainedwithinitslimitsgovidedconditionsa.through
- d. above are maintained. As noted on Figure 474=4s RCS flow rate and l F
H may be "traded off" against one another (i.e., a low measured RCS flow rateisacceptableifthemeasuredFhisalsolow)toensurethatthe calculated DNBR will not be below the design DNBR value. The relaxation of Fh as a function of THERMAL POWER allows changes in the radial power shape for all permissible rod insertion limits. 3.Z-2 R, as calculated in 3.2.3 and used in Figure M accounts for F H less than or equal to g This value is used in the various accident analyses where F 3g influences parameters other than DNBR, e.g., peak. clad temperature and thus is the maximum "as measured" value allowed. Fye(rodbowigreduces he value fONBr(io. Cred' isavailbleto generic e generic design mar ins, kOk offs t this redu to ling 9.1% ion in t BR, compl ely offs ( argin. any rod ow penait es." This margin
- idM cludes the oilowing
+ 1) sign limi NBR of 1. vs. 1.2 2 59 GridSpacig(K)of0046vs.[00.038vs.,d.059
) Thermal ffusi8nCo ficient o
- 4) DNBR H tiplier of .86 vs. O. 8 3f 5) Pite reduction OY The appli able value of rod bow pe alties is r ferenced i the FSA,R.
When an F measurement is taken, an allowance for both experimental error q anc manufacturing tolerance must be made. An allowance of 5% is appropriate for a full core map taken with the incore detector flux mapping system and a g 3% allowance is appropriate for manuf acturing tolerance, th ra 1 pe king t yhZ) is ure peri al to h vi g ass anc sthat he h ch nel actk.F( n rema ns wi in i i li t. e g
"Th ericdrgins 1soodetthdenait asso ed w he therma Td A J,gg,redu on in t 45 ueedf^menem o tny eenny $
l SUMMER - UNIT 1 B 3/4 2-4
l POWER DISTRIBIJTION LIMIT BASES l HEAT FLUX HOT CHANNEL FACTOR and RCS FLOWRATE and NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR (Continued) _ lim t for ated hers 1 Po r(F TP) x as rovi d in he Ra ial P) kin gf Fa or Li it Re rt p spe ific tion 9.1.1 was eters ned f a exp cteI i' powe cent 1 man vers over he 1N11 r ge o burn con itions in th core. h+M N When RCS flow rate and Fg are measured, no additional allowances are necessary prior to comparison with the limits of Figures 3.2-3. Measurement P9- N errors of 2.1% for RCS total flow rate and 4% for F g have been allowed for s 1 in determining the limits of Figure 3.2-3. The 12 hour periodic surveillance of indicated RCS flow is sufficient to detect only flow degradation which could lead to operation outside the acceptable region of operation shown on Figure 3.2-3. 3/4.2.4 OUADRANT POWER TILT RATIO The quadrant power tilt ratio limit assures that the radial power distribution satisfies the design values used in the power capability analysis. Radial power distribution measurements are made during startup s testing and periodically during power operation. The limit of 1.02, at which corrective action is required, provides DNB and linear heat generation rate protection with x-y plane power tilts. A limiting tilt of 1.025 can be tolerated before the margin for uncertainty in . F is deplet.ed. The limit of 1.02 was selected to provide an allowance for 9 the uncertainty associated with the indicated power tilt. The two hour time allowance for operation with a tilt condition greater than 1.02 but less than 1.09 is provided to allow identification and cor-rection of a dropped or misaligned control rod. In the event such action does not correct the tilt, the margin for uncertainty on Fg is rei,u,tated by reducing the maximum allowed power by 3 percent for each percent of tilt in excess of 1.0. For purposes of monitoring QUADRANT POWER TILT RATIO when one excore detector is inoperable the movable incore detectors are used to confirm that the normalized syneetric power distribution is consistent with the QUADRANT POWER TILT RATIO. The incore detector monitoring is done with a full incere flux map or two sets of 4 syneetric thimbles. These locations are C-8, E-5, E-11, H-3, H-13, L-5, L-11, N-8. 3/4.2.5 DNB PARAMETERS The limits on the DNB related parameters assure that each of the , parameters are maintained within the normal steady state envelope of operation ; assumed in the transient and accident analyses. The limits are consistent SUW ER - UNIT 1 B 3/4 2-5
INSERT 5 O-Margin is maintained between the safety analysis limit ENER and the design limit Et!BR. Thi:: margin is more than sufficient to offset any rod bow penalty and transition core penalty. The remaining margin is available for plant design flexibility. O O
t I?EERT 6 O . The hot channel and height dependentfactor powe [h(z) factor is measured appropriate periodically to either RAOC or and Baseincreased by a cycle Load operation, t provide assurance that the limit on .the hot channel factor, W(z) er Fn(z) is met W(z)B.L,W(z) accounts for the effects of normal operation transients and was datemined from expected power control maneuvers over the full rarge of burnup conditions in the core. W(z) accounts for the more restrictive operating limits allowed by Base Load operatio$gwhich result in less severe transient values. The functions described above for normal operation are provided in the W(z) Peaking and W(z)bbt Report per Specification 6.9.1.11. Facto O . i { I O x 1 I i
" ~ ~'
)
3/4.5 EMERGENCY CORE COOLING SYSTEMS BASE 5 3/4.5.1 ACCUMULATORS The OPERABILITY of each Reactor Coolant System (RCS) accumulator ensures that a sufficient volume of borated water will be immediately forced into the reactor core through each of the cold legs in the event the RCS pressure falls below the pressure of themechanism accumulators. duringThisinitialsurgeofwaterinto(thecoreprovides the initial cooling large RCS pipe ruptures. htADD 7 } The limits on accumulator volume, L ,ron concentration and pressure ensure that the assumptions used for accumulator injection in the safety analysis are met. The accumulator power operated isolation valves are considered to be "operating bypasses" in the context of IEEE Std. 279-1971, which requires that bypasses of a protective function be removed automatically whenever permissive conditions are not met. In addition, as these accumulator isolation valves fail to meet single fai. lure criteria, removal of power to the valves is required. The limits'for operation with an accumulator inoperable for any reason
- except an isolation valve closed minimizes the time exposure of the plant to a l'
' LOCA event occurring concurrent with failure of an additional accumulator which may result in unacceptable"peak cladding temperatures. If a closed i isolation valve cannot be immediately opened, the full capability of one J accumulator is not available and prompt action is required to place the reactor in a mode where this capability is not required. l I l 3/4.5.2 and 3/4.5.3 EMERGENCY CORE COOLING SYSTEM (ECCS) SUBSYSTEMS ( ' The OPERABILITY of two independent ECCS subsystems ensures that sufficient emergency core cooling capability will be available in the event of a LOCA ,i assuming the loss of'one subsystem through any single failure consideration. ! Either subsystem operating in conjunction with the accumulators is capable of supplying sufficient core cooling to limit the peak cladding temperatures within acceptable limits for all postulated break sizes ranging In from the addition, double ended break of the largest RCS cold leg pipe downward. each ECCS subsystem provides long term core cooling capability in the recirculation mode'during the accident recovery period. With the RCS temperature below 350*F, one OPERABLE ECCS subsystem is acceptable without single failure consideration on the basis of the stable reactivity condition of the reactor and the limited core cooling requirements. SUMER - UNIT 1 8 3/4 5-1 ;
ItGERT 7 . In addition, the borated water serves to limit the maximum power which may be reached during large secondary pipe ruptures. . 0 O e e O l l
ADNINISTRATIVE CONTROLS O e. Type of container (e.g., LSA, Type A, Type B, Large Quantity), and
- f. Solidification agent (e.g., cement, urea formaldehyde).
The radioactive affluent release reports shall include unplanned releases from site to unrestricted areas of radioactive materials in gaseous and liquid 1 effluents on a quarterly basis. - The radioactive affluent release reports shall include any changes to the Process Control Program (PCP) made during the reporting period. MONTHLY OPERATING REPORT 6.9.1.10 Routine reports of operating statistics and shutdown experience, in-ciuding documentation of all challenges to the PORV's or safety valves, shall be submitted on a monthly basis to the Director, Office of Resource Management, [ U.S. Nuclear Regulatory Commission, Washington, D.C. 20555, with a copy to the Regional Office of Inspection and Enforcement, no later than the 15th of each month following the calendar month covered by the report. Any changes to the OFFSITE DOSE CALCULATION MANUAL shall be submitted with the Monthly Operating R6 port within 90 days in which the change (s) was made effective. In addition, a report of any major changes to the radioactive waste treatment
- systems shall be submitted with the Monthly Operating Report for the period in which the evaluation was reviewed and accepted as set forth in 6.5 above, b M L PEAKING FACTOR LIMIT REPORT limit for Rated Thermal Power (F P) sha'll be pr ' ed to 6.9.1.11 or of the Regional Office of Insp n and Enforcement, the Regional Adminis Q with a copy to the Direct uclear Reactor Regulat , Attention Chief of the onnission, Washington, D.C.
Core Performance Branch, U. S. ear Regulato 2 20555 for all core planes containing control rods and all unre.1ded core ! W criticality. In the event that 4 j planes at least 60 days prior to cyc the limit would be subsitted at other time ing core life, it shall be submitted 60 days prior to date t,he limit would e effective unless tu l V otherwise exempted b e Commission. RTP Any inform ' n needed to support F will be by request from the d ne ot be included in this report SPECIAL REPORTS 6.9.2 Special reports shall be submitted to the' Regional Administrator of the Office of Inspection and Enforcement Regional Office within the time period specified for each report. 6.10 RECORD RETENTION In addition to the applicable record retention requirements of Title 10, Code O of Federal Regulations, the follewing records shall be retained for at least the minimum period indicated. Sui'ER - UNIT 1 6-18
- - . ~ . . . . ..
+ INSERT 8 ,,
- 6.9 1.11 PEAKING FACTOR LIMIT REPORT .
Be NgFD limits, the W(z) Functions for RAOC and Base Lead operation and the value for APL (as required) shall be established for each reload core and implemented prior to use. - he methodology used g generate the W(z) functions for RAOC and Base Load Operation and the value fer APL shall be those previously reviewed and approved by the NRC#, If changes to these methods are deemed necessary they will be evaluated in 4 accordance with 10CFR50.59 and submitted to the NRC for review and approval prior to
~ their use if the change is determined to involve an unreviewed safety question or if i such a change would require amendment of previously submitted docmentation.
s the W z) functions for RAOC and Base load A report containing operation and the valuetheforAFD APLliTD (,as requ(ired) shall be provided to the NRC l doc ment control desk with copies to the regional administrator and the resident , inspector within 30 days of their implementation. ; D and APL will be by request from I
'the AnyNiC information and needneeded to support not be included in W(z), W(z)ht.
this rep b i O l lf i
'WCAP-10210 P-A "Relaxation of Constant Axial Offset Control-F q Surveillance Technical Specification".
O
ATTACHMENT 3 , NON-LOCA ACCIDENT ANALYSIS , FOR THE - V. C. SUMMER PLANT TRANSITION TO 17x17 VANTAGE 5 FUEL i O !~ i l 1 l O 00681:6/880504
1 TABLE OF CONTENTS o
- Description Page Section 15.0 ACCIDENT ANALYSES 15.0-1 15.1 CONDITION I - NORMAL OPERATION AND OPERATIONAL TRANSIENTS _
15.1-1 15.1.1 Optimization of Control Systems 15.1-3 15.1.2 Initial Power Conditions Assumed in Accident Analyses 15.1-3 15.1.2.1 Power Rating 15.1-3 15.1.2.2 Initial Conditions 15.1-4 15.1.2.3 Power Distribution 15.1-5 15.1.3 Trip Points and Time Delays to Trip Assumed in Accident Analyses 15.1-5 15.1.4 Rod Cluster Control Assembly Insertion Characteristic 15.1-7 15.1.5 Reactivity Coefficients 15.1-8 15.1.6 Fissior. Product Inventories , 15.1-9 15.1.7 Residual Decay Heat 15.1-9 /7 Fission Product Decay 15.1-9 V 15.1.7.1 15.1.7.2 Decay of U-238 Capture Products 15.1-10 i 15.1.7.3 Residual Fissions 15.1-11 j 15.1.7.4 Distribution of Decay Heat Following i Loss of Coolant Accident 15.1-12 _. l 15.1.8 Computer Codes Utilized 15.1-12 15.1.8.1 FACTRAN 15.1-13 15.1.8.2 LOFTRAN 15.1-14 ; 15.1.8.3 LEOPARD 15.1-14 15.1.8.4 TURTLE 15.1-15 15.1.8.5 TWINKLE 15.1-15 15.1.8.6 THINC _ 15.1-15 15.1.9 References 15.1-16 l 15.2 CONDITION II - FAULTS OF MODERATE FREQUENCY 15.2-1 15.2.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical Condition 15.2-3 15.2.1.1 Identification of Causes and Accident Description 15.2-3 15.2.1.2 Analysis of Effects and Consequences 15.2-5 O issav:1o/ose2ss 15.0-i
TABLEOFCONTENTS(Cont) Section Description Page 15.2.1.3 Results 15.2-7 15.2.1.4 Conclusions 15.2-7 15.2.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal at Power - 15.2-8 15.2.2.1 Identification of Causes and Accident Description 15.2-8 15.2.2.2 Analysis of Effects and Consequences 15.2-10 15.2.2.3 Results 15.2-11 15.2.2.4 Conclusions 15.2-14 15.2.3 Rod Cluster Control Assembly Misoperation 15.2-15 15.2.3.1 Identification of Causes and Accident Description 15.2-15 15.2.3.2 Analysis of Effects and Consequences 15.2-17 15.2.3.3 Results 15.2-18 15.2.3.4 Conclusions 15.2-20 15.2.4 Uncontrolled Baron Dilution 15.2-22 15.2.4.1 Identification of Causes and Accident Description 15.2-22 15.2.4.2 Analysis of Effects and Consequences 15.2-24 15.2.4.3 Conclusions 15.2-28 15.2.5 Partial Loss of Forced Reactor Coolant Flow 15.2-31 15 2.5.1 Identification of Causes and Accident Description 15.2-31 15.4.5.2 15.2.5.3 Analysis of Effects and Consequences Results 15.2-32 15.2-33 h 15.2.5.4 Conclusions 15.2-33 15.2.6 Startup of an Inactive Reactor Coolant Loop 15.2-34 15.2.6.1 Identification of Causes and Accident Description 15.2-34 _ 15.2.6.2 Analysis of Effects and Consequences 15.2-35 15.2.6.3 Results 15.2-36 15.2.6.4 Conclusions 15.2-36 15.2.7 Loss of External Electrical Load and/or Turbine Trip 15.2-37 15.2.7.1 Identification of Causes and Accident Description 15.2-37 15.2.7.2 Analysis of Effects and Consequences 15.2-38 15.2.7.3 Results 15.2-40 15.2.7.4 Conclusions 15.2-41 15.2.8 Loss of Normal Feedwater 15.2-43 15.2.8.1 Identification of Causes and Accident Description 15.2-43 15.2.8.2 Analysis of Effects and Consequences 15.2-44 6 1322v:10/ose2sa 15.0-ii .
l TABLEOFCONTENT.}(Cont)
. Section Description Page i
15.2.8.3 Results
- 15.2-45 15.2.8.4 Conclusions 15.2-46 15.2.9 Loss of Offsite Power to the Station Auxiliaries (Station Blackout) 15.2-47 15.2.9.1 Identification of Causes and Accident Description 15.2-47 15.2.9.2 Analysis of Effects and Consequences 15.2-48 15.2.9.3 Conclusions 15.2-49 15.2.10 Excessive Heat Removal Due to Feedwater System Malfunctions 15.2-50 15.2.10.1 Identification of Causes and Accident Description 15.2-50 15.2.10.2 Analysis of Effects and Consequences 15.2-50 15.2.10.3 Results 15.2-52 15.2.10.4 Conclusions 15.2-53 15.2.11 Excessive Load Increase Incident 15.2-54 15.2.11.1 Identification of Causes and Accident Description 15.2-54 15.2.11.2 Analysis of Effects and Consequences 15.2-54 15.2.11.3 Results 15.2-56 15.2.11.4 Conclusions 15.2-57 f' 15.2.12 15.2.12.1 Accidental Depressurization of th'e Rea: tor Coolant System .
Identification of Causes and Accident Description 15.2-58 15.2-58 15.2.12.2 Analysis of Effects and Consequences 15.2-58 15.2.12.3 Results 15.2-59 15.2.12.4 Conclusions 15.2-59 , i 15.2.13 Accidental Depressurization of the l Main Steam System 15.2-60 i 15.2.13.1 Identification of Causes and Accident Description 15.2-60 , 15.2.13.2 Analysis of Effects and Consequences, 15.2-61 15.2.13.3 Results 15.2-63 15.2.13.4 Conclusions 15.2-63 15.2.14 Spurious Operation of the Safety injection System at Power 15.2-64 15.2.14.1 Identification of Causes and Accident Description 15.2-64 15.2.14.2 Analysis of Effects and Consequences 15.2-66 15.2.14.3 Results 15.2-67 15.2.14.4 Conclusions 15.2-67 15.2.15 References 15.2-68 l 1332v;1o/eso2ss 15.0-iii l
1 l TABLE OF CONTENTS (Cont) l t Section Description Page 15.3 CONDITION III - INFREQUENT FAULTS 15.3-1 15.3.2 Minor Secondary System Pipe Breaks 15.3-2 [ 15.3.2.1 Identificatir , of Causes and Accident Description 15.3-2 l 15.3.2.2 Analysis of cffects and Consequences - 15.3-2 15.3.2.3 Conclusions 15.3-2 l ) i 15.3.3 Inadvertent Loading of a Fuel Assembly Into An i Improper Position 15.3-3 l 15.3.3.1 Identification of Causes and Accident Description 15.3-3 s 15.3.3.2 Analys4.s of Effects and Consequences 15.3-4 15.3.3.3 Results 15.3-4 15.3.3.4 Conclusions 15.3-5 l l l 15.3.4 Compicte Loss of Forced Reactor Coolant Flow 15.3-6 , l 15.3.4.1 Identification of Causes and Accident Description 15.3-6 15.3.4.2 Analysis of Effects and Consequences 15.3-7 , 15.3.4.3 Results 15.3-8 i 15.3.4.4 Conclusions 15.3-8 l 15.3.6 Single Rod Cluster Control Assembly Withdrawal at Full Power 15.3-9 1 Identification of Causes and Accident Description j 15.3.6.1 15.3.6.2 15.3.6.3 Analysis of Effects and Consequences Results. 15.3-9 15.3-10 15.3-10 hj l l 15.3.6.4 Conclusions 15.3-11 l 15.3.7 References - 15.3-12 1 15.4 CONDITION IV - LIMITING FAULTS 15.4-1 ! 15.4.2 Major Secondary System Pipe Rupture 15.4-2 15.4.2.1 Rupture of a Main Steam Line 15.4-2 15.4.2.2 Major Rupture of a Main Feedwater Pipe 15.4-10 1
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15'.4.4 Single Reactor Coolant Pump Locked Rotor 15.4-16 15.4.4.1 Identification of Causes and Accident Description 15.4-16 l 15.4.4.2 Analysis of Effects and Consequences 15.4-16 l 15.4.4.3 Results 15.4-19 l 15.4.4.4 Conclusions 15.4-19 15.4.6 Rupture of a Control Rod Drive Mechanism Housing (Rod Cluster Control Assembly Ejection) 15.4-21 15.4.6.1 Identification of Causes and Accident Description 15.4-21 15.4.6.2 Analysis of Effects and Consequences 15.4-24 15.4.6.3 Results 15.4-29 15.4.6.4 Conclusions 15.4-31 15.4.7 References 15.4-32 1332v:1oteso2ss 15.0-iv
L. _. _ ,.. - l l
+
l TABLE OF CONTENTS (Cont) 4 TABLES Table Title l 15.0-1 Equipment Available for Transient and Accident Conditions , ci l 15.1-1 Nuclear Steam Supply System Power Rating _ l 15.1-2 Trip Points and Time Delays to Trip Assumed in Accident Analyses : 15.1-4 Summary of Initial Conditions and Computer Codes Used l 15.2-1 Time Sequence of Events for Condition II Events i 15.3-3 Time Sequence of Events for Condition III Events 15.4-8 Time Sequence of Events for Major Secondary System Pipe Ruptures 15.4-9 Summary of Results for Locked Rotor Transients 15.4-10 Parameters Used in the Analysis of the Rod Cluster Control Assembly Ejection Accident l-l l
- O l
I 1 9 O 1332v:1o/cso2ss 15.0-v 1:
TABLE OF CONTCNTS (Cont) FIGURES g, Figure Title 15.1-1 Overtemperature and Overpower Delta-T Protection 15.1-2 Rod Position versus Time on Reactor Trip
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15.1-3 Normalized RCCA Reactivity Worth versus Percent Insertion 15.1-4 Normalized RCCA Bank Worth versus Time After Trip 15.1-5 Doppler Power Coefficient Used in Accident Analyses 15.1-6 Residual Decay Heat 15.1-7 1979 ANS Decay Heat 15.1-8 Fuel Rod Cross Section 15.2.1-1 Uncontrolled Red Withdrawal from A Suberitical Condition - Nuclear Power and Core Heat Flux Versus Time 15.2.1-2 Uncontrolled Rod Withdrawal from A Suberitical Condition - Hot Spot Fuel Average and Clad Temperature Versus Time 15.2.2-1 Uncontrolled Rod Withdrawal From 100*.' Power Tenainated by High heutron Flux Trip - Pressurizer Pressure and Nuclear Power Versus Time 15.2.2-2 Uncontrolled Rod Withdrawal From 100% Power Terminated by High Neutron Flux Trip - DNBR and T avg Versus Time . 15.2.2-3 Uncontrolled Rod Withdrawal From 100% Power Terminated by Overtemperature Delta-T Trip - Pressurizer Pressure and Nuclear Power Versus Time 15.2.2-4 Uncontrolled Rod Withdrawal From 100% Power Terminated by Overtemperature Delta-T Trip - DNBR and T,yg Versus Time 15.2.2-5 E'ffect of Reactivity Insertion Rate on Minimum DNBR For a Rod Withdrawal Accident at 100% Power O 1332vao/oso2as 15.0-vi
TABLE OF CONTENTS (Cont) O FIGURES Figure Title 15.2.2-6 Effect of Reactivity Insertion Rate on Minimum DNBR For a Rod Withdrawal Accident at 60% Power
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15.2.2-7 Effect of Reactivity Insertion Rate on Minimum DNBR For a Rod Withdrawal Accident at 10% Power 15.2.3-1 Transient Response to A Dropped RCCA - Nuclear Power and Heat Flux Versus Time 15.2.3-2 Transient Response to A Dropped RCCA - T'V9 and Pre'ssurizer Pressure Versus Time 15.2.5-1 All Loops Operating, One Loop Coasting Down - Vessel Flow and Faulted Loop Flow Versus Time 15.2.5-2 All Loops Operating, One Loop Coasting Down - Nuclear Power and Heat Flux Versus Time 15.2.5-3 All loops Operating, One Loop Coasting Down - DNBR Versus Time 15.2.6-1 Startup of an Inactive Loop - Nuclear Power Versus Time 15.2.6-2 Startup of an Inactive Loop - Average and Hot Channel Heat Flux Versus Time 15.2.6-3 Startup of an Inactive Loop - Pressurizer Pressure and Core T,yg Versus Time 15.2.6-4 Startup of an Inactive Loop - Core Flow and DNBR Versus Time 15.2.7-1 Loss of Load With Pressurizer Spray and Power-operateo telief Valves at Beginning of Life - Nuclear Power and DNBR Ve sus Time 15.2.7-2 Loss of Load With Pressurizer Spray and Power-operated Relief Valves at Beginning of Life - Pressurizer Pressure and Water Volume Versus Time ! 15.2.7-3 Loss of Load With Pressurizer Spray and Power-operated Relief Valves at Beginning of Life - Core T ayg and Steam Temperature Versus Time 15.2.7-4 Loss of Load With Pressurizer Spray and Power-operated Relief Valves at End of Life - Nuclear Power and DNBR Versus Time 15.2.7-5 Loss of Load With Pressurizer Spray and Power-operated Relief Valves at End of Life - Pressurizer Pressure and Water Volume Versus Time O V 1332mio/oso2sa 15.0-vii l
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TABLE OF CONTENTS (Cont) FIGURES h. Figure Title 15.2.7-6 Loss of Load With Pressurizer Spray and. Power operated Relief Valves at End of Life - Core T,yg and Steam Temperature Versus Time 15.2.7-7 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at Beginning of Life - Nuclear Power and DNBR Versus Time 15.2.7-8 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at Beginning of Life - Pressurizer Pressure and Water Volume Versus Time 15.2.7-9 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at Beginning of Life - Core T avg and Steam Temperature Versus Time 15.2.7-10 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at End of Life - Nuclear Power and DNBR Versus Time 15.2.7-11 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at End of Life - Pressurizer Pressure and Water Volume Versus Time 15.2.7-12 Loss of Load Without Pressurizer Spray and Power-operated Relief Valves at End of Life - Core T,yg and Steam Temperature Versus h Time 15.2.8-1 Loss of Normal Feedwater - Loops 1 and 2 Temperatures, Loop 3 Temperatures, Steam Generator Water Hass Versus Time - 15.2.8-2 Loss of Normal Feedwater - Nuclear Power, Pressurizer Water Volume and Pressurizer Pressure Versus Time 15.2.10-1 Feedwater System Malfunction - Nuclear Power, Core Heat Flux and Pressurizer Pressure Versus Time 15.2.10-2 Feedwatar System Halfunction - Loop Delta-T, Core T avg and DNBR Versus Time 15.2.11-1 Excessive Load Increase Without Control, Minimum Feedback - Nuclear Power and Pressurizer Pressure Versus Time 9 1332<:1o/aso2ss 15.0 viii l
g TABLE OF CONTENTS (Cont) FIGURES Figure Title 15.2.11-2 Excessive Load Increase Without Control', Minimum Feedback - T,yg and DNBR Versus Time 15.2.11-3 Excessive Load Increase Without Control, Maximum Feedback - Nuclear Power and Pressurizer Pressure Versus Time 15.2.11-4 Excessive Load Increase Without Control, Maximum Feedback - T,yg and DNBR Versus Time 15.2.11-5 Excessive Load Increase With Control, Minimum Feedback - Nuclear Power and Pressurizer Pressure Versus Tiine 15.2.11-6 Excessive Load Increase With Control, Minimum Feedback - T'V9 and DNBR Versus Time 15.2.11-7 Excessive Load Increase With Control, Maximum Feedback - Nuclear Power and Pressurizer Pressure Versus Time 15.2.11-8 Excessive Load Increase With Control, Maximum Feedback - Tavg and DNBR Versus Time 15.2.12-1 Accidental Depressurization of the Reactor Coolant System - Nuclear Power and Core T,yg Versus Time 15.2.12-2 Accidental D'epressurization of the Reacter Coolant System - Pressurizer Pressure and Water Volume Versus Time 15.2.12-3 Accidental Depressurization of the Reactor Coolant System - DNBR , Versus Time 15.2.13-1 Main Steam Depressurization - Variation of K,ff with Core l Temperature j 15.2.13-2 Main Steam Depressurization - Safety Injection Flowrate ! 15.2.13-3 Transient Response For A Steam Line Break Equivalent to 255 lb/see at 1100 psia With Offsite Power Available 15.2.13-4 Transient Response For A Steam Line Break Equivalent to 255 lb/see at 1100 psia With Offsite Power Available l 15.2.14-1 Spurious Actuation of the Safety Injection System - Nuclear l Power, Steam Flow and Core T,yg Versus Time l 15.2.14-2 Spurious Actuation of the Safety Injection System - Pressurizer I Water Volume and Pressurizer Pressure Versus Time l (hG iss2v:1o/osozes 15.0-ix
HBLEOFCONTENTS(Cont) FIGURES Figure Title 15.3.3-1 Inadvertent Fuel Misloading - Interchange of L.1.ion 1 and Region 3 Assembly 15.3.3-2 Inadvertent Fuel Misloading - Interchange of Region 1 and Region 2 Assembly, Poison Rods Retained in Region 2 Assembly 15.3.3-3 Inadvertent Fuel Misloading - Interchange of Region 1 and Region 2 Assembly, Poison Rods Transferred to Region 1 Assembly 15.3.3-4 Inadvertent Fuel Misloading - Enrichment Error, A Region 2 Assembly, loaded into the Coro Center 15.3.3-5 Inadvertent Fuel Misloading - A 9egion 2 Assembly Loaded into A Region 1 Position Near Core Periprery , 15.3.4-1 All loops Operating, All Loops Coast;ng Down - Vessel Flow and Heat Flux Versus Time 15.3.4-2 All Loops Operating, All loops Coasting Dswn - Nuclear Power and DNBR Versus Time 15.4.2-1 Variation of Reactivity with Power at Constant Core Average Temperature 15.4.2-2 Transient Response to a Steam Line Break Double Enoad Rupture with Offsite Power Available (Case A) Transient Response to a Steam Line Break Double Ended Rupture 15.4.2-3 with Offsite Power Available (Case A) 15.4.2-4 Transient Response to a Steam Line Break Double Ended Rupture with No Offsite Power Available (Case B) 15.4.2-5 Transient Response to a Steam Line Break Double Ended Rupture with No Offsite Power Available (Case B) 15.4.2-6 Main Feedline Rupture with Offsite Power - Nuclear Power and Core Heat Flux Versus Time 15.4.2-7 Main Feedline Rupture with Offsite Power - Pressurizer Pressure and Water Volume Versus Time 15.4.2-8 Main Feedline Rupture with Offsite Power - Faulted and Intact Loop Coolant Temperatures Versus Time 1 1 0 l 1332v:10/050288 15.0-x l l 1
TABLE OF CONTENTS (Cont)
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FIGURES Figure Title 15.4.2-9 Main Feedline Rupture with Offsite Power - Steam Generator Pressure and Water Mass Versus Time. . ~ 15.4.2-10 Main Feedline Rupture without Offsite Power - Nuclear Power and Core Heat Flux Versus Time 15.4.2-11 Main Feedline Rupture without Offsite Power - Pressurizer Pressure and Water Volume Versus Time 15.4.2-12 Main Feedline Rupture without Offsite Power - Faulted and Intact Loop Temperatures Versus Time 15.4.2-13 Main Feedline Rupture without Offsite Power - Steam Generator Pressure and Water Mass Versus Time 15.4.4-1 All Loops Operating, One Locked Rotor - RCS Pressure, RCS Flow and Faulted Loop Flow Versus Time . 15.4.4-2 All Loops Operating, One Locked Rotor - Nuclear Power, Heat Flux ! and Clad Temperature Versus Time ,
'N 15.4.6-1 Rod Ejection Accident, BOL HFP - Nuclear Power, Hot Spot F.ue' and Clad Tamperature Versus Time ,
15.4.6-2 Rod Ejection Accident, BOL HZP - Nuclear Power, Hot Spot Fuel and Clad Temperature Versus Time i I 1 l ! I l 1 l l l 1 l 13:2wto/osores 15.0-xi l
^ Chapter 15 ACCIDENT ANALYSES Since 1970, the ANS classification of plant conditions has been used to divide plant conditions into four categories in accordance with anticipated frequency of occurrence and potential radiological consequences to the public. The four categories are as follows: (1) Condition 1: Normal Operation and Operational Transients (2) Condition II: Faults of Moderate Frequency (3) Condition III: Infrequent Faults Limiting Faults. (4) Condition IV: The basic principle applied in relating design requirements to each of the conditions is that the most frequent occurrences must yield little or no radiological risk to the public, and those extreme situations having the potential for the greatest risk to the public shall be those least likely to occur. Where applicable, reactor trip system and engineered safety features functioning is assumed, to the extent allowed by considerations such as the - single failure criterion, in fulfilling this principle. In the evaluation of the radiological consequences associated with initiation of a spectrum of accident conditions, numerous assumptions must be postulated. In many instances these assumptions are a product of extremely conservative judgments. This is due to the fact that many physical phenomena, in particular fission product transport under accident conditions, are not understood to the extent that accurate predictions can be made. Therefore, the set of assumptions postulated would predominantly determine the accident classification. i 1 t 0' 1332v:1o/050288 15.0-1 __________L
The specific accident sequences analyzed in this chapter include those required by Revision 1 of Regulatory Guide 1.70, Standard Format and Content . of Safety Analysis Reports for Nuclear Power Plants, and others considere'd significant for V. C. Sumer. Because the V. C. Sumer design differs from other plants, some of the accidents identified in Tab 1e 15-1 of Regulatory Guide 1.70, Revision 1, are not applicable to this~ plant; some coments on these items are as follows: (Item 10) - There are no pressure regulators or regulating instruments in the Westinghouse pressurized water reactor (PWR) design whose failure could cause heat removal greater than heat generation. (Item 11) - Reactor coolant flow controller is not a feature of the Westinghouse PWR design. Treatment of the performance of the reactivity controller in a number of accident conditions is offered in this chapter. (Item 12) - The analysis of specific effects of internal and external events such as major and minor fires, floods, storms, or earthquakes are generally discussed in Chapter 3. Refer to Section 3.1.2.1 for guidance on which FSAR sections specifically address GDCs 2, 3 and 4.
- l. (Item 22) - No instrument lines from the RCS boundary in the V. C. Sumer .
design penetrate the containment (a) , l (a) For definition of the RCS boundary, refer to the 1972 issue of ANS N18.2, Nuclear Safety Criteria for the Design of Stationary PWR Plants. l l l tas2cio/osissa 15.0-2 L _
(Item 26) - Habitability of the control room following accident conditions is discussed. In addition, Chapter 7 contains an analysis showing that the plant , can be brought to, and maintained in, the hot shutdown condition from outside the control room. (Item 27) - Overpressurization of the residual heat removal system (RHRS) is considered extremely unlikely due to the isolation valve interlocks described in Section 7.6. (Item 28) - This event is covered by the analyses of Section 15.2.7, Loss of External Electrical Load and/or Turbine Trip. (Item 29) - Same as item 28 above. (Item 30) - Loss of the service water system is discussed in Section 9.2. (Item 31) - Loss of one DC system is discussed in Chapter 8. (Item 33) - The effects of turbine trip on the RCS are presented in Section 15.2.7. Turbine trip with failure of the generator breakers to open g is discussed in Chapter 10.
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(Item 34) - Halfunctions of this system are discussed in Chapter 9. (Item 35) - The radiological effects of this event are not significant for PWR plants. O 1332rio/ose2 s 15.0-3
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[ N v Sheet 1 cf 2 V) i TABLE 15.0-1 FOUIPMENT AVAILABLE FOR TRANSIENT AND ACCIDENT CONDITIONS Incident Reactor Trip Functions ESF Actuation Functions Other Equipment ESF Equipment
- t. Uncontrolled RCAA Power range high Bank withdrawal from flux (Iow s.p.), manual - -
- 2. Uncontrolled RCCA Power range high flux - Pressuetror safety valves. -
Bank withdrawal at OTdT. ht pressurtzer steam generator safety Power pressure, manual valves
- 3. RCCA Power range negative - -
Mtsalignment flux rate, manual 4 Uncontrolled Source range high flux. Low insertion limit - Boron Otlution power range high flux, annunctators for OTar manuai boration
- 5. Startup of an Power range high flux. - -
Inactive Reactor manual , Coolant Loop
- 6. Less of External High pressurtzer pressure -
Pressurtzer safety valves. -
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Electrical Load and/ OTar, manuai . team generator safety or Turbine Trip valves
- 7. Loss of Normal Steen generator lo-lo Steam generator 1o-10 -
One motor driven Feedwater level, manual level emergency feedwater pump
- 8. Loss of Offette Same as 7 Some as 7 Same as 7 Same as 7 :
Power to the Station , Aux 14tartes
- 9. Excess Heat Re- Power range high flux. High steam generator Feedwater teolation -
moval due to Feed- high steam generator level produced feed- valves water System Mal- level, manual water isolation and functions turbine trip i
- 10. Excessive Load Power range high flux. -
Pressurizer self-actuated - i Increase Incident OTAT. oPAT. manuai safety vaives, steam i generator safety valves it. Accidental Depres- Pressurt er low - surtratton of the RCS pressure, OTAT. manual
- 12. Major Rupture of Main SIS. manual - Low pressurizer Feed itne isolation valves. Emergency feed-Steam Line pressure. Iow cosp- steam Itne isolation water system. SI ensated steam Itne valves equipment minus pressure, ht-t con- either one Si tainment pressure, charging pump, or manual one diesel generator.
1332v:1D/OS0388 t . - - ~ ~ .- _ . a, .. _,,
Sheet 2 ttf 2 TABLE 15.0-1 (Cont'd) EQUIPMENT AVAILABLE FOR TRANSIENT AND ACCIDENT CONDITIONS Incident Reactor Trip Functions ESF Actuation Ft.nctions Other Equipment ESF Equipment
- 13. Complete Loss of Low flow, undervoltage -
Forced Reactor underfrequency, manual Coolant Flow
- 14. Rupture of a Control Power range high flux.
