ML20195C300
| ML20195C300 | |
| Person / Time | |
|---|---|
| Site: | Point Beach  |
| Issue date: | 10/31/1988 |
| From: | WISCONSIN ELECTRIC POWER CO. |
| To: | |
| Shared Package | |
| ML20195C298 | List: |
| References | |
| NUDOCS 8811020435 | |
| Download: ML20195C300 (98) | |
Text
. e FINAL REPORT FOR INCREASED PEAKING FACTORS AND FUEL UPGRADE ANALYSIS Point Beach Nuclear Plant Units 1 and 2 Wisconsin Electric Power Company Dockets 50-266 and 50-301 October 1988 i
- 8P58ss 8?s8ljy P
. o TABLE OF CONTENTS Section Title Page Number
1.0 INTRODUCTION
1-1 2.0 DESIGN FEATURES 2-1 2.1 Introduction 2-1 2.2 Upgraded Fuel Product Features 2-2 2.3 PLEX Fuel Management Features 2-4 2.4 Thimble Plugging Devices 2-6 3.0 NUCLEAR DESIGN 3-1 3.1 Introduction 3-1 3.2 Methodology 3-2 3.3 Design Evaluation - Physics Characteristics 3-3 and Key Safety Parameters 3.4 Design Evaluation - Power Distributions 3-3 and Peaking Factors 3.5 Conclusion 3-6 4.0 THERMAL-HYDRAULIC DESIGN 4-1 4.1 Introduction 4-1 4.2 Calculational Methods 4-1 4.3 Hydraulic Compatibility - Transition Core 4-2 4.4 Effects of Fuel Rod Bow on DNBR 4-3 4.5 DNBR E ' -' +5e Upgraded Fuel 4-3 4.6 Fuel Tem r .. , for Safety Analysis 4-4 4.7 Thimble Plug Removal 4-4 4.8 Conclusion 4-8 5.0 FUEL R00 DESIGN d-1 5.1 Introduction 5-1 5.2 Methodology and Input Assumptions 5-1 5.3 Fuel Rod Design Criteria Evaluation Results 5-5 5.4 Conclusion 5-5 6.0 REACTOR PRESSURE VESSEL SYSTEM EVALVATIONS 6-1 6.1 Introduction 6-1 6.2 Internal Pressure Losses 6-2 6.3 Core Bypass Flow 6-3 6.4 Homentum Flux 6-5 6.5 Closure Head fluid Temperature 6-5 6.6 Hydraulic Lift Forces 6-6 6.7 Rod Control Cluster Assembly Drop Times 6-6 6.8 Conclusions 6-7 i
TABLE OF CONTENTS (continued)
Section Title Page Number 7.0 ACCIDENT ANALYSIS 7-1 7.1 Non-LOCA Accidents 7-1
- 7. 2 Small-Break LOCA Analysis 7-11 7.3 Steam Generator Tube Rupture 7-15 Accident 11
e .
LIST OF TABLES Tab Title Page Number
'. 1 Key Safety Parameters 3-8 3.2 Representative Nuclear Design 3-10 Parameters 4.1 Point Beach Units 1 and 2 Thermal 4-10 and Hydraulic Design Parameters 4.2 DNBR Margin Summary 4-14 7.1 Summary of Non-LOCA Events 7-23 7.2 Non-LOCA Safety Analysis Assumptions 7-24 7.3 Input Assumptions Used in the Small 7-26 Break LOCA Analysis 7.4 Point Beach SGTR Results 7-27 111
LIST OF FIGURES Figure Title Boe Number 2-1 Removable Top Nozzle Joint 2-7 Illustration 2-2 Typical Axial Zoning in an IFBA Fuel 2-8
' Rod and Loading of IFBA Fuel Rods in a Fuel Assembly 2-3 Debris Filter Bottom Nozzle vs Current 2-9 Bottom Nozzle Flow Hole Illustration 2-4 Core Locations for Peripheral Power 2-10 Suppression Assemblies
- 3-1 Upgraded Core Assembly Average Burnup 3-11 2
Distribution and Absorber Placement at BOL (O MWD /MTU) 4
?' Upgraded Core Normalized Power Distribution 3-12
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! 3-3 Upgraded Core Normalized Power Distribution 3-13 i
at 2000 MWD /MTU, Unrodded, HFP, Equilibrium i X6aon, Peak F-Delta-H = 1.574 i
I 3-4 Upgraded Core Normalized Power Distribution 3 14
l 3-5 Upgraded Core Normalized Power Distribution 3-15
- at 10500 MWD /MTU, Unrodded, HFP, Equilibrium Xenon, Peak F-Delta H = 1.503 1 3-6 Maximum F T* P versus Axial Core Heicht 3-16 OuringNobalObeOperation 3-7 Normalized F Limit Versus Core Height 3-17
] q 3-8 Flux Difference Operating Envelope for 3-18 Upgraded Core 3-9 Upgraded Core Proposed Rod Insertion Limits 3-19 6-1 Thimble Plugging Device 6-8 6-2 Schematic View of Typical Reactor Vessel 6-9 1
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LIST OF FIGURES (continued)
Figure Title Page Number 7-1 Overtemperature and Overpower Delta-T 7-28 Setpoint Equations 7-2 Small Break LOCA Axial Power Shape 7-29 7-3 High Head SI Flow Versus RCS Pressure 7-30 v
1.0 INTRODUCTION
Wisconsin Electric Power Company has contracted with Westinghouse Electric Corporation to perform analyses and evaluations to support a license amendment for the Point Beach Nuclear Plant, Units 1 and 2, which will accommodate a
, number of proposed changes to the units, hereafter referred to as "upgraded core features." The term "upgraded core features" is used to mean a composite of the eight (8) design features listed in Section 2.1 of this report and the ;
increase of the Technical Specifications allowable core power peaking factors (Fq and F3g). A subset of the upgraded core features will be referred to as "upgraded fuel product features." These five (5) features are identified in Section 2.1 and discussed in detail in Section 2.2.
Increase of the peaking factors will allow implementation of a Low-Low Leakage Loading Pattern (L4P) fuel management scheme, which will result in a reduction in neutron fluence to the reactor vessel. Fluence reduction is part of the Plant Life Extension (PLEX) program, which will enhance the ability to extend the useful life cf the Point Beach reactor vessels. To accomp1'sh maximum fluence reduction, an increase in the values for Fqand F AH is required.
As a result of the plan to refuel and operate the Point Beach Nuclear Plant, Units 1 and 2 with Westinghouse upgraded core features, future core loadings will have fuel assemblies consisting of Removable Top Nozzles and Debris Filter Bottom Nozzles, with the capability of achieving extended discharge burnups and higher peaking factors than previous reload designs. Future reload fuel may also have axial blankets and Integral fuel Burnable Absorbers. The reactor 4 core can consist of Optimized fuel Assemblies (OFA), upgraded fuel product assemblies, and previously-depleted Low-Parisitic (LOPAR) Assemblies also i known as standard (STD) fuel assemblies. Since the previously-depleted STD fuel will be at a lower power than the fresh 0FA fuel, it will be bounded by the OFA. The use of previously-depleted STD fuel assemblies will be justified
! by cycle-specific reload analysis.
t l'1
e .
Some of the upgraded fuel product features are also VANTAGE 5 design features, generically approved by the NRC via their review of WCAP-10444-P-A (Proprietary) and WCAP-10445-A (Non-Proprietary) (reference 2 3). A brief summary of the VANTAGE 5 design features and major advantages of the upgraded fuel design are given in section 2.0 of this report.
Following the design features section (Section 2.0) are the sections which describe the analyses that were performed to support implementation of the upgraded core features. Removai of thimble plugging devices from the fuel assemblies and increase of the peaking ', actors required re-analysis of a number of design basis accidents described in Chapter 14 of the Point Beach Final Safety Analysis Report (FSAR). Results of the re-analysis of the affected accidents are described in section 7.0 of this report. In some cases, it will be seen that analytical assumptions were made which will allow for future design or operational flexibility without requiring comprehensive re-analysis.
This report is intended to serve as a base reference evaluation / analysis report for the transition from the present Point Beach Units 1 and 2 cores to cores containing the upgraded core features.
The analyses were performed at a core thermal power level of 1518.5 megawatt thermal (MWt) for 2000 psia and 2250 psia operation, with the following conservative assumptions made in the evaluations: a nuclear entnalpy rise hot channel factor (Fj ) of 1.70, an increase ir, the total core peaking factor (Fq ) to 2.50, removal of the third line segment of the K(Z) curve, and an increase in the uniform steam generator tube plugging level to 13 and 14 1-2
percent for Units 1 and 2, respectively (Refer to the note on page 1-4). In j addition, the small-break LOCA analysis supports a constant kw/f t limit, i.e. ,
elimination of an elevation-dependentqF limit, i
4 'Following is a comparison of the current and proposed design parameters for !
Point Beach. Point Beach currently operates with the Westinghouse STO and 0FA fuel designs. The current parameters are. .
i i
Core power of 1518.5 W t i Average linear power density of 5.7 kw/ft System pressure 2000 (or 2250) psia r i Core inlet temperature of 545.0*F (or 545.3'F) 4 Enthalpy rise hot channel peaking factor limit of 1.58 (FAH) !
\ ital peaking factor limit of 2.21 (F g)
- Reactor Coolant System (RCS) Thermal Design Flow of 178,000 gpm.
I The proposed design parameters for the core upgrade are: ;
! Cor<. power of 1518.5 Wt l
- Average linear power density of 5.7 kw/ft t
- System pressure 2000 (or 2?50) psia ,
Core inlet temperature of $45.0 F (or 544.8'F) [,
l Enthalpy rise hot channel peaking factor liheit of 1.70 (FAH) l Total peaking factor limit of 2.50 (Fg ) l 1 RCS Thermal Design Flow of 178,000 gpe. l
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l The analyses comprising this report utilize the standard reload design methods j j described in reference 3-1, and this report will be used as a base reference I j document in support of future Point Beach Reload Safety Evaluations (RSEs) for i f upgraded fuel reloads. Sections 3.0 through 7.0 of this report summarize the l analyses performed for nuclear design, thermal-hydraulic design, and fuel rod j
! design by the Westinghouse Commercial Nuclear Fuel Division, and the reactor {
l pressure vessel system evaluations and the safety analyses performed by the !
! Westinghouse Power Systems Division.
l i 1-3
e e Consistent with the Westinghouse standard relcad methodology, parameters are chosen to maximize the applicability of the analyses / evaluations for future cycles. The objective of subsequent cycle-specific RSEs will be to verify that applicable safety limits are satisfied, based upon the reference evaludtion/ analyses established in this report.
This report will demonstrate the following:
- 1. Removal of thimble plugging devices from the Westinghouse fuel assemblies (containing STO, OFA, or upgraded fuel) for the Point Beach Nuclear Pla.'t, Units 1 and 2, will satisfy the new design bases and safety limits as documented in this report. Operation with thimble plugs installed has also been bounded by the analyses.
- 2. Changes in the thermal-hydraulic and core design characteristics, due to the transition to upgrrded fuel ptoduct features, will be wit'in the range normally seen from cycle to cycle, due to fuel management effects. The change from the current fuel to the upgraded fuel will not cause changes to the current nuclear design bases.
- 3. The core design, fuel rod design, and safety analyses results documented in this report show the core's capability for operating safely at the current rated Point Beach Units 1 and 2 design thermal power with an FaH I 1.70, ar F qof 2.50, the current Thermal Design Flow of 89,000 gpm/ loop *,
and any combination of p mposed upgraded core features.
- Note: The non-LOCA analyses were performed for uniform tube plugging levels of 13% and 14%, respectively, for Units 1 and 2, at the current Thermal Design Flow. The LOU, and steam generator tube rupture (SGTR) analyses were conservatively ierformed to bound up to a 25% uniform tube plugging level and the associsted reduction in Thermal Design Flow.
1-4
- 4. When the upgraded core features are impleastnted in the Point Beach units, the previously reviewed and licensed safety analysis conclusions remain valid. Plant operating limitations given in the Technical Specifications will be satisfied with the proposed changes requested in License Amendment Request No. 127. This report establishes a reference upon which to base Westinghouse reload safety evaluations for future reloads with the upgraded core features.
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- 2. 0 DESIGN FEATURES 2.1 Introduction The Point Beach units are presently operating with Westinghouse standard (STD or LOPAR) assembly and Optimized Fuel Assembly (OFA) fuel designs in the core. The analyses discussed in the following sections were performed to support future operation of the units with any of the following Westinghouse VANTAGE 5 fuel modifications incorporated into the Point Beach fuel assembly design as part of an upgraded fuel product features package:
Removable Top Nozzles Integral Fuel Burnable Absorbers Axial Blankets Extended Burnup Geometry The upgraded fuel product features will also include a Debris Filter Bottom Nozzle.
Each of these features is added to the Westinghouse 14x14 0FA to produce an upgraded fuel product. The Westinghouse 14x14 0FA was reviewed and generically approved by the NRC via reference 2-1, and specifically approved by the NRC for Point Beach via reference 2-2. As indicated previously, the VANTAGE 5 fuel features have been generically reviewed and approved by the NRC via reference 2-3.
Also planned for Point Beach Units 1 and 2 is operation incorporating the following:
Low-Low-Neutron-Leakage Loading Patterns (L4P)
Peripheral Power Suppression Assemblies or PPSAs (Part-length hafnium absorbers)
Removal of fuel assembiy thimble plugging devices i
All of these upgraded fuel product and reactor core operation features together are referred to as upgraded core features. A brief description of each of the
- above features is given in the following sections.
, 2-1
2.2 Upgraded Fuel Product Features 2.2.1 Removable Top Nozzle Removable Top Nozzle ('lTN) refers to a top nozzle attached to the fuel assembly as shown in Figure 2-1, in a manner that facilitates the removal and replacement of the nozzle to allow access to the fuel rods for ir.3pection, removal, or replace. ment. The RTN inccrporates dimensional changes that allow for greater fuel rod growth and greatep fuel rod plenum length.
P A stainless steel insert at each thimble location acts as a transition piece between the zircaloy thimble top and the corresponding adaptor plate hole, as illustrated in Figure 2-1. The stainless steel adaptor pitte contains an array of through-hole =, one at each thimble location, each wich a m.tchined circumferential groove. The top end of the stainless steel insert has a circumferential ledge and is axially slotted. The axial slots permit the I insert to flex, or "collapse," into the mating adaptor plate hole and slide 1
until the circumferential ledge of the insert aligns with the circumferential l groove of the hole. The insert then snaps into place, j
The bottom portion of the insert concentrically overlaps the top of the L thimble. Concentric four-lobe bulges in the overlapping tubes joln the j
thimble to the insert, as illustrated in Figure 2-1. An internally concentric
! lock tube placed in the insert top end extends below the axial slots in the insert and prevents insert deformation when the joint is loaded axially, j Local "dimples" projecting from the outside surface secure the lock tube, 1 Lock tube removal is accomplished by a remotely inserted tool which overpowers the securing dimples. After the lock tubes are removed, a moderate axial load will depress the insert ends as a group, and the nozzle can be lifted away i from the fuel assembly proper. All fuel rods are then directly accessible for i
inspection, removal, or replacement, i
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I l 2-2 l
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0 0 The top nozzle can then be reattached to the fuel assembly by reinstalling the nozzle and reinserting the lock tubes. The RTN design also incorporates the "extended burnup," or low profile, top nozzle geometry. This low profile nozzle geometry is designed to accommodate a longer fuel rod and to provide room for greater fuel rod growth within the 14x14 fuel assembly. Increased fuel rod plenums and rod growth gaps accommodate increased fission gas releases and fuel rod growths associated with extended discharge fuel burnups.
