ML18150A106

Proposed Tech Specs Supporting Planned Fuel Desing Change from Westinghouse Low Parasitic 15x15 (Lopar) Fuel Assembly to Modified 15x15 Optimized Fuel Assembly

ML18150A106
Person / Time
Site:	Surry
Issue date:	05/26/1987
From:	VIRGINIA POWER (VIRGINIA ELECTRIC & POWER CO.)
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Shared Package
ML18150A105	List: ML18150A104 ML18150A106
References
NUDOCS 8706010287
Download: ML18150A106 (58)

Text

e ATTACHMENT 1 Proposed Technical Specifications Changes

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PDR ADOCK 05000280 p PDR

TS 2.1-2

4. The reactor thermal power level shall not exceed 118% of rated power.

B. The safety limit is exceeded if the combination of Reactor Coolant System average temperature and thermal power level is at any time above the appropriate pressure line in TS Figures 2.1-1, 2.1-2 or 2.1-3; or the core thermal power exceeds 118% of the rated power.

Basis To maintain the integrity of the fuel cladding and prevent fission pro-duct release, it is necessary to prevent overheating of the cladding under all operating conditions. This is accomplished by operating within the nucleate boiling regime of heat transfer, wherein the heat transfer coefficient is very large and the clad surface temperature is only a few degrees Fahrenheit above the reactor coolant saturation temperature. The upper boundary of the nucleate boiling regime is

/

termed Departure From Nucleate Boiling (DNB) and at this point there is a sharp reduction of the heat transfer coefficient, which would result in high clad temperatures and the possibility of clad failure. DNB is not, however, an observable parameter during reactor operation.

Therefore, DNB has been correlated to thermal power, reactor coolant temperature and reactor coolant pressure which are observable parameters. This correlation has been developed to predict the DNB flux and the location of DNB for axially

I~

TS 2.1-3 uniform and non-uniform heat flux distributions. The local DNB heat flux ratio, DNBR, defined as the ratio of the DNB heat flux at a particular core location to the local heat flux, is indicative of the margin to DNB. The DNB basis is as follows: there must be at least a 95% probability with 95% confidence that the minimum DNBR of the limiting rod during Condition I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The correlation DNBR limit is based on the entire applicable experimental data set to meet this statistical criterion. (l)

The curves of TS Figure 2.1-1 which show the allowable power level decreasing with increasing temperature at selected pressures for constant flow (three loop operation) represent limits equal to, or more conservative than, the loci of points of thermal power, coolant system average temperature, and coolant system pressure for which the calculated DNBR is not less than the design DNBR limit or the average enthalpy at the exit of the vessel is equal to the saturation value. The area where clad integrity is assured is below these lines. The temperature limits are considerably more conservative than would be required if they were based upon the design DNBR limit alone but are such that the plant conditions required to violate the limits are precluded by the self-actuated safety valves. on the steam generators. The three loop operation safety limit curve allows for heat flux peaking effects due to fuel densification and applies to 100% of design flow. The effects of rod bowing are also considered in the DNBR analyses.

The curves of TS Figure 2.1-2 and 2.1-3 which show the allowable power level decreasing with increasing temperature at selected pressures for constant flow (two loop operation), represent limits equal to, or more

e TS 2.1-4 conservative, than the loci of points of thermal power, coolant system average temperature, and coolant system pressure for which either the calculated DNBR is equal to the design DNBR limit or the average enthalpy at the exit of the core is equal to the saturation value. At low pressures or high temperatures the average enthalpy at the exit of the core reaches saturation before the calculated DNBR reaches the design DNBR limit and, thus, this arbitrary limit is conservative with respect to maintaining clad integrity. The plant conditions required to violate these limits are precluded by the protection system and the self-actuated safety valves on the steam generator. Upper limits of 70% power for loop stop valves open and 75% with loop stop valves closed are shown to completely bound the area where clad integrity is assured. These latter limits are arbitrary but cannot be reached due to the Permissive 8 protection system setpoint which will trip the reactor on high nuclear flux when only two reactor coolant pumps are in service.

Operation with natural circulation or with only one loop in service is not allowed since* the plant is not designed for continuous operation with less than two loops in service.

N TS Figures 2.1-1 through 2.1-3 are based on a FLrn. of 1. 55, a 1.55 cosine axial flux shape and a DNB analysis procedure including 4

t h e f ue 1 d ens1"f"1cat1on

• power sp1"k"1ng ( ) as part o f t h e generic .

5 6 margin to accommodate rod bowing. ( )( ) TS Figure 2.1-1 is also valid for the following limit of the enthalpy rise hot channel N

factor: F~H = 1.55 (1 + 0.3 (1-P)) where P is the fraction of rated power. TS Figures 2.1-2 and 2.1-3 include a 0.2 rather than 0.3 part power multiplier for the enthalpy rise hot channel factor.

These hot channel factors are higher than those calculated at full power over the range between that of all control rod assemblies fully withdrawn to

TS 2.1-5 maximum allowable control rod assembly insertion. The control rod assembly insertion limits are covered by Specification 3.12.

Adverse power distribution factors could occur at lower power levels because additional control rod assemblies are in the core; however, the control rod assembly insertion limits dictated by TS Figures 3.12-lA (Unit 1) and 3.12-lB (Unit 2) ensure that the DNBR is always greater at partial power than at full power.

The Reactor Control and Protection System is designed to prevent any anticipated combination of transient conditions for Reactor Coolant System temperature, pressure and thermal power level that 3

would result in a DNBR less than the design DNBR limit( ) based on steady state nominal operating power levels less than or equal to 100%, steady state nominal operating Reactor Coolant System average temperatures less than or equal to 574.4°F and a steady state nominal operating pressure of 2235 psig. Allowances are made in initial conditions assumed for transient analyses for steady state errors of +2% in power, +4°F in Reactor Coolant System average temperature and +/-30 psi in pressure. The combined steady state errors result in the DNB ratio at the start of a transient being 10 per cent less than the value at nominal full power operating conditions. The steady state nominal operating parameters and allowances for steady state errors given above are also applicable for two loop operation except that the steady state nominal operating power level is less than or equal to 60%.

