ML20080H127
| ML20080H127 | |
| Person / Time | |
|---|---|
| Site: | Harris  |
| Issue date: | 09/01/1983 |
| From: | Rubenstein L Office of Nuclear Reactor Regulation |
| To: | Novak T Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML20079F427 | List:
|
| References | |
| FOIA-84-35 NUDOCS 8309210116 | |
| Download: ML20080H127 (69) | |
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SEP 1 1983 MEMORANDlM FOR:
T. M. Novak, Assistant Director s
for Licensing, DL FROM:
L. S. Rubenstein, Assistant Director for Core and Plant Systems, DSI
SUBJECT:
SHEARON HARRIS SAFETY EVALUATION REPORT 5
Plant Name:
Shearon Harris Units 1 & 2 Docket Numbers:
50-400, 50-401 Licensing Stage:
Operating License
. Responsible Branch:
Licensing Branch Number 3 Project Manager:
B. Buckley i
DSI Review Branch:'
Core Perfomance Branch Review Status:
Two Confimatory Issues in Section 4.2 Three Open Issues in Section 4.2 s
Two Open Issues in Section 4.4 One Open Issue in Section 4.3 The Core Perfomance Branch has prepared the enclosed SER input of Secticns 4.2, 4.3, 4.4, 15.4.1, 15.4.2, 15.4.3,15.4.7, and 15.4.8 on
' the Shearon Harris FSAR.
The confimatory issue and open issues are identified as follows:
Confirmatory Issue 1.
Resolution of the use of irradiation-strengthened Zircaloy yield strengths (see Paragraphs 4.2.1.1 and 4.2.3.1).
2.
Confimation that the predicted cladding collapse time exceeds the expected lifetime of the fuel (see Paragraph 4.2.3.2).
Open Issues 1.
A detemination that the fuel assembly mechanical response to seismic and LOCA forces meets the requirements of NUREG-0609, Appendix E (see Paragraph 4.2.3.3).
2.
A description of plans for on-line fuel system conitoring (see Paragraph 4.2.4.2).
3.
A description of plans for post-irradiation poolside surveillance of fuel (see Paragraph 4.2.4.3).
gsaana
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SEP 1.1997
-2 T. M. Novak.
The inconsistency between the peaking factor used in 4.
w resolved.
(Section 4.3)
The appli: ant must sutait the remainder of the documentation required by NUREG-0737 and the complete documentation must be reviewed and approved 5.
prior to the licensing of SHNPP.
(Section4.4)
The applicant must complete the evaluation of the SHHPP loose parts nonito system for conformance to Regulatory Guide 1.133 pr 6.
ui Original algned by
. S. Rubehshefn, k*sYshnt Director for Core and Plant Systems
/ Division of Systems, Integration,
]
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Enclosure:
As stated cc:: R. f'attson D. Eisenhut 800904600009 B. Buckley 1
T. Kadambi B. Sheron G. Knighton t
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. 4.2 Fuel Design The Shearon Harris fuel assembly described in the FSAR is a 17x17 array of fuel rods having a diameter of 0.374 in. This design will be referred to as the standard fuel assembly (SFA) in the pa'ragraphs below.
FSAR Section 4.2 presents the design bases for the SFA.
For the Westinghouse analysis, plant design conditions are divided into four categories of operation that are consistent with traditional industry classification (ANSI Standards N18.2-1973 and N-212-1974):
Condition I is Nomal Operation, Condition II,
' Incidents of Moderate Frequency, Condition III, Infrequent' Incidents; and Condition IV, Limiting Faults.
Fuel damage is related to these conditions of operation, which are coupled to the fuel design bases and design limits. The subsections of the design bases section address such topics such as (1) clad-ding, (2) fuel material (3) fuel rod perfomance, (4) spacer grids, (5) fuel assemblies, (6) reactivity control and burnable poisons (" core components"),
and (7)' testing, irradiation, and surveillance. As part of the discussion of the cladding design bases, material and mechanical properties, stress-strain
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limits, vibration and fatigue, and chemical properties are also presented.
A similar approach is taken for the other major su'btopics.
The staff review and safety evaluation follow SRP 4.2.
The objectives of this fuel system safety review are to provide assurance that (1) the fuel system is not damaged as a result of normal. operation and anticipated operational occur-rences, (2) fuel system damage is never so severe as to prevent control rod insertion when it is required. (3) the number of fuel rod failures is not underestimated for postulated accidents, and (4) coolability is always maintained.
"Not damaged" is defined as meaning that fuel rods do not fail, that fuel system dimensions remain within bperational tolerances, and that functional capabilities are not reduced below those assumed in the safety analysis. This objective implements GDC 10 and the design limits that
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accomplish this are called spe'cified acceptable fuel design limits (SAFDLs).
" Fuel rod failure".means that the fuel rod leaks and that the first fission 4
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p'roduct barrier (the cladding) has,' therefore, been breached.
Fuel rod failures must be accounted for in the dose analysis required by 10 CFR Part 100 for postulated accidents.
"Coolability," which is sometimes temed "coolable geometry," means, in general, that the fuel assembly retains its rod-bundle geometrical configuration with adequate coolant channeling to pemit removal of' residual heat after a severe accident.
The general requirements to maintain control rod insertability and core coolability appear repeatedly in the General Design Criteria (GDC 27, and 35).
Specific coolability requirements for the loss-of-coolant accidents are given in 10 CFR Part 50.46.
To meet the above-stated objectives of the fuel system review, the following specific areas are critically examined:
(a) design bases (b) description and design drawings, (c) design evaluation, and (d) testing, inspection, and sur-veillance plans.
In assessing the adequacy of the design, several items involving operating experience, prototyre testing, and analytical predictions ar's weighed in terms of specific acceptance criteria for fuel system damage, fuel rod failure, and fuel coolability.
Recently Westinghouse developed an improved fuel assembly design, which is described in WCAP-9500 and is called Optimized Fuel Assembly (OFA). WCAP-9500 was approved by,NRC (Rubenstein 1981).
The OFA design also consists of a 17x17 array of fuel rods having a diameter j
of 0.360 in., which is somewhat smaller than the standard. assembly.
Because the formst of WCAP-9500 followed Regulatory Guide 1.70, some of the fuel design bases and. design limits for the OFA were not presented in WCAP-9500 in a fom that pemitted cross-checking with the acceptable criteria provided in Section 4.2 of the SRP. Therefore several questions were issued (Rubenstein, August 8,1980) to clarify the design bases and limits.
Responses to those questions are contained in letters from Westinghouse (Anderson, August 15, 1980 and April 21,1981). These responses are applicable to the Standard Assembly as well (Petrick, September 9,1981).
Reference to these questions and answers will be made at several places in the review that follows.
4.2.1 Design Bases Design ~ bases for the safety analysis address fuel system damage mechanisms and suggest li,miting values for important parameters such that danage will be e
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For convenience, acceptance criteria for these design limits are grouped into three categories in the Standard Review Plan:
(a) fuel system damage criteria, which are most applicable to nomal operation (W plant Condition I), including anticipated operational occurrences (W plant Condition II), (b) fuel rod failure criteria, which apply to nomal operation (W plant Condition I), anticipated operational occurrences (W plant Condition II),
and accidents (W plant Conditions III End IV), and (c) fuel coolability criteria, which apply to accidents (W plant Conditions III and IV).
4.2.1.1 Fuel System Damage Criteria The following paragrap s discuss the NRC staff's evaluation of the design bases cnd carresponding design limits for the damage mechanisms Tisted in the SRP.
