ML16256A216
| ML16256A216 | |
| Person / Time | |
|---|---|
| Site: | Waterford  |
| Issue date: | 08/25/2016 |
| From: | Entergy Operations |
| To: | Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML16256A115 | List:
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| References | |
| W3F1-2016-0053 | |
| Download: ML16256A216 (80) | |
Text
WSES-FSAR-UNIT-34.2-14.2FUEL SYSTEM DESIGN4.2.1DESIGN BASESThe bases for fuel system design are discussed in the following subsections. Additional information forthe current fuel cycle is discussed in Appendix 4.3A.4.2.1.1Fuel AssemblyThe fuel assemblies are required to meet design criteria for each design condition listed below to assurethat the functional requirements are met. Except where specifically noted, the design bases presented in this section are consistent with those used for previous designs.a)Condition I: Non-operation and Normal OperationCondition I situations are those which are planned or expected to occur in the course of handling,initial shipping, storage, reactor servicing and power operation (including maneuvering of theplant). Condition I situations must be accommodated without fuel assembly failure and withoutany effect which would lead to a restriction on subsequent operation of the fuel assembly. The guidelines stated below are used to determine loads during Condition I situations:1)Handling and Fresh Fuel ShippingLoads correspond to the maximum possible axial and lateral loads and accelerationsimposed on the fuel assembly by shipping and handling equipment during these periods, assuming that there are no abnormal contact between the fuel assembly and any surface, nor any equipment malfunction. Irradiation effects on material properties are considered when analyzing the effects of handling loads which occur during refueling. Additional information regarding shipping and handling loads is contained in Subsection 4.2.3.1.5.2)StorageLoads on both new and irradiated fuel assemblies reflect storage conditions oftemperature, chemistry, means of support, and duration of storage.3)Reactor ServicingLoads on the fuel assembly reflect those encountered during refueling and reconstitution.4)Power OperationLoads are derived from conditions encountered during transient and steady-stateoperation in the design power range. (Hot operational testing, system startup, hot standby, operator controlled transients within specified rate limits and system shutdown are included in this category.)
WSES-FSAR-UNIT-34.2-25)Reactor TripLoads correspond to those produced in the fuel assembly by control element assembly(CEA) motion and deceleration.b)Condition II: Upset ConditionCondition II situations are unplanned events which may occur with moderate frequency during thelife of the plant. The fuel assembly design should have the capability to withstand any upset condition with margin to mechanical failure and with no permanent effects which would prevent continued normal operation. Incidents classified as upset conditions are listed below:1)Operating basis earthquake (OBE)2)Uncontrolled CEA withdrawal3)Uncontrolled boron dilution 4)Partial loss-of-coolant flow5)Idle loop startup (in violation of established operating procedures)6)Loss of load (reactor-turbine load mismatch)7)Loss of normal feedwater8)Loss of offsite power 9)Excessive heat removal (feedwater system malfunction) 10)CEA drop11)Accidental depressurization of the Reactor Coolant System (RCS)c)Condition III: Emergency ConditionsCondition III events are unplanned incidents which might occur very infrequently during plant life.Fuel rod mechanical failure must be prevented for any Condition III event in any area not subject to extreme local conditions (e.g., in any fuel rod not immediately adjacent to the impact surfaceduring fuel handling accident).The Condition III incidents listed below are included as a category to provide assurance that underthe occurrence of a Condition III event, rod damage is minimal.1)Complete loss or interruption of primary coolant flow at 100% power, excluding reactorcoolant pump locked rotor2)Steam bypass malfunction WSES-FSAR-UNIT-3 4.2-3 Revision 14 (12/05) 3) Minor fuel handling accident (fuel assembly and grapple remain connected) 4) Inadvertent loading of fuel assembly into improper position d) Condition IV: Faulted Conditions Condition IV incidents are postulated events whose consequences are such that the integrity and operability of the nuclear energy system may be impaired. Mechanical fuel failures are permitted, but they must not impair the operation of the Engineered Safety Features (ESF) systems to
mitigate the consequences of the postulated event. Condition IV incidents are listed below: 1) Safe shutdown earthquake (SSE)
- 2) Loss-of-coolant accident (LOCA)
- 3) Locked coolant pump rotor
- 4) Major secondary system pipe rupture
- 5) CEA ejection
- 6) Major fuel handling accident (fuel assembly and grapple are disengaged) (DRN 03-2058, R14) See Sections 3.6.2.1.1.1(d) and 3.6.3 for discussions on pipe break criteria and leak-before-break.(DRN 03-2058, R14) 4.2.1.1.1 Fuel Assembly Structural Integrity Criteria For each of the design conditions, there are criteria which apply to the fuel assembly and components with the exception of fuel rods. These criteria are listed below and give the allowable stresses and
functional requirements for each design condition. a) Design Conditions I and II Pm S m P m + P b F s S mUnder cyclic loading conditions, stresses must be such that the cumulative fatigue damage factor does not exceed 0.8. Cumulative damage factor is defined as the sum of the ratios of the number of cycles at
a given cyclic stress (or strain) condition to the maximum number permitted for that condition. The selected limit of 0.8 is used in place of 1.0 (which would correspond to the absolute maximum damage
factor permitted) to provide additional margin in the design. Deflections must be such that the allowable trip time of the control element assemblies is not exceeded.
WSES-FSAR-UNIT-34.2-4Revision 11 (05/01)b)Design Condition III Pm 15.S m P m P b 1.5 F sS mDeflections are limited to a value allowing the CEAS to trip, but not necessarily within theprescribed time.c)Design Condition IV P m S mP m + P b F s S mwhere S' m = smaller value of 2.4 S m or 0.7 S u.1)If the equivalent diameter pipe break in the LOCA does not exceed the largest line connected tothe main reactor coolant lines, the fuel assembly deformation shall be limited to a value notexceeding the deformation which would preclude satisfactory insertion of the CEAS.2)For pipe breaks larger in equivalent diameter than the largest lines connected to the main reactorcoolant lines, deformation of structural components is limited to maintain the fuel in a coolablearray. CEA insertion is not required for these events as the appropriate safety analyses do not take credit for CEA insertion.d)NomenclatureThe symbols used in defining the allowable stress levels are as follows:
P m = Calculated general primary membrane stress (a)P b = Calculated primary bending stress S m = Design stress intensity value as defined by Section III, ASME Boiler and Pressure Vessel Code(b)S u =Minimum unirradiated ultimate tensile strength (DRN 00-644)
F s =Shape factor corresponding to the particular cross section being analyzed(c) (DRN 00-644)
S'm =Design stress intensity value for faulted conditions WSES-FSAR-UNIT-3 4.2-5 Revision 15 (03/07)
The definitions of S' as the lesser value of 2.4 S m and 0.7 S u is contained in the ASME Boiler and Pressure Vessel Code (19/4)Section III, Appendix F-1323.1.
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(a) P m and P b are defined by Article NB-3000,Section III, ASME Boiler and Pressure Vessel Code, 1971.
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(b) With the exception of zirconium base alloys, the design stress intensity values, S m , of materials not tabulated by the Code are determined in the same manner as the Code. The design stress intensity of zirconium base alloys shall not exceed two-thirds of the unirradiated minimum yield
strength at temperature. Basing the design stress intensity on the unirradiated yield strength is conservative because the yield strength of zircaloy increases with irradiation. The use of the two-thirds factor ensures 50% to component yielding in response to primary stresses. This 50%
margin together with its application to the minimum unirradiated properties and the general conservatism applied in the establishment of design conditions is sufficient to ensure an
adequate design.
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(c) The shape factor, F s, is defined as the ratio of the "plastic" moment (all fibers just at the yield stress) to the initial yield amount (extreme fiber at the yield stress and all other fibers stressed in proportion to their distance from the neutral axis). The capability of cross sections loaded in
bending to sustain moments considerably in excess of that required to yield the outermost fibers
is discussed in Timoshenko.
(1) 4.2.1.1.2 Material Selection (DRN 02-1538, R12)
The fuel assembly grid cage structure consists of 10 Zircaloy-4 spacer grids, 1 Inconel 625 spacer grid (at the lower end), 5 Zircaloy-4 CEA guide tubes, 2 stainless steel end fittings, and 5 Inconel X-750 coil springs. Beginning with Batch U, some grid cages will have 9 Zircaloy-4 grids and 2 Inconel 625 grids (at
the upper and lower ends) Zircaloy-4, selected for fuel rod cladding, guide tubes and spacer grids, has a
low neutron absorption cross section, high corrosion resistance to reactor water environment and there is
little reaction between the cladding and fuel or fission products. As described in Subsection 4.2.3, Zircaloy-4 has demonstrated its ability as a cladding, CEA guide tube, and spacer grid material.
(DRN 06-1059, R15)
Beginning with the Region Y fuel assemblies in Cycle 15, ZIRLO TM is introduced as a fuel rod cladding material to provide added corrosion resistance and fuel reliability. ZIRLO TM is a zirconium-based alloy that improves fuel assembly corrosion resistance and dimensional stability under irradiation. (DRN 06-1059, R15)
The bottom spacer grid is of Inconel 625 and is welded to the lower end fitting. For the assembly designs with the Inconel 625 top grids, the grid is retained by 10 Zircaloy-4 sleeves that are welded to the Zircaloy-4 guide tubes. In these regions of higher turbulence, Inconel 625 was selected rather than Zircaloy-4 to provide additional strength and relaxation resistance. Inconel 625 is a very strong material with good ductility, corrosion resistance and stability under irradiation at temperatures below 1000F. (DRN 02-1538, R12)
WSES-FSAR-UNIT-3 4.2-6 Revision 309 (06/16)
(DRN 02-1538, R12, LBDCR 15-025, R309)
The fuel assembly lower end fitting is of cast stainless steel (Grade CF-3) and the upper end fitting assembly consists of two cast stainless steel plat es and five Type 304 stainless steel machined alignment posts. This material was selected based on c onsiderations of adequate strength and high-corrosion resistance. Also, Type 304 stainless steel has been us ed successfully in almost all pressurized water reactor environments, including all currently operating C-E reactors. (DRN 02-1538, R12, LBDCR 15-025, R309)
(EC-9533, R302; LB DCR 13-014, R309)
With the introduction of the Next Generation Fuel (NGF) design in Region Z (Cycle 16), the fuel assembly grid cage structure consists of 13 spacer grids (an Inconel-718 t op grid, six vaned Optimized ZIRLO TM mid grids, three unvaned Optimized ZIRLO TM mid grids, two vaned Optimized ZIRLO TM Intermediate Flow Mixing grids (IFMs), and one Inconel-625 bottom gr id), five Stress-Relief Annealed (SRA) ZIRLO TM CEA guide tubes, two stainless steel end fittings, and five Inconel X-750 coil springs. Instead of welds, the NGF grid cage structure utilizes bulges to secure the spacer grids to the CEA guide tubes. Similarly, the CEA guide tube flange-to-tube connection is bulged for NGF instead of welded. (LBDCR 13-014, R309)
The use of Optimized ZIRLO for the mid grid and IF Ms improves the corrosion resistance dimensional stability of the grids, thereby reducing grid growth and improving fretting resistance. Similarly, the use of ZIRLO for the CEA guide tubes improves the corrosi on resistance and dimensional stability of the guide tubes. (LBDCR 13-014, R309)
The use of Inconel-718 for the top grid maintains t he ductility, strength, and st ability benefits of the prior Inconel-625 top grid while utilizing a design that is compatible with the NGF rod diameter and basically the same as used in Westinghouse reactors for m any years with excellent performance results. (LBDCR 13-014, R309)
The NGF fuel rod design utilizes Optimized ZIRLO TM for the cladding and includes several geometric changes (see Section 4.2.2.2 for description of the NGF fuel rod geometry). The use of Optimized ZIRLO TM improves the corrosion re sistance of the cladding. (EC-9533, R302) 4.2.1.1.3 Control Element Assembly Guide Tubes
All CEA guide tubes are manufactured in accordanc e with Grade RA-2, ASTM B353, Wrought Zirconium and Zirconium Alloy Seamless and Welded Tubes for Nu clear Service, with the following exceptions and/or additions:
a) Chemical Properties (EC-9533, R302)
Chemical analyses are performed for the alloyi ng elements. For Zircaloy-4 guide tubes, the analyses check for tin, iron, chromium, oxygen, and zirconium. For ZIRLO TM guide tubes, the analyses check for tin, iron, niobium, oxygen, and zirconium. (EC-9533, R302) b) Mechanical Properties (EC-9533, R302)
The guide tubes are fabricated from Zirc aloy-4 or, starting with NGF, ZIRLO TM in the stress-relief annealed (SRA) condition and are tested for yield st rength, ultimate strength, and elongation at room temperature and elevated temperature conditions.
(EC-9533. R302) c) Dimensional Requirements Permissible Tolerance Dimension (in.)
ID 0.005 (thru Batch y) +/- 0.002 (Batch Z and beyond) (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-7 Revision 302 (12/08) 4.2.1.1.4 Zircaloy-4 Bar Stock (DRN 00-644)
All Zircaloy-4 bar stock is fabricated in accor dance with Grade RA-2, ASTM B351, Hot-Rolled and Cold-Finished Zirconium and Zirconium Alloy Bars, Rod and Wi re for Nuclear Application, with the following exceptions or additions: (DRN 00-644) a) Chemical Properties
Additional limits are placed on oxygen and silicon content.
b) Metallurgical Properties
- 1) Grain Size
The maximum average grain size is restricted. (EC-9533, R302) 4.2.1.1.5 Zirconium-B ased Alloy Strip Stock (EC-9533, R302)
All Zircaloy-4 strip stock is f abricated in accordance with Grade RA-2, ASTM B352, Zirconium Alloy Sheet, Strip and Plate for Nuclear Application, with the following exceptions or additions: (DRN 00-644) a) Chemical Properties (EC-9533, R302)
Chemical analyses are performed for the alloying elements. For Zircaloy-4 strip stock, the analyses check for tin, iron, chromium, o xygen, and zirconium. For Optimized ZIRLO TM strip stock, the analyses check for tin, iron, niobium, oxygen, and zirconium. (EC-9533, R302) b) Metallurgical Properties
(1) Grain Size
The maximum average grain size is restricted.
c) Mechanical Properties
- 1) Bend (DRN 06-895, R15)
Spacer and perimeter strips for spacer grids are to be free of cracks. Strips from each material lot are penetrant inspected in a ccordance with a quality control plan that ensures, with 95% confidence, that at least 95% of the strips are free of cracks. The method used is capable of detecting known cra cks in a standard specimen grid strip. All strips found to have cracks shall be rejected. (DRN 06-895, R15) d) Coefficient of Thermal Expansion
Axial direction - See Reference 2
e) Irradiation Properties:
The yield and tensile strengths are enhanced by irradi ation. The stress relaxation with irradiation at operating temperatures proceeds at a rapid ra te until nearly complete. The irradiation induced growth is documented.
WSES-FSAR-UNIT-3 4.2-8 Revision 309 (06/16) 4.2.1.1.6 Stainless Steel Castings (DRN 02-1538, R12, LBDCR 15-025, R309)
Stainless steel castings are fabricated in acco rdance with Grade CF-3, ASTM A744/A744M, with the following addition:
(DRN 02-1538, R12) a) Chemical Properties Cobalt content is limited.
Starting in 2015, stainless steel end fitting castings are fabricated in accordance with Westinghouse Specification MACASS01. MACASS01 duplicates t he requirements of ASTM A744 except for two changes. The required heat treatment (i.e., solution anneal) is not necessary for this application and is not specified. The specification to control delta ferri te level is removed given the relatively low carbon content of the CF-3 cast stai nless steel. Casting soundness and cast ing mechanical properties and other properties affected by the ferrite level are cont rolled by non-destructive examination and mechanical property measurements of representative samples. (LBDCR 15-025, R309) 4.2.1.1.7 Stainless Steel Tubing (EC-9533, R302; EC-30663, R307)
Stainless steel tubing is fabricated in accordance wi th ASTM A269 (with additional requirements) for wear sleeves and the GuardianTM1 grid inserts, and in accordance with either ASTM A213 or A249 (both with additional requirements) for the top Inconel grid sleeve. (EC-9533, R302; EC-30663, R307) a) Chemical Properties Carbon content is limited on tubing to be welded. Cobalt content is limited.
4.2.1.1.8 Inconel X-750 Compression Springs (DRN 02-1538, R12; EC-9533, R302)
All Inconel springs are fabricated in accor dance with AMS 5699, with the following addition: (DRN 02-1538, R12; EC-9533, R302) a) Chemical Properties Cobalt content is limited. (DRN 02-1538, R12) 4.2.1.1.9 Inconel 625 Spacer Grid Strip Material (EC-9533, R302)
Inconel spacer grid strip material is procured in accordance with t he specification for nickel-chromium-molybdenum-columbium alloy plate, sheet, and st rip (ASTM B443) for Inconel 625 strip and age-hardenable nickel-chromium-iron alloy sheet, strip, and plate (ASTM B670) for Inconel 718, both with the following additional requirements: (DRN 02-1538, R12; EC-9533, R302) a) Chemical Properties Cobalt content is limited.
b) Special Tests A check analysis and a bend test are required.
4.2.1.2 Fuel Rod 4.2.1.2.1 Fuel Cladding Design Limits The fuel cladding is designed to sustain the effect s of steady-state and expec ted transient operating conditions without exceeding acceptabl e level of stress and strain. Exc ept where specifically noted, the design bases presented in this section are consistent with those used for previous core designs. The fuel rod design accounts for cladding irradiation growth, ex ternal pressure, differential expansion of fuel and clad, fuel swelling, clad
1 Guardian is a trademark or registered trademark of We stinghouse Electric Company LLC, its affiliates and/or its subsidiaries in the United States of America and may be registered in other countries throughout the world. All rights rese rved. Unauthorized use is strictly prohibited. Other names may be trademarks of their respective owners.
WSES-FSAR-UNIT-3 4.2-9 Revision 302 (12/08) creep, fission and other gas releases, initial inter nal helium pressure, thermal stress, pressure and temperature cycling, and flow-induced vibrations. T he structural criteria discussed below are based on the following for the normal, upset, and emergency loading combinations identified in Subsection 4.2.1.1.
For a discussion of the thermal/hydrau lic criteria, see Subsection 4.4.1. (DRN 06-1059, R15; EC-9533, R302) a) During normal operating and upset conditions, the maximum primary tensile stress in the Zircaloy, ZIRLO TM, or Optimized ZIRLO TM clad shall not exceed two-thirds of the minimum unirradiated yield strength of the material at t he applicable temperature. The corresponding limit under emergency conditions is the material yield st rength. The use of the unirradiated material yield strength as the basis for allowable stress is conservative because the yield strength of zircaloy increases with irradiation. The use of the two-thirds factor ensur es 50 percent margin to component yielding in response to primary stresse
- s. The 50 percent margin, together with its application to the minimum unirradiated properties and the general conservatism applied in the establishment of design conditions, is sufficient to ensure an adequate design. (DRN 06-1059, R15) b) Net unrecoverable circumferential strain shall not exceed one percent as predicted by computations considering clad creep and fuel cl ad interaction effects. In addition, the incremental total strain induced during a transient is also limited to one percent, as described in Reference 82 for Zircaloy-4 cladding, Reference 80 for ZIRLO TM cladding, and Reference 84 for the NGF design with Optimized ZIRLO TM cladding. (EC-9533, R302)
Data from O'Donnell and Weber were used to det ermine the present one percent strain limit. (See References 4 & 5.) O'Donnell developed an anal ytical failure curve for Zircaloy cladding based upon the maximum strain of the material at its point of plastic instability. O'Donnell compared his analytical curve to circumferential strain data obtained on irradiated coextruded Zr-
U metal fuel rods tested by Weber. The corre lation was good, thus substantiating O'Donnell's instability theory. Since O'Donnell performed his analysis, additional data have been derived at Bettis and AECL. (See References 6, 7, 8, 9 & 10.)
These new data are shown in Figure 4.2-1, along with O'Donnell's curve and Weber's data. This curve was then adjusted because of differences in anisotropy, stress and strain rates; and the design limit was set at one percent. (DRN 06-1059, R15; EC-9533, R302)
The conservatism of the clad strain calculations is provided by the selection of adverse initial conditions and material behavior assumptions, and by the assumed operating history. The acceptability of the 1.0 percent unrecoverable circumferential strain limits is demonstrated by
data from irradiated Zircaloy-clad fuel rods whic h show no cladding failures (due to strain) at or below this level, as illustrated in Figure 4.2-1. (DRN 06-1059, R15)
The ductility of ZIRLO TM is expected to be at least equivalent to Zircaloy-4 (Reference 80, Section 5.3.5). Section B.7 of Reference 83 docum ents that the ductility of Optimized ZIRLO TM and ZIRLO TM are indistinguishable from each other at temperatures above room temperature, so the ductility of Optimized ZIRLO TM is also at least equivalent to that of Zircaloy-4. Ductility is a function of irradiation and hydride formation in the cladding. Since the corrosion rates of ZIRLOTM and Optimized ZIRLO TM are significantly less than that of Zircaloy-4, fewer hydrides will be formed at high burnup levels. Therefore the 1% st rain capability limit criterion will continue to be applied and satisfied in Westinghouse fuel mechanical design analysis.
c) The clad will be initially pressurized with he lium to an amount sufficient to prevent gross clad deformation under the combined effects of exter nal pressure and long-term creep. The clad design can rely on the support of fuel pellets (Reference 82) or the holddown spring (Reference
- 81) to prevent gross deformation. (EC-9533, R302) d) Cumulative strain cycling usage, defined as the sum of the ratios of the number of cycles in a given effective strain range () to the permitted number (N) at that range, as taken from Figure 4.2-2, will not exceed 0.8.
WSES-FSAR-UNIT-3 4.2-10 Revision 302 (12/08)
(EC-9533, R302)
The cyclic strain limit design curve shown on Figure 4.2-2, is based upon the Method of Universal Slopes developed by S.S. Manson and has been adjusted to provide a strain cycle margin for the effects of uncertainty and irradiation. The resulting curve has been compared with known data on the cyclic loading of Zircaloy and has been show n to be conservative. (See Reference 11.)
Specifically, it encompasses all the data of O'Donnell and Langer. (See Reference 12). The application of the curve to ZIRLO TM is documented in Reference 80 and Appendix B.10 of Reference 83 documents that there is no distinguishable difference in the fatigue characteristics
of ZIRLO TM and Optimized ZIRLO TM. (EC-9533, R302)
As discussed in Subsection 4.2.3.2.5, the fatigue calculation method includes the effect of clad
creep to reduce the pellet to clad diametral gap during that portion of operation when the pellet and clad are not in contact. The same model is used for predicting clad fatigue as is used for predicting clad strain. Therefore, the effe cts of creep and fatigue loadings are considered together in determining end-of-life cumulative fatigue damage factor and the end-of-life cumulative fatigue damage factor and the end-of-life clad strain. Moreover, the current fatigue
damage calculation method includes a factor of tw o which is applied to the calculated strain before determining the allowable number of cycles associated with that strain. This, in combination with the allowable fatigue usage fa ctor of 0.8 ensures a considerable degree of conservatism (see Figure 4.2-2).
e) There is no specific limit on lateral fuel rod defle ction for structural integr ity considerations except that which is brought about through application of cladding stress criteria. The absence of a specific limit on rod deflection is justified because it is the fuel assembly structure, and not the individual fuel rod, that is the limiting fa ctor for fuel assembly lateral deflection.
f) Fuel rod internal pressure increases with increasing burnup and toward end-of-life the total internal pressure, due to the combined effects of the initial helium fill gas and the released fission gas, can approach values comparable to the exter nal coolant pressure. The maximum predicted fuel rod internal pressure will be c onsistent with the following criteria.
- 1) The primary stress in the cladding resulting from differential pressure will not exceed the stress limits specified earlier in this section. (DRN 02-1538, R12)
- 2) The internal pressure will not cause the cl ad to creep outward from the fuel pellet surface while operating at the design peak linear heat rate for normal operation. In determining compliance with this criterion, internal pre ssure is calculated for the peak power rod in the reactor, including accounting for the ma ximum computed fission gas release. In addition, the pellet swelling rate (to which t he calculated clad creep rate is compared) is based on the observed swelling rate of "restrai ned" pellets (i.e., pelle ts in contact with clad), rather than on the greater observed sw elling behavior of pellets which are free to expand. (DRN 02-1538, R12)
WSES-FSAR-UNIT-3 4.2-11 Revision 302 (12/08)
(DRN 03-1821, R13)
The criteria discussed above do not limit fuel r od internal pressure to values less than the primary coolant pressure, and the occurrence of positive differential pressures would not adversely affect normal operation so long as appropriate criteria for cladding stress, strain, and strain rate were satisfied. The fuel rod maximum pressure criterion for allowing fuel rods to operate in reactors with internal hot gas pressure in excess of
reactor coolant system pressure is provided in Reference 79. (DRN 03-1821, R13)
g) The design limits of the fuel rod cladding, wi th respect to vibrati on considerations, are incorporated within the fuel assembly design. It is a requirement that the spacer grid intervals, in conjunction with the fuel rod stiffness, be such that fuel rod vibrati on, as a result of mechanical or flow induced excitation, does not result in excessi ve wear of the fuel rod cladding at the spacer grid contact areas.
4.2.1.2.2 Fuel Rod Cladding Properties
4.2.1.2.2.1 Mechanical Properties
a) Modulus of Elasticity (DRN 06-1059, R15; EC-9533, R302)
Young's Modulus x 10 6 = value specified in Reference 13 for Zircaloy-4, in Reference 80 for ZIRLO TM , and in Reference 83 for Optimized ZIRLO TM. (DRN 06-1059, R15; EC-9533, R302) b) Poisson's Ratio (DRN 06-1059, R15; EC-9533, R302) n = value specified in Reference 13 for Zircaloy-4, in Reference 80 for ZIRLO TM , and in Reference 83 for Optimized ZIRLO TM. (DRN 06-1059, R15; EC-9533, R302) c) Thermal Coefficient of Expansion (DRN 06-1059, R15; EC-9533, R302) diametral direction = value specified in Refe rence 13 for Zircaloy-4, in Reference 80 for ZIRLO TM , and in Reference 83 for Optimized ZIRLO TM. (DRN 06-1059, R15; EC-9533, R302) d) Yield Strength (DRN 06-1059, R15; EC-9533, R302)
Yield strength in the non-irradiated condition is shown in Figure 4.2-20 of Reference 13 for Zircaloy-4, in Section 5.3.