Rod Drive Mechanisa manual Housing
- 15. Single RCCA With- OTAT. manuai -
drawal at Full Power
- 16. Major Rupture of a to steam generator level High containment Steam 11ne isolatton Emergency feed-Main Feedwater Line plus steam / feed mismatch, pressure, high valves, feed Itne isolation water pumps SIS manual pressurtzer pressure, pressurtzer self-actuated steam generator low- safety valves, steam gen-low water level. Iow erator safety valves compensated steam line pressure
- 17. Large Break LOCA Reactor trip system Engineered safety Service water system. Faergency core features actuation component cooling cooling system.
system water system containment heat removal system. . emergency power system
- 18. Small Break LOCA Reactor trip system Engineered safety Service water system, Ersrgency core features actuation component cooling cooling system, system water system, generator emergency feedwater safety and/or reitef valves system containment heat removal system emergency power system
- 19. Steam Generator Reactor trip system Engineered safety Service water system. Emergency core Tube Rupture features actuation component cooling water cooling system.
system system, steam generator emergen;y feed-shell side fluid operattnD water system, system, steam generator emergency power
. safety and/or reitef valve, systems steam Itne isolation valves 1332v:lD/050388 9 ,
9 O
t 15.1 CONDITION I - NORMAL.0PERATION AND OPERATIONAL TRANSIENTS O G Condition I occurrences are those that are expected frequently or regularly in the course of power operation, refueling, maintenance, or maneuvering of the ; plant. As such, Condition 1 occurrences are accommodated with margin between l any plant parameter and the value of that parameter which would require either l automatic or manual protective action. Inasmuch as Condition I occurrences j occur frequently or ' regularly, they trust be considered from the point of view ! of affecting the consequer.ces of fault conditions (Conditions II, III and IV). In this regard, analysis of each fault condition is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions that can occur during Condition I operation. A typical list of Condition I events is shown below: - (1) Steady state and shutdown operations i Mode 1 - Power operation (> 5% of rated thermal power)
,FT 4 Mode 2 - Startup (K,ff 3 0.99, 5 5% of rated thermal power)
Mode 3 - Hot standby (X,ff < 0.99, T ,yg 3 350'F) _ Mode 4 - Hot shutdown (suberitical, residual heat removal system in operation, K,ff < 0.99, 200'F < T,yg < 350'F) Mode 5 - Cold shutdown (suberitical, residual heat removal system in operation, K,ff < 0.99, T avg 5 203*F) Mode 6 - Refueling (k,ff 5 0.95, T,yg 5 140*F) b'
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1332v:1oicsere 15.1-1
(2) Operation with permissible deviations Various deviations that may occur during continued operation as permitted by the plant Technical Specifications (1) must be considered in conjunction with other operational modes. These include: (a) Operation with components or systems out of service (b) Leakage from fuel with cladding defects-(c) Activity in the reactor coolant
- 1. Fission products
- 2. Corrosion products
- 3. Tritium (d) Operation with steam generator leaks up to the maximum allowed by the Technical Specifications (e) Testing .s allowed by the Technical Specifications -
(3) Operational transients (a) elant heatup and cooldown (up to 100*F/ hour for the reactor ccolant system (RCS); 200*F/ hour for the pressurizer) (b) Step load changes (up to +10%) (c) Ramp load changes (up to 5% per minute) (d) Load rejection up to and including design load rejection transient O t u 2v:1o/osozes 15.1-2
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i m 15.1.1 Optimization of Control Systems kh A setpoint study W has been performed in order to simulate performance of the reactor control and protection systems. Emphasis was placed on the development of a control system that will automatically maintain prescribed conditions in the plant even under the most conservat'ive set of reactivity , parameters with respect to both system stability and trans.ient performance. i i For each mode of plant operation, a group of optimum controller setpoints is determined. In areas where the resultant setpoints are different, compromises based on the optimum overall performance are made and verified. A consistent set of control system parameters is derived satisfying plant operational requirements throughout the core life and for power levels between 15 and 100%. The study comprises an analysis of the following control systems: rod cluster assembly control, steam dump, steam generator level, pressurizer pressure, and pressurizer level. 15.1.2 Initial Power Conditions Assumed in Accident Analyses Reactor power-related initial conditions assumed in the accident analyses b LJ presented in this chapter are described in this section. 15.1.2.1 Power Rating Table 15.1-1 lists the principal power rating values that are assumed in . analyses performed in this section. Two ratings are given: (1) The guaranteed nuclear steam supply system (NSSS) thermal power output. This power output includes the thermal power generated by the reactor coolant pumps. l (2) The engineered safety features (ESF) design rating. The Westinghouse-supplied ESFs are designed for a thermal power higher than the guaranteed value in order not to preclude realization of future potential power capabilty. This higher thermal power value I is designated as the ESF design rating. This power output includes the thermal power generated by the reactor coolant pumps. O v l 1332v:to/oso:ss 15.1-3
Where initial power operating conditions are assumed in accident analyses, the guaranteed NSSS thermal power output (plus allowance for errors in steady ' state power determination for some accidents) is assumed. Where demonstration of the adequacy of the containment and ESF is concerned, the ESF design rating plus allowance for error is assumed. The thermal power values for each transient analyzed are given in Table 15.1-4. 15.1.2.2 Initial Conditions For most accidents which are DNB limited, nominal values of initial conditions are assumed. The allowances on power, temperature, and pressure are determined on a statistical basis and are included in the limit DNBR, as described in Reference 3. This procedure is known as the "Improved Thermal Design Procedure" (ITDP) and these accicents utilize the WRB-1 and WRB-2 DNB correlations (References 4 and 5). ITDP allowances may be more restrictive than non-ITDP allowances. The initial conditions for other key parameters are selected in such a manner to maximize the impact on DNBR. Minimum measured flow is used in all ITDP transients. For accident evaluations that are not CNS-limited, or for which the Improved Thermal Design Procedura is not employed, the initial conditions are obtained by adding maximum steady state errors to rated values. The following steady state errors are considered: (1) Core power +2.0%/-2.1% allowance calorimetric error (2) Average RCS temperature +4.0*F/-4.3*F allowance for deadband and measurement error (3) Pressurizer pressure +33 psi /-45 psi allowance for steady state fluctuations and measurement error. I O m 2 n o/ese2ss 15.1-4
_ m._ ._ . __ 15.1.2.3 Power Distribution The transient response of the reactor system is dependent on the initial power - distribution. The nuclear design of the reactor core minimizes adverse power distribution through the placement of fuel assemblies, control rods, and by operation instructions. The power distribution may be characterized by the radial peaking factor F3g and the total peaking factor F .g _ The peaking factor limits are given in Technical Specification 3/4.2. For transients that may be DNB-limited, the radial peaking factor is of i importance. The radial peaking factor increases with decreasir.g power level due to rod insertion. This increase in F 3g is included in the core limits illustrated on Figure 15.1-1. All transients that may be DNB-limited are assumed to begin with an FAH consistent with the initial power level defined in the Technical Specifications. The axial power shape used in the DNB calculation is discussed in Section 4.4.3. For transients that may be overpower-limited, the total peaking factor Fg is of importance. The value of Fgmay increase with decreasing power level so that the full power hot spot heat flux is not exceeded, i.e., F0 x Power = l t design hot spot heat' flux. All transients that may be overpower-limited are
~;
i assumed to begin with a value of Fg consistent with the initial power level as defined in the Technical Specifications. The value of peak kW/ft can be directly related to fuel temperature as illustrated on Figures 4.4-1 and 4.4-2. For transients that are slow with t respect to the fuel rod thermal time constant (approximately 5 seconds), the fuel temperatures are illustrated on Figures 4.4-1 and 4.4-2. For transients ! that are fast with respect to the fuel rod thermal time constant (for example, rod ejection), a detailed heat transfer calculation is made. 15.1.3 Trip Points and Tinie Delays to Trip Assumed in Accident Analyses f A reactor trip signal acts to open two trip breakers connected in series feeding power to the control rod drive mechanisms. The loss of power to the O 1 22w1o/eso2ss 15.1-5 i , I
mechanism coils causes the mechanism to release the rod cluster control assemblies (RCCAs) which then fall by gravity into the core. There are various instrumentation delays associated with each trip function, including g, delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to tr,ip is defined as the time delay from the time that trip conditions are reached to the time the rods are free and begin to fall. Limiting trip setpoints assumed in accident analyses and the time delay _ assumed for each trip function are given in Table 15.1-2. Reference is made in that table to the overtemperature and overpower AT trip shown on Figure 15.1-1. This figure presents the allowable reactor coolant loop average temperature and aT for the design flow and the NSSS Design Thermal Power distribution as a function of primary coolant pressure. The boundaries of operation defined by the Overpower AT trip and the Overtemperature AT trip are represented as "protection lines" on this diagram. The protection lines are drawn to include all adverse instrumentation and setpoint errors so that under nominal conditions trip would occur well within the area bounded by these lines. The utility of this diagram is in the fact that the limit imposed by any given DNBR can be represented as a line. The DNB lines represent the locus of conditions for which the DNBR equals the safety analysis limit values (1.44 and 1.48 for Standard thimble. cell and typical cells, respectively; 1.60 and 1.68 for V-5 thimble cell and typical cells, respectively) for ITDP accidents. All points below and to the lef t of a DNB line for a given pressure have a DNBR greater - than the limit values. The diagram shows that DNB is prevented for all cases if the area enclosed with the maximum protection lines is not traversed by the applicable DNBR line at any point. The area of permissible operation (power, pressure and temperature) is bounded by the combination of reactor trips: high neutron flux (fixed setpoint); high pressurizer pressure (fixed setpoint); low pressurizer pressure (fixed setpoint); Overpower and Overtemperature AT (variable setpoints). O 1 1332tio/csc2ss 15.1-6 i l l l
The limit values, which were used as the DNBR limits for all accidents Q analyzed with the Improved Thermal Design Procedure are conservative compared to the actual design DNBR values required to meet the DNB design basis. The difference between the limiting trip point assumed for the analysis and the normal trip point represents an allowance for' instrumentation channel
. error and setpoint error. During startup tests, it is demonstrated that ,
actual instrument errors and time delays are equal to or less than the assumed values. 15.1.4 Rod Cluster Control Assembly Insertien Characteristic The negative reactivity insertion following a reactor trip is a function of the acceleration of the RCCA and the variation in rod worth as a function of rod position. With respect to accident analyses, the critical parameter is the time of insertion up to the dashpot entry or approximately 85% of the rod cluster travel. For accident analyses, the insertion time to dashpot entry is conservatively taken as 2.7 seconds. The RCCA position versus time assumed in accident analyses is shown on Figure 15.1-2. L Figure 15.1-3 shows the fraction of total negative reactivity insertion for a :
~'
core where the axial distribution is skewed to the lower region of the core. This curve is used as input to all point kinetics core models used in transient analyses. There is inherent conservatism in the use of this curve in that it is based on a skewed axial power distribution that would exist relatively infrequently. . j For cases other than those associated with xenon oscillations, significant i negative reactivity would have been inserted due to the more favorable axial power distribution existing prior to trip. l t 1s32v:to/ose2sa 15.1-7 I
i The normalized RCCA negative reactivity insertion versus time is shown on Figure 15.1-4. The curve shown in this figure was obtained from Figures 15.1-2 and 15.1-3. A total negative reactivity insertion following a trip of h. 4.8% ak is assumed in the transient analyses except where specifically noted otherwise. This assumption is conservative with respect to the calculated trip reactivity worth available as shown in Tables 4.3-2 and 4.3-3. The normalized RCCA negative reactivity insertion versus time curve for an axial power distribution skewed to the h ttom (Figure 15.1-4) is used in transient analyses. Where special analyses require the use of three-ditensional or axial one-dimensional core models, the negative reactivity insertion resulting from reactor trip is calculated directly by the reactor kinetic code and is not separable from other reactivity feedback effects. in this case, the RCCA position versus time of Figure 15.1-2 is used as a code input. 15.1.5 Reactivity Coefficients The transient response of the reacter coolant system is dependent.on reactivity feedback effects, in particular the moderator temperature g coefficient and the Doppler power coefficient. These reactivity coefficients and their values are discussed in detail in Chapter 4. In the analysis of certain events, conservatism requires the use of large reactivity coefficient values, whereas in the analysis of other events, conservatism requires the use of small reactivity coefficient values. Some analyses, such as loss of reactor coolant from cracks or ruptures in the RCS, do not depend on reactivity feedback effects. The values used are given in Table 15.1-4; reference is made in that table to Figure 15.1-5 that shows the upper and lower Doppler power coefficients, as a function of power, used in the transient analysis. The justification for use of conservatively large versus small reactivity coefficient values is treated on an event-by-event basis. O 1222v:1o/eso2ss 15.1-8
15.1.6 Fission Product Inventories The fission product inventories existing in the core and fuel rod gaps are described in Section 15.1.7 of the FSAR. The description of the modals used j for calculating fuel gap activities is included in Section 15.1.7.2 of the l FSAR. l 15.1.7 Residual Decav Heat Residual heat in a subtritical core consists of: (1) Fission product decay energy (2) Decay of neutron capture products (3) Residual fissions due to the effect of delayed neutrons. These constituents are discussed separately in the following paragraphs.
/ 15.1.7.1 Fission Product Decay l For short times (<103 seconds) after shutdown, data on yields of short-half-life isotopes is sparse. Very little experimental data is available for the gamma ray contributions and even less for the beta ray contribution. _
Several authors have compiled the available data into a conservative estimate , of fission preduct decay energy for short times after shutdown, notably Shure , Dudziak , and Teage(8) . Of these tnree selections, Shure's curve is the highest and is based on the data of Stehn and Clancy(9) and Obenshain and Foderaro(10) . The fission product contribution to decay heat that has been assumed in the LOCA accident analyses is the curve of Shure increased by 20% for conservatism. This curve with the 20% factor included is shown on Figure 15.1-6. For the non-LOCA analyses the 1979 ANS decay heat curve is usedIII'. Figure 15.1-7 presents this curve as a function of time after shutdown. l (3 I 1332v:10/880517 15.1-9
,15.1.7.2 Decay of U-238 Capture Products Betas and gammas from the decay of U-239 (23.5-minute half-life) and Np-239 g' (2.35-day half-life) contribute significantly to the heat generation after shutdown. The cross sections for production.of these isotopes and their decay schemes are relatively well known. For long irradiation times their contribution can be written as:
P1 /P, =(E 1+E 31 )c(1+a) e -lit watts / watt (15.1-1) 200 MeV 1 (,-12t _ , -l t)i , ,-12 3 twatts / watt E P /P = (Eg 2+ 32 )c(1+a) [ 2 2 o 1 -1 1 2 200 MeV (15.1-2) - where: P1 /P, is the energy from U-239 decay P2 /P, is the energy from Np-239 decay h t is the time after shutdown (seconds) c(1+a) is the ratio of U-238 captures to total fissions = 0.6(1 + 0.2)
-4 -1 1 = the decay constant of U-239 = 4.91 x 10 seconds 1
-6 -1 1 = the decay constant of Np-239 = 3.41 x 10 seconds 2
E = total r ray energy from U-239 decay = 0.06 HeV 1 E = total r ray energy from Np-239 decay = 0.30 MeV 2 O isartic/csores 15.1-10 t
0 n E = total 8 ray energy from U-239 decay = 1/3(a) x 1.18 MeV l 3 0 , .l E = total B ray energy from Np-239 decay = 1/3(a) x ,43 gey 2 B This expression with a margin of 10% is shown on Figure 15.1-6 as it is used in the LOCA analysis. The 10% margin, compared te 20% for fission product decay, is justified by the availability of the basic data required for this analysis. The decay of other isotopes, produced by neutron reactions other than fission, is neglected. For the non-LOCA analysis, the decay of U-238 l capture products is included as an integral part of the 1979 decay heat curve ! presented as Figure 15.1-7. l. 15.1.7.3 Residual Fissions The time dependence of residual fission power after shutdown depends on core ! properties throughout a transient under consideration. Core average conditions are acre conservative for the calculation of reactivity and power level than actual local conditions as they would exist in hot areas of the core. Thus, unless otherwise stated in the text, static power shapes have r p been assumed in the analysis and these are factored by the time behavior of d core average fission power calculated by a point kinetics model calculation with six delayed neutron groups.
~
For the purpose of illustration, only one delayed neutron group calculation, with a constant shutdown reactivity of -4% ak is shown on Figure 15.1-6. (a) Two-thirds of the potential 6-energy is assumed to escape by the accompanying neutrinos. O i 1 szw1o/csezas 15.1-11
}
15.1.7.4 Distribution of Decay Heat Following loss of Coolant Accident During a loss-of-coolant accident (LOCA), the core is rapidly shut down by . void formation or RCCA insertion, or both, and long-term shutdown is assured by the borated ECCS water. A large fraction of the heat generation to be considered comes from fission product decay gamma rays. This heat is it distributed in the same manner ss steady state fission power. Local peaking effects that are important for the neutron dependent part 6f the heat generation do not apply to the gama ray source contribution. The steady state factor of 97.4% that represents the fraction of heat generated within the cladding and pellet drops to 95% for the hot rod in a LOCA. For example, consider the transient re 'ino from the postulated double-ended break of the largest RCS pipe; 1/2 r id arter the rupture, about 30% of the heat generated in the fuel rods is from gamma ray absorption. The gamma power shape is less peaked than the steady state fission power shape, reducing the energy deposited in the hot rod at the expense of adjacent colder rods. A conservative estimate of this effect is a reduction of 10% of the gama ray contribution or 3% of the total. Since the water density is considerably reduced at this time, an average of 98% of the available heat is deposited in the fuel rods, the remaining 2% being absorbed by water, thimbles, sleeves, and grids. The net effect is a factor of 0.95, rather than 0.974, to be. applied to the heat production in the hot rod. 15.1.8 Comouter Codes Utilized Summaries of some of the principal computer codes used in transient analyses are given below. Other codes, in particular, very specialized codes in which the modeling has been developed to simulate one given accident, such as the SATAN-VI code used in the analysis of the RCS pipe rupture (Section 15.4), and which consequently have a direct bearing on the analysis of the accident itself, are summarized in their respective accident analyses sections. The codes used in the analyses of each transient are listed in Table 15.1-4. O l 1332v:io/oscass 15.1-12 l l l
1.1, i_ L . _. . - . I l i 15.1.8.1 FACTRAN FACTRAN calculates the transient temperature distribution in a cross section -l of a metal clad 002 fuel rod (see Figure 15.1-8) and the transient heat flux l at the surface of the clad using as input the nuclear power and the time-dependent coolant parameters (pressure, flow, te'mperature and density). The code uses a fuel model that exhibits the following features simultaneously: i I 3 (1) A sufficiently large number of finite difference radial space { increments to handle fast transients such as rod ejection accidents l { (2) Material properties that are functions of temperature and a sophisticated fuel-to-clad gap heat transfer calculation (3) The necessary calculations to handle post-DNB transients: film boiling heat transfer correlations, zircaloy-water reaction and partial melting of the materials. The gap heat transfer coefficient is calculated according to an elastic pellet model. The thermal expansion of the pellet is calculated as the sum of the radial (one-dimensional) expansions of the rings. Each rir.g is assumed to expand freely. The clad diameter is calculated based on thermal expansion and internal and external pressures. _ If the cutside radius of the expanded pellet is smaller than the inside radius of the expanded clad, there is no fuel-clad contact and the gap conductance is calculated on the basis of the thermal conductivity of the gas contained in the gap. If the pellet outside radius so calculated is larger than the clad inside radius (negative gap), the pellet and the clad are pictured as exerting upon each other a pressure sufficient to reduce the gap to zero by elastic deformation of both. This contact pressure determines the heat transfer coefficient. FACTRAN is further discussed in Reference 12. s O
,,s _ m ss 15.1-u ;
I
l l 15.1.8.2 LOFTRAN The LOFTRAN program is used for studies of transisat rasponse of a PWR system to specified perturbations in process parameters. LOFTRAN simulates a multiloop system by modeling the reactor core and vessel, hot and cold leg piping, steam generator (tube and shell-sides), reactor coolant pumps and the pressurizer with up to four reactor coolant loops.- The pressurizer heaters, spray, relief and safety valves are also considered in the' program. Point model neutron kinetics and reactivity effects of the moderator, fuel, boren, and rods are included. The secondary side of the steam generator utilizes a homogeneous, saturated mixture for the thermal transients and a water level correlation for indication and control. The reactor protection system is simulated to include reactor trips on neutron flux, overpower and overtemperature reactor coolant AT, high and low pressure, low flow, and high pressurizer level. Control systems are also simulated including rod - control, steam dump, feedwater control, and pressurizer pressure control. The safety injection system (SIS), including the accumulators, is also modeled. LOFTRAN is a versatile program that is suited to both accident evaluation and control studies as well as parameter sizing. LOFTRAN also has the capability of calculating the transient value of DNB based on the input from the core limits illustrated on Figure 15.1-1. The core limits represent the minimum value of DNBR as calculated for a typical or thimble cell. LOFTRAN is further discussed in Reference 13. 15.1.8.3 LEOPARD The LEOPARD computer program determines fast and thermal spectra using only basic geometry and temperature data. The code optionally computes fuei depletion effects for a dimensionless reactor and recomputes the spectra before each discrete burnup step. LEOPARD is further discussed in Reference 14. O 1332rioecso2:s 15.1-14
15.1.8.4 TURTLE TURTLE is a two group, two-dimensional neutron diffusion code featuring direct treatment of the nonlinear effects of xenon, enthalpy, and Doppler feedback. Fuel depletion is allowed. TURTLE was written for the study of azimuthal xenon oscillations, but the code is useful for general analysis. The input is simple, fuel management is handled directly, and a bcron criticality search is allowed. TURTLE is further described in Reference 15. 15.1.8.5 TWINKLE The TWINKLE program is a multidimensional spatial neutron kinetics code which was patterned after steady state codes presently used for reactor core design. The code uses an implicit finite-difference method to solve the two group transient neutron diffusion equations in one, two, and three dimensions. The code uses six delayed neutron groups and contains a detailed multirecion fuel-clad-coolant heat transfer model for calculating pointwise Doppler and moderator feedback effects. The code handles up to 2000 spatial points and performs its own steady state initialization. Aside from basic cross section data and thermal-hydraulic parameters, the code accepts as input basic driving functions such as inlet temperature, pressure, flow, boren _ concentration, control rod motion, and others. Various edits provide channelwise power, axial offset, enthalpy, volumetric surge, pointwise power, fuel temperatures, and so on. The TWINKLE code is used to predict the kinetic behavior of a reactor for transients that cause a major perturbation in the spatial neutron flux distribution. TWINKLE is further described in Reference 16. 15.1.8.6 THINC The THINC code is described in Section 4.4.3 of the FSAR. O 1 32v:ivoso2ss 15.1-15
O 15.
1.9 REFERENCES
- 1. Technical Specifications, V. C. Summer Nuclear Station, Appendix A to License No. NPF-12, as amended through Amendment Number 66.
- 2. D. A. Reed and J. L. Little, Setooint Study SCE&G V. C. Summer Nuclear Plant, WCAP-9399, December 1978. -
- 3. Chelmer, H., et al., "lecroved Thermal Design Procedure," WCAP-8567 (Proprietary) and WCAP-8568 (Non-Droprietary), July 1975.
{
- 4. Hotleg, F. E., et al., "New Westinghouse Correlations WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids," I WCAP-8762-P-A and WCAP-8763-A, July 1984.
l l
- 5. Davidson, S. L. and Kramer, W. R.; (ed.) "Reference Core Report Vantage 5 l Fuel Assembly," Appendix A.2.0, September 1985. l
- 6. K. Shure, Fission Product Decay Energy in Bettis Technical Review, WAPD-BT-24, December 1961, pp. 1-17. h
- 7. K. Shure and D. J. Dudziak, "Calculating Energy Released by Fission Products," Trans. Am. Nucl. Soc. 4 (1) 30, 1961.
- 8. U.K.A.E.A. Decay Heat Standard.
I 9. J. R. Stehn and E. F. Clancy, "Fission-Product Radiocctivity and Heat I Generation," Proceeding of the Second United Nations International Conference on the Peaceful Uses of Atomic Energy, Geneva,1958, Volume 13, United Nations, Geneva, 1958, pp. 49-54. l
- 10. F. E. Obenshain and A. H. Federaro, Energy from Fission Product Decay, WAPD-P-652, 1955.
O 1332<io/esozas 15.1-16
- 11. ANSI /ANS-5.1-1979, "Decay Heat Power In Light Water Reactors",
August 23, 1979. .l
- 12. H. G. Hargrove, FACTRAN - A Fortran IV Code for Thermal Transients in a UO Fuel Rod, WCAP-7908, June 1972. -
2
~
- 13. T. W._T. Burnett et al, LOFTRAN Code Description, WCAP-7907-P-A (Proprietary), WCAP-7907-A (Non-Proprietary), April 1984.
l
- 14. R. F. Barry, LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code
- for the IBM-7904, WCAP-3269-26, September 1963.
- 15. R. F. Barry and S. Altomare, The TURTLE 24.0 Diffusion Deoletion Code, WCAP-7213-P-A (Proprietary), WCAP-7758-A (Non-Proprietary), January 1975.
2
- 16. D. H. Risher, Jr. and R. F. Barry, TWINKLE - A Multi-Dimensional Neutron Kinetics Computer Code, WCAP-7979-P-A (Proprietary), WCAP-8028-A 4
(Non-Proprietary), January 1975. ji
- O
. A l
I I i lO insav:1o/esoiss 15.1-17
TABLE 15.1-1 NUCLEAR STEAM SUPPLY SYSTEM POWER RATING
- h. l' Core thermal power (license level) 2775 Thermal power generated by the reactor coolant pumps 12 Nuclear steam supply system thermal power output 2787
. Engineered safety features design ratinggjximumcalculatedturbine rating) 2912 O (a) The unit will not be operated at this rating because it exceeds the license rating. 9 1332v:1o/Os0288
I l TABLE 15.1-2 Sheet 1 of 2 0 TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSES l 4 Limiting Trip Trip Point Assumed ' Time Delay, Function In Analyses see Power range high neutron flux, , high setting 0.5 118% Power range high neutron flux, low setting 35% 0.5 Overtemperature AT Variable, see 8.5(a) Figure 15.1-1 Overpower AT Variable, see 8.5(a) Figure 15.1-1 High pressurizer pressure 2440 psig 2 Low pressurizer pressure 1760 psig 2 l Low reactor coolant flow (from loop flow detectors) 87% loop flow 1 Undervoltage trip (b) 1.5 4 O i 1332v:10/050288
TABLE 15.1-2 Sheet 2 of 2 O Limiting Trip Trip Point Assumed Time Delay, Function In Analyses see Turbine trip Not applicable - 2 Low-low steam generator level 0% of narrow 2 range level span High-high steam generator level trip 96% of narrow 2 of the feedwater pumps and turbine; range level span 13* (for closure of feedwater system feedwater valves
- isolation)
(a) Total time delay (including RTD and thermowell time response, trip circuit and channel electronics delay) from the time the temperature difference in the coolant loops exceeds the trip setpoint until the rods are free to fall. (b) A specific undervoltage setpoint was not assumed in the safety analysis. O 1332v;1o/050288
O O O : Sheet 1 of 4 l-TABLE 15.1-4 ! SUle4ARY OF INITIAL CONDITIONS AND COMPUTER CODES USED i I" "I " Assumed Reactivity Coefficients Thermal Power Moderator Moderator Computer Temp.Ih), DensityI *I, Assumed (c), Faults Codes Utilized pcm/*F Ak/gm/cc Doppler (b) MWt CONDITION II Uncontrolled RCCA bank TWINKLE, +7 - Consistent 0 withdrawal from a subcritical FACTRAN, THINC with lower condition limit on Fig. 15.1-5 i Uncontrolled RCCA bank LOFTRAN +7 0.50 Lower and 2790 withdrawal at power Upper [ RCCA misoperation THINC, TURTLE - - - 2787 LOFTRAN Uncontrolled boron dilution 0 and 2787 Partial loss of forced reactor LOFTRAN 0 - Upper 2787 coolant flow FACTRAN, THINC Startup of an inactive reactor LOFTRAN, - 0.50 Lower 1672 coolant loop FACTRAN, TilINC . Loss of external electrical load LOFTRAN +7 0.50 Lower and 2787 and/or turbine trip Upper loss of normal feedwater LOFTRAN +7 - Upper 2790 Loss of offsite power to the LOFTRAN +7 - Upper 2790 plant auxiliaries (plant blackout) 1332v:1D/042788
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Sheet 2 cf 4 TABLE 15.1-4 i Assumed Reactivity Coefficients n ai M Moderator Moderator t Computer Temp. , DensityI '}, Assumed IC) , Faults Codes Utilized pcm/*F Ak/gm/cc Doppler (b) NWt CONDITION II (Cont'd) Excessive heat removal due to LOFTRAN - 0.50 Lower 0 and 2787 feedwater system malfunctions Excessive load increase LOFTRAN - 0 and 0.50 Lower and 2787 Upper Accidental depressurization of LOFTRAN +7 - Lower 2787 the reactor coolant system Accidental depressurization of LOFTRAN - Function of See Figure 0 the main steam system the modera- 15.4.2-1 (Subcritical) tor density. See Sec. 15.2.13 (Figure 15.2.13-1) Spurious operation of the SIS LOFTRAN +7 0.50 Lower and 2787 at power Upper , CONDITION III Loss of reactor coolant from small NOTRUMP - - - 2775I *) ruptured pipes or from cracks in SBLOCTA large pipe which actuate emergency core cooling 10/03308s
I O O O I i Sheet 3 of 4 } TAGLE 15.1 ' 4 i Assumed Reactivity Coefficients Initial NSSS Thermal Power Moderator Moderator Computer Temp.I* DensityI *I, c) , I I, faults Codes Utilized pcm/*F ak/gm/cc Doppler (b) MWt . CONDITION III (Cont'd) Inadvertent loading of a fuel LEOPARD, - - - 2775I *) assembly into an improper position TURTLE Complete loss of force reactor LOFTRAN, 0 - Upper 2787 coolant flow FACTRAN, THINC Single RCCA withdrawal at TURTLE, THINC, - - - 2787 full power LEOPARD CONOITION IV Major rupture of pipes containing SATAN-VI Function of - Function 2775I *) reactor coolant up to and including C0C0 moderator of fuel double ended rupture of the largest BASH density. temp. See pipe in the reactor coolant system WREFLOOD See Sec. Sec. 15.4.1 (loss-of-coolant accident) LOCBART 15.4.1 , Major secondary system pipe rupture LOFTRAN - Function of See Figure 0 up to and including double-ended the Modera- 15.4.2-1 , (Subcritical) rupture (rupture of a steam pipe) tor Density see Section 15.2.13 (Figure i 15.2.13-1)' 1332v:10/033088
Sheet 4 cf 4 TABLE 15.1-4 Assumed Reactivity Coefficients Initial NSSS Thermal Power Moderator Moderator DensityI *I, Output (c) Computer Temp.(*IfI, Assumed , Faults Codes Utilized pcm/*F Ak/gm/cc Doppler (b) NWt CONDITION IV (Cont'd) Major secondary system pipe rupture LOFTRAN - 0.50 Upper 2912 up to and including double ended rupture (rupture of a feedline) Single reactor coolant pump locked LOFTRAN 0 - Upper 2787 rotor FACTRAN, THINC Rupt TWINKLE, +7.1 BOL - Consistent 0 and 2775 g of a control rod mechanism housing (RCCA ejection) FACTRAN, -23. EOL with lower LEOPARD limit on Fig. 15.1-5 1 (a) Only one is used in analysis, i.e., either moderator temperature or moderator density coefficient. (b) Reference Figure 15.1-5. (c) Appropriate calorimetric error considered where applicable. (d) Pcm means percent mille. See footnote Table 4.3-1. (e) Core power. e
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L1 - 0 i a i e i e i U U U U U TDE F701 Rein 5E (EID2CS) V. C. Sumer Figure 15.1-2 Rod Position vs. Time On Reactor Trip
13 il - SJ - 1.7 - 0.5 - 6.5 - O Il.4-IJ - 1.2 - . 1.1 - IJ ~ i s. I i i i i i 1.0 1.2 1.4 0.1 IJ lJ sem 2Nsperzm peQNT DEDMD) V. C. Summer Figure 15.1-3 Normalized RCCA O Reactivity Worth vs. Percent Insertion
i 1.0 0.9 - IJ - l I 1.7-1 1.1 - l I1.5-a i l l0,4-0.3 - 1 I h-1 l 1.2 - N l ~ l 0.1 - 1 I I 0.0 i i i i i i i 0J 1.0 2.0 1.0 4.0 Tot not xc ac2Asz (szcxs) V. C. Sumer Figure 15.1-4 Normalized RCCA Bank Worth vs. Time After Trip
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0 , i i i i i i i i 0 20 40 40 to 100 PDCDU P0MR V. C. Sumer Figure 15.1-5 Doppler Power O Coefficient Used In Accident Analyses
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i j 10-3 I l 11l_,_l l I i 1 I l ll l 1 'l I , 3 I 10-1 10 0 101 102 10 TIME AFTER SHt/TDOWN (SECONDS) i V. C. Sumer j Figure 15.1-7 1979 ANS Decay Heat
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15.2 CONDITION II - FAULTS OF MODERATE FREQUENCY , p , d These faults result at worst in reactor shutdown with the plant being capable of returning to operation. By definition, these faults (or events) do not I-propagate to cause a more serious fault, i.e., a Condition III or IV fault. ;
~
In addition, Condition II events are not expected to result in fuel rod j failuresorreactorcoolantsystem(RCS)overpressurization. For the purposes l of this report the following faults have been grouped into these categories: i Uncontrolled rod cluster control assembly (RCCA) bank withdrawal (1) from a suberitical condition (2) Uncontrolled RCCA bank withdrawal at power (3) RCCA misoperation (4) Uncontrolled boren dilution (5) Partial loss of forced reactor cociant flow : l (6) Startup of an inactive reactor coolant loop (7) Loss of external electrical load and/or turbine trip l (8) Loss of normal feedwater l (9) Loss of offsite power to the station auxiliaries (station blackout) l l -(10) Excessive heat removal due to feedwater s'ystem malfunctions l (11) Excessive load increase i (12) Accidental RCS depressurization O is2ov:to/os2ssa 15.2-1 ! L
l (13) Accidental main steam system depressurization (14) Spurious operation of safety injection system (SIS) at power. Each of these faults of moderate frequency are analyzed in this section. In general, each analysis includes an identification of causes and description of the accident, an analysis of effects and consequences, a presentation of results, and relevant conclusions. An evaluation of the reliability of the reactor protection system actuation following initiation of Condition 11 events has been completed and is presented in Reference 1 for the relay protection logic. Standard reliability engineering techniques were used to assess the likelihood of the trip failure due to random component failures. Common-mode failures were also qualitatively investigated. It was concluded from the evaluation that the likelihood of no trip following initiation of Condition 11 events is extremely
-7 small (2 x 10 derived for random component failures). The reliability of the solid-state protection system has also been evaluated using the same methods. The calculated reliability is of the same order of magnitude as that obtained for the relay protection logic.
Hence, because of the high reliability of the protection system, no special provision is included in the design to cope with the consequences of Condition 11 events without trip. The time sequence of evnnts for the Condition II faults are shown in Table 15.2-1. O 132ov:t o/o412as 15.2-2
a.2.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical Condition ' g* . 15.2.1.1 Identification of Causes and Accident Description An RCCA withdrawal accident is defined as an uncontrolled increase in reactivity in the reactor core caused by withdrawal of RCCAs resulting in a power excursion. Such a transient could ha caused by a malfunction of the reactor control or control rod drive systems. This could occur with the reactor at either suberitical, hot zero power, or at power. The at power case is discussed in Section 15.2.2. Although the reactor is normally brought to power from a suberitical condition by means of RCCA withdrawal, initial startup procedures with a clean core call for boron dilution. The maximum rate of reactivity increase in the case of boron dilution is less than that assumed in this analysis (see Section 15.2.4). The RCCA drive mechanisms are wired into preselected bank configurations that are not altered during core reactor life. These circuits prevent the assemblies from being withdrawn in other than their respective banks. Power supplied to the banks is controlled so that no more than two banks can be l withdrawn at the same time. The RCCA drive mechanisms are of the magnetic latch type and coil actuation is sequenced to provide variable speed travel. l The maximum reactivity insertion rate analyzed in the detailed plant analysis , is that occurring with the simultaneous withdrawal of the two control banks l having the maximum combined worth at maximum speed. l l The neutron flux response to a continuous reactivity insertion is [ characterized by a very fast rise terminated by the reactivity feedback effect l. i of the negative Doppler coefficient. This self-limitation of the power burst i is of primary importance since it limits the power to a tolerable level during the delay time for protection action. Should a continuous RCCA withdrawal accident occur, the transient will be terminated by the following automatic features of the reactor protection system: 1stovao/os2ssa 15.2-3 I
15.2.1.1.1 Source Range High Neutron Flux Reactor Trip The source range high neutron flux reactor trip is actuated when either of two independent source range channels indicates a neutron flux level above a ' preselected manually adjustable setpoint. This trip function may be manually bypassed when either intermediate range flux channel indicates a flux level above a specified level. It is automatically reinsta'ted when both intermediate range channels indicate a flux level below a specified level. 15.2.1.1.2 Intermediate Range High Neutron Flux Reactor Trio The intermediate range high neutron flux reactor trip is actuated when either of two independent intermediate range chennels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when two of the four power range channels give readings above approximately 10% of full power and is automatically reinstated when three of the four channels indicate a power below this value. 15.2.1.1.3 PowerRangeHighNeutronFluxReactorTrio(LowSetting The power range high neutron flux trip (low setting) is actuated when two-out-of-four power range channels indicate a power level above approximately 25% of full power. This trip function may be manually bypassed when two of the four power range channels indicate a power level above h approximately 10% of full power and is automatically reinstated when three of the four channels indicate a power level below this value. 15.2.1.1.4 Pcwer Range High Neutron Flux Reactor Trip (High Setting) The power range high neutron flux reactor trip (high setting) is actuated when two-out-of-four power range channels indicate a power level above a preset setpoint. This trip function is always active. In addition, control rod stops on high intermediate range flux level (one-of-two) and high power range flux level (one-out-of-four) serve to discontinue rod withdrawal and prevent the need to actuate the intermediate range flux level trip and the power range flux level trip, respectively. O 1 132oao/o32ses 15.2-4
J . 15.2.1.1.5 High Neutron Flux Rate Trip p The high neutron flux rate trip is actuated when the rate of change in power V exceeds the positive or negative setpoint in two-out-of-four power range channels. This function is always active. 15.2.1.2 Analysis of Effects and Consequences The analysis of the uncontrolled red withdrawal from suber.itical accident is performed in three stages: first a core average nuclear power transient calculation is performed, followed by an average core heat transfer calculation, and finally a DNBR calculation. The core average nuclear power transient calculation is performed using a spatial neutron kinetics code, TWINKLE (2) , to determine the average power generation with time including the various total core feedback effects, i.e., Doppler and moderator reactivity. The average heat flux and temperature transients are determined by performing a fuel rod transient heat transfer calculation in FACTRAN(3) , The average heat flux is next used in THINCN) for the transient DNBR calculation. The core axial power distribution is severely peaked to the bottom of the core O fer 18e 14mitine transfent. T8e w-3 Dhs cerreiatien is usee to evaieate DheR in the span between the lower non-mixing vane grid. The WRB-1 correlation (LOPAR fuel) and the WRB-2 correlation (VANTAGE 5 fuel) remain applicable for the rest of the fuel assembly. In order to give conservative results for a startup accident, the following assumptions are made concerning the initial reactor conditions: (1) Since the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly dependent on the Doppler coefficient, conservative values (1ew absolute magnitude) as a function of power are used. See Section 15.1.5 and Table 15.1-4. l (2) Contribution of the moderator reactivity coefficient is negligible during the initial part of the trsnsient because the heat transfer ti a between the fuel and the moderator is much 1cnger than the O . Istovao/o422ss 15.2-5
ceutron flux response tine. However, after the initial neutron flux peak, the succeeding rate of power increase is affected by the moderator reactivity coefficient. A conservative value, given in Table 15.1-4, is used in the analysis to yield the maximum peak heat flux. (3) The reactor is assumed to be at hot zero power. , This assumption is more conservative than that of a lower initial system temperature. The higher initial system temperature yields a 1. rger fuel-water heat transfer coefficient, largar specific heats, and a less negative (smaller absolute magnitude) Doppler uoefficient, all of which tend to reduce the Doppler feedbs.ch erfect thereby increasing the neutron flux peak. The initial effective multiplication factor is assumed to be i since this results in maximum neutron flux peaking. (4) Reactor trip is assumed to be initiated by power range high neutron flux (low setting). The mest adverso combination of instrument and setpoint errors, as well as delays for trip signel actuation and RCCA release, is taken into account. A 10% increase is assumed for the power range flux trip setpoint, raising it from the nominal g value of 25 to 35%. Previous results, however, show that the rise in neutron flux is so rapid that the effect of errors in the trip setpoint on the actual time at which the rods are released is negligible. In addition, the reactor trip insertion characteristic is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn position. See Section 15.1.4 for RCCA insertion characteristics. (5) The maximum positive reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two control banks having the greatest combined worth at maximum speed (45 inches / minute). Control rod drive mechanism design is discussed in Section 4.2.3 of the FSAR. 1 inotio/o422ss 15.2-6 l l l
^
(6) The initial power level was assumed to be below the power level
- expected for any shutdown condition. The combination of highest ,
reactivity insertion rate and lowest initial power produces the highest peak heat flux. (7) Two reactor coolant pumps are assumed to be operating. l 15.2.1.3 Results The calculated sequence of events for this accident is shown on Table 15.2-1. Figures 15.2.1-1 and 15.2.1-2 show the transient behavior for the indicated reactivity insertion rate with the accident terminated by reactor trip at 35% nominal power. This insertion rate is greater than that for the two highest 4 worth control banks, both assumed to be in their highast incremental worth region. Figure 15.2.1-1 shows the nuclear power transient. The nuclear power overshoots the full power nominal value but this occurs for only a very short lp time' period. Hence, the energy release and the fuel temperature increase are
\._/ relatively small. The thermal flux response, of interest for departure from ,
nucleate boiling (DNB) considerations, is shown on Figure 15.2.1-1. The ; beneficial effect on the inherent thermal lag in the fuel is evidenced by a , peak heat flux less than the full power nominal value. ( l
. Figure 15.2.1-2 shows the response of the hot spot fuel average and clad l
temperatures. The hot spot fuel average temperature increases to a value ! lower than the nominal full power value. j 15.2.1.4 Conclusions , j In the event of an RCCA withdrawal accident from the suberitical condition, the cora and the RCS are not adversely affected since the combinatien of l thermal power and the coolant temperature result in a departure from nucleate i boiling ratio (DNBR) greater than the design limit value. Thus, no fuel or clad damage is predicted as a result of DNB. i j !O . inov.to/o42:ss 15.2-7 I
15.2.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal at Power 15.2.2.1 Identification of Causes and Accident Description j Uncontrolled RCCA bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from the steam g.enerator lags behind the core power generation until the steam generator pressure reaches the reli f or safety valve setpoint, there is a net increase in the reactor coolant temperature. Unless terminated by manual or automatic action, the power i mismatch and resultant coolant temperature rise would eventually result in DNB. Therefore, in order to avert damage to the cladding, the reactor protection system is designed to terminate any such transient before the DNBR falls below the safety analysis limit values. The automatic features of the reactor protection system that prevent core damage following the postulated accident include the following: (1) The power range neutron flux instrumentation acte:tes a reactor trip if two-out-of-four channels exceed a high flux setpoint; (2) The reactor trip is actuated if any two-out-of-three AT channels exceed an overtemperature oT setpoint. This setpoint is automatically varied with axial power iebalance, coolant temperaturo, and pressure to protect against DNB; (3) The reactor trip is actucted if any two-out-of-three oT channels exceed an overpower AT setpoint to ensure that the allowable heat' generation rate (kw/ft) is not exceeded; (4) A high pressurizer pressure rac.; tor trip actuated fro.n any two-out-of-three pressure channels th6t are set at a fixed point. This set pressure is less than the set pressure for the pressurizer safety valves; O 1320v:io/osissa 15.2-8
(5) A high pressurizer water level reactor trip actuated from any two-out-of-three level channels that are set at a fixed point. g . In addition to the above listed reactor trips, there are the following RCCA withdrawal blocks: , (1) High neutron flux (one-out-of-four); . (2) Overpower aT (two-out-of-three);
-(3) Overtemperature AT (two-out.-of-three).