These design limits along with certain criteria that define failure (see Section 4.2.1.2 of this SER) constitute the SAFDLs required by GDC 10. The design limits in this section should not be exceeded during nomal operation inc'luding anticipated operational occurrences.
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(1)
Cladding Design Stress j
The design basis for fuel rod cladding stress as given in the response to Q 231.2* is that the fuel system will not be damaged due to excessive fuel rod cladding stresses.
The design limit for fuel rod cladding stress under Condition I and II modes of operation is that the volume-averaged effective stress calculated with the von Misas equation, considering interference due to unifom cylindrical pellet-to-cladding contact (caused by pellet themal expansion and swelling, unifom cladding creep, and fuel rod / coolant system pressure differences), is less than the Zircaloy 0.2 percent offset yield stress as affected by temperature and irradiation. This is a traditional l
limit consistent with previous Westinghouse design practice and is, therefore, acceptable without further comment except with respect to the credit that is taken by Westinghouse for irradiation-induced strengthening.
- All questions and responses referred to in this manner will be found in the correspondence cited above.
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The NRC does not routinely grant credit for irradiation-induced strengthening of the cladding, although the Standard Review Plan does not specifically precludei such practice. Moreover, in a response to an NRC question on the Westinghouse topical report on fuel material properties WCAP-9179, it was stated that no credit is taken for irradiation strengthening of Zircaloy. This restriction was subsequently a condition of approval (Thomas, September 29, 1982) in the WCAP-9179 SER.
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Because current staff understanding is that typical design cladding stresses under Condition I and II modes of operation are significantly below the 0.2 percent offset yield stress, the difference between irradiated and unirradiated values may be unimportant in this application.
However, we are unable to make that detemination. Hence, this issue is considered confimatory and the applicant should either justify the particular application 'of irradiated Zircaloy yield strength or discontinue use of this credit.
(2)
Cladding Design Strain With regard to cladding strain, a design limit for fuel rod cladding plastic tensile creep (due to unifom cladding creep and unifom cylindrical fuel
- /
. pellet swelling and themal expansion) of less than 1 per' cent from the unirra-diated condition is given in response to Q 231.2, Furthemore, the otal tensile strain transient limit (due to unifom cylindrical pellet themal expansion during the transient) is stated to be less than I percent from the pretransient value. While the staff has not explicitly reviewed the supporting i
data for nomal operation (Condition I), that value appears to be consistent with past practice (no numerical value for nomal operation cladding strain is provided a's an acceptance criterion in the Standard Review Plan), and thus there is reasonable assurance that I percent total plastic cre2p strain is an acceptable design limit for nomal operation, including Condition I power changes (load following).
For transient-induced defomation, the Standard Review Plan indicates that 1 percent unifom cladding strain is'an acceptable damage limit that should preclude some types of pellet / cladding interaction (PCI) failures. Such a limit, however, while consistent with past practice, should'no't be construed to be a broadly applicable PCI damage limit because there is ample evidence,(Tokar, November 14,1979) that PCI failures can occur
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' Westinghouse has indicated it.
at less than 1 percent uniform cladding strain.
response to staff question 231.24 that 1 percent plastic strain from the pretransient value'is not meent to serve as a broadly applicable PCI criterion.
Nevertheless, the staff finds the I percent cladding transient plastic strain criterion to be an acceptable design limit for the type of application indicated in SRP Section 4.2.
For fuel assembly structural design, Westinghouse set design limits on stresses and defomations due to various nonoperational, operational, and accident loads.
As indicated.in the FSAR, the stress categories and strength theory presented in Section III of the ASME Code are used as a general guide. This is consistent with acceptance criterion II.'A.1(a) of SRP Section 4.2 and is acceptable.
r, (3)
Strain Faticue The strain fatigue criteria given in response to Q 231.2 are the same as those described in SRP Section 4.2, viz., a safety factor of 2 on stress amplitude or i)f 20 on the number cycles and are, therefore, acceptable.
(4)
Fretting Wear
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j While the Standard Review Plan does not provide numerical bounding-value acceptance criteria for fretting wear,'it does stipulate that the allowable fretting wear should be stated in the safety analyis report and that the stress and fatigue limits should presume the existence of this wear.
l From the response to Q 231.5, it can be seen that the Westinghouse design basis for fretting wear is that fuel rods shall not fail during Condition I and II events.
Furthemore, Westinghouse does not use an explicit fretting wear limit in their stress and fatigue analysis for fuel rods.
- However, Westinghouse does use a value (proprietary) of wall thickness as a general guide in evaluating cladding imperfections, including fretting wear.
Cladding imperfections including fretting wear are thus considered in the stress and fatigue analysis, albeit in a very qualitative, nonrigorous manner.
In view of the apparently small effects of these defects and large stress and fatigue margins (see Section 4.2.3.1(4) of this Safety Evaluation Report), this design method is acceptable.
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e The design basis for guide thimble tubes is treated differently by Westinghouse, as described in the response to Q 231.41. The design basis is that the thinning of the guide thimble tube walls should not result in the failure of th-fuel assembly structural integrity or functio'nability of the guide thimble wbes.
The staff finds this to be an acceptable design basis.
With regard to a design limit for guide thimble tube wear, Westinghouse has
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detemined from stress analyses that the most limiting load on the fuel assembly struc'ture is that which might occur during a fuel handling accident.
For the analysis of thic accident, Westinghouse uses a design criterion of 6 g.
This design limit is therefore used for degraded guide thimble tubes and has been previously accepted for Westinghouse fuels.
(5)
Oxidation and Crud Buildup The SFA design basis for cladding oxidation and crud buildup is that the increase in cladding temperature due to cladding oxidation and crud buildup is not excessive (see Overheating of Cladding, below).
I Section 4.2 of the Standard Review Plan identifies cladding oxidation and crud l/
buildup as potential fuel system damage mechanisms.
Because of the increased themal resistance of these layers, th'ere is an increased potential for elevated temperature within the fuel as well as' the cladding.
Because the effect of oxidation and crud layers on fuel and cladding tenperature is a function of
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several different parameters (such as heat flux and ther:aal-hydraulic boundary, conditions), a design limit on oxide or crud layer thickness does not, per se, preclude fuel damage as a result inf these layers. Rather, it is necessary that these layers be appropriately considered in other temperature-related fuel system damage and failure analyses.
This is, indeed, the approach taken by Westinghouse in the design of the Standard Fuel Assembly.
Tne staff finds this approach acceptable.
l (6)
Rod Bowing Fuel rod. bowing is a phenomenon that alters the pitch dimensions between adjacent fuel rods.
Bowing affects local nuclear power per. king and the local l
a heat transfer to the coolant.
Rather than placing design limits on the amount of bowing that is permitted, the effects of bowing are included in the safety analysis. This is consistent with the Standard Review Plan and is acceptable.
The methods used for predicting the degree of rod bowing are evaluated in Section 4.2.3.1(6), and the impact of the resulting bow magnitude is evaluated in Sections 4.3 and 4.4.
(7)
Ayial Growth In the SFA design the core components requiring axial-dimensional analyses are l
the control rods, neut'ron source rods, burnable poison rods, fuel rods, and fuel assemblies (thimble plugging rods are omitted because,they are short and Yiot axial-growth limited). The axial growth of the first three of these components is primarily dependen't upon the behavier of poison, source, or spacer pellets and their 304 stainless-steel cladding. The growth of the latter two is mainly governed by 'the behavior of fuel pellets, Zircaloy-4 l
cladding, and Zircaloy-4 guide thimble tubes.