7.1 of Reference 80 for ZIRLO TM , and in Figure B.7-6 of Reference 83 for Optimized ZIRLO TM. (DRN 06-1059, R15; EC-9533, R302)
The cladding stress limits identifi ed in Subsection 4.2.1.2.1 are based on values taken from the minimum yield strength curve at the appropriate temperatures. The limits are applied over the entire fuel lifetime, during conditions of reac tor heatup and cooldown, st eady state operation, and normal power cycling. Under these conditions, cladding temperatures and fast fluences can range from 70 to 750 F and from 0 to 1 x 10 22 nvt, respectively.
WSES-FSAR-UNIT-3 4.2-12 Revision 309 (06/16) e) Ultimate Strength (DRN 06-1059, R15; EC-9533, R302) Ultimate tensile strength in the non-irr adiated condition is shown in Figure 4.2-21 of Reference 13 for Zircaloy-4, in Section 5.
3.7.2 of Reference 80 for ZIRLO TM , and in Figure B7-7 of Reference 83 for Optimized ZIRLO TM. (DRN 06-1059, R15; EC-9533, R302) f) Uniform Tensile Strain (DRN 02-1538, R12)
Uniform tensile strain in the irradiated condi tion approaches one percent and remains relatively constant (Subsection 4.2.1.2.1). (DRN 06-1059, R15; EC-9533, R302)
Ductility is a function of irradiation and hydride formation in the cladding wall. The ductility of ZIRLO TM and Optimized ZIRLO is expected to be at least equivalent to Zircaloy-4 because the waterside corrosion is significantly lower for ZIRLO TM and Optimized ZIRLO and will result in less hydrogen uptake and less hydride formation.
Total strain capability of ZIRLO TM and Optimized ZIRLO is projected to be in excess of 1% at burnup levels of 60 MWd/kgU. (DRN 02-1538, R12; 06-1059, R15)
Note: No flare test currently done on production cladding.
g) Hydrostatic Burst Test
Hydrostatic burst tests are conducted on Zircaloy
-4 cladding to verify that burst pressure and circumferential elongation exceed prescribed minimum values. (EC-9533, R302)
The procedures originally used by CE for the hy drostatic tests were described by D.G. Hardy, J.R. Stewart and A.L. Lowe, Jr., "Development of a Closed End Burst Test Procedure for Zircaloy Cladding," Zirconium in Nuclear Applications, ST P-551, ASTM, 1974, pp. 14-30. This information was incorporated into ASTM B353-77. The present procedure for CE cladding is essentially the same procedure as described in B353-77. Typical burst pressures of 35 samples from three lots ranged from 16.6 to 18.8 ksi. (EC-9533, R302) h) Corrosion (DRN 03-2058, R14; 06-1059, R 15, LBDCR 15-035, R309)
The Zircaloy-water reaction rate correlation used for non-LOCA applications is given in Reference 74. Note that for current analyses, the cladding corrosion rates are described in
Reference 82 for Zircaloy-4 and in Reference 85 for ZIRLO TM and in Optimized ZIRLO TM. (DRN 03-2058, R14; 06-1059, R15; EC-9533, R 302, LBDCR 15-035, R309) (DRN 06-992, R15)
The maximum allowed fuel rod cladding corrosion will be limited to 100 microns. The corrosion thickness will be calculated using the best es timate models and methods described in CENPD-404-P. (DRN 06-992, R15) 4.2.1.2.2.2 Dimens ional Requirements
a) Tube straightness is limited to 0.010 in./ft, and in side diameter and wall thickness are tightly controlled. (DRN 06-1059, R15; EC-9533, R302) b) Ovality is measured as the difference between maximum and minimum inside diameters and is acceptable if within the diameter tolerances. Outside diameter is specified as 0.382 0.002 in.
Inside diameter is specified as 0.332 0.0015 in. NGF valves are 0.374 .0015 in. for the outside diameter and 0.329 .0015 in. for the inside diameter. (DRN 06-1059, R15; EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-13 Revision 302 (12/08) c) Eccentricity is defined as the difference between maximum and minimum wall thickness at a cross section and is specified as 0.004 inches maximum. (EC-9533, R302) d) Wall thickness is specified as 0.023 in. mini mum (the nominal value reported elsewhere is based on the nominal O.D. and I.D.). Minimum wall thickness specified in NGF is 0.0207 in. (EC-9533, R302) 4.2.1.2.2.3 Metallurgical Properties
a) Hydride Orientation
A restriction is placed on the hydr ide orientation factor for any third of the tube cross-section (inside, middle, or outside). The hydride orient ation factor, defined as the ratio of the number radially oriented hydride platelets to the total num ber of hydride platelets shall not exceed 0.3.
The independent evaluation of three portions of t he cross section is included to allow for the possibility that hydride orientation may not be uniform across the entire cross section.
4.2.1.2.2.4 Chemical Properties (DRN 02-1538, R12)
All fuel rod cladding is manufactured in accor dance with Grade RA-2, ASTM B811, Wrought Zirconium and Zirconium Alloy Seamless and Welded Tubes for Nucl ear Service, except additional limits are placed on oxygen, silicon, and iron content. (DRN 02-1538, R12) 4.2.1.2.3 Fuel Rod Component Properties
4.2.1.2.3.1 Zircaloy-4 Bar Stock (DRN 02-1538, R12)
All Zircaloy-4 bar stock is fabricated in accor dance with Grade RA-2, ASTM B351, Hot-Rolled and Cold-Finished Zirconium and Zirconium Alloy Bars, Rod and Wi re for Nuclear Application, with the following exceptions and/or additions:
(DRN 02-1538, R12) a) Chemical Properties
Additional limits are placed on oxygen and silicon content.
b) Metallurgical Properties
- 1) Grain Size
The maximum average grain size is restricted.
4.2.1.2.3.2 Stainless Steel Compression Springs (DRN 02-1538, R12)
All stainless steel springs are fabric ated in accordance with AMS 5688. (DRN 02-1538, R12)
(DRN 02-1538, R12)
(DRN 02-1538, R12)
WSES-FSAR-UNIT-34.2-14Revision 12 (10/02)4.2.1.2.4UO 2 Fuel Pellet Properties 4.2.1.2.4.1 Chemical Composition Salient points regarding the structure, composition, and properties of the UO 2 fuel pellets are discussed in the following subsections. Where the effect of irradiation on a specific item is considered to be of sufficient
importance to warrant reflection in the design or analyses, that effect is also discussed.a)Chemical analyses are pe rformed for the following constituents:1)Total Uranium 2)Carbon 3)Nitrogen 4)Fluorine 5)Chlorine and Fluorine 6)Iron 7)Thorium 8)Nickel(DRN 02-1538)9)Aluminum10)Chromium 11)Silicon 12)Calcium 13)Magnesium 14)Erbium(DRN 02-1538)b)Limits are placed on the oxygen-to-uranium ratio.c)The sum of the calcium + aluminum + silicon contents shall not exceed 300 ppm by weight.
d)The sum of the thermal neutron capture cross-sections of the following impurities shall not exceed a specified equivalent thermal-neutron capture cross-section of natural boron:1)Boron 2)Silver 3)Cadmium 4)Gadolinium 5)Europium WSES-FSAR-UNIT-34.2-15Revision 12 (10/02)6)Samarium7)Dysprosium(DRN 02-1538)8)Erbium(DRN 02-1538)e)The total hydrogen content of finished ground pellets is restricted.f)The nominal enrichment of the fuel pellet will be specified and shall be held within
+/- 0.05 wt percent U 235.4.2.1.2.4.2 Microstructure(DRN 02-1538)a)Acceptable porosity distribution will be determined by comparison of approved visual standards withphoto-micrographs from each pellet lot.(DRN 02-1538)b)The average grain size shall exceed a specified minimum size.4.2.1.2.4.3Density(DRN 02-1538)a)The density of the sintered pellet after grinding shall be between 94.0 and 96.5 percent of theoretical density (TD), based on a UO 2 theoretical density of 10.96 g/cm 3.b)The in-pile stability of the fuel is ensured by the use of an NRC-approved out-of-pile test during production.(DRN 02-1538)c)The effects of irradiation on the density of sintered UO 2 pellets are discussed in Reference 14.
4.2.1.2.4.4 Thermal Propertiesa)Thermal Expansion The thermal expansion of UO 2 is described by the following temperature dependent equations:
(15)(16)%Linear Expansion= (-1.723 x 10
-2) + (6.797 x 10
-4 T)+ (2.896 x 10
-7 T 2)(25 T 2200)%Linear Expansion = 0.204 + (3 x 10
-4T) + (2 x 10
-7 T 2)(10-10 T 3)(T > 2200) where T = fuel temperature, degrees Celsius.
WSES-FSAR-UNIT-3 4.2-16 Revision 14 (12/05)b) Thermal Emissivity A value of 0.85 is used for the thermal emissivity of UO 2 pellets over the temperature range 800 to 2600K. (See References 17, 18 and 19.) c) Melting Point and Thermal Conductivity The fuel temperature required to incur melting is linearly dependent on local burnup as given by: T = 5080 -290 x (Burnup) 50,000 melt(DRN 04-1096, R14) where, T melt is in F and burnup is in MWD/MTU. This equation T melt is based on UO 2 melt data given by Reference 76. In addition, the fuel melting temperature may be reduced depending on the amount and type of burnable poison in the fuel as described in Reference 78. (DRN 04-1096, R14) The variation of the thermal conductivity of UO 2 with burnup is not explicitly treated, but is implicitly taken from the porosity relationship discussed in Subsection 2.2.5 of Reference 14. d) Specific Heat of UO 2 The specific heat of UO 2 is described by the following temperature dependent equations.
(20)(DRN 04-1096, R14)
T F 22402 6 3-p 460)+T (10 x 3.2432-10 x 2.2784+49.67=C T(DRN 04-1096, R14)
T F 2240 )T 10 x (2.483-)T 10 x (3.1786+)T 10 x (1.399-)T (0.2621+126.07-C 4 12-8-2-4 p 3where: Cp = specific heat, BTU/lbm-F T = fuel temperature, F4.2.1.2.4.5 Mechanical Properties a) Young's Modulus of Elasticity The Young's modulus of elasticity for UO 2 is used in the analytical model for prediction of the effects of pellet clad interaction. Its value may be found in Reference 21. Subsection 4.2.3.2.11 discusses the pellet clad interaction model.
WSES-FSAR-UNIT-3 4.2-17 Revision 302 (12/08) b) Poisson's Ratio (DRN 02-1538, R12)
Poisson's ratio = 0.32 - (1.8 x 10
-5 (T-25)) for the range of temperature between 25 C to 1800 C, Poisson's ratio is assumed constant at 0.29 where T = fuel temperature, C. (DRN 02-1538, R12) c) Yield Stress (not applicable)
d) Ultimate Stress (not applicable)
e) Uniform Ultimate Strain (not applicable)
4.2.1.2.5 Fuel Rod Pressurization
Fuel rods are initially pressurized with helium for two reasons:
a) Preclude clad collapse during the design life of the fuel. The internal pressurization, by reducing stresses from differential pressure, extends the time required to produce creep collapse beyond the required service life of the fuel.
b) Improve thermal conductivity of the pellet-to-c lad gap within the fuel rod. Helium has a higher coefficient of conductivity than the gaseous fission products.
In unpressurized fuel, the initially good helium conduc tivity is eventually degraded through the addition of the fission product gases released from the pellets.
The initial helium pressurization results in a high helium to fission products ratio over the design life of the fuel with a corresponding increase in the gap conductivity and heat transfer.
The effect of fuel rod power level and pin bur nup on fuel rod internal pressure has been studied parametrically. Figures 4.2-3 and 4.2-4 show predicted variation of fuel rod internal pressure with pin
burnup and pin peaking factor for minimum pressure r ods and maximum pressure rods, respectively, for a full power core.
(DRN 03-2058, R14; 06-1059, R15; EC-9533, R302)
The initial helium fill pressure will be 395 15 psia for UO 2 and Erbia rods. Due to the design changes associated with the NGF rods, the initial helium fill pressure for UO 2 rods of the NGF design is specified as 275 +/- 15 psig at 75F. This initial fill pressure will be sufficient to prevent clad collapse as discussed in Subsection 4.2.3.2.
- 8. The calculational methods employed to generate internal pressure histories are discussed in Reference 14. (DRN 03-2058, R14)
The ZrB 2 IFBA rod for both non-NGF and NGF designs is pre-pressurized at a lower helium fill pressure (approximately 150 psig) to prevent an unacceptabl e maximum pressure due to an increased helium release. (DRN 06-1059, R15; EC-9533, R302) 4.2.1.2.5.1 Capacity fo r Fission Gas Inventory The greater portion of the gaseous fission products remain either within the lattice or the microporosity of the UO 2 fuel pellets and do not contribute to the fuel rod internal pressure. However, a fraction of the fission gas is released from the pellets by diffusi on and pore migration and thereafter contributes to the internal pressure. (DRN 06-1059, R15)
The annular pellets provide additional void volume to help control the rod pressure increases due to release of Helium from the thin IFBA c oating during the lifetime of such rod. (DRN 06-1059, R15)
WSES-FSAR-UNIT-3 4.2-18 Revision 15 (03/07)
The determination of the effect of fission gas generated in and released from the pellet column is discussed in Subsection 4.2.3.2.2. The rod pressure increase which results from the release of a given quantity of gas from the fuel pellets depends upon the amount of open void volume available within the fuel rod and the temperatures associated with the various void volumes. In the fuel rod design, the void
volumes considered in computing internal pressure are:
Fuel rod upper end plenum
Fuel-clad annulus
Fuel pellet-end dishes and chamfers
Fuel pellet open porosity
(DRN 06-1059, R15)
Hollow center of annular pellets (DRN 06-1059, R15)
These volumes are not constant during the life of the fuel. The model used for computing the available volume is a function of burnup and power level and accounts for the effects of fuel and clad thermal expansion fuel pellet densification, clad creep, and irradiation induced swelling of the fuel pellets.
4.2.1.2.5.2 Fuel Rod Plenum Design
The fuel rod upper end plenum is required to serve the following functions:
a) Provide space for axial thermal expansion and burnup swelling of the pellet column.
b) Contain the pellet column holddown spring.
c) Act as a plenum region to ensure an acceptable range of fuel rod internal pressures. (DRN 02-1538, R12)
Of these functions, listing c is expected to be the most limiting constraint on plenum length selection, since the range of temperatures in fuel rod, together with the effects of swelling, thermal expansion, and fission gas release, can produce a wide range of internal pressure during the life of the fuel. The fuel rod plenum pressure will be consistent with the pressurization and clad collapse criteria specified in
Subsection 4.2.1.2.1. (DRN 02-1538, R12) 4.2.1.2.5.3 Outline of Procedure Used to Size the Fuel Rod Plenum
a) A parametric study of the effects of plenum length on maximum and minimum rod internal pressure is performed. Because the criteria pertaining to maximum and minimum rod internal
pressure differ, the study is divided into two sections:
WSES-FSAR-UNIT-3 4.2-19 Revision 309 (06/16)
- 1) Maximum Internal Pressure Calculation
Maximum rod pressure is limited by the stre ss criteria. Maximum end-of-life pressure is determined for each plenum length by includi ng the fission gas released, selecting conservative values for components dim ensions and properties, and accounting for burnup effects on component dimensions.
The primary cladding stress produced by each maximum pressure is then compared to t he stress limits to find the margin available with each plenum length. Stress limits are listed in Subsection 4.2.1.2.1.
- 2) Minimum Internal Pressure/Collapse Calculation
Minimum rod pressure is limited by the crit erion that no rod will be subject to collapse during the design lifetime. The minimum pr essure history for each plenum length is determined by neglecting fission gas release, se lecting a conservative combination of component dimensions and properties, and a ccounting for dimension changes during irradiation. Each minimum pressure histor y is input to the cladding collapse model to establish the acceptability of the asso ciated plenum length (see Reference 22).
b) For each plenum length, there is a resultant range of acceptable initial fill pressures. The optimum plenum length is generally considered to be the shortest which satisfies all criteria related to maximum and minimum rod internal pressure including a range sufficient to
accommodate a reasonable manufacturing tolerance on initial fill pressure. (LBDCR 13-014, R309) c) Additional information on those factors wh ich have a bearing on determination of the plenum length are discussed below:
- 1) Creep and dimensional stability of the fuel rod assembly influence the fission gas release model and internal pressure calculations , and are accounted for in the procedure of sizing the fuel rod plenum length. Creep in the cladding is accounted for in a change in clad inside diameter, which in turn influenc es the fuel/clad gap. The gap change varies the gap conductance in the FATES computer code with resulting change in annulus temperature, internal pressure, and fission gas release (see Reference 14). In addition, the change in clad inside diameter causes a change in the internal volume, with its resulting effect on temperature and pressure.
Dimensional stability considerations affect the internal volume of the fuel rod, causing changes in internal pressure and temperature. Fuel pellet densification r educes the stack height and pellet diameter.
Irradiation-induced radial and axial swelling of the fuel pellets decreases the internal volume within the fuel rod. In-pile growth of the fuel rod cladding contributes to the internal volume. Axial and radial elastic deformation calculations for the cladding are based on the differential pressure the cladding is exposed to, resulting in internal volume changes. Thermal relocation, as well as diffe rential thermal expansion of the fuel rod materials also affect the internal volume of the fuel rods. (LBDCR 13-014, R309)
WSES-FSAR-UNIT-3 4.2-20 Revision 302 (12/08)
(DRN 02-1538, R12; 04-1096, R14)
- 2) The maximum expected fission gas release in the peak power rod is calculated using the FATES computer code. Rod power history input to the code is consistent with the design limit for peak linear heat rate set by LOCA considerations, and therefore the gas release used to size the plenum represents an upper limit. Because of time-varying gap conductance, fuel depletion, and expected fuel management, the release rate varies as a function of burnup. (DRN 02-1538, R12; 04-1096, R14) 4.2.1.2.6 Fuel Rod Performance
Steady state fuel temperatures are determined by the FATES computer program. The calculational procedure considers the effect of linear heat rate, fuel relocation, fuel swelling, densification, thermal expansion, fission gas release, and clad deformati ons. The model for predicting fuel thermal performance including the specific effects of fuel densification on increased linear heat generation rate (LHGR) and stored energy is discussed in detail in Reference 14. (DRN 02-1538, R12)
Significant parameters such as cold pellet and cl ad diameters, gas pressure and composition, burnup and void volumes are calculated and used as initial conditions for subsequent calculations for stored energy during the ECCS analysis. The coupling me chanism between FATES calculations and the ECCS analysis is described in detail in Reference 23.
(DRN 02-1538, R12)
Discussions of uncertainties associated with the model, and of comparative analytical and experimental results, are also included in Reference 14. (DRN 02-1538, R12; EC-9533, R302)
The methodology for modeling the NGF design is de scribed in the CE 16x16 Next Generation Fuel Topical Report, Reference 84. (EC-9533, R302) 4.2.1.2.7 Fuel Rod with Erbia (Er 2 O 3) Addition Some fuel rods in the fuel assembly may contain pellets which incorporate erbia (Er 2 O 3) as a burnable absorber into the central portion of the pellet column.
These fuel rods are analyzed by the same methods and subject to the same design criteria as fuel rods containing only urania pellets.
The urania-erbia pellets are fabricated by mechanically blending erbia powder with urania powder to
produce a homogenous mixture, followed by pressing and sint ering. These fuel pellets may contain up to 2.5 weight percent erbia. (DRN 02-1538, R12)
(DRN 04-1096, R14)
The addition of erbia to urania fuel pellets may influenc e the thermal properties of the fuel. Of particular importance are the properties that are used in fuel performance analyses. These properties are: 1) solidus temperature, 2) specific heat, 3) density, 4) thermal expansion, and 5) thermal conductivity. The effect of erbia addition on these properties of urania is discussed in detail in Section 2.2 of Reference 78. (DRN 04-1096, R14)
(DRN 06-1059, R15) 4.2.1.2.8 Fuel Rod with IFBA (ZrB 2 coated) Pellets
The Zirconium Diboride (ZrB
- 2) integral fuel burnable absorber (IFBA) fuel design commences with Batch Y for Cycle 15. The ZrB 2 is applied as a very thin uniform coat ing on the outer surface only of the solid UO 2 pellet stack prior to loading into the fuel rod claddi ng tube. The coating is applied over the center of the UO 2 pellet stack length, consistent with positioning of the Erbia (Er 2 O 3 - UO 2) burnable absorber pellets in the prior batches present in Cycle 15 (T , U, W and X) and does not extend to either end of the fuel rod (see Figure 4.2-11A). Pellets at the ends of the pellet stack (cutback zones) are of an annular design. (EC-9533, R302)
The annular pellets have the same pellet outside di ameter (.3250 inch for pre-NGF batches and .3225 inch for NGF batches) and pellet edge chamfer as t he corresponding enriched solid fuel pellets, but have no dish on the pellet ends. The annular pellets are also longer (DRN 06-1059, R15; EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-21 Revision 302 (12/08)
(DRN 06-1059, R15; EC-9533, R302) than the solid fuel pellets (.500 inch versus .
390 inch for pre-NGF batches and 0.387 inch for NGF batches). The diameter of the annulus is 0.1625 inc hes (pre-NGF) or 0.1550 inches (NGF) which results in about 25% annular volume to accommodate gas rel ease in the IFBA rods. The fully-enriched annular pellets in the IFBA rods increase t he void volume for gas accommodation within the fuel rod compared to the previous burnable absorber fuel rod design (Erbia), thereby providing sufficient margin to meet the rod internal pressure criterion. Also, to compensate for the additional helium released from the ZrB 2 coating, the initial fill gas pressure, designed to reduce pressure differences across the cladding, is reduced as compared to non-IFBA rods. (EC-9533, R302)
Introduction of the ZrB 2 IFBA fuel rod design has influenced fuel rod pressurization as discussed in the Topical Report, Reference 81. During irradiation, the B-10 isotope absorbs a neutron and fissions into Helium and Lithium. Much of the Helium may be released from the thin coating into the fuel rod void by the time complete burnout is attained, thus additionally increasing the rod internal pressure at end of life.
The released Helium compensates for the initial r eduction in helium fill gas and mitigates the potential impact of less helium fill gas on the thermal heat transfe r from the fuel pellets to the cladding and into the coolant. Thus, the IFBA coating and corresponding Helium release have no significant impact on the heat
transfer characteristic of the fuel rod. (DRN 06-1059, R15) 4.2.1.3 Burnable Poison Rod (DRN 06-1059, R15)
The earlier cycles poison rods containing the Al 2 O 3 burnable poison pellets were replaced by fuel rods with Erbia (Er 2 O 3 - UO 2) burnable absorber pellets (Section 4.2.1.
2.7) during the late 1990's. Most recently, the ZrB 2 IFBA fuel rod design (Section 4.2.1.2.8) is being introduced beginning with Batch Y in Cycle 15, such that the current core design uses only Erbia or IFBA burnable absorbers rods as poison rods. Hence, the previous design of the poison rods containing the Al 2 O 3 burnable poison pellets, as presented in subsections 4.2.1.3.1 th rough .2.1.3.3.3, is only relevant to those poison rods, if any, that are being kept in long term storage outside of the current core. (DRN 06-1059, R15) 4.2.1.3.1 Burnable Poison Rod Cladding Design Limits
The burnable poison rod design accounts for external pr essure, differential expansion of pellets and clad, pellet swelling, clad creep, helium gas release, initia l internal helium pressure, thermal stress, and flow-induced vibrations. Except where specifically noted, the design bases presented in this section are consistent with those used for previous designs. The structural criteria for the normal, upset and emergency loading combinations identified in Subsection 4.2.1.1. are as follows:
a) During normal operating and upset conditions, the ma ximum primary tensile stress in the Zircaloy clad shall not exceed two-thirds of the minimum uni rradiated yield strength of the material at the applicable temperature. The corresponding limit under emergency conditions is the material yield strength.b) Net unrecoverable circumferential strain shall not exceed one percent as
predicted by computations considering cl ad creep and poison pellet swelling effects.
c) The clad will be initially pressurized with he lium to an amount sufficient to prevent gross clad deformation under the combined effects of exter nal pressure and long-term creep. The clad design will not rely on the support of pellets or the holddown spring to prevent gross deformation.
4.2.1.3.2 Burnable Poison Rod Cladding Properties
Cladding tubes for burnable poison rods are purchas ed under the specification for fuel rod cladding tubes. Therefore, the mechanical metallurgical c hemical, and dimensional properties of the cladding are as discussed in Subsection 4.2.1.2.2.