2.2.3 Integral Fuel Burnable Absorbers The Integral Fuel Burnable Absorber (IFBA) is a fuel rod design that is based upon the 14x14 fuel rod design and incorporates a burnable absorber inside the rod. The IFBA is a section of fuel pellets coated by a thin film (less than 0.001-inch thick) of zirconium diboride (ZrB 2) burnable absorber material.
The zirconium diboride IFBA provides a nominal boron (isotope 10) burnable absorber loading of 1.67 milligrams of B10 per inch of coated pellet length.
The ZrB 2 c ated section of the fuel stack is typically centered axially and can be sized to meet the requirements of the cycle-specific core design. The IFBA fuel rod design is illustrated in Figure 2-2.
2.2.4 Axial Blankets Axial blankets are nominally six-inch sections of natural uranium pellets at the top and bottom of the fuel stack of each fuel rod. The axial blanket zones in an IFBA fuel rod are illustrated in Figure 2-2.
2.2.5 Debri; Filter Bottom Nozzle Debris Filter Bottom Nozzle (DFBN) refers to a modified VANTAGE 5 bottom nozzle incorporating a flow hole pattern designed to limit the passage of debris above the top plate, as illustrated in Figure 2-3. The DFBN also incorporates dimensional changes that allow for greater fuel rc 4 growth and greater fuel rod plenum length. It is fabricated from stainless steel, which differs from the VANTAGE 5 Inconel nozzle described in reference 2-3. The modified stainless steel nozzle meets all design requirements for 14x14 VANTAGE 5 fuel.
2-3
, o The re-designed top plate of the nozzle includes a revised pattern of flow l holes that are specifically designed to:
- a. Reduce the passage of flow-entrained debris into the fuel assembly,
- b. Maintain the structural integrity of the nozzle, and,
- c. Maintain the hydraulic performance of the fuel assembly.
2.2.6 Extended Burnup Geometry The terminology of "Extended Burnup Geometry" applies to those aspects of the i fuel assembly geometry, including shorter nozzles, longer fuel rod growth gaps, and longer fuel rod plenums, that increase the capability of the fuel assembly to achieve extended discharge burnups. Increased fuel rod plenums and rod growth gaps accommouate the increased fission gas roleases and fuel rod growths associated with extsnded discharge fuel burnups. ;
2.3 PLEX Fuel Management Features The Point Beach cores are currently reloaded each cycle with fresh 0FA feed i assemblies in a pattern known to reduce the leakage of neutrons from the core, !
called a loeleakage-loading pattern. The current hot channel enthalpy rise l peaking factor limit F 3g for Point Beach is 1.58. i i ,
! To achieve improvements in fuel utilization and to achieve reductions in i
reactor vessel fast neutron fluence, it is desirable to further reduce neutron l flux and associated neutron leakage at the core periphery. This is achieved i
- by designing reload cores with lower peripheral assembly powers, a higher hot l I channel peaking factor (F AH limit = 1.70), higher total peaking factor The lower assembly i l (Fg = 2.50), and PPSAs (part-length hafnium absorbers).
powers at the core periphery reduce core-wide neutron leakage from the core, ;
and the peripheral power suppression assemblies reduce local neutron leakage at the cardinal flats of the core near critical reactor vessel welds, i
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-~ --
2.3.1 Low-Low-Neutron-Leakage Loading Patterns Low-Low-Neutron-Leakage Loading Patterns (L4P) are core loadins '.atterns incorporating reduced peripherci assembly powers and hot channel peaking
. factors of up to a 1.70 limit.
L4P, designed to a hot channel peaking limit of 1.70, result in a significant j reduction in the leakage of neutrons from the core. At Point Beach, the benefits of such patterns are:
L
- a. Improved fuel utilization, and
- b. Reduced fast neutron exposure of the reactor pressure vessel with the corresponding reduction in the irradiation induced embrittlement rete !
of the vessel material.
2.3.2 Peripheral Power Suppression Assemblies (PPSAs)
$ It is also desirable to further suppress assembly powers at the flats of the
. core to shield key areas of the vessel from fast nettron flux. The use of !
core components containing a part-length hafnium neutron absorber provides the !
l additional local power suppression. The ccre component is of a mechanical <
design similar to the Westinghouse hafnium Rod Cluster Control Assembly (RCCA) ;
designs currently in use and will fit into the thimble tubes of the fuel assemblies. PPSAs are neutron-absorbing core component assemblies which locally suppress the power at the periphery of the fuel core near critical reactor vessel welds. The core locations for the PPSAs are provided in Figure 2-4.
The use of L4P and PPSAs requires an increased peaking factor limit. The following sections of this report document the analyses and evaluations that establish the acceptability of increasing the hot channel peaking factor limit at Point Beach from the current 1.58 to a new licensing limit of 1.70 and the total core peaking factor from 2.21 to 2.50.
2-5
2.4 Thimble Plugging Devices I
Thimble plugging devices are used to minimize the degree of core bypass flow passing through fuel assembly thimble tubes. These devices, described in more detail in Section 6.0, are inserted into the guide tubes of those fuel assemblies not already occupied by RCCAs, neutron source rods, burnable neutron absorber rods, or water displacer rods.
The analyses dccumented in this report, in particular those of sections 4.0, 6.0, and 7.0, also include verification that the operation of the Point Beach units with the guide thimble plugging devices removed is acceptable.
RJi RENCES 2-1. Oavidson, S.L., and lorii, J.A., "Reference Core Report - 17x17 Optimized Fuel Assembly," WCAP-9500-A, May 1982.
2-2. Letter f rom J. R. Miller (USNRC) to C. W. Fay (WEPCO), subject:
i Revised Technical Specifications to Allow Use of Westinghouse OFAs in Point B3ach Reloads, October 5, 1984, i
- 2-3. Davidson, S.L., and Kramer, W.L., Ed., "VANTAGE 5 Reference Core f Report VANTAGE 5 Fuel Assembly," WCAP-10444-P-A (Proprietary) and l WCAP-10445-A (Non-Proprietary), September 1985.
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l 3.0 NUCLEAR DESIGN
! 3.1 Introduction l
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The nuclear design portion of this report has two objectives. First, the impact on the key safety parameters due to the upgraded core features will be f evaluated. These safety pa',ameters are used as input to the FSAR Chapter 14 accident analyses. Second, the plant Technical Specifications that apply to nuclear design must. be reviewed to determine if they remain applicable or must be revised to accommodate a core containing the upgraded core features.
l i
To satisfy these objectives, a representative core model which contained the [
l upgraded core features was developed using fuel management techniques typical i of anticipated Point Beach fuel cycles. The upgraded core features include !
l the use of L4P with six-inch natural uranium axial blankets at the ends of (
each fuel rod and the use of highly burned fuel in all periphoral assembly l
l locations. In addition, partial-length non-burnable absorber rods (PPSAs), '
composed of hafnium, are loaded in the guide tubes of the assemblies located i on the core flats to further decrease neutron leakage and provide reduced fast neutron fluence to the reactor vessel critical welds. To reduce core power ;
peaking, selected interior fresh fuel assemblies contain partial-length IFBA i rods.
Key safety parameters, listed in Table 3.1, were evaluated to determine the j expected ranges of variation of these parameters. The safety parameters I referred to here are those described in the Westinghouse standard reload design methodology, Reference 3-1. The majority of these parameters are ,
insensitive to fuel type and are primarily loading pattern-dependent, e.g., I control rod worths and peaking factcrs. The observed variations in these loading pattern dependent parameters for the core containing the upgraded f features are typical of the normal cycle-to-cycle variations for core reloads, l
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A Point Beach core containing the upgraded core features will necessitate some i Technical Specification changes as a result of increased peaking factors. The increased peaking factor limits are needed because of reduced neutron leakage fuel management schemes required primarily to achieve reduced reactor vessel
! fluence. The upgraded core features will reduce leakage both radially and axially and provide improved fuel economy and increased nuclear design flexibility.
3.2 Methodology The methocit used in this analysis are based upon the Westinghouse standard reload design methodology and the methods related to VANTAGE 5 fuel described in Reference 3-2. The core models are based primarily upon improved three-dimensional nodal methods described in Reference 3-3. These licensed methods and models have been used for Point Beach and other previous Westinghouse reload designs. No change to the nuclear design philosophy, methods, or models are necessary because of the upgraded core features.
Increased emphasis will be placed on the use of three-dimensional nuclear models because of the axially-heterogeneous nature of the fuel design when axial blankets and part length absorbers (IFBAs and PPSAs) are used.
The reload design philosophy employed includes the evaluation of the reload core key safety parameters which make up the nuclear design-dependent input to the FSAR safety evaluation for each reload cycle. This philosophy is described in References 3-1 and 3-2. These key safety parameters will be evaluated for each Point Beach reload cycle. If one or more of the key parameters fall outside the bounds assumed in the safety analysis, the reload core will be re-designed or the affected transient will be re-evaluated and the results documented in the RSE for that cycle. The primary objective of the Point Beach upgrade analysis is to determine, prior to the cycle-specific reload design, if the previous key safety parameters will continue to remain applicable. The results of this upgrade core analysis are described in Section 3.4.
3-2
3.3 Design Evaluation - Physics Characteristics and Key Safety Parameters As previously mentioned, a representative model containing the upgraded core features was generated for nuclear design evaluation. A low-low-leakage loading pattern consisting of 28 fresh fuel assemblies was generated to achieve a cycle length of approximately 10,500 MdD/MTV. The feed fuel inventory consists of 16 assemblies enriched to 4.0 w/o U-235 and 12 assemblies with an enrichment of 3.8 w/o U-235. A total of 288 IFBA rods is used in the twelve 3.8 w/0 assemblies. Each assembly contains 24 IFbA rods. All fuel assemblies, with the exception of the center assembly, contain axial blankets at the top l and bottom of each fuel rod. The twelve fuel assemblies on the core flats each contain a PPSA, consisting of hafnium in the lower 6 feet of the guide tubes.
Table 3.2 summarizes the core loading pattern nuclear design parameters.
Figure 3-1 shows the quarter core beginning of cycle assembly burnups, IFBA placement, and PPSA locations for this pattern. The representative loading l
pattern utilizing the upgraded core features has been shown to be within the key safety parametera listed in Table 3.1.
3.4 Design Evaluation - Power Distributions and Peaking Factors 3.4.1 Evaluation Discussion The implementation of L4P, PPSAs, and axial blankets will have impact on core power distributions and peaking factors experienced in the Point Beach cores.
- The use of axial blankets, where the top and bottom of the enriched fuel stack l are replaced by natural uranium pellets, and the enrichment of the remaining fuel is increased slightly, reselts in higher axial peaking factors. The use of L4P and PPSAs results in reduced fluence to the reactor vessel and improved fuel utilization by placing less reactive fuel on the periphery of the core.
The reduction in power carried by the peripheral assemblies is offset by increases in power in the remaining assemblies. The increased radial and axial peaking is accommodated by increasing the core peaking factor limits, i F aH and Fg .
3-3 J
A representative loading pattern was developed and modeled based upon the anticipated upgraded core features. Results of calculations show an increase in radial peaking from previous cycles, which is not unexpected. This results from the reduced power carried by the more highly-burned assemblies placed on the core periphery to reduce neutron leakage, as well as from use of axial blankets which reduce power at the extreme top and bottom of the fuel, thereby j reducing axial leakage. To achieve acceptable radial power peaking near beginning of cycle, selected fresh fuel assemblies were loaded with IFBA rods. The coated lengths of the IFBA rod fuel stack were reduced to 96 inches and axially centered at the core midplane to obtain reasonable normal-operation, elevation-dependent Fq values. The coated IFBA length will be adjusted on a cycle-by-cycle basis to optimize Fqmargin.
Figures 3-2 through 3-5 give the quarter core power distributions at hot f ull power (HFP), all rods out (ARO), at various cycle burnup steps. The assembly ;
powe-s are typical values and do not necessarily represent bounding values for j future designs. These power distributions are for illustrative purposes and i do not represent the full scope of the nuclear analysis that was performed. !
The total peaking factor, Fg, was evaluated as a function of core height for the loading pattern consisting of the upgraded core features. Various operating conditions were imposed to achieve variations in power i
distributions. The limiting values of F qtimes relative power were i maintained below the Fg limit of 2.50 times K(Z) as shown in Figure 3-6.
- The Fg limit it based upon the assumption of the removal of the third line segment from the K(Z) curve. The calculated Fq values resulted from Relaxed l
3 Axial Offset Control (RAOC) analyses performed to determine an allowable flux j difference operating envelope based on the upgraded fuel product features and f the increased Fg limit as discussed in the following subsection. (
t 3.4.2 Technical Specification Changes Relative to Nuclear Design l There are three major areas related to nuclear design that will affect Technical Specifications: increase in F g limits, increase in Fq ;
limits, and a change to the allowable flux difference operating envelope f (RAOC AI band). The F limit will be increased from 1,58 to 1.70, j AH i.
1 3-4
t4' 1 including uncertainties at hot full power conditions. Similarly, the calculated Fg over the power range from HZP to HFP becomes F g X U 1 (Fg limit) [1 + 0.3 (1-P)].
where P is the fraction of full power, F g limit = 1.70, and U = 1.08 to allow for an eight (8) percent calculational uncertainty.
The F limit will be increased from 2.21 to 2.50 at relative power levels q
greater than 50 percent. The F q limit at power levels of 30 percent or less will be increased from 4.42 to 5.00. The Fq functions becomes 2 50 Fq (Z) 1 p x K(Z) for P > 0.5 Fq (Z) 1 [x K(Z) for P 1 0.5
- where P is the fraction of full power at which the core is operating, K(Z) is the normalized F (Z) function as shown in Figure 3-7, and Z is the core
)
q height location of Fq.
The K(Z) curve shown in Figure 3-7 is based upon the assumption of the removal of the third line segment from the previous Technical Specification curve.
The removal of this line segment allows a larger acceptable Fq at the top o'T l the core, which results in a more flexible core design and plant operation.
I Small-break LOCA analyses have confirmed the acceptability of this revised K(Z) curve. This assumption is contingent, of course, upon satisfactory results from the on going large-break LOCA analysis.