The fuel overpower design limit is 118% of rated power. The over-power limit criterion is that core power be prevented from reaching a value at which fuel pellet melting would occur. The value of 118% power allows substantial margin

e TS 2.3-8 will prevent the minimum value of the DNBR from going below the applicable design limit during normal operational transients and anticipated transients when only two loops are in operation and the overtemperature ~T trip setpoint is adjusted to the value specified for three-loop operation. During two-loop operation with the loop stop valves in the inactive loop open, and the overtemperature ~T trip setpoint is adjusted to the value specified for this condition, a reactor trip at 60% power will prevent the minimum value of DNBR from going below the applicable design limit during normal operational transients and anticipated transients when only two loops are in operation. During two-loop operation with the inactive loop stop valves closed and the overtemperature ~T trip setpoint is adjusted to the value specified for this condition, a reactor trip at 65% power will prevent the minimum DNBR from going below the applicable design limit during normal operational transients and anticipated transients.

Although not necessary for core protection, other reactor trips provide additional protection. The steam/feedwater flow mismatch which is coincident with a low steam generator water level is designed for and provides protection from a sudden loss of the reactor's heat sink. Upon the actuation of the safety injection circuitry, the reactor is tripped to decrease the severity of the accident condition. Upon turbine trip, at greater than 10% power, the reactor is tripped to reduce the severity of the ensuing transient.

References (1) FSAR Section 14.2.1 (2) FSAR Section 14.2 (3) FSAR Section 14.5 (4) FSAR Section 7.2 (5) FSAR Section 3.2.2 (6) FSAR Section 14.2.9 (7) FSAR Section 7.2

TS 3.1-Sa

b. With one Reactor Vessel Head vent path inoperable; startup and/or power operation may continue provided the inoperable vent path is maintained closed with power removed from the valve actuator of both isolation valves in the inoperable vent path.

c. With two Reactor Vessel Head vent paths inoperable; maintain the inoperable vent path closed with power removed from the valve actuator of all isolation valves in the inoperable vent paths, and restore at least one of the vent paths to operable status within 30 days or be in hot shutdown within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> and in cold shutdown within the following 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.

Basis Specification 3.1.A-l requires that a sufficient number of reactor coolant pumps be operating to provide coastdown core cooling flow in the event of a loss of reactor coolant flow accident. This provided flow will maintain the DNBR above the applicable design limit. (l) Heat transfer analyses also show that reactor heat equivalent to approximately 10% of rated power can be removed with natural circulation; however, the plant is not designed for critical operation with natural circulation or one loop operation and will not be operated under these conditions.

When the boron concentration of the Reactor Coolant System is to be reduced, the process must be uniform to prevent sudden reactivity changes in the reactor. Mixing of the reactor coolant will be sufficient to maintain a uniform concentration if at least one reactor coolant pump or one residual heat removal pump is running while the change is taking place. The residual heat removal pump will circulate the equivalent of the reactor coolant system volume in approximately one half hour.

TS 3.12-8

~T and Overtemperature ~T trip settings shall be reduced by the equivalent of 2% power for every 1% quadrant to average power tilt.

c. Inoperable Control Rods

1. A control rod assembly shall be considered inoperable if the assembly cannot be moved by the drive mechanism or the assembly remains misaligned from its bank by more than 12 steps.

Additionally, a full length control rod shall be considered inoperable if its rod drop time is greater than 2.4 seconds to dashpot entry.

2. No more than one inoperable control rod assembly shall be permitted when the reactor is critical.

3. If more than one control rod assembly in a given bank is out of service because of a single failure external to the individual rod drive mechanism (i.e., programming circuitry), the provisions of Specifications 3.12.C.l and 3.12.C.2 shall not apply and the reactor may remain critical for a period not to exceed two hours provided immediate attention is directed toward making the necessary repairs. In the event the affected assemblies cannot be returned to service within this specified period, the reactor will be brought to hot shutdown conditions.

4. The provisions of Specifications 3.12.C.l and 3.12.C.2 shall not apply during physics tests in which the assemblies are intentionally misaligned.

5. The insertion limits in TS Figure 3.12-2 apply:

a. If an inoperable full-length rod is located below the 200 step level and is capable of being tripped, or

TS 3.12-13 in service, the effects of malpositioned control rod assemblies are observable from nuclear and process information displayed in the Main Control Room and by core thermocouples and in-core movable detectors. Below 50% power, no special monitoring is required for malpositioned control rod assemblies with inoperable rod position indicators because, even with an unnoticed complete assembly misalignment (full length control rod assembly 12 feet out of alignment with its bank), operation at 50% steady state power does not result in exceeding core limits.

The specified control rod assembly drop time is consistent with safety analyses that have been performed.

An inoperable control rod assembly imposes additional demands on the operators. The permissible number of inoperable control rod assemblies is limited to one in order to limit the magnitude of the operating burden, but such a failure would not prevent dropping of the operable control rod assemblies upon reactor trip.

Two criteria have been chosen as a design basis for fuel performance related to fission gas release, pellet temperature, and cladding mechanical properties. First, the peak value of fuel centerline temperature must not exceed 4700°F. Second, the minimum DNBR in the core must not be less than the applicable design limit in normal operation or in short term transients.