WSES-FSAR-UNIT-3 4.2-22 Revision 15 (03/07)
(DRN 03-2058, R14) 4.2.1.3.3 A2 O 3-B 4 C Burnable Poison Pellet Properties (DRN 02-1538, R12)
The A2 O 3-B 4C burnable poison pellets used in C-E designed reactors consist of a relatively small volume fraction of fine B 4 C particles dispersed in a continuous A2 O 3 matrix. The boron loading is varied by adjusting the B 4 C concentration in the range from 0.7 to 4.0 w/o (1 to 6.0 v/o). Typical pellets have a bulk density of about 90 percent of theoretical. Many properties of the two-phase A2 O 3-B 4 C mixture, such as thermal expansion, thermal conductivity, and specific heat are very similar to the properties of the A2 O 3 major constituent. In contrast, properties such as swelling, helium release, melting point and corrosion are dependent on the presence of B 4 C. The operating centerline temperature of burnable poison is less than 1100F, with maximum surface temperatures close to 750F. (DRN 02-1538, R12) 4.2.1.3.3.1 Thermal-Physical Properties
a) Thermal Expansion
The mean thermal expansion coefficients of A2 O 3 and B 4 C from 0 to 1850F are 4.9 and 2.5 in/in.-F x 10-6, respectively (see References 24 and 25). The thermal expansion of the A2 O 3-B 4 C two-phase mixture can be considered to be essentially the same as the value for the continuous A2 O 3 matrix, as the dispersed B 4 C phase has a lower expansion coefficient and occupies no more than 6 v/o of the available volume. The low temperature (80 to 250F) thermal expansion coefficient of A2 O 3 irradiated at 480, 900, and 1300F does not change as a result of irradiation (see Reference 26). The expansion of a similar material, beryllium oxide, up to 1900F has also been reported to be relatively unchanged by irradiation (see Reference 27). It is therefore appropriate to use the values of thermal expansion
measured for A2 O 3 for the burnable poison pellets: (DRN 03-2058, R14)
Temperature Range Linear Expansion (F) (percent)
400 0.12
600 0.23
800 0.30
1000 0.40 b) Melting Point
(DRN 03-2058, R14) The melting points of A2 O 3 (3710F) and B 4 C (4440F) are higher than the melting point of the Zr-4 cladding (see References 28 and 29). No reactions have been reported between the component which would lower the melting point of the pellets to any significant extent. As the
B 4C burns up, the lithium atoms formed occupy interstitial sites randomly distributed within the B 4 C lattice, rather than forming a lithium-rich phase (see Reference 30). The solid solution of lithium in B 4 C should not appreciably influence the melting point of the A2 O 3-B 4 C pellets, as only a small quantity of lithium compounds (0.5 w/o) forms during irradiation. It is concluded that the melting point of A2 O 3-B 4 C will remain considerably above the maximum 1100F operating temperature. (DRN 03-2058, R14)
WSES-FSAR-UNIT-3 4.2-23 Revision 15 (03/07)
(DRN 03-2058, R14) c) Thermal Conductivity The thermal conductivity of A2 O 3-B 4 C was calculated from the measured values for A2 O 3 and B 4 C using the Maxwell-Buckan relationship for a continuous matrix phase (A2 O 3) with spherical dispersed phase (B 4C) particles (see Reference 31). Because of the high A2 O 3 content of these mixtures and the similarity in thermal conductivity, the resultant values for A2 O 3-B 4 C were essentially the same as the values for A2 O 3. The measured, unirradiated values of thermal conductivity at 750F are 0.06 cal/sec-cm-K for B 4 C and 0.05 cal/sec-cm-K for A2 O 3. The thermal conductivity of A2 O 3 after irradiation decreases rapidly as a function of burnup to values of about one-third the unirradiated values (see Reference 26). The irradiated values of A2 O 3-B 4 C calculated from the above relationships are given below as a function of temperature (see References 26 and 32). (DRN 03-2058, R14)
Temperature Thermal Conductivity (F) (cal/sec-cm-K) 400 0.015
600 0.013
800 0.010
1000 0.008
d) Specific Heat
(DRN 03-2058, R14)
The specific heat of the A2 O 3-B 4 C mixture can be taken to be essentially the same as pure A2 O 3 since the concentration of B 4 C is low (6.0 v/o maximum). In addition, the effect of irradiation on specific heat is expected to be small based on experimental evidence from similar materials which do not sustain transmutations as a function of neutron exposure. (DRN 03-2058, R14) (DRN 03-2058, R14)
The values for A2 O 3 measured on unirradiated samples (32)(33) are given below:
Temperature Specific Heat (F) (cal/gm-F) 250 0.12
450 0.13
800 0.14
1000 and above 0.15 (DRN 03-2058, R14)
WSES-FSAR-UNIT-3 4.2-24 Revision 15 (03/07)
(DRN 03-2058, R14) 4.2.1.3.3.2 Irradiation Properties
a) Swelling
A2 O 3-B 4 C consists of B 4 C particles dispersed in a continuous A2 O 3 matrix, which occupies more than 94 percent of the poison pellet. The swelling of A2 O 3-B 4 C depends primarily upon the neutron fluence on the continuous A2 O 3 matrix and, secondarily, on the B 10 burnup of the dispersed B 4 C phase. Recent measurements performed on material containing about two w/o B 4 C irradiated in a C-E PWR to 100 percent B 10 burnup at a fluence of 2.4 x 10 21 nvt (E>0.8 MeV) revealed a diametral swelling of about one percent. Pellets similar to the burnable poison
used in C-E reactors with up to 3 w/o B 4 C also sustained about 100 percent B 10 burnup.
Experimental data (34) on A2 O 3 reveal a diametral swelling of about 0.7 percent at a fluence of 2.4 x 10 21 nvt (E>0.8 MeV). Swelling of A2 O 3 increases linearly with fluence to 1.8 percent diametral after an exposure of 6x10 21 nvt (E>0.8 MeV).
These data show that A2 O 3-B 4 C swells somewhat more than A2 O 3 up to a burnup of 100 percent B 4 C (about 2 x 10 21 nvt, E>0.8 MeV).
The C-E design value of A2 O 3-B 4 C swelling rate for fluences less than 2 x 10 21 is greater than the swelling rate of A2 O 3 , while after 100 percent B 10 burnup the swelling rate for A2 O 3-B 4 C is considered equal to that of A2 O 3. The data and considerations presented above result in best-estimate diametral swelling values at end-of-life (7 x 10 21 nvt, E>0.8 MeV) of about two percent for A2 O 3 and from two to three percent for A2 O 3-B 4 C depending on B 4 C. b) Helium Release
Experimental measurements reveal that less than five percent of the helium formed during irradiation will be released.
(35) These measurements were performed on A2 O 3-B 4 C pellets irradiated at temperatures to 500F and, subsequently, annealed at 1000F for five days. The helium release in a burnable poison rod which operated for (DRN 03-2058, R14) one cycle in a ABB CE PWR was calculated from internal pressure measurements to be less than five percent. The design is based on a release of three to ten percent of the helium generated. The design of the burnable poison rod will not be limited by helium pressure despite
the conservative use of 10 percent release.
4.2.1.3.3.3 Chemical Properties
(DRN 03-2058, R14) a) A2 O 3-B 4 C Coolant Reactions (DRN 02-1538, R12)
The stability of A2 O 3-B 4 C in contact with reactor coolant has been investigated before and after irradiation. Prior to irradiation no significant boron loss was observed after testing for hundreds of hours at 650F in borated water at 2250 psig. Visual and metallographic evaluations showed no erosion of the A2 O 3 matrix. In addition, pellet measurements showed no change in diameter or length as a result of exposure to the borated water. (DRN 02-1538, R12; 03-2058, R14)
WSES-FSAR-UNIT-3 4.2-25 Revision 15 (03/07)
(DRN 03-2058, R14) A series of tests were performed to assess the compatibility of irradiated A2 O 3-B 4 C with reactor coolant. The results of these tests show that A2 O 3-B 4 C pellets irradiated to 100 percent B 10 burnup retain their mechanical integrity after 350 hours0.00405 days <br />0.0972 hours <br />5.787037e-4 weeks <br />1.33175e-4 months <br /> in 650F, 2250 psig water.
Visual and metallographic observations indicate that the A2 O 3 matrix does not sustain significant erosion of micro-cracking, although the B 4 C particles are leached out of portions of the pellet. No diameter or length changes were noted in the pellets. The amount of B 4 C loss is primarily dependent upon the accessibility of the B 4 C particles to the reactor coolant, and the time of exposure. B 4 C particles that are completely enclosed in the A2 O 3 matrix do not corrode, as the A2 O 3 matrix material has relatively good corrosion resistance.
Should irradiated B 4 C particles be exposed to reactor coolant, the primary corrosion products that would be produced are H 3 BO 3 and Li 2 O, which are soluble in water, and free carbon. The presence of these products in the reactor coolant would not be detrimental to the operation of the
plant.
b) Chemical Compatibility
Chemical compatibility between the A2 O 3-B 4 C pellets and the burnable poison rod cladding during long-term normal operations has been demonstrated by examinations of a burnable poison rod from the Maine Yankee Reactor. The rod had been exposed to an axial average
fluence in excess of 2 x 10 21 nvt (>0.821 MeV). No evidence of a chemical reaction was observed on the cladding I.D.
Short term chemical compatibility during upset and emergency conditions is demonstrated by the fact that conditions favorable to a chemical reaction between B 4 C and A2 O 3 are not present at temperatures below 1300F (36) This temperature is higher than that which will occur at burnable poison pellet surfaces during Condition II and III occurrences (Subsection 4.2.1.1). The action between Zr-4 and A2 O 3 described by Idaho Nuclear (37) was observed to occur rapidly only at temperatures in excess of 2500F, well above the peak Condition IV Zr-4 temperatures in the higher energy fuel rods described in Chapter 15. (DRN 03-2058, R14) 4.2.1.4 Control Element Assembly
Except where specifically noted, the design bases presented in this section are consistent with those
used for previous designs.
The mechanical design of the control element assemblies is based on compliance with the following
functional requirements and criteria:
a) To provide for or initiate short term reactivity control under all normal and adverse conditions experienced during reactor start-up, normal operation, shutdown, and accident conditions.
b) Mechanical clearances of the CEA within the fuel and reactor internals are such that the requirements for CEA positioning and reactor trip are attained under the most adverse
accumulation of tolerances.
c) Structural material characteristics are such that radiation induced changes to the CEA materials will not impair the functions of the reactivity control system.
WSES-FSAR-UNIT-3 4.2-26 Revision 15 (03/07) 4.2.1.4.1 Thermal-Physical Properties of Absorber Material (DRN 00-644; 01-1103, R12) The primary control rod absorber materials consist of boron carbide pellets (B 4 C) and silver-indium-cadmium bars (Ag-In-Cd). Refer to Figures 4.2-5, 4.2-6, and 4.2-7 for the specific application and orientation of the absorber materials. The significant thermal and physical properties used in mechanical
analysis of the absorber materials are listed below: (DRN 00-644; 01-1103, R12) a) Boron Carbide (B 4 C) Configuration Right cylinder
Outside diameter in. 0.737 0.001 Pellet length, in. nominal 2
End chamber 0.03 in. by 45 Density gm/cc 1.84
w/o boron, minimum 77.5
Percent open porosity in 27 pellet
Ultimate tensile strength, psi N/A Yield strength, psi N/A
Elongation, percent N/A
Young's modulus, psi N/A
Thermal conductivity (cal/sec-cm-C): Irradiated Unirradiated 800F 8.3 x 10
-3 28 x 10-3 1000F 7.9 x 10
-3 24 x 10-3 Melting point, F 4400 Percent thermal linear 0.23% @ 1000F expansion
b) Silver-Indium-Cadmium (Ag-In-Cd)
Configuration Cylindrical bars with central hole
Outside diameter, in. 0.734 0.003 Inside diameter, in. 1/4
WSES-FSAR-UNIT-3 4.2-27 Revision 15 (03/07)
Length of bar, in. nominal 12.5 (for 5 element CEAs),
5 (for 4 element CEAs)
Density, lb/in.3 0.367
Ultimate tensile strength, psi N/A
Yield strength, psi N/A
Elongation, percent N/A
Young's modulus, psi N/A
Thermal conductivity (cal/sec-cm-C): Irradiated Unirradiated at 300C 0.14 0.182 at 400C 0.148 0.196 Melting point, F 1,470
Linear thermal expansion (in./in.-F0 12.5 x 10
-6 c) Inconel Alloy 625 (Ni-Cr-Fe)
Configuration (as absorber) Cylindrical bar
Outside diameter, in. 0.816 0.002 Inside diameter, in. Solid
Length of cylinder, in. See Figures 4.2-5, 4.2-6, 4.2-7
Density, lb/in.
3 0.305 Ultimate tensile Strength, psi 120-150
Specified minimum yield strength @ 650F, ksi 65
Elongation in two in.,
percent 30
Young's modulus, psi
at 70F 29.7 x 10 6 at 650F 27.0 x 10 6 Thermal conductivity (Btu/hr-ft-F):
WSES-FSAR-UNIT-3 4.2-28 Revision 15 (03/07) 70F 5.7 600F 8.2 (DRN 00-644) Linear thermal expansion 7.4 x 10
-6 (in./in.-F) (70 to 600F) (DRN 00-644) 4.2.1.4.2 Compatibility of Absorber and Cladding Materials
The cladding material used for the control elements is Inconel Alloy 625. The selection of this material for use as cladding is based on considerations of strength, creep resistance, corrosion resistance, and dimensional stability under irradiation and also upon the acceptable performance of this material for this application in other ABB CE reactors currently in operation.
a) B 4 C/Inconel 625 Compatibility
Studies have been conducted by HEDL(38) on the compatibility of Type 316 stainless steel with B 4C under irradiation for thousands of hours at temperatures between 1300 and 1600F. Carbide formation to a depth of about 0.004 in. in the Type 316 stainless steel was measured after 4400 hours0.0509 days <br />1.222 hours <br />0.00728 weeks <br />0.00167 months <br /> at 1300F. Similar compound formation depths were observed after ex-reactor bench testing. After testing at 1000F, only 0.0001 in/yr of penetration was measured. Since Inconel 625 is more resistant to carbide formation than 316 stainless steel, and the expected pellet/clad interfacial temperature in the Waterford 3 design is below 800F, it is concluded that B 4 C is compatible with Inconel.
4.2.1.4.3 Cladding Stress-Strain Limits
The stress limits for the Inconel Alloy 625 cladding are as follows:
Design Conditions I and II (Non Operation, Normal Operation, and Upset Conditions)
P m S m P m P b F sS m Design Condition III (Emergency Conditions)
P m 1.5 S m P m P b 1.5 F sS m Design Condition IV (Faulted Conditions)
P m S'm P m + P b F sS m' where S'm is the smaller of 2.4S m or 0.7S u For definition of P m , P b , S m , S'm , S u , and F s see Subsection 4.2.1.1.1. For the Inconel 625 CEA cladding, the value of S m is two-thirds of the minimum specified yield strength at temperature.
For Inconel 625, the specified minimum yield strength is 65,000 psi at 650F.
WSES-FSAR-UNIT-3 4.2-29 Revision 15 (03/07)
F s = Mp/My where Mp is the bending moment required to produce a fully plastic section and My is the bending moment which first produces yielding at the extreme fibers of the cross section.
The capability of cross-sections loaded in bending to sustain moments considerably in excess of that required to yield the outermost fiber is discussed in Reference 1. For the CEA cladding
dimensions, F s = 1.33.
The strain of the cladding is limited to a value which will permit the CEAs to trip within the allowable time and which is less than the irradiated uniform elongation of the material.
The values of uniform and total elongation of Inconel Alloy 625 cladding are as follows:
Fluence (E>1 MeV), nvt 1 x 10 22 3 x 10 22 Uniform elongation, percent 3 1
Total elongation, percent 6 3
4.2.1.4.4 Irradiation Behavior of Absorber Materials
a) Boron Carbide Properties
- 1) Swelling. The linear swelling of B 4 C increases with burnup according to the relationship:
%L = (0.1) B 10 Burnup, a/o (DRN 00-644;06-895, R15)
This relationship was obtained from experimental irradiations on high density (90 percent theoretical density) wafers (39) and pellets with densities ranging between 71 and 98 percent TD.
(38)(40) Dimensional changes were measured as a function of burnup, after irradiating at temperatures expected in the Waterford 3 design. (DRN 00-644;06-895, R15)
- 2) Thermal Conductivity. The thermal conductivity of unirradiated 73 percent dense B 4 C decreases linearly with temperatures from 300 to 1600F, according to the relationship:
= 1 2.17(6.87 + 0.017 )cal/cm-K-sec T This relationship was obtained from measurements performed on pellets ranging from 70
to 98 percent TD.
(41) The relationship between the thermal conductivity of irradiated 73 percent TD B 4 C pellets and temperature given below was derived from measured values (41) on higher density pellets irradiated to fluences out to 3 x 10 22 nvt (E > 1 MeV).
= 1 2.17(38 + 0.025 ) cal/cm-K-sec T where T = temperature, K Thermal conductivity measurements of 17 B 4 C specimens with densities ranging from 83 to 98 percent TD, irradiated at temperatures from 930 to 1600F showed that thermal conductivity decreased significantly after irradiation. The rate of decrease is high at the lower irradiation temperatures, but saturates rapidly with exposure.
WSES-FSAR-UNIT-3 4.2-30 Revision 15 (03/07)
- 3) Helium Release. Helium is formed in B 4 C as B 10 burnup proceeds. The fraction of helium released from the pellets is important for determining rod internal gas pressure.
The relationship between helium release and irradiation temperature given below was
developed at ORNL (42) to fit experimental data obtained from thermal reactor irradiations. (DRN 00-644) 5RT-Q e e 1.85D-A release He % (DRN 00-644) where:
A = Constant, 6.69 for ABB CE pellets
D = Fractional density, 0.73 for ABB CE pellets
Q = Activation energy constant, 3600 cal/mole (DRN 06-895, R15)
R = Gas constant, 1.98 cal/mole -K (DRN 06-895, R15)
T = Pellet temperature, K This expression becomes
5 + 208 = release He %(-1820/T)e when the above parameters are substituted. In this form, design values for helium release as a function of temperature are generated. The five percent helium release allowance (the last term in the expression) was added to ensure that design values lie above all reported helium release data. Calculated values of helium release obtained from the recommended design expression lie above all experimental data
points (38)(43)(44) obtained on B 4 C pellet specimens irradiated in thermal reactors.
- 4) Pellet Porosity. Experimental evidence is available (45) which shows that for pellet densities below 90 percent, essentially all porosity is open at beginning-of-life. Irradiation induced swelling does not change the characteristics of the porosity, but only changes
the bulk volume of the specimens. Therefore, the amount of porosity available at end-of-life is the same as that present at beginning-of-life.
b) Silver-Indium-Cadmium Properties
- 1) Swelling. Measurements performed on Ag-In-Cd rods irradiated at fluences up to 6.2 x 10 21 nvt (E>0.6 eV) were employed to develop the following expression to predict the volumetric swelling for silver-indium-cadmium alloy:
% V = 0.3 10 21 where = fluence, nvt (E>0.6 eV).
WSES-FSAR-UNIT-34.2-31Linear swelling is approximately one-third of the volumetric swelling.2)Thermal Conductivity. The increase in cadmium content from five to perhaps 10 w/o, andthe formation of two to three w/o tin as a result of long-term exposures, is expected to decrease the thermal conductivity from the accepted(46) unirradiated values. Publisheddata for unirradiated Ag-Cd binary alloys shows that thermal conductivity was decreasd by about 20 percent by increasing the cadmium content from 5.0 to 10.0 w/o.(46) Sinceirradiated Ag-In-Cd is expected to perform in much the same fashion, the unirradiatedvalues of thermal conductivity are decreased by 25 percent to account for irradiation.3)Linear Thermal Expansion. The coefficient of linear thermal expansion for unirradiatedAg-In-Cd material is 12.5 x 10
-6 in./in.-°F over the temperature range of 70 to 930
°F(47)Published data on unirradiated(46) Ag-Cd binary alloys reveal that a cadium increase offive percent will result in about a five percent increase in thermal expansion coefficient.
The small changes in indium and tin content do not influence the thermal coefficientappreciably. For simplicity, the irradiated value of 13.1 x 10
-6 in./in.-°F is used in alldesign calculations.4)Melting Point. The melting point of unirradiated Ag-In-Cd has been measuredas 1470 +/- 30°F(46) (800 +/- 17°C). The formation of three w/o tin due to the transmutationof indium and the increase in cadmium content to about 10 w/o due to the transmutation of silver may result in a small decrease in the melting point.c)Inconel 625 Properties1)Swelling. Available information indicates that Inconel 625 is highly resistant to radiationswelling. Exposure of Inconel 625 to a fluence of 3 x 10 22 nvt (E>0.1 MeV) at atemperature of 400
°C (752°F) showed no visible cavities in metallographicexaminations(48) so that swelling, if any, would be very minor. Direct measurementsmade after exposure of Inconel 625 to fluence of 5 x 10 22 nvt (E>0.1 MeV) at LMFBRconditions showed no evidence, of swelling.(49) Thus, Inconel 625 after fluences of 3 x 10 22 nvt (>0.1 MeV) is not expected to swell.2)Ductility. The ductility of Inconel 625 decreases after irradiation. Extrapolation of lowerfluence data on Inconel 625 and 500 indicates that the values of uniform and total elongation of Inconel 625 after 1 x 10 22 nvt (E>1 MeV) are three and six percent,respectively.
WSES-FSAR-UNIT-34.2-32 Revision 12 (10/02)4.2.1.5 Surveillance Program 4.2.1.5.1 Requirements for Surveillance and Testing of Irradiated Fuel Rods(DRN 02-1538)High burnup performance experience, as described in Subsection 4.2.2 has provided evidence that the fuel will perform satisfactorily under the design conditions. The current core design bases do not include a specific requirement for testing of irradiated fuel rods. However, the fuel assembly design allows
disassembly and reassembly to facilitate such inspections, should the need arise.(DRN 02-1538)A fuel rod irradiation program has been developed to evaluate the performance of the fuel rods designed for use in the 16 x 16 fuel assembly. The program includes the irradiation of six standard 16 x 16 assemblies, two each for one, two, and three cycles, respectively, in the Arkansas Nuclear One Unit 2 reactor (ANO-2).
Each assembly will contain a minimum of 50 precharacterized, removable rods distributed within the assembly to obtain a spectrum of exposure levels for evaluation purposes in interim and terminal examinations. Interim examination of all six assemblies is planned during refueling shutdowns after each cycle.(DRN 02-1538)The ANO-2 fuel rods and specific components of the fuel rods will receive detailed precharacterizations. The program calls for substantial cladding characterization to include mechanical properties, texture, hydride orientation and out-of-reactor low strain rate behavior. In addition to the ID and OD dimensional data normally obtained on the clad tubing material, a minimum of 300 fuel rods will be profiled to obtain as-loadeddimensions. Sufficient fuel rods will be pr ofiled to obtain diameter and quality measurements such that changes in these parameters can be tracked by similar measurements during interim inspections. Also, a
random selection of approximately 100 UO 2 pellets from each lot per batch used will be characterizeddimensionally and the density distribution will be determined. About one-half of these pellets will be placed
in known axial locations in selected fuel rods while the remainder will be set aside as archives.A poolside non-destructive examination will be made during each of the first three refuelings at ANO-2.
The six 16 x 16 assemblies with characterized rods will be removed from the reactor at each refueling and moved to the spent fuel pool for leak testing (if failed fuel is in the core) and for visual inspection. The length of the assembly and peripheral rods will be measured. During the shutdown, a target of 20 precharacterized rods per batch will be scheduled for examination and measurement. At some time after the refueling outage, pre-characterized rods retained in discharged assemblies will be measured. A target of 100 rods will be eddy
current tested after each shutdown.(DRN 02-1538)A post irradiation fuel surveillance program for Waterford 3 is planned. This program shall consist of a visual inspection of a minimum of six irradiated assemblies prior to replacement of the Reactor Vessel Head at each of the first three refueling outages. The six assemblies inspected shall consist of two assemblies of
each fuel type and will be from core locations which are non-adjacent. Visual inspections shall consist of viewing the top and sides of each fuel assembly via an underwater TV camera or periscope.
WSES-FSAR-UNIT-34.2-33 Revision 12 (10/02)The visual inspection will include observation with special attention to gross problems involving cladding defects, spacer grid damage and other major structural abnormalities. No special measurement devices for
these effects are intended to be provided for this visual inspection.
If major abnormalities are detected during this visual inspection or if plant instrumentation indicates gross fuel failures, the fuel vendor will be informed and further inspections shall be performed. Depending on the nature of the observed condition, further examination could include fuel sipping, single rod examination and other examinations. The 16 x 16 fuel design enables reconstitution. Individual fuel rods and other structural components may be examined and replaced, if required. Under unusual circumstances, destructive examination of a fuel rod may be required but this would not be accomplished on site or during the refueling outage.The NRC shall be contacted regarding gross fuel failure detected by plant instrumentation or major abnormalities observed during the post irradiation inspections described a bove.(DRN 02-1538)The post fuel irradiation fuel surveillance program shall be continued following the first three cycles of operation of Waterford 3. Six assemblies shall be visually inspected during each refueling outage, not necessarily prior to replacement of the reactor vessel head. The visual inspection shall consist of viewing the tops and sides of each fuel assembly via an underwater TV camera or periscope. The visual inspection
will include observation with special attention to gross problems involving cladding defects, spacer grid damage, and other major structural abnormalities. The NRC will be notified of major abnormalities noted as
a result of these inspection activities.(DRN 02-1538)4.
2.2 DESCRIPTION
AND DESIGN DRAWINGSThis subsection summarizes the mechanical design characteristics of the fuel system and discusses the design parameters which are of significance to the performance of the reactor. A summary of mechanical design parameters is presented in Table 4.2-1. These data are intended to be descriptive of the design; limiting values of these and other parameters will be discussed in the appropriate sections.4.2.2.1Fuel AssemblyThe fuel assembly (Figure 4.2-8) consists of 236 fuel and poison rods, five control element assembly guide tubes, 11 fuel rod spacer grids, upper and lower end fittings, and a hold-down device. The outer guide tubes, spacer grids, and end fittings form the structural frame of the assembly.(DRN 02-1538)The fuel spacer grids (Figure 4.2-9) maintain the fuel rod array by providing positive lateral restraint to the fuel rod but only frictional restraint to axial fuel rod motion. The grids are fabricated from pre-formed Zircaloy or Inconel strips (the bottom, and in some cases the top, spacer grid material is Inconel) interlocked in an egg crate fashion and welded together. Each cell of the spacer grid contains two leaf springs and four arches. The leaf springs press the rod against the arches to restrict relative motion between the grids and
the fuel rods. The perimeter strips contain features designed to prevent hangup of grids during a refueling operation.(DRN 02-1538)
WSES-FSAR-UNIT-3 4.2-34 Revision 309 (06/16)
(DRN 02-1538, R12;04-502, R13)
The Zircaloy-4 spacer grids are fastened to the Zi rcaloy-4 guide tubes by welding, and each grid is welded to each guide tube at eight locations, four on t he upper face of the grid and four on the lower face of the grid, where the spacer strips contact the guide tube surface. T he lowest spacer grid (Inconel) is not welded to the guide tubes due to material differenc es. It is supported by an Inconel 625 skirt which is welded to the spacer grid and to the perimeter of t he lower end fitting. For the assembly design with an Inconel top spacer grid, the grid is retained by ten Zircaloy-4 sleeves (five above and five below the grid) that are welded to the guide tubes at four locations per sleeve. (DRN 02-1538, R12;04-502, R13)
(LBDCR 15-025, R309)
The upper end fitting is an assembly consisting of two ca st stainless steel plates, five machined posts and five helical Inconel X-750 springs, which attaches to the guide tubes to serve as an alignment and locating device for each fuel assembly and has features to permit lifting of the fuel assembly. The lower cast plate locates the top ends of the guide tubes and is designed to prevent excessive axial motion of the fuel rods. (LBDCR 15-025, R309)
The Inconel X-750 springs are of conventional coil design having a coil diameter of 1.844 in., a wire diameter of 0.299 in., and approximately 14 active coils. Inconel X-750 was selected for this application because of its previous use for coil springs and good resistance to relaxation during operation.