1 l-The Westinghouse design bases for core component rods are that (a) dimensional stability and cladding integrity are maintained during Condition I and II
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events and (b) these components do not interfere with shutdown during Condition III and IV events.
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Westinghouse does not, per se, have design limits on the axial growth of their control, source, and burnable poison rods.
However, allowances are nude to accommodate (a) pellet swelling due to gas production and (b) relative themal expansion between the stainless-steel ~ cladding and the encapsulated material.
Westinghouse does not account for irradiation growth of the stainless-steel cladding and has cited experiments (Foster and Strain, October 1974) as justifi-cation for the insignificance of irradiation growth of stainless-steel at PWR operating conditions.
For the Zircaloy cladding and fuel assembly components, the axial-dimensional j
behavior is governed by creep (due to mechanical or hydraulic loading) and l
irradiation growth. The critical tolerances that require controlling are l
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l (a) the spacing between the fuel rods and the fuel assembly (shoulder gap) and (b) the spacing between the fuel assenblies and the core internals.
Failure l
to adequately design for the.former may result..in fuel rod bowing, and for the latter may result in collapse of the holddown springs. With regard to inadequ-ately designed shoulder caps, problems have been reported (Schenk, October 1973; l
Kuffer and Lutz,1973; and FSAR of R. E. Ginna Unit 1,1972) in foreign i.
(Obrigheim and Beznau) and donestic (Ginna) plants that have necessitated pre-discharge modifications to fuel assenblies.
With regard to a design basis for shoulder gap spacing, Westinghouse stated in the responses to Q 231.2, 231.8, 231.25, and 231.40 that interference is precluded by having clearance between the fuel rod end and the top and bottom
-nozzles.
The design clearance accommodates the differences in growth,' fabrication tolerances, and the differences in themal expansion between the fuel cladding i
and the thimble tubes.
Westinghouse does not have specific limits on growth, but'does provide a gap spacing that is equal to or greater than a percentage (the specific value is proprietary) of the fuel rod length. The percentage value used by Westinghouse provides gap spacings that are similar to those employed in other fuel vendor designs.
j With regard.to fuel assembly growth, Westinghouse has a design basis that there shall be no axial interference between the fuel assenbly and upper and lower core plates caused by temperature or irradiation. As a design limit, Westinghouse provides a minimum gap (proprietary value that is a fraction of the fuel assembly length) between the fuel assembly and the reactor internals.
The above design bases and limits-dealing with axial growth are acceptable.
(8)
Fuel and Poison Rod Pressure For Condition I and II events, the mechanical design basis for core component rods described in the FSAR is that dimensional stability and cladding integrity are maintained. A necessary corollary cf this design basis is that the driving force, rod internal pressure, is never so great as to result in loss of dimensional stability and cladding integrity.
g Section 4.2 of the Standar'd Review Plan identifies rod internal pressure as a potential fuel system damage mechanism.
In this sense, damage is defined as an increased potential for elevated temperatures within the rod as well as an increased potential for cladding failure.
Although the Standard Review Plan mentions only fuel and burnable poison rods, the mechanism also applies to control rods, neutron source rods, and other core component rods.
Because rod internal pressure is a driving force for, rather than a direct mechanism of, fuel system damage, it is not necessary that a damage limit be specified.
It is only necessary that the phenomenon be appropriately. considered in other fuel system damage and fuel failure analyses.
In other words, rod internal pressure must be considered in calculating the temperature of the rod internals, cladding defonnation, and cladding bursting.
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I'n order to simplify the analysis of fuel system damage due to excessive rod internal pressure, the Standard Review Plan states that rod internal gas pressure should remain below the nominal system pressure during normal operation unless otherwise justified.
Westinghouse has elected to justify limits other than that provided in the Standard Review Plan.
For the fuel rods, revised internal rod pressure criteria as described in an approved topical report (WCAP-8963) were used in the FSAR.
Briefly stated.
these criteria allow the fuel rod internal pressure to exceed the system pressure under certain conditions:
(a) the internal pressure is limited such l
that the fuel-to-cladding gap does not increase during steady-state operation, and (b) extensive departure from nucleate boiling (DNB) propagation does not occur for postulated transients and accidents. These criteria have'been previously approved and remain acceptable.
For nonfueled rods, the rod internal pressure is limited such that the mechanical design limits, discussed in Section 4.2.1.5 of the FSAR, are not exceeded for Condition I and II events. This ' implies a stress limit of 2/3 of the material yield stress and a strain limit of 1 percent. These limits are unchanged from previously approved Westinghouse fuel' designs and remain acceptable for this FSAR.
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(9)
Assembly Liftoff The Standard Review Plan calls for the fuel assembly holddown capability (gravity and springs) to exceed worst-case hydraulic loads for nomal operation, which includes anticipated operational occurrences. The SFA design basis provides for positive holddown for Condition I, but allows momentary liftoff during one Condition II event. This design basis is acceptable provided that 4
it can be shown that the affected fuel assemblies will reseat properly without damage and without other adverse effects during the event. The ability of the affected fuel assemblies to satisfy this provision will be discussed in paragraph 4.2.3.1.
f10)
Control Material Leaching
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The Standard Review Plan and General Design Criteria requi that control rod reactivity be maintained.
Control rod reactivity can sometimes be lost by leaching of certain poison materials if the control rod cladding has been
. breached. The mechanical design basis for the control rods is stated in the FSAR to be consistent with the loading conditions of Section III of the ASME Code.
Thus, the design basis for the SFA control rods is to neintain cladding j
integrity; because cladding integrity would ensure that~ reactivity is maintained, this design basis might appear to be acceptable.
However, under some circum-stances, unexpected breaches might go undetected, so the staff does not nomal,ly l
accept control rod cladding integrity as a sufficient design basis. A discussion will be presented under Design Evaluation, paragraph 4.2.3.1, that shows that adequate surveillance will be provided to ensure maintenance of reactivity.
l 4.2.1.2 Fuel Rod Failure Criteria The NRC staff's evaluation of fuel rod failure threshold for the failure mechanisms listed in the SRP is presented in the following paragraphs.
When 1
these failure thresholds are applied to nomal or transient operation, they are med as limits (and hence SAFDLs), since fuel failures under those conditions should not occur (according to the traditional conservative interpretation of GDC 10).,. When these. thresholds are applied to accident analyses, the
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-s fuel failures must be detem.ed fer u.put to the radiological dose calculations required by 10 CFR 100. The basis or reason for establishing these failure t
thresholds is thus predetemined,.and only the threshold values are reviewed below.
t (1)
Internal Hydridino Hydriding as a cladding failure mechanism is precluded by controlling the I
level of moisture and other hydrogenous impurities during fabrication.
As described in the revised response to Q 231.6, the moisture levels in the i
uranium dioxide fuel are limited by Westinghouse to less than or equal to j
20 ppm. This specification is compatible with the ASTM specification for sintered uranium dioxide pellets, which allows 2p g hydrogen per gram of i
uranium (2 ppm), and they are the same as the limits provided in the Standard Review plan; they are therefore acceptable.
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l (2)
Claddino Collapse If axial gaps in the fuel pellet column were to
~ _ "- due to densification, the cladding would have the potential of collaps into a gap (flattening).
Because of the large local strains that would reru;t from collapse, the cladding is assumed to fail. As indicated in the FSAR and responses to Q 231.2, 231.9 and 231.34, it is a Westinghouse design basis that cladding collapse is precluded during the fuel rod design lifetime. This design basis is the same as that in l
the Standard Review Plan and is therefore acceptable.