An increase in the Fq limits and the other upgraded core features for the
- Point Beach cores require a change to the allowable flux difference operating envelope (RAOC AI band). The RAOC analysis methodology is described in i
detail in Reference 3-4. Explicit calculations were performed for the represen-tative core consisting of the upgraded features.
i 3-5 i
l Initial calculations assumed current rod insertion limits and full-length IFBA I rods. This resulted in a very conservative, narrow AI band near full power.
To obtain a wider and more realistic AI band, the coated length of the IFBA i rods was reduced to 96 inches, in order to achieve lower calculated Fq values. l In addition, shallower rod insertions were employed. The resulting AI envelope is shown in Figure 3-8. The corresponding power shapes were confirmed by thermal-hydraulic analyses to meet DNB safety limits (see Section 4.0). The ;
reduced coated length of the IFBA rods is necesssary, due to the presence of axial blankets. The coating length is loading pattern-dependent and will be adjusted on a cycle-by-cycle basis to obtain optimized Fq values. The coating length will be reduced or increased, depending upon the number of IFBA rods and the core loading arrangement. In cycles in which full-length discrete burnable absorber rods are used, the fuel assemblies containing these absorbers will not have axial blankets. This will result in lower Fq values and, therefore, the al envelope will be conservative.
1 i In order to ensure that the RAOC al band is conservative for actual upgraded
- core cycles, a change to the rod insertion limit is required. The current !
insertion limits will be raised 14 steps (approximately 6 percent) at all L
power levels. The new rod insertion limits are shown in Figure 3-9 and pose j 4
no adverse impact on other safety parameters, t i
3.5 Conclus un i
The key safety parameters evaluated for the conceptual design show that the !
l
- expected ranges of variation for many of the parameters will lie within the j normal cycle-to-cycle variations observed for reload designs.
fa In addition to the normal variations experienced with different loading l 6
patterns, power distributions and peaking factors show some changes as a result of the incorporation of the upgraded fuel product features and I increased peaking factor limits. The usual methods of loaaing pattern shuffling and enrichment variation can be employed in future cycles using the upgraded core features to ensure compliance with the Point Beach revised
]
i Technical Specifications. ,
) .
i 3-6 l
1' Therefore, nuclear design-related Technical Specifications changes will be limited to increases in F AH and Fq and the as-noted changes to the K(Z) curve, al band, and the rod insertion limits. Outside of these specific areas, there will be no further core upgrade-related changes to the Technical Specifications required by the nuclear design for the proposed changes to Point Beach.
In summary, the change from the ct,r< tnt core to a core containing the upgraded >
core features will not cause changes to the current nuclear design bases given in the Point Beach FSAR. The evaluation of the Point Beach upgrade demonstrates that the impact of implementing the upgraded core features does not cause a i significant change to the physics characteristics of the Point Beach core beyond the normal range of variations seen from cycle to cycle.
1 l REFERENCES 3-1. Davidson, S. L. (Ed.), et. al., "Westinghouse Reload Safety Evaluation Methodology," WCAP-9272-P-A (Proprietary) and WCAP-9273-A (Non-Proprietary),
- July 1985.
l l- 3-2. Davidson, S. L. and Kramer, W. R. (Ed.), "Reference Core Report VANTAGE l I 5 fuel Assembly," WCAP-10444-P-A (Proprietary) and WCAP-10445-A '
(Non-Proprietary), September 1985. I i
i
) 3-3. Davidson, S. L. (Ed.), at. al., "ANC: Westinghouse Advanced Nodal .
I Computer Code " WCAP-10965-P-A (Proprietary) and WCAP-10966-A (Non-Proprietary), September 1986.
3-4. Miller, R. W., et. al., "Relaxation of Constant Axial Offset Control,"
WCAP-10216-P-A (Proprietary) and WCAP-10217-A (Non-Proprietary), June :
1983.
I i
1 j 3-7 L
TABLE 3.1 KEY SAFETY PARAMETERS Safety Parameter Upgraded Core Reactor Core Power, (MWt) 1518.5 Vessel Average Coolant Temp. 573.9 HFP, (Deg F)
Core Average Linear Heat Rate. 5.7 (Kw/ft)
- Most Positive Moderator Temperature 5
- Coefficient (MTC), (pcm/0eg F)
Most Positive Moderator Density .43 Coefficient (MDC), (Ak/g/cm3 )
Doppler Temperature Coefficient. -0.91 to -2.90 (pcm/*F)
Doppler Only Power Coefficient
- Least Negative, (pcm/% Power) -6.05 to -9.55
- Most Negative (pcm/% Pcwer) -12.6 to -19.4 Least Negative Doppler Only Power Defect -0.90
, at BOL, (% delta-rho)
Beta-Effective .0043 to .0075 I:inimum Shutdown Margin at EOL, (% delta-rho) 2.77 Normal Operation Fgg (without Uncertainties) 1.574 3-8 l
TABLE 3.1 (Continued)
KEY SAFETY PARAMETERS Safety Parameter Upgraded Core Fq (including uncertainties) 2.50 Differential Rod Worth for Rod Withdrawal from Suberitical, (pcm/in) 130 Ejected Rod Worth, (% delta-rho)/ Maximum Fq
- BOL HFP 0.40/4.50
- BOL HZP 0.79/12.00
- EOL HFP 0.42/5.69
- EOL HZP 0.95/12.00 i
I 3-9
O
- TABLE 3.2 REPRESENTATIVE NUCLEAR DESIGN PARAMETERS LOW-LOW-LEAKAGE LOADING PATTERN (L4P) FEATURES:
Cycle Length 10,500 MWD /MTU Feed Fuel 16 assemblies at 4.0 w/o U-235 12 assemblies at 3.8 w/o U-235 I
Axial Blankets 6-inch natural uranium at top and bottom of each !
fuel rod "
IFBAs 24 rods in each of 12 assemblies i (Coated Length = 96 inches) !
Hafnium Rods 6 feet in lower half of the guide tubes in the 12 [
assemblies on core flats l
j ,
r l
I t
l I
i I
l 3-10 !
r
o
- 34
_ 30000 11864 37330 11464 19800 4 M343 M
11984 30533 13791 34033 37900 0 41338 H
s.
27325 13751 20035 14433 34 l 0 p 30301 l
11854 39033 14433 37753 0 30318 34 13eH 37800 0 0 40030 34 0 0 38 M1 M3tt N temsta trea n005 M AVERAa3 SW484J#
_ H PPSA 34343 41338 M M FIGURE 3-1 UPGRADED CORE ASSEMBLY AVERAGE SURNUP DISTRitUTION AND ABSORSER PLACEMENT AT SOL (0 MWD /MTU) 3 11 ,
o e
- 9.910 1.393 1.160 1. 4M 1. NS 1. 3 M 0.303 i
l 1.393 0.Mt 1.360 1. 3M 1.979 1.1 M 9.331 l
1.180 1.394 1.213 1.396 1.381 0.460 ]
-- -" 5 1.498 1.215 1.384 1. M7 1.193 0.310 ,
l w
l 1. M S 1.079 1.381 1.191 0.405 l
l -
i .n. i . in ...., ..Si.
AP AYtRASE P0wtR x
e.ses e.331 4 m Coanem suoicATis Assaisel.Y ANO QUA04 ANT 08 Plat FIGURE 3-2 UPGRADED CORE NDRMAUZED POWER Ol5TRitUTION AT 150 MWD /MTU.
UNRODDED. HPP. EQUILitRIUM XENON PEAK F-DELTA-H = 1.555 3 12
o .
I' l
i t
_ S.004 1.364 1.147 1.431 1. He 1.388 9. 3M I
1.350 S. Del 1.390 1.180 1. M 3 1.153 0.333 f
1.147 1.383 1.1H 1.M1 1.444 0.514 1
1.431 1.105 1.343 1. M9 1.308 8.335 i 1.344 1. M 5 1.444 1.307 0.431 a d
l 1.345 1.1H 8.554 l 9.338 AP AVERAGE P0wth s
- 9. 3M t.333 x W comNan peoicAfts Asseusty ANo ouADRANT OF PGAn F10VRE 3-3 UPGRADED CORE NORMAll2ED POWER DISTRIBUTION AT 2000 MWD /MTU.
UNRODDED. HFP. EQUluBRIUM XENON PEAK F-DELTA-H = 1.574 3 13
o e I
- 8.039 1.398 1.130 1.T79 1.334 1.303 9.318 -
I 1.395 9.983 1,
- 8 1.188 1. M 3 1.170 0.351 l 1.130 1.731 1.183 1.311 1.411 0.644 x
1.378 1.173 1.2 13 1.881 1.213 0.M3 ,
1 i
i 1.334 1.949 1.411 1.313 S.448 j i
r 1.3;3 1.171 0.544 9.381 (
AP AvtRAtt Powlt !
K O.314 0.311
' K 181 cospett IMOICA718 AS$PMSLY A4 QUADe&N7 08 PtAE I
f I
f i
i i
P10011tt 3-4 l l UPGRADED CORE NORMAllZED POWER DISTRIBUTION AT 6000 MWD /MTU, UNRODDED, MFP. IQUILIBRIUM XENON PEAK F-DELTA-H = 1.551 i e
3-14 t 3
$ t
(L ,
1 9
e
- 8.074 1.379 1.183 1. M4 1.304 1.371 0.844
_ l ,
1.370 0.083 1. M3 1.158 1. M 3 1.11 3 0.300 1.133 1.306 1.163 t.383 1. MB 4.574 x
- 1. M8 1.143 1.082 1. Mt 1.310 4. Se t 1 _- -
i i i 1.204 1. M 3 1.340 1.300 S.440 l
l 1.371 t.173 e.s74 e.M6 I !
g ,
an avemaN powsa m i s.34e e.See l t
x a coswesa escHearss assaan6LY AND QUADRA8tf Of PEM !
, i 1
4 i
i i
t i
F10URE 3-5 l UPGRADED CORE NORMALIZED POWER DISTRIBUTION AT 10500 MWD /MTV, 6 3
UNRODDED HFP, fouluBRIUM XENON PEAK F-DELTA-H = 1.503 ,
f 3-15 !
l l
3.
I10.0,2.50 l f6.0,2.501l
'I**' '
2'6 e- a ,e s==- e sT
.e 8e ,,* s %*
e e
- 2. ,
'e e
"a E
e.
(,1.5 5
m 1.
e e
e
.S 00 1 2 3 4 5 s i s e to 11 12 CORE HIG N (FT)
FIGURE 3-6 T
MAXIMUM Fa , p EL R vensus AXIAL Cone HEIGHT DURING NonMAL Come OpenATION 3-16
e .
1.2 l
(0.0,1.0)I (6.0,1.0) l g,g 3, (12.0. 92)l
.9
.7 - - -
g .6
.5
.4
.3
.2
.1
'o 1 2 3 4 5 6 7 4 9 to 11 12 CDRE E31G57 (PT)
FIGURE 3-7 NORMALIZED Fg LIMIT VERSUS CORE HEIGHT 3 17
a .
110 g 106 100 i .g,im j 9.100 1 95 to i . g , g, j 9,00 E \
70
/ \
/ \
( .M ,50 l IY.50 l
-1S -10 0 5 to is 20 23 30 M.3o 25 =20 -5 DCLTA 1 (f.)
FIGURE 3 8 FLUX DIFFERENCE OPERATING ENVELOPE FOR UPGRADED CORE 3 18 1
l 7 24o ---
- /s '7d '/ %
1, 94 */ s
,r e. c, rao. E */,-*
/ '
,' /
2co
/
/ ,
i /
/ / . .n.
)
f t.B,TK 8i ,
< 5 sei (176 ' < '
/ , j i d1 52] ,
/ -
/ '
{O ,
/ ,
/ (171)
/ -
.' /
'/
5 ,
/ -
s' / / /
m 120 i f i SR* C I , /
/ / r a e' / ,
/
y s' / ,
l lso i / .' / EprK O i
/ f j
/ s /
/ i
/
(50 l/
, / '
/
so
/ ,
' /
!?G ) O.RRDIT R!L'S e
/ - - - PRCKED RIL'S __
l /
( 28. 5% >
~/(2 4% )
O too o 2o e so ao PUtGNT & F1JLL PotOI FIGURE 3-9 UPGRADED CORE PROPOSED ROD INSERTION LIMITS 3-19
4.0~ THERMAL-HYDRAULIC DESIGN 1
4.1 Introduction .
This section describes the thermal-hydraulic analyses performed to support the .
use of the. upgraded core features in the Point Beach Units 1 and 2. The thermal-hydraulic design for the upgraded core features was analyzed for a nuclear enthalpy use hot channel factor (FAH) limit increase from 1.58 to 1.70. The increase in F3g and the offect of thimble plug removal were accommodated by using the Departure from Nuclear Boiling Ratio (DNBR) design margin available in the safety analysis DNBR. Table 4.1 summarizes the thermal-hydraulic design parameters used in these analyses. The thermal-hydraulic design criteria and methods remain the same as those !
presented in the Point Beach FSAR, with the exceptions noted in the following section. All of the current thermal-hydraulic design criteria are satisfied with the upgraded core features.
4.2 Calculational Methods The existing thermal-hydraulic analysis for 0FA fm1 currently used in the Point Beach units is based upon the Improved Thermal Design Procedure (ITDP),
- reference 4-1, and the Westinghouse Critical Heat Flux (WRB-1) correlation, reference 4-2, as described in the Point Beach FSAR. The analysis of the upgraded core features is based upon the Revised Thennal Design Procedure (RTDP), reference 4-3, and the Westinghouse Critical Heat Flux (WRB-1)
] correlation. In addition, the W-3 correlation, reference 4-9, is used where appropriate. (See Section 7.0.)
The RTDP removes some of the conservatism in the ITDP methodology while i satisfying the design criterion that protects against DNB in the core. The -
DNB thermal design criterion is that the probability that DNB will not occur on the most limiting fuel rod is at least 95% (at a 95% confidence level) during normal operation and operational transients and during transient f conditions arising from faults of moderate frequency (Condition I and II events, as defined in ANSI N18.2).
4-1
.With the ITDP methodology, system uncertainties are statistically combined separately from DNB correlatior. uncertainty. The two are combined directly to determine the DNBR limit. With tne RTDP methodology, the system and correlation uncertainties are statistically combined to predict the DNBR limit.
For this application, the system uncertainties, correlation uncertainty, and calculation of DNBR limit are considered for the typical'and thimble cells, respectively. The design value is conservatively increased to include 8.6%
DNBR margin. The safety analysis DNBRs are calculated as follows:
Typical Cell Safety Analysis DNBR = Typical Cell Design Limit DNBR 1.0 - Margin 1.217 1.0 - 0.086 = 1.33 Thimble Cell Safety Analysis DNBR = Thimble Cell Design Limit DNBR 1.0 - ;iargin 1.212
= = 1.33 1.0 - 0.086 The THINC IV computer program was used to perform thermal and hydraulic calculations. THINC IV calculates coolant density, mass velocity, enthalpy, vapor void, static pressure, and DNBR distributions along flow channels within a reactor core under all expected operating conditions. The THINC IV code is described in detail in references 4-4 and 4-5, including models and correlations used. In addition, a discussion on experimental verification of THINC IV is given in reference 4-5.