ATTACHMENT 2 Safety Evaluation For Change To Fuel Assembly Design

e

1.0 INTRODUCTION

AND

SUMMARY

Surry Units 1 and 2 have been operating with Westinghouse 15x15 low-parasitic (LOPAR) fueled cores. For Surry Unit 1 Cycle 10 and Unit 2 Cycle 10 and subsequent cycles, it is planned to refuel both units with 15x15 Surry Improved Fuel (SIF) regions supplied by the Westinghouse Electric Corporation. As a result, future core loadings would range from approximately a 1/3 SIF - 2/3 LOPAR mixed core to eventually a 100% SIF fueled core. The 15x15 SIF assembly has the same intermediate Zircaloy grids and fuel rod diameter as the Westinghouse 15x15 VANTAGE 5 assembly and the 15x15 optimized fuel assembly (OFA). In addition, the SIF assembly incorporates two additional VANTAGE 5 features, an increased high discharge burnup capability and a reconstitutable top nozzle. The SIF assembly fuel rods retain the use of the standard bullet nose end plug, and the fuel rods are slightly longer than the 15x15 LOPAR and 15x15 OFA fuel rods. The SIF assembly will decrease neutron parasitic capture and thereby permit more efficient fuel usage.

Significant operating experience on Zircaloy clad fuel with Zircaloy grids has been obtained from a number of commercial reactors operating with regions of OFAs with 14x14, 15x15 and 17x17 arrays and some high burnup 14x14 and 17x17 demonstration OFAs. Approximately nine regions (five of which were 15x15 fuel arrays) have completed one cycle of irradiation, and two regions (14x14 and 17x17 fuel arrays) have completed two cycles of irradiation. No Zircaloy grid abnormalities or fuel rod abnormalities related to the Zircaloy grids have been observed. Ten demonstration OFAs (14x14 and 17x17 fuel arrays) have been irradiated up to 40,000 MWD/MTU 43-NPW-2016S

e e assembly burnups. These demonstration assemblies were in good condition with the exception of one 14x14 assembly (Reference 1). This one assembly had failed rods due to fretting wear at the bottom Inconel grid caused by an improper manufacturing procedure which has no generic implication for the OFA (or SIF) design.

Four VANTAGE 5 demonstration assemblies (17x17) were loaded into the V.C.

Summer Unit 1 Cycle 2 core and began power production in December of 1984.

These assemblies completed one cycle of irradiation in October of 1985 with an average burnup of 11,357 MWD/MTU. Post-irradiation examinations showed the four demonstration assemblies were of good mechanical integrity. No mechanical damage or wear was evident on any of the VANTAGE 5 components.

The four demonstration assemblies were reinserted into V.C. Summer 1 for a second cycle of irradiation. This cycle was completed in March of 1987, at which time the demonstration assemblies had achieved an average burnup of 30,000 MWD/MTU.

In addition to the fuel assembly modifications, thimble plug assemblies will be removed from Surry Unit 1 Cycle 10, Surry Unit 2 Cycle 10 and subsequent cycles.

2.0

SUMMARY

AND CONCLUSIONS Consistent with the Westinghouse and Virginia Electric and Power Company standard reload methodologies for analyzing cycle specific reloads (References 2 and 3), parameters are chosen to maximize the applicability 43-NPW-2016S

e e of the transition evaluations for future cycles and to facilitate subsequent determinations of the applicability of 10 CFR 50.59. The mechanical, thermal/hydraulic, nuclear and accident evaluations considered the transition core effects described for mixed cores in Chapter 18 of Reference 4 and the impact of thimble plug removal.

The transition design and safety evaluations consider the following nominal conditions which are consistent with 100% of FSAR thermal design flow.

These nominal conditions are 2441 MWt core power, 2250 psia system pressure, 543.0°F core inlet temperature, and 265,500 gpm RCS thermal design flow.

The results of evaluation/analyses and tests described herein lead to the following conclusions:

1. The Westinghouse SIF assemblies are mechanically and hydraulically compatible with the current LOPAR fuel assemblies, control rods and reactor internals interfaces.

2. Commercial operating and/or demonstration experience with 14x14, 15x15 and 17x17 OFAs and 17x17 VANTAGE 5 assemblies containing Zircaloy grids provides evidence of satisfactory operation of 15x15 SIF Zircaloy grids.

3. Demonstration experience with 17x17 VANTAGE 5 assemblies containing removable nozzles provides evidence of satisfactory operation of 15x15 SIF removable nozzles.

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4. Thimble plug removal will increase the core by-pass flow slightly from 4.5% to 6%. This value will be used in the transient and thermal/hydraulic analyses.

5. Changes in the nuclear characteristics due to the transition from LOPAR to SIF will be within the range normally seen from cycle-to-cycle due to fuel management effects.

6. Plant operating limitations given in the Technical Specifications will be satisfied with the proposed changes noted in Section 7.0 of this report.

3.0 MECHANICAL EVALUATION The SIF assemblies have been designed to be compatible with the LOPAR assemblies, reactor internals interfaces, and fuel handling and refueling equipment. Figure 1 gives a side-by-side comparison of the fuel assemblies. The grid elevations for the two assembly designs match, minimizing mechanical and hydraulic interaction. The assembly envelopes, fuel rod diameters, and the top and bottom Inconel grids are the same.

Changes from the LOPAR to the SIF assembly design include: decreases in guide thimble and instrumentation tube diametral dimensions, the change from use of the five intermediate LOPAR Inconel grids to a SIF Zircaloy grid design, the change to the SIF removable top nozzle (RTN) with modified holddown spring, the change to a low profile SIF removable bottom nozzle, and the increased lengths for the SIF fuel rod and fuel assembly structure.