(3)
.0verheatinc of Claddino The design basis as given in the FSAR for the prevention of fuel failures due to overheating is that there will be at least 95 percent probability that departure from nucleate boiling (DNB) will not occur on the limiting fuel rods during nonnal operation or any transient conditions arising fran faults of moderate frequency (Condition I and II events) at a 95 percent confidence level. This design basis is consistent with the thermal margin criterion of SRP Section 4.2 and is, thus, acceptable. The specific DNBR limits and methods of analysis are reviewed in Section 4.4.
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(4)
Overheating of Fuel Pellets As a second method of avoiding cladding failure due to overheating, Westing-house avoids centerline fuel pellet melting as a design basis. This design basis is the same as given in the Standard Review Plan and is thus acceptable.
The design limit corresponding to the design basis given above is that, during modes of operation associated with Condition I and Condition II events, there is at least a 95 percent probability that the peak kW/ft fuel rod will not 7,
exceed the UO2 melting tenperature. This design limit is an acceptable repre-sentation of the design basis given previously.
(5)
Pellet / Cladding Interaction As indicated in SRP Ser. tion 4.2, there are no generally applicable criteria for PCI failure.
However,. two acceptance criteria of limited application are presented in the SRP for PCI:
(a) less than 1 percent transient-induced clad-ding strain and (b) no centerline fuel melting. The response to Q 231.2 indicates that the I percent cladding plastic strain limit is met for the SFA design, and as stated in Section 4.2.1.2 of the FSAR, the SFA design ensures that U02 centerline melting will not occur through selection of a calculated fuel centerline temperature of 4700*F as an overpower limit. Thus the SFA design basis and limits agree with the only existing licensing criteria for PCI.
(6)
Cladding Rupture In the LOCA analysis for SFA-designed plants, an empirical model is used to predict the occurrence of cladding rupture. The failure temperature is expressed
-as a function of differential pressure across the cladding wall. There are no specific design limits associated with cladding rupture, and the rupture model is a portion of the ECCS evaluation model, uhich is documented in WCAP-8301 and WCAP-8302.
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4.2.1.3 Fuel Coolability Criteria For major accidents in which severe fuel damage might occur, core coolability must be maintained as required by several GDCs (e.g., GDC 27 and 35). The following paragraphs discuss the staff's evaluation of limits that will assure that coolability is maintained for the severe damage mechanisms listed in Section 4.2 of the SRP.
(1) ' Fragmentation of Embrittled Cladding For LOCA analysis. Westinghouse uses the acceptance criteria of 2200'F on peak cladding temperature and 17 percent on maximum cladding oxidation as prescribed by 10 CFR 50.46.
For events other than the LOCA, the NRC staff does not have separately estab-lished temperature or oxidation criteria. Yet it is clear that for short-term i
events such as locked rotor, the 2200'F peak cladding temperature and 17 percent oxidation LOCA criteria are not really meaningful, because the temperature history for such an event is much shorter than that of a LOCA. For events such as locked rotor, therefore, Westinghouse uses a unique peak-cladding-temperature (PCT) criterion of 2700'F.
The Westinghouse 2700*F PCT limit as selected taking into consideration the i
short time (a few seconds) that the fuel is calculated to be in DNB for a i
locked-rotor type event and the fact that the PCT and total metal-water reaction l
at the fuel hot spot would not be expected to impact fuel coolable geometry.
l While this linit has been used by Westinghouse for several years, the basis for the limit has only recently been reviewed. However, a recent assessment l
(Van Houten February 23,1981) of the available experimental infonnation indicates that fuel rod cladding will, indeed, retain its rod-like geometry after exposure to short-tem (a few seconds) peak cladding temperature of 2700*F. That conclusion is based on four Japanese nports (Shiozawa, March 1979; Hoshi, May 1980; JAERI-M-9011 Septenber 1980; and Fukishiro, October 1980) that describe experimental results for reactor test programs reported since 1979. The staff, therefore, concludes that there is reasonable assurance that the 2700'F PCT limit for short-tem events such as locked rotor is an acceptable coolability limit for the Westinghouse SFA design.
t It should be noted that staff acceptance of the 2700*F PCT limit for fuel rod coolability is currently restricted to undercooling events such as locked rotor. For overpower events such as control rod ejection, which involve a pellet-to-cladding mechanical interaction, the staff has not detemined the applicability of a PCT limit and currently uses a fuel rod enthalpy criterion of 280 cal /g for coolability of a rod-ejection accident.
(2)
Violent Expulsion of Fuel Material The design bases that there should be little or no possibility of fuel dispersal in the coolant, gross lattice distortion, or severe shock waves are given in Section 15.4.8.1.2 and are equivalent to those in the Standard Review Plan.
The design limits given in the FSAR are:
(a)
Average fuel pellet enthalpy at the hot spot will be below 225 cal /g for unirradiated fuel and 200 cal /g for irradiated fuel.
(b)
Average cladding temperature at the hot spot will be below the temperature
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at which cladding embrittlement may be expected (2700'F).
(c)
Peak reactor coolant pressure will be less than that which could cause pressures to exceed the faulted condition stress limits.
(d)
Fuel melting will be limited to less than 10 percent of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits above.
These limits are more conservative than the single 280 cal /g limit given in Regulatory Guide 1.77, they have been previously approved in the review of WCAP-7588, and they remain acceptable.
(3)
Cladding Ballooning and Flow Blockage i
In the LOCA analyses for SFA-designed plants, empirical models are used to predict the degree of cladding circumferential strain and assably flow blockage
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.2 at the time of hot-rod and hot-assembly burst. These models are each expressed as functions of differential pressure across the cladding well. There are no specific design limits associated with ballooning and blockage, and the balloon-ing and blockage models are portions of the ECCS evaluation model, which is documented in EAP-8301 and EAP-8302.
(4)
Structural Damage from External Forces Section 4.2.3.5 of the FSAR states that the fuel assembly will maintain a l
, geometry that is capable of being cooled under the worst case accident Condition IV event and that no interference between control rods and thimble tubes will occur during a safe shutdown earthquake. This is equivalent to the design basis as presented in the Standard Review Plan and is therefors
(
acceptable.
4.2.2 Description and Design Drawings l
l l
The description of fuel system components, including the fuel rods, bottom and top nozzles, guide and instrument thimble tubes, grid assemblies, rod cluster control assemblies, burnable poison rods, neutron sources, and thimble plugs, is contained in the FSAR Section 4.2.2.
In addition, in Table 4.3-1 numerical values are provided for various core component parameters. While each parameter listed in SRP subsection 4.2.2 is not provided in the FSAR, enough information is provided in sufficient detail to provide a reasonably accurate representation cf the SFA design and this information is thus acceptable.
4.2.3 Design Evaluation t
i Design bases and limits were presented and discussed in SER Section 4.2.1.
In this section we review y_ methods of demonstrating that the SFA fuel design l
meets the design criteria that have been established. This SER subsection will, therefore, correspond to subsection 4.2.1 of the SER point by point. The methods of demonstrating that the design criteria have been met include operating experience, prototype testing, and analytical predictions, i
4.2.3.1 Fuel System Damage Evaluation The following paragraphs discuss the NRC staff's evaluation of the ability of the SFA fuel to meet the fuel system damage criteria described in Section 4.2.1.1.
Those criteria apply only to nomal operation and anticipated trans-ients.
(1)
Cladding Desian Stress As indicated in the response to Q 231.2, Westinghouse used its Perfomance-Analysis and Design (PAD) code to analyze cladding stress (WCAP-8720). That code has been reviewed and found acceptable (Stolz, February 9,1979).