4.3 Hydraulic Compatibility-Transition Core For thermal-hydraulic purposes, the upgraded fuel product is hydraulically identical to the 14x14 0FA fuel currently used in Point Beach Units 1 and 2, and no transition core penalty is required. The use of STD fuel assemblies requires a small DNBR penalty on all the fuel.
4-2
J 4.4 Effects of Fuel Rod Bow on DNBR a
The phenomenon of fuel rod bowing must be accounted for in the DNBR safety analysis of Condition I and Condition II events. Currently, a rod bow penalty of less than 3% on the DNBR is assessed, based upon references 4-6, 4-7, and 4-8. This penalty is the maximum rod bow penalty for 14x14 optimized fuel at an assembly average burnup of 24,000 MWD /MTV. For burnups greater than 24,000 MWD /MTU, credit is taken for the effect of FAH burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory. Therefore, no additional rod bow penalty is required at burnups greater than 24,000 MWD /MTU.
4.5 DNBR Effect of the Upgraded Fuel The proposed change to the Point Beach Units 1 and 2 Technical Specifications which impacts DNBR, is the increased design nuclear enthalpy rise hot channel factor as defined by the following equation:
N F AH = 1.70 [1.0 + 0.3 (1.0 - P)]
where P = fractional core power level at less than 100% rated power O_I P= 1.0 for core power level greater than or equal to 100% rated power.
The radial peaking factor limit increase from 1.58 to 1.70 has a direct impact on DNBR calculations. The impact on DNBR due to the increased Fg was offset by the additional margin resulting from RTOP methodology and reduction in the margin which defines the ONBR limit value for the safety analysis. The new safety analysis DNBR values were selected to retain a margin sufficient to cover rod bow penalty, transition core penalty when STO fuel is used, and still retain some margin for future use, such as to address a flow shortfall or other minor problem.
4-3
A summary of the DNBR limits for the upgraded fuel is presented in Table 4.2.
The impact of the changes to the core limits on the non-LOCA analyses is addressed in Section 7.1.2.
The axial blankets and the increased allowable Fq , associated with the fuel upgrade, affect the axial power distribution and, therefore, the DNBR analyses. These effects were accounted for in the non-Overtemperature Delta-T (non-0 TDT) DNBR analyses by means of a limiting axial power distribution.
Cycle-specific limiting power shapes will be e/aluated against this limiting l axial power distribution for each Reload Safety Evaluation (RSE). The impact on OTDT accident analyses is discussed in Section 7.1.
4.6 Fuel Temperatures for Safety Analysis The fuel temperatures (as a function of linear heating rate) for use in safety ,
analysis calculations for the upgraded fuel are the same as those used for the current fuel. The PAD fuel performance code, references 5-1 and 5-2, was used for the calculations. The use of IFBAs reduces the beginning-of-life (BOL) fuel temperature, compared to non-IFBA fuel. This effect is a result of the reduced fuel-to-cladding gap because of the presence of the IFBA coating. The effect of axial blankets was considered in the analysis.
4.7 Thimble Plug Removal 4.7.1 Introduction Coincident with implementation of the upgraded fuel product features, Wisconsin Electric plans to remove thimble plugging devices from the Point Beach cores. Thimble plugging devices are currently utilized in Point Beach units to limit the core bypass flow. All fuel assembly guide thimble tubes that are not in RCCA locations or are not equipped with sources or burnable ,
absorbers currently have thimble plugs inserted in them. A net gain of approximately 2% in DNBR margin is realized due to their presence. The evaluation is described in the following sections.
I 4-4 [
i
4.7.2 Bypass Flow i The main impact of thimble plug removal is the increase in core bypass flow.
Calculations performed by Westinghouse have shown that the design limit of i core bypass flow needs to increase from 4.5% to 6.5L This increase is based upon having only RCCAs in the fuel assembly thinibles, with the remaining thimbles open.
4.7.3 Primary System Flow Rate [
Thimble plug 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 upon these tests, ,c is estimated that there will be a slight increase in primary system flow rate due to thimble plug removal from Point Beach cores. No mechanical design criteria are impacted by j this slight increase in flow rate.
4.7.4 Fuel Assembly Hydraulic Lift Forces i l
The hydraulic lift force on the fuel assembly can be represented by the i following function: !
l F
LIFT aK FA x (core flow)2 Westinghouse has performed hydraulic tests to quantify the magnitude of the effect of thimble plug removal on fuel assembly hydraulic loss coefficient (KFA).
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 I
more than compensates for the slight increase in vessel flow rate. Thimble plug removal is therefore acceptable from a fuel assembly lift force standpoint.
P 4-5 I
r: .
4.7.5 Effect of Outlet Hydraulic Mismatch on DNB Current DNB analyses are performed assuming the presence of a 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 DNB zone. Westinghouse has performed numerous sensitivity studies to demonstrate the insensitivity of as-calculated DNBRs to non-uniform outlet pressure distributions and to variations in outlet loss coefficients. '
The effect of thimble plug removal on the core-wide distribution of outlet ;
loss coefficients for the Point Beach cores has been evaluated. It was demonstrated that the variations in outlet loss coefficient due to thimble j plug removal are within the bounds of the sensitivity studies that have been performed. Therefore, it is concluded that removal of all or any combination of thimble plugs will not result in the reduction of DNBR margin due to mismatches in core outlet pressure gradients and loss coefficients.
4.7.6 Mechanical Design Evaluation 4.7.6.1 Fuel Rod Fretting Wear 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 cross flow induced fuel rod vibration and wear due to this increased PF0 mismatch is addressed.
j The maximum PF0 mismaten that exists in the Point Beach core after removal of all or any combination of thimble plugs is much smaller than that found by Westinghouse in recently performed fuel rod vibration tests with two 17x17 fuel assemblies. The results showed that there was no significant difference I in fuel rod response between the tests performed with and without this large PF0 mismatch. This conclusion can be extended to 14x14 0FA based on the l following similarities:
l l
l l 4-6
- c. ..
- a. Lateral flow area / axial flow area ratio
- b. PF0 values
- c. The change in PF0 due to thimble plug removal, and,
- d. The axial velocity in the fuel rod bundle region Because of these similarities, it is judged that the PF0 mismacch and maximum crossflow velocities associated with thimble plug removc1 for any 14x14 0FA will not exceed the test values. Therefore, <t is concluded that thimble plug removal will not have a detrimental effect on fuel rod vibration and wear.
- 4.7.6.2 Control Rod Wear Westinghouse studies on control rod wear have shown that wear tends to occur 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.
However, since the core PF0 distribution changes when thimble plugs are removed, the effect of potential control rod vibration, due to interassembly 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 interassembly crossflows through the gap (~0.75")
between the top nozzle and upper core plate. For the Point Beach reactor upper internals configuration, it was concluded that the maximum PF0 mismatch between an RCCA 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 ro.4 caused by this crossflow will not be increased.
Based upon the above evaluation, thimble plug removal will not have an adverse impact on control rod wear for the Point Beach reactors.
4-7
4.7.7 Summary 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 implementation of the upgraded core features. (See Sections 7.1 and 7.2.)
Based upon the assessment of the impact of the thimble plug removal on core plant safety, it is concluded that, from a thermal-hydraulic standpoint, it is acceptable to remove all or any combination of these devices from the Point Beach cores. The evaluation also bounds the use of any combination of dually-compatible thimble plugs (thimble plugs that can be used in either STD or 0FA fuel assemblies), PPSAs, sources assemblies, and burnable absorber assemblies.
4.8 Conclusion The thermal and hydraulic analysis has shown that the DNBR penalties resulting from the increase in peaking factors and the removal of thimble plugs are offset by the present DNBR margin and the margin provided by RTDP methodology. More than sufficient DNBR margin in the safety limit DNBR exists to cover rod bow penalty. All of the current thermal-hydraulic design criteria are satisfied.
REFERENCES 4-1. Chelemer, H. et. al., "Improved Thermal Design Procedure," WCAP-8567-P (Proprietary) and WCAP-5868 (Non-Proprietary) July 1975.
4-2. Motley, F. E. et. al., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids,"
WCAP-8262-P-A (Proprietary) and WCAP-8763-A (Non-Proprietary), July 1984.
4-3. Friedland, A. J., Ray, S., "Revised Thermal Design Procedure,"
WCAP-11397 (Proprietary) and WCAP-11398 (Non-Proprietary), February 1987.
4-8
. o l
l l
4-4. Hochreiter, L. E., Chelemer, H., Chu, P. T., "THINC IV, An Improved I Program For Thermal Hydraulic Analysis of Rod Bundle Cores," WCAP-7956, i June 1973.
l l
4-5. Hochreiter, L. E., "Application of the THINC IV Program to PWR Design,"
WCAP-8054 (Proprietary) and WCAP-8195 (Non-Proprietary), October 1973. l 4-6. Skaritka, J. , (Ed.) "Fue' '.od Bow Evaluation," WCAP-8691, Revision 1 (Proprietary) and WCAP-86Y2 (Non-Proprietary), July 1979. l l
l 4-7. "Partial Response to Request . Number 1 for Additional Information on WCAP-8691, Revision 1" letter, E. P. Rahe, Jr. (Westinghouse) to J. R. Miller (NRC), NS-EPR-2515, dated October 9, 1981; "Remaining Response to Request Number 1 for Additional Information on WCAP-8691, )
Revision 1" letter, E. P. Rahe, Jr. (Westinghouse) to J. R. Miller (NRC). NS-EPR-2572, dated March 16, 1982.
I 4-8. Letter C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse), "Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum Rod Bow Penalty," June 18, 1986. !
l l
4-9. Tong, L. S., "Prediction of Departure frem Nucleate Boiling for an l Axially Nonuniform Heat Flux Distribution," Journal of Nuclear Energy, Vol. 21, pp. 241-248, 1967.
I l
l l
4-9 .
l
i
'f I
-TABLE 4.1 POINT' BEACH UNITS 1 AND 2 THERMAL AND HYDRAULIC DESIGN PARAMETERS Current Analyses Upgraded Analyses Thermal and Hydraulic Design Parameters (ITDP Methods) (RTDP Methods)
Reactor core heat output, MWt 1518.5 1518.5 i
6 Reactor core heat output, Btu /hr 5182. :10 6 5182.6x10 Heat generated in fuel, % 97.4 97.4 f
, [
Reactor Coolant System pressure, psia 2000 (2250) 2000 (2250)
Radial power distribution 1.58[1+0.3(1-P)] 1.70[1+0.3(1-P)] i Minimum DNBR at nominal conditions-r Typical flow channel 2.49 2.12 -
r il Thimble (cold wall) flow channel 2.39 2.03 :
Minimum DNBR for design transients:
l Typical flow channel 11.66 11.33
! Thimble flow channel 11.65 11.33 4
DNB correlation WRB-1[a] WRB-1[a] (
i
) 4-10 1
i 1
I
. r .-
TABLE 4.1 (Cont.)
POINT BEACH UNITS 1 AND 2 THERMAL AND HYDRAULIC DESIGN PARAMETERS Current Analyses Upgraded Analyses Thermal and Hydraulic Design Parameters (ITDP Methods) (RTDP Methcas)
Coolant Flow 6 6 Total thermal flow rate, lb/hr 68.6x10 68.5x10 Average velocity along fuel rods, 14.1 13.7 ft/sec 6 6 Core average mass velocity,1b/hr-ft 2 2.28x10 2.22x10 Coolant Temperature Nominal inlet, *F 545.0 545.5 Average rise in vessel, *F 56.8 56.7 Average rise in core, *F 58.4 59.7 Average in core, *F 575.4 576.7 Average in vessel, *F 573.4 573.9 4-11
t
-TABLE 4.1 (Cont.)
POINT BEACH UNITS 1 AND 2 THERMAL AND HYDRAULIC DESIGN PARAMETERS t
Current Analyses Upgraded Analyses Thermal and Hydraulic Desian Parameters (ITDP Methods) (RTDP Methods)
Heat Transfer l i
2 Active heat transfer surface area, ft 27,161 27,161 l
[
Average heat flux, Btu /hr-ft 2 -185,850 185,850 Maximum heat flux, for normal operation, 410,728[b] 464,625[c]
Btu /hr-f t 2 .
Average thermal output, Kw/ft 5.7 5.7 Maximum thermal output, for normal 12.6[b] 14,3[c]
l operation, Kw/ft i
f
. -1 Fuel Centerline Temperature.
l ,
} !
I- Peak at peak linear power for prevention 4700 4700 j of centerline melt, 'F f j i i
t a ;
r i
' 1 i
- 4-12 l
O O TABLE 4.1 (Cont.)
POINT BEACH UNITS 1 AND 2 THERMAL AND HYDRAULIC DESIGN PARAMETERS Current Analyses Upgraded Analyses Thermal and Hydraul;c Desian Parameters (ITDP Methods) (RTDP Methods)
Pressure Drop Across core, psi 22.6 20.8
[a] Where appropriate, the Standard Thermal Design Procedure (STDP) ar.d the W-3 DNB correlation are used. (See Table 7.1.)
[b] This limit is associated with the value of Fq = 2.21 at 100 percent power
[c] This limit is associated with the value of Fq = 2.50 at 100 percent power i
l 4-13
o .
b s
TABLE 4.2 DNBR MARGIN
SUMMARY
Standard Upgraded Fuel Assembly [a] Current Analyses Analyses (STDP[b]) (ITDP) (RTDP)
Correlation Limit 1.30[c] 1.17 N/A Design Limit Typical Cell 1.30 1.33 1.22 Thimble Cell 1.30 1.32 1.21 Safety Limit Typical Cell 1.30 1.66 1.33 Thimble Cell 1.30 1.65 1.33
[a] Limited to previously-deplettd STD fuel. The use of previously-depleted STD fuel assemblies will be justified by cycle-specific reload analysis.
[b] Standard Thermal Design Procedure
[c] W-3 L-Grid.
4-14
I L
5.0 FUEL ROD DESIGN f
5.1 Introduction f The fuel rod design evaluation to support increased allowable core power peaking factors fo9 Point B ach Nuclear Plant, Units 1 and 2, is based upon meeting the fuel rad design criteria for the most limiting fuel rod design considered for the Point Beach units. Fuel rod features bounded by these performance evaluations include all combinations of Westinghouse STD and 0FA !
fuel, as currently used in the Point Beach Units, and upgraded fuel product features as described in Section 2.2. l l
l l
Increased core power peaking factors affect fuel rod design through increases :
! in the steady-state fuel rod power histories and in the fuel rod transient [
duty. The fuel rod design criteria affected by this more severe fuel duty are
~
f the rod internal pressure, cladding stress and strain, and cladding surface ]
temperature. The evaluation of these design criteria for the bounding Point Beach fuel rod designs and duty, discussed below, shows that the criteria are f
^
satisfied. i i
[
i 5.2 Methodology and Input Assumptions ;
f i
5.2.1 Fuel Rod Design Criteria j The relevant fuel rod design criteria used by Westinghouse to ensure reliable fuel service for all operations consistent with ANSI N18.2 Condition I and/or !