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e e The 15x15 SIF assembly guide thimbles are similar in design to their counterparts in the LOPAR fuel assemblies except for a 13 mil ID and OD reduction above the dashpot and an increased length due to an increase in fuel assembly length. The diameter reduction is due to a reduced grid cell. This results from use of thicker grid strap material for the Zircaloy grids. Below the dashpot the SIF and LOPAR assembly guide thimble dimensions are identical. The SIF guide thimble tube ID continues to provide an adequate nominal diametral clearance for the control rods and other core components. However, due to the reduced clearance, the time to the dashpot for accident analyses has conservatively been determined to increase from 1.8 seconds (to dashpot) for the LOPAR assemblies to 2.4 seconds for the SIF assemblies. The increase in rod drop time required accident reanalyses as described in Section 6.0 of this report.

The five intermediate SIF Zircaloy grids have thicker and wider straps than the LOPAR Inconel grids to compensate for differences in material strength properties. Impact tests that have been performed at 600°F to obtain the dynamic strength data verify that the Zircaloy grid strength data at reactor operating conditions is structurally acceptable. The 15x15 SIF Zircaloy grid design has approximately 7 percent less crush strength than the 15x15 Inconel grid design. The 15x15 SIF Zircaloy grid design is identical to the 15x15 OFA Zircaloy grid for which substantial operational experience exists.

The fuel assembly top nozzle for the SIF assembly differs from the current design in two ways: a groove is provided in each thimble thru-hole in the nozzle plate to facilitate removal; and the nozzle adapter plate is 15 mils 43-NPW-2016S

• e e thinner than the LOPAR non-removable top nozzle. In the LOPAR design the Zircaloy thimbles are mechanically joined through expanded bulges to the uppermost grid stainless steel sleeve which is welded to the top nozzle adapter plate. The SIF assembly top nozzle design and methods for removal/reattachment are the same as the Westinghouse VANTAGE 5 assembly removable top nozzle (RTN) design and methods for removal/reattachment.

The VANTAGE 5 RTN design, functional usage and NRC approval are given in Reference 5.

The 15x15 SIF bottom nozzle assembly is shorter by 320 mils when compared to the existing LOPAR assembly (see Figure 1). The difference in length is due to the SIF s nozzle plate being 195 mils thinner and the bottom nozzle 1

legs being 125 mils shorter. The SIF bottom nozzle retains the reconstitutable feature found on the LOPAR design, which uses a locking cup to lock the thimble screws on the guide thimble assembly in place, instead of the lockwire used in earlier LOPAR designs.

Both SIF and LOPAR fuel rod designs retain the nominal pellet stack height of 144 inches; however, the SIF fuel rod length increases by 0.335 inches due to an increased gas plenum for additional fission gas releases to higher burnups. Fuel rod performance is shown to satisfy the UFSAR fuel rod design bases on a region by region basis. These same bases are applicable to the Westinghouse LOPAR fuel design and the new fuel design for the Surry units. Fuel performance evaluations are completed for each fuel region to demonstrate that the design criteria will be satisfied for the fuel rod types in the core under the planned operating conditions.

43-NPW-2016S

e e The SIF assembly has a 4.5 percent increase in hydraulic resistance to flow compared to a LOPAR assembly, primarily due to the thicker and wider OFA Zircaloy grid straps. This results in an increased SIF assembly lift force and requires the use of 3-leaf holddown springs instead of the LOPAR assembly 2-leaf springs (See Figure 1). The 3-leaf spring has the same height and provides additional holddown force margin when compared to the LOPAR 2-leaf spring. The 3-leaf spring design has been successfully used in the 17xl7 optimized fuel assemblies (OFAs) and other 15xl5 LOPAR assemblies which have had up to 3 cycles of operation in other plants. The change to the 3-leaf spring is compatible with the LOPAR assembly and the handling tools at the Surry plants.

The rod bow magnitude for the SIF assemblies is expected to be less than that of the 15x15 LOPAR assemblies. The SIF assembly will have reduced grid forces (due to the Zircaloy grids) and the same fuel tube thickness-to-diameter ratio (t/d) as the LOPAR assembly, which should tend to decrease SIF rod bow compared to LOPAR fuel. For a given burn-up, the magnitude of rod bow DNBR penalty for the SIF assembly is conservatively taken to be the same as that applied to the 15x15 LOPAR fuel assembly (Reference 6).

The wear of a fuel rod cladding is dependent on both the support provided by the assembly skeleton and the flow environment to which it is subjected.

Of concern is the existence of crossflow caused by the difference in axial pressure distribution of the SIF and LOPAR fuel assemblies due to different grids and the existence of crossflow due to changes in the distribution of core outlet loss coefficients (PFO) resulting from thimble plug removal.

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Hydraulic flow tests, which are described in Section 5 of this report, were performed to verify the compatibility of the two assemblies. The results of the tests show that no significant SIF or LOPAR fuel rod wear occurs due to the small amount of crossflow between fuel assemblies. Fuel rod vibration tests have shown that there was no significant difference in fuel rod response between the tests performed with or without PFO mismatches much larger than those that exist in the Surry core after thimble plug removal. Therefore, it is concluded that thimble plug removal will not have a detrimental impact on fuel rod vibration and wear.

Westinghouse studies on control rod wear have shown that most of the wear tends to be in the upper internals region. When thimble plugs are removed, the hydraulic resistance at the outlet for these assemblies is reduced.

This in turn causes the flow through the RCCA guide tubes to be reduced, because more flow is now going through the outlet of the assemblies which were previously fitted with thimble plugs. This reduction of flow through the RCCA guide tubes is in the direction that would tend to reduce control rod wear. In addition, it was concluded that the maximum PFO mismatch between an RCC location and an adjacent assembly does not increase with thimble plug removal for the Surry reactor upper internals configuration.

As a result, the magnitude of the crossflow seen by the control rods through the gap between the top nozzle and the upper core plate and the vibration of the rods caused by this crossflow will not be increased.