Typical calculated design values for cladding effective stress provided in response to Q 231.2 are stated to be considerably below the 0.2 percent offset field stress design limit. However, as discussed in Section 4.2.1.1(1), there is a discrepancy in that credit has been taken for irradiation-induced strengthening and this is contradictory to a restriction on the approved '
Westinghouse topical report WCAP-9179. Hence, this issue is considered confimatory.
(2)
Claddina Design Strain The NRC-approved Westinghouse fuel performance code (PAD) was used in the l
strain analysis, as indicated in the response to Q. 231.2. Typical design l
values of steady-state and transient creep strain, as calculated by that code, are found to be below the 1 percent strain criterion. Henct., the staff con-cludes that the SFA cladding strain design limits have been met.
l 1
(3)
Strain Fatique t
l-As indicated in the response to Q 231.2, Westinghouse used their approved PAD code for the strain range and strain fatigue life usage analysis.
Experimental l
data (proprietary) obtained from Westinghouse testing programs were used to l
derive the Westinghouse Zircaloy' fatigue design curve, according to the response to Q 231.4.
For a given strain range, the number of fatigue cycles is less l
m
~
.m s'.
than that required for failure, considering a minimum safety factor of 2 on stress amplitude or a minimum safety factor of 20 on the number of cycles, (the fatigue usage factor is less then 1.0). And the computations were per-formed with an approved code. Therefore, the staff concludes that the SFA fatigue design basis has been met.
(4)
Fretting Wear With regard to the Westinghouse fretting analysis of the fuel cladding, the staff concludes the following:
(a)
The out-of-pile flow tests and analyses (WCAP-9401) to detennine the magnitude of fretting wear that is anticipated for the OFA design have been previously reviewed and found acceptable (Rubenstein, April 23,1981).
These analyses are also acceptably conservative for SFA applications.
-(b)
LWR operating experience demonstrates that the number of fretting-induced fuel failures is insignificant.
(c)
There should be only a small dependence of cladding stresses on fretting wear because this type of wear is local at grid-contact locations and relatively shallow in depth.
(d)
The built-in conservatisms (that is, safety factors of 2 on the stress amplitudes and 20 on the number of cycles) in the strain fatigue analysis as well as the calculated margin to fatigue life limit adequately offset the effect of fretting wear degradation.
Therefore, the staff concludes that the SFA fuel rods will perform adequately with respect to fretting wear.
Fretting war has also been observed on the inner surfaces of guide thimble tubes where the fully withdrawn control rods reside.
Significant wear is limited to the relatively soft Zircaloy-4 guide thimble tubes because the Inconel or stainless steel control rod claddings are relatively wear-resistant.
The extent of the wear is both time-dependent and plant-dependent and has, in some non-Westinghouse cases, extended coupleteTy through the guide thimble i
tube wall.
r,,-
Westinghouse has predicted that an SFA can operate under a rod cluster control assembly (RCCA) for a period of time (proprietary) that exceeds the amount of rodded time expected with current 3-cycle fuel schemes before fretting wear degradation would result in exceeding the present margin to the 6 g load criterion for the fuel handling accident. However, the NRC required several applicants to perform a surveillance program because of the uncertainties in predicting wear rates for the standard 17x17 fuel assembly design. The objective of this program as to demonstrate that there as no occurrence of hole femation in rodded guide thimble tubes, thus providing some confidence i
that scranmability is ensured. These applicants fomed an owners' group, which has ' submitted a generic report (Leasburg, March 1,1982) that provides post-irradiation examination results on guide thimble tube wear in the. Westing-house 17x17 fuel assembly design. Based on this report, the staff has con-cluded (Rubenstein, April 19,1982) that the Westinghouse 17x17 fuel assembly design is resistant to guide thimble tube wear.
(5)
Oxidation and Crud Buildup In the FSAR, there is no explicit discussion of cladding oxidation, hydriding, and crud buildup. The applicable models for cladding oxidation and crud buildup are discussed in the supporting documentation (Salvatori, January 4, 1973) for the Westinghouse fuel perfomance code PAD-3.1. These models were previously approved by the NRC staff. A new temperature-dependent cladding oxidation model is also presented in WCAP-9179. Because the temperature-independent model in PAD-3.1 is conservative with respect to the approved model in WCAP-9179, the staff continues to find the older models applicable.
These models affect the cladding-to-coolant heat transfer coefficient and the temperature drop across the cladding wall. Mechanical properties and analyses
[
of the cladding are not significantly impacted by oxide and crud buildup.
On the basis of the Westinghouse discussion (Anderson, January 12,1981) of the impact of cladding hydriding on fuel perfomance, and on previous staff review of the oxidation and crud buildup models, the staff concludes that these effects have been adequately accounted for in the Standard Fuel Design.
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(6)
Rod Bowino The NRC has previously approved (Meyer, March 2,1978) the rod bowing correlation (Anderson, April 19,1978) that was used by the applicant.
(7)
Axial Growth Relative to the discussion above (4.2.1) on stainless steel growth, the staff
.is aware of supporting infomation (Bloom, April 1972, and Appleby, April 1972) that was not cited by Westinghouse, but which also implies that irradi-ation growth of stainless steel should not be significant at the temperatures and fluences that are associated with PWR operation. Furthemore, because the suff is unaware of any operating experience that indicates axial-growth-related i
problems in Westinghouse NSSS plants, the staff concludes that Westinghouse l
has made sufficient accomnodations for control, source, and burnable poison rod axial rod growth in their NSSS designs.
The Westinghouse analysis of shoulder gap spccing for the SFA has found that interference will not occur until achieving burnups beyond traditional values.
The staff, therefore, finds that the required shoulder gap spacing has been reasonably accommodated. However, for extended burnup applications, the adequacy of the spacing should be reverified.
Furthemore, because stress-free irradiation growth of zirconium-bearing alloys is sensitive to texture (preferred cystallographic orientation) and retained cold work, which, in turn, are strongly dependent on the specific fabrication techniques that are employed during component production, reverification of the design shoulder gap should be perfomed if Westinghouse current fabrication specifications are significantly altered.
Finally, the staff finds the destinghouse analysis of fuel assembly growth to l
be acceptable.
However, as stated in the above discussion on shoulder gap spacing, reverification of the fuel assembly growth should be perfomed if significant changes are made in the Westinghouse current fabrication techniques.
l
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(8)
Fuel and poison Rod Pressure The analysis of fuel rod internal pressure for the Standard Fuel Design is described in an approved topical report, WCAP-8963-A. The evaluation relies on the Westinghouse PAD-3.3 fuel performance code, which has also been approved (Stolz, February 9,1979) by the staff.
The analysis of nonfueled rod internal pressure for the SFA is generally based on Section III, Article NG-3000, of the ASME Code. Control rod, neutron source rod, and burnable poison rod cladding is 10 percent cold-worked 304 stainless steel, which is not covered by the Code. Westinghouse therefore defines as the stress limit an intensity value Sm equal to 2/3 of the material yield stress. The yield for this material occurs at about 62,000 psi. A strain limit of 1 percent also applies to the cladding. Predicted maximun values of rod internal pressure have been provided in an answer to NRC question Q 231.2 and they are well below those imposed by the cladding stress and strain limits.
lae staff concludes that there is adequate assurance that nonfueled core component rods can operate safety during Conditions I and II modes of operation even though maximum internal rod pressure may exceed system pressure because appropriate stress and strain limits are met.