(DRN 02-1538, R12)
The upper cast plate of the assembly, called the hol d-down plate, together with the helical compression springs, comprise the hold-down device. The hold-dow n plate is movable, acts on the underside of the fuel alignment plate, and is loaded by the compressi on springs. Since the springs are located at the upper end of the assembly, the spring load combines with the fuel assembly weight to counteract upward hydraulic forces. The determination of upward hydr aulic forces includes factors accounting for flow maldistribution, fuel assembly component toleranc es, crud buildup, drag coefficient, and bypass flow.
The springs are sized and the spring preload selected such that a net downward force will be maintained for all normal and anticipated transient flow and temper ature conditions. The design criteria limit the maximum stress under the most adverse tolerance c onditions to below yield strength of the spring material. The maximum stress occurs during cold conditions and decreases as the reactor heats up.
The reduction in stress is due to a decrease in spri ng deflection resulting from differential thermal expansion between the Zircaloy fuel bundles and the stainless steel internals. (DRN 02-1538, R12)
During normal operation, a spring will never be compress ed to its solid height. However, if the fuel assembly were loaded in an abnormal manner such that a spring were compressed to its solid height, the
spring would continue to serve its function when the loading condition returned to normal.
The lower end fitting is a single piece stainless steel ca sting consisting of a plate with flow holes and four support legs which also serve as a lignment posts. Precision drilled holes in the support legs mate with four core support plate alignment pins, thereby pr operly locating the lower end of the fuel assembly.
WSES-FSAR-UNIT-3 4.2-35 Revision 309 (06/16)
(DRN 02-1538, R12)
The four outer guide tubes have a widened region at t he upper end which contains an internal thread.
Connection with the upper end fitting is made by passing the male threaded end of the guide posts through holes in the lower cast flow plate and into the guide tubes. When assembled, the flow plate is secured between flanges on the guide tubes and on t he guide posts. The connection with the upper end fitting is locked with a mechanical crimp. Each outer guide tube has, at its lower end, a welded Zircaloy-4 fitting. This fitting has a female threaded portion which accepts a stainless steel bolt, which passes through a hole in the lower end fitting, to secure it. This joint is secured with a stainless steel locking ring
tack welded to the lower end fitting in four places. (DRN 02-1538, R12)
The central guide tube inserts into a socket in the upper end fittings and is thus retained laterally by the relatively small clearance. The upper end fitting socke t is created by the center guide tube post which is threaded into the lower cast flow plate and tack welded in two places. (EC-9533, R302)
The NGF design incorporates many of the same feat ures and geometry as the st andard fuel assembly, but incorporates a full complement of innovative components to improve fuel reliability, fuel cycle economics, fuel duty, manufacturability, burnup capability, and ther mal performance. The major differences between the two designs are the following: The NGF assembly uses bulged joints to build the grid cage versus welded joints and uses a pull rod loading process versus the current push loading process. These process changes were selected for NGF to improve the fabricabilit y of the design while preserving the rigidity of the fuel assembly structure. The guide thimbles are made of SRA Zircaloy-4 in the prior designs and SRA ZIRLO TM in the NGF design. This change was made because of ZIRLO TM's improved corrosion resistance and dimensional stability under irradiation. The NGF guide tube flange, which includes an anti-rotation feature to prevent the transmission of torque to the grids during post installation/re moval, is connected to the guide tube by bulging instead of by welding as in the standard des ign. The bulged flange to guide tube connection retains adequate strength and is necessary to co mpensate for the axial shrinkage of the guide tubes due to bulging. The NGF top grid is made of Inconel-718 and has vertical springs and horizontal dimples.
Stainless steel sleeves are brazed into the gr id at guide tube locations and are bulged with the guide tubes during cage fabrication to secure the grid to the guide tubes. The design is comparable to others that have an extensive history of succe ssful operation in Westinghouse NSSS nuclear power plants. (LBDCR 15-035, R309) The standard design Mid grids (HID-1L, Figure 4.2-5) are made using wavy strap Zircaloy-4, while the NGF Mid grids use straight strap Optimized ZIRLO TM. The material change was made because of Optimized ZIRLO TM's improved corrosion resistance and dimensional stability under irradiation. The straight str aps allow the incorporation of t he "I-spring" design and mixing vanes for improved fretting and thermal performance, res pectively. Sleeves fabricated from Optimized ZIRLO TM are laser-welded into the guide tube openings and secured to the guide tubes by bulges both above and below the grid. (LBDCR 15-035, R309) Two IFM grids are included to improve thermal per formance in two critical grid spans near the top for active core. These grids are short, non-stru ctural grids that are made from straight strap Optimized ZIRLO TM with side-supported mixing vanes and opposing dimples with small grid-to-rod gaps in lieu of an active (preloaded spring-dimple) support system. The IFM grids have sleeves that are similar to the Mid grid sleeves , except the protrusion of the sleeve above the IFM grids is less than above the Mid grids becaus e the IFM sleeves are only bulged to the guide tubes below the grid. (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-36 Revision 307 (07/13)
(EC-9533, R302; EC-30663, R307)
The lower portion of the NGF assembly includes several changes to accommodate rod push loading. In lieu of welding, the NGF Guardian TM grid is retained by inserts that are laser-welded to the four outer guide tube openings and then clamped between the bottom of the guide tube and the lower end fitting. To facilitate the installation of the lower end fitting after the rods have
been pulled into the grid cage, a small gap remain s between the bottoms of the NGF fuel rods and the bottom nozzle. This gap, in combinati on with associated changes to the lower end cap design, result in the bottom of the active f uel column being 0.165 inches higher than the prior design. The head of the NGF bolt has a skirted regi on that is crimped into recesses in the lower end fitting to secure the bolt, rather than using a separate locking disc that is welded to the lower end fitting to secure the bolt. The NGF bolt also includes a hole through the center of the
bolt to allow water to drain out of the guide tubes after washing the fuel assemblies during fabrication, or prior to the installation of the fuel assemblies in dry casks for spent fuel storage. (EC-30663, R307)
The NGF fuel rod design includes several changes relative to the standard fuel rod design, the most significant of which are the reduced diam eter/thickness of the cladding, a modified pellet geometry, the use of Optimized ZIRLO TM cladding, and an increase in the overall rod length.
These changes, as well as the other design changes associated with the NGF fuel rods, are detailed in Section 4.2.2.2. (EC-9533, R302)
The five guide tubes have the effect of ensuring that bowing or excessive swelling of the adjacent fuel rods cannot result in obstruction of the cont rol element pathway. This is so because:
a) There is sufficient clearance between the fuel rods and the guide tube surface to allow an adjacent fuel rod to reach rupture strain without contacting t he guide tube surface.
b) The guide tube, having considerably greater di ameter and wall thickness (and also being at a lower temperature) than the fuel rod, is consider ably stiffer than the fuel rod and would, therefore, remain straight, rather than be deflected by cont act with the surface of an adjacent fuel rod.
Therefore, the bowing or swelling of fuel rods woul d not result in obstruction of the control element channels such as could hinder CEA movement.
The fuel assembly design enables reconstitution, i.
e., removal and replacement of fuel and poison rods, of an irradiated fuel assembly. The fuel and pois on rod lower end caps are conically shaped to ensure proper insertion within the fuel assembly grid cage structure; the upper end caps are designed to enable grappling of the fuel and poison rod for purposes of removal and handling. Threaded joints which mechanically attach the upper end fitting to the c ontrol element guide tubes will be properly torqued and locked during service, but may be removed to provide access to the fuel and poison rods.
Loading and movement of the fuel assemblies is c onducted in accordance with strictly monitored administrative procedures and, at the completion of fuel loading, an independent check as to the location and orientation of each fuel assemb ly in the core is required.
(DRN 00-644; 02-1538, R12)
Markings provided on the fuel assembly upper end fitting enable verification of fuel enrichment and orientation of the fuel assembly. Identical ma rkings are provided on the lower end fitting to ensure preservation of fuel assembly identity in the event of upper end fitting removal. Additional markings are provided on each fuel rod during the manufacturing pr ocess to distinguish between fuel enrichments and burnable poison rods, if present. (DRN 00-644; 02-1538, R12)
WSES-FSAR-UNIT-3 4.2-37 Revision 302 (12/08)
(DRN 00-644; 02-1538, R12)
During the manufacturing process, each fuel rod is mark ed in order to facilitate a means of maintaining a record of pellet enrichment, pellet lot and fuel sta ck weight. In addition, a quality control program specification requires that measur es be established for the identific ation and control of materials, components, and partially fabricated subassemblies.
These means provide assurance that only acceptable items are used and also provide a method of relating an item or assembly from initial receipt through fabrication, installation , repair, or modification to an applicable drawing, specification, or other pertinent technical document. (DRN 00-644; 02-1538, R12) 4.2.2.2 Fuel Rod (DRN 02-1538, R12; 06-1059, R15)
The fuel rods consist of slightly-enriched UO 2 cylindrical ceramic pellets, a round wire Type 302 stainless steel compression spring, and an alumina spacer di sc located at each end of the fuel column, all encapsulated within a Zircaloy-4 tube seal welded with Zircaloy-4 end caps. The upper alumina disc was removed in the Batch S rod assemblies, and both spacers were removed from the Batch U and
subsequent reload fuel. Beginning with Batch U, a T ungsten Inert Gas (TIG) welding is utilized, using a friction fit of the cladding on a reduced diameter pedestal section of the end cap. Beginning with Batch Y, the ZIRLO TM cladding tubes are used and are TIG welded with the Zircaloy-4 end caps. The fuel rods are internally pressurized with helium during assemb ly. Figure 4.2-10 depicts the fuel rod design. (DRN 02-1538, R12; 06-1059, R15)
Each fuel rod assembly includes a unique serial num ber. The serial number ensures traceability of the fabrication history of each fuel rod component. Fini shed fuel rods, prior to being loaded into bundles, are processed through a rod scanner to check pellet enrichment. (EC-9533, R302)
The fuel cladding is cold-worked and stress relief annealed Zircaloy-4 tubing 0.025 in. thick. The actual tube forming process consists of a series of cold working and annealing operations, the details of which are selected to provide the combination of properties discussed in Subsection 4.2.1.2.2. (EC-9533, R302)
The UO 2 pellets are dished at both ends in order to better accommodate thermal expansion and fuel swelling. The initial density of the UO 2 pellets is 10.44 g/cm 3 , which corresponds to 95.25 percent of the 10.96 g/cm 3 theoretical density (TD) of UO
- 2. However, because the pellet di shes and chamfers constitute about three percent of the volume of the pellet stack, the average density of the pellet stack is reduced to 10.11 g/cm
- 3. This number is referred to as the "stack density."
(DRN 06-1059, R15)
Note that the initial pellet density and stack density for Erbia pellets used in Erbia fuel rods (see Section 4.2.2.3) are slightly lower (respectively 10.41 and 10.09 g/cm
- 3) due to Erbia content. These densities for the ZrB 2 coated UO 2 (IFBA) pellets that were first introduced in Batch Y for Cycle 15 are consistent with those for the solid UO 2 pellets. However, the pellet stack dens ity for the annular pellets used in the cutback zones of the pellet stack is lowered to 7.80 g/cm 3 (due to hollow center of 0.1625" diameter). (DRN 06-1059, R15)
(DRN 02-1538, R12)
The compression spring located at the top of the fuel pellet column maintains the column in its proper position during handling and shipping. The fuel rod plenum, which is located above the pellet column, provides space for axial thermal differential expansi on of the fuel column and accommodates the initial helium loading and evolved fission gases. (See Subsecti on 4.2.1.2.5.1 and 4.2.1.2.
5.2). The specific manner in which these factors are taken into accoun t, including the calculation of temperatures for the gas contained within the various types of rod inte rnal void volume, is discussed in Reference 14. (EC-9533, R302)
Starting with Batch U, fuel rod fabrication was moved from Hematite, MO, to the Columbia, SC, facility.
Figure 4.2-10A compares the Hematite and Columbia production urania rod assembly features. Figure 4.2-11A compares the corresponding erbia rod assemblies.
(DRN 02-1538, R12; EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-38 Revision 309 (06/16)
(EC-9533, R302)
The basic configuration of the NGF fuel rod (Figure 4.2-10B) and IFBA rod (Figure 4.2-11B) are comparable to the prior rod designs, but there are si gnificant differences in the detailed design of the rods. The NGF rods have a smaller outside diameter than prior designs (0.374" versus 0.382") to compensate for some of the pressure drop incr ease associated with the NGF spacer grids.
The 0.374" diameter rod is the same as the standard Westinghouse 17x17 design, which precipitated the use of the 17x17 cladding di mensions and pellet geometry for the NGF design. Therefore, the cladding outside/insi de diameters are 0.374" and 0.329", while the fuel and IFBA pellets have a diameter of 0.3225", a length of 0.387", and a spherical dish at each end instead of a truncated dish. The blanket pellet associated with the IFBA rod has a
diameter of 0.3225", a length of 0.500", and a central hole of 0.155". The corresponding stack densities for these pellet configurations are 10.31 g/cc for the fuel and IFBA pellets, and 8.00 g/cc for the blanket pellets. (LBDCR 15-035, R309) Optimized ZIRLO TM fuel cladding has been used to replace the ZIRLO TM fuel cladding. The topical report, Reference 83, summarizes the materi al properties as they pertain to fuel rod cladding, design and licensing activities.
The difference between Optimized ZIRLO TM fuel cladding and ZIRLO TM cladding is that Optimized ZIRLO TM has a slight reduction in Tin content for improved corrosion resistanc e (0.6% minimum for Optimized ZIRLO TM versus 0.8% minimum for ZIRLO TM). Reference 85 updates the cladding corrosion model. (LBDCR 15-035, R309) The overall length of the NGF fuel rod is in creased by 0.7" to minimize the loss of void volume associated with the diameter reduction of t he rod. To further offset the effect of the diameter reduction, the initial fill gas pre ssure of the fuel rods has been reduced to approximately 275 psig. (EC-13881, R304)
The nominal active length remains 150" for bot h the fuel and IFBA rods. The fuel rod stack continues to exclude any cutback/blanket pelle ts, while the IFBA rod stack has a cutback/
blanket zone at each end of the center pellet column (See Fig. 4.3A-19b). (EC-13881, R304)
The bottom end cap has been modified to accommodate a recess in the bottom end that is necessary for pull-loading the rods into the f uel assemblies. The length of the upper end cap has been reduced and the "acorn" removed to allow as large an increase as possible to the
plenum to facilitate the accomm odation of fission gas release.
4.2.2.3 Burnable Poison Rod (DRN 02-1477, R12)
Fixed burnable neutron absorber (poi son) rods, Figure 4.2-11, will be included in selected fuel assemblies to reduce the beginning-of-life moderator coe fficient. They will replace fuel rods at selected locations. The poison rods will be mechanically similar to fuel rods. The poison material will be alumina with uniformly-dispersed boron carbide particles. The bal ance of the column will c onsist of two Zircaloy-4 spacers with the total column length the same as t he column length in fuel rods. The burnable poison rod plenum spring is designed to produce a smaller preload on the pellet column than that in a fuel rod
because of the lighter material in the poison pellets. (DRN 02-1477, R12; EC-9533, R302)
Each burnable poison rod assembly includes a unique serial number. The serial number is used to record fabrication information for each component in the rod assembly.
(DRN 02-1477, R12; 06-1059, R15; EC-9533, R302)
(DRN 02-1477, R12; 06-1059, R15; EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-39 Revision 302 (12/08) 4.2.2.4 Control Element Asse mbly Description and Design Drawings (DRN 01-1103, R12; 02-1477, R12)
The Waterford 3 reactor contains a total of 87 CEAs. These are distributed among the fuel assemblies as shown in Figure 4.2-12. The CEA is shown in Figure 4.2-5. CEAs have four control elements arranged in a 4.050-in. square array plus one element at the cent er of the array. Each CEA interfaces with the guide tubes of one fuel assembly. (DRN 01-1103, R12; 02-1477, R12)
(DRN 02-1477, R12)
(DRN 02-1477, R12)
(DRN 02-1477, R12)
The control elements of a CEA consist of an Inconel 625 tube loaded with a stack of cylindrical absorber pellets. The absorber material consis ts of 73 percent TD boron carbide (B 4 C) pellets, with the exception of the lower portion of the element s, which contain silver-indium-cadm ium (Ag-In-Cd) alloy cylinders. (DRN 02-1477, R12)
Two design objectives are realized by the us e of Ag-In-Cd in the element tip zones:
a) CEA Cladding Dimensional Stability
Because of its high ductility and low strength, t he Ag-In-Cd will not deform the CEA cladding. Buffering of the CEA following scram, which occurs when the corner element tips enter a reduced diameter portion of the fuel assembly guide tubes, is not degraded with long term exposure of the
CEA to reactor operating conditions.
b) Adequate CEA Worth
Although some reduction in CEA worth arises because of the substitution of B 4 C with Ag-In-Cd, the effect is small and is accounted for.
During normal powered operation, most of the CEAs ar e expected to be in the fully withdrawn position.
Above the poison column is a plenum which provi des expansion volume for helium released from the B 4 C. The plenum volume contains a Type 302 stainl ess steel hold-down spring, which restrains the absorber material against longitudinal shifting with respect to the clad while allowing for differential expansion between the absorber and the clad. The sp ring develops a load sufficient to maintain the position of the absorber material during shipping and handling. (DRN 02-1477, R12)
Each control element is sealed by welds which join the tube to an Inconel 625 nose cap at the bottom, and an Inconel 625 end fitting at the top. The end fittings, in turn, are threaded and pinned to the spider structure which provides rigid lateral and axial suppor t for the control elements. The spider hub bore is specially machined to provide a point of attachment for the CEA extension shaft. (DRN 02-1477, R12)
(DRN 01-1103, R12)
(DRN 01-1103, R12)
(DRN 00-644; 01-1103, R 12; 02-1477, R12) Each CEA is positioned by a magnetic jack control element drive mechanism (CEDM) mounted on the reactor vessel closure head. The extension shaft join s with the CEA spider and connects the CEA to the CEDM. CEAs may be connected to any extension shaft depending on control requirements. Mechanical reactivity control is achieved by positioning groups of CEAs by the CEDMs. (DRN 00-644; 02-1477, R12)
In the outlet plenum region, all CEAs are enclos ed in CEA shrouds which provide guidance and protect the CEA and extension shaft from coolant cross flow.
Within the core, each element travels in a Zircaloy guide tube. The guide tubes are part of the fuel asse mbly structure and ensure pr oper orientation of the control elements with respect to the fuel rods.
(DRN 01-1103, R12)
WSES-FSAR-UNIT-3 4.2-40 Revision 302 (12/08)
When the extension shaft is released by the CEDM, the combined weight of the shaft and CEA causes the CEA to insert into the fuel assembly. (DRN 01-1103, R12; 02-1477, R12)
The lower ends of the four outer fuel assembly guide tubes are tapered gradually to form a region of reduced diameter which, in conjunction with the outer control element on the CEA, constitutes an effective hydraulic buffer for reducing the decelerati on loads at the end of a trip stroke. This purely hydraulic damping action is augmented by a spring and plunger arrangement on the CEA spider. When fully inserted, CEAs rest on the central pos t of the fuel assembly upper end fitting. (DRN 01-1103, R12; 02-1477, R12)
The capability of the CEAs to scram within the allowable time is demonstrated as part of the flow testing discussed in Subsection 4.2.4.4.
4.2.3 DESIGN EVALUATION
4.2.3.1 Fuel Assembly
4.2.3.1.1. Vibration Analyses
Three sources of periodic excitation are recognized in evaluating the fuel assembly susceptibility to vibration damage. These sources are as follows:
a) Reactor Coolant Pump Blade Passing Frequency
Precritical vibration monitoring on previous C-E reactors indicates the peak pressure pulses are expected at the pump blade passing frequency, and a lesser but still pronounced peak at twice this frequency.
b) Core Support Plate Motion
Experience with earlier C-E reactors indicates that random lateral motion of the core support plate is expected to occur with an amplit ude of 0.001 to 0.002-in. and a frequency range of between 2 and 10 Hz.
c) Flow-induced vibration resulting from coolant flow through the fuel assembly.
The capability of the Waterford 3 16 x 16 fuel a ssembly to sustain the effects of flow-induced vibration without adverse effects has been demons trated in a dynamic flow test performed in CE's TF-2 flow test facility. The test utiliz ed prototypical 16 X 16 reactor components consisting of a 16 X 16 type fuel assembly, a CEA shr oud, control element drive mechanism, and a simulation of surrounding core internal support components and was performed under extreme
flow and temperature conditions. The success of this test, similar previous tests of 16 X 16 fuel assemblies and the operation of CE's ANO-2 plan t, demonstrate that flow-induced vibration will have no adverse effects on the Waterford 3 fuel assemblies. (EC-9533, R302)
The NGF fuel assembly design was designed to have a lateral stiffness comparable to the prior
designs that have operated successfully in t he Waterford plant. In addition, the NGF configuration was tested to confirm the hydraulic stability of the fuel assembly design and to demonstrate the acceptability of the fretting perfo rmance of the fuel assembly design. The testing included full scale single and dual bundle tests with both the NGF design and the
standard design. The single bundle tests demonstr ated the hydraulic stability of both designs over the expected range of flow rates. The dual bundle test was an endurance test that provided additional confirmation of the hydraulic stab ility of the designs and showed a significant improvement in the fretting performance of t he NGF design compared to the standard design.
These results indicate that the NGF design is ev en less susceptible to any vibration effects than the prior designs. (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-41 Revision 302 (12/08)
These sources of periodic motion are not expected to have an adverse effect on the performance of the Waterford 3 fuel assembly.
4.2.3.1.2 CEA Guide Tube
The CEA guide tubes were evaluated for structural adequacy using the criteria given in Subsection 4.2.1.1 in the following areas:
a) Steady axial load due to the combined effect s of axial hydraulic forces and upper end fitting holddown forces.
For normal operating conditions, the resultant guide t ube stress levels satisfy the criteria given in Subsection 4.2.1.1.1.
b) Short-term axial load due to the impact of the spring loaded CEA spider against the top of the fuel assembly at the end of a CEA trip.
For trips occurring during normal power operation, solid impact is not predicted to occur due to the kinetic energy of the CEA being dissipated in the hydraulic buffer and by the CEA spring.
c) Short-term differential pressure load occurring in the hydraulic buffer regions of the outer guide tubes at the end of each trip stroke.
The buffer region slows the CEA during the last fe w inches of the trip stroke. The resultant differential pressure across the guide tube in this region is predicted to be 300 psi, and this gives rise to circumferential stresses of 3300 psi, which is less than one quarter of the yield stress, for a
very short term. The trip is assumed to be r epeated daily. However the resultant stress is too small to have a significant effect on fatigue usage.
For conditions other than normal operation, the additional mechanical loads imposed on the fuel assembly by an OBE (equivalent to one-half SSE), SSE, and large break LOCA and their resultant effect on the control element guide tubes are discussed in the following paragraphs:
4.2.3.1.2.1 Operating Basis Earthquake
During the postulated OBE, the fuel assembly is subj ected to lateral and axial accelerations which, in turn, cause the fuel assembly to deflect from its normal shape. The method of calculating these deflections is described in Subsection 3.7.3.14. The magnitude of the lateral deflections and resultant
stresses are evaluated for acceptabilit
- y. The method for calculating stresses from deflected shapes is described in Reference 50. The results of the st ress analysis demonstrate t hat the equipment stresses are less than the allowable values discussed in Subsection 4.2.1.1.
4.2.3.1.2.2 Safe Shutdown Earthquake
The axial and lateral loads and deformation sustained by the fuel assembly during a postulated SSE have the same origin as those discussed above for the OBE, but they arise from initial ground accelerations twice those used for the OBE. The analytical met hods used for the SSE are identical to those used for the OBE. The predicted component stresses were less than the allowable values discussed in Reference
- 50.
4.2.3.1.2.3 Loss-of-Coolant Accident (DRN 03-2058, R14)
In the event of a large break LOCA, there will occur rapid changes in pressure and flow within the reactor vessel. Associated with the transient are relatively large axial and lateral loads on the fuel assemblies.
(DRN 03-2058, R14)
WSES-FSAR-UNIT-3 4.2-42 Revision 302 (12/08)
(DRN 03-2058, R14)
The response of a fuel assembly to the mechanica l loads produced by a LOCA is considered acceptable if the fuel rods are maintained in a coolable array, i.e., acceptably low grid crushing. The methods used for analysis of combined seismic and LOCA loads and stresses is described in Reference 50. See Sections 3.6.2.1.1.1(d) and 3.
6.3 for discussions on pipe break criteria and leak-before-break. (DRN 00-644) (DRN 03-2058, R14)
To qualify the complete fuel assembly, full-scale hot loop testing has been conducted. The tests were designed to evaluate fretting and wear of components, refueling procedures, fuel assembly uplift forces, holddown performance and compatibility of the fuel assembly with interfacing reactor internals, CEAs and
CEDMs under conditions of reactor water chemistry, fl ow velocity, temperature, and pressure. Additional information on the test is given in Subsection 4.2.