Therefore, thimble plug removal will not have an adverse impact on control rod wear for the Surry reactors.

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e 4.0 NUCLEAR EVALUATION The transition from 15x15 LOPAR fuel assemblies to 15x15 SIF assemblies will not result in changes from the current nuclear design bases as given in the UFSAR and applied to subsequent Surry Units 1 and 2 Reload Safety Evaluations. Although the physics characteristics are slightly different for a core of SIF fuel when compared to 15x15 LOPAR fuel, evaluations show that the differences are well within the normal range of variations seen from cycle to cycle.

Thus any significant changes in nuclear characteristics found in the LOPAR/SIF transition cores or eventually in a 100% SIF core will be due to fuel management variables (number of feed assemblies, feed enrichment, cycle length, etc.) and not due to the change in fuel assembly design. As in current practice, each reload core design will be evaluated to assure that design and safety limits are satisfied according to the reload methodology (Reference 2).

5.0 THERMAL AND HYDRAULIC DESIGN Hydraulic compatibility of the 15x15 SIF assembly with the 15x15 LOPAR assembly is demonstrated by hydraulic tests performed on the 15x15 OFA and the 15xl5 LOPAR assembly. The fuel rod and thimble tube diameters, and the Zircaloy mixing vane geometry of the SIF assembly are identical to that of the 15x15 OFA. Therefore, the results of hydraulic compatibility tests of the 15x15 OFA/LOPAR assemblies (i.e. mixing vane loss coefficients) are 43-NPW-2016S

applicable to 15x15 SIF/LOPAR assemblies. A side-by-side OFA and LOPAR fuel assembly arrangement was tested at the Westinghouse Fuel Assembly Test System (FATS) facility. Test results provided lift forces, pressure drops, cross flow and fuel vibrations, and fuel rod clad wear. Based on these test results, it was concluded that hydraulic compatibility exists between OFA and LOPAR assemblies and therefore, it is concluded that hydraulic compatibility exists between SIF and LOPAR assemblies.

The same calculational methods currently used on the 15x15 LOPAR fuel assembly and described in the FSAR and fuel densification documents are applicable to the evaluation of a core containing both 15x15 LOPAR and 15x15 SIF assemblies except for the application of the WRB-1 DNB correlation for the SIF assemblies. The present DNB safety evaluations for LOPAR fuel use the W-3 based L-grid DNB correlation and the design limit DNBR of 1.30. The evaluations contain a generic DNBR margin of 18 percent for the LOPAR fuel (Reference 7). The SIF assemblies are analyzed with the THINC-I code and the WRB-1 CHF correlation for which the 95/95 limit is 1.17.

Additionally, a plant specific DNB margin allowance has been considered in the Surry Units 1 and 2 SIF analysis. In particular, a design DNBR limit of 1.46 was employed in the safety analyses, resulting in 20 percent safety analysis DNBR margin as defined in the following equation.

DNBR LIMIT = 1.17 = 1.46 1-Margin 1-0.20 43-NPW-2016S

e e The plant allowance available between the design DNBR used in the safety analyses and the WRB-1 correlation limit DNBR of 1.17 is not required to meet the design basis, but will be used for flexibility in the design, operation, and analyses for Surry Units 1 and 2. For instance, the allowance may be used for improved fuel management or increased plant availability. In the Surry Units 1 and 2 transition cycles, the 20 percent margin is more than sufficient to accommodate the maximum rod bow penalty at Loss of Flow Conditions (References 6 and 8) and the 3 percent transition core penalty. The 3 percent penalty, which accounts for the lower flow in the higher resistance SIF assemblies, was determined by analyzing 15x15 OFA (same mixing vane geometry as the SIF assembly) and LOPAR assembly loading patterns at various core conditions. The analyses were performed in the same manner as for the 17x17 OFA/LOPAR fuel which was reviewed and approved by the NRC (Reference 9). When the full transition is complete (LOPAR assemblies removed from core), the transition core penalty will no longer apply to the SIF assemblies.

The main impact of thimble plug removal is the increase in core bypass flow. Calculations performed by Westinghouse have shown that the design value of core bypass flow assumed in the evaluations needs to increase from 4.5% to 6%. This results in an approximate loss of 2% in DNBR margin due to the 1.5% flow decrease in the fuel rod channels. This loss of margin is accommodated by retained DNB margin for both the SIF (WBR-1) .and the LOPAR (W-3) fuel.

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e Thimble plug removal also results in a reduction to the fuel assembly hydraulic loss coefficient. Westinghouse has performed hydraulic tests to quantify the magnitude of this effect. Based on these tests, the effect on the primary system flow rate due to thimble plug removal has been calculated. The results show that the primary system flow rate will increase by a maximum of 0.2% for the SIF assemblies without thimble plugs.

The reduced fuel assembly loss coefficient also results in a net reduction in the hydraulic lift force on the assembly, which more than compensates for the slight increase in core flow rate. Thimble plug removal is therefore acceptable from a fuel assembly lift standpoint.

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. Since the variations in outlet loss coefficient due to thimble plug removal for the Surry cores are within the bounds of the sensitivity studies that have been performed, it is concluded that thimble plug removal will not result in the reduction of DNBR margin due to mismatches in core outlet pressure gradients and loss coefficients.

The current Technical Specification core safety limits (T.S. Figures 2.1-1, 2.1-2, and 2.1-3) and the associated set points continue to be applicable for the transition cores and the all-SIF cores.