Condition II events are given below. The fuel rod design is judged to have [
met these criteria when it is demonstrated that the performance of a fuel [
region is within the limits specified by the criteria for these events. (
i 5.2.1.1 Rod Internal Pressure l
The design basis for fuel rod internal pressure is that the fuel system will ;
not be damaged due to excessive fuel rod internal pressure. The current l
! NRC-approved design limit for fuel rod internal pressure is that the internal ,
i
! i l
! 5-1
! t I
,, o I
pressure of the' lead rod in the reactor will be limited to a value below that which could cause
- a. The diametrical gap to increase due to outward cladding creep during steady-state operation, and, l
- b. Extensive DNB propagation to occur.
The rod internal pressure limit value which precludes gap increase and DNB propagation is a function of (and significantly higher than) system pressure.
5.2.1.2 Cladding Stress The design basis for fuel rod cladding stress is that the fuel system will not be damaged due to excessive fuel cladding stresses. The design limit for fuel rod cladding stress is that the volume average effective stress is less than the zircaloy 0.2% offset yield strength, with due consideration to temperature and irradiation effects under Condition I and II n. odes of operation. While the cladding has some capability for accommodating plastic strain, the yield stress has been accepted as a conservative design limit.
NRC-approved Westinghouse fuel performance codes are used for evaluating cladding stress limits. (See Section 5.2.2.1.) Both steady-state (Condition I) and transient (Condition II) conditions are evaluated. However, analyses have shown that transient cladding stresses are more limiting than steady-state stresses and that the limiting stresses occur during the second cycle of operation.
5.2.1.3 Cladding Strain The design basis for fuel rod cladding strain is that the fuel system will not be damaged due to excessive fuel cladding strain. The design limit for fuel rod cladding strain during steady-state operation is that the total plastic tensile creep and uniform cylindrical fuel pellet expansion due to fuel swelling and thermal expansion are less than 1 percent from the unirradiated 52
I t
condition. For Condition II transients, the design limit for cladding strain is that the total tensile strain due to uniform cylindrical pellet thermal expansion during the transient is less than 1 percent from the pre-transient value. These limits are consistent with proven practice. l During steady-state operation, tensile cladding creep strain results primarily ;
from cladding stresses caused by pellet swelling and thermal expansion, i following the closing of the pellet-clad gap. For Condition II events, transient analyses have shown that the limiting tensile creep strains, [
resulting from pellet thermal expansion, occur during the second cycle of ;
operation when transient cladding stress is also most limiting. Resulte also (
show that transient strain criteria are less limiting than transient stress ;
criteria and, therefore, transient strain limits are always met when transient !
stress limits are met, f
, t
' 5.2.1.4 Cladding Surface Temperature /0xidation l The design limit applied to cladding oxidation evaluations is that the l
calculated cladding temperature (oxide-to-metal interface) shall be limited 4
during steady-state operatio and Condition II transients to preclude a !
l condition of accelerated oxidation. i
. l The factors that are considered to control in-reactor corrosion for Zircaloy-4
(
are the temperature (metal-to-oxide interface) and irradiation enhancement to J
! the degree applicable. Extensive experience has shown that irradiation enhancement of the corrosion rate does not occur under normal operating conditions in current Westinghouse PWRs. The controlling factor for the (
, in-reactor zircaloy corrosion rate is therefore the oxide-to-metal interface r
! temperature, which is controlled by the above cladding temperature criteria. f
- i 5.2.2 Fuel Rod Design Methodology I f l The fuel rod design criteria are evaluated on a best-estimate plus f
! uncertainties basis. Best-estimate results are obtained using best-estimate i I fuel performance models, nominal fuel fabrication attributes, and 1
I t
k
{
, 5-3 p
~ _ - - - - - - - - - - , -
.y. - .
1 E
i best-estimate pouars, fluxes, and fluences. Uncertainties with respect to the I design criteria are calculated separately for the significar.t model, fabrication, and nuclear uncertainties. Typical model uncertainties considered in fuel performance evaluations are fission gas release, helium ,
release, rod growth, cladding creep, fuel densification and swelling, and cladding corrosion. Typical fabrication uncertainties considered are fuel OD, j .
cladding ID, cladding OD, fuel density, plenum size, and backfill pressure. !
Nuclear design uncertainties in the power, flux, and fluence are also l considered. The total uncertainty is obtained by a statistical convolution of ;
the individual uncertainties. The evaluations have been based upon the .
assumption that coolant chemistry is within the limits specified by ;
Westinghouse to ensure reliable plant operation.
l 5.2.2.1 Fuel Rod Performance Models !
Fuel rod performance results are obtained from the PAD fuel rod performance l code. The fuel performance evaluations for the Point Beach upgraded core features analysis have been done with the NRC-approved PA03.3 and PAD 3.4 fuel f performance models, references 5-1 and 5-2, respectively.
5.2.3 Fuel Rod Design Input !
Fuel Rod Fabrication Data: Standard nominal values for Westinghouse 14x14 0FA and VANTAGE 5 fuel were used in the fuel rod design evaluations. The most i Ifmiting combination of VANTAGE 5 features, i.e., boron loading and axial (
power shape peaking due to the axial blankets, has been used. f I
Fuel Duty: The standard Westinghouse methodology for the fuel rod power histories used in fuel design evaluations was used. No single steady-itate p I
power history is limiting for all the design criteria, and appropriate power histories which define limiting duty for each of the crfteria were used. The l
Condition II transient power limits used in the fuel rod design evaluation f were updated to accrunt for the increased core power peaking factors. [
l
( Evaluations have shown that storage of fuel in the spent fuel pool for several {
i cycles, followed by re-insertion in the core, does not affect fuel performance. t i
- 5-4 I
Reactor System Conditions: The design evaluations were performed for the most limiting values of the coolant inlet temperature and flow rate. The Point Beach Nuclear Units have the option of operating at either 2000 or 2250 psi system pressure. The most limiting value of the system pressure for each of the design criteria evaluations was used.
5.3 FLe1 Rod Design Criteria Evaluation Results Evaluations of the rod internal pressure and cladding stress and strain criteria show that these design criteria will be satisfied for the desired increased allowable core power peaking factort and the desired fuel rod design features.
The cladding surface temperature criteria were shown to be satisfied for fuel rods operated through five cycles.
5.4 Conclusion Evaluation of the fuel rod design criteria affected by the upgraded core features for Point Beach Nuclear Plant, Units 1 and 2, has shown that these j design criteria will be satisfied.
REFERENCES 5-1. Miller, J. V., (Ed.), "Improved Analytical Models Used In Westinghouse Fuel Rod Design Computations," WCAP-8720, October 1976 (Proprietary) and WCAP-8785, October 1976 (Non-Proprietary).
5-2. Weiner, R. A., et al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-PA, (Proprietary) and WCAP-11873-A (Non-Proprietary), August 1988.
a 5-5
6.0 REACTOR PRESSURE VESSEL SYSTEM EVALUATIONS 6.1 Introduction The evaluations presented in this section were performed to ensure that eliminating thimble plugs from those core locations not already occupied by burnable absorbers, neutron source rods, or rod control clusters will not violate reactor pressure vessel internals system design requirements.
Thimble plugging devices are used to minimize the degree of core bypass flow passing through fuel assembly thimble tubes. These devices are inserted into the tops of those fuel assemblies not already occupied by rod control cluster assemblies (RCCAs), neutron source rods, or burnable neutron absorber rods.
Figure 6-1 presents a typical plugging device. The stainless steel rods are inserted about 6-8 inches into the fuel assembly. As such, they do not extend below the top of the active fuel region. The rods are bolted to a holddown device baseplate. The relatively high resistance flow path formed by the annular space between the rods and the fuel assembly thimble tubes n.11mizes the amount of flow that passes through these thimble tubes.
Thimble plug removal results in a reduction in core hydraulic resistance, and a related increase in the portion of core bypass flow passing through the fuel assembly thimble tubes. These direct consequences lead to secondary effects within the reactor pressure vessel internals system. Such effects were evaluated for fluid system pressure drops, core bypass flow, baffle gap coolant jetting momentum flux, closure head fluid temperature, internals component lift forces, and RCCA drop time.
The evaluations described herein were performed using operating, geometric, and hydraulic characteristics specific to Point Beach Units 1 and 2 with Westinghouse 14x14 optimized fuel and the thimble plugging devices removed.
Two different system pressures were considered (2000 and 2250 psia).
6-1
'l t
6.2 Internal Pressure
- asses [
Figure 6-2 presents a schematic view of the reactor pressure vessel and internals system common to Point Beach Units 1 and 2. The principal flow route through this system begins at the two inlet nozzles. At this point, flow turns downward through the reactor vessel / core barrel annulus. After passing the thermal shield within this downcomer region, the flow enters the lower reactor vessel dome region. This region is occupied by the internals energy absorber structure, lower support columns, bottom-mounted instrumentation columns, and supporting tie plates. From this region, flow passes upward through a diffuser plate, then through the lower core plate, and into the core region. After passing up through the core, the coolant flows into the upper plenum, turns, and exits the reactor vessel through the two outlet nozzles. Note that the upper plenum region is occupied by support columns and RCCA guide columns.
The first part of this evaluation was performod to assess the impact of thimble plug removal on the pressure losses experienced by the primary coolant as it flows through the principal flow route just described. By solving the various continuity and momentum flux equations that describe the system, flows and corresponding pressure drops across the parallel and series flow paths were calculated. This was performed for the following four cases:
- 1. Point Beach Units 1 & 2; thimble plugs in, system pressure = 2000 psia
- 2. Point Beach Units 1 & 2; thimble plugs in, system pressure = 2250 psia
- 3. Point Beach Units 1 & 2; thimble plugs out, system pressure = 2n00 psia
- 4. Point Beach Units 1 & 2; thimble plugs out, system pressure = 2250 psia The results of this evaluation show that the change in internals pressure drop associated with thimble plug removal occurs in the core region. This change, less than 10% of the estimated pressure drop through the core with thimble plugs in, applies to both Point Beach units. Both units are characterited by essentially the same set of internals series pressure drops. The total reactor vessel internal pressure drop (inlet nozzle to outlet nozzle inclusive) would decrease apprc,timately 5% due to the decrease in core pressure drop.
6-2
6.3 Core Bypass Flow Core bypass flow is defined as the total amount of reactor coolant flow which
]
bypasses the core region, and is not considered effective in the core heat transfer process. Consequently, the effect of increasing bypass flow is a reduction in core power capability. Previous Westinghouse work has shown that
- the previous total core bypass flow lim..,of 4.5% (of total reactor vessel flow) must be increased to 6.5% to accommodate the effects of thimble plug j removal. The purpose of this bypass flow evaluation is to ensure that the 6.5% limit, applicable when thimile plugs are removed, will not be violated.
The principal core bypass flow paths are:
l
] a. Baffle / Barrel Region
- The baffle / barrel region consists of vertical baffle plates that i follow the periphery of the core. They are joined to the core barrel
! by horizontal former plates spaced along the elevation of the baffle
} plates. At b h i Point Beach Units, all of the former plates have flow l holes machined in them. Some flow from the lower plenum enters this l region at the lower core plate and exits above the core. There will also be some flow axchange between the baffle / barrel region and the j core through the baffle plate gaps.
l
- b. Vessel Head Coolina Spray Norries l
l l These nozzles are flow paths between the reactor vessel and core l barrel annulus and the fluid volume in the vessel closure head region l above the upper support plate. A fraction of the flow that enters the
) vessel inlet nozzles and into the vessel / barrel downcomer passes through these nozzles and into the vessel closure head region.
f j c. Core Barrel - Reactor Vessel Outlet Nozzle Gap i
Some of the flow that enters the vessel / barrel downcomer will leak through the gaps between the core barrel outlet nozzles and the l
6-3
. a reactor vessel outlet nozzles and merge with the vessel outlet nozzle flow.
- d. Fuel Assembly - Baffle Plate Cavity Gap This is the core bypass flow path between the peripheral fuel assemblies and the core baffle plates,
- e. Fuel Assembly Thimble Tubes These tubes are physically part of each fuel assembly and flow within them is partially effective in removing core heat. Hovever, such flow is analytically not considered to be effective in heat rooval, and is consequently treated as core bypass flow. Elimination of thimble plugs at Point Beach Units 1 and 2 will result in a significant increase in this component of core bypass flow. This expected increase is responsible for increasing the total core bypass flow design limit from 4.5% to 6.5% of total reactor vessel flow.
Calculations were performed to estimate the impact of thimble plug removal on total core bypass flow. An iterative process was used to estimate bypass flow levels in the various bypass flow paths with the exception of the fuel assembly thimble tubes. The bypass flow through the fuel assembly thimble tubes was estimated separately and is discussed in Section 4.7.
The results of this evaluation indicate that there would virtually be no change to the sum of the core bypass flow through paths (a) through (d), as a result of thimble plug removal. In fact, this sum was estimated to decrease by 3%. This is logically consistent with the fact that thimble plus removal reduces core hydraulic resistance.
Also, the results indicate that revised total core bypass flow design limit of 6.5% will, indeed, be bounding.
6-4
( ,
p].
l
.' (
i' 6.4 Momentum Flux !
i 6 . Westinghouse has previously ptrformed a number of experimental tests to study j the interaction between coolar t jetting through core baffle plate gaps and the !
I vibratory response of fuel
- ods and assemblies near such gaps. These tests [
fndicated that there ar2 two vibration levels that can result in fuel rod damage. Lower levels of vibration amplitude can inflict damage in the fom of ,
- vibration wear at the rod / grid interface. Large amplitude vibration l (whirling), caused by fluid elastic instability, can result in fuel rod damage l due to cladding fatigue failure, rod-to rod contact, or even rod-to-baffle <
l plate wall contact, f
- i j The results of this evaluation show that the removal of fuel assembly thimble l l plugs does not adversely impact coolant jetting momentum flux through the >
l Point Beach Units 1 and 2 baffle plate gaps. This result was also etident in l l the core bypass flow evaluation results where it was shown that the e.hange in !
j the sum of the bypass flows in paths (a) through (d) was only 3%. Coupled !
with a reduced core hydraulic resistance, this tends to reduce baffle plate f j pressure differentials. ;
l 6.5 Closure Head riuid Tempc*ature l
The average temperature of the primary coolant fluid that occupies the reactor vessel closure head volume is an important initial condition for certain dynamic Loss-of-Coolant Accident (LOCA) analyses. Therefore, it was necessary to confirm that the removal of fuel assembly thimble plugs will not alter this average closure head temperature for Point Beach Units 1 and 2.