The manner in which the combination of overpower and overtemperature AT trips provide protection over the full range of RCS conditions is described in Chapter 7. Figure 15.1-1 presents allowable reactor coolant loop average temperature and ai for the design power distribution and flow as a functier, of primary coolant pressure. The boundaries of operation defined by the overpower AT trip an.<i the overtemperature AT are represented as protection lines on this diagram. The protection lines are drawn to include all ' adverse instrumentation and setpoint errors so that under nominal conditions a trip l would occur well within the area bounded by these lines. The utility of this l diagram is in the fact that the limit imposed by a given DNBR can be l represented as a line. The DNB lines represent the locus of conditions for which the DNBR equals the safety analysis limit value. All points below and l to the left of a DNB line for a given pressure have a DNBR greater than the L limit. The diagram shows that DNB is prevented for all cases if the area enclosed with the maximum protection lines is not traversed by the applicable l DNBR line at any peint. The area of permissible operation (power, pressure, and temperature) is L bounded by the combination of reactor trips: high neutron flux (fixed i setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint); overpower and overtemperature AT (variable setpoints). i O 132ov:1o/o32ssa 15.2-9 , t !
15.2.2.2 Analysis of Effects and Consecuences The uncontrolled RCCA bank withdrawal at power transient is analyzed by the LOFTRAN code (4) This code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. The core limits as illustrated on Figure 15.1-1 are used as input to LOFTRAN to determine the minimum DNBR during the transient. This accident is analyzed with the Improved Thermal Design Procedure as described in Reference 5. In order to obtai anservativo results, the following assumptions are made: (1) Initial conditions of nominal core power and reactor coolant average temperatures and nominal reactor coolant pressure are assumed. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5; (2) Reactivity Coefficients - two cases are analyzed: (a) Hinimum reactivity feedback. A positive moderator coefficient O of reactivity of +7 pcm/*F is assumed. A variable Doppler power coefficient with core power is used in the analysis. A conservatively small (in absolute magnitude) value is assumed; (b) Maximum reactivity feedback. A conservatively large positive I moderator density coefficient and a large (in absolute magnitude) r:agative Doppler power coefficie- t are assumed; (3) The reactor trip on high neutron flux is assumed to be actuated at a l conservative value of 118% of nominal full power. The AT trips include all adverse instrumentation and setpoint errors, while tne delays for the trip signal actuation are assumed at their maximun values; O iazov;1oto412ss 15.2-10
(4) The RCCA trip insertion characteristic is based on the assumption e that the highest worth assembly is stuck in its fully withdrawn pos.i tion; (5) The maximum positive reactivity insertion rate is greater than that ; which would be obtained from the simultaneous withdrawal of the two control red banks having the maximum combined we'rth at maximum speed. i The effect of RCCA movement on the axial core power distribution is accounted for by causing a decrease in overtemperature and overpower AT trip setpoints proportional to a decrease in margin to DNB. 15.2.2.3 Results Figures 15.2.2-1 and 15.2.2-2 show the response of nuclear' power, pressure, average coolant temperature, and DNBR to a rapid RCCA withdrawal starting from i full power. Reactor trip on high neutron flux occurss 'hortly after the start of the accident. Since this is rapid with respect to the chermal time constants of the plant, small changes in T avg and pressure result and a large margin to DNB is maintained. The response of nuclear power, pressure, average coolant temperature, and DNBR for a slow control rod assembly withdrawal from full power is shown on Figures l 15.2.2-3 and 15.2.2-4. Reactor trip on overtemperature aT occurs after a ! longer period and the rise in temperature and pressure is consequently larger i than for rapid RCCA withdrawal. Again, the minimum DNBR is never less than the safety analysis limit values. Figure 15.2.2-5 shows the minimum DNBR as a function of reactivity insertion ! rate from initial full power operation for the minimum and for the maximum reactivity feedbacks. It can be seen that two reactor trip channels provide .- protection over the whole range of reactivity insertion rates. These are the high neutron flux and overtemperature AT trip channels. The minimum DNBR is l l never less than the safety analysis limit values. l l l O l l 1320w1o/o413ss 15.2-11 4 l
~~
-T ._ . -- - _ _ __ .
Figures 15.2.2-6 and 15.2.2-7 show the minimum DNBR as a function of reactivity insertion r:te for RCCA withdrawal incidents starting at 60% and 10% power, respectively. The results are similar to the 100% power case, - except that as the initial power is decreased, the range over which the overtemperature AT trip is effective is increased. In neither case does the DNBR fall below the safety analysis limit values. The shape of the curves of minimum DNB ratio versus reactivity insertion rate in the reference figures is due both to reactor core and coolant systea transient response and to protection system action in initiating a reactor trip. Referring to Figure 15.2.2-7, for example, it is noted that:
- 1. For reactivity insertion rates ~ abo <e 20 pcm/sec reactor trip is initiated by the high neutron flux trip for the minimum reactivity feedback cases. The neutron flux level in the core rises rapidly for these insertion rates while core heat flux and coolant system temperature lag behind due to the thermal capacity of the fuel and coolant system fluid. Thus, the reactor is tripped prior to a significant increase in heat flux or water temperature with resultant high minimum DNB ratios h
during the transient. As the reactivity insertion rate decreases, core heat flux and coolant temperatures can remain more nearly in equilibrium with the neutron flux. Minimum DNBR during the transient thus decreases with decreasing insertion rate.
- 2. The overtemperature AT reactor trip circuit initiates a reactor trip when measured coolant loop AT exceeds a setpoint based on me.asured Reactor Coolant System average temperature and pressure. It is important to note that the average temperature contribution to the circuit is lead-lag compensated in order to decrease the effect of the thermal capacity of the Reactor Coolant System in response to power increases.
- 3. For reactivity insertion rate below ~ 20 pcm/see the overtemperature AT trip terminates the transient.
O 132cv:io/o32ssa 15.2-12
~'
1 For reactivity insertion rates between ~ 20 pcm/see and ~ 5 pcm/sec the effectiveness of the overtemperature AT trip' increases (in terms of [] V increased minimum DNBR) due to the fact that with lower insertion rates the pcwer increase rate is slower, the rate of .-ise of average coolant temperature is slower and the system lags and de, lays become less significant.
- 4. For reactivity insertion rates less than ~ 5 pcm/sec, the rise in the reactor coolant temperature is sufficiently high so that the steam
^
generator safety valve setpoint is rea:hed prior to trip. Opening of these valves, which act as an additional heat load on the Reactor Coolant System, sharply decreases the rate of increase of Reactor Coolant System average temperature. This decrease in rate of increase of the average coolant system temperature during the transient is accentuated by the lead-lag compensation causing the overtemperature AT trip setpoint to be reached later with a resulting lower minimum DNBR. l For' transients initiated from higher power levels (for example, see Figure L 15.2.2-5) the effect described in item 4 above, which results in the sharp peak in minimum DNBR at approximately 5 pcm/sec, does not occur since the ( steam generator safety valves are not actuated prior to trip. Figures 15.2.2-5,15.2.2-6, and 15.2.2-7 illustrate minimum DNBRs calculated i for minimum and maximum reactivity feedback. L l Since the RCCA withdrawal at power incident is an overpower transient, the l fuel temperatures rise during the transient until after reactor trip occurs. l For high reactivity insertion rates, the overpower transient is fast with respect to the fuel rod thermal time constant, and the core heat flux lags l behind the neutron flux response. Due to this lag, the peak core heat flux l does not exceed 118 percent of its nominal value (i.e., the high neutron flux trip setpoint assumed in the analysis). Taking into account the effect of the RCCA withdrawal on the axial core power distributio'n, the peak fuel centerline i temperature will still remain below the fuel melting temperature. l l 1 O 1 1320<:1o/os2ssa 15.2-13 l
~~
For slow reactivity insertion rates, the core heat flux remains more nearly in equilibrium with the neutron flux. The overpower transient is terminated by the overtemperature AT reactor trip before a DNB condition is reached. The - peak heat flux again is maintained below 118 percent of its nominal value. Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel centerline temperature will remain below the fuel melting temperature. _ Since DNB does not occur at any time during the RCCA withdrawal at power transient, the ability of the primary coolant to remove heat from the fuel red is not reduced. .Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient. The calculated sequence of events for this accident is shown on Table 15.2-1. With the reactor tripped, the plant eventually returns to a stable condition. The plant may subsequently be cooled down further by following normal plant shutdown procedures. 15.2.2.4 Conclusic'ns The high neutron flux and overtemperature AT trip channels provide adequate protection over the entire range of possible reactivity insertion rates; i.e., the minimum value of DNBR is always larger than the safety analysis limit values. I i O 132ov:io/o32ssa 15.2-14 1
15.2.3 Rod Cluster Control Assembly Misoperation This section discusses RCCA misoperation that can result either from system
/(q) malfunction or operator error.
15.2.3.1 Identification of Causes and Accident Description RCCA misalignment accidents include: (1) One or more dropped RCCAs within the same group; (2)AdroppedRCCAbank; (3) Statically misaligned RCCA. Each RCCA has a position indicator channel that displays the position of the assembly. The displays of assembly positions are grouped for the operator's convenience. Fully inserted assemblies are further indicated by a rod at bottom signal, which actuates a local alarm and a control room annunciator. Group demand position is also indicated. RCCAs are always moved in preselected banks, and the banks.are always moved in the same preselected sequence. Each bank of RCCAs is divided into two l l groups. The rods comprising a group operate in parallel through multiplexing l thyristors. The two groups in a bank move sequentially such that the first [ group is always within one step of the second group in the bank. A definite ( schedule of actuation (or deactuation of the stationary gripper, movable' gripper, and lift coils of a mechanism) is required to withdraw the RCCA attached to the mechanism. Since the stationary gripper, movable gripper, and l ; I lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure that would cause rod withdrawal would affect a ' minimum of one group. Mechanical failures are in the direction of insertion, or immobility. l l o 1s2ov:1o/os2ssa 15.2-15
~
A dropped RCCA, or RCCA bank, is detected by: (1) A sudden drop in the core power level as seen by the nuclear - instrumentation system; (2) Asymmetric power die h ibution as seen on oct-of-core neutron detectors or core-exit thermocouples; _ (3) Rod at bottom signal; (4) Rod deviation alarm; (5) Rod position indication; (6) Negative neutron flux rate trip circuitry. Misaligned RCCAs are detected by: (1) Asymmetric power distribution as seen on out-of-core neutron detectors or core-exit thermocouples; (2) Rod deviation alarm; (3) Rod position indicators. The deviation alarm alerts the operator whenever an individual rod position signal deviates from the other rods in the bank by a preset limit. If the rod deviation alarm is not operable, the operator is required to take action as required by the Technical Specifications (0) . If one or more rod position indicator channels should be out of service, detailed operating instructions are followed to ensure the alignment'of the nonindicated RCCAs. The operator is also required to take action as required by the Technical Specifications. O is2ov:to/os2ssa 15.2-16
15.2.3.2 Analysis of Effects and Consequences i g t,~ Method of Analysis (1) One or More Dropped RCCAs from the Same Group For evaluation of the dropped RCCA event, the transient system response .is calculated using the LOFTRAN code. The code simulates the neutron kinetics, RCS, pressurizer, pressurizer rA ief and j safety valves, pressurizer spray, steam generator, and steam ! I generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. Statepoints are calculated and nuclear models are used to obtain a { hot channel factor consistent with the primary system conditions and reactor power. By incorporating the primary conditions from the transient and the hot channel factor from tha nuclear analysis, the DNB design basis is shown to be met using the THINC code U) . The transient resp,onse, nuclear peaking factor analysis, and DNB design d basis confirmation are performed in accordance with the methodology described in Reference 10. (2) Dropped RCCA Bank Analysis is not required since the dropped RCCA bank results in a trip. (3) Statically Misaligned RCCA Steady state power distributions are analyzed u~ sing the computer codes as described in Table 4.1-2 of the FSAR. The peaking factors i are then used as input to the THINC code to calculate the ONBR. O V is20mio/o32ssa 15.2-17 l
15.2.3.3 Results
~
(1) One or More Dropped RCCAs Single or multiple dropped RCCAs within the,same group result in a negative reactivity insertion that may be detected by the power range negative neutron flux rate trip circuitry.- If detected, the reactor is tripped within approximately 2.7 seconds following the drop of the RCCAs. The core is not adversely affected during this period since power is decreasing rapidly. Following reactor trip, normal shutdown procedures are followed. The operator may manually retrieve the RCCA by following approved operating procedures. For those dropped RCCAs that do not result in a reactor trip, power may be reestablished either by reactivity feedback or control bank withdrawal. Following a dropped rod event in manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is renotonic, thus removing power overshoot as a ccncern and establishing the automatic rod control mode of operation as the limiting case. For a dropped RCCA event in the automatic rod control mode, the rod control system detects the drop in power and initiates control bank withdrawal. Power overshoot may occur due to this action by the automatic rod controller after which the control system will insert the control bank to restore nominal power. Figures 15.2.3-1 and 15.2.3-2 show a typical transient response to a dropped RCCA (or RCCAs) in automatic control. In all cases, the minimum DN3R remains above the safety analysis limit value. (2) Dropped RCCA Bank A dropped RCCA bank typically results in a reactivity insartion of greater than 500 pcm which will be detected by the power range O 132ov:1o/o32ssa 15.2-18
negative neutron flux rate trip circuitry. The reactor is tripped within approximately 2.7 seconds following the drop of a RCCA bank.
'v The core is not adversely affected during this period since power is decreasing rapidly. Following the reactor trip, normal shutdown procedures are followed to further cool down the plant. Any action required of the operator to maintain the pIant in a stabilized ,
condition will be in a' time frame in excess of 10 minutes following the incivent. (3) Statically Misaligned RCCA The most severe misalignment situations with respect to DNBR at significant power levels arise from cases in which one RCCA is fully inserted, or where Bank D is fully inserted with one RCCA fully withdrawn. Multiple independent alarms, including a bank insertion limit alarm, alert the operator well before the postulated conditions are approached. The bank can be inserted to its ! insertion limit with any one assembly fully withdrawn without the DNBR falling below the safety analysis limit value. l q l The insertion limits in the Technical Specifications may vary from l l time to time depending on a number of limiting criteria. It is i l preferable, therefore, to analyze the misaligned RCCA case at full l power for a position of ti'e control bank as deeply inserted as the l criteria on minimum DNBR and power peaking factor will allow. The full power 'nsertion limits on control Bank D must then be chosen to , be above that position and will usually be dictated by other criteria. Detailed results will vary from cycle to cycle depending on fuel arrangements. For this RCCA misalignment, with Bank D inserted to its full power insertion limit and one RCCA fully withdrawn, ONBR does not fall below the safety analysis limit value. This case is analyzed assuming the initial reactor power, pressure, and RCS temperatures are at their nominal values but with the increased radial peaking factor associated with the misaligned RCCA. O 1320ao/o32ssa 15.2-19 _. l
DNB calculations have not been performed specifically for RCCAs I missing from other banks; however, power shape calculations have g been done as required for the fully withdrawn analysis. Inspection W- l of the power shapes shows that the DNB and peak kW/ft situation is less severe than the Bank D case discussed above assuming insertion limits on the other banks equivalent to a Bank D full-in insertion limit. _ For RCCA misalignments with one RCCA fully inserted, the DNBR does not fall below the limit value. This case is analyzed assuming the initial reactor power, pressure, and RCS temperatures are at their nominal values, but with the increased radial peaking factor associated with the misaligned RCCA. DNB does not occur for the RCCA misalignment incident and thus the ability of the primary coolant to remove heat from the fuel rod is not reduced. The peak fuel temperature corresponds to a linear heat generation rate based on the radial peaking factor penalty associated with the misaligned RCCA and the design axial power distribution. The resulting linear heat generation is well below that which would cause fuel melting. Following the identification of an RCCA group misalignment condition by the coerator, the operator is required to take action as required by the plant Technical Specifications and operating instructions. 15.2.3.4 Conclusions For all cases of dropped RCCAs or dropped banks, fcr which the reactor is tripped by the pcwer range negative neutron flux rate trip, there is no reduction in the margin to core thermal limits and, consequently, the DNB design basis is met. It is shown for all cases which do not result in reactor trip that the DNBR remains greater than the safety analysis limit value and, therefore, the DNB design basis is met. O 132ovn o/os2ssa 15.2-20
. - . -_=_ __ -. - . _ - - - . - . . - . .. -
For all cases of any RCCA inserted, or Bank D inserted to its rod insertion limits with any single RCCA in that bank fully withdrawn (static O misalignment), the DNBR remains greater than the safety analysis limit value. J O . 1 l l 1 i 4 O 132ov:1o/cateaa 15.2-21
- . . ........ = = _ _
- _ :. .. . = u ._ - .,_ = - .._ - = .... . -
15.2.4 Uncontrolled Baron Dilution 15.2.4.1 Identification of Causes and Accident Description Reactivity can be added to the core by feeding primary grade water into the reactor coolant system (RCS) via the reactor makeup portion of the chemical and volume control system (CVCS). Boron dilution is 'a manual operation under strict administrative controls with procedures calling for_a limit on the rate and duration of dilution. A boric acid blend system is provided to permit the operator to match the baron concentration of reactor coolant makeup water during normal charging to that in the RCS. The CVCS is designed to limit, even under various postulated failure modes, the potential rate of dilution to a value which, af ter indication through alarms and instrumentation, provides the operator sufficient time to correct the situation in a safe and orderly manner. The opening of the primary water makeup control valve provides makeup to the RCS which can dilute the reactor coolant. Inadverter.t dilution from this source can be readily terminated by closing the control valve. In order for makeup water to be added to the RCS at pressure, at least one charging pump must be running in addition to a primary makeup water pump. The rate of addition of unborated makeup water to the RCS is limited by a flow limiting orifice between the reactor makeup water pumps and the boric acid blender. As demonstrated by tests at the plant, flow is within the bounds of unborated water used in analyses in this section. The boric acid from the boric acid tank is blended with primary grade water in the blender and the composition is determined by the preset flow rates of boric acid and primary grade water on the control board. In order to dilute two separate operations are required:
- 1. The operator must switch from the automatic makeup mode to the dilute mode;
- 2. The start /stop switch is in the start position.
l l 132ov:t o/o32ssa 15.2-22
Omitting either step would prevent dilution. The status of the RCS makeup is continuously available to the operator by:
- 1. Indication of the boric acid and blended flow rates,
~
- 2. CVCS and RMWS pump status lights,
- 3. D'eviation alarms if the boric acid or blended flow rates deviate by more than 10% from the preset values, Indication of a dilution event is available to the operator by:
- 1. Source Range Neutron Flux - when reactor is suberitical;
, a. High flux at shutdown alarm. A separate alarm will be provided for I each channel, n b. Indicated source range neutron flux count rates, and V .
- c. Audible source range neutron flux count rate. 3
- 2. With the reactor critical; 1 s
- a. Axial flux difference alarm (reactor power > 50% RTP),
- b. Control rod insertion limit low and low-low alarms, l
- c. Overtemperature AT alarm (at power),
- d. Overtemperature AT turbine runback (at power),
- e. Overtemperature AT reactor trip, and
- f. Power range neutron flux - high, both high and low setpoint Reactor Trips.
132ov:1o/o32ssa 15.2-23
^ ^ ^ ~'
15.2.4.2 Analysis of Effects and Consequences To cover all phases of plant operation, boren dilution during refueling, cold ' shutdown, hot standby, startup, and power operation are considered in this analysis. The hot shutdown case is bounded by the analysis for cold shutdown and hot standby. Table 15.2-1 contains the time sequence of events for this accident.
- 1. Dilution During Refueling An uncontrolled boron dilution accident based on a failure in the primary water makeup system cannot occur during refueling. This accident is
-v prevented by administrative controls which isolate the RCS from the potential sourc.e of unborated water. Valves 8454, 8441, 8430, and 8439 will be locked closed during refueling operations. These valves will block the flow paths which could allow unborated maket:p water to reach the RCS. Any makeup which is required during refueling , vill be added to the Reactor Coolant System by unlocking these valves as appropriate and initiating tbs required blended makeup water flow. After the required volume of blended makeup flow has been added, these valves will again be locked closed. An alternate source of h borated water that could be used is from the Refueling Water Storage Tank to the Charging Pump suction. The most limiting alternate source of unborated water is from the boron
. thermal regeneration system (BTRS). For this case, highly borated PCS water is depleted of boron as it passes through the BTRS and is returned via the volume control tank. The following conditions are assumed for an uncontrolled boron dilution during refueling.
Technical Specifications require the reactor to be borated to at least 2,000 ppm er shutdown by at least 5.0 percent ak/k at refueling. If an inadvertent dilution from the BTRS occurs during refueling with the reactor vessel head off and the refueling cavity filled with borated water O 1320c10/03298s 15.2-24
(i.e., in a condition to move fuel), the maximum dilution capability of ;
, the BTRS is insufficient to cause a return to criticelity. -
f The maximum dilution capability of the BIRS at these conditions is conservatively estimated te ce 250 ppm. However, the minimum change in boron concentration necessary to bring the react 6r critical at these conditions is conservatively estimated to be '800 ppm._ An initial boron [ concentration of 2500 ppm is assumed. Therefore, a dilution to criticality from the BTRS at these refueling conditions cannot occur.. ; The most limiting conditions for an inadvertent boron dilution from the BTRS during refueling occur when the reactor head is unbolted but in place and the reactor coolant level is at the vessel / head junction. The dilution capability of the BTRS at these conditions is sufficient to cause a return to criticality. The minimum volume in the reactor coolant system corresponding to this condition is conservatively estimated to be 3300 ft.3The critical boron concentration is conservatively estimated to be , 1700 ppm.
- 2. Dilution During Cold Shutdown Technical Specifications specify the required shutdown margin as a function of RCS boron concentration during cold shutdown. The specified l
P shutdown margin ensures sufficient time for the operator to terminate the dilution. For a boren concentration of 1000 ppm, the required shutdown f margin is 2.0% ak. If the reactor is in cold shutdown and on the
- residual heat removal system with RCS piping filled and vented, the following conditions are assumed for an uncontrolled baron dilution.
Dilution flow is assuined to be a maximum of 150 gpm, which is the capability of one primary water makeup pump to deliver unborated water to the RCS. Mixing of the reactor coolant is accomplished by the operation ( of one residual heat removal pump. l D 15.2-25 is2ov:to/o32ses l r .
~
3 A volume of 4816.2 ft in the reactor coolant system is used. This corresponds to the active volume of the' reactor coolant system minus the pressurizer volume, while on the residual heat removal system. If the reactor is in cold shutdown and the RCS water level is drained down from a filled and vented condition while on RHR, an inadvertent dilution is prevented by administrative controls which isolate the RCS from the potential source of unborated water. Valves 8454, 8441, 8430, and 8439 will be locked closed during operations in these conditions. These valves block all flow paths that could allow unborated makeup water to reach the RCS. Any make~up which is required will be added to the Reactor Coolant System by unlocking these valves as appropriate and initiating the required blended makeup water flow. After the required volume of blended makeup flow has been added, these valves will again be locked closed. An alternate source of borated water which may be used is from the Refueling Water Storage Tank to the Charging Nmp suction.
- 3. Dilution during Hot Standby Technical Specifications specify the required shutdown margin as a function of RCS boren concentration. For a boron concentration of 1500 ppm, the required shutdown margir. is conservatively estimated to be 2.85% ak.
The following conditions are assumed for a continuous boron dilution during hot standby: 1 Dilution flow is assumed to be a maximum of 150 gpm, which is the capability of one primary water makeup pump to deliver unborated water to I the RCS. 3 A minimum RCS water volume of 5050 ft is used. This is a conservative l estimate of the active RCS volume with'one reactor coolant pump operating. l l l 9 l 132omio/o32ssa 15.2-26
- 4. Dilution During Startup Prior to startup, the RCS is filled with borated water at a boron concentration of 2200 ppm. This is a conservative estimate with the reactor at a 1.77% ak/k shutdown otargin at 557'F.
Dilution flow is assumed to be a maximum of 150 gpm, which is the capability of one primary water makeup pump to deliver unborated water to the RCS, A minimum volume of 76823 ft in the reactor coolant system is used. This is a conservative estimate of the active volume of the RCS excluding the pressurizer.
- 5. Dilution During Full Power Operation During power operation, the plant may be operated two ways, under manual operator control or under automatic Tavg/ rod control. The Technical Specifications require three reactor coolant pumps operating and a shutdown margin of at least 1.77% ak/k. The RCS is conservatively
. assumed to be filled with borated water at a boron concentration of 2200 ppm.
I While the plant is in manual control, the dilution flow is assumed te be a l l maximum of 150 gpm, which is the capacity of one reactor makeup water pump .; to deliver unborated water to the RCS. When in automatic control, the dilution flow is limited by the maximum letdown flow (approximately l 125gpm). , 1 3 A minimum RCS water volume of 7682 ft is used. This is a conservative ; estimate of the active volume of the RCS excluding the pressurizer. j 1 O inzoao/osassa 15.2-27 i
- .=_ -_ - .,:_. , _:_.-. :_ - . . . - - - . _ . - - - . _
15.2.4.3 Conclusion Dilution During Refueling - During refueling, an inadvertent dilution from the reactor makeup water system is prevented by administrative controls which isolate'the RCS from the potential source of unborated makeup water. _ The most limiting conditions for an inadvertent dilution from the BTRS occur when the reactor vessel head is unbolted and the vessel water level is at the vessel /headjunction. The high flux at shutdown alarm, set at twice the background flux level measured by the source range nuclear instrumentation, is available at these conditions to alert the operator that a dilution event is in progress. For this case, the operator has 48 minutes from the high flux at shutdown alarm to recognize and terminate the dilution before shutdown margin is lost and the reactor becomes critical. Dilution During Cold Shutdown While in cold shutdown, the high flux at shutdown alarm set at twice the background flux level measured by the source range nuclear instrumentation, is available to alert the operator that a dilutien event is in progress. During the cold shutdown mode while operating on the residual heat remeval system (RHRS) with the RCS piping filled and vented, the shutdown margin requirement ensures that the operator has at least 13.6 minutes from the high flux at shutdown alarm to recognize and terminate the uncontrolled reactivity insertion before shutdown margin is lost. During the cold shutdown mode while operating on the RHRS with the RCS' drained down from a filled and vented condition, an inadvertent dilution is precluded by administrative controls which isolate the RCS from the potential source of unborated water. O 132ov:1o/o32ses 15.2-28
}
Dilution During Hot Shutdown (V3 Analysis for a dilution during hot' shutdown is bounded by the analysis for a dilution during cold shutdown and hot standby. Dilution During Hot Standby While in hot standby, the high flux at shutdown alarm, set at twice the background flux level measured by the source range nuclear instrumentation, is available to alert the operator that a dilution event is in progress. During hot standby, the shutdown margin requirement ensures that the operator . has at least 13.4 minutes from the high flux at shutdown alarm to recognize and terminate the uncontrolled reactivity insertion before shutdown margin is lost. Dilution During Startup in the event of an unplanned apprcach to criticality or dilution during power escalation while in the s'artup t mode, the operator is alerted to an ;
~
uncontrolled reactivity insertion by a reactor trip at the Power Range Neutron i Flux-High, low setpoint (nominally 25% RTP). After reactor trip there is at l least 20.6 minutes for operator action prior to return to criticality. . Dilution at Power i y During the at power mode with manual control, the operator is alerted to an uncontrolled reactivity insertion by an overtemperature AT trip. 19.0 l minutes are available from the trip for the operator to recognize and terminate the uncontrolled dilution. The sensitivity and alarm thresholds are already assumed to be degraded to the maximum extent allowable for the overtemperature AT trip function (see Section 15.2.2). L ,,,c _ ,,,.. 15.2 2, i l - L . _ .
During the at power mode with automatic control, the operator is alerted to an uncontrolled reactivity insertion by the rod insertion limit alarms. Two insertion limit alarms'are available: the first occurs when the rods are 10 steps above the insertion limit (LO Insertion Limit Alarm) and the second occurs at the insertion limit (Lo-Lo Insertion Limit Alarm). The analysis assumed that the operator is alerted to the need for action by the Lo-Lo Alarm although action would be taken when the first alarm occurs. Thus the analysis already assumes a 10 step allowance for rod position indicator inaccuracies. Even with this conservatism, there are still 23.0 minutes available from the time of alarm until all shutdown margin is lost. In addition to the above, other indications are available. The main indication would be a violation of the axial offset control band which could result in a reactor trip (reduction in overtemperature AT setpoint). O O 1320v;10/o32ssa 15.2-30
l i 15.2.5 Partial Loss of Forced Reactor Coolant Flow ~ l (V3 15.2.5.1 Identification of Causes and Accident Description A partial loss of coolant flow accident can result from a mechanical or
- i. electrical failure in a reactor coolant pump, or from a fault in the power supply to the pump. If the reactor is at power at the. time of the accident, the immediate effect of a loss of' forced reactor coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly.
The necessary protection against a partial loss of coolant flow accident is provided by the low primary coolant flow reactor trip that is actuated by two-out-of-three low flow signals in any reactor coolant loop. Above approximately 38% power (Permissive 8), low flow in any loop will actuate a reactor trip. Between approximately 10% power (Permissive 7) and the power level corresponding to Permissive 8 low flow in any two loops will actuate a reactor trip. Reactor trip on low flow is blocked below Permissive 7.
~
A reactor trip signal from the pump breaker position is also provided. When operating above Permissive 7,'a breaker open signal from any two pumps will actuate a reactor trip. This serves as a backup to the low flow trip. Reactor trip on reactor coolant pump breakers open is blocked below Permissive 7. Normal power for each pump is supplied through individual buses connected to the isolated phase bus duct between the generator circuit breaker and the main transformer. Faults in the substation may cause a trip of the main transformer high side circuit breaker leaving the generator to supply power to the reactor coolant pumps. When a generator circuit breaker trip occurs because of electrical faults, the pumps are automatically transferred to an alternate power supply and the pumps will continue to supply coolant flow to the core. Following any turbine trip where there are no electrical faults, the generator circuit breaker is tripped and the reactor coolant pumps remain connected to the network through the transformer high side breaker. Continuity of power to the pump buses is achieved without motoring the O is2 otto osassa 15.2-31
i generator since means are provided to isolate the generator without isolating the pump buses from the external power lines (e.g., a generator output breaker is provided as well as a station output breaker). I 15.2.5.2 Analysis of Effects and Consequences { 15.2.5.2.1 Method of Analysis _ The following case has been analyzed: All loops operating, one loop coasting down This transient is analyzed by three digital computer codes. First the LOFTRAN code is used to calculate the loop and core flow during the transient. The LOFTRAN code is also used to calculate the time of reactor trip, based on the calculated flows and the nuclear power transient following reactor trip. The FACTRAN code is then used to calculate the heat flux transier.t based on the nuclear power and flow from LOFTRAN. Finally, the THINC codeN) is used to calculate the mininum DNBR during the transien: based on the heat flux from FACTRAN and the flow from LOFTRAN. The DNBR transient presented represents-the minimum of the typical and thimble cells for Standard and VANTAGE 5 fuel, h 15.2.5.2.2 Initial Conditions The assumed initial operating conditions are the most adverse with respect to the margin to DNB, i.e., nominal steady state power level, nominal steady state pressure, and nominal steady state coolant average temperature. See Section 15.1.2 for an explanation of initial conditions. The accicent is analyzed using the Improved Thermal Design Procedure as described in Reference 5. l 15.2.5.2.3 Reactivity Coefficients A conservatively large absolute value of the Doppler-only power coefficient is used (see Table 15.1-4). The total integrated Doppler reactivity from 0 to , 100% power is assumed to be -0.016 Ak. l l O is2ovio/o32ssa 15.2-32
The least negative moderator temperature coefficient at full power (0 pcm/'F) is assumed since this results in the maximum hot spot heat flux during the ( initial part of the trar sient when the minimum DNBR is reached. 15.2.5.2.4 Flow Coastdown ' The flow coastdown analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance, and the pump characteristics and is based on high estimates of system pressuro losses to calculate the flow coastdown. 15.2.5.3 Results The calculated sequence of events is shown in Table 15.2-1. Figures 15.2.5-1 and 15.2.5-2 show the vessel flow coastdown, the faulted loop flow coastdown, the nuclear power and heat flux transient. The minimum DNBR is not less than the safety analysis limit value. A plot of DNBR vs. time is given in Figure 15.2.5-3 for the most limiting typical or thimble cell for Standard and VANTAGE 5 fuel. lj.2.5.4 Conclusions l The analysis shows that the DNBR will not decrease below the safety analysis limit values at any time during the transient. Thus, no core safety limit is l l violated. L l L 1 O is2ov1o/o414ss 15.2-33
. ,. --..e.-
15.2.6 Startup of an Inactive Reactor Coolant Loop In accordance with Technical Specification 3/4.4.1, V. C. Summer operation during startup and p:wer operation with less than three loops operating is not g permitted. This analysts is presented for completeness. 15.2.6.1 Identificati,on of Causes and Accident Description If a plant is operatirq eith one pump out of service, there is reverse flow through the loop due t> the pressure difference across the reactor vessel. The cold leg temperaturo in an inactive loop is identical to the cold leg temperature of the actir loops (the reactor core inlet temperature). If the reactor is operated at p<"er,.and assuming the secondary side of the steam generator in the inactive loop is not isolated, there is a temperature drop across the steam generator in the inactive loop and, with the reverse flow, the hot leg temperature of the inactive locp is lower than the reactor core inlet temoerature. Administrative procedures require that the unit be brought to a load of less than 25% of full power prior to starting a pump in an inactive loop in order to bring the inactive loop hot leg temperature closer to the core inlet temperature. Starting of an idle reactor coolant pump without bringing the inactive loop hot leg temperature close to the core inlet temperature would result in the injection of cold water into the core which causes a rapid reactivity insertion and subsequent power increase. . This event is classified as an ANS Condition II incident (an incident of moderate frequency) as defined in Section 15.0. Should the startup of an inactive reactor coolant pump at an incorrect
~
temperature occur, the transient will be terminated automatically by a reactor trip on low coolant loop flow when the power range neutron flux (two out of four channels) exceeds the P-8 setpoint, which has been previously reset for two loop operation. O u2cvoo/o42sas 15.2-34
i 15.2.6.2 Analysis of Effects and Consequences f'
'" ""' '"d *" d' * **"
- d**' '"' ' "^"
Code (4) is used to calculate the loop and core flow, nuclear power and core pressure and temperature transients following the startup of an idle pump. FACTRAN(3) is used to calcyste the core heat flux transient based on core flow and nuclear power f"or LOFTRAN. The THINC Code (7) is then used to calculate the DNBR during the transient based on system conditions (pressure, temperature, and flow) calculated by LOFTRAN and heat flux as calculated by FACTRAN. In order to obtain conservative results for the startup of an inactive pump accident, the following assumptions are made: (1) Initial conditions of maximum core power and reactor coolant average temperatures and minimum reactor coolant pressure resulting in minimum initial margin to DNB. These values are to be consistent with maximum steady state power level allowed with all but one loop in operation including appropriate allowances for calibration and s instrument errors. The high initial power gives the greatest
) temperature difference between the core inlet temperature and the inactive loop hot leg temperature.