(9)
Assembly liftoff In response to the staff's question on this topic, Westinghouse has confirmed that momentary liftoff will occur only during a turbine overspeed. Westing-house has further found that (a) proper reseating will occur after momentary liftoff, (b) damage to adjacent assemblies will not occur even if one assenbly is fully lifted and the adjacent ones remain seated, and (c) no ill con-sequences of momentary liftoff are expected. The staff concludes, therefore, that fuel assembly liftoff has been adequately addressed for the SFA design.
.u.
e e
(10) Control Material teachino While the design basis for the SFA control rods is to maintain cladding integrity, and while the probability of control rod cladding failures appears to be quite low, the staff has considered the corrosion behavior of the Shearon Harris control material and concludes that a breach in the cladding should not result in serious consequences because the Ag-In-Cd absorber material is relatively inert.
, 4.2.3.2 Fuel Rod Failure Evaluation l
The following paragraphs discuss the staff's evaluation of (a) the ability of the SFA fuel to operate without failure during nomal operation and antici-pated transients, and (b) the accounting for fuel rod failures in the applicant's accident analysis. The fuel rod failure criteria described in Section 4.2.1.2 were used for this evaluation.
(1)
Internal Hydridino l
Westinghouse has used moisture and hydrogen control limits in the manufacture of earlier fuel types and has found that typical end-of-life cladding hydrogen levels are less than 100 ppm -- a level below which hydride blister formation is not anticipated in fuel cladding.
The staff therefore concludes that reasonable evidence has been provided that hydriding as a fuel failure mechanism will not be significant in the SFA.
(2)
Claddino Collapse In calculating the time at which cladding collapse will occur, Westinghouse l
uses the generic methods described in WCAP-8377, which is approved (Stello, l
January 14,1975) for licensing applications.
Inputs to the analysis include cladding ovality, heliisa prepressurization, free volume of the fuel rod, and limiting power histories.
...x s.
s s Westinghouse adjusts the fuel rod pressure so that cladding collapse will not occur at a residence time that is less than the design lifetime.
Consequently, the staff expects that cladding collapse will not occur, but confimation should be provided '.ey the applicant by showing that the calculated cladding collapse time for Shearon Harris using WCAP-8377 methods is more than the expected lifetime of the fuel.
l (3)
Overheatino of Cladding As stated in SRP Section 4.2, adequate cooling is assumed to exist when the themal margin criterion to limit the departure from nuclear boiling (DNB) or boiling transition in the core is satisfied. The method employed to meet the i
DNB design basis is reviewed in Section 4.4 and will not be discussed here.
(4)
Overheating of Fuel Pellets The design evaluation of the fuel centerline melt limit is perfomed with the Westinghouse fuel perfomance code, PAD-3.3 (WCAP-8720). This code, which has been approved by the NRC (Stolz, February 9,1979), is also used to calculate initial conditions for transients and accidents described in Chapter 15 of the Standard Review Plan (see paragraph 4.2.3.3(1) below for further comments on PAD-3.3).
In applying the PAD-3.3 code to the centerline melting analysis, the melting temperature of the U0 is assumed to be 5081'F unirradiated and is decreased 2
by 58'F per 10,000 mwd /t. This relation has been almost universally adopted by the industry and has been accepted by the NRC staff in the past. The expressions for themal conductivity and gap conductance, described in Section 4.4.2.11 of the FSAR, are unchanged from that originally described in the PAD code. The staff considers it unnecessary to further review these models.
-In order to avoid using the PAD code to calculate a continuous set of burnup-dependent conditions necessary to cause centerline melt Westinghouse has I
perfomed the calculation for a single case. This was done by assuming a U0 2
_m._
melting temperature of 4701*F, which corresponds to the melting temperature at 65,000 mwd /t, and melting occurred at a linear power rating of approximately 21 kW/ft. The limiting local power for the worst Condition II transient, boron dilution with automatic rod control, is less than or equal to 18 kW/ft for Westinghouse plants with 17x17 fuel. Thus, the centerline melt criterion is satisfied in an acceptable manner.
(5)
Pellet / Cladding Interaction The only two PCI criteria in current use in licensing (1 percent cladding strain and no fuel melting), while not broadly applicable, are easily satisfied. As noted in the discussion of the clad' ding stress and strain evaluation, Westinghouse uses an approved code (PAD) to calculate creep strain, and the values calculated by that code are found to be below the 1 percent strain criterion. And, as indicated in the discussion on overheating failures, the no-centerline-melt criterion is satisfied based on an analysis (described in Chapter 15.4.6) of the boron dilution event, which is analyzed with an approved code. Therefore, the two existing licensing criteria for PCI have been satisfied.
In addition to the SRP-type treatment of PCI, however, responses to Q 231.23 andFSARSection4.2.3.3(a)addressPCIfromthestandpointofitseffecton fatigue life. Thus, PCI produces cyclic stresses and strains that can affect fatigue life of the cladding.
Furthemore, gradual compressive creep of the cladding onto the fuel pellet occurs due to the differential pressure exerted on the fuel rod by the coolant. Westinghouse contends that, by using prepres-surized fuel rods, the rate of cladding creep is reduced, thus delaying the time at which fuel-to-cladding contact first occurs. The staff agrees that fuel rod prepressurization should improve PCI resistance, albeit in a presently
~ 'unquantified amount.
In conclusion. Westinghouse has used approved methods to demonstrate that the present PCI acceptance criteria h' ave been met.
..a
..s (6)
Cladding Ruoture Although a revised cladding rupture temperature correlation has recently been approved (Miller. December 1.1981) as an integral part of the 1981 ECCS evaluation model, we previously concluded (NUREG-0630) that the old correlation in the earlier ECCS evaluation model us non-conservative over some regions of j
spplicability. To compensate for this deficiency, supplemental calculations l -
have been required for each plant application that uses the earlier Westinghouse l
ECCS evaluation model. Since the Shearon Harris analysis was done with this earlier model, supplemental calculations were provided to demonstrate that Shearon Harris would confonn to the ECCS acceptance criteria of 10 CFR 50.46 if the NRC staff cladding rupture temperature correlation (NUREG-0630) was substituted for the Westinghouse model contained in WCAP-8301.
l l
l This requirement for supplemental ECCS calculations is the same as the present requirement made for all operating license applications and all ECCS reanalyses j
of operating reactors (Eisenhut November 9,1979, and Denton Novmber 26,1979).
l The supplemental calculations for Shearon Harris are evaluated in Section l
4.2.3.3(3). The overall met of cladding rupture on the response of the SFA design to the loss-of-coc~
accident is evaluated in Section 15.6.5 and not reviewed further in this a:cion.
4.2.3.3 Fuel Coolability Evaluation The following paragraphs discuss the staff's evaluation of the ability of the SFA fuel to meet the fuel coolability criteria described in Section 4.2.1.3.
Those criteria apply to postulated accidents.
(1)
Fragmentation of Embrittled Cladding i
i The primary degrading effect of a significant degree of cladding oxidation is embrittlement of the cladding. Such embrittled cladding will have a reduced l
ductility and resistance to fragmentation. The most severe occurrence of such embrittlement is during a LOCA..The overall effects of cladding embrittlement on the SFA design for the loss-of-coolant accident are analyzed in Section 15.6.5 and are not reviewed further in this section.
g, One of the most significant analytical methods that is used to provide input to the analysis in Section 15.6.5 is the steady-state fuel perfomance code, which is reviewed in Section 4.2.
This code provides fuel pellet temperatures (stored energy) and fuel rod gas inventories for the ECCS evaluation model as prescribed by Appendix K to 10 CFR 50. The code accounts for fuel thermal conductivity, fuel densification, gap condu'ctance, fuel swelling, cladding creep, and other phenomena that affect the initial stored energy.