3.2.4.2. The test was run for approximately 2000 hours0.0231 days <br />0.556 hours <br />0.00331 weeks <br />7.61e-4 months <br /> and was completed in 1976. (DRN 00-644)
Mechanical testing of the fuel assembly and it s components has been performed to support analytical means of defining the assembly's structural characteristics. The test program consisted of static and dynamic tests of spacer grids and static and vi bratory tests of a full size fuel assembly.
4.2.3.1.2.4 Combined SSE and LOCA
It is not considered appropriate to combine the st resses resulting from the SSE and LOCA events.
Nevertheless for purposes of demonstrating margin in the design, the maximum stress intensities for each individual event were combined by a square root of the sum of the squares (SRSS) method. This was performed as a function of fuel assembly elevat ion and position, e.g., the maximum stress intensities for the center guide tube at the upper grid elev ation (as determined in the analysis discussed in Subsections 4.2.3.1.2.2 and 4.2.3.1.2.3) were combined by the SRSS method. The results demonstrated that the allowable stresses described in Reference 50 were not exceeded for any position along the fuel assembly, even under the added conservatism provided by this load combination.
4.2.3.1.3 Spacer Grid Evaluation
As discussed in Subsection 4.2.2.1 the function of the spacer grids is to provide lateral support to fuel and burnable poison rods in such a manner that the axial fo rces are not sufficient to buckle or bow the rods and that the wear resulting at the grid-to-clad contact points will be lim ited to acceptably small amounts.
It is also a criterion that the grid be capable of withstanding the lateral loads imposed during the postulated seismic and LOCA events.
(DRN 02-1538, R12)
With respect to the design criterion that the axial re straint offered by the grids during initial assembly be such that the axial forces on a fuel rod are not sufficient to cause the rod to bow or buckle, it is currently understood that the observed instances of fuel rod bowing have occurred because the axial restraint of
the spacer grids on the fuel rods was such that re lative motion between the fuel rods and the grids (e.g., differential thermal expansion) could not occur except at axial forces high enough to cause slight bowing of the fuel rods. Fuel assemblies, however, are designed such that the combination of fuel rod rigidity, grid spacing, and grid preload will not caus e significant fuel rod deformation under axial loads.
The long-term effects of clad creep (reduction in cl ad OD), the reduction of grid stiffness with temperature, and the partial relaxation of the grid ma terial during operation ensure that this criterion is also satisfied during all operating conditions. Moreov er, visual inspection of irradiated fuel assemblies from the Maine Yankee (14 x 14), Palisades (15 x 15) and Fort Calhoun (14 x 14) reactors has not shown any significant bowing of the fuel rods. In view of these factors and the similarity of these designs to the Waterford 3 design, it is concluded that the axial fo rces applied by the grids on the cladding will not result in a significant degree of fuel rod bow. Additional di scussion of the causes and effects of fuel rod bowing are contained in Subsection 4.2.3.2.6 and in References 53 and 75. (DRN 02-1538, R12)
WSES-FSAR-UNIT-3 4.2-43 Revision 302 (12/08)
(EC-9533, R302)
The capability of the grids to support the clad wit hout excessive clad wear has been demonstrated by out-of-pile flow testing, and by the re sults of post-irradiation examinati on of grid-to-clad contact points in Maine Yankee fuel assemblies which showed only negligible clad wear (51). An extensive flow test program was conducted to support the implementation of the NGF design. A full scale dual bundle test with a NGF fuel assembly and a standard fuel assemb ly was run to demonstrate the acceptability of the fretting performance of the NGF assembly design.
The dual bundle test was an endurance test that provided confirmation of the hydraulic stability of the designs and showed a significant improvement in the fretting performance of the NGF desi gn compared to the standard design. (EC-9533, R302)
The capability of the grid to withstand the lateral loads produced during the postulated seismic and LOCA events is demonstrated by impact testing of the refer ence grid design and comparing the test results with the analytical predictions of the seismic and LOCA l oads. The test methods are discussed in Reference
- 50. (DRN 00-644) (DRN 03-2058, R14)
For the original fuel design, the results of the load comparison were that under seismic loading no spacer grids in the core were subjected to loadings in exce ss of their capability based on test results. However, under LOCA conditions some fuel assemblies in the periphery of the core had spacer grids with predicted loads which exceeded the capability defined by testing.
An ECCS analysis was performed for the core locations occupied by these fuel assemblies, and the re sults confirmed that the ECCS acceptance criteria (10CFR50.46) were still satisfied. The methods used in the ECCS evaluation were the same as used in previous analyses (Reference 77). In order to demonstrate margin in the design, spacer grid loadings from the SSE and LOCA events were combined by a squar e root of the sum of squares (SRSS) method, and no additional fuel assemblies were found to have grids which exceeded the capability defined by
testing.
With the introduction of the HID-1L gird design, the grid strengths increased above those of the original fuel design such that the grid strengths exceeded the maximum grid impact loads. There was therefore no further need to perform the ECCS analysis to s how that the HID-1L grid design was acceptable.
For the power uprate condition, updated LOCA loadi ngs were determined which included a combination of power uprate and Leak-Before-Break effects.
Seismic loadings remained unchanged. Because the loadings used in the analyses performed for the orig inal fuel design are conservative and bounding with respect to the uprate loadings, there was no need to reevaluate the fuel assemblies. (DRN 03-2058, R14) (EC-9533, R302)
The NGF design utilizes straight-strip mid grids that have a higher spacer grid stiffness than the HID-1L spacer grids of the prior fuel designs. Due to this increased grid stiffness, the seismic and LOCA analyses were reevaluated for the mixed core and all NGF core cases. The results showed that the documented strengths of the NGF and HID-1L spacer grids exceeded predicted impact loads, but that the loading history simulated in the determination of the HID-1L grid strength did not bound the predicted loading history in some peripheral core locations. In stead of retesting the HID-1L grids with the predicted mixed core loading history, an E CCS evaluation was performed to demonstrate compliance with the ECCS acceptance criteria.
The Zircaloy-4 spacer grid material is of the sa me composition as the fuel rods and guide tubes with which it is in contact, thereby obv iating any problem of chemical inco mpatibility with those components.
For the same reason, adequate resistance to corrosi on from the coolant is assured (see Subsection 4.2.3.2.3, for additional information relative to the co rrosion resistance of Zircaloy
-4 in the primary coolant environment). Similarly, the NGF design is not susceptible to chemical incom patibility since it utilizes Optimized ZIRLO TM spacer grids and Optimized ZIRLO TM cladding. In addition, the use of Optimized ZIRLO TM for the spacer grids offers improved corrosion resi stance relative to the Zircaloy-4 spacer grids. (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-44 Revision 309 (06/16)
(DRN 02-1538, R12; 06-1059, R15, LBDCR 15-025, R309)
The Inconel-625 material used for the lowest, and in so me cases the uppermost, spac er grid is in contact with the coolant, the stainless steel lower end fitting (to which it is welded), the Zircaloy-4 or ZIRLO TM fuel and poison rods, and the Zircaloy-4 guide tubes. The mut ual chemical compatibility of these materials in a reactor environment has been demonstrated by the use of these materials in fuel assemblies that have been operated in other C-E reactors and for which pos t irradiation examination has yielded no evidence of chemical reaction between these components.
In addition, experiments have also been performed at C-E on Inconel type alloys and Zircaloy-4 which s howed the eutectic reactions did not occur below 2200 F, a temperature far in excess of that anticipated at the lower grid location in the event of a LOCA. (DRN 02-1538, R12; 06-1059, R15, LBDCR 15-025, R309)
(EC-9533, R302)
The Inconel-718 material used for the top spacer grid in the NGF design has similar material characteristics to the Inconel-625 material and has operated successfully in Westinghouse plants for many years with ZIRLO TM clad fuel rods. The slight reduction in tin content of the Optimized ZIRLO TM cladding compared to the ZIRLO TM cladding does not impact its compatibility with Inconel-718, as evidenced by the successful operati on of the NGF lead fuel assemblies in Waterford and in other Westinghouse reactors. (EC-9533, R302)
The only dissimilarity, between the fuel for which post-irradiation examination dat a are presently available and the Waterford 3 design (other than dimensional va riations), is that the Inconel-625 is used as a spacer grid for Waterford 3 and was used originally as a retention grid. However, the effect that such a change might have on fretting behavior has been evaluat ed in out-of-pile flow test programs (see Subsection 4.2.3.2.4.2). (EC-9533, R302) 4.2.3.1.4 Dimensional Stabilit y of Zirconium-Based Alloys (DRN 06-1059, R15)
Zircaloy components are designed to allow for dimens ional changes resulting from irradiation-induced growth. Extension analyses of in-pile growth data have been performed to formulate a comprehensive model of in-pile growth. The in-pile growth equations are used to determine the minimum axial differential growth allowance which must be included in the axia l gap between the fuel rods and the upper end fitting.
For determining the necessary fuel rod growth allow ance, the growth correlations for fuel rod and guide tube growth are combined statistically such that the minimum initial gap is adequate to accommodate the upper 95 percent confidence level of differential gr owth between fuel rods and guide tubes in the peak burnup assembly for Zircaloy, ZIRLO TM , and Optimized ZIRLO TM cladded rods. For the purpose of predicting axial and lateral growth of the fuel assemb ly structure (thereby estab lishing the minimum initial clearance with interfacing components), the equations are used in a conserva tive manner to ensure adequate margins to interference are maintained. (DRN 06-1059, R15; EC-9533, R302)
(DRN 02-1538, R12)
Inspection of fuel assemblies after two cycles of oper ation at the Arkansas Nuclear One, Unit 2 reactor has shown higher rates of gap closure than predicted by the method described in Reference (3). Closure rates predicted by Reference (3) may remain valid for the Waterford 3 fuel assemblies because of
differences in the Waterford and Arkansas des igns. Nonetheless, additional shoulder gap has been provided in those fuel assemblies sc heduled for three cycles of operation. (DRN 02-1538, R12)
The additional gap was selected to provide the maxi mum shoulder gap without violating other design criteria. Based on the shoulder gap reduction observed at ANO-2 at EOC2, the additional shoulder gap is expected to provide three cycle operation capability. (EC-9533, R302)
The NGF design incorporates material changes that improve the dimensional stability of the CEA guide tubes and the fuel rod cladding. These improvement s allow a reduction in the NGF shoulder gap while still providing adequate space to accomm odate rod burnups above 60,000 MWd/MTU. (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-45 Revision 302 (12/08) 4.2.3.1.5 Fuel Handling and Shipping Design Loads
Three specific design bases hav e been established for shipping and handling loads. These are as follows:
a) The fuel assembly, when supported in the new fuel shipping container, shall be capable of sustaining the effects of five g axial, lateral or vertical acceleration without sustaining stress levels in excess of those allowed for normal operation. The five g criterion was originally established
experimentally, and its adequacy is continually confir med by the presence of impact recorders as described in the following paragraph.
Impact recorders are included with each shipment wh ich indicate if loadings in excess of five g are sustained. A record of shipping loads in excess of five g indicates an unusual shipping occurrence in which case the fuel assembly is inspected for damage prior to releasing it for use.
The axial shipping load path is through either end fitting to the guide tubes. A five g axial load produces a compressive stress level in the guide tubes less than the two-thirds yield stress limit
that is allowed for normal condition events. T he fuel assembly is prevented from buckling by being clamped at grid locations. For lateral or ve rtical shipping loads, the grid spring tabs have an initial preload which exceeds five times the fuel rod weight. Therefore, the spring tabs see no
additional deflection as a result of five g lateral or vertical acceleration of the shipping container.
In addition, the side load on the grid faces produced by a five g lateral or vertical acceleration is less than the measured impact strength of the grids.
b) The fuel assembly shall be capable of sustaining a 5000 pound axial load applied at the upper end fitting by the refueling grapple (and resisted by an equal load at the lower end fitting) without sustaining stress levels in excess of those allowed for normal operation. The 5000 pound load was chosen in order to provide adequate lift capab ility should an assembly become lodged. This load criterion is greater than any lift load that has been encountered in-service.
c) The fuel assembly shall be capable of with standing a 0.125 in. deflection in any direction whenever the fuel assembly is raised or lower ed from a horizontal position without sustaining a permanent deformation beyond the fuel assembly inspection envelope.
Fuel handling procedures require the use of a str ongback to limit the fuel assembly deflection to a maximum of 0.125 in. in any direction whenever t he fuel assembly is raised or lowered to a horizontal position. This limits the stress and st rain imposed upon the fuel assembly to values well below the limits set for normal operating c onditions. The adequacy of the 0.125 in. criterion is based on the inclusion of this limitation in specifications and procedures for fuel handling equipment, which is thereby constrained to provi de support that lateral deflection is limited to 0.125 in.
4.2.3.2 Fuel Rod Design Evaluation
The evaluations discussed in this section are based on assumed fuel rod operation within certain linear heat rate limits related to avoiding excessive fuel clad temperatures. Information concerning the bases for these limits is contained in Section 4.4.
4.2.3.2.1 Results of Vibration Analyses
Three sources of periodic excitation are recognized in evaluating the fuel rod susceptibility to vibration damage. These sources are as de scribed in Subsection 4.2.3.1.1.
These sources of periodic motion are not expected to have an adverse effect on the performance of the fuel rod. Subsection 4.2.3.2.4 includes additional informaton on fuel rod response to the sources.
WSES-FSAR-UNIT-3 4.2-46 Revision 309 (06/16) 4.2.3.2.2 Fuel Rod Internal Pressure and Stress Analysis (DRN 02-1538, R12)
A fuel rod cladding stress analysis is conducted to determine the circumferential stress and strain resulting from normal, upset, and emergency conditi ons. The analysis includes the calculation of cladding temperatures and rod internal pressures dur ing each of the occurrences listed in Subsection 4.2.1.1. The design criteria to be used to evaluat e the analytical results are specified in Subsection 4.2.1.2.1. Fuel rod stresses re sulting from seismic events are calculated, using the methodology described in Reference 50. (DRN 02-1538, R12) 4.2.3.2.3 Potential for Chemical Reaction
a) Corrosion
Corrosion tests of Zircaloy-4 fuel rod t ubing which were conducted in excess of 4000 hours0.0463 days <br />1.111 hours <br />0.00661 weeks <br />0.00152 months <br /> exposure include 600 and 650 F autoclave tests and 600 F loop tests with borated lithium hydroxide additives to the water chemistry. The test results agree with long term corrosion tests in lithium hydroxide reported by Bettis.
(52) No deleterious effects have occurred. (DRN 02-1538, R12)
Experience at both Shippingport and Saxton Core I have shown under PWR conditions (hydrogen overpressure and chemical additives) t hat in reactor behavior with low heat flux was similar to autoclave behavior. Experience at the Saxton reactor in Cores II and III, however, have shown that with severe nucleate boiling, some accelerated corrosion was encountered. Similar accelerated corrosion with high crud deposits was al so reported at KWO, but was terminated by using hydrogen overpressure and chemical additives. (DRN 00-644; 06-1141, R15; EC-9533, R302, LBDCR 15-035, R309)
Batch Y fuel rods were fabricated with ZIRLO TM cladding to improve the corrosion resistance of the fuel. Section 4.5 of Reference 80 present s corrosion data at high burnup for both Zircaloy-4 cladding and ZIRLO TM cladding and concludes that the ZIRLO TM cladding offers a significant improvement in the corrosion resi stance of the cladding. NGF f uel rods are fabricated with Optimized ZIRLO TM cladding that has a slightly reduc ed tin content compared to ZIRLO TM specifically to improve its corrosion resistance.
Autoclave steam testing demonstrated almost a 20% corrosion resistance improv ement of Optimized ZIRLO TM compared to ZIRLO TM. Reference 85 presents additional cladding corrosion data and provides updated cladding corrosion models for both ZIRLOTM and Optimized ZIRLOTM cladding. (EC-9533, R302, LBDCR 15-035, R309)
Coolant chemistry parameters have been specified that minimize corrosion product release rates and their mobility in the primary system. Specif ically, the precore hot functional environment is controlled (ph and oxygen) to provide a thin, t enacious, adherent, protective oxide film. This approach minimizes corrosion product release and associated inventory on initial startup and subsequent operation. During operation, the spec ified lithium concentration range (0.2-3.5 ppm) effects a chemical potential gradient or drivi ng force between hot and cooler surfaces (refuel cladding and steam generator tubing, respectively) such that soluble iron and nickel species will preferentially deposit on the steam generator su rfaces. The associated ph also minimizes general corrosion product release rates from primar y system surfaces. Moreover, the specified hydrogen concentration range (10-50 cm 3 kg STP) ensures; reducing conditions in the core thereby avoiding low solubility Fe 3+. Additionally, dissolved hydrogen promotes rapid recombination of oxidizing species. Oxidizi ng species and a fast neutron flux are synergistic prerequisites to accelerated Ziraloy-4 corrosion. (DRN 00-644; 02-1538, R 12; 06-1141, R15) During operations lithium, dissolved oxygen, and dissolved hydrogen will be monitored at a frequency consistent with maintaining these parameters within their specifications. (EC-9533, R302)
Post-operational examinations of fuel cladding t hat has operated within t hese specifications, has shown no significant chemical or corrosive attack of the Zircaloy cladding. ZIRLO TM cladding and Optimized ZIRLO TM cladding are less sensitive to chemical or corrosive attack due to their better corrosion resistance. (EC-9533, R302)
WSES-FSAR-UNIT-3 4.2-47 Revision 309 (06/16) b) External Hydriding (EC-9533, R302)
During operation of the reactor with exposure to high temperature, high pressure water, Zirconium-based cladding will react to form a prot ective oxide film in accordance with the following equation. (EC-9533, R302)
(DRN 00-644) Zr + 2H 2 O = ZrO 2 + 2H 2 (DRN 00-644) Approximately 20 percent of the hydrogen is ads orbed by the Zircaloy. Based on data described in WAPD-MRP-107, the cladding would be expec ted to contain up to 250 ppm of hydrogen following three years of exposure.
A series of 600 F burst tests was performed on Zircaloy-4 tubes containing to 200 to 250 and 400 ppm of hydrogen precipitated as hydride platelet s in various orientations from radial to circumferential. Additional burst te sts have shown similar effects at 725 F. Little difference in burst test ductility was evident. Therefore, hydrogen normally adsorbed in Zr-4 tubing will not prove deleterious to the cladding integrity. (EC-9533, R302, LBDCR 15-035, R309)
The impact of hydrides and hydride reorientation in ZIRLO TM cladding is discussed in Section 4.4.2.5 of Reference 80, where it is c oncluded that the performance of the ZIRLO TM cladding will be similar to that of the Zircaloy-4 cladding since the hydride reorientation is primarily a function of the tensile stresses and temperatures in the cl adding. Due to the similarity of the material composition of Optimized ZIRLO TM and ZIRLO TM , the same conclusion applies for Optimized ZIRLO TM cladding. Reference 85 provides updated data and models for ZIRLOTM and Optimized ZIRLOTM cladding. (EC-9533, R302, LBDCR 15-035, R309) c) Internal Hydriding
A number of reported fuel rod failures have resulted from excessive moisture available in the fuel.
Under operation, this moisture would fl ash to steam and oxidize the Zircaloy.
The hydrogen, which was not absorbed during norma l oxidation, would then be absorbed into the Zircaloy through a scratch in the oxide film.
This localized hydrogen absorption by the cladding would shortly result in a localized fuel rod failu re. Work performed at the Institt for Atomenergi, Halden, Norway, of which C-E is a member, demonstrated that a threshold value of water moisture is required for hydride sunbursts to occu
- r. Through a series of in-pile experiments, the level of this threshold value was established.
The allowable hydrogen limit in the fuel complies with this requirement, ensuring that hydride sunbursts will not occur.
d) Crud
The slow general corrosion of out-of-core plant su rfaces will release corrosion products to reactor coolant, some of which will deposit on core surfac es as "crud". The major constituents of crud are iron and nickel, with lesser amounts of chromium and traces of manganese and cobalt, all
present as oxides. Crud is e ssentially a nickel ferrite (Ni x FE 3-x O 4) with "x" in the range 0.45-0.75. Chromium appears to enter the inverse sp inel substantially to give a composition CR y Ni x Fe 3-x-y O 4 (Reference 69). The porosity of core crud deposits is typically given as 80 to 85 percent (density 1.2 g/cm 3.)
Although there are significant efforts underway within the industry to develop mathematical models for crud transport in reactor coolant syst ems (e.g., see Reference 70), at present there are not analytical techniques available fo r estimating crud buildup on fuel surfaces.
WSES-FSAR-UNIT-3 4.2-48 Revision 302 (12/08)
Although heavy crud deposits have been observed in older plants (see Reference 71), measurements made on modern pressurized water r eactors indicate that crud buildup is low, ranging from < 0.02 mils up to a few tenths of a mil (see Reference 72). As discussed in Subsection 4.2.3.2.3 a) above, coolant chemistry parameters have been specified to minimize crud deposition. Visual inspection of fuel remo ved from CE's Calvert Cliffs I plant, which operated under these specifications, revealed relative ly light crud deposits, su ch that clad surface features from fabrication could be discerned (s ee Reference 73). Similar behavior is anticipated for the Waterford 3 core.
Enhanced corrosion of Zircaloy cladding should not occur under light deposits of porous crud.
Water is free to flow through porosity in the cr ud, providing heat transfer through convection.
Even heavy crud found in Yankee Rowe (Refer ence 71) was non-insulating because of its porosity.
e) Fuel-Cladding Chemical Reaction
An in-depth Post Irradiation Examination has been conducted wherein fuel-cladding chemical reactions were among those items studied. This study concluded that early unpressurized
elements containing unstable fuel were more susceptible to stress corrosion attack than are those of the current design t hat utilizes stable fuel and pre ssurized cladding. Since stress corrosion attack is the result of a combination of stress imposed by the fuel on the cladding and
the corrosive chemical species available to the cladding, irradiation programs are being pursued to define the conditions under which pellet-clad interaction will damage the cladding. These
programs are currently underway both at Hal den and in the Pathfinder test program being conducted jointly with KWU in the Obrigheim and Petten reactors.
4.2.3.2.4 Fretting Corrosion
The phenomenon of fretting corrosion, particularly in Zircal oy clad fuel rods supported by Zircaloy spacer grids, has been extensively investigated. Since i rradiation-induced stress relaxation causes a reduction in grid spring load, spacer grids must be designed for end-of-life conditions as well as beginning-of-life conditions to prevent fretting caused by flow - induced tube vibrations. To ensure this, out-of-pile fretting tests have been performed concentrating on the more severe end-of-life conditions. Two testing approaches have been used; i.e., autoclave vi bration tests and dynamic flow tests.
4.2.3.2.4.1 Autocl ave Vibration Tests
The autoclave tests were performed by vibrating a f uel rod sample supported by two rigidly held spacer grid sections. Test conditions ma tched reactor coolant chemistry, te mperature, and pressure. Variable parameters provided data to evaluate the effects of:
a) Frequency of tube vibration
b) Spacer grid spring load (preset)
c) Axial tube movement (simulating r eactor load following characteristics)
Data from such tests have indicated that wear star ts with a brief break-in period and then proceeds at a negligible rate. Changes in frequency, spring pres ent (including zero preset) and amplitude within representative limits do not significantly alter fretting characteristics.
At no time under any conditions was fretting significant.
WSES-FSAR-UNIT-3 4.2-49 Revision 302 (12/08) 4.2.3.2.4.2 Dynamic Flow Tests
Dynamic flow tests have been performed on four x four rod arrays (16 fuel rods) and on full size fuel assemblies. The four x four rod array te sting was conducted under the following conditions:
a) Flow velocities ranged from 14 ft/sec to 25 ft/sec b) Coolant temperature was 590 F c) Coolant pressure was 2150 psia
In addition, the four x four rod arrays were subj ected to cross-flow and a mechanically induced forced vibration of the lower end of the rod array at a frequency of 15 Hz and an amplitude of five mils (representing vibratory forces imparted by the reactor internals). The four x four rod array testing also included rod arrays with preset spacer grid springs ranging from approximately 10 mils interference to gaps of up to five mils, simulating both tightly held and l oose rods. The four x four rod arrays were tested for intervals from 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> up to 3182 hours0.0368 days <br />0.884 hours <br />0.00526 weeks <br />0.00121 months <br /> for a total accumulated test time of 18,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />.
The fuel rods in the four x four assemblies were eit her of a 0.413 in. diameter on a 0.550 in. pitch or of a 0.440 in. diameter on a 0.580 in. pitch, which are representative of a 15 x 15 and 14 x 14 fuel array, respectively. All fuel rods were visually inspected at each spacer grid interface. The depth of wear marks was accurately determined using an optical microm eter. The maximum depth of wear noted for the conditions above was less than 1/2 mil. In a special test where a fuel rod was completely unsupported at its lower end for a distance of 15 in., a depth of wear of three mils was noted after 2000 hours0.0231 days <br />0.556 hours <br />0.00331 weeks <br />7.61e-4 months <br /> of flow at
25 ft/sec. This test was not representative of any design condition, but was performed to demonstrate the need for supporting the lower end of the fuel rod. Bas ed on test results of four x four assemblies which follow the same trend as found in the autoclave vibr ation test, the maximum expected clad wear at end-of-life will be will be less than three mils.
Separate full scale flow tests at or exceeding reactor flow conditions were run with an array of four full-size prototypical 15 x 15 fuel assemblies, four fu ll size 14 x 14 fuel assemblies of which two were prototypical and two contained stainle ss steel fuel rods, several tests of individual full size 14 x 14 fuel assemblies and a prototype 16 x 16 fuel assembly. The test conditions were as follows:
a) Flow velocities ranged from 16 ft/sec for 15 x 15 fuel assemblies up to 23.7 ft/sec for some of the 14 x 14 fuel assemblies, and 22 ft/sec for the 16 x 16 f uel assembly. In all cases, the flow test velocities exceeded the maximum calculated velocity at operating conditions for fuel assemblies
in each particular reactor.
b) A large number of fuel rods (in some cases all rods within a fuel assembly) were tested with zero preset Zircaloy spacer grid spring loads to c onservatively represent end-of-life spacer grid conditions. A number of fuel rods were also l oosely supported at various spacer grid locations, and, in some cases, over t he entire length of the rod.
c) Test time accumulated exceeded 13,500 hours0.00579 days <br />0.139 hours <br />8.267196e-4 weeks <br />1.9025e-4 months <br />, wi th the longest single test 4000 hours0.0463 days <br />1.111 hours <br />0.00661 weeks <br />0.00152 months <br /> in length.