(2) Following the start of the idle pump, the inactive loop flow reverses and accelerates to its nominal full flow value. (3) A conservatively large (absolute value) negative moderator temperature coefficient associated with end of life conditions. (4) A conservatively low (absolute value) negative Doppler power coefficient is used. (5) The initial reactor coolant loop flows are at the appropriate values for one pump out of service. O 132owlo/o42688 15.2-35
(6) The reactor trip is assumed te occur on low coolant flow when the power rarge neutron flux exceeds the P-d setpoint. The P-8 setpoint is conservatively assumed to be 74 percent of rat G r wer. h l 15.2.6.3 Results The results following the startup of an idle pump with the above listed ) assumptions are shown in Figures 15.2.6-1 through 15.2.6-4. As shown in these curves, during the first part of the transient, the increase in coro flow with cooler water results in an increase in nuclear power and a decrease in core average temperature. The minimum DNBR during the transient is considerably greater than the safety analysis limit values. Reactivity addition for the inactive loop startup accident is due to the decrease in core water temperature. During the transient, this decrease is due both to (1) the increase in reactor coolant flow and, (2) as the inactive loop flow reverses, to the coldor water entering the core from the hot leg side (colder temperature side prior to the start of tha transient) of the steam generator in the inactive loop. Thus, the reactivity insertion rate for this transient changes with time. The resultant core nuclear power transient, computed with consideration of both moderator and Doppler reactivity feedback effects, is shown on Figure 15.2.6-1. g The calculated sequence of eve .ts for this accident is shown in Table 15.2-1. The transient results illustrated in Figures 15.2.6-1 through 15.2.6-4 indicate that a stabilized plant condition, with the reactor tripped, is approached rapidly. Plant cooldown may subsequently be achieved by following normal shutdown procedures. 15.2.6.4 Conclusions The transient results show that the core is not adversely affected. There is considerable margin to the safety analysis DNER limit values; thus, no fuel or clad damage is predicted.
\
132ov:io/o329ss 15.2-36 i __
15.2.7 Loss of External Electrical Lead and/or Turbine Trip !l O
'd 15.2.7.1 Identification of Causes and Accident Description A major load loss on the plant can result from either a loss of external electrical load or from a turbine trip. For either case, offsite power is available for the continued operation of plaat components such as the reactor I
coolant pumps. The case of loss of all ac power (station _blatkout) is i analyzed in Section 15.2.9. , For a turbine trip, the reactor would be tripped directly (unless it is below approximately 50% power) from.a signal derived from the turbine autostop oil pressure and turbine stop valves. The automatic steam dump system accommodates the excess steam generation. Reactor coolant temperatures and pressure do not significantly increase if the steam dump system and pressurizer pressure control system are functioning properly. If the turbine condenser were not available, the excess steam generation would be dumped to the atmosphere. Additionally, main feedwater flow would be lost if the turbine condenser were not available. For this situation, steam generator level would be maintained by the' emergency feedwater system. For a loss of external electrical load without subsequent turbins trip, no direct reactor trip signal would be generated. With full load rejection l: l capability the plant would be expected to continue operating without a reactor I~ trip. A continued steam load of approximately 5% would exist after total loss of external electrical load because of the electrical demand of plant auxiliaries. - In the event the steam dump valves fail to open following a large loss of load, the steam generator cafety valves may lift and the' reactor may be tripped by the high pressurizer pressure signal, the high pressurizer water 4 level signal, or the overtemperature AT signal. The steam generator j shell-side pressure and reactor coolant temperatures will increase rapidly. l The pressurizer safety valves and steam generator safety valves are, however, sized to protect the R M nd steam generator against overpressure for all load losses without assumi;; ise operation of the steam dump system, pressurizer O v tazov:1o/o412es 15.2-37 L L _
, -y i
spray, pressurizer power-operated relief valves, automatic RCCA control, or direct reactor trip on turbine trip. The steam generator safety valve capacity is sized to remove the steam flow at the engineered safeguards design rating (104.5% of steam flow at rated power) from the steam generator without exceeding 110% of th'e steam system design pressure. The pressurizer safety valve capacity is sized based on a complete loss of heat sink with the plant initially operating at the maximum calculated turbine load along with operation of the steam generator safety valves. The pressurizer safety valves are then able to maintain the RCS pressurn within 110% of the RCS design pressure without direct or immediate reactor trip action. A more complete discussion of overpressure protaction can be found in Reference 8. 15.2.7.2 Analysis of Effects and Consequences in this analysis, the behavior of the unit is evaluated for a complete loss of steam load from full p'ower without a direct reactor trip. This is done to show the adequacy of the pressure relieving devices and to demonstrate core protection margins. The reactor is not tripped until conditions in the RCS result in a trip. The turbine is assumed to trip without actuating all the turbine stop valve limit switches. This assumption delays reactor trip until conditions in the RCS result in a trip due to other signals. Thus, the analysis assumes a worst case transient. In addition, no eredit is taken for steam dump. Main feedwater flow is terminated at the time of turbine trip, with no credit taken for emergency feedwater (except for long-term recovery) to mitigate the consequences of the transient. The total loss of load transients are analyzed with the LOFTRAN crmputer program (see Section 15.1). The program simulates the neutron kinetics, RCS, pressurizer, pressurizar relief and safety valves, pressurizer spray, steam ; generator, and steam generator safety valves. The program computes pertinent plant variables including temperatures, pressures, and power level, j l l O i l u20<:io/o412ss 15.2-38
^ ~ '
- a. . . . _ .. . - . . . . .
Major assumptions are summarized below: 9 (1) Initial Operating Conditions
-The initial reactor power and RCS temperatures are assumed at their :,
1 maximum values consistent with the steady-state full power operation . including allowances for calibration and instrument errors. The j' initial RCS pressure is assumed at a minimum value consistent with l the steady-state full power operation including allowances for calibration and instrument errors. This results in the maximum power difference for the load loss, and the minimum margin to core protection limits at the initiation of the accident. (2) Moderator and Doppler Coefficients of Reactivity O The turbine trip is analyzed with both maximum and minimum reactivity feedback. The maximum feedback (EOL) cases assume a large negative moderator temperature coefficient and the most negative Doppler power coefficient. The minimum feedback (BOL) i-cases assume a minimum moderator temperature coefficient and the least negative Doppler coefficient. l (3) Reactor Control From the standpoint of the maximum pressures attained, it is ' conservative to assume that the reactor is in manual control. If l the reactor were in automatic control, the control rod banks would move prior to trip and reduce the severity of the transient. f (4) Steam Release No credit is taken for the operation of the steam dump system or steam generator power-operated relief valves. The steam generator pressure rises to the safety valve setpoint where steam release through safety valves limits secondary steam pressure at the setpoint value. 1320v:1D/032988 15.2-33 ,
~
-- .. . .- ~__--_: : - -. ___
(5) Pressurizer Spray ar.d Power-operated Relief Valves Two cases for both the BOL and EOL are analyzed: (a) Full credit is taken for the effect of pressurizer spray and power-operated relief valves in reducing or limiting the coolant pressure. Safety valves are also available. (b) No credit is taken for the effect of pressurizer spray and power-operated relief valves in reducing or limiting the coolant pressure. Safety valves are operable. (6) Feedwater Flow Main feedwater flow to the steam generators is assumed to be lost at the time of turbine trip. No credit is taken for emergency feedwater flow since a stabilized plant condition will be reached before emergency feedwater initiation is normally assumed to occur;
'however, the emergency feedwater pumps would be expected to start on a trip of the main feedwater pumps. The. emergency feedwate- flow would remove core decay heat following plant stabilization.
Reactor trip is actuated by the first reactor protection system trip setpoint reached with.no credit taken for the direct reactor trip on the turbine trip. 15.2.7.3 Results The transient responses for a total loss of load from full power operation are shown for four cases; two cases for the BOL and two cases for the EOL on Figures 15.2.7-1 through 15.2.7-12. Figures 15.2.7-1, 15.2.7-2 and 15.2.7-3 show the transient responses for the total loss of steam load at BOL assuming full credit for the prassurizer spray and pressurizer power-operated relief valves. No credit is taken for the steam dump. The reactor is tripped by the overtemperature AT trip channel. The minimum DNBR is well above the limit value. The pressurizer safety valves 9 122orio/o412ss 15,2-40 l 1
.c .
^-' ' ^ ~ ^ - ' ' - - ' - ^ ^ ~ . ~ . - .
are actuated for this case and maintain system pressure below 110 percent of , the design value. Th6 steam generator safety valves open and limit the secondary steam pressure increase. - Figures 15.2.7-4, 15.2.7-5 and 15.2.7-6 show the responses for the total loss of load at EOL assuming a large (absolute value) negative moderator temperature coefficient. All other plant paramet'rs e are the same as in the l above case. The reactor is tripped by the overtemperature AT trip channel. , The DNBR increases throughout the transient and never drops below its initial value. Total loss of load was also studied assuming the plant to be initially operating at full power with no credit taken for the pressurizer spray, pressurizer power-operated relief valves, or steam dump. The reactor is tripped on the high pressurizer pressure signal. Figures 15.2.7-7, 15.2.7-8 and 15.2.7-9 show the BOL transients. The nuclear power remains at or above full power until the reactor is tripped. The DNBR generally increases throughout the transient. In this case the pressurizer Jafety valves are actuated and maintain the system pressure below 110 percent of the design value. 3 Figures 15.2.7-10, 15.2.7-11 and 15.2.7-12 show the transient at EOL with the other assumptions being the same as on Figures 15.2.7-7 through 15.2.7-9. Again, the DNBR inervases throughout the transient and the pressurizer safety valves are actuated to limit the primary pressure. Reference 8 presents additional results for a complete loss of heat sink l including loss of main.feedwater. This report.shows the overpressure t protection that is afforded by the pressurizer and steam generator safety l valves. , 15.2.7.4 Conclusions Results of the analyses, including those in Reference 8, show that the plant design is such that a total loss of external electrical load without a direct or issnediate reactor trip presents no hazard to the integrity of the RCS or the main steam system. Pressure-relieving devices incorporated in the two O iszov:1o/o412ss 15.2-41 l
. . . . - - __ . . _ l
systems are adequate to limit the maximum pressures to within the design limits. , The integrity of the core is maintained by operation of the reactor protection system; i.e., the DNBR will be maintained above the safety analysis limit values. Thus, no core safety limit will be violated. O l l l 1 l 1 l l l 9 l 122cv;1o/o22ssa 15.2-42
15.2.8 Loss of Normal Feedwater O D 15.2.8.1 Identification of Causes and Accident Description A loss of normal feedwater (from pump failures, valve malfunctions, or loss of offsite ac power) results in a reduction in capability of the secondary system to remove the heat generated in the reactor core. If the reactor were not tripped during this accident, core' damage would possibly occur from a sudden loss of heat sink. If an alternativo supply of feedwater were not supplied to the plant, residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer would occur. Significant loss of water from the RCS could conceivably lead to core damage. Since the plant is tripped well before the steam generator heat transfer capability is reduced, the primary system variables never approach a DNB condition. The following provide the necessary protection against a loss of normal feedwater: (1)' Reactor trip on low-low water level in any steam generator; (2) Reactor trip on steam flow-feedwater flow mismatch in coincidence i with low steam generator water level; i (3) Two motor-driven emergency feedwater (EFW) pumps that are started on: (a) Low-low level in any steam generator, (b) Trip of all main feedwater pumps, (c) Any safety injection signal, (d) Loss of offsite power (automatic transfer to diesel generators), (e) Manual actuation. O inov;1ofonssa 15.2-43 L _ -_ _ n- . -- _ - _ . .
(4) One turbine-driven emergency feedwater pump that is started on: 1 (s) Low-low level in any two steam generators, (b) Loss of offsite power, . (c) Manual actuation. _ The motor-driven EFW pumps are connected to vital buses and are supplied by the diesels if a loss of offsite power occurs. The turbine-driven pump utilizes steam from the secondary system and exhausts it to the atmosphere. The controls are designed to start both types of pumps within 1 minute even if a loss of all ac power occurs simultaneously with loss of normal feedwater. The EFW pumps take suction from the condensate storage tank for delivery to the steam generators. The analysis shows that following a loss of normal feedwater, the EFW system is capable of removing the stored and residual heat thus preventing either overpressurization of the RCS or loss of water from the reactor core. 15.2.8.2 Analysis of Effects and Consecuences O A detailed analysis using the LOFTRAN codel4) is performed in order to determine the plant transient following a loss of normal feedwater. The code describes the plant thermal kinetics, RCS including natural circulation, pressurizer, steam generators, and feedwater system, and computes pertinent variables, including the pressurizer pressure, pressurizer water levei, and reactor coolant average temperature. Major assumptions are: (1) Reactor trip occurs on steam generator low-low level at 23.2% of narrow range span. (2) The plant is initially operating at 102% of the NSSS design rating. O 1320v:1o/o32ssa 15.2-44
(3) Conservative core residual heat generation based on long-term operation at the initial power level preceding the trip is assumed. , The 1979 decay heat ANSI 5.1 + 2 SIGMA was used for calculation of residual decay heat levels. (4) The emergency feedwater system is actuated by the low-low steam generator water level signal. - (5) The worst single failure in the emergency feedwater system occurs (turbine-driven pump) and one motor-driven pump is assumed to be unavailable. The emergency feedwater system is assumed to supply a total of 380 gpm to two steam generators from the available motor-driven pump. (6) The pressurizer sprays and PORVs are assumed operable. This maximizes the peak transient pressurizer water volume. (7) Secondary system steam relief is achieved through the self-actuated , safety valves. Note that steam relief will, in fact, be through the O- power-operated relief valves or condenser dump valves for most cases t of loss of normal feedwater. However, for the sake of analysis ! these have been assumed unavailable. (8) The initial reactor coolant average temperature is 4.0*F higher than the nominal value to allow for uncertainty on nominal temperature. The initial pressurizer pressure uncertainty is 33 psi. 15.2.8.3 Results Figurcs 15.2.8-1 and 15.2.8-2 show plant parameters following a loss of normal feedwater. Following the reactor and turbine trip from full load, the water level in the steam generators will fall due to the reduction of steam generator void fraction and because steam flow through the safety valves continues to O is2ovno/o412:s 15.2-45
_ _ _ ~ _ . dissipate the stored and generated heat. One minute following the initiation of the low-low level trip, the motor-driven EFW pump is automatically started, - reducing the rate of water level decrease. The capacity of the motor-driven EFW pump is such that the water level in the steam generator being fed does not recede below the lowest level at which sufficient heat transfer area is available to dissipate core residual heat without water relief from the RCS relief or safety valves. From Figure 15.2.8-2 it can be seen that at no time is there water relief from the pressurizer. If the emergency feed delivered is greater than that of one motor-driven pump, the initial reactor power is less than 102% of.the NSSS design rating, or the steam generator water level in one or more steam generators is above the low-low level trip point at the time of trip, then the results for this transient will be less limiting. The calculated sequence of events for this accident is listed in Table 15.2-1. As shown in Figures 15.2.8-1 and 15.2.8-2, the plant approaches a stabilized condition following reactor trip and emergency feedwater initiation. Plant-procedures may be followed to further cool down the plant. h 15.2.8.4 Conclusions Results of the analysis show that a loss of normal feedwater does not adversely affect the core, the RCS, or the steam system since the EFW capacity is such that the reactor coolant water is not relieved from the pressurizer relief or safety valves. O u 2ov:1o/o412ss 15.2-46
15.2.9 Loss of Offsite Power to the Station Auxiliaries (Station Blackout) : 15.2.9.1 Identification of Causes and Accident Description During a complete' loss of offsite power and a turbine trip there will be loss of power to the plant auxiliaries, i.e., the reactor . coolant pumps, condensate ; pumps, etc. The events following a loss of ac power with turbine and reactor trip are described in the sequence listed below: (1) Plant vital instruments are supplied by emergency power sources. (2) As the steam system pressure rises following the trip, the steam
- system power-operated relief valves are automatically opened to the -
atmosphere. Steam dump to the condenser is assumed not to be available. If the power-operated relief valves are not available, the steam generator self-actuated safety valves may lift to dissipate the sensible heat of the fuel and coolant plus the residual heat produced in the reactor. l (3) As the no-load temperature is approached, the steam system power-operated relief valves (or the self-actuated safety valves, if 7 the power-operated relief valves are not available) are used to dissipate the residual heat and to maintain the plant at the het standby condition. (4) The emergency diesel generaters started on loss of voltage on the plant emergency buses begin to supply plant vital loads. The EFW system is started automatically as discussed in the loss of normal feedwater analysis. The steam-driven emergency feedwater pump utilizes steam from the secondary system and exhausts to the atmosphere. The two motor-driven EFW pumps are supplied by power from the diesel generators. The pumps take suction directly from the condensate storage tank for delivery te i the steam generators. 1 O 132ovn o m ssa 15.2-47
Upon the loss of power to the reactor coolant pumps, m olant flow necessary for core cooling and the removal of residual heat is maintained by natural g' circulation in the reactor coolant loops. T 15.2.9.2 Analysis of Effects and Consequences A detailed analysis using the LOFTRAN codel4) is performed in order to determine the plant transient following a station blackout The code describes the plant thermal kinetics, RCS including natural circulation, pressurizer, steam generators, and feedwater system, and computes pertinent variables, including the pressurizer pressure, pressurizer water level, and reactor coolant average temperature. Major assumptions differing from those in a loss of normal feedwater are: (1) No credit is taken for immediate response of control rod drive mechanisms caused by a loss of offsite power. (2) A heat transfer coefficient in the steam generator associated with RCS natural circulation is assumed following the reactor coolant pump coastdown. The time sequence of events for the accident is given in Table 15.2-1. The first few seconds after the loss of power to the reactor coolant pumps will closely resemble a simulation of the complete loss of flow incident (seo Section 15.3.4); i.e., core damage due to rapidly increasing core temperatures is prevented by promptly tripping the reactor. After the reactor trip, stored and residual heat must be removed to prevent damage to either the RCS or the core. The LOFTRAN code results show that the natural circulation flow available is sufficient to provide adequate core decay heat removal following reactor trip and RCP coastdown. s O 132cv:ioto414s 15.2-48
15.2.9.3 Conclusions Results of the analysis show that, for the loss of offsite power to the station auxiliaries event, all safety criteria are met. Since the DNBR remains above the safety analysis limit, the core is not adversely affected. EFW capacity is sufficient to prevent water relief th, rough the pressurizer ; relief and safety valves; this assures that the RCS is not overpressurized. , e Analysis of the natural circulation capability of the RCS demonstrates that sufficient long-term heat removal capability exists following reactor coolant pump coastdown to prevent fuel or clad damage. I l l 132ov:1oto32ssa 15.2-49 i
15.2.10 Excessive Heat Removal Due to Feedwater System Malfunctions 15.2.10.1 Identification of Causes and Accident Description Excessive feedwater additions are a means of increasing core power above full power. Such transients are attenuated by the thermal capacity of the secondary plant and of the RCS. The overpower and overtemperature protection (high neutron flux, overtemperature AT, and overpower AT trips) prevent any power increase that could lead to a DNBR that is less than the DNBR limit. An example of excessive feedwater flow would be a full opening of a feedwater control valve due to'a feedwater control system malfunction or an operator error. At power, this excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generator. With the plant at no-load conditions, the addition of cold feedwater may cause a decrease in RCS temperature and thus a reactivity insertion due to the effects of the negative moderator coefficient of reactivity. Continuous excessive feedwater addition is prevented by the steam generator high-high level trip, which closes the feedwater valves. 15.2.10.2 Ar.alysis of Effects and Consecuences The excessive heat removal due to a feedwater system malfunction transient is analyzed with the LOFTRAN code. This code simulates a multiloop system, neutron kinetics, the pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. The system is analyzed to evaluate plant behavior in the event of a feedwater system malfunction. O 1320rlo/032988 15.2-50
I Excessive feedwater addition due to a control system malfunction or operator error that allows a feedwater control valve to open fully is considered. Two t cases are analyzed as follows: (1) Accidental opening of one feedwater control valve with the reactor just critical at zero load conditions assum'ing a conservatively , large moderator density coefficient characterist.ic of end-of-life conditions. . (2) Accidental opening of one feedwater control valve with the reactor in manual control at full power. The reactivity insertion rate following a feedwater system malfunction is esiculated with the following assumptions: (1) For the feedwater control valve accident at full power, one feedwater control valve is assumed to malfunction, resulting in a f l step increase of 250% of nominal feedwater flow to one steam i l generator. O (2) For the feedwater control valve accident at zero load conditions, a feedwater valve malfunction occurs that resulte in a step increase in flow to one steam generator from zero to the nominal full load i value for one steam generator. (3) Fo'r the zero load condition, feedwater temperature is at a conservatively low value of 70'F. L (4) No credit is taken for the heat capacity of the RCS and steam generator thick metal in attenuating the resulting plant cooldown. l (5) No credit is taken for the heat capacity of the steam and water in the unaffected steam generators. l l l O istor.to/ostsee 15.2-51 l
(6) The feedwater flow resulting from a fully open control valve is terminated by the steam generator high-high icvel signal that closes all feedwater control valves, closes all feedwater bypass valves, h-trips the main feedwater pumps, and shuts the feedwater isolation valves. The steam generator high-high levei signal also produces a signal to trip the turbine. _ 15.2.10.3 Results In the case of an accidental full opening of one feedwater control valve with - the reactor at zero power and the above mentioned assumptions, the maximum reactivity insertion rate is less than the maximum reactivity insertion rate analyzed in Section 15.2.1, Uncontrolled RCCA Bank Withdrawal from a Suberitical Condition, and its analysis is, therefera, covered by that of the latter. It should be noted that if the incident occurs with the unit just critical at no-load, the reactor may be tripped by the power range high neutron flux trip (low setting) set at approximately 25% of nominal full power. The full power case (end-of-life, without control) gives the largest reactivity feedback and results in the greatest power increase. Assuming the reactor to be in the automatic control mode results in a slightly less severe g transient. The rod control system is not required to function for an excessive feedwater flow event. A turbine trip is actuated when the steam generator level reaches the high-high level setpoint. For convenience, reactor trip is assumed to be initiated upon turbine trip. However, this function is not necessary. Should turbine tr1p not initiate a reactor trip signal, reactor trip will occur on power range high neutron flux. For all cases of excessive feedwater, continuous addition of cold feedwater is prevented by closure of all feedwater control valves, closure of all feee ter bypass valves, a trip of the feedwater pumps, and closures of the feedwater isolation valves on steam generator high-high level. Transient results (see Figures 15.2.10-1 and 15.2.10-2) show the core heat flux, pressurizer pressure, T,yg, and DNBR, as well as the increase in nuclear power and loop AT associated with the increased thermal load on the reactor. Steam generator level rises until the feedwater is terminated as a result of the high-high steam generator level trip. The DNBR does not drop below the limit safety analysis DNBR value. 1320v:10/0517ss 15.2-52
i 15.2.10.4 Conclusions The reactivity insertion rate that occurs at no-load following excessive O- feedwater addition is less than the maximum value considered in the analysis of the rod withdrawal from a suberitical condition. Also, the DNBRs encountered for excessive feedwater addition at power,are well above the safety analysis limit DNBR value. 1 t i lO . 1 i t l I I i f O 122orio/os2ssa 15.2-53
l 15.2.11 Excessive Load Increase Incident 15.2.11.1 Identification of Cause and Accident Description An excessive load increase incident is defined as a rapid increase in the steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The reactor control system is designed to accommodate a 10% step-load increase or a 5% per minute ramp load increase in the range of 15 to 100% of full pcwer. Any loading rate in excess of these values may cause a reactor trip actuated by the reactor protection system. This accident could result from either an administrative violation such as excessive loading by the operator or an equipment malfunction in the steam dump control or turbine speed control. During power operation, steam dump to the condenser is controlled by reactor coolant condition signals; i.e., high reactor coolant temperature indicates a need for steam dump. A single controller malfunction does not cause steam dump; an interlock is provided that blocks the opening of the valves unless a large turbine lead. decrease or a turbine trip has occurred. Protection against an excessive load increase accident is provided by the following reactor protection system signals: (1) Overpower AT, (2) Overtemperature AT, (3) Power range high neutron flux. 15.2.11.2 Analysis of Effects and Consequences This accident is analyzed using the LOFTRAN code I4) . The code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, presarizer spray, feedwater system, steam generator, and steam generator safety va ves. The code computes pertinent plant variables including l temperatures, pressures, and power level. O 1320v:io/o4csas 15.2-54
~ l Four cases are analyzed to demonstrate the plant behavior following a 10% step i q y
'oad increase from rated lead. These cases are as follows:
o (1) Reactor control in manual with BOL minimum moderator reactivity j feedback, , l-1 (2) Ra' actor control in manual ivith EOL maximum moderator reactivity feedback, (3) Reactor control in automatic with BOL minimum moderator reactivity feedback,
~
(4) Reactor control in automatic with EOL maximum moderator reactivity feedback. For the BOL minimum moderator feedback cases, the core has the least negative moderator temperature coefficient of reactivity and the least negative Doppler only power coefficient curve; therefore the least inherent transient response ' capability'. For the EOL maximum moderator feedback cases, the moderator temperature coefficient of reactivity has its highest absolute value and the , most negative Doppler only power coefficient curve. This results in the largest amount of reactivity feedback due to changes in coolant temperature. A conservative limit on the turbine valve opening is assumed, and all cases are studied without credit being taken for pressurizer heaters. This accident is analyzed with the Improved Thermal Design Procedure as described in Reference 5. Initial reactor power, RCS pressure and temperature are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5. Plant characteristics and initial conditions are further discussed in Section 15.1. O 132ov:to/os2ssa 15.2-55
--^
Normal reactor control systems and engineered safety systems are not required to function. The reactor protection system is assumed to be operable; however, reactor trip is not encountered for most cases due to the error g, allowances assumed in the setpoints. No single active failure will prevent the reactor protection system from performing its intended function. Th'e cases which assume automatic red control are analyzed to ensure that the worst case is presented. The automatic function is not required. 15.2.11.3 Results The calculated sequence of events for the excessive load increase incident are shown on Table 15.2-1. Figures 15.2.11-1 through 15.2.11-4 illustrate the transient with the reactor in the manual control mode. As expected, for the BOL minimum moderator feedback case, there is a slight power increase, and the average core temperature shows a large decrease. This results in a DNBR which increases abcve its initial value. For the EOL maximum moderator feedback manually controlled case, there ia a much larger increase in reactor pcwer due co the moderator feedback. A reduction in DNBR is experienced but DNBR remains above the limit value. Figures 15.2.11-5 through 15.2.11-8 illustrate the transient assuming the reactor is in the automatic control mode. Both the BOL minimum and EOL
, maximum moderator feedback cases show that core power increases, thereby reducing the rate of decrease in coolant average temperature and pressurizer pressure. For both of these cases, the minimum DNBR remains above the limit value.
For all cases, the plant rapidly reaches a stabilized condition at the higher powar level. Normal plant operating procedures would then be followed to reduce power. O 132aa o/o32ses 15.2-56
The excessive load ir. crease incident is an overpower transient for which the fuel temperatures will rise. Reactor trip does not occur for any of the cases analyzed, and the plant reaches a new equilibrium condition at a higher power , level corresponding to the increase in steam flow. Since DNB does not occur at any time during the excessive load increase transients, the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient, l 15.2.11.4 Conclusions j-The analysis presented above shows that for a 10% step load increase, the DNBR remains above the safety analysis limit value, thereby precluding fuel or clad damage. The plant reaches a stabilized condition rapidly, following the load increase, l' O l i 1 iO iszov:io/osassa 15.2-57 l
15.2.12 Accidental Depressurization of the Reactor Coolant System 15.2.12.1 Identification of Causes and Accident Description An accidental depressurization of the Reactor Coolant System could occur as a result of an inadvertent opening of a pressurizer relief or safety valve. Since a safety valve is sized to relieve approximately twice the steam flowrate of a relief valve, and will therefore allow a much more rapid depressurization upon opening, the most severe core conditions resulting from an accidental depressurization of the RCS are associated with an inadvertent opening of a pressurizer safety valve. Initially, the event results in a rapidly decreasing RCS pressure until this pressure reaches a value corresponding to the hot leg saturation pressure. At that time, the pressure decrease is slowed considerably. The pressure continues to decrease, however, throughout the transient. The effect of the pressure decrease would be to decrease the neutron flux via the moderator density feedback, but the reactor control system (if in the automatic mode) functions to maintain the power and average coolant temperature essentially constant throughout the initial stage of the transient. Pressurizer level increases initially due to expansion caused by depressurization and then decreases following reactor trip. The reactor will be tripped by the following reactor protection system signals: (1) Pressurizer low pressure, (2) Overtemperature AT. 15.2.12.2 Analysis of Effects and Consecuences The accidental depressurization transient is analyzed with the LOFTRAN code (4) The code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. This accident is analyzed with the Improved Thermal Design Procedure as described in Reference 5. 9 132crio/o32ssa 15.2-58
In calculating the DNBR the following conservative assumptions are made: (b (1) Plant characteristics and initial conditions are discussed in Section 15.1. Uncertainties and initial conditions are included in the limit DNBR as described in Reference 5. (2) A positive moderator temperature coefficient of-reactivity for BOL operation in order to provide a conservatively high amount of positive reactivity feedback due to changes in moderator temperature. The spatial effect of voids due to local or subcooled boiling is not considered in the analysis with respect to reactivity feedback or core power shape. These voids would tend to flatten the core power distribution. (3) A low (absolute value) Doppler coefficient of reactivity such that the resultant amount of negative feedback is conservatively low in order to maximize any power increase due to moderator reactivity feedback. 15.2.12.3 Results , l Figure 15.2.12-1 illustrates the nuclear power transient foll_owing the RCS ! depressurization accident. The flux increases until the time reactor trip occurs on overtemperature AT, thus resulting in a rapid decrease in the . t nuclear flux. The time of reactor trip is shown in Table 15.2-1. The pressure decay transient following the accident is given on Figure 15.2.12-2. I l The resulting DNBR.never goes below the safety analysis limit value as shown l on Figure 15.2.12-3. l' 15.2.12.4 Conclusions The pressurizer low pressure and the overtemperature AT reactor protection system signals provide adequate protection against this accident, and the minimum DNBR remains in excess of the safety analysis limit value. O 1 2ov:10/o412ss 15.2-59 i
15.2.13 Accidental Depressurization of the Main Steam System 1 2.13.1 Identification of Causes and Accident Description O. The most severe : ore conditions resulting from an accidental depressurization of the main steam system are associated with an inadv.ertent opening of a single steam dump, relief, or safety valve. The analyses, assuming a rupture of a main steam pipe, are discussed in Section 15.4. - The steam released as a consequence of this accident results in an initial increase in steam flow that decreases during the accident as the steam pressure falls. The energy removal from the RCS causes a reduction of coolant temperature and pressure. In the presence of a neg:.tive moderator temperature coefficient, the cooldown results in a reduction of core shutdown margin. The analysis is performed to demonstrate that the following criterion is satisfied: assuming a stuck RCCA and a single failure in the engineered safety features (ESF) the limit DNBR value will be met after reactor trip for a steam release equivalent to the spurious opening, with failure to close, of the largest of any single steam dump, relief, or safety valve. g The following systems provide the necessary mitigation of an accidental depressurization of the main steam system. (1) Safety injection system (SIS) actuation from any of the following: (a) Two-out-of-three low pressurizer pressure signals, (b) High differential pressure signals between steam lines, (c) Two-out-of-three high-1 containment pressure signals. (2) The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal. O
$ 320a ceo32ssa 15.2-F0
l (3) Redundant isolation'of.the main feedwater lines: sustained high feedwater flo>. would cause additional cooldown. Therefore, a safety injection signal will rapidly close all feedwater control valves, trip the main feedwater pumps, and close the feedwater isolation valves. 15.2.13.2 Analysis of Effects and Consecuences The following analyses of a secondary system steam release are performed: (1) A full plant digital simulation using LOFTRANU) to determine RCS temperature and pressure during cooldown. (2) An analysis to ascertain that the reactor does not exceed the limit DNBR value. The following conditions are assumed to exist at the time of a secondary system break accident. (1) EOL shutdown margin at no-load, equilibrium xenon conditions, and with the most reactive assembly stuck in its fully withdrawn position. Operation of RCCA banks during core burnup is restricted in such a way that addition of positive reactivity in a secondary i system break accident will not lead to a more adverse condition than the case analyzed. (2) A negative moderator coefficient corresponding to tha EOL redded core with the most reactive RCCA in the fully withdrawn position. The variation of the coefficient with temperature -and pressure is l included. The k,ff versus te Terature curve at 1150 psia corresponding to the negative moderator temperature coefficient plus the Doppler temperature effect used is shown on Figure 15.2.13-1. O u2otio/csus 15.2-61
(3) Minimum capability for injection of high concentration beric acid solution corresponding to the most restrictive single failure in the g. safety injection system. The injection curve is shown on Figure
~15.2.13-2. This corresponds to the flow delivered by one charging pump delivering its full contents to the cold leg header. No credit has been taken for the low concentration' boric acid that must be swept from the safety injection lines downstream of the refueling watar storage tank (RWST) isolation valves prior to the deliverj of high concentration boric acid (2300 ppm) to the reactor coolant loops.
(4) The case studied is an initial total steam flow of 255 lb/se: at 1100 psia from one steam generator with offsite power available. This is the maximum capacity of any single steam dump or safety valve. Initial hot shutdown conditions at time zero are assumed since this represents the most pessimistic initial condition. Should the reactor be just critical or operating at power at the time of a steam release, the reactor will be tripped by the normal overpower protection when the power level reaches a trip point. g Following a trip at power the RCS contains more stored energy than at no-load, the average coolant temperature is higher than' at no-load, and there is appreciable energy stored in the fuel. Thus, the additional energy stored is removed via the cooldown caused by the steam line break before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored ener;y has been removed, the cooldown and reactivity intertions proceed then in the same manner as in the analysis which assumes no-load condition at time zero. However, since the initial steam generator water inventory is greatest at no-load, the magnitude and duration of the RCS cooldown are less for steam line breaks occurring at power. (5) In computing the steam flow, the Hoody Curve for fL/D = 0 is used. O 1 2ovio/o32ss: 15.2-62
_ .. = -- _ 9 (6) Perfect moisture separation in the steam generator is assumed. 15.2.13.3. Results j The results presented are a conservative indication of the events that would , occur assuming a secondary system steam release since.it is postulated that < all of the conditions described above occur simultaneously. Figures 15.2.13-3 and 15.2.13-4 show the transient arising as the result of a steam release having an initial steam flow of 255 lb/see at 1100 psia with 7 steam release from one safety valve. The assumed steam release is the maximum
; capacity of any single steam dump or safety valve. In this case, safety L injection is initiated automatically by low pressurizer pressure. Operation of one centrifugal charging pump is considered, Boron solution at 2300' ppm entera the RCS providing sufficient negative reactivity to prevent core damage. The reactivity transient for the case shown on Figure 15.2.13-4 is more severe than that of a failed steam generator safety or relief valve that is terminated by steam line differantial pressure, or a failed condenser dump valve that is terminated by low pressurizer pressure and level. The transieht is quite conservative with respect to cooldown since no credit is taken for the energy stored in the system metal other than that of the fuel elements or i the energy stored in the other steam generators. Since the transient occurs over a period of about 5 minutes, the neglected stored energy is likely to have a significant effect in slowing the cooldown.
i I 15.2.13.4 Conclusient ( The analysis has shown that the criteria stated earlier in this section are satisfied. For an accidental depressurization of the main steam system, the DNB design basis is met. This case is less limiting than the rupture of a main steam pipe case presented in Section 15.4. O istor. tons 2ssa 15.2-63
l l 15.2.14 Snurious Ooeration of the Safety Iniection System at Power - 15.2.14.1 Identification of Causes and Accident Descriotion Spurious SIS operation at power could be caused by opbrator error or a false electrical actuating signal. A spurious signal in any of the following channels could cause this accident. j l (1) High containment pressure, (2) Low pres'arizer pressure, (3) High steam line differential pressure, (4) Low steam line pressure, (5) Hanual actuation. Following the actuation signal, the suction of the coolant charging pumps is diverted from the volume control tank to the refueling water storage tank (RHST). The charging pumps then force highly concentrated (2300 ppm) boric acid solution from the RHST through the header and injection line and into the cold legs of each loop. The safety injection pumps also start automatically but provide no flow w1en the reactor coolant system (RCS) is at normal pressure. The passive injection systern and the low-head system also provide no flow at normal RCS pressure. A safety injection system (SIS) signal normally results in a reactor trip followed by a turbine trip. However, it cannot be assumed that any single fault that actuates the SIS will also produce a reactor trip. Therefore, two different courses of events are considered. Case A: Trip occurs at the same time spurious injection starts. O 1320v:1D/880517 15.2-64
1 Case B: The reactor protection system produces a trip later in the f transient. ,l For Case A, the operator should determine if the spurious signal was transient or steady state in nature, i.e., an occasional occurrence or a definite fault. The operator will determine this by following approved procedures. In the transient case, the operator would stop the safety injection and bring the ! plant to the hot shutdown condition. If the SIS must be disabled for repair, l l boration should continue and the plant brought to cold shutdown. ; j I For Case B, the reactor protection system does not produce an immediate trip and the reactor experiences a negative reactivity excursion due to the injected boron causing a decrease in the reactor power. At beginning of life, the power mismatch causes a drop in T,yg and consequent coolant shrinkage, and pressurizer pressure and level drop. Load will decrease due to the effect of reduced steam pressure on load when the turbine throttle valve is fully l open. If automatic rod control is used, these effects will be lessened until I the rods have moved out of the core. The transient is eventually terminated ; j by the reactor protection system low pressure trip or by manual trip. Results at end of life are similar except that mcderator feedback effects result in a slower transient. The pressurizer pressure and level increase slowly and the coolant T,yg decreases slowly. The transient is eventually terminated by the reactor protection system high pressurizer pressure or high pressurizer level trip or by manual trip. 1 The time to trip is affected by initial operating conditions including core burnup history that affects initial boron concentration, rate of change of boron concentration, and Doppler and moderator coefficients. 1 Recovery from this incident for Case B is in the same manner as for Case A. The only difference is the lower T,yg and pressure associated with the power imbalance during this transient. The time at which reactor trip occurs is of no concern for this occurrence. At lower loads coolant contraction will be slower resulting in a longer time to trip. O V i 1320v:1o/o32ssa 15.2-65 i
15.2.14.2 Analysis of Effects and Consequences The spurious operation of the SIS system is analyzed with the LOFTRAN g' code (4) The code simulates the neutron kinetics, RCS, pressurizer,
. W pressurizer relief and safety valves, pressurizer spray, steam generator, steam generator safety valves, and the effect of the SIS. The program
~
computes pertinent plant variables including temperatures, pressures, and power level. - Because of the power and temperature reduction during the transient, operating conditions do not approach the core limits. Analyses of several cases show that the results are relatively independent of time to trip. A typical transient is considered represer. ting conditions at BOL. This accident is analyzed with the Improved Thermal Design Procedure as described in Reference 5. The assumptions made in the analysis are: (1) Initial Operating Conditions The initiai reactor power, pressure and RCS temperatures are assumed h? to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5. (2) Moderator and Doppler Coefficients of Reactivity A positive BOL moderator temperature coefficient was used. A low absolute value Doppler power coefficient was assumed. (3) Reactor Control The reactor was assumed to be in manual control. (4) Pressurizer Heaters Pressurizer heaters were assumed to be inoperative in order to increase the rate of pressure drop. O 132ov:1o/o32ssa 15.2-66
. . . _ ~ ..