Westinghouse uses a relatively new fuel performance code called PAD-3.3 (WCAP-8720). This new Westinghouse code was approved with four restrictions as described in the staff's safety evaluation (Si.olz. February 9.1979).
Three of those restrictions dealt with numerical limits and have been met.
The fourth restriction related to the use of the PAD-3.3 code for the analysis of fission gas release from U0 for power-increasing conditions during nomal 2
operation This restriction applied to the SFA. However Westinghouse pre-pared and subnitted a detailed analysis (Anderson October 22,1979) of this restriction in an Addendtn to WCAP-8720. We have completed our review and issued (Rubenstein. June 30,1982) a safety evaluation of the Addendum.
In that evaluation, we concluded that the fourth restriction on the use of the PAD-3.3 code is unnecessary. As a result, we find the analysis described for the SFA acceptable as docketed for all cycles of operation.
At this time, the staff can therefore state that for the first cycle operation at full power, the restriction for PAD-3.3 is not significant and the analyses presented in the FSAR are acceptable. The staff anticipates completion of its review of the Westinghouse evaluation prior to the attainment of extended burnup at the Shearon Harris plant.
For non-LOCA events, the locked rotor accident (one-pump seizure with three loops operating) is the most severe undercooling event that is analyzed. This event is analyzed in Section 15.3.3 of the FSAR. where it is found that the peak cladding tenperature is 2250*F. which is well below the 2700*F design limit. The analysis of this event is reviewed in Section 15.3.3 of this report, but it is clear that the SFA meets the non-LOCA peak cladding temperature design limit.
m j
- l.sf (2)
Violent Expulsion of Fuel Material l
The analysis that demonstrates that the design limits we met for this event for the SFA is presented in Section 15.4.8 of the FSAR and is reviewed in that l'
section of the report.
(3)
Cladding Ballooning and Flow Blockage Revised cladding rupture temperature, strain, and assenbly flow blochge correlations have recently been approved (Miller December 1,1981) as integral parts of the 1981 ECCS evaluation model. We had previously concluded (NUREG-0630) that these correlations were non-conservative over some regions of applicability in the earlier ECCS evaluation model. To compensate for these deficiencies, supplemental calculations have been required for each plant application that uses the earlier Westinghouse ECCS evaluation model.
Since the Shearon Harris analysis was done with this earlier model, supple-mental calculations were provided to demonstrate that Shearon Harris would conform to the ECCS acceptance criteria of 10 CFR 50.46 if the NRC staff cladding rupture temperature, strain, and assenbly flow blockage correlations (NUREG-0530) were substituted for the Westinghouse models contained in WCAP-8301 and WCAP-8302.
l The applicant's calculation responses to Q 490.2 and 440.101 also accounted for a non-conservatism identified ( Anderson, November 16,1979) by Westing-house in their February, 1978 ECCS evaluation model, which used a fast-heatup-rate rupture-temperature correlation for slow transient analyses. Based on an Fg of 2.11 the applicant's submittal assessed the combined impact of tnis calculational error and the NUREG-0630 correlations to be worth 855'F peak l
cladding temperature above that previously calculated.
l l
l Subsequently, Westinghouse calculated that a reduction in total peaking factor Fg of 0.157 would offset the portion of the 855'F increase in peak cladding tempeature that exceeded 2200*F.. However Westinghouse also identified a
. margin in F available through the use of UHI themohydraulic model improvements g
that are generically approved for the Shearon Harris type of 3-loop plant.
This margin is worth 0.15 in F.
g
- o':
Consequently, an F reduction of only 0.01 is required for Shearon Harris, and q
the applicant's Technical Specifications should reflect a new F of 2.10. We, q
therefore, concluded that the applicant has satisfied our concerns related to the swelling and rupture issue.
The overall impact of cladding ballooning and assembly flow blockage models on the responses of the SFA design to the loss-of-coolant accident is evaluated in Section 15.6.5 and is not reviewed further in this section.
(4)
Structural Demace from External Forces 1
Section 4.2.3.5 of the FSAR refers to WCAP-8236 for this analysis. The staff has reviewed and approved another report (WCAP-9401) which essentially augments the information presented in WCAP-8236 because both WCAP reports apply to similar assemblies.
For.he Shearon Harris application, however, the applicant must demonstrate compliance with Appendix A of SRP Section 4.2.
The applicant may make reference to WCAP-8236 and WCAP-9401 to accomplish this.
4.2.4 Testing, Inspection, and Surveillance Plans 4.2.4.1 Testing and Inspection of New Fuel As required by SRP Section 4.2, testing and inspection plans for new fuel should include verification of significant fuel design parameters. While details of the nanufacturer's testing and inspection programs should be docu-mented in quality control reports, the programs for onsite inspection of ned fuel and control assemblies after they have been delivered to the plant shoald also be described in the FSAR.
The Shearon Harris FSAR discussion of the Westinghouse quality control program addresses fuel system components and parts, pellets, rod inspection, assemblies, process control, and so forth. Fuel system component inspection depends on
_ the component parts and includes dimensions, visual appearance, audits of test
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l reports, material certification, and nondestructive examinations.
Pellet i
inspections, for example, are perfonned for dimensional characteristics such as diameter, density, length and squareness of ends. Fuel rod, control rod.
burnable poison, and source rod inspection reportedly consist of nondestructive' examination techniques such as leak testing, weld inspection, and dimensional measurements.
Process control procedures are described in detail.
In addition, Westinghouse states in FSAR Section 4.2.4.4 that if any tests and inspections are to be perfonned by others on behalf of Westinghouse Westinghouse will review and approve the quality control procedures, inspection plans, and so forth, to ensure that they are equivalent to the description provided in WCAP-9500 and are perfonned properly to meet all Westinghouse requirements.
The staff concludes, based on the infonnation provided in FSAR Section 4.2.4 and che commitment by Westinghouse to-ensure the acceptability of any tests and inspections perfonned by others on behalf of Westinghouse, that the fuel testing and inspection program for new fuel is acceptable.
4.2.4.2 Online Fuel Failure Monitoring The applicant should provide on-line fuel rod failure detection methods.
These are needed to' satisfy the guidelines described in paragraph II.D.2 of the SRP.
4.2.4.3 Post-irradiation Surveillance A commitment to do visual examination of some discharged fuel assemblies from each refueling has not been made by the applicant. This is needed to satisfy the guidelines described in paragraph II.D.3 of the SRP regarding the need for post-irradiation surveillance.
4.2.5 Evaluation Findings The following have not yet been provided for by the applicant.
(1)
Resolution of the use of irradiation strengthened Zircaloy yield strengths (see Paragraphs 4.2.1.1(1) and 4.2.3.1(1)).
~
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. 5, (2)
Confirmation that the predicted cladding collapse time exceeds the expected lifetime of the fuel (see Paragraph 4.2.3.2(2)).
(3)
A determination that the fuel assembly mechanical response to seismic and LOCA forces meets the guidelines of Appendix A of SRP Section 4.2 (see paragraph 4.2.3.3(4))
(4)
. A description of plans for on-line fuel system monitoring (see paragraph
'4.2.4.2).
(5)
A description of plans for post-irradiation poolside surveillance of fuel (see paragraph 4.2.4.3).