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e 6.0 ACCIDENT EVALUATION Those accidents analyzed and reported in the UFSAR (Reference 10) which could potentially be affected by the SIF reload have been reviewed. As discussed in Section 3.0 of this report, rod drop time is increased from 1.8 to 2.4 seconds (to dashpot) for control/shutdown rods in SIF guide thimbles. Accident transients significantly affected are 11 fast 11 transients for which the protection system responds by tripping the reactor within a few seconds after the transient begins. The transients that fall into this category are Loss of Flow, Locked Rotor, and Rod Ejection. In these reanalyses, the impact of the increased drop time was considered for all transition core configurations, including an all-LOPAR fueled core. Other non-LOCA accidents analyzed in the UFSAR are evaluated to be minimally affected by the increased rod drop time.

The simultaneous Loss of Flow from the three coolant pumps, Locked Rotor, and Rod Ejection accidents were reanalyzed to account for the increased rod drop time. Results for these accident reanalyses (Tables 1 - 3) showed that the safety limits and criteria are satisfied for the increased rod drop time. These reanalyses will be incorporated in the next UFSAR update.

The 1.5% increase in bypass flow resulting from thimble plug removal will result in approximately a 2% reduction in DNBR margin for the reanalyzed accidents. This reduction will be accommodated by the retained DNBR margin discussed in Section 5.0 of this report.

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e The large break loss-of-coolant accident (LOCA) analysis for Surry (VPA),

applicable to a full core of 15xl5 SIF, was analyzed to develop specific peaking factor limits. This is consistent with the methodology employed in the Reference Core Report 17xl7 Optimized Fuel Assembly (OFA) for 17xl7 OFA Transition (WCAP-9500). The assumptions and initial operating conditions used in this reanalysis were the same as those used in the LOCA-ECCS analysis (Reference 11) currently under review, with the following exception:

A 15xl5 SIF fuel assembly was modeled with a core power of 2441 MWt, Zircaloy grids, a rod length increase of 0.33 inches, thimble plug removal, and increased core bypass of 6%.

With the above changes incorporated, the analysis performed to determine the sensitivity to SIF fuel resulted in a peak clad temperature (PCT) of 1969.2°F at peaking factors of FQ=2.32 and F~H=l.62.

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 resistance mismatch may exist only for transition cores and is the only unique difference between a complete core of either fuel type and the transition core. The difference in fuel assembly resistance (K/A 2) for the two assembly designs (15xl5 SIF versus 15xl5 LOPAR) may impact the blowdown and reflood transients of the large break LOCA analysis model. Specific blowdown heat up calculations have been performed to determine the clad temperature effect on the OFA 43-NPW-2016S

e design (hydraulically the same as the SIF assembly) for mixed core configurations. The results of these calculations have shown that no peak clad temperature (PCT) penalty is observed during blowdown. In addition, design specific analyses have been performed which accurately model mixed core cases during reflood. Westinghouse transition core designs including specific 14x14, 15x15 and 17x17 LOPAR to OFA transition core cases were analyzed. The various fuel assembly specific transition core analyses performed resulted in peak clad temperature increases of up to 10°F. Since the SIF assembly is hydraulically the same (see Section 5.0) as the 15x15 OFA, the maximum PCT penalty possible for 15x15 SIF assembly during transition cores is 10°F. Once a full core of 15x15 SIF is achieved, the PCT for a 100% SIF core will be 1969°F for a FQ peaking factor of 2.32.

From the Reference 11 analysis performed for Surry, a direct sensitivity showing an increase in PCT of 27°F can be demonstrated for 15x15 SIF compared to the existing fuel at Surry. Since the increase in PCT was greater than the 20°F criterion specified in 10 CFR 50 Appendix K, the results of this analysis are provided in Appendix A to the Safety Evaluation. Only the resulting tables and figures are provided in Appendix A since the analysis is a sensitivity to the Reference 11 submittal.

7.0 TECHNICAL SPECIFICATION CHANGES Based on the preceding evaluations the following technical specification changes for Surry Units 1 and 2 are required to support the transition to SIF:

43-NPW-2016S

1.

Modifications to Specification 3.12.C and Specification Bases B3.2 to permit an increase in the shutdown and control rod drop time to 2.4 seconds.

2. Modifications to Specification Bases 2.1, 2.3, 3.1, 3.2 and 3.12 to incorporate the design DNBR limit for SIF.

These changes are given in the proposed Technical Specification page changes (Attachment 1). The changes necessary to change the LOCA FQ limit from the currently approved limit of 2.18 to 2.32 have been included in Reference 11.

In addition to the changes mentioned above, minor editorial changes are included in this submittal on some of the affected Technical Specification pages. The editorial changes to the Technical Specifications are limited to: reformating the paragraphs to be consistent with the outline format; removing the word 11 DELETED 11 where a specification had previously been deleted; placing in parentheses a supplementary 11 i.e. 11 clause; replacing the words 11 hot standby 11 with 11 hot shutdown 11 ; and adding commas where necessary. Other editorial changes to the bases include correcting verb tenses.

43-NPW-2016S

e 8.0 10 CFR 50.59 EVALUATION The proposed changes to the fuel assembly design have been determined not to result in an unreviewed safety question as defined in 10 CFR 50.59. The basis for this determination is as follows:

1. The probability of occurrence or the consequence of an accide~t or malfunction of equipment important to safety previously evaluated in the safety analyses is not increased. The proposed changes involve parameters and equipment which are not accident initiators and therefore, they will not increase the probability of occurrence of any malfunction or accident previously addressed. The reanalyzed large break LOCA analysis verifies that operation under the revised specifications would not result in any increase in LOCA accident consequences. The Loss of Flow, Locked Rotor, and Rod Ejection accidents were reanalyzed to limits based on the WRB-1 DNB correlation. The WRB-1 correlation limit, like the W-3 limit, was developed on the basis that if the calculated DNBR for a rod was greater than or equal to the design limit there was 95% probability with 95% confidence that the rod was not in DNB. On this basis, the reanalysis of the non-LOCA transients indicate that there will be no increase in the consequence of non-LOCA transients previously analyzed.