I (5) Boron Injection At time zero, two charging pumps inject 2300 ppm borated water into the cold legs of each loop. (6) Turbine Load Turbine load was assumed constant until the governor drives the throttle valve wide open. Then turbine load drops as steam pressure drops. (7) Reactor Trip Reactor trip was initiated by low pressurizer pressure. The trip was conservatively assumed to be delayed until the pressure reached 1775 paia. 15.2;14.3 Results The transient response for the minimum feedback case is shown on Figures 15.2.14-1 through 15.2.14-2. Nuclear power starts decreasing immediately due to baron injection,-but steam flow does not decrease until 25 seconds into the transient when the turbine throttle valve goes wide open. Tha mismatch between load and nuclear power causes T,yg, pressurizer water level, and pressurizer pressure to. drop. The low pressure trip setpoint is reached at 54 seconds and rods start moving into the core at 56 seconds. 3
*)
15.2.14.4 Conclusions Results of the analysis show that spurious safety injection with or without immediate reactor trip presents no hazard to the integrity of the RCS. DNBR is never less than the initial value. Thus, there will be no cladding damage and no release of fission products to the reactor coolant system. If the reactor does not trip immediately, the low pressure reactor trip will be actuated. This trips the turbine and prevents excess conidown thereby expediting recovery from the incident. O ta2ov:1o/osassa 15.2-67
= = . . _ _ - - _ - - . = . = = = . . . ==- _ ::- - . _ - - _ _ _ . . _ . . _ ,
l i t [ 15.2.15 References
- 1. W. C. Gangloff, @ Evaluation of Anticipated Operational Transients in h.
- Westinghouse Pressurized Water Reactors, WCAP-7486, May 1971.
i
- 2. Risher, D. H. Jr. and Barry, R. F., "TWINKLE-A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary) and WCAP-8028-A (Non proprietary), January 1975. j i
- 3. H. G. Hargrove, _F_ACTRAN - A Fortran IV Code for Thermal Transients in A
( UO2 Fuel Red, WCAP-7908, June 1972.
- 4. T. W. T. Burnett, et al., LOFTRAN Code Description, WCAP-7907-P-A (Proprietary), WCAP-7907-A (Non proprietary), April 1984.
- 5. Chelemer, H., et al., "Improved Thermal Design Procedure," WCAP-8567 (Proprietary) and WCAP-8568 (Non proprietary), July 1975.
l 6' , Technical Specifications, V. C. Sumer Nuclear Station Appendix A to l License No. NPF-12, as amended through Amendment Number 66.
- 7. Chelemer, H. et al, Subchannel lhermal Analysis of Rod Bundle Cores, WCAP-7015. Rev. 1, January 1969.
l
- 8. M. A. Mangan, Overoressure Protection for Westinghouse Pressurized Water f Reactor, WCAP-7769, October 1971.
- 9. J. S. Shefcheck, Application of the THINC Program to PWR Design, WCAP-7359-L, August 1969 (Proprietary) and WCAP-7838, January 1972.
l l 10. Morita, T., et. al., "Dropped Rod Methodology for Negative Flux Rate Trip Plant," WCAP-10297-P-A (Proprietary) and WCAP-10298-A (Non proprietary), June 1983. O inov:10/o40ssa 15.2-68
~ ~ ~ '
. .: ~ ~ '
~ ~~ - ' ~ - - ^ ^^ ~ -
. . . -. .. - -. ~ .
TABLE 15.2-1 Sheet 1 of 14
, y\
TIME SEQUENCE OF EVENTS FOR CONDITION 11 EVENTS 4 Accident Event _ Time, sec
, Uncontrolled RCCA Initiation of uncontrolled
-4 Withdrawal from a rod withdrawal 9.0 x 10 Suberitical Ak/sec reactivity insertion Condition rate from 10'9 of nominal power 0.0 Power range high neutron flux low setpoint reached 8.8 Peak nuclear power occurs 8.9 t.
Rods begin to fall into core 9.3 Peak heat flux occurs 11.8 Peak hot spot average clad temperature occurs 11.8 L 1 Peak hot spot average fue' temperature occurs 12.1-l l A U 1 132ov:10/o40588 l l-
=.:-:-.__.. ---- r- -
l l l TABLE 15.2-1 Sheet 2 of 14 l el Accident Event Time, sec l l Uncontrolled RCCA Withdrawal at Power
- 1. Case A Initiation of uncontrolled RCCA withdrawal at a high reactivity insertion rate (7.5 x 10'4 ak/sec) 0.0 Power range high neutron flux high trip setpoint reached 1.5 Rods begin to fall into 2.0 core Minimum DNBR occurs 2.8
- 2. Case B Initiation of uncontrolled RCCA withdrawl at a small reactivity insertion rate
-5 (5.0 x 10 ak/sec)
Overtemperature AT reactor trip signal initiated 19.6 Rods begin to fall into core 20.1 Minimum DNBR occurs 20.7 e 1320v:1D/032988
~ ~ '^~ ~ '~
j I i i TABLE 15.2-1 Sheet 3 of.14 Accident Event , Time, see Uncontrolled Baron _ Dilution
- 1. Dilution during refueling Dilution begins 0 Operator receives high flux at shutdown alarm, set at twice background 1791 Operator irolates source of dilution; minimum margin to criticality occurs 4680 0 2. Dilution during.
cold shutdown Dilution begins 0 l l
- Operator receives high flux at shutdown alarm set at twice background 1358 Operator isolates source of f.
dilution; shutdown margin is lost 2181 ;
- 3. Dilution during hot standby Dilution begins 0 Operator receives high flux at shutdown alarm, set at twice background 1405 O
1320v:10/032988
\
TABLE 15.2-1 Sheet 4 of 14 1 9 Accident Event Time, see Operator isolates source of dilution; shutdown margin is-lost 2273
- 4. Dilution during startup Power Range-low setpoint Reactor trip due to dilution 0 Shutdown margin lost (if dilution continues after trip) 1236
- 5. Dilution during full power operation
- a. Automatic reactor Operator receives lo-lo rod insertion control limit alarm due to dilution . O Shutdown margin lost 1380
- b. Manual reactor Overtemperaturo AT reactor control trip due to dilution 0 Shutdown margin lost (if dilution continues after trip) 1140 0
1320v:1D/032988
- . . . . . . . ... -. .... . ~ ..
9 TAB L E '.'55. 2-1 Sheet 5 of 14 O : Accident Event Time, sec !
~
j Partial Loss of Forced l Reactor Coolant Flow
, All loops operating, one pump coasting down Coastdown begins 0.0 Low-flow reactor trip 1.49 Rods begin to drop 2.49 3.4
~ Minimum DNBR occurs O i i i 1 l i O 1320v:f o/032988
, . = ... . _. - -.. _
TABLE 15.2-1 Sheet 6 of 14 Accident Event Time, see Startup of an Inactive Reactor Coolant Loop Initiation of pump startup - 0.0 Power reaches high nuclear flux trip 3.1 Rods begin to drop 3.6 Minimum DNBR occurs 4.2 O O l l 9 l 1320v:1D/032988 l 1 l i
TABLE 15.2-1 Sheet 7 of 14 0 ; Accident Event Time, see Loss of External _ 4 Electrical Load l 1. (
- 1. With pressurizer i control (BOL) Loss of electrical load 0.0 !
Overtemperature AT 6.6 8.1 : Rods begin to drop Minimum DNBR occurs 9.0 Initiation of steam O rei <re t rator safety valves a"- 10.0 Peak pressurizer pressure 10.6 l occurs
- 2. With pressurizer control (E0L). Loss of electrical load 0.0 Overtemperature AT trip setpoint 6.9 reached Rods begin to drop 8.4 L
Peak pressurizer pressure occurs 8.2
- O 1320v:1o/032988 r
1 l l TABLE 15.2-1 Sheet 8 of 14 O1i Accident Event Time, see Initiation of steam release l from steam generator
~
l safety valves 9.8 l Minimum ONBR occurs (a)
- 3. Without l
pressurizer control (BOL) Loss of electrical load 0.0 High pressurizer pressure reactor trip setpoint reached 5.2 Rods begin to drop 7.2 Peak pressurizer pressure occurs 8.8 Initiation of steam release from steam generator safety valves 10.2 Minimum DNBR occurs (a) O 1320v:10/032988
~
. ~ . . . _ . . .
l t l TABLE 15.2-1 Sheet 9 of 14 0 [. Accident Event Time, sec l l 4. Without
~
pressurizer > control (EOL) Loss of electrical load 0.0 High pressurizer pressure reactor trip setpoint reached 5.4 1
- Rods begin to drop 7.4 l
I Peak pressurizer pressure occurs 8.2 Initiation of steam release
' fro.m steam generator 10.0 I safety valves Minimum DNBR occurs .
(a) O 1320v:1o/032988 l i
- + 1
TABLE 15.2-1 Sheet 10 of 14 9. Accident Event Time, see Loss of Normal W/ Power W/0 Power Feedwater and Loss - of Offsite Power to the Station Auxiliaries (StationBlackout) Main feedwater flow stops 10 10 Low-low steam generator water level reactor trip 56.3 56.3 . Rods begin to drop 58.3 58.3 Reactor coolant pumps begin to coast down - 60.3 Peak water level in pres-surizer occurs 62 62 Two steam generators begin to receive emergency feed-water from one motor driven emergency feedwater pump 116.3 116.3 Cold emerge'ncy feedwater is delivered to the steam generators 128 128 Core decay heat plus pump j heat decreases to emergency feedwater heat removal I capacity ~3600 ~1200 i O 1320v:10/032988
ein F TABLE 15.2-1 Sheet 11 of 14 0 - Accident Event Time, see Excessive feedwater _ Flow at Full Load One main feedwater control valve fails fully open 0 High-high steam generator level signal generated 27.6 Turbine trip occurs duu to high-high steam generator level 29.6 Minimum DNBR occurs 30.0 Reactor trip due to turbine trip (b) 31.6 Feedwater isolation valves fully closed 40.6 l l l O 1320v:1D/o42688 l
i l TABLE 15.2-1 Sheet 12 of 14- l
-l Accident Event Time, see Excessive Load Increase
- 1. Manual reactor control (BOL minimum moderator feedback) 10% step lead increase 0.0 Equilibrium conditions reached (approximate timesonly) 300
- 2. Manual reactor control (EOL maximum moderator feedback) 10% step load increase 0.0 Equilibrium conditions reached (approximate timesonly) 100
- 3. Automatic reactor l control (BOL minimum moderator feedback) 10% step load increase 0.0 j l
Equilibrium conditions ! reached (approximate times only) 200 : Ol l 1320v:10/032988 ;
. . . . . ~ .
l TABLE 15.2-1 Sheet 13 of 14 Accident Event Time, sec
- 4. Automatic reactor _
control (EOL , maximum moderator feedback) 10% step load increase 0.0
~
Equilibrium conditions reached (approximate times only) 100 Accidental Depressuri-zation of the Reactor Coolant System Inadvertent opening of one RCS safety valve 0.0 O Overtemperature AT Trip Setpoint Reached 22.8 Rods begin to drop 24.3 Minimum DNBR occurs 24.8 Accidental Depressuri-zation of the - Main Steam System Inadvertent opening of one main steam safety ' or r'elief valve 0.0 Pressurizer empties 193 Boron from the RWST reaches RCS loops 262 1320v:10/032988 \ . . - - .. . .
TABLE 15.2-) Sheet 14 of 14 O1 Accident Event Time, sec l Inadvertent Operation of ECCS During Power Operation Charging pumps begin injecting borated water 0.0 l Low pressure trip setpoint reached 54 Rods begin to drop 56 (a) DNBR does not decrease below its initial value. (b) Not a required safety function. g I O 1320v:10/042688
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TIME (SEC) V. C. Sumer Figure 15.2.1-1 Uncontrolled Rod Withdrawal From A Suberitical Condition O Nuclear Power and Core Heat Flux vs. Time l
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> Spray and PORVs at BOL l
Core Tavg and Steam Temperature vs. Time 1
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TIME ISEC1 V. C. Summer Figure 15.2.12-2 Accidental Depressurization of the Reactor Coolant System & W Pressurizer Pressure and , Water Volume vs. Time l I 1 l l
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I O l ZERO POWER. 2150 PSIA END OF LIFE R000ED CORE WITH ONE RCCA STUCK j FULL QUT > 137 - l.N - 2 12 - 8 s u 134 - E 8 [ 133 - 5 5 2; IJ2 - s lJ1 - l.M - 2-W 1- i i i i i M M M M CORE AVERAGE TEMPERATURE (*F) V. C. Sumer Figure 15.2.13-1 Main Steam Depressurization Variation of Keff with g Core Temperature
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FROM ONE CENTRZFUGAL CHARGING PUMP V. C. Summer Figure 15.2.13-2 Main Steam Depressurization Safety Injection Flowrate I
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- To 255 lb/sec at 1100 psia w/ Offsite Power Available
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i l 15.3 CONDITION III - INFRE0 VENT FAULTS
~'N (O
By definition, Condition III occurrences are faults which may occur very infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods althnugh sufficient fuel damage might occur to preclude resumption of W operation for a considerable outage time. The release of radioactivity will not be sufficient to interrupt or restrict public use of those areas beyond the exclusion radius. A Condition .III fault will not, by itself, generate s Condition IV fault or result in a consequential loss of function of the raaetor coolant system or containment Larriers. For the purpose of this report the following faults have been gro; ped into this category: l 1. Loss of reactor coolant, from small ruptured pipes er from cracks in large pipes, which actuates the emergency core cooling system, i 1
- 2. Minor secondary system pipe breaks.
- 3. Inadvertent loading of fuel assembly into an improper position.
f]
- 4. Complete loss of forced reactor coolant flow.
- 5. Singis rod cluster control assembly withdrawal at full power.
1 Each of these infrequent faults are analyzed in this section. In ganeral, i each analysis includes an identification of causes and description of the accident, an analysis of effects and consequences, a presentation of results, ) and relevant conclusions. l The time sequence of events during applicable Condition III faults 1 and 4 above is shown in Table 15.3-1. l 1243c1c/041288 15.3-1 l
15.3.2 Minor Secondary System Pice Breaks 9 15.3.2.1 Identification of Causos and Accident Description Included in this grouping are ruptures of secondary system lines which would result in steam release rates equivalent to a 6-inch-diameter break or smaller. 15.3.2.2 Analysis of Effects and Consequences Minor secondary system pipe breaks must be accomodated with the failure of only a small fraction of the fuel elements in the reactor. Since the results of analysis presented in Section 15.4.2 for a major secondary system pipe rupture also meet these criteria, separate analyses for minor secondary system pipe breaks is not required. The analyses of the more probable accidental opening of a secondary system steam dump, relief, or safety valve is presented in Section 15.2.13. These aralyses are illustrative of a pipe break equivalent in size to a single valve cpening. 15.3.2.3 Conclusions The analysis presented in Section 15.4.2 ~ demonstrates that the consequences of h a minor secondary system pipe break are acceptable since a departure from nucleate boiling ratio (DNBR) of less than the design basis values does not occur even for a more critical major secondary system pipe break. l l 1 l 1 9 1319v lo/041188 15.3-2
- _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ ~ _
l I 15.3.3 Inadvertent Loadino of a Fuel Assembly Into an Improper Position O 15.3.3.1 Identification of Causes and Accident Description Fuel and core loading errors such as can arise from the inadvertent loading of one or more fuel assemblies into improper positions,, loading a fuel rod during manufacture with one or more pellets of the wrong. enrichment will lead to , increased heat fluxes if the error results in placing fuel-in core positions calling for fuel of lesser enrichment. Also included among possible core loading errors is the inadvertent loading of one or more fuel assemblies requiring burnable poison rods into a new core without burnable poison rods. Any error in enrichment, beyond the normal manufacturing tolerances, can cause power shapes which are more peaked than those calculated with the correct enrichments. There is a 5 percent uncertainty margin included in the design value of power peaking factor assumed in the analysis of Condition I and Condition II transients. The incere system of movable flux detectors which is used to verify power shapes at the start of life is capable of revealing any assembly enrichment e ror or loading error which causes power shapes to be peaked in excess of the design value. To reduce the probability of core loading errors, each fuel assembly is marked with an identification number and loaded in accordance with a core loading diagram. During core loading, the identification number will be checked j before each assembly is moved into the core. Serial numbers read during fuel movement are subsequently recorded on the loading diagram as a further check on proper placing after the loading is completed. i The power distortion due to any combination of misplaced fuel assemblies would significantly raise peaking factors and would be readily observable with ; incore flux monitors. In addition to the flux monitors, thermocouples are located at the outlet of about one third of the fuel assemblies in the core. there is a high probability that these thermocouples would also indicate any abnormally high coolant enthalpy rise. Incore flux measurements are taken during the startup subsequent to every refueling operation. O V 1319v:1o/04198: 15.3-3
15.3.3.2 Analysis of Effects and Consequences Steady-state power distribution in the x y plane of the core are calculated using the TllRTLE(4) Code based on macroscopic cross section calculated by the LEOPARDU) Code. A discrete representation is used wherein each individual fuel rod is described by a inesh interval. The power distribution in the x y plane for a correctly loaded core assembly are also given in Chapter 4 of the FSAR based on enrichments given in that section. For each core loading error case analyzed, the percent deviations from detector readings for a normally loaded core are shown at all incere detector locations (see Figures 15.3.3-1 to 15.3.3-5 inclusive). 15.3.3.3 Results The following core loading error cases have been analyzed:
- 1. Case A Case in which a Region 1 assembly is interchanged with a Region 3 assembly. The particular case considered was the interchange of two adjacent assemblies near the periphery of the core (see Figure 15.3.3-1).
- 2. Case B Case in which a Region 1 assembly is interchanged with a neighboring Region 2 fuel assembly. Two analyses have been performed for this case (see Figures 15.3.3-2 and 15.3.3-3).
In Case B-1, the interchange is assumed to take place with the burnable poison rods transferred with the Region 2 assembly mistakenly loaded into Region 1. l In Case B-2, the interchange is assumed to take place closer to core center and with burnable poison rods located in the correct Region 2 , position but in a Region 1 assembly mistakenly loaded into the Region 2 { position. 1 1319v:10/041288 15.3-4
~- _ _ _- ______ - _ - ___ _ ____ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ -
. 3. - Case C O Enrichment error: Case in which a Region 2 fuel assembly is loaded in the core central position (see Figure 15.3.3-4).
- 4. Case D Case in which a Region 2 fuel assembly instead of a Region 1 assembly is loaded near the core periphery (see Figure 15.3.3-5).
15.3.3.4 Conclusions Fuel assembly enrichment errors would be prevented by administrative ll procedures implemented in fabrication. j i in the event that a single pin or pellet has a higher enrichment than the nominal value, the consequences in terms of reduced DNBR and increased fuel and clad temperatures will be limited to the incorrectly loaded pin or pins. O Fuel assembly loading errors are prevented by administrative procedures implemented during core loading. In the unlikely event that a loading error occurs, analyses in this section confirm that resulting power distribution effects will either be readily detected by the incere movable detector system or will cause a sufficiently small perturbation to be acceptable within the uncertainties allowed between nominal and design power shapes. O isisv:1o/o411:s 15.3-5
l I l l 15.3.4 Comolete loss of Forced Reactor Coolant Flow .! 15.3.4.1 Identification of Causes and Accident Description A complete loss of forced reactor coolant flow may res, ult from a simultaneous loss of electrical supplies to all reactor coolant. pumps. If the reactor is at power at the time of the accident, the immediate effect of a loss of forced coolant flow is a rapid increase in the coolant temperature. This increase could result in departure from nucleate boiling (DNB) with subsequent fuel damage if the reactor were not tripped promptly. The following provide necessary protection against a loss of coolant flow accident: (1) Undervoltage or underfrequency on reactor coolant pump power supply buses; (2) Low reactor coolant locp flow. . The reactor trip on reactor coolant pump bus undervoltage is provided to protect against conditions that can cause a loss of voltage to all reactor coolant pumps, i.e., station blackout. The reactor trip on reactor coolant pump underfrequency is provided to open the reactor coolant pump breakers and h trip the reactor for an t.nderfrequency condition, resulting from frequency disturbances on the major power grid. The trip disengages the reactor coolant pumps from the power grid so that the pumps' flywheel kinetic energy is available for full coastdown. Both trips are blocked below approximately 10% power (Permissive 7). The reactor trip on low primary coolant loop flow is provided to protect against loss-of-flow conditions that affect only one reactor coolant loop. It also serves as a backup to the undervoltage and underfrequency trips. This function is generated by two-out-of-three low-flow signals per reactor coolant loop. Above approximately 38% power (Permissive 8), low flow in any loop will actuate a reactor trip. Between approximately 10 and 38% power (Permissive 7 and Permissive 8), low-flow in any two loops will actuate a reactor trip. O 1319v:1o/051788 15.3-6
Normal power for each pump is supplied through individual busses connected to the isolated phase bus duct between the generator circuit breaker and the main transformer. Faults in the substation may cause a trip of the main transformer high side circuit breaker leaving the generator to supply power to the reactor coolant pumps. When a generator circuit breaker trip occurs because of electrical faults, th6 pumps are autom'atically transferred to an alternate power supply and the pumps will continue to supply coolant flow to the core. Following any turbine trip where there are no electrical faults, the generator circuit breaker is tripped and the reactor coolant pumps remain connected to the network through the transforcer high side breaker. Continuity of power to the pump buses is achie.ed without motoring the generator since means are provided to isolate the generator without isolating the pump buses from the external power lines (e.g., a generater output breaker is provided as well as a station output breaker). 15.3.4.2 Analysis of Effects and Consecuences This transient is analyzed by three digital computer codes. First, the LOFTRAN code (1) is used to calculate the loop and core flow during the transient. The LOFTRAN code is also used to calculate the time of reactor trip based on the calculated flows and the nuclear power transient following reactor trip. The FACTRAN code (2) is then used to calculate the heat flux transient based on the nuclear power and flow from LOFTRAN. Finally, the THINC codeI3) is used to calculate the minimum DNBR during the transient based on the heat flux from FACTRAN and flow from LOFTRAN. The transients presented represent the minimum of the typical and thimble cells for Standard and VANTAGE 5 fuel. The following case has been analyzed: All loops operating, all loops coasting down. The method of analysis and the assumptions made regarding initial operating conditions and reactivity coefficients are identical to those discussed in Section 15.2, except that following the loss of supply to all pumps at power, a reactor trip is actuated by either bus undervoltage or bus underfrequency. O 1319v:1o/041188 15.3-7
- + - - - -
15.3.4.3 Results The calculated sequence of events is shown in Table 15.3-3. Figures 15.3.4-1 and 15.3.4-2 show the flow coastdown, nuclear power and heat flux transients lll) and minimum DNBR for the limiting complete loss of flow event. The reactor is assumed to trip on the undervoltage signal. The DNBR.versus time plot represents the limiting cell for the three-loop coastdown. 15.3.4.4 Conclusions The analysis performed has demonstrated that for the complete loss of forced reactor coolant flow, the DNBR does not decrease below the safety analysis limit values during the transient, and thus, no core safety limit is violated. O 9 131sv:1o/0411ss 15.3-8
. ~ . .
f p 15.3.6 Single Rod Cluster Control Assembly Withdrawal at Full Power
~
h Identification of Causes and Accident Description 15.3.6.1 No single electrical or mech.nical failure in the rod control system could cause the accidental withdrawal of a single RCCA from'the inserted bank at full power operation. The operator could deliberately wit.hdraw a single RCCA in the control bank; this feature is necessary in order-to retrieve an assembly should one be accidentally dropped. In the extremely unlikely event of simultaneous electrical failures that could result in single RCCA withdrawal, rod deviation and rod control urgent failure would both be displayed on the plant annunciator, and the rod position indicators would indicate the relative positions of the assemblies in the bank. The urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single RCCA by operator action, whether deliberate or by a combination of errors, would result in activation of the same alarm and the same visual indications. Each bank of RCCAs in the system is diviced into two groups of four mechanisms (O) each. The rods comprising a group operate in parallel through multi;:lexing thyristors., The two groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A definite j schedule of actuation and deact'uation of the stationary gripper, movable ;. gripper, and lift coils of a mechanism is required to withdraw the RCCA { attached to the mechanism. Since the four stationary grippers, movable b grippers, and lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure that would cause rod withdrawal would affect a minimum of one group, or four RCCAs. Mechanical failures are either l in the direction of insertion 'or immobility, l In the unlikely event of multiple failures that result in continuous ! withdrawal of a single RCCA, it is not possible, in all cases, to provide assurance of automatic reactor trip so that core safety limits are not violated. Withdrawal of a single RCCA results in both positive reactivity l insertion tending to increase core power, and an increase in local power density in the core area covered by the RCCA. l 1319v.10/041988 15.3-9 m.
15.3.6.2 Analysis of Effects and Consecuences Power distributions within the core are calculated by the TURTLE code based on a macroscopic cross section generated by LEOPARD. The peaking factors calculated by TURTLE are then used by THINC to calculate the minimum DNB for the event. The plant was analyzed for the case of the worst rod withdrawn from Bank C inserted at the insertion limit, with the reactor initially at full power. 15.3.6.3 Results Two cases have been considered as follows: (1) If the reactor is in the automatic control mode, withdrawal of a single RCCA will result in the immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as Case 2 described below. For such cases as above, a trip will ultimately ensue, although not sufficiently fast in all cases to prevent a minimum DNBR in the core of less than the safety limit. (2) If the reactor is in the manual control mode, continuous withdrawal of a single RCCA results in both an increase in core power and coolant temperature, and an increase in the local hot channel factor in the area of the failed RCCA. In terms of the overall system response, this case is similar to those presented in Section 15.2; however, the increased local power peaking in the area of the withdrawn RCCA results in lower minimum DNBR than for the withdrawn bank cases. Depending on initial bank insertion and location of the withdrawn RCCA, automatic reactor trip may not occur sufficiently fast to prevent the minimum core ONBR from falling below the safety limit value. Evaluation of this case at the power and coolant condition at which overtemperature 6T trip would be expected to trip the plant shows that an upper limit for the number of rods with a DNBR less than the safety limit value is 5%. O 131stio/o41ssa 15.3-10
I i 15.3.6.4 Conclusions For the case of one RCCA fully withdrawn, with the reactor in either the
,(O) automatic or manual control mode and initially operating at full power with ' Bank D at the insertion limit, an upper bound of the number of fuel rods ,
experiencing DNBR less than the design limit is 5% on less of the total fuel , rods in the core. . For both cases discussed, the indicators and alarms mentioned would function to alert the operator to the malfunction before DNB could occur. For Case 2 discussed above, the insertion limit alarms (low and low-low alarms) would also serve in this regard. O ; O isisv:to/o41ssa 15.3-11
, I
15.3.7 References
- 1. T. W. T. Burnett, et. al., LOFTRAN Code Deteription, WCAP-7907-P-A 0
(Proprietary), WCAP-7907-1 (Non-Proprietaiy), April 1984.
- 2. H. G. Hargrove, FACTRAN-A Fortrar IV,, Cede for Thermal Transients in a U0 Fuel Rod, WCAP-7908, June 1972.
2
- 3. J. S. Shefcheck, Aeolication of the THINC Program to PWR Design, WCAP-7359-L, August 1969 (proprietary), WCAP-7838, January 1972
-(Non-Proprie tary) .
- 4. R. F. Barry and S. Altomare, The TURTLE 24.0 Diffusion Deoletion Code, WCAP-7213-P-A (Proprietary), WCAP-7758-A (Non-Proprietary), January 1975.
- 5. R. '. Barry, LEOPARD-A Spectrum Decendent Non-Soatial Decletion Code for the IBM-7904, WCAP-3269-26, September 1963.
O O uiscio/c41ssa 15.3-12
TABLE 15.3-3 , O TIME SEQUENCE OF EVENTS FOR CONDITION 111 EVENTS l c Event see Accident i
^
Complete Loss of Forced Reactor Coolant Flow All loops operating, Coastdown begins 0.0 all pumps coasting Rod motion begins 1.5 down Minimum DNBR occurs 3.4 l O 1319v:1o/032988
~ ~ ~ '
O - D C 8 A R P N N L K J H G F E 1
-191
-201 2
~
-si -158 -175 -21f 3
-itt, 4
-si -ist
-221 5
+0! -55 -Il3 - 20 ' -211 g - il d -16% 6 9- -si -lis - i s ', -2if 7
-125 -19 1 - 207. 8 16! 151 95 t
-20 ?. 9 17J -55 -95 104 -lan -17 5 10 355 vit af. -5% -se il f
-9% -l F 12 71% 151 I .
> N N /
s1 it 13 '!
>Xw 225. 4% lq 15 10!
CASE A i r ' V. C. Sumer i Figure 15.3.3-1 f Inadvertent Fuel Misloading Interchange of Region 1 l and Region 3 Assembly i
- * - ^ -
4 J 0 F E D C B A R P N N L K H , 4 N'.a_ I 15 __ 2 25 23 3 05 15 15 4 Of 15 24 3 ', 3: 5
-ot 01 .5 31 2
ot 15 3 *- 6
-ot of 21 st 2' 7
- li -14 -o' -os ss 2f it a
.,s
., 2, x -34 -l'
... e 10
-11 :.s
-it - 25 -3% -si 11
- 2f. - 2' - 21 12
-l%
13
- 25 -2t It!
- 25 -2%
-11 15 I i CASE 8-1 m
V. C. Sumer Figure 15.3.3-2 Inadvertent Fuel Misloading Interchange of Region 1 and Racion 2 Assa. ably, Poison Rods Retained In Region 2 Assembly wc - , - - . , - - . , - , , - - , - -,-v. - - - -
l O 4. L J 6 C 8 A R P N N X H F E D f
=- l, 1 J 23 l.~
2r 2 l 25 3 2t 3% 35 4 21 31 3%
- 21. 41 24 2f. I #. 5 It 21 st 45 6 75 - 15 +0' 7 l +c t 75
-og -ot it jk[ ~f 2 -01 -o' 8
- 2' -11 -l' 9
-lt
-25 -l' IO
-l$ -57.
-st -st -si -35 ll ,
l'
-3t -is -Si -35 12
-55 -a f. 13
-55 14
-es
- - - =
-54 15
- CASE 8-2 V. C. Sumer ,
Figure 15.3.3-3 Inadvertent Fuel Misloading Interchange of Region I O and Region 2 Assembly, Poison Rods Transferred Te Region 1 Assembly i s . - - . - ,-
,....T.---...:. .,LT*.-,--._,--
, , , , ,,- t ,
-l 1
1 F. E D C B A A P M M L K J H G T 1 I -q% . 2
-is
- 3?.
3
- 3 ', - 3 '.
-34 4
- 21 -11 - 2?.
-4' 5 2t - 1.5 -24
- 3 *- - l '.
6 0' 73 34
- 18 -48 7 lit 625
- 3' 7% -3! -4 ' 8
-t' - 3 '- 2t
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-it 21 ot -11 11
-ti - 35 -3' 12
-35
-3 '; 13
-45 14
-e,r 15
-4t CASE C V. C. Summer Figure 15.3.3-4 Inadvertent Fuel Hisloading Enrichment Error, A Ragion 2 Assembly, Loaded Into The Core Center
-l J & F E D C 8 A It P M M L ~K H I j
-18%
j'
-19.t 2
-t '. -15% -17% -W 3 1
-ti -15% -188 4
-E -185 .in -z; 5
- 21 -lif
-M -Il' -151 6
- I I ', -171 - 19 ' 7 5' -4 *-
ll* ll' 7% -l it - i7' - 18 ' 8 Int -4t -81 - 17 < 9 O 3; 10 -il5 H51 10 a', M -M -4f II so. [ In -77 -#I 12 lit it 13 361 7% 14
- 15) 15 L., . . . - CASE D a
V. C. Sumer Figure 15.3.3-5
' Inadvertent Fuel Misloading A Region 2 Assemoly Loaded Into A Region 1 Position Near Core Periphery
. ..,.__.a. .=-:-.__-_.:- .. --
_ -l g l n.4 o
. 1. 2 - -
E g 1. - ~ n .8 3 L .6< E .4 m u
.2 m
0 0' W O. 1. 2. 5. 4 5. 6. 7. 8. 9. 10. E TIME (SEC1 1.4 E g i.2 u O I. - W .0 a E .6< i x 3 u
.4 E .2 W
~
0. O. I '. 2. 5. 4 5. 6. 7. B. 9. 10. s TIME (SECl V. C. Sumer Figure 15.3.4-1 All Loops Operating All Loops Coasting Down Vessel Flow and Heat Flux vs. Time
- - > . = - . _ . ,
l l I
,I
= i lJ g' t , 4 , 'l 2 .
i .2 ; b - 1. a h g .8 <
.6 <
o Q. 4 Cd l l 6 _a 2< l V D D
- 9. 10.
- 0. 1. 2. 5. 4 5. 6. 7. B.