The results of these tests are similar to those of the four x four fuel a ssemblies with a few exceptions. On the 14 x 14 fuel assemblies subjected to 4000 hours0.0463 days <br />1.111 hours <br />0.00661 weeks <br />0.00152 months <br /> of continuous testing at 23.7 ft/sec and 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> at 19.1 ft/sec, the maximum depth of wear on one assembly was 1.7 mils, while on the other assembly, one wear mark was found to be 2.2 mils deep and a fe w others ranged from 1.6 to 1.8 mils. The only incidence of significant wear on a full size fuel assembly occurred in special test of an off-design condition
where the lower end of the fuel assembly was essentia lly unrestrained laterally. In this test, the depth of wear of one fuel rod was 10.9 mils after only 1188 hours0.0138 days <br />0.33 hours <br />0.00196 weeks <br />4.52034e-4 months <br /> of testing at 23.7 ft/sec. A gain, this test, as in the case of the cantilevered fuel rod test in a f our x four fuel assembly, showed the need for laterally restraining the lower end of the fuel assembly.
WSES-FSAR-UNIT-3 4.2-50 Revision 302 (12/08)
(EC-9533, R302)
Results for the 16 x 16 fuel assembly test, where the Zi rcaloy spacer grid springs were preset to zero interference and the Inconel grid springs were preset to a small interference with the fuel rods, showed no evidence of fretting on any fuel rod after 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> of testing.
An extensive flow test program was conducted to support the implementation of the NGF design. The NGF configuration was tested to confirm the hydrau lic stability of the fuel assembly design and to demonstrate the acceptability of the fretting performanc e of the fuel assembly design. These tests included full scale single bundle tests of the NGF and standard designs, a full scale dual bundle test with a NGF fuel assembly and a standard fuel assembly, and a full cross-section/short length bundle test of the NGF design. The single bundle tests were run to evaluate the hydraulic stability of the fuel assemblies. The tests demonstrated the hydraulic st ability of both designs ov er the expected range of flow rates. The dual bundle test was an endurance te st to evaluate fretting performance of the two designs. This test provided additional confirmation of the hydraulic stability of the designs and showed a
significant improvement in the fretting performance of the NGF design compared to the standard design. The short length bundle test was run to confirm the absence of flow-induced strip vibration within the
spacer grids, which it did. The successful results of the flow test program demonstrate that the hydraulic performance of the NGF design is acceptable and superior to that of the prior designs. (EC-9533, R302)
4.2.3.2.5 Cycling and Fatigue
A fatigue analysis is performed to determine the cu mulative fatigue damage of fuel rods exposed to lifetime power cycling conditions. The fatigue cycl e is determined by considering combinations of normally anticipated events that would produce conservative estimates of st rain in the clad. Some of the major conservative assu mptions are as follows:
a) Hot spot fuel radii are used in the calculations
b) The most adverse tolerance conditions on the fuel and cladding dimensions are chosen to produce maximum interactions and hence maximum clad strains.
The chosen fatigue cycle represents daily operation at both full and reduced power. Clad strains are
calculated from the primary creep rate of the clad and used to calculate the effective strain ranges. The cumulative fatigue damage fraction is determined by su mming the ratios of the number of cycles at a given effective strain range to the permitted number at that range as taken from the fatigue curve
presented in Figure 4.2-2.
4.2.3.2.6 Fuel Rod Bowing
Analysis of bowing data has shown that the bowing expected in the 16 x 16 design will have no effect on the margin to DNB beyond the allowance provided by the pitch, bowing and clad diameter enthalpy rise factor given in Table 4.4-1 and discussed in Section 4.4. A more complete discussion of the cause and effects of rod bowing is presented in References 53 and 75.
4.2.3.2.7 Irradiation St ability of Fuel Rod Cladding
The combined effects of fast flux and cladding temper ature are considered in three ways as discussed below:
a) Cladding Creep Rate
WSES-FSAR-UNIT-3 4.2-51 Revision 302 (12/08)
(EC-9533, R302)
The in-pile creep performance of Zircaloy-4, ZIRLO TM , and Optimized ZIRLO TM are dependent upon both the local material temperature and the loca l fast neutron flux. The functional form of the dependencies for Zircaloy-4 cladding is pr esented in Reference 14 for gap conductance calculations, and in Reference 22 for cladding collapse time predictions. The corresponding functional form of the dependencies is presented in Reference 80 for ZIRLO TM cladding, while Reference 83 documents that the similarities between the ZIRLO TM cladding and the Optimized ZIRLO TM cladding result in the same correlation for the two materials. (EC-9533, R302) b) Cladding Mechanical Properties
The yield strength, ultimate strength, and duc tility of Zircaloy-4 are dependent upon temperature and accumulated fast neutron fluence. The temperature and fluence dependence is discussed in
Subsection 4.2.1.2.2.1. Uni rradiated or irradiated properties were used depending upon which is more restrictive for the phenomenon being evaluated.
c) Irradiation Induced Dimensional Changes (EC-9533, R302)
Zirconium-based alloys have been shown to sustain dimensional changes (in the unstressed condition) as a function of the accumulated fast fluence. These changes are considered in the appropriate clearances between the various core components. The irradiation induced growth correlation method is discussed in Reference 3 (s ee Subsection 4.2.3.1.4), with the rod growth correlations specified in Reference 82 for Zircloy-4 cladding, Reference 80 for ZIRLO TM cladding, and Reference 84 for Optimized ZIRLO TM cladding.
Zircaloy-4 fuel cladding has been utilized in pressuri zed water reactors at temperatures and burnups anticipated in current designs with no failures attri butable to radiation damage. Mechanical property tests on Zircaloy-4 cladding exposed to neutron irradiation of 4.7 x 10 21 nvt (estimated) have revealed that the cladding retains a significant amount of ductility (in excess of four percent elongation). Typical results are shown in Table 4.2-2. It is believed that the fluence of 4.7 x 10 21 nvt is at saturation so that continued exposure to irradiation will not change these properties.
(54) Similar performance has been experienced with ZIRLO TM cladding and Optimized ZIRLO TM cladding, as detailed in Reference 80 and 83, respectively. (EC-9533, R302) 4.2.3.2.8 Cladding Collapse Analysis (DRN 02-1538, R12)
A cladding collapse analysis is performed to ensure that no fuel rod in the core will collapse during its design lifetime. The clad collapse calculation method (22) itself does not include arbitrary safety factors.
However, the calculation inputs are deliberately select ed to produce a conservative result. For example, the as-built clad dimensional data are chosen to be worst case combinations which result in 95 percent confidence in minimum predicted collapse time; the internal pressure history is based on minimum fill
pressure with no assistance from released fission gas; and the flux and temperature histories are based on conservative assumptions. The combined effect of using conservative inputs in the clad buckling analysis method is to produce computed collapse ti me in excess of three cycles of operation. (DRN 02-1538, R12) 4.2.3.2.9 Fuel Dimensional Stability
Fuel swelling due to irradiation (accumulation of solid and gaseous fission products) and thermal expansion results in an increase in the fuel pellet diameter. The design makes provision for
accommodating both forms of pellet growth. The f uel-clad diametral gap is more than sufficient to accommodate the thermal expansion of the fuel. To accommodate irradiation-induced swelling, it is conservatively assumed that the fuel-clad gap is reduced by the thermal expansion and that only the volume due to fuel porosity and the dishes on each end of the pellets are available. Thermal and irradiation induced creep of the restrai ned fuel results in redistribution of fuel so that the swelling due to irradiation is accommodated by the free volume (eight percent of the fuel volume).
WSES-FSAR-UNIT-3 4.2-52 Revision 302 (12/08)
For such restrained pellets, and at a total fissi on-product-induced swelling rate of 0.7 percent V/V per 10 20 fissions/cm 3 , 0.54 percent would be accommodated by the fuel porosity and dished pellet ends through fuel creep, and 0.16 percent would increase the fuel diameter. Assuming peak burnup, this would correspond to using up a void volume equal to approximately 7.4 percent of the fuel volume and increasing the fuel rod diameter by a maximum of
< 0.0025 in. (< 0.7 percent clad strain). When these numbers were compared to the minimum available vo lume and the maximum allowable strain, it was concluded that sufficient accommodation volume has been provided even under the most adverse burnup and tolerance conditions.
Demonstration of the margin which exists is seen in the large seed blanket reactor (LSBR) irradiation.
Two rods which operated in the B-4 loop of the MT R offer an interesting simulation for current PWR design.(6) (7) (55)
Both rods were comprised of 95 percent theoretical density pellets with dished ends and clad in Zircaloy. The first of these, No. 79-21, was operated successfully to a burnup of 12.41 x 10 20 fissions/cm 3 (>48.00 MWd/MTU). The second fuel pin, No.
79-25, operated successfully to 15.26 x 10 20 fissions/cm 3 (>60,000 MWd/MTU). The linear heat rating ranged from 7.1 to 16.0 KW/ft. The wall thickness for the latter pin was 0.028 in. as compared with 0.016 in. for the former. All other parameters were essentially identical. The two rods were assembled by shrink ing the cladding onto the fuel. The maximum diametral increase measured at the ridge heights for rod 79-21 wa s 0.005 in., while it was less than 0.002 in. for rod 79-25. From post-irradiation examination, it was concluded that approxim ately 84 percent of the total fuel swelling was accommodated by the porosity and dishes , while 16 percent caused diametral expansion of the clad and ridging at pellet interfaces. These results indicate that a comparable irradiation of the fuel elements for Waterford 3 (cold diametral gap 0.007 in
., wall thickness of 0.025 in., density 94.75 percent TD) would allow adequate margin for swelling accommodation.
The successful combined VBWR-Dresden irradiation of Zircaloy-clad uranium dioxide pellets provides additional confidence with respect to the design conditions for the fuel rods for this core.
(56)(57) Ninety-eight rods which had been irradiated in VBWR to an average burnup of about 10,700 MWd/MTU were
assembled in fuel bundles and irradiated in Dres den to a peak burnup greater than 48,000 MWd/MTU.
The reported maximum heat rating for these rods is 17.3 KW/ft which occurred in VBWR. Post-irradiation examination (58) revealed that diametral increases in the fuel rods ranged from 0.001 to 0.003 in.
maximum. The maximum diametrical change corresponds to 1.42 percent V/V, (or 0.12 percent V/V per 10 20 fission/cm
- 3) for these 0.424 in. diameter rods. The relevant fuel parameters are listed below for the above test and the Waterford 3 design. (DRN 00-644)
Fuel Density Cold Diametral Peak Burnup
%TD Gap (in.) (MWd/MTU)
VBWR-Dresden 95 0.004 to 0.008 >48,000
LSBR-MTR 95 0.001 50,000; 61,000
Waterford 3 94.75 0.007 55,000
A comparison of the design parameters above, relative to the test results, provi des a demonstration of the clad strains resulting from swelling of fuel.
4.2.3.2.10 Potential for Waterloggi ng Rupture and Chemical Interaction
The potential for waterlogging rupture is considered remote. Basically, the necessary factor or combination of factors, include the presence of a sma ll opening in the cladding, time to permit filling of the fuel rod with water, and finally a rapid power trans ient. The size of the opening necessary to cause a problem falls within a fairly narrow band. Above a cert ain defect size, the rod can fill rapidly, but during a WSES-FSAR-UNIT-3 4.2-53 Revision 302 (12/08) power increase it also expels water or steam readily without a large pressure buildup. Defects which could result in an opening in cladding are scrupulously checked for during the fuel rod manufacturing process by both ultrasonic and helium leak testing.
Clad defects which could develop during reactor operation due to hydriding are also controlled by lim iting those factors (e.g., hydrogen content of fuel pellets) which contribute to hydriding.
The most likely time for a waterlogging rupture incident would be after an abnormally long shutdown period. After this time, however, the startup rate is controlled so that even if a fuel rod were filled with coolant, it would "bake out", thus minimizing t he possibility of additional cladding rupture. The combination of control and inspection during the m anufacturing process and the limits on the rate of power change restrict the potential for waterlogging r upture to a very small number of fuel rods.
The UO 2 fuel pellets are highly resistant to attack by reactor coolant in the event cladding defects should occur. Extensive experimental work and operating experience have s hown that the design parameters chosen conservatively account fo r changes in thermal performance during operation and that coolant activity buildup resulting from cladding rupture is limit ed by the ability of uranium dioxide to retain solid and gaseous fission products.
4.2.3.2.11 Fuel-Cladding Interaction
An analytical model to evaluate cladding res ponse to pellet-clad interaction has been developed (21). This analysis which is based on an advanced versi on of the FATES computer code, considers generalized plane-strain of a unit section of fuel and clad. All of the physical phenomena calculational
methods and input variables of the present FATES pr ogram(14) are included in the new version; and in addition, models are included for elastic and plastic stresses and strains in the clad and fuel, and fuel
creep. A compatible interface modeled between the fuel and clad ensures that interaction is accurately accounted for.
The treatment of power history, axial power shapes and other operating parameters is handled similar to the current FATES version with the exception that power ramp rates and cycling can be considered. The response of the fuel and clad is calculated through an iteractive process, and the interaction between the
fuel and clad is established.
The resulting analytical predictions of temperatur es, stresses, strains and geom etric configuration are thus made available for use in conjunction with operating experience and irradiation test results in
demonstrating the acceptability of the various operating conditions to which the fuel may be subjected. A detailed discussion of the methods and capabilities of the pellet-clad interaction model is contained Reference 21.
4.2.3.2.12 Fuel Burnup Experience (EC-9533, R302)
Design bases for the Zircaloy-4, ZIRLO TM , and Optimized ZIRLO TM cladding have been established which are conservative with respect to the reported data. Evidence currently available indicates that these claddings and UO 2 fuel performance is satisfactory to exposures in excess of 60,000 MWd/MTU. (EC-9533, R302) a) High Linear Heat Rating Irradiation Experience (DRN 00-644) The determination of the effect of linear heat rating and fuel cladding gap on the performance of Zircaloy-clad UO 2 fuel rods was the object of two ex perimental capsule irradiation programs conducted in the Westinghouse Test Reactor (WTR).
(59) In the first program, 18 rods containing 94 percent theoretical density UO 2 , pellets were irradiated at 11, 16, 18, and 25 kW/ft with cold diametral gaps of 0.006 in., 0.012 in., and 0.025 in. T he wall thickness to diameter ratio (t/OD) of the Zircaloy cladding was 0.064 which is comparable to the 0.066 value in this design. Although (DRN 00-644)
WSES-FSAR-UNIT-34.2-54 Revision 12 (10/02)theseirradiations were of short duration (about 40 hours4.62963e-4 days <br />0.0111 hours <br />6.613757e-5 weeks <br />1.522e-5 months <br />) significant results applicable to this designwere obtained. No significant dimensional changes were found in any of the fuel rods. Only one
rod, which operated at a linear heat rate of 24 kW/ft with an initial diametral gap of 0.025 in.,
experienced center melting. Rods which operated at 24 kW/ft with cold gaps of 0.006 in. and 0.012 in. did not exhibit center melting. On these bases, the initial gap of 0.007 in. and the maximum
linear heat ratings for this design provide adequate margin against center melting, even when 112 percent overpower conditions are considered. These results also indicate that an initial diametral
gap of 0.007 in. is adequate to accommodate radial thermal expansion without inducing cladding dimensional changes even at a linear heat rate of 24 kW/ft. This margin with respect to thermal
expansion will be diminished with increasing burnup at a rate of 0.16 percent V/V per 10 20 fissions/cm
- 3. However, the linear heat rating will decrease with burnup and thus limit the sum of the strains to values below the allowable.Further substantiation of the capability of operation at maximum linear heat ratings in excess of those in this design is obtained from later irradiation tests in WTR.
(59) Fuel rods 38 in. long and 6 in. long were irradiated at linear heat ratings of 19 kW/ft and 22.2 kW/ft to burnups of 3450 and 6250MWd/MTU. The cold diametral gaps in these Zircaloy clad rods containing 94 percent dense UO 2were 0.002 in., 0.006 in., and 0.012 in. The cladding t/OD was 0.064. No measurable diameterchanges were noted for the 0.006 in. or 0.012 in. fuel clad gap rods. Only small changes were
observed for the rods with a 0.002 in. diametral gap.b)Shippingport Irradiation Experience Zircaloy clad fuel rods have operated successfully (three defects have been observed which were aresult of fabrication defects) in the Shippingport blanket with burnups of about 37,000 MWd/MTU and maximum linear heat ratings of about 13 kW/ft.
(59)(60)(61) Although higher linear heat ratings will be experienced, swelling (primarily burnup dependent) and thermal expansion (linear heat ratingdependent) provide the primary forces for fuel cladding strain at the damage limit. Thus, the Shippingportirradiations have demonstrated that Zircaloy clad rods with a cladding t/OD less than that for this plant (0.066) can successfully contain the swelling associated with 37,000 MWd/MTU burnup while at the same time containing the radial thermal expansion associated with peak linear heat ratings. Irradiation test programs in support of Shippingport in-reactor loops demonstrated
successful operation at burnups of 40,000 MWd/MTU and linear heat ratings of about 11 kW/ft with cladding t/OD ratios as low as 0.053.
(62)c)Saxton Irradiation Experience(DRN 02-1538)
Zircaloy-4 clad fuel rods containing UO 2-PuO 2 pellets of 94 percent theoretical density have beensuccessfully irradiated in Saxton to peak burnups of 31,800 MWd/MTU at 15 kW/ft linear heat rate under USAEC Contract AT (30-1)-3385 (63). The t/OD of the cladding was 0.059 which is less thanthat of this design. The amount of PuO 2 , 6.6 percent is considered as insignificant with respect to providing any difference in performance when compared with that for UO
- 2. Subsequent tests on two of the above rods (18,000 MWd/MTU at 10.5 kW/ft) successfully demonstrated the capability of these rods to undergo power transients from 16.8 kW/ft to 18.7 kW/ft.(DRN 02-1538)
WSES-FSAR-UNIT-34.2-55 Revision 12 (10/02)d)Vallecitos Boiling Water Reactor (VBWR) - Dresden Experience(DRN 02-1538)
The combined VBWR - Dresden irradiation of Zircaloy clad oxide pellets provides additional confidence with respect to the design conditions for the fuel rods for this core.
(55)(57)(64) Ninety-eight rods which had been irradiated in VBWR to an average burnup of about 10,700 MWd/MTU were assembled in fuel bundles and irradiated in Dresden to a peak burnup greater than 48,000MWd/MTU. The reported maximum heat ratings for these rods is 17.3 kW/ft which occurred in VBWR. The t/OD cladding ratio of 0.052, and the external pressure of about 1000 psia are
conditions which are all in the direction of less conservatism with respect to fuel rod integrity when compared with the design values of 0.066 cladding t/OD ratio and an external pressure of 2250 psia.
Ten of these VBWR - Dresden rods representing maximum com binations of burnup, linear heat rating, and pellet density have been examined in detail and found to be in satisfactory condition.
The remaining 88 rods were returned to Dresden and successfully irradiated to the termination of the
program.(DRN 02-1538)e)Large Seed Blanket Reactor (LSBR) Rods Experience Two rods operated in the B-4 loop at the Materials Testing Reactor (MTR) provide a very interesting simulation for current PWR designs (6)(7)(52). Both rods were comprised of 95 percent theoreticaldensity pellets with dished ends, clad in Zircaloy. The first of these No. 79-21, was operated successfully to a burnup of 12.41 x 10 20 fiss/cm 3 (48,000 MWd/MTU) through several power cycleswhich included linear heat rates from 5.6 to 13.6 kW/ft. The second fuel pin, No. 79-25, operated
successfully to 15.26 x 10 20 fiss/cm 3 (60,000 MWd/MTU). The basic difference in this rod was the0.028 in. wall thickness as compared to 0.016 in. (t/OD = 0.058) in the first rod. All other
parameters were essentially identical.
The linear heat rating ranged from 7.1 to 16.0 kW/ft. After the seventh interim examination, the rod operated at a peak linear power of 12.9 kW/ft at a time when the peak burnup was 49,500MWd/MTU. These high burnups were achieved with fuel elements which were assembled byshrinking the cladding onto the fuel and indicate that a comparable irradiation of the fuel elements for this reactor (cold diametral gap of 0.007 in.) would allow a considerable increase in swelling life
at a given clad strain.f)Central Melting in Big Rock Point Experience(DRN 00-644)
As part of a joint U.S. - Euratom Research and Development Program, Zircaloy clad UO 2 pellet rods (95 percent theoretical density) were irradiated under conditions designed to induce central melting
in the Consumers Power Co. Big Rock Point Reactor (65). The test includes 0.7 in. diameter fuel rods (cladding t/OD = 0.057; fuel clad gap of about 0.012 in.) at maximum linear heat ratings of
about 27 kW/ft and 22 kW/ft with peak burnups up to 30,000 MWd/MTU. Result of these irradiations provide a basis for incorporating linear heat ratings well in excess of those calculated for
this reactor, and show that the presence of localized regions of fuel melting is not catastrophic to
the fuel rod.(DRN 00-644)
WSES-FSAR-UNIT-3 4.2-56 Revision 309 (06/16) g) KWU Irradiations-Kraftwerk Un ion Reactor, Obrigheim, Germany
C-E has entered into a technical agreement with Kraftwerk Union (KWU) for the complete exchange of information and technology relating to pressurized water reactor systems including fuel.
This agreement makes available to C-E the exper ience of eight years su ccessful operation of the KWO reactor at Obrigheim, Germany. (DRN 00-644; 02-1538, R12) In the area of nuclear fuel performance, t he experience at Obrigheim has shown successful operation through seven operating cycles. Fuel batches of 95 percent TD, both pressurized and nonpressurized, have been irradiated.
Substantial testing has been performed in the reactor on the load following ability of both pressurized and nonpr essurized fuel rods. Selected rods were subjected to power changes from 50 to 100 percent at rates of 20 percent/min for more than 900 cycles. Peak power densities in the rods we re 15 kW/ft with maximum burnups in excess of 30,000 MWd/MTU. No failures have been observed to date. This experiment demonstrates the load-following capability of a design similar to C-E's in an operating PWRs. (DRN 00-644; 02-1538, R12) h) Long Term Irradiation Testing
As indicated, C-E has several self-sponsored f uel irradiation programs in progress and several cooperative fuel development programs with Kraftw erk Union as part of a technical agreement.
In addition, C-E has access to all data and result s of Kraftwerk Union's own fuel development programs. (EC-9533, R302) i) High Burnup Combustion Engineering Operational Experience
Reference 82 presents fuel performance data obt ained during poolside examinations and hot cell examinations of high burnup fuel utilizing Zircal oy-4 cladding in Combustion Engineering cores.
The data demonstrates the acceptability of the f uel's performance to rod average exposures in excess of 60,000 MWD/MTU. j) Westinghouse Experience
ZIRLO TM cladding material is in widespread use domesti cally in at least 38 nuclear power plants (Reference 80, Section 3.3). ZIRLO TM has been shown to have improved corrosion resistance compared to Zircaloy-4. Also, no oxide spalling has been observed in current ZIRLO TM fuel rods for normal operation. (LBDCR 15-035, R309)
Optimized ZIRLO TM cladding has a slightly lower allowed tin level than ZIRLO TM (lower by 0.2%)
with the remainder of the material composit ion requirements being the same. The reduced tin level is to further enhance the corrosion resistanc e of the cladding. Reference 83 documents that ZIRLO TM material properties currently utilized in various models and methodologies are applicable to analyses for Optimized ZIRLO TM and shows the differences are negligible with no impact on any design or safety analyses. Re ference 85 presents additional cladding corrosion data and provides updated cladding co rrosion models for both ZIRLO TM and Optimized ZIRLO TM cladding. Therefore, in addition to the oper ational experience of Optimized ZIRLO TM , the ZIRLO TM operational experience discussed above is applicable to Optimized ZIRLO TM. (EC-9533, R302, LBDCR 15-035, R309) 4.2.3.2.12.1 Combustion Engineer ing Fuel Development Programs (DRN 02-1538, R12)
Since mid-1972, C-E has performed an extensive irradi ation test program on fuel densification. When fuel densification became apparent, C-E immediately init iated an irradiation test program to determine the causes of densification and to defi ne the specifications and processes required to limit densification of fuel. The first irradiation test program in the sequence confirmed that the phenomena is real and defined (DRN 02-1538, R12)
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(DRN 02-1538, R12) the parameters important in the effect. An imm ediate response was a change in the C-E fuel pellet specification and a modification of the fuel fabric ation process to provide densification resistant UO 2 fuel. The irradiation tests are continuing to establish conc lusively that the current specification and process used is effective in minimizing densification. (DRN 02-1538, R12)
C-E is also a participating member of the Halden Reac tor Project in Halden, Norway. The Halden project has underway a spectrum of fuel development programs from which C-E can further verify present fuel design models and continually evaluat e advanced fuel design concepts.
4.2.3.2.12.2 Combustion Engineering/Kra ftwerk Union Fuel Development Programs
The primary objectives of the cooperat ive fuel development programs are:
a) To assess the causes of fuel densificati on and provide process changes which will preclude densification. Then subsequently to verify thr ough irradiation testing that the process changes have been effective. b) To obtain long term data to further verify fuel performance models.
c) To evaluate advanced fuel design concepts in-reactor.
C-E and KWU currently have three densification test programs in progress in both United States and European test reactors. In addition, C-E and KWU are participating extensively in the densification test program under primary sponsorship of the Edison Electric Institute.
4.2.3.2.12.3 Kraftwerk Uni on Fuel Development Programs (DRN 00-644) The design of the C-E fuel rods is very similar to the KWU fuel utilized in the Obrigheim reactor. The
Obrigheim core has operated with peak power densitie s up to 15 kW/ft with maximum burnups in excess of 46,000 MWd/MTU without observed life limiting failures. Several fuel rods, both pressurized and
unpressurized, from the Obrigheim reactor hav e undergone detailed hot cell examination under the direction of KWU. The results of all nondestructive fuel examinations performed during shutdowns and
the complete results of the hot cell program are available to C-E under the technical agreement with KWU. (DRN 00-644)
In addition to the programs to routinely examine high burnup standard fuel, KWU also has comprehensive fuel development programs underway which utilize special test assemblies in the Obrigheim reactor. Under this program, fuel rod design parameters have been varied over significant ranges to experimentally establish the basis for further design optimization.