This confirmation was relatively simple, as it stemred from the evaluations used to assess core bypass flow. The interactico between all different flow paths into and out of the closure head region were modelled. Based upon +his interaction, the core bypass flow into the head region and average head fluid temperatures for different flow path conditions were calculated to bound expected normal reactor operation.
6-5
E.
, a The results show that the average closure head temprature will be essentially unaffected by thimble plug removal at Point Beach Units 1 and 2.
6.6 Hy&aulia Lif t Forces An evaluation of the effect of thimble plug removal on lift forces on the Paint Beach Units 1 and 2 reactor internals was performed to ensure that the
.nternals assembly would remain seated and stable, based upon the upper-bound reactor vessel flow rate for normal reactor operation. One function of the calculation performed is to estimate uplift forces on reactor internals compor,ents. Such forces ganerally are combined with other mechanical and body forces to evaluate the resultant preload of the core barrel flange against the reactor vessel.
The results of this e a .ation show, however, that thimble plug elimination at both Point Beach linits will actually incur an approximate 3.3% reduction in total lift force against the reactor internals. Consequently, it can be said without further calculatlon that thimble plug removal will certainly not impact reactor internals seating.
6.7 Rod Control Cluster Assembly Drop Times Consideration of the removal of thimble plugs from Point Beach Units 1 and 2 included an evaluation of the potential impact on the limit for rod control cluster assembly (RCCA) drop time-to-dashpot entry. This limit is 2.2 seconds for the Point Beach units with Westinghouse 14x14 0FA.
Actual RCCA orop times are governed by hydraulic lift ferces on the RCCA/ drive rod assembly, assembly weight, a:,sembly buoyancy force, and mechanical i friction. A mjor part of the total hydraulic resistance to RCCA drop motion occurs in the core region. The removal of thimble plugs reduces core hydraulic resistance and, therefore, core pressure drop. The Technical Specification drop time-to-dashpot entry limit is based upon an estimated upper bound normal rt -tar operation flow rate. Since this flow upper bourd is not impacted by thimble plug ren. oval, the Technical Specification RCCA drop time limit is not impacted by thimble plug elimination. With respect to 6-6
actual drop times under various operating conditions, the reduction in the core pressure drop associated with thimble plug elimination actually would
^
reduce the hydraulic litt force against a falling RCCA.
t Since thimble plug removal will not affect RCCA weight, buoyancy forces, and j mechanical friction, the net effect (albeit small) would be to lower actual f drop times. Most importantly, though, thimble plug elimination will not <
impact the currently applicable Technical Specification RCCA drop time limit ,
of 2.2 seconds for the Point Beach units with Westinghouse 14x14 0FA.
6.8 Conclusions l
The impact of thimble plug removal at Point Beach Units 1 and 2 on reactor
(
internal pressure losses, coolant jetting through core baffle plate gaps, and i closure head average fluid temperature would be essentially inconsequential. !
Thimble plug elimination at Point Beach Units 1 and 2 will not result in the f total core bypass flow exceeding the revised design limit of 6.5% of total j reactor vessel flow. l i
Thimble plug elimination at Point Beach Units 1 and 2 will result in an !
insignificant reduction in total reactor internals lif t forces. As such, [
internals lif t-of f . not an issue. !
Thimble plug elimination at Point Beach Units 1 and 2 will not impact the !
Technical Specification RCCA drop time-to-dashpot entry limit o.* 2.2 seconds (applicable to these units with Westinghouse 14x14 0FA),
l I
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wth FIGURE 6 2 SCHEMATIC VIEW OF TYPICAL REACTOR VESSEL 6-9
i 7.0 ACCIDENT ANALYSIS 1
This section addresses the impact of the upgraded core features (defined in Section 1.0) on the following analyses presented in Chapter 14 of the Point Beach Nuclear Plant FSAR: Non-Loss of Coolant Accidents (non-LOCA), small break LOCA, and steam generator tube rupture (SGTR).
t 7.1 Non-LOCA Accidents 7.1.1 Introduction This evaluation addresses the impact of the upgraded core features on non-LOCA ,
accident analyses. The analyses performed to support this evaluation are listed in Table 7.1. In addition, this section also addresses use of the Revised Thermal Design Procedure (RTOP) and the new Dropped Rod Methodology (references 7-1 and 7-2, respectively). Table 7-2 summarizes the nominal and initial values of plant parameters assumed in the non-LOCA safety analysis.
i 7.1. 2 The Effecta of an Increase in F3g An increase in the power-dependent F AH limit does not directly affect the system transient response of the non-LOCA events presented in the Point Beach '
FSAR. Rather, the power level-dependent F 3g limit is used in the determination of the DNBR for those events for which DNB is the safety acceptance criterion. (The F 3g is not relevant for the non-DNB-related non-LOCA events.)
ONBR calculations fall into two categories:
(a) Those events in which the power level-dependent value of FAH I5 indirectly accounted for via the core limits, and, (b) Those events which directly assume the power level-dependent value of F 3g in the analysis.
l l
l 7-1 l
( -- --_----- - .-__ _
For those events in the former category, revised core limits were generated, reflecting the increased F 3g limit of 1.70. Based upon the new core limits, !
new Overtemperature and Overpower Delta-T (OTDT/0PDT) setpoint equations havo
- been calculated. (See Figure 7-1.) The events which rely on the OTDT/0PDT setpoints for protection have been reanalyzed. These FSAR events include
- ,
Section Event 14.1.2 Uncontrolled RCCA Withdrawal at Power 14.1.6 Reduction in Feedwater Enthalpy Incident j 14.1.7 Excessive Load Increase Incident i 14.1.9 Loss of External Electricel Load t
The results of the analysis of these events show that the minimum DNBR value :
for each event is greater than the Safety Analysis Limit value. These events i are identified by their ANSI N18.2 classification (i.e., condition i designation) in the discussions that follow, i
7.1.2.1 Uncontrolled RCCA Withdrawal at Power r
i This Condition II event is analyzed at various power levels and reactivity i insertion rates, for both minimum and maximum reactivity Teedback cases. The transient is terminated by a High Neutron Flux or Overtemperature Delta-T f reactor trip. Using the new Overtemperature Delta-T setpoint and core limits, the new analysis shows that the ONBR value never falls below the Safety j Analysis Limit value. Therefore, the conclusions of the FSAR remain valid for this event.
I i 7.1.2.2 Reduction in feedwater Enthalpy Incident 1
This Condition II event is bounded by the Excessive Load Increase Incident (see ;
l 7.1.2.3). Based upon the results presented for the Excessive Load Increase !
Incident, the DNBR design basis has been met. Therefore, the conclusions of i
i 7-2 l t
I . .
7.1.2.3 Excessive Load Increase Incident '
For this Condition II event, cases are analyzed at beginning- and end-of-life conditions, with and without automatic rod cor. trol. In all cases, the transient approaches an equilibrium condition and a reactor trip does not result. Using the revised core limits, the new analysis shows that the DNBR value never falls below the Safety Analysis Limit value. Therefore, the [
conclusions of the FSAR remain valid for this event. !
7.1.2.4 Loss of External Electrical Load This Condition II event is analyzed at beginning- and end-of-life conditions, ,
with and without pressurizer control. Using the new Overtemperature Delta-T setpoint and new core limits, the new analysis shows that the DNBR value never falls below the Safety Analysis Limit value. In addition, the peak RCS i pressure and the peak secondary side pressure never exceed 110% of their ;
design values of 2500 psia and 1100 psia, respectively. Therefore, the conclusions of the FSAR remain valid for this event.
For those events in the second category (i.e. , those events which directly assume the power level-dependent value of F 3g in the analysis), the increased value for Fg was used in the analysis of the following FSAR events:
I Section Evert i
14.1.1 Uncontrolled RCCA Withdrawal from a Subcritical Condition 14.1.3 RCCA Drop i 14.1.5 Startup of an Inactive Reactor Coolant Loop f
14.1.8 Loss of Reactor Coolant Flow An increase in Fg results in a decrease in the DNBR value for a given set of thermal-hydraulic conditions. However, the results of the analysis of these events, assuming the revised limit value for FAH, show that the i
7-3 ;
I
_ _ - _ _ . . . , , - - ..~-.-----<-m-
calculated DNBR value for each event is greater than Safety Anelysis Limit value. Each of these events is discussed below.
7.1.2.5 Uncontrolled RCCA Withdrawal from a Subcritical Condition This Condition II event is defined as an uncontrolled addition of reactivity to the reactor core caused by a withdrawal of RCCAs. This results in a power excursion and increase in core heat flux. The DNBR calculation was performed, assuming the most limiting axial and radial power shapes associated with having the two highest combined-worth banks in their high-worth position. The results show that the DNBR never falls below the Safety Analysis Limit value.
Therefore, the DNB design basis has been met for this event.
7.1.2.6 RCLA Drop This event hetually consists of two separate events: the RCCA Drop event, and the Statically Hisaligned Rod event. Both of these events are classified ;
as Condition II events. The analysis of the RCCA Drop event was performed, using the methodology described in reference 7-2. The DNBR calculations for both events were performed, assuming the increased value for Fg of 1.70.
The results show that the DNBR never falls below the Safety Analysis Limit value. Therefore, the DNB design basis has been met for both events.
7.1.2.7 Startup of an Inactive Reactor Coolant Loop This Condition II event is dufined as an uncontrolled addition of reactivity to the reactor core caused by the startup of an inactive reactor coolant pump at 10% power. This results in a power excursion and increase in core heat flux. The DNBR calculation was performed, assuming the power level-dependent <
value of F 3g associated with the full power design limit of 1.70. The results show that the DNBR never falls below the Safety Analysis Limit value. I Therefore, the DNB design basis has been met for this event.
{
l.
l t
7-4 l
o .
7.1.2.8 Loss of Reactor Coolant Flow Three individual analyses make up this event: Complete Loss of Flow, Partial Loss of Flow, and Underfrequency loss of Flow. The Complete Loss of Flow is a l
Condition III event, while the other two are Condition II events. The DNBR calculations for all three events were performed, assuming an increased value for F 3g of 1.70. The results of each analysis show that the DNBR never falls below the Safety Analysis Limit value. Therefore, the DNB design basis has been met for all three events.
7.1.2.9 Steamline Break The Rupture of a Steam Pipe event in FSAR Section 14.2.5 is a Condition IV event. For the Core Response event, it is shown that the DNBR design basis is ;
met. The analysis is performed at zero power conditions, assuming the most :
reactive rod stuck in its fully withdrawn position. An increase of the power-dependent Fg limit results in an increase in the zero power stuck rod peaking factor. The impact of the increase in the zero power stuck rod peaking factor on the DNBR calcelation has been evaluated, and it has been shown that the DNBR design basis has been met. In addition, the increase in F3g will not change the primary-to-secondary heat transfer characteristics in the Mass and Energy Release to Containment event. Therefore, this event is I not impacted by the increase in F 3g. l Given the discussion above, the increase in the power-dependent F3g limit is acceptable, and the conclusiont presented in the FSAR Chapter 14 non-LOCA safety analyses remain valid. ,
7.1.3 The Effects of an Increase in F q To ensure that cladding integrity and fuel melting at the "hot spot" are !
maintained within the applicable safety analysis limits, the two transients affected by an increase in the Fq limit were analyzed: j I,
t i
7-5 i
I
. s' FSAR ,
S_ection Event 14.1.8 Locked Rotor 14.2.6 Rod Ejection i The Locked Rotor event is classified as a Condition IV event. The results of the Locked Rotor analysis show that the maximum RCS pressure is 2744 psia, the j maximum cladding temperature is 2166'F, the amount of Zr-water reaction is 1.30% ;
by weight, and less than 86 percent of the fuel rods in the core undergo DNB. ;
Since the rsak RCS pressure reached during any of the transients is less than l that which.would cause stresses to exceed the faulted condition stress limits, (
the integrity of the primary coolant system is not endangered. Since the peak j cladding surface temperature calculated for the hot spot during the transient [
remains considerably less than 2700"F and the amount of zirconium-water reaction is small, the core remains in place and intact with no consequential loss of l core cooling capability. Therefore, all applicable safety criteria are met for ;
the Locked Rotor event, e
The Rod Ejection event is classified as a Condition IV event. The results of i the Rod Ejection analysis show that in all cases analyzed, the maximum cladding temperature is less than 2700'F, the maximum fuel stored energy is less than f
200 cal /ge, and the maximum fuel melt at the hot spot is less than 10%. There- [
fore, all applicable safety criteria are met for the Rod Ejection event.
Since the safety criteria have been met for the Locked Rotor and Rod Ejection !
events, the increase in Fgto 2.50 is acceptable with respect to the non-LOCA analyses. '
, i 7.1.4 Effects of Thimble Plug Deletion [
I Thimble plugs are currently installed in all fuel assemblies which are not (
[
under RCCA locations or are not equipped with sources or burnable poisons. f Placement of the plugs in these locations serves to limit the flow through the !
guide thimble tubes. The removal of the plugs will allow coolant flow through .
the guide thimble tubes, thus reducing the amount of flow available for core f heat removal. This is reflected in the increase in the core bypass flow h F
7-6 i
! o e assumed in the safety analyses. The events reanalyzed (see Table 7-1) have incorporated the effects of the increase in core bypass flow. The Steamline Break event was not reanalyzed. However, an increase in core bypass flow and the resultant reduction in core flow would reduce the severity of the core cooldewn in this transient. This would result in a lower peak heat flux, which is a benefit with respect to DNB. Therefore, the increase in core bypass flow would not invalidate the conclusions of the Steamline Break, Core Response event. In addition, the reduction in core flow would not significantly change the primary to secondary heat transfer characteristics of the Mass and Energy Release to Containment event. Therefore, thimble plug removal will not impact the Steamline Break, Mass and Energy Release to Containment event.
The removal of the thimble plugs would also have a slight impact on the vessel prest.ure drops. The effects of this change have been incorporated into those events which were reanalyzed. For the Steamline Break events, the change in vessel pressure drops would have an insignificant impact on the results of the event. Therefore, the change in the vessel pressure drops would not invalidate the conclusions of the Core Response event, nor would this change impact the Mass and Energy Release to Containment event.
7.1. 5 Utilization of Revised Thermal Design Procedure Methodology The Revised Thermal Design Procedure (RTOP) methodology involves the extension of the Improved Thermal Design Procedure (ITOP) methodology. RTOP differs from ITOP in that it statistically combines both the ONB correlation uncertainty and the system uncertainties to predict the ONBR limit value (see reference 7-1). The Safety Analysis ONBR limit value is 1.33 for both thiaible and typical cells.
The events which used RTOP in the new analysis are shown in Table 7-1. In all cases, it was shown that the minimum DNBR was above the Safety Analysis Limit value. Therefore, DNB criteria for these events have been met. Events which did not use RTOP continued to use the Standard Themal Design Procedure (STOP) and the W-3 correlation where applicable.