2. The possibility for an accident or malfunction of a different type than any evaluated in the safety analysis is not created since the SIF assemblies satisfy the current design criteria. In addition, the 43-NPW-2016S

e e changes to the assembly design do not involve any alteration to plant equipment or procedures which would introduce any new or unique accident precursors.

3. The margin of safety, as defined in the basis for the affected Technical Specifications, is not reduced. The revised ECCS analysis meets the requirements of 10 CFR 50.46. Additionally, the reanalyses of the non-LOCA accidents using the WRB-1 DNB correlation meet the design limits. As stated above, the WRB-1 and W-3 correlation limits were developed on the same basis. Therefore, since the reanalyses meet the design limits using the WRB-1 correlation there is no reduction in the margin of safety.

The editorial changes being made to these Technical Specifications have also been determined not to result in an unreviewed safety question as defined in 10 CFR 50.59.

43-NPW-2016S

e 9.0 10 CFR 50.92 NO SIGNIFICANT HAZARDS DETERMINATION The proposed changes do not involve a significant hazards consideration because operation of Surry Units 1 and 2 in accordance with these changes would not:

(1) involve a significant increase in the probability or consequence of an accident previously evaluated. The revised LOCA analysis which supports these changes demonstrated that the ECCS acceptance criteria of 10 CFR 50.46 were met. Also the non-LOCA analyses were re-evaluated and met limits based on the WRB-1 DNB correlation. The WRB-1 correlation limit, like the W-3 limit, was developed. on the basis that if the calculated DNBR for a rod was greater than or equal to the design limit there was 95% probability with 95% confidence that the rod was not in DNB. Therefore, the probability or consequence of an accident previously evaluated will not increase.

(2) create the possibility of a new or different kind of accident from any accident previously identified. 'The proposed fuel assembly design meets all the design criteria applied to the current 15x15 LOPAR design. In addition, the changes in the assembly design do not change plant systems or procedures. Therefore the changes do not produce any accident precursors.

(3) involve a significant reduction in the margin of safety. The revised ECCS analysis meets the requirements of 10 CFR 50.46. Additionally, 43-NPW-2016S

e e the reanalyses of the non-LOCA accidents using the WRB-1 DNB correlation meet the analysis limits. As stated above, the WRB-1 and W-3 correlation limits were developed on the same basis. Therefore, since the reanalyses meet the analysis limits using the WRB-1 correlation there is no reduction in the margin of safety.

We have also reviewed the examples of types of amendments which the NRC considers not likely to involve significant hazards considerations (51 FR 7744, March 6, 1986) and found that Example (iii) was directly applicable to the proposed change. Example (iii) states: 11 For a nuclear power reactor, a change resulting from a nuclear reactor core reloading, if no fuel assemblies significantly different from those found previously acceptable to the NRC for a previous core at the facility in question are involved. This assumes that no significant changes are made to the acceptance criteria for the technical specifications, that the analytical methods used to demonstrate conformance with the technical specifications and regulations are not significantly changed and that the NRC has previously found such methods acceptable. 11 The analyses show that the demonstration assemblies will meet the original design criteria of other assemblies in reload cores. The analytical methods used will remain unchanged. In addition, Example (i) was found to be applicable to the proposed editorial changes. Example (i) states: 11 A purely administrative change to technical specifications: for example, a change to achieve consistency throughout the technical specifications, correction of error, or a change in nomenclature. 11 The editorial changes reflect an effort to make the Technical Specifications Reactor Operating Conditions consistent and correct several grammatical errors.

43-NPW-2016S

e

10.0 REFERENCES

1. Letter from E. P. Rahe (Westinghouse) to J. Lyons (NRC),

Subject:

11 Transmittal of Operational Experience with Westinghouse Cores through December 31, 1985, 11 ( non-proprietary) , NS-NRC-86-3184, November 26, 1986.

2. 11 Reload Nuclear Design Methodology, 11 VEP-FRD-42 Revision 1-A, September 1986.

3. Bordelon, F. M. et al., 11 Westinghouse Reload Safety Evaluation Methodology, 11 WCAP-9272 (Prop.) and WCAP-9273 (Non-Prop.), March 1978.

4. Davidson, S. L., Iorii, J. A., 11 Reference Core Report - 17x17 Optimized Fuel Assembly, 11 Westinghouse Report WCAP-9500-A, dated May 1982.

5. Davidson, 11 Reference Core Report - VANTAGE 5 Fuel Assembly, 11 Westinghouse Report WCAP 10444-P-A, dated September 1985.

6. Letter from C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse),

Subject:

11 Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum Rod Bow Penalty, 11 June 18, 1986.

7. 11 Reduction of Rod Bow Penalty for North Anna and Surry, 11 NFE Technical Report No. 409, K. L. Basehore, September 1984.

8. Stolz, J. F., NRC letter to T. M. Anderson, Westinghouse, 11 Staff Review of WCAP-8691, 11 April 5, 1979.

9. Letter from C. 0. Thomas (NRC) to E. P. Rahe (Westinghouse),

Subject:

Supplemental Acceptance Number 2 for Referencing of Licensing Topical Report WCAP-9500, NRC SER letter dated January 24, 1983.

10. Surry Nos. 1 and 2 Updated Final Safety Analysis Report.

11. Letter from W. L. Stewart, to the Nuclear Regulatory Commission, Serial No.87-001 dated February 23, 1987.

e 15X15 SURRY UPGRAOEO FUEL ASSEMBLY 1""1--------------------- 169.955 R E F . - - - - - - - - - - - - - - - - - - - - - - . - ,

2.418 REF. _ _,.__ __.

~iu;-

\ :

-,._J:

--:*:..-r---

REF. REF. REF.