One assembly has been irradiated annually since October 1973. Also included in this s pecial assembly are segmented rods or "rodlets" which are connected to form a complete fuel rod. T hese rodlets are preirradiated in a test reactor. The test reactor irradiations provide data on fuel rod performance under transient conditions.
In summary, C-E has in process, or in the planni ng stages, fuel development pr ograms that will provide additional assurance of fuel design adequacy.
4.2.3.2.13 Temperature Tr ansient Effects Analysis
4.2.3.2.13.1 Waterlogged Fuel
The potential for a fuel rod to become waterlogged during normal operation is discussed in Subsection 4.2.3.2.10. In the event that a f uel rod does become waterlogged at low or zero power, it is possible that a subsequent power increase could cause a buildup of hydrostatic pressure. It is unlikely that the pressure would build up to a level that could cause cladding rupture because a fuel pin with the potential
for rupture requires the combination of a very small defect together with a long period of operation at low or zero power.
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Tests which have been conducted using intentionally waterlogged fuel pins (capsule drive core at SPERT)(66)(67) showed that the resulting failures did eject some fuel material from the rod and greatly deformed the test specimens. However, these test r ods were completely sealed, and the transient rates used were several orders of magnitude greater than those allowed in normal operation.
In those instances where waterlogged fuel rods hav e been observed in commercial reactors, it has not been clear that waterlogging was the cause, and not ju st the result, of associated cladding failures; and C-E has not observed any case in which mate rial was expelled from waterlogged fuel rods.
It is therefore, concluded that the effect of normal power transients on waterlogged fuel rods is not likely to result in cladding rupture and even if rupture does o ccur it will not produce the sort of postulated burst failures which would expel fuel material or damage adj acent fuel rods or fuel assembly structural components.
4.2.3.2.13.2 Intact Fuel
The thermal effects of anticipated operational occurrences on fuel rod integrity are discussed in the
following paragraphs.
a) Fuel rod thermal transient effects are basically manifested as the change in internal pressure, the changes in clad thermal gradient and thermal st resses, and the differential thermal expansion between pellets and clad. These effects are discussed in Subsections 4.
2.3.2.2 and 4.2.3.2.11.
b) Another possible effect of transients would be to cause an axial expansion of the pellet column against a flattened (collapsed) section of the clad.
However, the fuel rod design includes specific provisions to prevent clad flattening, and, t herefore, such interactions will not occur.
4.2.3.2.14 Energy Release During Fuel Element Burnout
The Reactor Protection System provides fuel clad protection so that the probability of fuel element burnout during normal operation and anticipated operational occurrences is extremely low. Thus, the potential for fuel element burnout is restricted to faulted conditions. The LOCA is the limiting event since it results in the larger number of fuel rods exper iencing burnout; thus the LOCA analysis, which is very conservative in predicting fuel element burnout, pr ovides an upper limit for evaluating the consequences of burnout. The LOCA analysis explicitly accounts for the additional heat release due to the chemical reaction between the Zircaloy clad and the coolant following fuel element burnout in evaluating the consequences of this accident.
LOCA analysis results are di scussed in Subsection 15.6.3.
4.2.3.2.15 Energy Release on Rupture of Waterlogged Fuel Elements
A discussion of the potential for waterlogging fuel r ods and for subsequent energy release is presented in Subsection 4.2.3.2.10.
4.2.3.2.16 Fuel Rod Behavior Effe cts from Coolant Flow Blockage
An experimental and analytical program was conducted to determine the effects of fuel assembly coolant flow maldistribution during normal reactor operation.
In the experimental phase, velocity and static pressure measurements were made in cold, flowing water in an oversize model of a C-E 14 x 14 fuel
assembly in order to determine the three-dimensional WSES-FSAR-UNIT-34.2-59Revision 11 (05/01)flow distributions in the vicinity of several types of flow obstructions. The effects of the distributions onthermal behavior were evaluated, where necessary, with the use of a preliminary version of the TORC thermal and hydraulic code(68). Subjects investigated included:a)The assembly inlet flow maldistribution caused by blockage of a core support plate flow hole.Evaluation of the flow recovery data indicated that even the complete blockage of a core support plate flow hole would not produce a W-3, Burnout Heat Flux Correlation, DNBR of less than 1.0 even though the reactor might be operating at a power sufficient to produce a DNBR of 1.3 without the blockage.b)The flow maldistribution within the assembly caused by complete blockage of one to ninechannels was also evaluated. Flow distributions were measured at positions upstream and downstream of a blockage one to nine channels. The influence of the blockage diminished very rapidly in the upstream direction. Analysis of the data for a single channel blockage indicated that such a blockage would not produce a W-3 DNBR of less than 1.0 downstream of the blockage even though the reactor might be operating at a power sufficient to produce a DNBR of 1.3 without the blockage. (DRN 00-644)The results presented above were obtained through flow testing an oversize model of a standard 14 x 14fuel assembly. Because of the great similarity in design between the Waterford 3 16 x 16 assembly, and the earlier 14 x 14 array, these test results also constitute an adequate demonstration of the effects thatflow blockage would have on the 16 x 16 assembly. This conclusion is also supported by the fact that the16 x 16 assembly has been demonstrated to have a greater resistance to axial flow than the 14 x 14 assembly. The higher flow resistance of the 16 x 16 arrangement would lead to more rapid flow recovery downstream of any blockage than would occur with the 14 x 14 array. The effect of the higher flow resistance is to produce a more rapid flow recovery (i.e., more nearly uniform flow) and is analogous to thecommon use of flow resistance devices (screens or perforated plates) to smooth non-uniform velocityprofiles in ducts or process equipment. (DRN 00-644)4.2.3.2.17Fuel TemperaturesSteady state fuel temperatures are determined by the FATES computer program. The calculationalprocedure considers the effect of linear heat rate, fuel relocation, fuel swelling, densification, thermal expansion, fission gas release, and clad deformations. The model for predicting fuel thermal performance is discussed in detail in Reference 14. (DRN 00-644)Two sets of burnup and axially dependent linear heat rate distributions are considered in the calculation.One is the hot rod, time averaged, distribution expected to persist during long term operation, and the other is the envelope of the maximum linear heat rate at each axial location. The long term distributionsare integrated over selected time periods to determine burnup, which is in turn used for the various burnupdependent behavioral models in the FATES computer program. The envelope accounts for possible variations in the peak linear heat rate at any elevation which may occur for short periods of time and is used exclusively for fission gas release calculations. (DRN 00-644)
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The power history used assumes continuous 100 percent power from beginning-of-cycle. Using this history, the highest fuel temperatures occur at beginning-of-life. It has been shown that fuel temperatures
for a given power level and burnup are insensitive to the previous history (e.g., operating power transients, length and number of shutdowns, etc.) used to arrive at the given power level.
Fuel thermal performance parameters are calculated for the hot rod. These parameters for any other rod in the core can be obtained by using the axial location in the hot rod, whose local power and burnup corresponds to the local power and burnup in the rod being examined. this procedure will yield conservatively high stored energy in the fuel rod under consideration.
The maximum power density, including the local peaking as affected by anticipated operational occurrences, is discussed in Sections 4.3, 4.4, and Chapter 15.
4.2.3.3 Burnable Poison Rod
4.2.3.3.1 Burnable Poison Rod Internal Pressure and Cladding Stress (DRN 02-1538, R12; 06-1059, R15)
A poison rod cladding analysis will be performed to determine the stress and strain resulting from the various normal, upset, and emergency conditions discussed in Subsection 4.2.1.1. Specific accounting
will be made for differential pressure, differential thermal expansion, cladding creep, and irradiation induced swelling of the burnable poison material. Owing to a lower linear heat generation rates in these rods, the cladding analyses can be accomplished using conventional strength of materials formula, except for determining clad collapse resistance which will be done using the CEPAN computer
model (22). (DRN 06-1059, R15)
The design criteria used to evaluate the analytical results are specified in Subsection 4.2.1.3.1. (DRN 02-1538, R12) 4.2.3.3.2 Potential for Chemical Reaction
A discussion of possible chemical reaction between the poison material and the coolant was presented in
Subsection 4.2.1.3.3.3, along with information on chemical compatibility between poison material and cladding. Since the cladding material is identical to that of the fuel rod (Subsection 4.2.1.3.2), the description of potential chemical reactions between cladding and coolant in Subsection 4.2.3.2.3 is
applicable to both fuel and poison rods. (DRN 00-644) The potential for waterlogging rupture in poison rods is much lower than that in fuel rods because of the smaller thermal and dimensional changes that occur in a poison rod during reactor power increases.
Refer to Subsection 4.2.3.2.10 for a discussion of the potential for waterlogging rupture in fuel rods. (DRN 00-644) 4.2.3.4 Control Element Assembly
The CEAs are designed for 10 effective full power years based on estimates of neutron absorber burnup, allowable plastic strain of the Inconel 625 cladding and the resultant dimensional clearances of the
elements within the fuel assembly guide tubes.
WSES-FSAR-UNIT-34.2-61Revision 11 (05/01)a)Internal PressureThe value of internal pressure in the control element is dependent on the following parameters:
1)Initial fill gas pressure 2)Gas temperature 3)Helium generated and released 4)Available volume including B 4C porosity (DRN 00-644)Of the absorber materials utilized in the CEA design, only the B 4C contributes to the total quantityof gas which must be accommodated within the control element. The helium is produced by thenuclear reaction on o n 1 + 5 B 103 Li 7 + 2 He 4, and the fraction of the quantity generated whichis actually released to the plenum is temperature dependent and is predicted by the empirical equation discussed in Subsection 4.2.1.4.4.A.3. Temperatures used for release fractioncalculations are the maximum predicted to occur during normal operation. (DRN 00-644)b)Thermal Stability of Absorber MaterialsNone of the materials selected for the control elements are susceptible to thermally inducedphase changes at reactor operating conditions. Linear thermal expansion, thermal conductivity,and melting points are given in Subsection 4.2.1.4.c)Irradiation Stability of Absorber MaterialsIrradiated properties of the absorber materials are discussed in Subsection 4.2.1.4. Irradiationinduced chemical transmutations are produced in both the B 4 C and theAg-In-Cd. Neutron bombardment of B10 atoms results in the production of lithium and helium.
The percent of helium released is given by the expression in Subsection 4.2.1.4.Ag-In-Cd alloy, which has an initial chemical composition of 79 w/o minimum Ag, 15
+/- 0.35 w/o In, 5+/- 0.35 w/o Cd and 0.2 w/o maximum impurities, is expected to undergo small changes incomposition. Formation of 3 w/o tin due to the transmutation of indium and an increase incadmium content to about 10 w/o due to the transmutation of silver is expected. These affect the thermal conductivity and linear expansion characteristics of the alloy and are accounted for in thedesign of the control elements.Irradiation enhanced swelling characteristics of the absorber materials are given in Subsection4.2.1.4. Accommodations for swelling of the absorbers have been incorporated in the design of the control elements and include the following measures:
WSES-FSAR-UNIT-34.2-62Revision 11 (05/01)1)All B 4C pellets have chamfered edges to promote sliding of the pellets in the cladding dueto differential thermal expansion and irradiation enhanced swelling.2)Dimensionally stable Type 304 stainless steel spacers are located at the bottom of allabsorber stacks adjacent to the nose cap to minimize strain at the weld joint.3)A hole is provided in the center of the Ag-In-Cd cylinder to accommodate swelling inexcess of the amount expected over the life of the control element.d)Potential for and Consequences of CEA Functional FailureThe probability for a functional failure of the CEA is considered to be very small. This conclusionis based on the conservatism used in the design, the quality control procedures used during manufacturing and on testing of similar full size CEA/CEDM combinations under simulated reactor conditions for lengths of travel and numbers of trips greater than that expected to occur during the Waterford 3 design life. The consequences of CEA/CEDM functional failure are discussed in Chapter 15. (DRN 00-644)A postulated CEA failure mode is cladding failure. In the event that an element is assumed topartially fill with water under low or zero power conditions, the possibility exists that upon returning to power, the path of the water to the outside could be blocked. The expansion of the entrappedwater could cause the element to swell. In tests, specimens of CEA cladding were filled with aspacer representing the poison material. All but nine percent of the remaining volume was filledwith water. The sealed assembly was then subjected to a temperature of 650
°F and an externalpressure of 2,250 psia followed by a rapid removal of the external pressure. The resulting diametral increases of the cladding were on the order of 15 to 25 mils and were not sufficient to impair axial motion of the CEA, which has a 0.084 in. diametral clearance with the fuel assembly guide tubes. This test result, coupled with the low probability of a cladding failure leading to a waterlogged rod, demonstrates that the probability for a CEA functional failure from this cause is low. (DRN 00-644)Another possible consequence of failed cladding is the release of small quantities of CEA fillermaterials, and helium and lithium (from the neutron-boron reactions). However, the amounts which would be released are too small to have significant effects on coolant chemistry or rod worth.4.2.4TESTING AND INSPECTION PLAN Fuel bundle assembly and control element assembly quality assurance is attained by adherence to theANS Quality Assurance Program Requirements for Nuclear Power Plants, ANSI N45.2-1971.
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Vendor product certifications, process surveillance, inspections, tests, and material check analyses are performed to ensure conformity of all fuel assembly and control element assembly components to the
design requirements from material procurement thr ough receiving inspection at the plant site. The following are basic quality assuranc e measures which are performed.
4.2.4.1 Fuel Assembly
A comprehensive quality control plan is established to ensure that dimensional requirements of the drawings are met. In those cases where a lar ge number of measurements are required and 100 percent inspection is impractical, these pl ans shall ensure with 95 percent c onfidence that 95 percent of these dimensions are within tolerance.
Sensitivity and accuracy of a ll measuring devices are within 10 percent of the dimensioned tolerance. The basic quality assurance measures which are performed in addition to dimensional inspections and material veri fications are described in the following sections.
4.2.4.1.1 Weld Quality Assurance Measures
The welded joints used in the fuel assembly desi gn are listed below in a series of paragraphs which describe the type and function of each weld, and include a brief description of the testing (both
destructive and non-destructive) performed to ensure the structural integrity of the joints. The welds are listed from top to bottom in the fuel assembly.
The CEA guide tube joints (between the tube and threaded upper and lower ends) are butt welds
between the two Zircaloy subcomponents. The weld s are required to be full penetration welds and must not cause violation of dimensional or corrosion resistance standards.
The upper end fitting center guide post to lower cast flow plate joint has a threaded connection which is prevented from unthreading by tack weld ing the center guide post to the bottom of the lower cast plate using the gas tungsten arc (GTA) process. Each weld is inspected for compliance with a visual standard.
(DRN 02-1538, R12)
The spacer grid welds at the intersection of perpendicu lar Zircaloy-4 grid strips are made by the laser processes. Each intersection is welded at the t op and at the bottom, and each weld is inspected by comparison with a visual standard. (DRN 02-1538, R12)
For the spacer grid to CEA guide tube weld (both components Zircaloy-4), each grid is welded to each guide tube with eight small welds, evenly divided bet ween the upper and lower faces of the grid. Each weld is required to be free of cracks and burnthrough and each weld is inspected by comparison to a visual standard. Also, sufficient testing of sample welds is required to establish acceptable corrosion resistance of the weld region. Each guide tube is in spected after welding to ensure that welding has not affected clearance for CEA motion.
The bottom spacer grid welds at spacer strip inte rsections and between spacer and perimeter strips (all components Inconel 625) have the same configuration as for the Zircaloy and are all inspected for
compliance with appropriate visual standards.
(DRN 02-1538, R12; EC-9533, R302)
The bottom spacer grid (Inconel 625) to Inconel skirt weld was made using the GTA process. Each weld was inspected to ensure compliance with a visual standard. The debris-filtering bottom spacer grid has eliminated this weldment.
(DRN 02-1538, R12; EC-9533, R302)
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(LBDCR 15-025, R309)
The Inconel skirt to lower end fitting (stainless steel) weld is made using the GTA process and each weld is inspected to ensure compli ance with a visual standard. (LBDCR 15-025, R309)
The lower end fitting is fastened to the Zircal oy guide tubes using threaded connections. The connections are prevented from unthreading by stainl ess steel locking rings which are welded to the lower end fitting. Each ring is tack welded to the end fitting in four places using the GTA process, and each weld is inspected for compliance with a visual standard. The inspection requirements and acceptance standards for each of the welds ar e established on the basis of providing adequate assurance that the connections will perform their required functions. (EC-9533, R302)
The implementation of the NGF design eliminates four weld types while introduci ng three new weld types.
As discussed in Section 4.2.2.1, the welds between the flange and the CEA guide tubes, between the Zircaloy-4 spacer grids and the guide tubes, bet ween the bottom grid and the lower end fitting, and between the locking disc and the lower end fitting have a ll been eliminated. The three new weld types are discussed below: The Inconel top grid is a brazed design composed of Inconel-718 inner and outer straps and stainless steel sleeves. The braze joints are ins pected for length and the absence of cracks. ZIRLO TM sleeves are laser welded to the Optimized ZIRLO TM inner straps of the mids grids and the IFMs. The welds are inspected for length and the absence of cracks. Stainless steel inserts are laser welded to the Inc onel inner straps of the bottom grid. The welds are inspected by comparison to a vis ual standard and the absence of cracks. (EC-9533, R302) 4.2.4.1.2 Other Quality Assurance Measures
All guide tubes are internally gaged ensuring free passage within the tubes including the reduced diameter buffer region.
Each upper end fitting post to guide tube joint is in spected for compliance with a visual standard.
The spacer grid to fuel rod relationship is carefully examined at each grid location.
Stainless steel inserts are laser welded to the Inconel inner straps of the bottom grid. The welds are inspected by comparison to a vis ual standard and the absence of cracks. (EC-9533, R302)
For NGF assemblies, inspections of the bulges are performed for size, location, and absence of cracks. (EC-9533, R302)
(DRN 06-895, R15)
Each completed fuel assembly is inspected fo r cleanliness, wrapped to preserve its cleanliness and loaded within shipping containers. DRN 06-895, R15)
Visual inspection of the conveyanc e vehicle, shipping container, and f uel assembly are performed at the reactor site. Approved procedures are provided fo r unloading the fuel assemblies. Following unloading, exterior portions of the fuel assembly components are inspected for shipping damage and cleanliness. If damage is detected, the assembly may be repaired ons ite or returned to the manufacturing facility for repair. In the event the repair process were other than one normally used by the manufacturing facility, or that the repaired assembly did not meet the standard requirements for new fuel, the specific process or assembly would be reviewed by the appropriate des ign department before the process or assembly would be accepted.
4.2.4.2 Fuel Rod
4.2.4.2.1 Fuel Pellets (DRN 00-644; 02-1538, R12)
Beginning with Batch U, all urania fuel pellets w ill be fabricated at the Columbia, SC, manufacturing facility. (DRN 00-644; 02-1538, R12)
WSES-FSAR-UNIT-3 4.2-65 Revision 302 (12/08)
(DRN 00-644; 02-1538, R12)
During the conversion of source material to ceramic grade uranium dioxide powder, the UO 2 powder is divided into lots blended to form uniform isotopic, c hemical and physical characteristics. Samples are tested from each powder blend to verify compliance with the specification limits for the blend. Additional finished pellets are tested for the final enrichment certification of the pellets.
Pellets are divided into lots during fabrication with all pellets within the lot being processed under the same conditions, as defined per the pellet specification. Representat ive samples are obtained from each lot for product acceptance tests. Hydrogen content of the finished ground pellets is restricted. The pellets' diameters are inspected and certified to meet the design tolerance requirements at a 95/99
confidence level. All other pellet dimensions meet a 90/90 confidence level. Density requirements of the sintered pellets must meet a 95/95 confidence level.
Sample pellets from each pellet lot are prepared for metallographic examination to ensure conformance to microstructural requirements. Surface finish of
ground pellets is restricted and meets a 90/90 confidence le vel. Pellet surfaces are inspected for chips, cracks, and fissures in accordance with approved standards. (DRN 00-644; 02-1538, R12)
(DRN 02-1538, R12) (DRN 02-1538, R12) 4.2.4.2.2 Cladding
(DRN 02-1538, R12)
Lots are formed from tubing produced from the same ingot, annealed in the same final vacuum annealing charge and fabricated using the same procedures. Samples randomly se lected from each lot of finished tubing are chemically analyzed to ens ure conformance to specified chem ical requirements, and to verify tensile properties and hydride orientation. Samples from each lot are also used for metallographic tests, and burst tests. Each finished tube is ultrasonically tested for internal soundness; visually inspected for cleanliness and the absence of acid stains, surfac e defects, and deformation; and inspected for inside dimension and wall thickness. The follo wing summarizes the test requirements: (DRN 02-1538, R12)
a) Test (refer to Subsection 4.2.1.2.2)
- 1) Chemical Analysis
(DRN 02-1538, R12)
Ingot analysis is required for top, mi ddle, and bottom of each ingot. Finished intermediate TREX or finished tube product is tested for hydrogen, nitrogen, and oxygen per ASTM E353. (DRN 02-1538, R12)
- 2) Tensile Test at Room Temperature (ASTM E8-69)
- 3) Corrosion Resistance Test (ASTM G2-67)
- 4) Grain Size (ASTM E112-63)
(DRN 02-1538, R12)
- 5) Deleted (DRN 02-1538, R12)
- 6) Surface Roughness
- 7) Visual Examination WSES-FSAR-UNIT-34.2-66Revision 13 (04/04)8)Ultrasonic Test9)Wall Thickness 10)Straightness 11)Inside Diameter 4.2.4.2.3 Fuel Rod Assembly(DRN02-1538, R12)Immediately prior to loading pellets must be capable of passing approved visual standards. Each fuel pellet stack is weighed to within 0.1 percent accuracy. The loading process is such that cleanliness and dryness of all internal fuel rod components are maintained until after the final end cap weld is completed. Loading
and handling of pellets is carefully controlled to minimize chipping of pellets.(DRN02-1538, R12)The following procedures are used during fabrication to assure that there are no axial gaps in fuel rods.
4.2.4.2.3.1 Stack Length Gage(DRN02-1538, R12)
The pellet stacks for Batches A through T were preassembled in "V" troughs that had been gauge marked to the proper length. They were then pushed into cladding tubes and the distance from the end of the tube to the end of the pellet stack checked with a gauge. The rods for Batches U and later are fabricated at the Columbia facility, which builds its pellet stacks directly in the cladding tubes. Their stacks are built up, 25
at a time, from a series of shorter preassembled segments that are fed into a like number of tubes by a vibratory feeder. Before feeding the last row of segments into the tubes, the distance from the end of the tube to the end of the pellet stack is checked with a gauge. If necessary, an appropriate number of pellets are added to or removed from each segment in the row. As before, the distance from the end of the tube to
the end of the pellet stack is then checked with a gauge.
4.2.4.2.3.2Rod Scanner Before being loaded into bundle assemblies, the finished fuel rods are gamma scanned to ensure that no gaps exist within them.
Loaded fuel rods are pressurized with helium to a prescribed pressure as determined for the fuel batch.Impurity content of the fill gas shall not exceed 0.5 percent.
In Batches A through T, the fuel rod upper end cap to cladding tube weld is a Magnetic Force (i.e., resistance) Weld whose outer surface is subsequently machined (i.e., deflashed). Beginning with Batch U, thejoint was converted to Tungsten Inert Gas (TIG) welding. The latter also utilizes a separate (TIG) seal weld to close the opening through which the rod is pressurized. Quality assurance on the end cap weld is
as follows:(DRN 04-502, R13)a)Non-destructive examination in accordance with approved procedures of all end cap welds (Batches U and later only) to certify bond length and to detect porosity or undercut.(DRN 04-502, R13)b)Visual examination of all end cap welds to establish freedom from cracks, seams, inclusions and foreign particles (Note: In Batches A through T, this examination was performed after final
machining of the weld).c)Destructive examination of a sufficient number of weld samples to establish that the allowable percent of unbonded wall thickness and the maximum allowable continuous unbonded region are satisfied.(DRN02-1538, R12)
WSES-FSAR-UNIT-34.2-67Revision 12 (10/02)(DRN 02-1538)d)Helium Leak checking of all end cap welds to establish that no leak rate greater than 10-8cc/sec is present.e)Corrosion testing of a sufficient number of samples to establish that weld zones do not exhibit excessive corrosion compared to a visual standard.(DRN 02-1538)All finished fuel rods are visually inspected to ensure a proper surface finish (scratches that measure greater than 0.001 in. depth, cracks, slivers and other similar defects are not acceptable).
Each fuel rod is marked to provide a means of identification.
4.2.4.3 Burnable Poison Rod 4.2.4.3.1 Burnable Poison Pellets B 4 C powder is sampled to verify particle size and w/o boron requirements prior to its use in pellet production. Finished pellets are 100 percent inspected for diameter and must satisfy a 90/90 confidence levelon other dimensions. Samples are taken from each of the pellet lots and examined for uniform dispersion of the B 4 C in A1 2 O 3. Conformance with density range requirements is demonstrated at a 95/95 confidence level and with B 4C loading requirements at a 90/90 level. Samples are drawn from each lot to verify acceptable impurity levels. Finally, all pellets are inspected for conformance with surface chip and
crack standards.
4.2.4.3.2 CladdingThe testing and inspection plan for burnable poison rod cladding is identical to that for fuel rod cladding (Subsection 4.2.4.2.2).4.2.4.3.3Bur nable Poison Rod AssemblyThe moisture content of poison pellets prior to loading is limited. The loading process is such that cleanliness and dryness of all internal p oison rod components are maintained until the final end cap weld is completed.