TIME ISEC1 l
- 2. 6 <
t
- 2. 4 -
- 2. 2 <
i w 1 E 2. O
- 1. 8 <
l.6 . t
- 1. 4 <
l' 1.2 4 O. 1. 2. 5. 4 5. 6. 7. B. 9. 10. TIME ISECl V. C. Sumer Figure 15.3.4-2 All loops Operating All Loops Coasting Down O. Nuclear Power and ONBR vs. Time
=
M - ~ - - - w .w'--9 I R e 9 - , - , -.-,w-(w - w gw%e-yg- - + y w gg- e,.- 9 .+,9 ,g.,3y se--%--y.c.y.gwa,-+cwy- 7-, y---=-y P
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15.4 CONDITION IV - LIMITING FAULTS
!3 V Condition IV occurrences are faults that are not expected to take placa, but are postulated because their consequer:es would include the potential for the release of significant amounts of radioactive material. These are the most
~
drastic occurrences that must be designed against and represent limiting design cases. Condition IV faults shall not cause a fission product release to the environment resulting in an undue risk to public health and safety in excess of guideline values of 10 CFR 100. A single Condition IV fault shall not cause a consequential loss of required functions of systems needed to cope with the fault. including those of the emergency core cooling system (ECCS) and the containment. For the purposes of this report the following faults have been classified in this category: (1) Major rupture of pipes containing reactor coolant up to and including double-ended rupture of the largest pipe in the reactor coolantsystem(RC*1,i.e., loss-of-coolantaccident(LOCA); O (2) Major secondary system pipe ruptures; V (3) Steam generator tube rupture; (4) Single reactor coolant pump (RCP) locked rotor, , I (5) Rupture of a control rod mechanism housing (rod cluster control assembly (RCCA] ejection). , i Each of these five limiting faults is analyzed in Section 15.4. In general, each analysis includes an identification of causes and description of the accident, an analysis of effects and consequences, a prescatation of results, and relevant conclusions. 1 0 1319v:1D/042988 15.4-1
15.4.2 Major Secondary System Pioe Ruoture Two major secondary system pipe ruptures are analyzed in this section: rupture of a main steam line and rupture of a main feedwater pipe. The time sequence g of events for each of these events is pro /ided in Table 15.4-8. 15.4.2.1 Ruoture of a Main Stee7. Line 15.4.2.1.1 Identification of Causes and Accident Description The steam release arising from a rupture of a main steam pipe would result in an initial increase in steam flow that decreases curing the accident as the steam pressure falls. The energy removal from the RCS causes a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in a reduction of core shutdown margin. If the most reactive RCCA is assumed stuck in its fully. withdrawn position af ter reactor trip, there is an increased possibility that the core will become critical and return to pcwer. A return to power following a steam pipe rupture is a potential problem mainly because of the high power peaking factors that exist assuming the most reactive RCCA to be stuck in its fully withdrawn position. The core is ultimately shut down by the teric acid injection delivered by the SIS. The analysis of a main steam pipe rupture is performed to demonstrate that the following criteria are satisfied: (1) Assuming a stuck RCCA, with or without offsite power, and assuming a single failure in the engineered safety features (ESF) there is no consequential damage to the primary system and the core remains in place and intact; I ('.) Energy release to containcient from the worst steam pipe break does not l cause failure of the containment structure. l l Although DNB and possible cladding perforation following a steam pipe rupture are not necessarily unacceptablo, the following analysis, in fact, shows that no DNB occurs for any rupture assuming the most reactive assembly stuck in its fully withdrawn position. 1 O l isistierosasse 15.4-2
The following functions provide the necessary protection against a steam pipe , rupture: , (1) SIS actuation from any of the following: (a) Two-out-of-three low pressurizer pressure signals; l l' (b) High differential pressure signals between steam lines; , (c) Two-out-of-three low steam line pressure signals; i (d) Two-out-of-three high-1 containment pressure signals. (2) The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal. sustained high (3) Redundant isolation of the main feedwater lines: p feedwater flow would cause additional cooldown. Therefore; a safety V injection signal will rapidly close all feedwater control valves, trip the main'feedwater pumps, and close the' feedwater isolation valves that backup the control valves. (4) Trip of the rain steam line isolation valves on: (See Technical SpecificationsII) Table 3.3-5) (a) High steam flow in two-out-of-three main steam lines in coincidence with two-out-of-three low-low T,yg signals; (b) High-2 containment pressure signal; (c) Two-out-of-three low steam line presswa signals. ( t 1319v:1o/032968 15.4-3 e
For breaks downstream of the isolation valves, closure of all valves would completely terminate the blowdown. For any break, in any location, no more than one steam generator would blow down even if one of the isolation valves fails to close. A description of steam line isolation is included in Chapter 10 of the FSAR. . Steam flow is measured by monitoring dynamic head inside ths steam pipes. Nozzles that are of considerably smaller diameter than the main steam pipe are located in the steam generators and serve to limit the maximum steam flow for any break at any location. 15.4.P.1.2 Analysis of Effects and Consecuences The analysis of the steam pipe rupture has been pe-formed to determine: (1) The core heat flux and RCS temperature and pressure resultin from the cooldown following the steam line break. The LOFTRAN code (2 has been used. (2) The thermal and hydraulic behavior of the core following a steam line break. A detailed thermal and hydraulic digital-computer code, THINC O) , has been used to determine if DNB occurs for the core conditions computed in (1) above. The following conditions were assumed to exist at the time of a main steam line break acr.ident. (1) End of life (EOL) shutdown margin at no-load, equilibrium xenon conditions, and the most reactive assembly stuck in its fully withdrawn position. Operation of the control rod banks during core burnup is restricted in such a way that addition of positive rea:tivity in a steam line break accident will not lead to a more adverse condition than the case analyzed. 1 0 1319c10/032988 15.4-4
(2) The negative moderator coefficient corresponding to the EOL rodded core with the most reactive red in the fully withdrawn position. The variation of the coefficient with temperature and pressure has been included. The k ,ff versus temperature at 1150 psia corresponding to the negative moderator temperature coefficient plus the Doppler temperature effect used is shown on Figure- 15.2.13-1. The effect of power generation in the core on overall reactivity'is shown on Figure 15.4.2-1. The core properties associated with the sector nearest the affected steam' generator and those associated with the remaining sector were conservatively combined to obtain average core properties for reactivity feedback calculations. Further, it was conservatively assumed that the core power distribution was uniform. These two conditions cause underprediction of the reactivity feedback in the high power region near the stuck rod. To verify the conservatism of this method, the reactivity as well as the power distribution was checked. These core analyses considered the Doppler re. activity from the high fuel temperature near the stuck RCCA, moderator feedback from O. the high water enthalpy near the stuck RCCA, power redistribution and nonuniform core inlet temperature effects. For cases in which steam ;, generation occurs in the high flux regions of the core, the effect of void formation was also included. It was determined that the reactivity employed in the kinetics analysis was always larger than the true reactivity verifying conservatism; i.e., underprediction of negative reactivity feedback from power generation. l-(3) Minimum capability for injection of.high concentration boric acid (2300 ppm) solution corresponding to the most restrictive single failure in the SIS. The characteristics of the injection unit used I are shown on Figure 15.2.13-2. This corresponds to the flow delivered by one charging pump delivering its full flow to the cold leg header. No credit has been taken for the low concentration of boric acid that O V , 1siscio/osass 15.4-5 N -- :, . - . - _ _ _ . _ _ _ . . . _ . _ _ _ _ __ ____ _
must be swept from the safety injection lines downstream of the refueling water storage tank (RWST) isolation valves prior to the , delivery of highly concentrated boric acid to the reactor coolant locps. This effect has been allowed for in the analysis. The l l modeling of the SIS in LOFTRAN is described in Reference 2. For the case where offsite power is assumed, the se~quence of events in the SIS is the following: After the generation of the safety injection signal (appropriate delays for instrumentation, logic, and signal transport included), the apprepriate valves begin to operate and the high-head injection pump starts. In 27 seconds, the valves are assumed to be in their final position and the pump is a:sumed to be at full speed. The volume containing the low concentration borated water is swept before the 2300 ppm be on reaches the core. This delay is inherently included in the modeling. In cases where offsite pcwer is not available, an additional 10-second delay is assumed to be required to start the diesels and to load the necessary safety injection equipment onto them. That is, after a total of 37 seconds following an SIS signal, the SIS is assumed to be capable of delivering flow to the RCS, h (4) Two cases have been considered in determining the core power and RCS transients: (a) Complete severance of a pipe with the plant initially at no-load conditions, full reactor coolant flow with offsite power available, (b) Complete severance of a pipe with the plant initially at no-load conditions with offsite power unavailable. (5) Power peaking factors corresponding to one stuck RCCA and nonuniform core inlet coolant temperatures are determined at EOL. The coldest core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local void in the region of the stuck control assembly during the 9 1319v.1o/03:ssa 15.4-6
~
s return to power phase following the steam line break. This void in conjunction with the large negative moderator coefficient partially - offsets the effect of the stuck assembly. The power peaking factors - depend on the core power, operating history, temperature, pressure, and flow, and thus are different for each case studied. Both cases assums initial hot shutdown conditions at time zero since this represents the most pessimistic initial condition. Should the reactor be just critical or operating at power at the time of a steam line break, the reactor will be tripped by the normal overpower protection system when power level reaches a trip point. Following a trip at power the RCS contains more stored energy than at no-load, the average coolant temperature is higher than at no-load, and there is appreciable energy stored in the fuel. Thus, the additional stored energy is removed via the cooldown caused by the steam line break before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the ecoldown and reactivity insertions proceed in the same manner as in the analysis which assumes no-load condition at time zero. , However, since the initial steam generator water inventory is greatest at no-load, the magnitude and duration of the' RCS cooldown are less , j for steam line breaks occurring at power. 1 (6) In' computing the steam flow during a steam line break, the Moody Curve (4) for fL/D = 0 is used. The Moody Multiplier is 1 with a
~
discharge at dry saturated steam conditions. (7) Perfect moisture separation in the steam generator is assumed. The assumption leads to conservative results since, in fact, considerable l water would be discharged. Water carryover would reduce the magnitude of the temperature decrease in the core and the pressure increase in the containment. l meao/onses 15.4-7
15.4.2.1.3 Results The results presented are a conservative indication of the events that would occur assuming a steam line rupture since it is postulated that all of the conditions described above occur simultanecusly. Figures 15.4.2-2 and 15.4.2-3 show the response of pertinent system parameters following a main steam pipe rupture. Offsite power is atsuined to be available such that full reactor coolant flow exists. The trensient shown assumes an uncontrolled steam release frem only one staam generator. As can be soen, the core attains criticality with RCCAs inserted (with the design shutdown assuming one stuck RCCA) befora boric acid solution at 2300 ppm enters the RCS from the SIS which is drawing frcm the RWST. The delay time consista of the time to receive and actuate the safety injection signal and the time to completely open valve trains in the safety injection lines. The safety injection pumps are then ready to deliver flow. At this stage, a further delay is incurred before 2300 ppm boron solution can be injected to the RCS due to the low concentration solution being swept from the safety injectien lines. Should a partial loss of offsite power occur such that power is lost to the ESF functions while the reactor coolant pumps remain in operation, an additional safety injection delay of 10 seconds would occur while the diesel generators startup and the necessary safety injection equipment is loaded onto them. A peak core power well below the nominal full power value is attained. The calculation assumes the boric acid is mixed with and diluted by the water , flowing in the RCS prior to entering the reactor core. The concentration after mixing depends on the relative flowrates in the RCS and the SIS. The variation of mass ficwrate in the RCS due to water density changer is included in the calculation as is the variation of flowrate from the SIS and the accumulator due to changes in the RCS pressure. The SIS flow calculation includes the line losses in the system as well as the pump head curve. The accumulators provide an additional source of berated water after the RCS pressure has decreased to belew 600 psia. O isisv ioecusas 15.4-8
4 Should the core be critical at near zero power when the rupture occurs, the
- initiation of safety injection by high differential pressure between any steam line and the remaining steam lines or low steam line pressure will trip the reactor. Steam release from more than one steam generator will be prevented by automatic trip of the isolation valves in the steafn lines by low steam line ,
pressure or the high steam flow signal in coincidence with low-low RCS _ l temperature. The steam line isolation valves are designed to be fully closed ' in less than 5 seconds after receipt of closure signal. Figures 15.4.2-4 and 15.4.2-5 show the responses of the salient parameters for the case discussed above with a total loss of offsite power at the time of the rupture. This results in a coastdown of the reactor coolant pumps. In this case, the core power increases at a slower rate and reaches a lower peak value than in the cases in which offsite power is available to the reactor coolant pumps. The ability of the emptying steam generator to extract heat from the RCS is reduced by the decreased flow in the RCS, It should be noted that following a steam line break. only one steam generator blows down completaly. Thus, the remaining steam generators are still available for dist.ipation of decay heat after the initial transient is over. In case of a loss of offsite power, this heat is removed to the atmosphere via the steam line safety valves. ll 15.4.2.1.4 Conclusion . A DNS analysis was performed for the above cases. It was found that the DNS design basis (16) is met. O isterwestss: 15.4-9
15.4.2.2 Major Rupture of a Main Feedwater Pipe 15.4.2.2.1 Identification of Causes and Accident Description A major feedwater line rupture is defined as a break in a feedwater pipe large enough to prevent the addition of sufficient feedwater to the steam generators to maintain shell-side fluid inventory in the steam generators. If the break is postulated in a feedline between the check valve, the forward flush valve, or the reverse flush valve and the steam generator, fluid from the steam generator may also be discharged through the break. (A break upstream of the feedline check valve, or downstream of the forward or reverse flush valves would affect the nuclear steam supply system (NSSS) only as a loss of feedwater. This case is covered by the evaluation in Section 15.2.8.) Depending on the size of the break and the plant operating conditions at the time of the break, the break could cause either an RCS cooldown (by excessive energy discharge through the break), or an RCS heatup. The potential RCS cooldown resulting from a secondary pipe rupture is evaluated in Section 15.4.2.1, Rupture of a Main Steam Pipe. Therefore, only the RCS heatup effects are evaluated'for a feedline rupture. A feedline rupture reduces the ability to remove heat generated by the core from the RCS for the following reasons: (1) Feedwater to the steam generators is reduced. Since feedwater is l l subcooled, its loss may cause reactor coolant temperatures to increase prior to reactor trip; (2) Liquid in the steam generator may be discharged through the break, and would then not be available for decay heat removal after trip; (3) The break may be large enough to prevent the addition of any main feedwater after trip. O 121stto/o41ssa 15.4-10
~ '-^
An emergency feedwater system is provided to assure that adequate feedwater !! will be available such that: . lf i-. (1) No substantial overpressurization of the reactor coolant system shall occur; and - (2) Liquid in the reactor coolant system shall be sufficient to cover the reactor core at all times. b The following provide the necessary protection against a main feedwater line l! rupture. (1) A reactor trip on any of the following conditions: ; (a) High pressurizer pressure; (b) Overtemperature AT;
. i hr (c) Low-low steam generator water level in any steam generator; l
(d) Low steam generator level plus steam /feedwater flow mismatch in l any steam generator; (e) Safety injection signals from any of the following: L
- 1. Low steam line pressure, -
l
- 2. High containment pressure (Hi-1),
{
- 3. High steamline differeritial pressure.
(Refer to Chapter 7 for a description of the actuation system.) O i 131scio/c32sse 15.4-11
, , . - , - - . - - , , - -. ,. , . .,.- - .mr- ---- - - - . , , , , -
(2) An emergency feedwater system to provide an assured source of feedwater to the steam generators for decay heat removal. (Refer to FSAR Section 10.4.9 for a description of the emergency feedwater g system.) 15.4.2.2.2 Analysis of Effects and Consecuences A detailed analysis using the LOFTRAN W code is performed-in order to determine the plant transient following a feedline rupture. The code describes the plant thermal kinetics, RCS including natural circulation, pressurizer, steam generators, and feedwater system, and computes pertinent variables, including the pressurizer pressure, pressurizer water level, and reactor coolant average temperature. Major assumptions are: (1) The plant is initially eperating at 102% of the ESF design rating. (2) Initial reactor coolant average temperature is 4.0*F above the nominal value, and the initial pressurizer pressure is 33 psi above its nominal value. (3) A conservatively high initial pressurizer level is assumed; initial steam generator water level is at the nominal value plus 5% in the faulted steam generator, and at the nominal value minus 5% in the intact steam generators. (4) No credit is taken for the pressurizer pcwer-operated relief valves or pressurizer spr,ay. (5) No crodit is taken for the high pressurizer pressure reactor trip. Note: This assumption is a,ade for calculational convenience. Pressurizer power-operated relief valves and spray could act to delay the high pressure trip. Assumptions 3 and 4 permit evaluation of one hypothetical, limiting case rather than two 9 131 set o/o32sas 15.4-12
possible cases: one with a high pressure trip and no pressure v control; and one with pressure control but no high pressure trip. - (6) Main feed to all steam generators is assumed to stop at the time the break occurs. (All main feedwater spills out'through the break.) (7) A conservative feedline break discharge quality is assumed prior to the time the reactor trip occurs, thereby maximizing the time the trip setpoint is reached. After the trip occurs, a saturated liquid discharge is assumed until all water inventory is discharged from the affected steam generator. This minimize the heat removal capability of the affected steam generator. (8) Reactor trip is assumed to be initiated when the low-low level trip setpoint in the ruptured steam generator is reached. A low-lew level setpoint of 0% narrow range span is assumed. (9) The worst possible break area which minimizes the steam generator fluid inventory at the time of trip and is assumed maximizes the blowdown discharge rate following the time of trip, and thereby ; maximizes the resultant heatup of the reactor coolant. . t (10) No credit is taken for heat energy deposited in RCS metal during the RCS heatup. (11) No credit is taken for charging or letdown. (12) Steam generator heat transfer area is assumed to decrease as the shell-side liquid inventory decreases. (13) Conservative core residual heat generation based on long-term operation at the initial power level preceding the trip is assumed. The 1979 ANS 5.1(5) decay heat standard plus uncertainty was used for calculation of residual decay heat levels. O 1ais<.to/cs:ssa 15.4-13
(14) The emergency feedwater system is actuated by the low-low steam generator water level signal. The emergency feedwater system is assumed to suoply a total 380 gpm to the unaffected steam h. generators. A 60 second delay following reactor trip is assumed to allow time for startup of the emergency diesel generators and the emergency feed pumps. Before the relatively cold (120*F)
~
emergency feedwater enters the unaffected steam generators, additional time is modeled to allow for the purging of 5 cubic feet of hot water contained in the emergency feedwater system lines. 15.4.2.2.3 Results Results for two feedline break cases are presented. Results for a case in which offsite power is assumed to be available are presented in Section 15.4.2.2.3.1. Results for a case in which offsite power is assumed to be lost following reactor trip are presented in Sectien 15.4.2.2.3.2. The calculated sequence of events for both cases is listed in Table 15.4-8. 15.4.2.2.3.1 Feedline Ruoture with Offsite Power Available The system response following a feedwater line rupture, assuming offsite power is available, is presented in Figures 15.4.2-6 through 15.4.2-9. Results presented in Figures 15.4.2-7 and 15.4.2-9 show that pressures in the RCS and main steam system remain below 110% of the respective design pressures. Pressurizer pressure decreases after reactor trip on low-low steam generator water level due to the reduction of heat input. Following this initial decrease, pressurizer pressure increases to the pressurizer safety valve setpoint. This increase in pressure is the result of coolant expansion caused by the reduction in heat transfer capability in the steam generators. Figure 15.4.2-7 shows that the water volume in the pressurizer increases in response to the heatup, pressurizer water relief begins at 1512 seconds. At approximately 3300 seconds, decay heat generation decreases to a level such that the total RCS heat generation (decay heat plus pump heat) is less than emergency feedwater heat removal capability, and RCS pressure and temperature begin to decrease. O 1319c1c/c32988 15.4-14
i The results show that the core remains covered at all times and that no boiling occurs in the reactor coolant loeps. . 15.4.2.2.3.2 Feedline Ruoture with Offsite Power Unavailable The system response folicwing a feedwater line rupture without offsite power available is similar to the case with offsite power available. However, as a result of the loss of offsite power (assumed to occur at r'eactor trip), the reactor coolant pumps coast down. This results in a reduction in total RCS heat generation by the amount produced by pump ol vation. The reduction in total RCS heat generation produces a milder transient than in the case where offsite power is available. Results presented in Figures 15.4.2-11 and 15.4.2-13 show that pressure in the RCS and main steam system remain below 110% of the respective design pressures. Pressurizer pressure decreases after reactor trip on low-low steam generator water level due to the reduction of heat input .Following this initial decrease, pressurizer pressure increases to a peak pressure of 2502 psia at 746 seconds. This increase in pressure is the result of coolant expansion caused by the reduction in heat transfer capability in the steam generators. Figure O 15.4.2-11 shows that the water volume in the pressurizer. increases in response to the heatup but does not fill the pressurizer. At approximately 1200 seconds, decay heat generation decreases to a level less than the emergency feedwater heat removal capability, and RCS temperatures begin to decrease. The results show that the core remains covered at all times since the l pressurizer does not empty. l 15.4.2.2.4 Canclusion ! Results of the analysis show that for the postulated feedline rupture, the assumed emergency feedwater system capacity is adequats to remove decay heat, to prevent overpressurizing the RCS, and to prevent uncovering the reactor core. i l l i O isiscio/es2ssa 15.616
15.4.4 Single Reactor Coolant Pumo Locked Rotor 15.4.4.1 Identitication of Causes and Accident Description The accident postulated is an in:tantaneous seizure of an RCP rotor. Flow through the affected reactor coolant loop is rapidly reduced, leading to an initiation of a reactor trip on a low flow signal.. Following initiation of the reactor trip, heat stored in the fuel rods continues to be transferred to the coolant causing the coolant to expand. At the same time, heat transfer to the shell-side of the steam generators is reduced, first because the reduced flow results in a decreased tube-side film coefficient and then because the reactor coolant in the tubes cools down while the shell-side temperature increases (turbine steam flow is reduced to zero upon plant trip). The rapid exoansien of the coolant in the reactor core, combined with reduced heat transfer in the steam generators causes an insurge into the pressurizer and a pressure increase throughout the RCS. The insurge into the pressurizer compresses the steam volume, actuates the automatic spray system, opens the power-cperated relief valves, and opens the pressurizer safety valves in that sequence. The three power-cperated relief valves are
~
designed for reliable operation and would be expected to function properly during the accident. However, for conservatism, their pressure-reducing effect as well as the pressure-reducing effect of the spray is not included in the analysis. 15.4.4.2 Analysis of Effects and Consecuences Two digital computer codes are used to analyze this transient. The LOFTRAN(2) code is used to calculate the resulting loop and core coolant flow following the pump seizure. The LOFTRAN code is also used to calculate the time of reactor trip, based on the calculatted flow, the nuclear power following reactor trip, and to determine the peak pressure. The thermal behavior of the fuel located at the core hot spot is investigated using the FACTRANIO) code, using the core flow and the nuclear power calculated by LOFTRAN. The FACTRAN code includes the use of a film boiling heat transfer coefficient. O
~
1319rlo/041288 15.4-16
The following case is analyzed: (3 ' All loops operating, ene locked rotor. At the beginning of the postulated locked rotor accident, i.e., at the time i the shaf t in one of the RCPs is assumed to seize, .the plant is assumed to be , operating under the most adverse steady state operating conditions, i.e., maximum steady state power level, maximum steady state pressure, and maximum , steady state coolant average temperature. When the peak pressure is evaluated, the initial pressure is conservatively estimated as 33 psi above nominal pressure, 2250 psia, to allow for errors in the pressurizer pressure measurement and control channels. This is done to I obtain the highest possible rise in the. coolant pressure during the transient. To obtain the maximum pressure in the primary side, conservatively high it,co pressure drops are added to the calculated pressurizar pressure. The pressure response is shown on Figure 15.4.4-1. 15.4.4.2.1 Evaluation of the Pressure Transient I After pump seizure and reactor trip, the neutron' flux is rapidly. reduced by i control rod insertion effect. Rod motion is assumed to begin 1 second after the flow in the affected loop reaches 87% of nominal flow. No credit is taken l for the' pressure-reducing effect of the pressurizer relief valves (PORVs), pressurizer spray, steam dump, or controlled feedwater flow after plant trip. Although these operations are expected to occur and would result in a lower peak pressure, an additional degree of conservatism is provided by ignoring their effect. , l l l The pressurizer safety valves are assumed to initially open at 2500 psia and ' actinve rated flow at 2575 psia (3% accumulation). O tster.to/osuss 15.4-17 . l' i
15.4.4.2.2 Evaluation of the Effects of CNB in the Core Durine the Accident - For this accident, DNS is assumed to occur in the core and, therefore, an evaluation of the consequences with respect to fuel red thermal transients is l perfor1ned. Results obtained from analysis of this hot spot condition j represent the upper limit with respect to cladding temperaturo and zirconium-water reaction. . In the evaluation, the red power at the hot spot is conservatively assumed to be approximately 2.6 times the average rod power at the initial core power level. 15.4.4.2.3 Film Boiling Coefficient The film boiling coefficient is calculated in the FACTRAN code using the Bishop-Sandberg-Tong film boiling correlation The fluid preperties are evaluated at film te perature (average between wall and bulk temperatures). The program calculates the film coefficient at every time step bcsed on the actual heat transfer conditions at the time. The neutron flux, system pressure, bulk density, and mass flowrate as a function of time are used as program input. h For this analysis, the initial values of the pressure and the bulk censity are used throughout the transient since they are the most conservative with respect to cladding temperature response. For conservatism, DNS was assumed to start at the beginning of the accident. 15.4.4.2.4 Fuel Cladding Gao Coefficient The magnitude and time dependence of the heat transfer coefficient between fuel and cladding (gap coefficient) has a pronounced influence on the thermal results. The larger the value of the gap enefficient, the more heat is transferred between pellet and cladding. Based on investigations on the effect of the gap coefficient upon the maximem cladding temperature during the transient, the gap coefficient was assumed to increase from a steady state 2 value consistent with the initial fuel temperature to 10,000 BTU /hr-f t ..F O 121sv. ion 22ssa 15.4-18
l at the initiation of the transient. This assumption causes energy stored in L/ the fuel to be released to the cladding at the initiation of the transient and maximizes the cladding temperature during the transient. 15.4.4.2.5 Zirconium-steam Reaction The zirconium-steam resetion can become significant above.1800*F (cladding temperature). The Baker-Just parabolic rate equation shown below is used to define the rate of the zirconium-steam reaction. 2 d(w ) = 33.3 x 106 exp - ( ) (15.4-1) et where: 2 w = amount reacted, mg/cm t = time, see
~
T = temperature, 'K The reaction heat is 1510 cal /gm. 15.4.4.3 Results Transient values of RCS pressure, RCS flow, faulted loop flow, nuclear power i het channel heat flux, and clad temperature are shown in Figure 15.4.4-1 and Figure 15.4.4-2. ,. i; Maximum RCS pressure, maximum cladding temperature, and amount of zirconium water reaction are contained in Table 15.4-9. l 15.4.4.4 Conclusions (1) Since the peak RCS pressure reached during any of the transients is less than that which would cause stresses to exceed the faulted condition stress limits, the integrity of the primary coolant system is not endangered. O tais<:1ote:2ss: 15.4-19
(2) Since the peak cladding surface temperature calculated fer the hot spot during the worst transient remains considerably less than 2700*F and the amount of zirconium-water reaction is small, the core will remain in place and intact with no consequential loss of core cooling capability. (3) The results of the transient analysis show that less than 15.0% of the fuel rods will have DNBRs below the safety analysis limit values. O O 1sts a o/c329ss 15.4-20
i 15.4.6 Rupture of a Control Rod Drive Mechanism Housing (Red Cluster Control Assembly Ejection) , 15.4.6.1 Identification of Causes and Accident Description This accident is defbed as the mechanical failure of a control rod mechanism pressure housing resulting in the eje: tion of a rod cluster control assembly (RCCA) and drive shaft. The consequence of this mechanical failure is a rapid positive reactivity insertion and system depressurization together with an adverse core power distribution, possibly 16ading to localized fuel rod damage. 15.4.6.1.1 Design Precautions and Protection Certain features of the VCSPP are intended to preclude the possibility of a rod ejection accident, or to limit the consequences if the accident were to occur. These include a sound, conservative mechanical design of the rod housings, together with a thorough quality centrol (testing) program during assembly, and a nuclear design that lessens the potential ejection worth of RCCAs, and minimizes the number of assemblies inserted at high power levels. 15.4.6.1.2 Mechanical Design The mechanical design is discussed in FSAR Section 4.2. Hechanical design and quality control procedures intended to preclude the possibility of an RCCA [ drive mechanism housing failure are listed below: , I e (1) Each full length control rod drive mechanism housing is completely assembled and shop tested at 4100 psi. (2) The nehanism housings are individually hydrotested after they are i
, attached to the head adapters in the reactor vessel head, and checked during the hydrotest of the completed reactor coolant system.
l l i (3) Stress levels in the mechanism are not affected by anticipated l system transients at power, or by the thermal o vement of the coolant loops. Moments induced by the design-basis earthquake can f be accepted within the e' owable primary working stress range specified by the ASME C a , Section Ill, fer Class ! components. O isistio/o41ssa 15.4-21
=-=e
(4) The latch mechanism housing and rod travel housing are each a single length of forged Type-304 stainless steel. This matarial exhibits excellent notch toughness at all temperatures that will be encountered. A significant margin of strength in the elastic range ,together with the large energy absorption capability in the plastic range gives additional assurance that gross failure of the housing will not occur. The joints between the latch mechanism housing and head adapter, and between the latch mechanism housing and rod travel housing, are threaded joints reinforced by canopy-type red welds which are subject to periodic inspections. '15.4.6.1.3 Nuclear Design Even if a rupture of an RCCA drive mechanism housing is postulated, the operation of a plant utilizing chemical shim is such that the severity of an ejected RCCA is inherently limited. In genersi, the reactor is operated with the RCCAs inserted only far eneagh to permit lead follow. Reactivity changes caused by core depletion and xenon transients are compensated by boron changes. Further, the location and grouping of control RCCA banks are selected during the nuclear design to lessen the severity of an RCCA ejection accident. Therefore, should a RCCA be ejected from its normal position during full power operation, only a minor reactivity excursion, at worst, could be expected to occur. However, it may be occasionally desirable to operate with larger than normal insertions. For this reason, a red insertion limit is defined as a function of power level. Operation with the RCCAs above this limit guarantees adequate shutdown capability and acceptable power distribution. The position of all RCCAs is continuously indicated in the control room. An alarm will occur if a bank of RCCAs approaches its insertion limit or if one RCCA deviates from its bank. There are low and low-low level insertion monitors with visual and audio signals. Operating instructions require boration at icw-level alarm and emergency boration at the low-low alarm. O 131 sti ceo41ssa 15.4-22
15.4.6.1.4 Reactor Protection The reactor protection in the event of a rod ejection accident has been , described in Reference 7. The protection for this accident is provided by the power range high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. These protection functions are described in detail in FSAR Section 7.2. - 15.4.6.1.5 Effects'en Adjacent Housings Disregarding the remote possibility of the occurrence of an RCCA mechanism housing failure, investigations have shown that failure of a hout. .; due to either longitudinal or circumferential cracking is not expected to cause damage to adjacent housings leading to increased severity of the initial accident. 15.4.6.1.6 Limiting Criteria Due to the extremely low probability of an RCCA ejettien accident, limited fuel damage is censidered an acceptable consequence. Ccmprehensive studies of the threshold of fuel failure and of the threshold of significant. conversion of the fuel thermal energy to mechanical energy have been c.arried out as part of the SPERT project by the Idaho Nuclear Corporation (8) . Extensive tests of zirconium-clad UO 2 fuel rods representative of those in PWR-type cores have demonstrated failure thresholds in the range of 240 to 257 cal /gm. However, other rods of a slightly different design have exhibited failures as low as 225 cal /gm. These results differ significantly from the TREAT (9) results, which indicated a failure , threshold of 280 cal /gm. Limited results have' indicated that this threshold decreases by about 10% with fuel burnup. The cladding failure mechanism appears to be melting for zero burnup rods and brittle fracture for irradiated rods. Also important is the conversion ratio of thermal to mechanical energy. This ratio becomes marginally detectable above 300 cal /gm for unirradiated rods and 200 cal /gm for irradiated rods; catastrophic failure, (large fuel dispersal, large pressure rise) even for irradiated rods, did not occur belew 300 cal /gm. O 1319v:10/0419 s 15.4-23 i min me- muni--m a i....-.___.. . i siis
In view of the above experimental results, conservative criteria are applied l to ensure that there is little or no possibility of fuel dispersal in the . coolant, gross lattice distortion, or severe shock waves. These criteria are: 1 (1) Average fuel pellet enthalpy at the hot spot below 225 cal /gm for l unirradiated fuel and 200 cal /gm for irradiate'd fuel; (2) Average cladding temperature at the hot spot below the temperature at which cladding e:rbrittlement may be expected (2700'F); (3) Peak reactor coolant pressure less than that which would cause stresses to exceed the faulted condition stress limits; (4) Fuel melting will be limited to less than 10% of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of Criterien (1) above. 15.4.6.2 Analysis of Effects and Consecuences The analysis of the RCCA ejoction accident is performed in two stages: (a) an average core nuclear power transient calculation and (b) a hot spot heat transfer calculation. The average core calculation is performed using spatial neutron kinetics methods to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity. Enthalpy and temperature transients in the hot spot are then determined by multiplying the average core energy generation by the hot channel factor and performing a fuel rod transient heat transfer calculation. The power distribution calculated without feedback is pessimistically assumed to persist throughout. the transient. A detailed discussion of the method on anclyiis can be found in Reference 10. 15.4.6.2.1 Averace Core Analysis The spatial kinetics computer code, TWINKLE @ ), is used for the av u age core transient analysis. This code solves the two group neutron diffusion theory kinetic equations in one, two, or three spatial dimensions (rectangular 9 131sa o/oc ssa 15.4-24
coordinates) for six delayed neutron groups and up to 2000 spatial points. ).O The computer code inchides a detailed multiregion, transient fuel-clad-coolant heat transfer model for calculating pointwise Doppler, and moderator feedback j , effects. In this analysis, the code is used as a one-dimensional axial kinetics code since it allows a more realistic representation of the spa ~tial effects of
- axial moderator feedback and RCCA movement and the elimination of axial feedback weighting factors. However, since the radial dimension is missing, it is still necessary to employ very conservative methods (described below) of calculating the ejected red worth and hot channel factor. A further description of TWINKLE appears in Saction 15.1.8.
15.4.6.2.2 Hot Scot Analysis The average core energy addition, calculated as described above, is multiplied by the appropriate hot channel factors, and the hot spot analysis is performed using the detailed fuel and cladding transient heat transfer computer code, FACTRAN(6) This computer code calculates the transient temperature distribution in a cross section of a metal clad U0 2fuel rod, and the heat flux at the, surface of the rod, using as input the nuclear power versus time and the local coolant cenditions. The zirconium-water reaction is explicitly represented, and all material properties are represented as functions of temperature. A parabolic radial power generation is used within the fuel rod. FACTRAN uses the Dittus-Boelter(12) or Jens-Lottes(13) correlation to determine the film heat transfer before ONB, and the Bishep-Sandberg-Tong correlation (14) to determine the film boiling coefficient after DNB. The DNB heat flux is not calculated; instead the code is forced into DNB by specifying a conservative ONB heat flux. The gap. heat transfer coefficient can be calculated by the code; however, it is adjusted in order to force the full power steady state pellet temperature distribution to agree with that predicted by design fuel heat transfer codes. O 4 1319v;1o/041968 15.4-25
For full power cases, the design initici hot channel factor (F g ) is input to the code. The hot channel factor during the transient is assumed to increase from the steady state design value to the maximum transient value in 0.1 seconds, and remain at the maximum for the duration of the transient. This is conservative, since detailed spatial kinetics models show that the hot channel factor decreases shortly af ter the nuclear power peak due to power flattening caused by preferential feedback in the hot channel. Further description of FACTRAN appears in Section 15.1.8. if.4.6.2.3 System Overpressure Analysis Because safety limits for fuel damage specified earlier are not exceeded, there is little likelihood of fuel disperni into the coolant. The pressure surge may therefore be calculated on the bacis of conventional heat transfer from the fuel and prompt heat generation in the coolant. The pressure surge is calculated by first performing the fuel heat transfer calculation to determine the average and het spot heat flux versus time. Using this heat flux data, a THINC calculation is conducted to determine the volume surge. Finally, the volume surge is simulated in a plant transient computer code. This code calculates the pressure transient taking into account fluid transport in the system, heat transfer to the steam generators, and the action of the pressurizer spray and pressure relief valves. No credit is taken for the possible pressure reduction caused by the assumed failure of the control rod pressure housing (15) , 15.4.6.2.4 Calculation of Basic Parameters Input parameters for the analysis are conservatively selected on the basis of calculated values for this type of core. The more important parameters are discussed below. Table 15.4-10 presents the parameters used in this analysis. 15.4.6.2.5 Eiected Rod Worths and Hot Channel Factors The values for ejected rod worths and hot channel factors are calculated using three dimensional calculations. Standard nuclear design codes are used in the analysis. No credit is taken for the flux-flattening effects of reactivity feedback. The calculation is performed for the maximum allowed bank insertien O 131stic/o41sas 15.4-26 1
l l at'a given power level as determined by the rod insertion limits. Adverse ! y) xenon distributions are considered in the calculations. . Appropriate margins are added to the results to allow for calculational uncertainties, including an allowance for nuclear power peaking due to ; densification. l 15.4.6.2.6 Reactivity Feedback Weighting Factors The largest temperature rises, and hence the largest reactivity feedbacks, occur in channels where the power is nigher than average. Since the weight of regions is dependent on flux, these regions have high weights. This means that the reactivity feedback is larger than that indicated by a simple single channel analysis. Physics calculations were carried out for temperature changes wi t h a flat temperature distribution, and with a large number of axial and radial temperature distriDutions. Reactivity changes were compared and effective weighting factors determined. These weighting factors take the form of multipliers that, when applied to single channel feedbacks, correct them to affective whole core feedbacks for the aopropriat.e flux shape, in this analysis, since a one-dimensional (axial) spatial kinetics method is e'mployed, axial weigh. ting is not used. In addition, no weighting is applied to the moderator feedback. 'A conservative radial weighting factor is applied to the transient fuel temperature to obtain an effective fuel temperature as a function of time accounting for the missing spatial dimension. These weighting factors were shown to be conservative compared to three-dimensional analysis. 15.4.6.2.7 Moderator and Doooler Coefficient The critical boron concentrations at the beginning-of-life (BOL) and end-of-life (EOL) are adjusted in the nuclear code in order to obtain moderator density coefficient curves which are conservative compared to actual design conditions for the plant. As discussed above, no weighting factor is applied to these results. 4 1319v:1o/041988 15.4-27
.__. . _ . . _ ~ . _ ~1
The Doppler reactivity defect is determined as a function of power level using the one-dimensional steady state computer code with a Doppler weighting factor of 1. The resulting curve is conservative compared to design predictions for this plant. The Doppler weighting factor should be larger than 1 (approximately 1.2) just to make the present calculation agree with design predictions before ejection. This weighting factor will increase under accident conditions, as discussed above. The Doppler defect used as an ! I initial condition is 900 pcm at BOL and 840 pcm at EOL. 15.4.6.2.8- Celayed Neutron Fraction Calculations of the effective delayed neutron fraction (B,ff) typically yield values of 0.70% at BOL and 0.50% at EOL for the first cycle. The accicent is sensitive to b if the ejected rod worth is nearly equal to or I greater than 8 as in zero power transients. In order to allow for future fuel cycles, pessimistic estimates of S of 0.54% at beginning of cycle and 0.44% at end of cycle were used in the analysis. 15.4.6.2.9 Trio Reactivity Insertion The trip reactivity insertion assumed is given in Table 15.4-10 and includes the effect of one stuck rod. These values are reduced by the ejected rod reactivity. The shutdown reactivity was simulated by dropping a red of the required worth into the core. The start of rod motion occurred 0.5 seconds after the high neutron flux trip point was reached. This deley is assumed to consist of 0.2 seconds for the instrument channel to product a signal, 0.15 seconds for the trip breaker to cpen, and 0.15 seconds for the coil to release the rods. The analyses presentad are applicable for a rod insertion time of l 2.7 seconds from coil release to entrance to the dashpot. The choice of such a conservative insertion rate means that there is over 1 second after the trip l point is reached before significant shutdown reactivity is inserted into the l core. This is a particularly important conservatism for hot full power accidents. The rod insertion versus time is described in Section 15.1.4. O isistiomissa 15.4-28
15.4.6.3 Results The values of the parameters used in the analysis, as well as the results of , the analysis, are presented in Table 15.4-10 and discussed below.
}i 15.4.6.3.1 Beginning of Cycle, Full Power .
Control Bank D was assumed to be inserted to its insertion limit. The worst ejected rod worth and hot channel factor were conservatively assumed to be O.20% AK and 6.5, respectively. The peak hot spot clad average temperature was 2524'F. The peak hot spot fuel center temperature exceeded the 80L melting temperature of 4900*F. However, melting was restricted to less than 10% of the pellet'. 15.4.6.3.2 Beginning of Cycle, Zero Power For this condition, control Bank 0 was assumed to be fully inserted and C was at its insertion limit. The weret ejected red is located in control Bank D-and was conservatively assumed to have a worth of 0.855% AK and a hot channel factor of 13. The peak hot spot clad average temperature reached 2476*F. The peak hot spot fuel center temperature reached 4697'F.. 15.4.6.3.3 End-of-Cycle, Full Power Control Bank D was assumed to be inserted to its insertion limit. The ejected , rod worth and het channel factors were conservatively assumed to be 0.21% AK - and 7.0, respectively. This resulted in a peak hot spot clad average temperature of 2414*F. The peak hot spot fuel center temperature exceeded the EOL melting temperature of 4800'F. However, melting was restricted to less than 10% of the pellet. 15.4.6.3.4 End of Cycle, Zero Power , The ejected rod worth and hot channel factor for this case were obtained l assuming control Bank D to be fully inserted and Bank C at its insertion limit. The results were 0.90% AK and 22.5, respectively. The peak clad average and fuel center temperatures were 2344 and 4104*F, respectively. O isiscio/o41ssa 15.4-29
,,,,,,,,n----. --
A summary of the cases presented is given in Table 15.4-10. The nuclear power and hot spot fuel clad temperature transients for the worst cases (BOL full power and zero power) are presented on Figures 15.4.6-1 through 15.4.6-2. 15'.4.6.3.5 Fission Product Release It is assumed that fission products are released from the gaps of all rods entering DNB. In all cases considered, less than 10% of the rods entered DNB based on a detailed three-dimensional THINC analysis. Although limited fuel melting at the hot spot was predicted for the full power cases, in practice melting is not expected since the analysis conservatively assumed that the hot spots before and after sjection were coincident. 15.4.6.3.6 Pressure Surce A detailed calculation of the pressure surge for an ejection worth of one dollar at BOL, hot full power, indicates that the peak pressure does not exceed that which would cause stress to exceed the faulted condition stress limits. Since the severity of the present analysis does not exceed this worst case analysis, the accident for this plant will not result in an excessive pressure rise or further damage to the RCS. 15.4.6.3.7 Lattice Deformations A large temperature gradient will exist in the region of the hot spot. Since the fuel rods are free to move in the vertical direction, differential expansion between separate rods cannot produce distortion. However, the temperature gradients across individual rods may produce a force tending to bow the midpoint of the rods toward the hot spot. Physics calculations indicate that the net result of this would be a negative reactivity insertion. In practice, no significant bowing is anticipated, since the structural rigidity of the core is more than sufficient to withstand the forces produced. Boiling in the hot spot region would produce a net flow away from that region. However, the heat from fuel is released to the water relatively slowly, and it is considered inconceivable that cross flow will be sufficient to produce significant lattice forces. Even if massive and rapid boiling, sufficient to distort the lattice, is hypothetically postulated, the large void fraction in the hot spot region would produce a reduction in the O 131stio/o41sas 15.4-30
l l
\
total core moderator to fuel ratio, and a large reduction in this ratio at the , hot spot. The net effect would therefore be a negative feedback. It can be , concluded that no conceivable mechanism exists for a net positive feedback resulting f. rom lattice deformation. In fact, a small negative feedback may result. The effect is conservatively ignored in the analyses.
~
15.4.6.4 Conclusions ' Even on a conservative basis, the analyses indicate that the described fuel and cladding limits are not exceeded. It is concluded that there is no danger - of sudden fuel dispersal into the coolant. Since the peak pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the reactor coolant system. The analyses have demonstrated that the upper limit in fission product release as a result of a number of fuel rods entering DNS amounts to 10%(15) , O i i i 131ev:ioro41ss; 15.4-31
15.4.7 References
- 1. Technical Specifications, V. C. Sumar Nuclear Station Appendix A to License No. NPF-12, as amended through Amendment Nurber 66.
4
- 2. T. W. T. Burnett, et al., LOFTRAN Code Description, WCAP-7907-P-A (Proprietary), WCAP-7907-A (Non-Preprietary), April 1984.
- 3. J. S. Shefcheck, Acolication of the THINC Program to PWR Design, WCAP-7359-L, August 1969 (Proprietary), and WCAP-7838, January 1972.
~4. F. S. Moody, "Transactions of the ASME," Journal of Heat Transfer, February 1965, Figure 3, page 134.