15X15 LOPAR FUEL ASSEMBLY

- - - - - - - - - - - - - - - - - - 159.71 R E F . - - - - - - - - - - - - - - - - - - - - - - -

2.738 REF.---......- -

---

151.85 REF.-----------------

FUEL ROD LENGTH REF. REF.

Figure 1

TABLE I LOSS OF FLOW RESULTS Calculated Minimum Value Limit DNBR (SIF) 1.480 1.46 DNBR (LOPAR FUEL) 1.374 1.30 TABLE 2 LOCKED ROTOR RESULTS Calculated Value Limit System Pressure {psia) 2614 2750 Peak Clad Temperature (°F) 1654 2700

% of Rods Below DNBR Limit <5 -:<;5 *

• Current UFSAR design basis TABLE 3

SUMMARY

OF ROD EJECTION ANALYSIS PARAMETERS AND RESULTS Case BOL-HZP*** BOL-HFP EOL-HZP EOL-HFP LIMIT Max. Fuel Centerline Temperatures (°F) 3917 Melt** 3726 4789 10% Melt Max. Clad Inner Temperature (°F) 2353 2383 2228 2278 2700 Max. Fuel Enthalpy (Btu/lb) 250 320 234 301 360

• • Less than 10% of fuel melt at fuel rod hot spot

• • • BOL-HZP Beginning of Life - Hot Zero Power 43-NPW-2016S

APPENDIX A LARGE BREAK ECCS LOSS OF COOLANT ACCIDENT FOR SURRY 43-NPW-2016S

TABLE 1 INITIAL CORE CONDITIONS ASSUMED FOR THE DOUBLE-ENDED COLD-LEG GUILLOTINE BREAK (DECLG}

Calculational Input Core Power (MWt) 102% of 2441 Peak linear power (kW/ft) 102% of 14.39 Heat flux hot-channel factor (FQ) 2.32 Enthalpy rise hot-channel factor (F ~H) 1.62 Accumulator water volume (ft 3 , each) 1000 Reactor vessel upper head temperature equal to Thot Limiting Fuel Region and Cycle Cycle Region Unit 1 All All regions Unit 2 All A11 regions

- TABLE 2 CONTAINMENT DATA (DRY CONTAINMENT)

Net Free Volume 1.863 X 10 6 ft 3 Initial Conditions Pressure (total), psia 9.35 Temperature, °F 80 RWST temperature, °F 45 Service water temperature, °F 32.5a Outside temperature, °F 9 Spray System I - Containment Spray System Number of pumps operating 2 Runout flowrate, gpm 3200 Actuation time, sec 59 Spray System II - Inside Recirculation Spray Subsystem Number of pumps operating 2 Runout flowrate (each), gpm 3500 Actuation time, sec 190 Heat exchanger (UA (per pump)), Btu/hr-°F 5 .18 X 10 6 Service water flow (per exchanger), gpm 6900 Spray System II - Outside Recirculation Spray Subsystem Number of pumps operating 2 Runout flowrate (each), gpm 3500 Actuation time, sec 365 6

Heat exchanger (UA (per pump)), Btu/hr-°F 5 .18 X 10 Service water flow (per exchanger), gpm 6900 aSensitivity analyses provided in Reference 15 demonstrate that service water temperature levels as low as 25°F will have a negligible impact on the limiting results of the LOCA-ECCS analyses.

TABLE 2 (Continued)

CONTAINMENT DATA (DRY CONTAINMENT)

Structural Heat Sinks Type/thickness (in.) Area (ft 2), with uncertainty Concrete/6 8,393 Concrete/12 62,271 Concrete/18 55,365 Concrete/24 11,591 Concrete/27 9,404 Concrete/36 3,636 Carbon steel/0.375 Concrete/54 46,489b Carbon steel/0.50 Concrete/30 25,652b Concrete/26 (floor) 12,110 Carbon steel/0.239 158,059b Stainless Steel/0.306 17,519 Aluminum/0.0091 3,911 bcredit for painted surfaces was taken only for the nominal surface area.

TABLE 3 TIME SEQUENCE OF EVENTS DECLG c0 = 0.4 (sec)

Start 0.0 Reactor trip 0.427 Safety injection signal 2.52 Accumulator injection 15.60 Pump injection 27.52 End of bypass 30.90 End of blowdown 30.90 Bottom of core recovery 43.41 Accumulator empty 53.78

TABLE 4 RESULTS FOR DECLG Peak clad temperature, OF 1979.2*

Peak clad location3 ft 6.25 Local Zr/H 2o reaction (max),% 4.68 Local Zr/H 20 location, ft 6.00 Total Zr/H 20 reaction, % <0. 3 Hot-rod burst time, sec 40.2 Hot-rod burst location, ft 6.25

• Includes 10°F transition core penalty.

e TABLE 5 REFLOOD MASS AND ENERGY RELEASES DECLG (CD= 0.4)

Total Mass Total Egergy Time (sec) Flow Rate (lb/sec) Flow Rate (10 Btu/sec) 43.411 0.0 0.0 44.361 6.93 0.090 53.980 75.01 0.941 67.480 82.00 1.022 84.280 86.45 1.075 102. 230 244.90 1.481 120.930 293.37 1.550 140.830 298.73 1.504 262.130 351. 90 1.372 TABLE 6 BROKEN LOOP ACCUMULATOR FLOW TO CONTAINMENT DECLG (CD= 0.4)

Time (sec) Mass Flow Ratea(lbm/sec) 0.00 4214.45 1.01 3809.62 3.01 3267.32 5.01 2907.05 7.01 2643.06 10.01 2346.95

15. 01

• 2004.47 20.01 1766.24 25.01 1592.09 27.01 1537.02 aFor energy flowrate, multiply mass flow rate by a constant of 59.62 Btu/lbm.

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