The following procedure is used during fabrication to assure that there are no axial gaps in poison rods:
The operator stacks pellets onto V troughs that are gage marked to the proper column height. When pellet stacking is completed, all column heights are checked by Quality Control. The pellets are subsequently loaded into tubes. After loading, the distance from the end of the tube to the end of the pellet column is
checked with a gage.Loaded poison rods are evacuated and backfilled with helium to a prescribed level. Impurity content of the fill gas must not exceed 0.5 percent.(DRN 02-1538)End cap weld integrity and corrosion resistance is ensured by a Quality Control plan identical to that used in fuel rod fabrication (Subsection 4.2.4.2.3).(DRN 02-1538)
WSES-FSAR-UNIT-34.2-68 Revision 12 (10/02)
All finished rods are visually inspected to ensure a proper surface finish (scratches greater than 0.001 in. in depth, cracks, slivers, and other similar defects are not acceptable).
4.2.4.4 Control Element AssembliesTheCEAs are subjected to numerous inspections and tests during manufacturing and after installation inthe reactor. A general product specification controls the fabrication, inspection, assembly, cleaning, packaging, and shipping of CEAS. All materials are procured to AMS, ASTM or C-E specifications. In
addition, various CEA hardware tests have been conducted or are in progress.
During manufacturing, the following inspections and tests are performed:(DRN 01-1103; 02-1477)a)The loading of each control element is carefully controlled to obtain the proper amounts and types of filler materials.(DRN 01-1103; 02-1477)b)All end cap welds are liquid penetrant examined and helium leak tested. A sampling plan is used to section and examine end cap welds.(DRN 02-1477)c)Deleted.d)Each CEA has unique serialization on the spider. See Figures 4.2-5.(DRN 02-1477)(DRN 00-644)e)Fully assembled CEAs are checked for proper alignment of the neutron absorber elements using aspecial fixture. The alignment check ensures that the frictional force that could result from adverse
tolerances is below the force which could significantly increase trip time.(DRN 00-644)
In addition to the basic measurements discussed above, the manufacturing process includes numerous other quality control steps for ensuring that the individual CEA components satisfy design requirements for
material quality, detail dimensions, and process control.
After installation in the reactor, but prior to criticality, each CEA is traversed through its full stroke andtripped. A similar procedure will also be conducted at refueling intervals.The integrity of each CEA will be tested at the beginning of the initial fuel cycle by performing a CEAsymmetry test as part of the low power physics testing. The CEA symmetry test will determine whether the
reactivities of symmetric CEAs are equal within the measurement limitation. The successful completion of these tests will demonstrate that no core loading or fabrication errors or loss in rod integrity exist that are sufficient to result in measurable CEA asymmetries. A CEA symmetry test will also be performed at the
beginning of each subsequent fuel cycle, as a minimum, on an abbreviated number of CEAS.
Hardware tests to date have been performed using CEA components developed primarily for CEs 800 MWeclass reactors which use 14 x 14 fuel assemblies.
WSES-FSAR-UNIT-3 4.2-69 Revision 309 (06/16)
(DRN 01-1103, R12; 02-1477, R12)
CEAs used in the Waterford 3 reactor are essentia lly similar in design and construction to the 800 MWe class CEA. The CEA spider arms are shorter and t he neutron absorber elements are smaller in diameter for compatibility with the 16 x 16 fuel assembly gui de tube dimensions employed in Waterford 3. (DRN 01-1103, R12) (LBDCR 15-039, R309)
Safety analyses assume the average CEA position is at least 90% inse rted at 3.2 seconds after trip breakers open. CEAs meet the 3.2 second aver age even under worst case conditions to reach 90 percent insertion in 3.2 seconds to agr ee with assumptions in Section 15.0.2. (DRN 02-1477, R12)
The reactivity worth of a CEA depends on the power (i.e., neutron flux) surrounding the CEA. During a reactor trip faster CEAs move in to higher flux regions sooner and, t hus, add more negative reactivity than slower CEAs. Note, CEAs do not nece ssarily fall with the same insertion times or at the same rate during a reactor trip. Therefore, the amount of negative reactivity inserted correlates to the average CEA insertion rate rather than the slowest CEA insert ion rate. This relation between CEA insertion and reactivity insertion is cycle independent if the mec hanical design, CEDM design, plus core physics and core thermohydraulics (pertinent to the CEAs) remain unchanged. (DRN 02-1477, R12)
CE performed three-dimensional space-time calc ulations with the NRC approved HERMITE computer program. The calculations adequately cover possibl e operating conditions and limits on the as-measured CEA distributions (Safety Evaluation Report for Amendment 58, dated October 31, 1989). The calculations show that for any reasonable dist ribution around an average CEA position during a trip, CEAs add negative reactivity at a rate directly rela ted to the average CEA position. Thus, Technical Specification limits should exist for the average CEA position. Thus, Te chnical Specification limits should exist for the average CEA drop time and Safety Analysis should assume that all CEAs fall in a "window shade" pattern with the average CEA drop time. Howe ver, if the time between the fastest and slowest CEA becomes too large, or the CEA distribution dev iates from the one modeled by CE, then the "window shade" may not necessarily represent the time dependent negative reactivity insertion. Therefore, besides the 3.2 second average insertion time limit, t he Technical Specifications limit the maximum drop time for the slowest CEA to 3.5 seconds. (DRN 02-1477, R12, LBDCR 15-039. R309)
WSES-FSAR-UNIT-34.2-70SECTION 4.2:REFERENCES1.Timoshenko, S., Strength of Materials, Part II Chapter IX, D. Van Nostrand Co., Inc., New York, 1956.2."High Temperature Properties of Zircaloy and UO 2 for use in LOCA Evaluation Models,"Combustion Engineering, Inc., CENPD-136 (Proprietary).3."Zircaloy Growth-In-Reactor Dimensional Changes in Zircaloy-4 Fuel Assemblies." CombustionEngineering, Inc., CENPD-198P (Proprietary), December 1975.4.O'Donnell, W.J., "Fracture of Cylindrical Fuel Rod Cladding due to Plastic Instability," WAPD-TM-651, April 1967.5.Weber, J.M., "Plastic Stability of Zr-2 Fuel Cladding, Effects of Radiation on Structural Metals,"ASTM STP 426, AM. Soc. Testing Mats., pp 653-669, 1967.6.Engle, J.T. and Meieran, H.B., "Performance of Fuel Rods Having 97 Percent Theoretical Density UO 2 Pellets Sheathed in Zircaloy-4 and Irradiated at Low Thermal Ratings," WAPD-TM-631, July 1968.7.Duncombe, E., Meyer, J.E., and Coffman, W.A., "Comparisons with Experiment of CalculatedDimensional Changes and Failure Analysis of Irradiated Bulk Oxide Fuel Test Rods Using the CYGRO-1 Computer Program," WAPD-TM-583, December 1966.8.McCauley, J.E., et al., "Evaluation of the Irradiation Performance of Zircaloy-4 Clad Test RodContaining Annular UO 2 Fuel Pellets (Rod 79-19)," WAPD-TM-595, December 1966.9.Notley, M.J.F., Bain, A.S., and Robertson, J.A.L., "The Longitudinal and Diametral Expansion of UO 2 Fuel Elements," AECL-2143, November 1964.10.Notley, M.J.F., "The Thermal Conductivity of Columnar Grains in Irradiated UO 2 Fuel Elements,"AECL-1822 July 1962.11.Manson, S.S., "Fatigue: A Complex Subject - Some Simple Approximations," ExperimentalMechanics, Vol. 22, No. 2, pp 193-226, July 1965.12.O'Donnell, W.J. and Langer, B.F., "Fatigue Design Basis for Zircaloy Components,'Nuc. Sci. Eng., Vol. 20, pp 1-12, 1964.13.CESSAR Proprietary Appendix, Docket 50-470.14. "C-E Fuel Evaluation Model Topical Report," Combustion Engineering, Inc., CENPD-139(Proprietary), CENPD-139 Rev. 01 (Non-Proprietary) CENPD-139 Supplement 1 (Proprietary),CENPD-139 Supplement 1, Rev. 01 (Non-Proprietary), July 1974.
WSES-FSAR-UNIT-34.2-71SECTION 4.2: REFERENCES (Cont'd)15.Conway, J.B., "The Thermal Expansion and Heat Capacity of UO 2 to 2200°C",GE-NMPD-TM-63-6-6.16.Christensen, J.A., "Thermal Expansion of UO 2", HW-75148, 1962.17.Jones, J.M., et. al., "Optical Properties of Uranium Oxides," Nature, 205, 663-65, 1965.18.Cabannes, F. and Stora, J.P., "Reflection and Emission Factors of UO 2 at High Temperatures",C.R. Acad. Sci., Paris, Ser. B. 264 (1) 45-48, 1967.19.Held, P.C. and Wilder, D.R., "High Temperature Hemispherical Spectral Emittance of UraniumDioxide at 0.65 and 0.70 m," J.Am. Cer. Soc., Vol. 52, No. 4, 1969.20.Brassfield, M.C., "Recommended Property and Reaction Kinetics Data for Use in Evaluating aLight Water Cooled Reactor Loss-of-Coolant Incident Involving Zircaloy-4 or 324-53 Clad UO 2 ,"GEMP-482, 1968.21."C-E Thermo-Structural Fuel Evaluation Method," Combustion Engineering, Inc., CENPD-179,April 1976.22."CEPAN, Method of Analyzing Creep Collapse of Oval Cladding," Combustion Engineering, Inc.CENPD-179, April 1976.23."STRIKIN-II, A Cylindrical Geometry Fuel Rod Heat Transfer Program," Combustion Engineering,Inc., CENPD-135P (Proprietary), CENPD-135 (Non-Proprietary), August 1974.24.Deverall, J.E., LA-2669 USAEC, Vol. 62 , 1954.25.Rudkin, R.L., Parker, J.W., and Jenkins, R.J., ASD-TDR-62-24, Vol. 1, p. 20, 1963.26.Thorne, R.P. and Howard, V.C. "Changes in Polycrystalline Alumina by Fast Neutron Irradiation,"p. 415, Proceedings, of the British Ceramic Society, No. 7, February 1967.27.Simnad, M.T. and Meyer, R.S., "BeO Review of Properties for Nuclear Reactor Applications,"Proceedings of the Conference on Nuclear Applications of Nonfissionable Ceramics, p 209-210,May 9-11, 1966.28.Rason, N.S. and Smith, A.W., "NAA-SR-862", Vol. 37 (AD85006), 1954.29.Saba, W.G. and Sterret, K.F., "J. Am. Chem. Soc." Vol. 79, pp 3637-38.30."Fuels and Materials Development Quarterly Progress Report," pp 38-58, ONRL-TM-3703,December 31, 1971.
WSES-FSAR-UNIT-34.2-72SECTION 4.2: REFERENCES (Cont'd)31.Kingery, W.D., "Introduction to Ceramics," John Wiley & Sons, pp 486-504.32.Toulookan, Y.S., "Thermophysical Properties of High Temperature Solid Materials," Vol. 4 and 5,MacMillan.33.Moore, G.E. and Kelley, K.K. , "J. Am. Chem. Soc.", Vol. 69, pp 309-16, 1947.34.Keilholtz, G.W. Moore, R.E., and Robitson, M.E., "Effects of Fast Neutrons on PolycrystallineAlumina and Other Electric Insulators at Temperatures From 60C-1230C" ORNL 4678, May 1971.35.Burian, R.J., Fromm, E.O., and Gates, J.E, "Effect of High Boron Burnups on B 4C and ZrBDispersions in A1 2 O 3 and Zircaloy-2," BM1-1627, April 24, 1963.36.Cunningham, G.W., "Compatibility of Metals and Ceramics, "Proceedings of Nuclear Applicationsof Nonfissionable Ceramics, pp 279-289, May 1966.37.Graber, M.J., "A Metallurgical Evaluation of Simulated BWR Emergency Core Cooling Tests,'Idaho Nuclear Corporation, IN-1453, March 1971.38.Pitner, A.L., "The WDC 1 Instrumental Irradiation of Boron Carbide in a Spectrum-HardenedETR Flux" HEDL-TME-73-38, April 1973.39.Gray, R.G. and Lynam, L.R., "Irradiation Behavior of Bulk B 4 C and B 4C-SiC Burnable PoisonPlates," WAPD-261, October 1963.40."HEDL Quarterly Technical Report for October, November, and December 1974," Vol. 1,HEDL-TME-74-4, pp A-51 to A-53, January 1975.41.Mahagan, D.E., "Boron Carbide Thermal Conductivity," HEDL-TME-73-78, September 1973.42.Homan, F.J., "Performance Modeling of Neutron Absorbers," Nuclear Technology, Vol. 16, pp216-225, October 1972.43.Pitner, A.L. and Russcher, G.E., Irradiation of Boron Carbide Pellets and Powders in HanfordThermal Reactors," WHAN-FR-24, December 1970.44.Pitner, A.L. and Russcher, G.E., "A Function of Predict LMFBR Helium Release Bound on BoronCarbide Irradiation Data from Thermal Reactors," HEDL-TME-71-127, September 30, 1971.45. HEDL-73-6, "Materials Technology Program Report for October, November, and December1973," pp A-69 to A-72.
WSES-FSAR-UNIT-34.2-73SECTION 4.2: REFERENCES (Cont'd)46.Cohen, I., "Development and Properties of Silver-Base Alloys as Control Rod Materials forPressurized Water Reactors," WAPD-214, December 1959.47.Tipton, C.R., "Reactor Handbook," Vol. 1, Materials, Interscience, p 827, 1960.48."National Alloy Development Program Information Meeting," pp 39-63, TC-291, May 22, 1975.49."Quarterly Progress Report - Irradiation Effects on Structural Materials," HEDL-TME-161, pp GE GE-10.50."Structural Analysis of Fuel Assemblies for Combined Seismic and Loss of Coolant AccidentLoadings," Combustion Engineering, Inc., CENPD-178, August 1976.51."Joint C-E/EPRI Fuel Performance Evaluation Program, Task C, Evaluation of Fuel RodPerformance on Maine-Yankee Core I," Combustion Engineering, Inc., CENPD-221, December 1975.52."Pressurized Water Reactor Project Period January 24, 1964 to April 23, 1964," WAPD-MRP-108.53."Fuel and Poison Rod Bowing," Combustion Engineering, Inc., CENPD-225-P (Proprietary),October 1976.54.Caye, T.E., "Saxton Plutonium Project, Quarterly Progress Report for the Period Ending March31, 1972," WCAP-3385-31, November 1972.55.Berman, R.M., Meieran, H.B., and Patterson, P., "Irradiation Behavior of Zircaloy-Clad Fuel RodsContaining Dished End UO 2 Pellets," (LWBR-LSBR Development Program), WAPD-TM-629, July 1967.56.Baroch, S.J., et al., "Comparative Performance of Zircaloy and Stainless Steel Clad Fuel RodsOperated to 10,000 MWd/MTU in the VBWR," GEAP-4849, April 1966.57.Megerth, F.H., Zircaloy-Clad UO 2 Fuel Rod Evaluation Program," Quarterly Progress Report No.8, August 1969-October 1969. GEAP-10121, November 1969.58.Megerth, F.H., "Zircaloy-Clad UO 2 Fuel Rod Evaluation Program," Quarterly Progress Report No.1, November 1967-January 1968, GEAP-5598, March 1968.59.Indian Point Nuclear Generating Unit No. 2, Preliminary Safety Analysis Report - Appendix A,Docket No. 50-247.60.Stiefel, J.T., Feinroth, H., and Oldham, G.M., "Shippingport Atomic Power Station OperatingExperience, Developments and Future Plans," WAPD-TM-390, April, 1963.
WSES-FSAR-UNIT-34.2-74SECTION 4.2: REFERENCES (Cont'd)61.Question V.B. 2, Prairie Island Nuclear Generating Plant, Preliminary Safety Analysis Report,Docket No. 50-306.62.Anderson, T.D., "Effects of High Burnup on Bulk UO 2, Fuel Elements," Nuclear Safety Vol. 6, No.2, Winter 1964-65 pp 164-169.63.Miller, R.S., et.al., "Operating Experience with the Saxton Reactor Partial Plutonium Core - II,"paper presented at AEC Plutonium Meeting in Phoenix, August, 1967.64.Megerth, F.H., "Zircaloy-Clad UO 2 Fuel Rod Evaluation Program," Quarterly Progress Report No.2, February 1968 April 1968, CEAP-5624, May, 1968.65.Blakely, J.P., "Action on Reactor and Other Projects Undergoing Regulatory Review ofConsideration," Nuclear Safety, Vol. 9, No. 4, p 326, July-August, 1968.66.Stephan, L.A., "The Response of Waterlogged UO 2 Fuel Rods to Power Bursts," IDO-ITR-105,April, 1969.67.Stephan, L.A., "The Effects of Cladding Material and Heat Treatment on the Response ofWaterlogged UO 2 Fuel Rods to Power Burst," IM-ITR-111, January, 1970.68."TORC Code: A Computer Code for Determining the Thermal Margin of a Reactor Core,"Combustion Engineering, Inc., CENPD-161-P, (Proprietary) July 1, 1975.69.Sandler, Y.L., "Structure of PWR Primary Corrosion Products," presented during NACECorrosion/78, March, 1978, Houston, Texas. Published in NACE-CORROSION, Vol. 35, No. 5, May, 1979.70.Lister, D.H., "The Accumulation of Radioactive Corrosion Products in Nuclear Steam Generators,"presented during NACE Corrosion/76, March, 1976, Houston, Texas.71.Yankee Core Evaluation Program,-Final Report, WCAP-3017-6094, 1971.72.Solomon, Y., Roesmer, T., "Measurement of Fuel Element Crud Deposits in Pressurized WaterReactors," Nuclear Technology, Vol 29, May, 1976, pp 166-173.73.Bessette, D.E., et al., CE/EPRI Fuel Performance Evaluation Program, RP-586-1, Task A,Examination of Calvert Cliffs I Test Fuel Assemblies at End of Cycles 1 and 2, September, 1978.74.Hillner, E., "Corrosion of Zirconium Base Alloys - An Overview," Zirconium in the Nuclear Industry,ASTM STP 633, pp.211-235, 1977.75. "Fuel and Poison Rod Bowing," Combustion Engineering, Inc., CENPD--225-P, Supplement 3-P(Proprietary), August 1979.
WSES-FSAR-UNIT-3 4.2-75 Revision 309 (06/16)
SECTION 4.2: REFERENCES (Cont'd)
- 76. J.A. Christensen, et. al., "Melting Point of Irradiated Uranium Dioxide," ANS Transactions, Volume 7:2, November 1964, p 390.
- 77. Final Safety Analysis Report, San Onofre Nu clear Generating Station Units 2 and 3, NRC Docket Nos. 50-361 and 50-362, Response to NRC Question 231.26.
- 78. CEN-382-P-A, "Methodology for Core Designs Containing Erbium Burnable Absorbers," ABB Combustion Engineering Nuclear Fuel, August 1993.
(DRN 03-1821, R13)
- 79. "Fuel Rod Maximum Allowable Gas Pressure," CEN-372-P-A, May 1990. (DRN 03-1821, R13)
(DRN 06-1059, R15)
- 80. CENPD-404-P-A, "Imp lementation of ZIRLO TM Material Cladding in CE Nuclear Power Fuel Assembly Designs," November 2001.
- 81. WCAP-16072-P-A, "Implementation of Zirconium Diboride Burnable Absorber Coatings in CE Nuclear Power Fuel Assembly Designs," August 2004. (DRN 06-1059, R15) (EC-9533, R302)
- 82. CEN-386-P-A, "Verification of the Acceptability of a 1-Pin Burnup limit of 60 MWD/kg for Combustion Engineering 16x16 PWR Fuel", August 1992.
- 83. WCAP-12610-P-A and CENPD-404-P-A Addendum 1-A, "Optimized ZIRLO TM", July 2006.
- 84. WCAP-16500-P-A, "CE 16x16 Next Generation Fuel Core Reference Report", August 2007. (EC-9533, R302) (LBDCR 15-035, R309)
- 85. WCAP-12610-P-A and CENPD-404-P-A Addendum 2-A, "Westinghouse Clad Corrosion Model for ZIRLO TM and Optimized ZIRLO TM ," October 2013. (LBDCR 15-035, R309)
WSES-FSAR-UNIT-3 TABLE 4.2-1 (Sheet 1 of 4) Revision 302 (12/08)
MECHANICAL DESIGN PARAMETERS Core Arrangement (EC-9533, R302)
NGF Number of fuel assemblies in core, total 217 (DRN 01-1103, R12)
Number of CEAs 87 (DRN 01-1103, R12)
Number of fuel rod locations 51,212
Spacing between fuel assemblies, fuel rod surface to surface, in. 0.208 0.216
Spacing, outer fuel rod surface to core shroud, in. 0.214 0.218
Hydraulic diameter, nominal channel, ft. 0.0394 0.0415
Total flow area (excluding guide tubes), ft 2 54.8 56.5
Total core area, ft 2 101.1
Core equivalent diameter, in. 136
Core circumscribed diameter, in. 143
Total fuel loading, Kg U 90 x 10 3 93x10 3 Total fuel weight, lbm. UO 2 224 x 10 3 234x10 3 Total weight of Zircaloy, lbm. 64,092 61,385
Fuel volume (including dishes), ft 3 356 359
Fuel Rod Array, square 16 x 16
Fuel Rod Pitch, in. 0.506 (EC-9533, R302)
WSES-FSAR-UNIT-3 TABLE 4.2-1 (Sheet 2 of 4) Revision 302 (12/08)
Fuel Assemblies (Cont'd) (EC-9533, R302)
Spacer Grid NGF Type - HID-1L Cantilever Spring Material Zircaloy-4 (DRN 02-1538, R12)
Number per assembly 11 (DRN 02-1538, R12)
Weight each, lb 1.7 (DRN 02-1538, R12; 06-1059, R15)
(DRN 02-1538, R12; 06-1059, R15)
Type Cantilever Spring Vertical Spring Material Inconel 625 Inconel-718 Number per assembly 1* 1 Weight each, lb 2.3 1.5
Type - Vaned Mid Grid I-Spring Material Optimized ZIRLO TM Number per Assembly 6
Weight, each, lb 2.8
Type - Unvaned Mid Grid I-Spring Material Optimized ZIRLO TM Number per Assembly 3
Weight, each, lb 2.7
Type - IFM Grid Co-planar Dimples Material Optimized ZIRLO TM Number per Assembly 2
Weight, each, lb 1.1
Type - Inconel Bottom Grid Cantilever Spring Cantilever Spring Material Inconel-625 Inconel-625 Number per Assembly 1 1
Weight, each, lb 2.6 2.3
Weight of fuel assembly, lbm. 1,435 1,416
Outside Dimensions Fuel rod to fuel rod, in. 7.972 x 7.972 7.96x7.964
Fuel Rod Fuel rod material (sintered pellet) UO 2 (DRN 06-1059, R15)
Pellet diameter, in., OD (annular ID) 0.325 (0.1625) 0.3225 (0.1550)
Pellet length, in., solid (annular) 0.390 (0.500) 0.387 (0.500)
Pellet density, g/cm 3 10.44 (DRN 06-1059, R15)
Pellet theoretical density, g/cm 3 10.96 (DRN 06-1059, R15)
Pellet density (% theoretical) 95.25
Stack density, g/cm 3, solid (annular) 10.11 (7.80) (10.31 (8.00) Clad material Zircaloy-4, ZIRLOTM Optimized ZIRLO TM (DRN 06-1059, R15)
Clad ID, in. 0.332 0.329 Clad OD, (nominal), in. 0.382 0.374 (DRN 06-1059, R15)
- some fuel assemblies in Batch U; all fuel assemblies beginning w/Batch W (DRN 06-1059, R15)
WSES-FSAR-UNIT-3 TABLE 4.2-1 (Sheet 3 of 4) Revision 302 (12/08)
Fuel Assemblies (Cont'd)
(EC-9533, R302)
NGF Clad thickness, (nominal), in. 0.025 0.0225
Diametral gap, (cold, nominal), in. 0.007 0.0065
Active length, in. 150 (DRN 02-1477, R12;04-502, R13)
Plenum length, in. 8.888 (Batch T) 10.013 9.138 (Batches U&W) (DRN 06-992, R15)
Uranium weight (nominal) grams 1830 1825 (DRN 04-502, R13;06-992, R15; EC-9533, R302)
Control Element (CEA)
(DRN 02-1477, R12)
(DRN 01-1103, R12)
Number 87
Absorber elements, No. per assy. 5
Type Cylindrical rods
Clad material Inconel 625
Clad thickness, in. 0.035
Clad OD, in. 0.816
Diametral gap, in. 0.009
Outside elements
Poison material B 4 C/Ag-In-CD
Poison length, in. 135.5/12.5
B 4 C Pellet Diameter, in. 0.737
Density, % of theoretical density
of 2.52 g/cm 3 73 Weight % boron, minimum 77.5
(DRN 01-1103, R12)
WSES-FSAR-UNIT-3TABLE 4.2-1 (Sheet 4 of 4)Burnable Poison RodAbsorber materialA1 2 0 3-B 4 CPellet diameter.307 Pellet length, min.1.000 Pellet density, (% theoretical), min.93Theoretical density, A1 2 0 3, g/cm 3 3.94Theoretical density, B 4C, g/cm 3 2.52Clad materialZircaloy-4 Clad ID, in.0.332 Clad OD, in.0.382 Clad thickness, (nominal), in.0.025 Diametral gap, (cold, nominal), in..025 Active length, in.136.0 Plenum length, in.11.090 WSES-FSAR-UNIT-3TABLE 4.2-2TENSILE TEST RESULTS ON IRRADIATEDSAXTON CORE III CLADDING (54)Fluence (1 MeV) 4.7 x 10 21 n/cm 2 (estimated)UniformStrainTotalLocationUltimateIn 2Strain FromTesting0.2% YieldTensilein. GageIn 2 in.Rod Bottom Temp StressStrength Length GageID(in.) (F)(psi x 103)(psi x 103) (%)LengthBO11-17 650 61.4 65.6 2.2 6.8 BO26-32 650 58.1 68.9 2.4 11.3 RD 3-9 650 62.2 70.0 2.0 4.2 RD12-18 650 60.5 65.4 1.7 5.8 MQ12-18 675 70.4 77.4 1.9 6.1 MQ28-34 675 66.0 75.1 1.6 6.2 FS28-34 675 57.2 71.4 3.9 12.9 GL12-18 675 60.5 71.5 2.4 9.3