- 5. ANSI /ANS-5.1-1979, American National Standard for Decay Hea+ Power in Light Water Reactors,1979. -
- 6. H. G. Hargrove, FA.TRAN-A Fortran IV Code for Thermal Transients ia_a,
@2 Fuel Rod, WCAP-7908, June 1972.
- 7. T. W. T. Burnett, Reactor Protection System Diversity in Westinghouse Pressurized Water Reactor, WCAP-7306, April 1969.
- 8. T. G. Taxelius, ed. "Annual Report - Spert Project, Octcber 1968 September 1969", Idaho Nuclear Corporation IN-1370, June 1970.
- 9. R. C. Liimatainen and F. J. Testa, Studies in TREAT of Zircaloy-2-Clad,
@2-Core Simulated Fuel Elements, ANL-7225, January - June 1966, p. 177, November 1966.
- 10. D. H. Risher, Jr., An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods,
,1 WCAP-7588, Revision 1, December 1971. j l 131sa o/o41 sis 15.4-32
- 11. D. H. Risher, Jr. and R. F. Barry, TWINKLE - A Multi-Dimensional Neutron XineticsComputerCode,WCAP-7979-P-A(Proprietary),WCAP-8028-A (Non-Proprietary), January 1975.
- 12. F. W. Dittus and L. M. K. Boelter, University of California (Berkeley),
Pubis. Eng. , 2,433, 1930,
- 13. W. H. Jens and P. A. Lottes, Analysis of Heat Transfer. Burnout, Pressure Drop, and Density Data for High Pressure Water, USAEC Report ANL-4627, 1951.
- 14. A. A. Bishop, et al., "Forced Convection Heat Transfer at High Pressure After the Critical Heat Flux," ASME 65-HT-31, August 1965.
- 15. Risher, D. H., Jr., "An Evaluation of the Red Ejection Accident in Westinghouse Pressurized Water Reactort. Using Spatial Kinactics Methods,"
WCAP-7588, Revision 1-A, January 1975.
'6. Westinghouse letter dated March 25,1986, NS-NRC-86-3116. "Westinghouse
(] Response to Additional Request on WCAP-9226-P/WCAP-9227-N-P, Reactor Core Response to Excessive Secondary Steam Release," (Non-Preprietary). O isis <;1orenses 15.4-33
TABLE 15.4-8 Sheet 1 of 3 TIME SEQUENCE OF EVENTS FOR NAJOR SECONDARY SYSTEM PIPE RUPTURES Accident Eeent Time (see) Wajor Steam Line Rupture A. Offsite power Steam line ruptures 0 available Criticality attained 19 Boron from RWST reaches core 65 Accumulators actuate 76 Peak heat flux attained 76
. Core becomes suberitical ~340 B. Without offsite Steam line ruptures O power Criticality attained 22 Boron from RWST reaches core 69 }; ! Peak heat flux attained ~270 Core becomes suberitical ~370 i
I l 1 I l l l lO l ,s,. ,o _ .s i l
TABLE 15.4-8 Sheet 2 of 3 Oll Accident Event Time, see ; Rupture of Main Feedline rupture occurs 10 feedwater Pipe (Offsite Power Available) Low-low steam generator level reactor trip setpoint reached in affected steam generator 32.4 Red begins to drop 34.4 E=ergency Feedwater is started 92.4 Feedwater lines are purged and emergency feedwater is delivered to two of three intact steam generators 104 O Low steamline pressure setpoint reached 225 Steamline and feedline isolation occurs 235 Steam generater safety valves lift in intact loeps 612 Pressurizer water relief begins 1512 Total RC3 heat generation (decay heat + pump heat) decreases to emergency feedwater heat removal capability ~3300 0 1319v:1o/o32s?8 ,
TABLE 15.4-8 Sheet 3 of 3 0 .; Event Time, see Accident Rupture of Main Feedline rupture occurs 10 feedwater Pipe (OffsitePower Unavailable) ' low-low steam generator level reactor trip setpoint reached in affected steam generator 32.4 Rod begins to drop 34.4 Reactor coolant pump coastdown 36.4 Emergency feedwater is started 92.4 1 , p Feedwater lines are purged and emergency feedwater is delivered to two of three intact steam generators. 104 4 Low steamline pressure setpoint reached 254 Steamline and feedline isolation occurs 264 Steam generator safety valves lift in intact loops 768 Total RCS heat generation decreases to emergency feedwater heat removal capability ~1200 Peak pressurizer water level reached -1470 0 1319. 10/o32988
.-___.__-___..-_________.___________.__.__l__-
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O O O TABLE 15.4-10 PARAMETERS USED IN THE ANALYSIS OF THE R00 CLUSTER CONT _ROL ASSEMBLY EJECTION ACCIDENT Time in Life 8eginning Eeginning End End Power level, % 102 0.0 102 0.0 Ejected rod worth, %ak D.20 0.855 0.21 0.90 Delayed neutron fraction, % 0.54 0.54 0.44 0.44 Feedback reactivity weighting 1.30 2.07 1.30 3.55 Trip reactivity, %ak 4 2 4 2 F before rod ejection 2.61 - 2.61 - O F after r d ejection 6.5 13 7.0 22.5 0 Number of operating pumps 3 3 3 3 Maximum fuel pellet average temperature. *F 4219 3888 4050 3463
^
*
- 4104 Maximum fuel center temperature, 'F 4697 Maximum clad average temperature. *F 2524 2476 2414 ' 2344 Maximum fuel stored energy, cal /gm 186 169 177 +
147
*Less than 10% fuel melt ,
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l l
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TIME 15tti V. C. Sumer Figure 15.4.2 3 l Transient Response To A Steam Line Break Double ! Ended Rupture With
- Offsite Power Available '
I (Case A) I
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.2-5 d .15 E
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- 0. 100, 200. 500. 400. 500, 600.
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- 0. Ig0. 200. 500. 400. 500. 600.
TINC (SCCI V. C. Sumer Figure 15.4.2-4 Transient Response To A Steam 1.ine Break Double Ended Rupture With No Offsite Power Available (Case B) l
)
f I .2 ' { b y ..
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T IP.E 15ECl V. C. Sumer Figure 15.4.2 5 Transient Response To A Steam Line Break Double Ended Rupture With No O Offsite Power Available (Case B) s . o+ eo s- w e h
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w '
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V. C. Sumer Figure 15.4.2 6 Main Feedline Rupture With Offsite Power Nuclear Power and Core Heat Flux vs. Time
~ -- -
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- 2422. -
u 2200.- M [ 2000.- V y 1800. <
- a. !
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# 1 scc. .
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S N g sco. < 8 w t c. 100 101 102 3c5 je e TIME (SEC) V. C. Sumer - Figure 15.4.2 7 Main Feedline Rupture With Offsite Power O Pressurizer Pressure and Water Volume vs. Time
C 4 ~ 700.< W 3 Saturation Hot Leg - 5 a 600. 8' Cold Leg
- ?.
d M
~
' ECD' c0 3 }cl 102 305 tc4 TIME ISCC1 E. ~ u 700. < b Saturetten 6H. g,g L, r - U a 60 0 . ' E a Cole Leg
~
O s50. < 5 E - 500. a 100 , 10 1 102 3c5 ic TIME (SCCI Y. C. Sumer Figure 15.4.2 8 Main Feedline Rupture With Offsite Power Faulted and Intact Loop Coolant Temperatures vs. Time
l 1 1 l
- l. n 2
f
- 1500.<
u ' Intact
$1250.-
0 - g 1000.< 750.- E W 500.. d \ g 250.< Faulted w m O. jg0 101 102 3e5 }c4 TIME ISEC1 E 9 14E
- 6 -
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E y .60E+5 < w :
" .40E.5 Intact {
E ! Faulted l' W .20E 5 e D. 100 101 102 105 }g4 TIME (SEC) V. C. Sumer Figure 15.4.2-9 Main Feedline Rupture With Offsite Power O Steam Generator Pressure and Water Mass vs. Time
I 1 E 1.4 t
- 1.2 u
- 1. N h_
g .8 h .6 ft .4-b
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0. 102 105 ggd ) 100 101 TIME (SEC1 l l V. C. Sumer I 1 Figure 15.4.2-20 ( Main Feedline Rupture Without Offsite Power Nuclear Power and Core Heat & l Flux vs. Time W
l g2ecc. g 2620.- ;
?n . $ 2400.
a g a 2200.- N y2000.< 8 ' g iec0. . c. 1600. 100 101 102 105 jed TIME (SEC)
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u O 8 g isec. < 5 g iccc. - i 5 N j 500.- 8 d
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100 10 1 102 105 10 4 , TIME (SEC1 ! V. C. Sunner Figure 15.4.2-11 Main Feedline Rupture Without Offsite Power O Pressurizer Pressure and Water Volume vs. Time
l l l C 9'l ( 700. s Saturation a Hot Leg a 600. 8
- Cold Leo 8 550.
500. - jed 100 101 102 105 TIME , tSEC) C. [700, s .- seturation g $650. a Hot Leg 600. 8 J Cold Leg y550.< E 500. # 101 102 105 10 100 TlhE ISEC) V. C. Sumer Figure 15.4.2 12 Main Feedline Rupture Without Offsite Power Faulted and intact Loop Temperatures vs. Time h
1 i b G ? e
- 1500.-
U Intact : h1250. I O ' g se00. - -
! 750.-
a W 5ea. O g 250, Faulted u G 0. d 100 10 1 102 305 3c TIME (SEC) E , g .14E 6 < l [ y w w
.32C 6 '
E .30E 6 m S .60E 5 E y .60E 5 w C Intact
,4cE.;
E Faulted W .20E 5 m O. 100 10! 102 105 jed TIME (SEC) l ~ l 1 V. C. Sumer l l Figure 15.4.2-13 l l Main Feedline Rupture i Without Offsite Power Steam Generator Pressure and l Water Mass vs. Time l l
~
~~~
1 l h c-2003. ' I M 1 2 g 2600. m g 2400. l E I M ~~ E 2223. , m 1 U j 2003. 4 5. 6 7. 8. 9. 10.
- 8. 1. 2. 5.
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- e. 1. 2. 5. 4 5.
T!nt ISCCI V. C. Sumer Figure 15.4.4-1 All Loops Operating One Locked Rotor RCS Pressure, RCS Flow and Faulted Loop Flow vs. Time 1
.. . . . . - . l l
- 1. 4 -
o ;,.2 g 3.
.9 ar .6<
4 I .2 < O.
- t. 1. 2. 5. 4 5. 6. 7. 8. 9 10. t TIME ISECl I
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- .
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ISECl
- 7. B. 9. 10.
TIME V. C. Sumer Figure 15.4.4 2 All Loops Operating n One Locked Rotor Nuclear Power, Heat Flux and IQ' Clad Temperature vs. Time n __
~
-_ : ._n.___.___
a G 5. 2., b _ y ,. t g i., gi. i
- g, 3, 3, 3. 4 6. 4. 7. 8. 9 II.
7,st GEC) 8008. c .,8_.., L 58e8. ..... - .......... T D P fRATL8tf doet. IM' CLAD DJTitt TDepu TU9E r FW.MM awn l AGE M. I 1988.
.. 2. 8. 4 5. 4. 7. 8.
! 3. 8. TDer CBEC) l V. C. Sumer l Figure 15.4.5 1 1 Rod Ejection Accident BOL HFP Nuclear Power, Hot Spot Fuel and Clad Temperature vs. Time 1 l
.,.s ,,. - - , - -
i 0 -l j ... . . R 1 ,, . Ii
, is i
15
' .. .s i, i.s a. 1.s s. s.s 4 TM (SECJ seer. '
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- g. .. 3 8. 4 5. 4. 7. O. 9. 10.
TM dBECJ V. C. Summer Figure 15.4.6 2 Rod Ejection. Accident BOL HIP Nuclear Power, Hot Spot Fuel and Clad Temperature vs. Time
** i.=+.
l l ATTACHMENT 4 ,; I LOCA ACCIDENT ANALYSES FOR THE - V. C. SUMMER PLANT i TRANSITION TO 17x17 VANTAGE 5 FUEL O IO 00681:6/880504 ,
- i, l
15.3 CONDITION III - INFRE0 VENT FAULTS p - O By definition, Condition III occurrences are faults which may occur very infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods although sufficient fuel damage might occur to preclude resumption of the operation for a considerable outage time. The release of radioactivity will not be sufficient to interrupt or restrict public use of those areas beyond the exclusion radius. A . Condition .III fault will not, by itself, generate a Condition IV fault or result in a consequential loss of function of the reactor coolant system or containment barriers. For the purpose of this report the following faults have been grouped into this category:
- 1. Loss of reactor coolant, from small ruptured pipes or from cracks in large pipes, which actuates the emergency core cooling system.
- 2. Minor secondary system pipe breaks.
O 3- t# i' vert #1 1 e4 9 er <" > 811 4"te 4 ar aer 9 4t4 #-
,i 1
- 4. Complete loss of forced reactor coolant flow. !
- 5. Single rod cluster control assembly withdrawal at full power.
l Each of these infrequent faults are analyzed in this section. In general, i each analysis includes an identification of causes and description of the f accident, an analysis of effects and consequences, a presentation of results, and relevant conclusions. The tire sequence of events during applicable Condition III faults 1 and 4 above is shown in Table 15.3-1. O) 1243mio/o412:a 15.3-1
15.3.1 LOSS OF REACTOR COOLANT FROM SMALL RUPTURED PIPES OR FROM CRACKS IN LARGE PIPES WHICH ACTUATES THE EMERGENCY CORE COOLING SYSTEM 15.3.1.1 Identification of Causes and Accident Description A loss of coolant accident is defined as a rupture of the reactor coolant system piping or of any line connected to the system. See Section 5.2 for a more detailed description of the loss of reactor coolant accident boundary limits. Ruptures of small cross section will cause expulsion of the coolant at a rate which can be accommodated by the charging pumps which would maintain an operational water level in the pressurizer, permitting the operator to execute an orderly shutdown. The coolant which would be released to the containment contains the fission products existing in it. The maximum break size for wnich the normal makeup system can maintain the pressurizer level is obtained by comparing the calculated flow from the reactor coolant system through the postulated break against the charging pump makeup flow at normal reactor coolant system pressure, i.e., 2250 psia. A makeup flow rate from one centrifugal charging pump is adequate to sustain pressurizer level at 2250 psia for a break through a 3/8 inch diameter hole. This break results in a loss of approximately 17.5 lb/sec. Should a larger break occur, depressurization of the reactor coolant system causes fluid to flow to the reactor coolant system from the pressurizer, resulting in a pressure and level decrease in the pressurizer. Reactor trip occurs when the pressurizer low pressure trip setpoint is reached. The safety injection system is actuated when the appropriate setpoint is reached. The consequence of the accident are limited in two ways:
- 1. Reactor trip and borated water injection complement void formation in causing rapid reduction of nuclear power to a residual level corresponding to the delayed fission and fission product decay.
- 2. Injection of borated water ensures sufficient ficoding of the core to prevent excessive clad temperatures.
O 1243cio/o412ss 15.3-2
i Before the break occurs, the plant is an equilibrium condition, i.e., the heat U generated in the core is being removed via the secondary system. During blowdown, heat from decay, hot internals and the vessel continues to be transferred to the reactor coolant system. The heat transfer between the reactor coolant system and the secondary system may b'e in either direction, depending on the relative temperatures. In the case of continued heat addition to the secondary, system pressure increases and steam dump may occur. Makeup to the secondary side is automatically provided by the emergency feedwater pumps. The safety injection signal stops norrel feedwater flow by closing the main feedwater line isolation valves and initiates emergency feedwater flow by starting the emergency feedwater pumm. The secondary flow aids in the reduction of reactor coolant system pressure. When the reactor coolant system depressurizes to 600 psia, the accumulators begin to inject water into the reactor coolant loops. The reacter coolant pumps.are assumed to be tripped at the initialization of the accident, and O fects of pump coastdown are included in the blowdown analyses. 15.3.1.2 Analysis of Effects and Consecuences O 15.3.1.2.1 Method of Analysis For loss-of-coolant accidents due to small breaks less than 1 square foot, the NOTRUMP computer code (13,14] is used to calculate the transient depressurization of the RCS as well as to describe the mass and enthalpy of flow through the break. The NOTRUMP computer code is a state-of-the-art one-dimensional general network code incorporating a number of advanced - features. Among these are calculation of thermal non-equilibrium in all fluid volumes, flow regime-dependent drift flux calculations with counter-current flooding limitations, mixture level tracking logic in multiple-stacked fluid nodes and regime-dependent heat transfer correlations. The NOTRUMP small-break LOCA emergency core cooling (ECCS) evaluaticn model was developed to determine the RCS response to oesign basis small break LOCAs, and to address NRC concerns expressed in NUREG-0611, "Generic Evaluation of Feedwater Transients and Small Break Loss-of-Coolant Accidents in Westinghouse-Designed Operating Plants". O 1243cio/o412:s 15.3-3
TFe reactor coolant system model is nodalized into volumes interconnected by flowpaths. The broken loop is modelled explicitly, while the two intact loops are lumped into a second . loop. Transient behavior of the system is determined h from the governing conservation equations of mass, energy, and momentum. The multinode capability of the program enables explicit,' detailed spatial representation of various system components; which, among .other capabilities, enables a proper calculation of the behavior of the loop seal during a loss-of-coolant accident. The reactor core is represented cs heated control volumes with associated phase separation models to permit transient mixture height calculations. Detailed descriptions of the NOTRUMP code and the evaluation model are provided in References 13 and 14. Safety injection systems consist of gas pressurized accumulater tanks and pumped injection systems. Minimum emergency core cooling system availability is assumed for the analysis. Assumed pumped safety injection characteristics as a function of RCS pressure used as boundary conditions in the analysis are shown in Figure 15.3-1. The injection rate is based upon the pump performance curves, but degraded for conservatism and to account for possible reduced injection rates due to pump cooling recirculation miniflow operation. Injection is delayed after the occurence of the injection signal as indicated in Table 15.3-1 to account for diesel generator startup and emergency power bus loading in case of a loss of offsite power coincident with an accident. Peak clad temperature calculations are performed with the LCCTA-IV code using the NOTRUMP calculated core pressure, fuel rod power h! story, uncovered core steam flow and mixture height as boundary conditions. Figure 15.3-2 depicts the hot rod axial power shape used to perform the smail break analysis. This shape was chosen because it represents a distribution with power concentrated in the upper regions of the core. Such a distribution is limiting for small-break LOCAs because it minimizes coolant level swell, while maximi:ing vapor superneating and fuel rod heat generation in the uncovered elevations. Figure 15.3-3 presents the normalized core power curve as a function of time after reactor trip. The scram delay times denoted in Table 15.3-1 reflect the assumotion that the core is assumed to continue to operate at full rated power until the control rods are completely inserted. O 1243v:10/041288 15.3-4 , 1
{} 15.3.1.2.2 Results
.i tO This section presents results of the limiting break size analysis as j determined by the highest peak fuel rod clad temperature for a range of break 1
sizes. The limiting break size was found to be a.3-inch diameter cold leg j break. The maximum temperature attained during the transient was 2095'F. ! Important parameters are summarized in Table 15.3-2, while the key transient l event times are listed in Table 15.3-1. Figures 15.3-4 through 9 show for the . three-inch. break transient, respectively: RCS pressure,
- Core mixture level, Peak clad temperature,
- Core outlet steam flow, Hot spot rod surface heat transfer coefficient, and Hot spot fluid temperature.
During the initial period of the small-break transient,'the effect of the O t.g break flow is not strong enough to overcome the flow maintained by the reactor i: recirculation cooling pumps as they coast down. Normal upward flow is maintained through the core and core heat is adequately removed. At the low heat generation rates following shutdown the fuel rods continue to be well cooled as long as the core is covered by a two phase mixture level. From the c1'ad temperature transient for the 3-inch break calculation shown in Figure 15.3-6, it is seen that the peak clad temperature occurs near the time at which the core is most deeply uncovered when the top of the core is steam cooled. This time is also accompanied by the highest vapor superheating above the mixture level. 15.3.1.2.3 Additional Break Sizes Studies documented in references 9 and 10 determined that the limiting small-break size occurred for breaks less than 10 inches in diameter. To insure that the 3-inch diameter break was limiting, calculations were run with breaks of 2 inches and 4 inches. The results of these calculations are shown O 1243v;1D/041288 15.3-5 i 1
in the Sequence of Events Table 15.3-1, and the Results Table 15.3-2. Plots of the following parameters are shown in Figures 15.3-10 through 15 for the h-2-inch break, and Figures 15.3-16 through 21 for the 4-inch break. RCS pressure, Core mixture level, _ Peak clad temperature,
- Corv outlat steam flow,
- Hot spot rod surface heat transfer coefficient, and Hot spot fluid temperature.
As seen in Table 15.3-2 the maximum clad temperatures were calculated to be less than that for the 3-inch break. 15.3.1.2.4 Additional Analysis NUREG-0737IIII, Section II.K.3.31, required plant-specific small break LOCA analysis using an Evaluation Model revised per Section II.K.3.30. In accordance with NRC Generic Letter 83-65(12[ generic analyses using NOTRUMP(17,18) were performed and are presented in KCAP-11145(15) Those g results demonstrate that in a comparison of cold leg, hot leg and pump suction leg break locations, the cold leg break location is limiting. Analyses of a LOCA in the pressurizer vapor space such as that caused by opening a pressurizer relief valve or a safety valve were provided in WCAP-9600(10) The conclusion presented in WCAP-9600 is that these breaks are not limiting since little or no core uncovery will take place. WCAP-9600 states that the analyses reported therein apply to all Wastinghou=e designed plants. Calculations were also performed for the Virgil Summer plant with the NOTRUMP I13'143 and LOCTA-IV I13 codes to examine the influence of initial loop fluid operating temperatures on small break LOCA peak clad temperature. The results showed that peak clad temperature decreased as loop operating temperature decreased. O 1243rio/c412ss 15.3-6
n 15.3.1.3 Conclusions () Analyses presented in this section show that the high head portion of the ; emergency core cooling system, together with accumulators, provide sufficient ; 1 core flooding to keep the calculated peak clad temperatures below required , limits of 10 CFR 50.46. Hence, afeguate protection is afforded by the j emergency core cooling system in the event of a small break loss of coolant l accident. i 15.3.2 MINOR SECONDARY SYSTEM PIPE BREAKS 15.3.2.1 Identification of Causes and Accident Description t Included in this grouping are ruptures of secondary system lines which would result in steam release rates equivalent to a 6 inch diameter break or smaller. 15.3.2.2 Analysis of Effects and Consequences
/~'s Minor secondary system pipe breaks must be accommodated with the failure of V
only a small fraction of the fuel elements in the reactor. Since the results of analysis presented in Section 15.4.2 for a major secondary system pipe rupture also meet this criteria, separate analysis far minor secondary system pipe breaks is not required. The analysis of 'he t more probable accidental opening of a secondary system steam dump, relief or safety valve is presented in Section 15.2.13. These analyses are illustrative of a pipe break equivalent in size to a single valve opening.
- 3. The iodine partition factor for activity released from the break is 0.1.
- 4. The concentration of radicactive nuclides in the reactor coolant is listed in Table 11.1-2 for the conservative case and in Table 11.1-5 for the realistic case.
O 1243cioec412ss 15.3-7
Using the previously listed assumptions, isotopic releases to the environment are determined to be those listed in Tables 15.3-6 and 15.3-7 for the . realistic and conservative cases, respectively. Gamma, beta and thyroid doses at the site boundary for the realistic case are
~4 7.30 x 10 -6 Rem, 6.81 x 10 Rem and 5.90 x 10 Rem, respectively.
~0
~7 Corresponding doses at the low population zone are 8.20 x 10 Rem, 7.65 x
~I Rem and 6.63 x 10 -5 Rem, respectively.
10 Gamma, beta and thyroid doses at the site boundary for the conservative case
-2 are 3.18 x 10 -2 Rem, 3.66 x 10 Rem and 6.72 x 10 Rem, respectively.
~1
-3 Corresponding doses at the low population zone are 1.85 x 10 Rem, 2.13 x
-2 Rem, respectively.
10'3 Rem and 3.90 x 10 Doses resulting from this accident are well within the limits defined in 10 CFR 20 (25 Rem whole body and 300 Rem thyroid). O l l 1 el 1243mio/0412sa 15.3-8 l
p 15.
3.8 REFERENCES
fS O '
- 1. Bordelon, F. M. , 'et al . , "LOCTA-IV Program: Loss of Coolant Transient Analysis," WCAP-8301 (Proprietary) and WCAP-8305 (Non-Proprietary), June, 1974.
- 2. Hellman, J. M., "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary) and WCAP-8219-A (Non-Proprietary), March,1975.
. 3. Altamore, S. and Barry, R. F., "The TURTLE 24.0 Diffusion Depletion Code,"
WCAP-7213-P-A (Proprietary) and WCAP-7758-A (Non-Proprietary), January, 1975.
- 4. Barry, R. F., "LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, September, 1963.
- 5. Baldwin, M.' S., Merrian, M. M., Schenkel, H. S. and Van De Walle, D. J.,
O ^m ev >e t4 # er ' e< rie ^cc48 "t> c e e er ee r s> te rreav ,cx Transients in Westinghouse PWR's," WCAP-8424, Revision 1, June, 1975.
- 6. Bordelen, F. M., "Calculation of Flow Coastdown After loss of Reactor Coolant Pump (PHOENIX Code)," WCAP-7973, September, 1972.
k
- 7. Burnett, T. W. T., et al., "LOFTRAN Code Description," WCAP-7907, June, L 1972.
- 8. Hargrove, H. G., "FACTRAN - A FORTRAN-IV Code for Thermal Transients in a UO Fuel Rod," WCAP-7908, June, 1972.
2 l
- 9. Salvatori, R., "Westinghouse ECCS Plant Sensitivity Studies," WCAP-8340 (Proprietary) and WCAP-8356 (Non-Proprietary), July, 1974.
O 1243v:1o/e412ss 15.3-9
- 10. "Report on Small Break Accidents for Westinghouse NSSS System, "WCAP-9600 g (Proprietary) and WCAP-9601 (Non-Proprietary), June, 1979. T'
- 11. "Clarification of THI Action Plan Requirements", HUREG-0737, November, 1983.
- 12. NRC Generic Letter 83-85 from D. G. Eisenhut, "Clarification of TMI Action Plan Item II.K.3.31", November 2,1983.
- 13. Meyer, P. E., "NOTRUMP - A Nodal Transient Small Break and General Network Code", WCAP-10079-P-A, August 1985.
- 14. Lee, N. et al., "Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code", WCAP-10054-P-A, August 1985.
- 15. Rupprecht, S. D., et al., "Westinghouse Small Break LOCA ECCS Evalu. tion Model Generic Study With the NOTRUMP Code"; WCAP-11145.
O O 1243tio/o412ss 15.3-10
'5 TABLE 15.3-1 gd .
TINE SEQUENCE OF EVENTS FOR CONDITION III EVENTS Small-break Loss of Coolant Accident Event Time (s) Break Size: 2-Inch 3-Inch 4-Inch Break occurs 0 0 0 Reactor trip signal 125.7 34.15 21 07 t Core power shutdown 131.4 39.85 26.77
~
Safety injection signal 143.3 49.84 33.04 O Safety injection begins 175.3 81.84 65.04 Top of core uncovered 1605 726 450 , Accumulator injection begins N/A 1183 660 Peak clad temperature occurs 2550 1374 752 i Core recovered 5575 3700 1875 O 124 v:1o/o412ss 15.3-11
, - - - - - - - - - - --,,-,-,-,,a-. ,, , -,r, --. . - , ,e - --p--- m.,n--.mre, ,
TABLE 15.3-1 (Continued) llgg i TIME SEQUENCE OF EVENTS FOR I CONDITION III EVENTS . e Time Accident Event (Seconds) f l
- Complete Loss of Forced Reactor Coolant Flow
- 1. Three pumps in operation, All operating pumps loos power l three pumps coasting down and begin coasting down 0 l l
l l Reactor coolant pump under-voltage trip pointed reached 0 Rods begin to drop 1.5 llll Minimum DNBR occurs 3.0 l Complete Loss of Forced l Reactor Coolant Flow I
- 2. Two pumps in operation, All operating pumps lose power l two pumps coasting down and begin coasting down 0 Reactor coolant pump under-voltage trip point reached 0 Reds begin to drop 1.5 Minimum DNBR occurs 2.7 0
124sv:io/o412ss 15.3-12
TABLE 15.3-2 SMALL-BREAK LOSS OF COOLANT ACCIDENT CALCULATION t RESULTS i PARAMETER VALUE Break Size: 2-Inch 3-Inch 4-Inch Peak clad temperature ('F) 1284 2095 1314 Elevation (ft) 11.75 12.00 11.50 Zr/H2O cumulative reaction Maximum local (%) . 0.28 9.44 0.11 Elevation (ft) 11.75 12.00 11.50 Total core (%) < 0.3 < 0.3 < 0.3 Rod Burst Nor.a None None SIGNIFICANT INPUT PARAMETERS Licensed core power 2775 MW Peak linear heat generation rate 13.303 kW/ft 4 Accumulator
- j. Tank water volume 1014 ft 3 i
Pressure 600 psi l O v i n to/o41:ss 15.3-13
..._.7,, . . .
i
/ 400.e .,
I sse.e i L . I m.e ... f i e s .~ 0. e a A O e a e a; 2ct.e e Q w a u e e c :4.e
\. a 0
e m tee.e ee.e
^
%' N .
*8.e m.e 2.w. e im.e 2m. o :m. . :w. . e Pressure (psi)
Figure 15.3-1. Pumped safety injection rate as a function of reacter coolant system pressure.
-_ ; .- - :_...- - - .. _ - - .-. ::.2. - .: :: - -.-...- - ._.._. -... .. . - _ - - - - - -
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.w.. ..s .. - ._
. . .z.a. . w.
. .a. .. : :. : .
. . ..J l . t 12,5 7.5 10.0 2.5 5.0 0.
Core Elevatien (ft) Tip,tre 15.3-2. PeaX red axial power shape as a function of core average linear ha.at generation rate, O
O . 10 0 e= 4 - TOTAL RESIOUAL HEAT (W.fH 4% SHUTDOWN) 2 - 1 - E to 8 ~ l
* =
4 4 - h E 2 - C
! ' 's
~
=
- E 4 -
O jo-2
= -
I i ! !!!!!! I I I 11111l 1 1 I I1111] I l l I t illl~ l l I I t ill 2 4 e s 10 3 2 4 e a 104 100 2 4 e a to i 2 4 e a 10 2 10-1 2 4 s a TIME AFTER SHUTDOWN (SECONOSI Figure 15.3-3. Normalized core heat generation rate following i shutdown (full rod insertion). iO l
~* -.- n, hI uupu
e. 2400. 2200. i 2000. l 1000. -- . 5 : E
- 1600. I
$ 1400. '
U l c. E 1200. .. 5 d
,1000.
U B 800. i 600. W _ _ 400. 200 500. 1020. 1500. 2000. 2500. 5000. 5500. 4000. B. ' TIME (SEC) l l 1 Figure 15.3-4. Reactor coolant systes depressurization transient (3-inch break).
]
. ei
~
U i O . 54 .
- 52.
50. C 28, 5 d All D 26. p, a w B E 24 E ' pg' ., Top of O s,2' u Core !
, / t 20.
J 18. 5000. 5500. 4000. D. 500. 1000. 1500. 2000. 2500. TIME ISEC) . Figure 15.3-3. Ct.,re mixture height (3-inch break). lO l
O 5000. , C -
,25se.
o
~ 382,.. -
A_ , L
,1523. 7 , ,
E / \ g , e % 12.0 FT.~ W- 18C2. i i te ~ s % V % as
/ '_ . . .. _ ._ ._ L_ _
~
9.5 FT.
'5'2 3 . 750.10:t.1258.15:8.1750. 2003. 2250. 2580. 2750. 5823. 5250. 55cc. 5752. 400.
7!MC 15CCI Figure 15.3-6. Clad tamparature transient at peak tamperature elevation and a selected lower elevation (3-inch break). l l l
\
l' 1400. t 1200. I 1000.
~
u M s g 800. B
.2 600.
w E
> 400.
d O ~ 8 200. I w v y% J 1 - .. _.L fn , j
.__..2,,_ :m3L 7-1 5
u 4 - 7- iir r, r l' 9. t
-200.
~400 D. 500. 1000. 1500. 2000, 2500. 5000. 5500. 4020.
l-T!KE (SCC) i; l l Figure 15.3-7. steam mass flowrate out the top of the core (3-inch break). l O
e 183 1 E It 6 9.5 FT. k l A % A N W. A lj m4 [ 3,j i i-N ,
~
. g i s x ni
-s. i s
- I/
P 'w M y ._m , ll ,
. _-:-u -
s 12.0 FT. 5 ' ' 331 I I i I L 9 3k80. 75C.182a.1250.1582.1750. TIMC 15CC) 2823. 2252. 2128. 2758. 5402. 5250 Figure 15.3-8. Clad surface heat transfer coefficiants on the peak power rod (3-inch break). i l 9i
1 I [' j 5888.
- 6 l
C E%S8.
= 3000. ,m N
1588.' i i g /
# e ,
12.0 TT.
# I E *% -se 1.
gi.... , , t W / ,\ /% I E ' ' '
# 1 8 .
'% ~'
ysea. -- y-- - 9.5 FT. 5'20, 752. 18:2. 1252. 1520. 1758. 2000. 2253. 2583. 2758. 8020. 5258. 5583 T!rt ISICI j Figure 15.3-9. Fluid temperature at the peak temperature elevation and a selected lower elevation (3-inch broaA) l0
1 4 I 2400. 2200. 2000. a: [1800. w 1600. e E r g 1400. w
$ 'Q D 1200.
( 1000. g 800. \ 600 2000. 5000. 4000. 5000, 6000. B. 1000. TIME ISEC1 Tigure 15.3-10. Reactor coolant systsu depressurization transient (2-inch break). O
y..Q 4 We y- 6%*W O .. 54. . I ' 52. - 50. C
- 29. .
a i W D I El/y W 26. 5 fh' n E w 24. D 8 O. -
- --- ---- - -- p_ ,e rg;
'g qv 20.
l kg l I8 2. 5000. 4000, 5000. 6000. 1000. 2000. l TIME (SEC) Figure 15.3-11. Core sixture height (2-inch break). 4 O
f l
)
5408. 1 Y,2548. 8 M
~ 2003.
H
,1888.
E __ 11.75 FT. k
- r. [-
. - N s w _
- 1882. ,
s
~, ~~ _
g
/
! ' ' . .__.__ L _ __ _ _
g,,,, 10.25 FT. I C 5500. 4t00. 45:2. ECOC. 550 . (COO. IIO2. 2002. 2120. 8:00. TIMC lltCI Figure 15.3-12. Clad temperature tranziant at peak temperature elevation and a selected lower elevation (2-inch break) . i ) O
2000. . 1000. 1600. U d s 1400. E a
~
1200. E d 1000. O }- 800. 8 , g 600. 8 s00. : I 200. P. ' N - -, ' -c'", ;, , - _ , _ 8 5000. 4000. 5000. 6000. i D. 1900. 2000. TIME ISEC) Figure 15.3-13. Steam mass flovrate out the top of the core (2-inch break). O
3 e O O Id e
' n I Il 4
e
- a. 10.25 FT. T
%ll 8
r .,.%+ . p .wy..,. .,y, j,p g ,,, , g a ' ;>. . t
- i
'9A - . _, ; ; w h, %_ ,
~
jL 11.75 FT. 1 E y O 4523. 5800. E100. 4000. 120 2500. 1 30. 5500. 42 3. 1530. 2230. 11MC f 5tti Figure 15.3-14. clad surface heat transfer coefficiants on the peak power rod (2-inch break). . I O 4
i O . \ l ssee. > C ; 35 1...
= 30tt.
1589. 11,75 FT. l ,-w
~
18., 6.. f w % - a r*%' W~-.+i,, _ - -w --- O
=
3 508.
" 10.25 FT.
5*:0. 5522- SCCO-5520. 4:00, 45:3.
$I: . 2C20. 25ce. 5820.
TIMC llCCi Figure 15.3-15. Fluidtemperatureatthepeaktemperatureelevation ' and a selected lower elevation (2-inch break) O i
2400. 2200. . 2000. 1800. 3 5 E
- tee 0. .
w g 1400. E 1200. g a
,1000.
\ g e
S 800. 600..
'- ,/N 400. vw v
200 B. 250. 500, 750, 1000. 1250. 1500. 1750. 2000. 2250. TIME ISEC) Figure 15.3-16. Reactor coolant system depressurization transient
.(4-inch break)
O
- 54. .
- 52. ; .
I - 50. {28. 2 W 26. l' y
\
u- 1 N Q 24. E of 22' __ -_.. _ -- __ 4E-- T 20.
\ IX /YiV y y
- 19. ,
16 B. 250. 500. 750. 1000. 1250. 1500. 1750. 2000. 2250. TIME (SEC) Core :nixture height (4-inch break) Figure 15.3-17. O I
1 . l t l l l L l 1 1
. 5808.
C
,2588.
w 8
- seco.
1 i l 8 .
* !!23.
11.5 FT.
^
y1:20. <' G l 1 s ,. A _
!t .__ __ __
9.5 FT. g l 0 *g . 223. 423, 600. 800. It:3. 1220. 1400. 1820. 1600. 2003. 32:0. 2403. TIMC ISCCI l Figure 15.3-18. Clad temperature transient at peak temperature elevation and a selected lower elevation (4-inch break) e
i O ' 1800. 1600. _ i l 1400.
~
u b g12e0. d 8 g le00. 8
> 600.
U . d - O 8 00-k' 8 400. , J 200. (J , M ~ %L y Lg _, 0 750. 1900. 1250. 1500. 1750. 2000. 2250.
- 9. 250. 500.
TIME' ISEC)
- 9. Steam mass flowrate out the top of the core fff[nchbrea)
O
'~
e e , 103 g
~
C .5 FT. h , Il l k '/h-I ' *- / w - _
.9 n [}f i 3,,
o:i a ' t 'A t e ~ L*" g( _
=c y Lj -
r f [. 11.5' FT. g. w 183 G E 3Cf
- 8. 202. 402. 622. 8:3. 1820. 1222. 14C2. 1622. 1622. 2922. 2290. 2400 TIMC ISCCI I
Figure 15.3-20. Clad surface heat transfer coefficients on the peak power rod (4-inch break) l l l
'~
5 i O. .ii 5488. C - m R$tt.
- 3883. I I
e 11.5 FT. E y l {IC20. w p g u=. . _ . .._.. _ .. .... 9.5 FT.
'8 . 220. 430. 623. gre, 1800. 1200. 14C2. !$80, 1930. SeCO. 22::. 24C2.