7-7
7.1.6 Utilization of New Dropped Rod Methodology The new analysis of the Dropped Rod event used the methodology described in reference 7-2. The transient response to the dropped rod event was analyzed assuming no protective action is provided by a reactor trip, turbine load runback or automatic rod withdrawal block. Nuclear models were used to obtain a hot channel factor consistent with the primary system conditions and reactor power. A DNB analysis was performed incorporating the primary conditions from the transient analysis and the hot channel factor from the nuclear analysis.
The results of the analysis show that the design basis was met. Therefore, neither a reactor trip from the reactor protection system, the turbine load runback feature, nor the rod withdrawl block is required for protection in a dropped rod event.
7.1.7 Effects of Upgraded Fuel Product Features I
7.1.7.1 Removable Top Nozzles and Debris Filter Bottom Nozzles !
Core flow areas and loss coefficients were preserved in the design of the RTN i and DFBN. As such, no parameters important to the non-LOCA safety analyses are impacted. Therefore, the conclusions of the non-LOCA safety analyses remain valid.
7.1.7.2 Integral Fuel Burnable Absorbers and Axial Blankets Axial blankets reduce power at the ends of the rod, which increases axial i I
peaking at the interior of the rod. Used alone, axial blankets reduce DNB and peak clad temperature margin, but the effect may be offset by the presence of IFBAt. which flatten the power distribution. The net effect on the axial shape is a function of the number and configuration of IFBAs in the core and [
time in life. The effect of axial blankets and IFBAs on the reload safety analysis parameters is taken into account in the reload design process. The axial power distribution assumptions in the safety analyses have been determined to be applicable for evaluating the use of axial blankets and IFBAs in the Point Beach units.
7-8
e O 7.1.7.3 Extended Burnup Reference 7-3 evaluates the impact of extended burnup on the design and operation of Westinghouse fuel. The major effect of the extended burnup rod design is on power shaping between fresh and burned assemblies. The effect of extended burnup on the reload safety analysis parameters is taken into account in the reload design process. The power distribution assumptions in the safety analyses have been determined to be applicable for evaluating the effects of extended burnup in the Point Beach units.
7.1.8 Low-Low-Leakage Loading Patterns and Part-Length Hafnium Absorbers The use of a low-low-leakage loading pattern and part-length hafnium absorbers will decrease the power at the periphery of the core. This will result in increased peaking factors. The reanalysis of the non-LOCA events has assumed an increase in F 3g to 1.70 and an increase in Fq to 2.50. Since all applicable rafety criteria were met with these assumptions, use of loading patterns ard core components that adhere to these new design limits is acceptable with respect to the non-LOCA safety analyses.
7.1.9 Miscellaneous Design Changes Rod Insertion Limits. A change to the rod insertion limits could impact the non-LOCA safety analyses in the following areas:
- b. Trip Reactivity
- c. Power Distribution Limits
- d. Ejected and Dropped Rod Worths
- e. Post Ejected Rod Peaking Factors
- f. Differential Rod Worth Nuclear design calculations have indicated that operation with the proposed insertion limits will ensure that the values for shutdown margin, trip reactivity, dropped rod worths, and differential rod worth previously assumed in the non-LOCA analyses will remain valid. The revised assumptions for power 7-9
O O h
distribution limits and rod ejection parameters used in the reanalysis of the events indicated in Table 7.1 have included the effect of the new rod insertion limits. In Lddition, the shutdown margin and power distribution >
assumptions used in the Steamline Break event remain valid with the proposed change to the rod insertion limits. Therefore, this change to the Point Beach 4
Technical Specifications is acceptable with respect to the non-LOCA safety analyses.
K(Z) Curve. Elimination of the third line segment of the K(Z) curve could impact the non-LOCA safety analyses assumptions for power distribution limits. However, nuclear design calculations have confirmed that the power distributions assumed in the reanalysis of the events in Table 7.1 will be ;
ensured with the adherence to the proposed K(Z) curve. Therefore, this change i to the Point Beach Technical Specifications is acceptable with resper,t to the !
non-LOCA safety analyses. ,
Flux Difference Operating Envelope. A change to the axial flux difftrence envelope could impact the non-LOCA safety analyses assumptions for pour ,
distribution limits. However, nuclear design calcolations have confirmd that i the power distributions assumed in the reanalysis of the events in Table .'.1 ,
i will be ensured with the adherence to the proposed Flux Difference Operating Envelope. Therefore, this change to the Point Beach Technical Specifications r is acceptable with respect to the non-LOCA safety analyses.
7.1.10 Results and Conclusions Using the revised safety analysis assumptions associated with the upgraded core features as documented in Table 7.2, the analyses and evaluations performed per ,
Table 7.1 show that all applicable safety criteria has been met. Therefore, i the conclusions of the non-LOCA safety analyses, as presented in Chapter 14 of ;
the Point Beach FSAR, remain valid.
l h
7-10 J
O *
- 7. 2 Small-Break LOCA Analysis 7.2.1 Introduction This section reports the results of an analysis of a small-break Loss-of-Coolant-Accident (LOCA) for the Point Beach Nuclear Plant, Units 1 and 2, assuming a 4-inch diameter cold leg break. The analysis incorporated the upgraded core features (as defined in Section 1.0), 25% uniform steam generator tube plugging, and reduced thermal design flow.
7,2.2 Methodology and Input Assumptions The analysis was performed, using the NRC-approved Westinghouse NOTRUMP Small Break Evaluation Model (reference 7-5) for a 4-inch break size. The Westinghouse NOTRUMP Emergency Core 'ooling System (ECCS) Small Break Evaluation Model, developed to deten ine the RCS response to design basis Small Break LOCAs, consists of the Nt TRUMP and LOCTA-IV computer codes, references 7-6 and 7-7, respectively The NOTRUMP code, a one-dimensional general network code with a number of advanced features, was used to calculate the system hydraulics throughout the transient. Among the features are the calculation of thermal non-equilibrium in all fluid volumes, flow r::gime-dependent drif t flux calculations with counter-current flow limitations, mixture level tracking logic in multiple-stacked fluid nodes, and regime-dependent heat transfer correlations. The LOCTA-IV code, which calculated the cladding thermal response, used the RCS pressure, fuel rod power history, steam flow past the uncovered part of the core, and mixture height history from the NOTRUMP code to determine the peak cladding temperature (PCT) during the small break LOCA.
Table 7.3 lists important input parameters and initial conditions used in the analysis. Major assumptions included a total peaking factor of 2.50, Fg of 1.70, 25% uniform steam generator tube plugging, thermal design flow of 85,200 gpm/ loop, and 102% of the core thermal power of 1518.5 MWt. The axial 7-11
0
- power distribution, shown in Figure 7-2, was chosen consistent with a LOCA Fq envelope of 2.50 at all core elevations. The pumped ECCS injection was assumed to be delivering to the RCS 25 seconds after the generation of a '
safety injection signal, which includes time required for diesel start-up, and loading of the ECCS pumps onto the emergency buses, pump start-up, and flow delivery. Minimum safeguards ECCS capability and operability was acsumed in the analysis, consistent with the loss-of-offsite power assumption aid the limiting singl<e failure of the loss of an entire train of pumped ECCS. due to the failure ef a diesel generator to start. Figure 7-3 shows the safe'y i injection flow from one high head safety injection (SI) pump, based upor.
performance curves degraded approximately 5% from the design head.
7.2.3 Results The analysis allows for the Point Beach units to operate at RCS pressures of either 2250 or 2000 psia. Since analyses at both pressures demonstrated the 2000 psia case to be limiting, the results shown here assume an RCS pressure of 2000 psia.
The analysis determined a PCT of 809*F for the 4-inch diameter cold-leg break, consistent with a LOCA qF envelope of 2.50 at all elevations. The transient was terminated when the hot rod cladding temperature was declinir.g dw to core recovery, the RCS total mass inventory was increasing, and no further core uncovery was calculated, due to accumulator injection and the SI flow exceeding the core boil-off rate due to decay heat removal. The PCT for the 4-inch break was well below the 2200*F Acceptance Criteria limit established by 10CFR50.46 and much lower than the value expected for the Large Break LOCA analysis being performed as part of the Best Estimate Evaluation Model Development program for 2-loop upper plenum injection plants.
1.2.4 Evaluation of Transition Core Impact The small-break LOCA was analyzed, assuming a full core of 14x14 0FA fuel with j
the upgraded fuel product features described in Section 2.2 to determine the PCT. This is consistent with the methodology employed in WCAP-10444-P-A l
l 7-12 t
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m
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l (reference 7-4) for VANTAGE 5 transition. Sensitivity studies performed using the WFLASH evaluation model, reference 7-8, have previously shown that 0FA fuel is more limiting than STD fuel in calculated ECCS performance. For the small-break LOCA, the eff2ct of the fuel difference is most pronount;ed 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 WFLASH studies regarding the fuel difference may be extended to this NOTRUMP analysis, i
When assessing the LOCA impact of transition cores, it must be determined whether the transition core can have a greater calculated PCT than either a complete core of the reference fuel design or a complete core of the improved fuel design. For a given peaking factor, the only mechanism available to cause a transition core to have a greater calculated PCT than a full core of either fuel is the possibility of flow redistribution due to fuel assembly hydraulic resistance mismatch. This hydraulic mismatch might exist only for transition cores and is the only unique difference between a complete core of [
either fuel type and a transition core.
The NOTRUMP computer code, reference 7-6, is used to model the core hydraulics during a s.nall-break event. Since the core flow during a small break is relatively slow, providing enough time to maintain flow equilibrium between fuel assemblies (i.e., no crossflow), only one core flow channel is modeled in the NOTRUMP code; therefore, hydraulic resistance mismatch is not a factor for the small-break event. Thus it is not necessary to perform a small-break evaluation for transition cores, and it is sufficient to reference the small-break LOCA for the complete core of the 14x14 0FA fuel with upgrade features as bounding for all transition cycles.
7.2.5 Summary This analysis demonstrates that the ECCS satisfies the acceptance criteria of 10CFR50.46 for a 4-inch diameter cold leg break.
7-13
That ist
- a. The calculated peak fuel element cladding temperature is below the requirement of 2200'F.
- b. The amount of fuel element cladding that reacts chemically with water or steam does not exceed one percent of the total amount of zircaloy in the reactor.
- c. The localized cladding oxidation limit of 17 percent is not exceeded during or after quenching.
i
- d. The core remains amenable to cooling during and af ter the break.
- e. The core temperature is reduced, and decay heat is removed for an extended period of time. This is required to remove the heat from the long-lived radioactivity in the core.
The small-break LOCA analysis for the Point Beach Nuclear Plant (Units 1 and 2), using the NOTRUMP evaluation model, resulted in a PCT of 809'F for the 4-inch diameter cold leg break case for an F envelope q based upon a total peaking factor of 2.50. The maximum local metal-water reaction was 0.07 percent, and the total metal-water reaction was less than 0.3 percent. The cladding temperature turned around at a time when the core geometry was still amenable to cooling. Criterion e, is addressed separately in a specific evaluation for each relcad cycle, i Mixed core hydraulie resistance mismatch is not a significant factor for small-a break LOCA analysis. Therefore, it is not necessary to perform any additional small-break evaluations for transition cores, and it is suf ficient to reference the small-break LOCA applicable to the complete core of the 14x14 0FA wtth upgraded core features, as bounding all transition cycles.
7-14
7.3 Steam Generator Tube Rupture Accident 7.3.1 Introduction ,
The steam generator tube rupture (SGTR) analysis for the Point Beach units was performed to evaluate the radiological consequences of a SGTR accident. A complete single tube break adjacent to the steam generator tube sheet was ,
assumed for the SGTR analysis. Since the RCS pressure is greater than the steam generator shell side pressure, radioactive reactor coolant is discharged into ,
the secondary system. For the Point Beach units, the major factors that affect the resultant offsite doses are the amount of fuel defects (level of reactor coolant contamination), the primary-to-secondary mass transfer through ,
the ruptured tube, and the steam released from the ruptured steam generator to the atmosphere. !
The proposed changes at Point Beach which were analyzed with respect to the FSAR SGTR analysis include the upgraded core features (as defined in Section 1.0) and 25% uniform steam generator tube plugging. Analyses were performed !
to bound operation at RCS pressures of 2000 and 2250 psia. Since the l' conservative assumption of 1% defective fuel is not impacted by the proposed changes, the variables which do impact the offsite radiation doses calculated for the FSAR SGTR analysis include the primary-to-secondary break flow and the ,
steam released from the ruptured steam generator to the atmosphere. Therefore, !
sensitivity SGTR analyses were performed to assess the impact of the proposed changes on the primary-to-secondary break flow and steam released to the atmosphere via the ruptured steam generator. The results of the sensitivity -
analyses were then used to determine the change to the offsite radiation doses ;
reported in the FSAR for the SGTR accident. [
7.3.2 Methodology f i
The results of the SGTR analysis in the Point Beach FSAR indicate that 70,000 !
lbs of RCS coolant would be discharged to the secondary side through the ruptured tube and 30,000 lbs of steam would be released to the atmosphere via the ruptured steam generator during the 30 minute interval considered.
7-15 t
However, the FSAR results reflect an assumption that operator action would be taken to shut off one (1) high head safety injection pump when water level returns to the pressurizer. This assumption is not consistent with the current SGTR recovery procedures. Specifically, the current SGTR recovery procedure for the Point Beach units does not include provisions for reducing safety injection flow until several requisite actions are completed and then safety injection flow is completely terminated. This is a necessary step to terminate the primary to secondary break flow. To conservatively model the SGTR accident consistent with SGTR recovery procedures, it is assumed that full safety injection flow is maintained to the RCS from the time of safety injection initiation until 30 minutes after the tube rupture, when the RCS and ruptured steam pressure are assumed to equilibrate, and break flow is assumed to be terminated. Therefore, the Point Beach FSAR analysis was re-evaluated to incorporate this assumption to provide a more accurate basis for performing the sensitivity studies.
The new base SGTR analysis was performed using the methodology and assumptions similar to the original FSAR SGTR analysis, with the exception of the safety injection termination, as explained above. The accident analyzed is a complete single tube break. The subsequent reactor coolant loss via the ruptured tube leads to RCS depressurization and a decrease in the pressurizer level. Reactor trip and safety injection are assumed to occur simultaneously at the low pressurizer pressure safety injection signal. A loss of off-site power is assumed to occur at the time of reactor trip. After safety injection actuation, flow from both high head safety injection pumps is assumed to be injected into the RCS until 30 minutes after the accident initiation, at which time it is assumed that the operator actions to terminate the break flow are completed. For the primary-to-secondary break flow to be terminated, the following actions must be performed:
- a. identify and isolate the affected steam generator
- b. cool down the RCS using the intact steam generator to provide subcooling margin
- c. depressurize the RCS to restore coolant inventory
- d. terminate safety injection.