Tirit 15CCI i Figure 15.3-21. Fluid temperature at the peak temperasure elevation and a selected lower elevation (4-inch break) i O
15.4 CONDITION IV - LIMITING FAULTS : O 1 Condition lY occurrences are faults that are not expected to take place, but are postulated because their consequences would include the potential for the
' release of significant amounts of radioactive material. These are the most drastic occurrences that must be designed against and represent limiting l
design cases. Condition IV faults shall not cause a fission product release [ d' to the environment resulting in an undue risk to public health and safety in excess of guideline values of 10 CFR 100. A single Condition IV fault shall not cause a consequential loss of required functions of systems needed to cope with the fault including those of the emergency core cooling system (ECCS) and the containment. For the purposes of this report the following faults have been classified in this category: (1) Major rupture of pipes containing reactor coolant up to and including double-ended rupture of the largest pipe in the reactor coolant system (RCS), i.e., less-of-coolant accident (LOCA); (2) Major secondary system pipe ruptures; (3) Steam generator tube rupture; (4) Single reactor coolant pump (RCP) locked rotor; (5) Rupture of a control rod mechanism housing (rod cluster control assembly (RCCA) ejection). Each of these five limiting faults is analyzed in Section 15.4. In general, each analysis includes an identification of causes and description of the I-accident, an analysis of effects and consequences, a presentation of results, j and relevant conclusions. O 1s43v:10/o42988 15.4-1 ,
e 15.4.1 MAJOR REACTOR COOLANT SYSTEM PIPE RUPTURES (LOSS OF COOLANT ACCIDENT) Q, 15.4.1.1 Identification of Causes and Frequency Classification A loss-of-coolant accident (LOCA) is the result of. a pipe rupture of the RCS pressure boundary. For the analyses reported here, a major-pipe break (large break) is defined as a rupture with a total cross-sectional area equal to or greater than 1.0 ft2. This event is considered an ANS Condition IV event, a limiting fault, in that it is not expected to occur during the lifetime of Virgil C. Summer, but is postulated as a conservative design basis. The results for the small break loss of coolant accident are presented in Section 15.3.1. The boundary considered for loss of coolant accidents are related to connecting pipe is defined in Section 3.6. The Acceptance Criteria for the LOCA are described in 10 CFR 50.46 (10 CFR 50.46 and Appendix K of 10 CFR 50 1974)(1) as follows:
- 1. The calculated peak fuel element clad temperature is below the requirement of 2,200*F. h
- 2. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1 percent of the total amount of Zircaloy in the reactor.
- 3. The clad temperature transient is terminaOd at a time when the core geometry is still amenable to cooling. The localized cladding oxidation limitof17percentisnotexceededduri}gorafterquenching.
- 4. The core remains amenable to cooling during and after the break.
- 5. The core temperature is reduced and decay heat is removed for an extended period of time, as required by the long-lived radioactivity remaining in the core.
O 1343rlo/0429ss 15.4-2
-a 1 . These criteria were established to provide significant margin in emergency ~
core cooling system (ECCS) performance following a LOCA. WASH-1400 (USNRC 1975)(10) presents a recent study in regards to the probability of occurrence of RCS pipe ruptures. , 15.4.1.2 Sequence of Events and Systems Operations - Should a major break occur, depressurization of the RCS results in a pressure . decrease in the pressurizer. The reactor trip signal subsequently occurs when the pressurizer low pressure trip setpoint is reached. A safety injection signal is generated when the appropriate setpoint is reached. These countermeasures will limit the consequences of the accident in two ways:
- 1. Reactor trip and borated water injection supplement void formation in causing rapid reduction of power to a residual level corresponding to fission product decay heat. However, no credit is taken in the LOCA 1
analysis for the boron content of the injection water. In addition, the insertion of control rods to shut down the reactor is neglected in the
~
large break analysis, i
- 2. Injection of borated water provides for heat transfer from the core and prevents excessive clad temperatures.
l 15.4.1.3 Description of Large Break Loss-of-Coolant Accident Transient The sequence of events following a large break LOCA is presented in Table 15.4.1-1. Before the break occurs, the unit is in an equilibrium condition; that is, the heat generated in the core is being removed via the secondary system. During blowdown, heat from fission product decay, hot internals and the ve n el continues'to be transferred to the reactor coolant. At the oeginning of the blowde.n phase, the entire RCS contains subcooled liquid which transfars heat from the core by forced convection with some fully developed nucleate boiling. After the break develops, the time to departure from nucleate u4sao/oossa 15.4-3
boiling is calculated, consistent with Appendix K of 10 CFR 50.(1) Thereafter the core heat transfer is unstable, with both nucleate boiling and film boiling occurring. As the core becomes uncovered, both turbulent and laminar forced convection and radiation are considered as core heat transfer mechanisms. The heat transfer between the RCS and the secondary system may be in either direction, depending on the relative temperatures, in the case of continued heat addition to the secondary system, the secondary system pressure increases and the main steam safety valves may actuate to limit the pressure. Makeup water to the secondary side is automatically provided by the emerSancy feedwater system. The safety injection signal actuates a feedwater isolation signal, which isolates normal feedwater flow by closing the main feedwater isolation valves, and also initiates emergency feedwater flow by starting the emergency feedwater pumps. The secondary flow aids in the reduction of RCS pressure. When the RCS d6 pressurizes to 600 psia, the accumulators begin to inject berated water into the reactor coolant loops. Ine conservative assumption is made that accumulator water injected bypasses the core and goes out througr h the break until the termination of bypass. This conservatism is again consistent with Appendix K of 10CFR50. Since loss of offsite pcwer (LOOP) is assumed, the RCPs are assumed to trip at the inception of the accident. The effects of pump coastdown are included in the blowdown analysis. The blowdown phase of the transient ends when the RCS pressure (initially assumed at 2280 psia) falls to a value approaching that of the containment atmosphere. Prior to or at the end of the blowdown, the mechanisms that are responsible for the emergency core cooling water injected into the RCS bypassing the core are calculated not to be effective. At this time (called end-of-bypass) refill of the reactor vessel lower plenum begins. Refill is completed when emergency core cooling water has filled the lower plenum of the reactor vessel, which is bounded by the bottom of the fuel rods (called bottom of core recovery time). O 1343rlo/042SBB 15.4-4
di The reflood phase of the transient is defined as the time period lasting from O the end-of-refill until the reactor vessel has been filled with water to the
~
extent that the core temperature rise has been terminated. From the latter stage of blowdown and then the beginning-of-reflood, the safety injection accumulator tanks rapidly discharge borated cooling water into the RCS, contributing to the filling of the reactor vessel downcomer. The downcomer water elevation head provides the driving force required for the reflooding of the reactor core. The low head and high head safety injection pumps aid in the filling of the downcomer and subsequently supply water to maintain a full downcomer and complete the reflooding process. Continued operation of the ECCS pumps supplies water during longterm cooling. Core temperatures have been reduced to longterm steady state levels associated with dissipation of residual heat generation. Af ter the water level of the residual water storage tank (RWST) reaches a minimum allowable value, coolant for long-term cooling of the core is obtained by switching to the cold recirculation phase of operation, in which spilled borated water is drawn from the engineered safety features (ESF) containment sumps by the low head safety injection ,(residual heat removal) pumps and returned to the RCS cold legs. , ( The. containment spray system continues to operate to further reduce containment pressure. ' Approximately 11 hours after initiation of the LOCA, the ECCS is realigned to supply water to the RCS hot legs in order to control the horic acid concentration in the reactor vessel. 15.4.1.4 Core and System Performance 15.4.1.4.1 Mathematical Model The requirements of an acceptable ECCS evaluation model are presented in Appendix K of 10 CFR 50 (Federal Register 1974).II) O 1343r1D/0429ss 15.4-5
.)
15.4.1.4.2 Large Break LOCA Evaluation Model The analysis of a large break LOCA transient is divided into three phases: (1) blowdown, (2) refill, ard (3) reflood. Thero are three distinct transients analyzed in each phase, including the thermal-hydraulic transient in the RCS, the pressure and temperature transient within the containment, and the fuel and clad temperature transient of the hottest fuel rod in the core. Based on these considerations, a system of interrelated computer codes has been developed for the analysis of the LOCA. l A description of the various aspects of the LOCA analysis methodology is given by Bordelen, Massie, and Zordon (1974).(6) This document describes the major phenomena modeled, the interfaces among the computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria. 1 The SATAN-VI, WREFLOOD, BASH, LOCBART, and CDC0 codes, which are used in the LOCA analysis, are described in detail by Bordelen ej aj. (1974)(5); g,))y ej al. (1974)(9); Young el al. (1987)(4); Bordelen and Murphy (1974)(3) Code modifications are specified in References 2, 7 and 14. These codes assess the core heat transfer geometry and determinc.if the core remains amenable to cooling throughout and subsequent to the blowdown, refill, and reflood phases of the LOCA. The SATAN-VI computer code analyses the thermai-hydraulic transient in the RCS during blowdown and the WREFLOOD computer code calculates this transient during the refill phase of the accident. The BASH code is used to determine the system response during the reflood phase of the transient. The COCO code is used for the complete containment pressure history for dry containments. The LOCBART computer code calculates the thermal transient of the hottest fuoi rod during the three phases. The Revised Pad Fuel Thermal Safety Model, described in Reference 14, generates the initial fuel rod conditions input to LOCBART. SATAN-VI calculates the RCS pressure, enthalpy, density, and the mass and energy flow rates in the RCS, as well as steam generator energy transfer between the primary and secondary systems as a function of time during the blowdown phase of the LOCA. SATAN-VI also calculates the a:ct.mulater water mass and internal pressure and the pipe break mass and enargy flow rates that O' isorto/onsas 15.4-6
i are asrumed to be vented to the containment during blowdown. At the end of q the blowdown, information on the state of the system is transferred to the V -l WREFLOOD code which performs the calculation of the refill period to bottom of l core (BOC) recovery time. Once the vessel has refilled to the bottom of the l core, the reflood portion of the transient begins. The BASH code is used to ! calculate the thermal-hydraulic simulation of the'RCS for .the reflood phase. l l i Information concerning the core boundary conditions is taken from all of the ! above codes and input to the LOCBART code for the purpose of calculating the core fuel rod thermal response for the entire transient. From the boundary conditions, LOCBART computes the fluid conditions and heat transfer coefficient for the full length of the fuel red by employing mechanistic models appropriate to the actual flow and heat transfer regimes. Conservative assumptions ensure that the fuel rods modeled in the calcula, tion represent the hottest reds in the entire core. The containment pressure analysis is performed with the COCO code (3), which is interactive with the WREFLOOD code. The transient pressure computed by the COCO code is then input to the BASH code for the purpose of supplying a backpressure at the break plane while computing the reflood transient. The large break analysis was performed with the December 1981 version of the Evaluation Model modified to incorporate the BASHI4) computer code. 15.4.1.4.3 Input Parameters and Initial Conditions . a The analysis presented in this section was performed with a reactor vessel upper head temperature equal to the RCS cold leg temperature. The bases used to select the numerical values that are input parameters to the i analysis have been conservatively determined from extensive sensitivity studies (Westinghouse 1974(12); Salvatori 1974(11)). In addition, the requirements of Appendix K regarding specific model features were met by selecting models which provide a significant overall conservatism in the analysis. The assumptions which were made pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA ] iustiomassa 15.4-7 li
occurs, and include such items as the core peaking factors, the containment g pressure, and the performance of the ECCS. Decay heat generated throughout T-the transient is also conservatively calculated. 15.4.1.4.4 Results Based on the results of the LOCA sensitivity studies (Westinghouse 1974(12); Salvatori 1974(11)) the limiting large breai was found to be the double ended cold leg guillotine (DECLG). Therefore, only the DECLG break is considered in the large break ECCS performance analysis. Calculations were performed for a range of Hoody break discharge coefficients. The results of these calculations are summarized in Tables 15.4.1-1 and 15.4.1-2. The hot spot is defined to be the location of the maximum peak clad temperature. This location is giv n in Table 15.4.1-2 for each break size analyzed. Containment data used to calculated ECCS back pressure is presented in Table 15.4.1-3. Figures 15.4-1 through 15.4-67 show transient plots of important parameters from the ECCS Evaluation Model calculations. Plots are grouped by break size as follows: Figures 15.4-1 through 15.4-16 Cd =0.4 MIN SI Figures 15.4-17 through 15.4-32 Cd =0.6 HIN SI Figures 15.4-33 through 15.4-48 Cd =0.8 MIN SI Figures 15.4-49 through 15.4-64 Cd =0.4 MAX SI For each break size, a series of plots is presented, showing the transients of the following parameters. I. For the blowdown portion of the transient: A. RCS Pressure B. Core inlet and outlet flow rates C. Cold leg accumulator delivery rate D. Core pressure drop g is4sv:io/o42ssa 15.4-8 , I
1 E. Break mass flow rate O F. Break energy discharge rate G. Normalized core power , II. For the reflood portion of the transient ! A. Core and Downcomer liquid levels B. The core inlet fluid velocity, as input to the rod thermal analysis code C. The accumulator flow rates D. Pumped safety injection flow rates III. From the fuel rod thermal analysis, at the peak temperature location: A. Fluid mass flux B. Rod heat transfer coefficient C. Clad peak temperature transient D. Temperature transient at the hot rod burst elevation E. Fluid tempeature For the most limiting break size, the containment pressure transient and the containment wall condensing heat transfer coefficient are presented in Figures 15.4-65 and 66, and the safety injection flow rate is presented in Figures 15.4-67. In addition to the above, Tables 15.4.1-4 and 15.4.1-5 present the reflood mass and energy release to the containment and the broken loop accumulator , mass and energy flowrate to the containment, respectiveiy. The maximum clad temperature calculated for a large break is 2141'F which is less than the Acceptance Criteria limit of 2200'F. The maximum local metal-water reaction is 10.128 percent, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The total core metal-water reaction is less than 0.3 percent for all breaks, as compared with the 1 percent criterion of 10 CFR 50.46. The clad temperature transient is O m scioto42ssa 15.4-9
terminated at a time when the core geometry is still amenable to cooling. As a result, the core temperature will continue to drop and the ability to remove h. decay heat generated in the fuel for an extended period of time will be provided. e 9 l l l O k unaom2ni "'
q REFERENCES FOR SECTION 15.4.1 V
- 1. "Acceptance Criteria for Emergency Core Cooling System for Light Water Cooled Nuclear Power Reactors," 10 CFR 50.46 and Appendix X of 10 CFR 50, Federal Register 1974, Volume 39, Number 3.
- 2. Rahe, E. P. (Westinghouse), letter to J. R. Miller (USNRC), Letter No. -
~ NS-EPRS-2679, November 1982.
- 3. Bordelon, F. M., and Murphy, E. T., "Containment Pressure Analysis Code (C0CO)," WCAP-8327 (Proprietary) WCAP-8326 (Non-Proprietary), June, 1974.
- 4. Young, M. Y, et al, "The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code," WCAP-10266-P-A Rev. 2 (Proprietary),1987.
- 5. Bordelen, F. M. g a_1,., "SATAN-VI Program: Comprehensive Space, Time Dependent Analysis of loss-of-Coolant " WCAP-8302 (Proprietary) and WCAP-8306 (Non-Proprietary), 1974.
O 6. Bordelen, F. M.; Massie, H. W.; and Zordon, T. A., "Westinghou'se ECCS Evaluation Medal - Summary," WCAP-8339, 1974.
- 7. Rahe, E. P., "Westinghouse ECCS Evaluation Model, 1981 Version,"
WCAP-9220-P-A(ProprietaryVersion),WCAP-9221-P-A(Non-Proprietary ', Version), Revision 1, 1981. 1
- 9. Kelly, R. D. ej a_1,., "Calculational Model for Core Reflooding After a l
Loss-of-Coolant Accident (WREFLOOD Code," WCAP-8170 (Proprietary) and WCAP-8171 (Non-Proprietary), 1974. l i l0 1:43v:1o/0429ae 15.4-11
- 10. V. S. Nuclear Regulatory Comission 1975, "Reactor Safety Study - An Assessment of Accident Risks in U. S. Commercial Nuclear Power Plant,"
WAfH-1400, NUREG-75/014.
- 11. Salvatori, R., "Westinghouse ECCS - Plant Sensitivity Studies," WCAP-8340 (Proprietary) and WCAP-8356 (Non-Proprietary), 1974. .
- 12. "Westinghouse ECCS Evaluation Model Sensitivity Studies," WCAP-8341 (Proprietary) and WCAP-8342 (Non-Proprietary), 1974.
- 13. "Bordelen, F. M., et al., "Westinghouse ECCS Evaluation Model -
Supplementary Information," WCAP-8471 (Proprietary) and WCAP-8472 (Non-Proprietary),1975.
- 14. Letter from J. F. Stoltz (NRC) to T. H. Anderson (Westinghouse),
Subject:
Review of WCAP-87?.0, Improved Anslytical Hodels used in Westinghouse Fuel Rod Cesign Computations. O O 1343v:10/0429ea 15.4-12
!l TABLE 15.4.1-1 O
, LARGE BREAK l I TIME SEQUENCE OF EVENTS l _ i DECLG (C D =0.8) DECLG (C D =0.6) DECLG (CD =0.4) (Sec) (Sec) (Sec) START 0.0 0.0 0.0 Reactor Trip Signal 0.433 0.440 0.452 S. I. Signal 0.62 0.70 0.85 Acc. Injection 8.96 11.25 15.10 End of Blowdown / Bypass 19.95 23.09 30.41 Bottom of Core Recovery 32.31 35.41 43.13 Pump Injection 32.62 32.70 32.85 Acc. Empty 42.269 45.234 50.527 0 i O 1343c1D/o42988 15.4-13
- . - . . - . _ _ . .n. - . . . - :=-..-..=.-....- - ...->
TABLE 15.4.1-2 LARGE BREAK DECLG (C D =0.8) DECLG(pD=0'.6) DECLG (C0 =0.4) Results Peak Clad Temp. 'F 1767. 1877. 2141. Peak Clad Location Ft, 7.0 8.0 7.0 Local Zr/H 2O Reaction 2.08 2.63 10.13 (max) % Local Zr/H2 0 Location Ft. 6.25 8.0 6.0 , Total Zr/H O Reaction % <0.3 <0.3 <0.3 2 Hot Rod Burst Time sec 43.37 41.46 40.2 Hot Rod Burst Location Ft. 6.25 6.0 6.0 Calculation NSSS Power Hwt 102% of 2775 h Peak Linear Power !:w/f t 102% of 13.340 Peaking Factor (At License Rating) 2.45 - n 3 Accumulator Water Volume (ft ) 1039 (minimum , 1;s line volume) Fuel regien + cycle analyzed Cycle Region UNIT 1 5 and Beyond ALL 15% Steam Generator Tube Plugging in each steam generator is assumed. O 1343<:1o/o42ssa 15.4-14
TABLE 15.4.1-3
\
4 LARGE BREAK CONTAINMENT DATA (DRYCONTAINMENT) 6 3 NET FREE VOLUME 1.9 x 10 ft ! INITIAL CONDITIONS Pressure 14.7 psia Temperature 90*F RWST Temperature 40*F Servica Water Temperature NA , Outside Temperature 19'F SPRAY SYSTEM Total Flow Rate 6000 gpm Actuation Time 52 secs i O SAFEGUARDS FAN COOLERS Number of Fan Coolers Operating 2 Fastests Post Accicent Initiation of Fan Coolers 43 secs l O 1343<:io/o42ses 15.4-15 1
TABLE 15.4.1-3 (Continued) g. LARGE BREAK CONTAINMENT DATA
~
(DRY CONTAINMENT) STRUCTURAL HEAT SINKS 2 Thickness (In) Area (Ft ) 0.348 Carbon Steel 48.0 Concrete 57,397 0.264 Carbon Steol 36.0 Concrete 20,241 0.125 Carbon Steel 24.0 Concrete 11,694 18.0 Concrete 315 22.56 Concrete 43,537 , 12.0 Concreta 10,811 48.0 Concrete 19.020 1.52 Stainless Steel 409 1.13 Stainless Steel 551
~0.6 Stainless Steel 1939 0.336 Stainless Steel 2194 g 0.06 Stainless Steel 88481 6.672 Carbon Steel 3300 3.504 Carbon Steel 130 2.376 Carbon Steel 2324 1.7568 Carbon Steel 4323 0.87 Carbon Steel 8787 0.744 Carbon Steel 17734 0.324 Carbon Steel 16929 0.06 Carbon Steel 654508 l
l 1 1 0 1343v:1o/o42ssa 15.4-16
TABLE 15.4.1-4 REFLOOD MASS / ENERGY RELEASES * (CD = 0.4) TOTAL MASS FLOWRATE TOTAL ENERGY FLOWRATE 5 TIME (SEC) (LBM/SEC) 10 BTU /SEC 43.00 0.0 0.0' : i 49.00 38.65 0.503 59.00 87.75 1.093 79.0 179.26 1.392 l 109.0 265.1 1.537 129.0 27*.31 1.491 169.0 306.53 1.468 1 1 3 i ! l
- Accumulator nitrogen was released between 50.0 and 70.0 seconds at a mass.
{ flow rate of 198.67 lbm/sec. l l i i l i l tusv:1o/o42ss 15.4-17 i; i .
TABLE 15.4.1-5 , BROKEN LOOP ACCUHULATOR FLONRATE TO CONTAINMENT FOR !- L1HITING CASE - DECLG (CD = 0.4) TIME (SEC) MASS FLONRATE* (LBM/SEC) 0.0 4763.3 3.01 3596.8 4.01 3369.0 5.01 3179.8 8.01 2752.3 11.01 2449.3 2 16.01 2093.6 21.01 1847.4 24.01 1734.0 25.55 0.0 O
*Enthalpy of accumulator water is 58 BTU /LBM b
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I ATTACHMENT 5 ! i< THIMBLE PLUG REMOVAL I EVALUATION : FOR THE V. C. SUMMER PLANT l TRANSITION TO 17x17 VANTAGE 5 FUEL O l I O i
- - . ~ - . . _ _ _ , _ _
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V THIHBLE PLUG REMOVAL EVALUATION 1.0 IliTRODUCTION AND
SUMMARY
Coincident with the fuel transition to the first region of-VANTAGE 5, it is ~ planned to remove thimble plugging devices from the core. This includes the , i removal of thimble plugs from the VANTAGE 5 assemblies, LOPAR assemblies, and 1 all new core component clusters (burnable absorbers and sources). Thimble plugging devices are currently utilized in the V. C. Summer reactors f to limit the core bypass flow. All guide thimble tubes that are not under RCC j locations or are not equipped with sources and burnable absorbers currently have thimble plugs inserted in them. A net gain of around 27. In DNBR margin is realized due to their presence. Westinghouse has evaluated the effect of thimble plug removal and has concluded that it is feasible to remove all or any combination of these devices from the V. C. Summer core. This evaluation is described in the following sections.and addresses the effect of thimble plug removal oni,y unless explicitly stated otherwise. , 2.0 THERMAL HYDRAULIC DESIGN EVALUATION 2.1 Bypass Flow The main impact of thimble plug removal is the Increase in core bypass flow. Calculations performed by Hestinghouse have shown that the design value of core bypass flow needs to increase from 6.47. to 8.97. (non-ITDP). This increase is due to the combined effect of thimble plug removal and the slightly higher pressure drop of the VANTAGE 5 fuel assembly. The safety analysis has been performed assuming that thimble plugs have been removed from the core, l 1 l
. ~
l 2.2 Primary System Flow Rate Thimble plugs removal also results in a reduction to,the fuel assembly hydraulic loss coefficient. Westinghouse has performed tests to quantify the magnitude of this effect. Based on these tests, it is est-imated that there will be a slight increase in primary system flow rate due to thimble plug removal from the V. C. Summer core. No mechanical design criteria are impacted by this slight increase in flow rate. 2.3 Fuel Assembly Hydraulic Lift The hydraulic lift force on the fuel assembly can be represented by the following function: FLIFT " KFA x(coreflow/ Westinghouse has performed hydraulic tests to quantify the magnitude of the g effect of thitnble plug removal on fuel assembly hydraulic loss coefficient W (K FA). The results show that there is a net reduction in F LIFT due to a reduced fuel assembly loss coefficient (caused by thimble plug removal) which more than compensates for the slight increase in vessel flow rate. Thimble plug removal is therefore acceptable from a fuel assembly lift standpoint, t 2.4 Effect of Outlet Hydraulic Hismatch on DNB Current DNB analyses are performed assuming the presence of an uniform static pressure distribution at the core outlet, even though pressure gradients and core outlet loss coefficient mismatches are known to exist. This is acceptable because these mismatch effects do not propagate upstream into the DNS zone. Westinghouse has performed numerous sensitivity studies to demonstrate the insensitivity of as calculated DNBR's to non-uniform outlet pressure distributions and to variations in outlet loss coefficients. The effect of thimble plug removal on the corewide distribution of outlet loss coefficients for the V. C. Summer cores has been evaluated. It was demonstrated that the variations in outlet loss coefficient due to thimble
/7 l O plug removal are within the bounds of the sensitivity studies that had been ;
, performed. Therefore, it is concluded that thimble plug removal will not result in the reduction of DNBR margin due to mismate.hes in core outlet pressure gradients and loss coefficients. j-i 3.0 MECHANICAL DESIGN EVALUATION l 3.1 Fuel Rod Fretting Hear , The removal of thimble plugging devices changes the distribution of core outlet loss coefficients. The core outlet loss coefficient (PFO) distribution shows an increase in PF0 mismatch after thimble plug removal. Therefore, the issue of crossflow induced fuel rod vibration and wear due to this increased PF0 mismatch is addressed. The maximum PF0 mismatch that exists in the V. C. Summer core after removal of , all or any combination of thimble plugs is isss than 1.5. Westinghouse, O V however, has recently performed fuel rod vibration tests with a PF0 mismatch I-of approximately 17 between two 17x17 fuel assemblies. The results showed that there was no significant difference in fuel rod response between the tests performed with and without this large PF0 mismatch. Therefore, it is concluded that thimble plug removal will not have a detrimental effect on fuel rod vibration and wear. 3.2 Control Rod Hear Westinghouse studies on control rod wear have shown that most of the wear tends to be in the upper internals region. When thimble plugs are removed the hydraulic resistance at the outlet for these assemblies is reduced. This in turn causes the flow through the RCCA guide tubes to be reduced, because more flow is now going through the outlet of the assemblies which were previously fitted with thimble plugs. This reduction of flow through the RCCA guide tubes is in the direction that would tend to reduce control rod wear. O
However, since the core PF0 distribution changes when thimble plugs are removed, the effect of potential control rod vibration due to inter assembly crossflows in the region of the control rod / fuel assembly guide thimble interface needs to be addressed. The control rods.can be directly affected in the core region only by inter assembly crossflows through the gap (-0.75") between the top nozzle and upper core plate. For the V. C. Summer reactor upper internals configuration, it was concluded that the maximum PF0 mismatch between an RCC location and an adjacent assembly does not increase with thimble plug removal. Therefore, the magnitude of the crossflow seen by the control rods and the vibration of the rods caused by this crossflow will not be increased. Based on the above evaluation, thimble plug removal will not have an adverse impact on control rod wear for the V. C. Summer reactor. 3.3 Seismic /LOCA Transient Loading The thimble plugging device (approximately 11 lbs.) contributes a very small percentage of the total fuel assembly weight. Therefore, the removal of these devices will have a negligible effect on the fuel structural responses to seismic or LOCA transient loading. 3.4 Reactor Internals Structural Adequacy There is a negligible impact of thimble plug removal on the internals time history analysis. Also, the effect of thimble plug removal on increased RCS flow will have a negligible effect on the structural adequacy of the internals. Thimble plug removal is therefore acceptable from a Reactor Internals standpoint. 4.0 SAFETY ANALYSIS The non-LOCA analyses (Attachrent 3) and LOCA analysis (Attachment 4) have conservatively assumed that all thimble plugs were removed from the core to produce the maximum core bypass flow and core consequences.
5.0 CONCLUSION
Detailed evaluations have shown that the main effect of thimble plug removal is the increase in core bypass flow. This increase has been incorporated into : the non-LOCA and LOCA safety analyses that have been performed in support of
. the VANTAGE 5/LOPAR fuel transition cores i:
Based on the assessment of the impact of the thimble plug removal on system and component structural adequacy and core plant' safety,.It is concluded that it is acceptable to remove all or any combination of these devices from the V. C. Summer core (s). The evaluation also bounds the use of any combination
. of dually compatible thimble plugs, HABAs and source rods.
J l I 1 I l-l l 'O L c l _ _ _ _ _ _. o r
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- I ATTACHMENT 6
~
~
1 i :( t SIGNIFICANT HAZARDS EVALUATION i , I~ FOR THE V. C. SUMMER PLANT _ i I TRANSITION TO HESTINGHOUSE 17x17 VANTAGE 5 FUEL ASSEMBLIES \ t f i l 4 l
- O t
l i 4 } 1 O i 00681:6/880516 i
2 g) (. SIGNIFICANT HAZARDS EVALVATION FOR VIRGIL C. SUMMER NUCLEAR STATION TRANSITION TO WESTINGHOUSE VANTAGE 5 FUEL Description of amendment request:
- i South Carolina Electric & Gas Company (SCE&G) requests an amendment to the Virgil C. Summer Nuclear Station (VCSNS) Technical Specifications to support (a) the Cycle 5 core reload to permit operatha with Westinghouse VANTAGE 5 (V-5) fuel assemblies in addition to the Westinghouse Low Parasitic (LOPAR) assemblies remaining in the core from Cycle 4 and (b) subsequent operating cycles with up to a full core of V-5. Design features of the V-5 fuel include assemblies with up to approximately 4.25 weight percent U-235, axial blankets, integral fuel burnable absorbers, intermediate flow mixers, reconstitutable top nozzles, and extended burnup capability. A debris filter bottom nozzle (DFBN) will be introduced to replace the standard V-5 bottom nozzle to reduce the possibility of fuel rod damage due to debris . induced fretting. This p requires changes to the Technical Specifications due to the use ,of the V-5
'd fuel and use of the following analytical methods and assumptions. o The Improved Thermal Design Procedure (ITDP) l o The WRB-1 and WRB-2 departure from nucleate boiling (DNB) correlations o The BASH large break loss-of-coolant accident (LOCA) model o The NOTRUMP small break LOCA model o The ANSI /ANS 5.1-1979 decay heat model for non-LOCA accidents and transients o Reir.xed Axial Offset Control (RAOC) , l-o Fg (z) Surveillance o Analysis baseline changes as outlined in Table 1 As a result of the above, changes to the following Technical Specifications and corresponding bases, as appropriate, are proposed: o Core Safety Limits o Reactor Coolant Flow Allowable Values J 1412v:1o/051788-1
._ _ . ~ - .-,
o Overtemperature delta T Reactor Trip Setpoints o Overpower delta T Reactor Trip Setpoints o Shutdown Margin for Modes 3, 4, and 5 o Moderator Temperature Coefficient o Rod Drop Time o Axial Flux Difference - o Heat Flux Hot Channel Factor - F0 (*) o Nuclear Enthalpy Rise Hot Channel Factor - F aH o DNS Parameters o Reactor Trip (s) Response Time o ECCS Accumul'ator water volume range o Borated Water Sources for Modes 1-4 o Reactor Trips and Emergency Safety Features Actuation System Drift Allowances for Determination of Operability o Charging Pump Flow Balance Surveillance Basis for proposed no significant hazards consideration determination: SCE&G has evaluated the proposed changes associated with the transition to V-5 fuel against the Significant Hazard *, Criteria of 10CFR50.92 and against the Commission guidance concerning application of this standard. VCSNS's proposed license amendment is closely related to an example (51 FR 7751) of action not likely to involve a significant hazard. Specifically, example (iii) of the guidance states: "For a nuclear power reactor, a change resulting from a nuclear reactor core reloading, if no fuel assemblies significantly different from ,those found previously acceptable to the NRC for a previous core at the facility in question are involved. This assumes.that no significant changes are made to the acceptance criteria for the Technical Specifications, that the analytical methods used to demonstrate conformance with Technical Specifications and regulations are not significantly changed, and the NRC has previously found such methods acceptable." 9 1412v:1o/Cs1788-2
The VCSNS proposed licensing amendment is directly related to the above , example in that the core reload uses V-5 fuel which is not significantly different from previous cores at VCSNS, the changes to the Technical Specifications are as a result of the core reload and not because of any significant change made to the acceptance criteria for Technical j Specifications, and the analytical methods used in the reqaired reload + analysis have been previously found acceptable by the NRC. Therefore, based on the above, SCE&G concludes that the proposed Technical Specifications , changes do not involve a significant hazard consideration. SCE&G has evaluated the proposed changes in design, analytical methodologies and Technical Specifications associated with the transition to V-5 fuel against the Significant Hazards Criteria of 10CFR50.92. The results of SCE&G's evaluations demonstrate that the changes do not involve any significant hazard as described below,
- a. The probability or consequences of an accident previously evaluated is not significantly increased. l The V-5 reload fuel assemblies are mechanically and hydraulically compatible with the current LOPAR fuel assemblies, control rods and reactor internals interfaces. Also, implementation of V-5 does not cause a significant change in the physics characteristics of the VCSNS cores beyond the normal range ~cf variation seen from cycle to cycle. Thus, both fuel types satisfy the design basis for VCSNS as proposed for this amendment.
Thimble plug removal has a negligible impact on the system and component structural adequacy but does cause core bypass flow to increase. The , revised core thermal-hydraulic design and safety analysis, however, show that the DNB penalty due to removal of the thirble plugs is more than offset by the increase in DNS margin resulting from the use of the ITOP and V-5 fuel. O 1412v;1o/051788-3 i
The proposed changes have been assessed from a core design and safety analysis standpoint. No increase in the prebability of occurrence of any accident was identified but an extensive reanalysis, as described in the Transition Safety Evaluation, was required to demonstrate compliance to the revised VCSNS Technical Specifications as proposed herein. These reanalyses app' tied methods which have been previously found acceptable by the NRC. Tne results, which include transition core effects, show changes in consequences of accidents previously analyzed. However, the results are all clearly within pertinent acceptance criteria and demonstrate the plants capability to opeirate safely at 100% power. Thus, it is concluded that there is not significant increase in the consequences of an accident previously evaluated.
- b. The possibility for an accident or malfunction of a different type than any evaluated previously in the safety analysis reports is not created.
These proposed changes do not significantly effect the overall method and manner of VCSNS operation and can be accommodated without compromising the performance or qualification of safety-related equipment. Thus, the creation of a new or different kind of accident from any previously evaluated accidelt is not considered a possibility,
- c. The margins of safety as defined in the bases of the Technical Specifications is not significantly reduced.
The evaluations and analyses described herein show some changes in the consequences of previcusly analyzed accidents. In some cases, an inerease in event consequences cccurs and may reduce margin. However, in all cases, the results of the changes are clearly within all pertinent design and safety acceptance criteria. Thus, there is no significant reduction in the margin of safety as a result of the proposed changes. O 1412v1o/051788-4
Table 1 VCSNS ANALYSIS BASEllHE f. 1 Proposed Value For Parameter Current Value ' Vantage 5 Transition NSS Power. HWt 2785 2787 Core Power Mdt 2775 2775 System Pressure. psia 225G 2250 Thermal Design Flow, gpm/ loop . ; ,0 92600*
. - _ . =
(. Core Bypass Flow, % 0.4 8.9'* TAVE. 'F 587.4 585.5 TH OT. 'F 618.7 618.7 F2H 1.55 1.62 F2H Multipier 0.2 0.3 LOCA FQ 2.25 2.45 . SG Tube Plugging, % , 16 15 AFO Control CAOC RAOC l. l Peaking Surveillance Fxy (z) FQ (2) , I' High Head Safety injection Recirculation Recirculation Isolated Hot Isolated Thimble Plugs Yes Optional
- Includes 2% additional flow margin for conservatism ,
** Non-ITOP 1'
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ATTACHMENT 7 ; RADIOLOGICAL IMPACT ASSESSMENT
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- i FOR THE V. C. SUMER PL5NT- - +
t TRANSITION TO HESTINGHOUSE 17x17 VANTAGE 5 FUEL ASSEMBLIES t J
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t O v Radioloaical Imoact Assessment The use of VANTAGE 5 fuel in the Virgil C. Summer Nuclear Station will result in a higher discharge region average burnup. While fission product inventories are. roughly proportional to operating power level, the level of fuel burnup has little impact except for isotopes with long half-lives. This is supported by the Westinghouse topical report, NCAP-10125-P-A (Proprietary), titled "Extended Burnup Evaluation of Westinghouse Fuel," which demonstrates that extension of fuel burnup to the higher discharge region average burnups evaluated in the topical report, which are bounding for V. C. Summer, would
'have only a small impact on the core fission product inventories. This change ! in fission product inventories would not significantly affect the radiological consequences of postulated accidents.
For the fuel handling accident, extending the discharge region average burnup to the maximum value esaluated in HCAP-10125-P-A would result in an increase of approximately four percent in the thyroid dose. This increase is based on ' continued use of the fuel handling accident analysis assumption, defined in Regulatory Guide 1.25, that ten percent of the core inventory of short-lived isotopes and thirty percent of core inventory of long-lived isotopoes are in ; the fuel rod gap. The short-lived isotopes are of greatest concern in regard 1 to radiological consequences of the accident, and analysis shows that the fraction of short-lived isotopes in the fuel rod gap, when at it's maximum, would be about one percent or less of the core inventory; not the value of ten percent assumed in Regulatory Guide 1.25. Extending the fuel burnup actually will result in a reduction in the gap inventories of short-lived isotopes due l to operation at lower power levels for the latter part of the residence time i l in the core which results in a concomitant reduced production rate for fission products. Also, with operating at a reduced power level, the fuel pellet temperature will be reduced resulting in a lower rate of diffusion of fission products into the rod gap. i f The radiological consequences of the fuel handling accident are also impacted l by the fact that the proposed fuel design has a F delta H (radial peaking l actor) of 1.68 specified. The fuel handling accident for V. C. Summer ! .00681:6/880516 1
i utilizes the guidance of Regulatory Guide 1.25 which specifies that a minimum 4I radial peaking factor of 1.65 be used in determining the maximum fuel assembly ! fission product inventory. Hith a radial peaking factor of 1.63, the deses ' reported in the FSAR would be increased by approximately two percent, ibe increase in calculated dose due to the combination of extended burnup and - increased radial peaking factor is not significant and the doses are still well within the acceptance criteria defined in the Standard Review Plan. The radiological consequences of accidents othe.r than the fuel handling accident are also impacted to a slight degree. As discussed in HCAP-10125-P-A, the impact of extended fuel burnup on the consequences of a rod ejection accident would be to slightly increase the thyroid dose (about two percent) and to decrease the whole body dose. This same effect on radiological consequences would also be seen in other accidents involving release of reactor coolant activity whether or not there is any fuel damage as a result of the accident. These increases in radiological consequences are insignificant, being within the uncertainty of the calculational assumptions. Thus, there is no need to recalculate the radiological consequences of the accidents due to extending the fuel burnup within the limits of the study reported in HCAP-10125-P-A. In addition, it is noted that the radiological consequences of accidents reported in the FSAR are well within the limits of 10 CFR 100; thus, if increases such as those discussed above were applied to the FSAR doses, there would be no impact on their acceptability. Other than the increased level of burnup and the increase in radial peaking factor discussed above, no features of the VANTAGE 5 fuel design have an impact on the radiological consequences of normal operation or of the postulated accidents. f 6 l , 00681:6/880516 2 l}}