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These actions are assumed to be completed by 30 minutes.
- For the determination of primary-to-secondary break flow prior to reactor trip and safety injection actuation, the RCS depressurization rate due to the !
primary-to-secondary break flow was calculated. The average break flow during the time period from tube rupture initiation to reactor trip was used to calculate an integrated break flow for this period. After reactor trip, the break flow rate is assumed to equilibrate to the point where the safety injection flow rate is exactly balanced by the outgoing break flow rate. This resultant equilibrium break flow rate is assumed to persist until safety
- injection is terminated at 30 minutes. These integrated break flow values are then summed to yield the total primary-to-secondary break flow for the 30 minutes.
Since off-site power is assumed to be lost coincident with reactor trip, the coMenser steam dump system would not be operable. Following reactor trip, the steam generator pressure increases rapidly due to the automatic turbine trip (at reactor trip) and lack of normal steam dump via the condenser.
Therefore, steam is relieved through the steam generator relief valves to dissipate the plant residual heat and the core decay heat. For the SGTR analysis, it was assumed that the steam is relieved via the safety valves and that the steam generators are maintained at the lowest safety valve setpoint. A mass and energy balance for the primary and secondary systems was used to calculate the steam released via the safety valve on the ruptured steam generator for the 30-minute time period considered.
To determine the impact of the proposed changes on the Point Beach SGTR, several sensitivity analyses were completed. Specifically, the relevant changes to the Point Beach parameters were made and the primary-to-secondary break flow and steam released via the ruptured steam generator for the 30-minute time period were recalculated.
The results of the Point Beach SGTR Base Case and the limiting sensitivity results are presented in Section 7.3.3.
7-17
e s 7.3.3 Break Flow and Mass Release Results The results for the Point Beach SGTR base case indicate that for the 30-minute time period considered, the primary to secondary break flow totaled 91,117 lbs, and the steam released via the ruptured steam generator was 54,321 lbs.
These results are greater than those reported in the FSAR due to the different analysis assumptien discussed previously.
The sensitivity analyses for the proposed changes at the Point Beach units show an increase in the primary-to-secondary break flow to 92,483 lbs and an increase in the steam released via the ruptured steam generator to 57,081 lbs. These are increases of 1.5% and 5.1%, respectively. Note that these increases are for the combined effect of the upgraded core features, 25%
uniform steam generator tube plugging, and RCS operating pressures of 2000 and 2250 psia. The results of the sensitivity analyses indicate that the 25%
steam generator tube plugging assumption was the foremost contributor to the increase in break flow and mass release.
These results, applicable to both Point Beach units, have been used to calculate the offsite doses for the revised Point Beach base case SGTR and also to determine the effect of the proposed changes at the Point Beach units.
7.3.4 Radiological Corsequences of a Steam Generator Tube Rupture The evaluation of the radiological consequences of an SGTR assumes that the i
reactor has been operating with a small percent of defective fuel and leaking l
generator tubes for sufficient time to establish equilibrium concentrations of radionuclides in the reactor coolant and in the secondary coolant. Radionu-i l clides from the primary coolant enter the steam generator via the ruptured tube and are released to the atmosphere through the steam generator safety or power-operated relief valves.
The radioactivity released to the environment due to an SGTR depends upon primary and secondary coolant activity, primary-to-secondary break flow, partitioning of elemental iodine activity between the steam generator liquid 7-18
r o e and steam, and the mass of steam discharged to the environment. All of these parameters were evaluated for a design basis failure of a single steam generator tube.
Assumptions The radiological assumptions used in this analysis are consistent with those used in the FSAR analysis, with the exception of the primary-to-secondary leak rate and the corresponding secondary coolant iodine activity. Specifically, the leak rate assumed in this analysis is 500 gpd (0.35 gpm) which is consistent with the plant Technical Specifications, rather than the 10 gpm leak rate assumed in the FSAR analysis,
- a. The iodine and noble gas concentrations in the reactor coolant are based on 1% fuel defects at 1520 Mwt as presented in FSAR Table 9.2-5.
- b. The equilibrium iodine concentrations in the secondary coolant are based on 1% defects in the primary coolant and 500 gpd (Technical Specification limit) primary-to-secondary leakage per generator as presented in FSAR Figure 14.2.5-22.
- c. Offsite power is not available.
- d. Atmospheric dispersion factors - FSAR Table 14.2.5-3 0-2 hr. Site boundary - 2.3 x 10'4 sec/m 3 Low population zone - 2.6 x 10 -5 2-4 hr. Low population zone - 1.3 x 10 -5
- e. Breathing rate - 3.47 x 10'4 m 3/sec
- f. Thyroid dose conversion factors - TID-14844 7-19
} l !
b, g. Elemental iodine partition coefficient for intact and faulted i n ,
generators - 0.1 i
- h. Iodine chemical species - elemental [
i
- 1. Mass transfer T
Base Case [
Re-evaluation Sensitivity Case [
f Steam released to atmosphere (1bs) !
l 0 - 2 hr 4
J Intact SG's 5.43 x 10 4 5.71 x 10 4
Faulted SG 5.43 x 10 4 5.71 x 10 ;
2 - 4 hr Intact SG's 2.70 x 10 5 2.70 x 10 5 (
! Faulted SG 0 0 !
I
, l Break flow (faulted SG - lbs) !
4
- O - 30 min 9.11 x 104 9.25 x 10 [
- l
! Primary to secondary leakage [
(both cases) 0.35 gpm (500 gpd) per SG f i
- 7.3.5 Doses to Receptor at the Site Boundary and Low Population Zone '
! Outer Boundary [
P
[
The potential radiological consequences resulting from a postulat.d SGTR have been analyzed using the assumptions described. The whole-body dose due to immersion and the thyroid dose due to inhalation have been analyzed for the 0 to 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> dose at the site boundary (S.B.) and for the duration of the [
accident (0-4 hr) at the low population zone outer boundary (LPZ). The doses f for the FSAR. Base and Sensitivity Cases are summarized in Table 7.4. !
I t
l
- 7-20 i
_ - _ - - _ . - - - - . . - - - - w
g e .
4 It should be noted that the resulting offsite doses are less than those reported in the FSAR.
The 10CFR100 guideline values are 300 rem thyroid and 25 rem whole body. All of the doses calculated for the SGTR are within a "small fraction" of the 10CFR100 exposure guideline. This "small fraction" is defined as 10% of the guideline values, that is, 30 rem thyroid and 2.5 rem whole body and is the smallest of the exposure limits defined by the NRC in KUREG-0800, 7.3.6 Summary Based upon the results of the Point Beach SGTR re-evaluation and the sensitivity analyses, the proposed changes (upgraded core features and 25%
uniform steam generator tube plugging) to the Point Beach units will cause the primary-to-secondary breakflow to increase by 1.5% and the steam release via the ruptured steam generator to increase by 5.1%. The impact of these increases on offsite dose consequences has been evaluated. The thyroid and whole body dose will increase by 6.6% and 1.5%, respectively. These slight I increases do not change the conclusion that the Point Beach SGTR radiological consequences are within a small fraction of the limits set forth in 10CFR100.
Table 7.4 is included to summarize the results of the SGTR analyses.
i J
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1 l
i 7-21 i
REFERENCES 7-1. Friedland, A. J., Ray, S., "Revised Thermal Design Procedure,"
WCAP-11397 (Proprietary), and WCAP-11398 (Non-Proprietary) February 1987.
7-2. Haessler, R. L. et al., "Methodology for the Analysis of the Dropped Rod Event," WCAP-11394 (Proprietary), WCAP-11395 (Non-Proprietary).
April 1987.
";- 3 . Davidson, S. L., Kramer, W. R., ed., "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A (Proprietary), and WCAP-10126-NP-A (Non-Proprietary), December 1985.
7-4. Davidson, S.L., and Kramer W.L., Ed., "VANTAGE 5 Reference Core Report VANTAGE 5 Fuel Assembly," VCAP 10444-P-A (Proprietary),
and WCAP-10445-NP-A (Non-Proprietary), September 1985.
7-5. Lee, N., et, al., "West'nghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," WCAP-10054-P-A (Proprietary) and
( WCAP-10081-A (Non-Proprietary), August 1985.
I 7-6. Meyer, P. E.. "NOTRUMP, A Nodal Transient Small Break and General Network Code," WCAP-10079-P-A (Propriotary) and WCAP-10080-A (Non-Proprietary), August 1985.
7-7. Bordelon, F. M., et al., "LOCTA-IV Program: Loss of Coolant Transient Analysis," WCAP-8301, (Proprietary) and WCAP-8305, (Non-Proprietary), June 1974.
7-8. Esposito, V. J., Kesavan, K., and Maul, B. J.; "W-FLASH-A Fortran-IV Computer Program for Sim1ation of Transients in a Multi-Loop PWR,"
WCAP-8200 (Proprietary), July 1973, and WCAP-8261-R1 (Non-Proprietary),
July 1974.
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O 4 l
l TABLE 7.1 SUtetARY OF NON-LOCA EVENTS l
)
l THERMAL ,
FSAR ANALYSIS / DESIGN DNB EVENT SECTION EVALUATION PROCEDtTRE CORRELATION Uncontrolled RCCA Withdrawal 14.1.2 A STDP W-3 from a Suberitical Condition )
l l Uncontrolled RCCA Withdrawl 14.1.2 A RTDP WRB-1 l at Power j RCCA Drop 14.1.3 A RTDP WRB-1 )
Boron Dilution
- 14.1.4 A NA NA Startup of an Inactive Reactor 14.1.5 A STDP W3 Coolant Loop Reduction in Feedwater Enthalpy 14.1.6 E NA NA Incident Excessive Load Increase 14.1.7 A RTDP WRB-1 Incident Loss of Reactor Coolant flow 14.1.8 A RTDP WRB-1 j Locked Rotor 14.1.8 A STDP NA l Loss of External Electrical 14.1.9 A RIDP WRB+1 Load Loss of Normal Feedwater* 14.1.10 A STDP NA Loss of All AC Power to the 14.1.11 A STDP NA Auxiliaries
- Steamline Break; Core Response 14.2.5 E STDP W-3 M/E Release to Containment 14.2.5 E NA NA Rod Ejection 14.2.6 A STDP NA 1
! *Vhile not directly affected by the changes described in Section 7.1, these events were reanalyzed coincident with the other analyses.
l l
1 7-23 l
l _ _ __ ___ - _ -_______- - _ _ _ - -
TABLE 7.2 NON-LOCA SAFETY ANALYCIS AUSUMPTIONS Nominal Conditions RTDP non-RTDP NSSS Power, HWt 151f.5 1518.5 RCS Vessel Average Temperature, 'F 573.9 573.9 RCS Pressure, psia 2000/2250 2000/2250 RCS Flow Rate,6gpm 181800 178000 Steam Flow, 10 lbm/hr 6.61 6.61 Steam Pressure, psia 783 783 Feedwater Temperature, 'F 435.7 435.7 Core Average Heat Flux, Btu /hr-ft 2 185850 185850 Initial Conditions NSSS Vessel RCS Prz.
Power Average Flow Press.
Event (MWt) Temp (*F) {ggm] (psia) l Uncontrolled RCCA 0 547 81880 2000
! Withdrawal from a Suberitical Condition Uncontrolled RCCA 1518.5 573.9 181800 2000 Withdrawal at Power 911.1 563.1 181800 2000 151.8 549.7 181800 2000 i
RCCA Drop 1518.5 573.9 181800 2000 Boron Dilution At Power 1518.5 NA NA NA Startup 0 NA NA NA Refueling 0 NA NA NA Startup of an 182.2 550.2 81400 1970 Inactive Reactor Coolant Loop Excessive load 1518.5 573.9 181800 2000 Increase Incident loss of Reactor Coolant Flow 1518.5 573.9 181800 2000 Locked Rotor 1548.9 577.9 178000 2280 7-24
[_______________-__ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _
iw o
TABLE 7.2 (Cont')
NON-LOCA SAFETY ANALYSIS ASSUMPTIONS Initial Conditions (cont.)
Power ' Average Flow' Press.
Event (MWt) Teup (*F) {ggel .(psia)
Loss of External 1518.5 573.9 181800 2000/
.a - Electrical Load 2250 Loss of Normal 1548.9 577.9/ 378000 2280 Feedwater/ Loss of 569.9 AC Power to the Auxiliaries Rod Ejection 0 547 81880 1970 1548.9 577.9 178000 1970 7-:5
E j;. .
TABLE 7.3 INPUT ASSUMPTIONS USED IN THE SMALL BREAK LOCA ANALYSIS Parameter Value Reactor Core Power [a] 1518.5 MWt Peak' Linear Power [a] 14.25 kw/ft Peaking Factors FQt = 2.50 Fgy = 1.70 Power Shape Figure 7-2
! Fuel Assembly 14x14 0FA[c]
Nominal Accumulator Water Volume (per accumulator) 1118 ft 3 Minimum Accumulator Pressure 714.7 psia Pumped Safety Injection (HHSI) (one pump in operation) Figure 7-3 Steam Gener: tor Tube Plugging Level (uniform) 25%
Thermal Design Flow 85,200 gpm/ loop Taverage 570*F i
Reactor Coolant Pressure 2000 or 2250 psia i
[a] Two percent is added to this power for calorimetric error. Reactor coolant pump heat is not modeled.
[b] The analysis supports an elevation-independent FQ limit, i.e., flat K(2) cu rve.
(c) Upgraded fuel product features are bounded by the analysis.
i 7-26 e
. c.
TABLE 7.4 POINT BEACH SGTR RESULTS Base Case FSAR Re-evaluation Sensitivity SGTR Breakflow 70,000 lbs 91,117 lbs 92,483 lbs Steam Release via 30,000 lbs 54,321 lbs 57,081 lbs Ruptured SG SB Thyroid Dose 2.13 0.52 0.55 LPZ Thyroid Dose 0.38 0.06 0.065 SB Whole Body Dose 5.9 x 10 -2 1.15 x 10 -1 1.17 x 10 -1
-2 LPZ Whole Body Dose 6.7 x 10 ~3 1.29 x 10'2 1.31 x 10 Note: SB dose period is 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />.
LPZ dose period is 0 to 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
The sensitivity case considers tha e graded core features, uniform steam s,enerator tube plugging, and RC operating pressures of 2000 and 2250 psia.
l SB and LPZ dose units are in rems.
4 7~27
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OVERPOWER 65~ _ _ _\ g. ,, g ' ~ g AT TRIP g g . p ,, , g
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CVERTEMPERATURE g
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35- LOCUS OF POINTS \ \ \
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gC 6de 160 578 500 590 See GIO 620 650 T evg (Orl FIGURE 7-1 OVERTEMPERATURE AND OVERPOWER DELTA-T SETPOINT EQUATIONS I
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0 , g ; g ;
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- 5. I. FLOW (1b/sec)
FIGURE 7 3 HIGH HEAD S1 FLOW VERSUS RCS PRESSURE 7-30