When the above are provided, the staff will conclude that the Shearon Harris fuel has been designed so that (a) the fuel system will not be damaged as a result of normal operation and anticipated operational occurrences, (b) fuel damage during postulated accidents would not be severe enough to prevent control rod insertion when it is required, and (c) core coolability will always be maintained, even after severe postulated accidents, and thereby meets the related requirements of 10 CFR Part 50.46; 10 CFR Part 50 Appendix A; GDC 10, 27, and 35; 10 CFR Part 50, Appendix K; and 10 CFR Part 100. This conclusion is based on the following:
(1)
The applicant has provided sufficient evidence that these design objectives will be met based on operating experience, prototype testing, and analytical predictions. Those analytical predictions dealing with structural response, control rod ejection, and fuel densification have been performed in accordance with (1) the guidelines of Regulatory Guide 1.77, and methods that the staff has reviewed and found to be acceptable alternatives to Regulatory Guides 1.60 and 1.126, and (b) the guidelines for " Evaluation of Fuel Assembly Structural Response to Externally Applied Forces" in Appendix A to SRP Section 4.2.
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(2)
The applicant has provided for testing and inspection of the fuel to ensure that it is within design tolerances at the time of core loadings. The applicant has made a commitment to perform on-line fuel failure monitoring and post-irradiation surve111ence to detect anomalies or confirm that the fuel has perfomed as expected.
The staff concludes that the applicant has described methods of adequately predicting fuel rod failures during postulated accidents so that radio-activity released are not underestimated and thereby meets the related requirements of 10 CFR Part 100.
In meeting these requirements, the applicant has (a) used the fission product release assumptions of Regulatory Guides 1.4,1.25, and 1.77, and (b) perfomed the analysis for fuel rod failures for the rod ejection accident in accordance with the guidelines l
dn the basis of its review, the staff concludes, that the applicant's fuel l
)
system design has met all the requirements of the applicable regulations, regulatory guides, and current regulatory positions.
l l
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l 4.3 Nuclear Desian The Shearon Harris Units 1 and 2 power plants have a reactor core consisting of 157 fuel assemblies of the Westinghouse 17X17 design. The core has a design heat output of 2775 thennal Megawatts and is similar to the Virgil C. Summer reactor and other recent Westinghouse 3 loop reactors. We i
have reviewed the nuclear design of the Shearon Harris thits 1 and 2 reactors.
x Our review us conducted in accordance with the guidelines provided by the 3
Standard Review Plan, Section 4.3.
- 4. 3.1 Desian Bases Design bases are presented which comply with the applicable General Design Criteria. Acceptable fuel design limits are specified (GDC 10), a negative prorot feedback coefficient is specified (GDC 11) and tendency toward divergent operation (power oscillation) is not pemitted (GDC 12).
Design bases a;e presented which require a control and monitoring system (GDC 13) which automatically initiates a rapid reactivity ' insertion to prevent exceeding fuel design limits in normal operation or anticipated transients i
(GDC20). The control system is required to be designed so that a single malfunction or single operator error will cause no violation of fuel design limits (GDC25). A reactor coolant boration system is provided which is capable of bringing the reactor to cold shutdown conditions (GDC 26) and the control system is required to control reactivity changes during accident conditions when combined with the engineered safety features (GDC 27).
l Reactivity accident conditions are required to be limited so that no damage to the reactor coolant system boundary occurs (GDC 28).
l We find the design bases presented in the FSAR to be acceptable.
4.3.2 Design Description The FSAR contains the description of the first cycle fuel loading which consists of three different enrichments and has a first cycle length of approximately one year. The enrichment distribution, burnable poison
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distribution, soluble poison concentration and higher isotope (actinide) content as a function of core exposure are presented. Values presented for the delayed neutron fraction and prompt neutron life-time at beginning and end of cycle are consistent with those normally used and are acceptable.
power Distribution The design bases affecting power distribution are:
'The generic design peaking factor fnr the reactor is 2.32. However, at present, the peaking factor in the core will not be greater than 2.11 during nomal operation of full power in order to meet the initial conditions assumed in the loss of coolant accident analysis.
Under nomal conditions (including maximum overpower) the peak fuel power will not produce fuel centerline melting.
The core will not operate during normal operation or anticipated-operational occurrences, with a power distribution that will cause the departure from nucleate boiling ratio to fall below 1.3 (W-3 correlation with modified spacer factor).
The first part of the following discussion assumes the Westinghouse generic design total peaking factor, F, of 2.32 provides the (only) required g
limiting peak power density.
This is, at present, the only value of F dis-O cussed in Section 4.3 of the FSAR, and the methodology and calculations and surveillance used to demonstrate limiting conditions of operation are appropriate only to that value of'2.32 (for first cycle). The F limit of g
2.11 will be discussed following that for 2.32.
The 2.32 Fg eaking factor is determined and maintained via calculations p
of extremes of allowed transient power distributions and periodically measured radial power distributions and radial peaking factors F and F xy AH*
These also provide maximum initial conditions for events described in 4.3 - 2
m.
..,s Section 15 which assure that peak fuel power does not cause center line fuel melting or result in departure from nucleate boiling during anticipated operational occurrences.
The applicant has described the manner in which the core will be operated and l
power distribution monitored so as to assure that these limits are met. The core will be operated in the Constant Axial Offset Control (CADC) mode which has been shown to result in peaking factors less than 2.32 for both constant power and load following operation. The applicant has elected to use an im-proved load follow package, developed by Westinghouse, in Shearon Harris i
Units 1 and 2.
1 CAOC is described in WCAP-8385 (Proprietary) and WCAP-8403 (non-Proprietary),
" Power Distribution Control and load Following Procedures." This report contains methodology for operation with and without part length control rods.
The former mode allows better return to power capability than the latter. Use of part length rods has been withdrawn from Westinghouse reactors. The improved load follow strategy provides a return to power capability during
(
operation without part length rods comparable to the level previously obtain-able from operation with part length rods.
The improved load follow strategy involves a redesigned control rod bank and modif%d overlap that allows greater reactivity insertion than the former design bank within the constraints of a widened, asymmetric CAOC band. The control bank has been changed from eight to four rods. The four rods removed from the control bank have been redesigned as a shutdown bank, thus maintaining shutdown margins.
(There are also an extra four rods assigned to a shutdown l
bank, compared to other Westinghouse three loop reactors.) The CAOC band has been changed from 25 to +3.-12, AI (delta flux difference). The greater inserted reactivity is available for return to power capability upon control red with-drawal. Another element in the load follow strategy is the use of moderator temperature reductions to augment return to power capability. The temperature reduction adds reactivity during rapid return to pcwer through the inherently negative moderator temperature coefficient.
l l
i 4.3 - 3
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The analysis used to calculate the maximum peaking factor which can occur y
using the improved strategy expands the set in the CAOC topical report to 18 calculational cases. However, with the redesigned control bank, maneuvers resulting in greater control rod insertion for a longer duration become operationally practical but tend to become slightly more limiting in tems of total peaking factors. Therefore, simulated load follow maneuvers which return AI to the target value (and thereby reduce control rod insertion) have been replaced by load follow strategies which maintain the deeper rod insertion. 3 As a result of our evaluation, we agree with Westinghouse's conclusion that substitution of these more conservative cases will naintain the limiting nature of the 18 case load following analysis.
.1 The analysis perfomed by Westinghouse indicated that the peaking factor limit could not be met at BOL of Cycle I due to the wide AI band. This resulted in y
1imiting the width of the band for the first 20% of the cycle typically, and until 3,000 MWD /MTU burnup for Shearon Harris Units I and 2 to the value of 5% AI. This AM.3I is the value previously justified by the CAOC analysis, p
These features will be incorporated in the Shearon Harris Technical Specific-h ations.
We conclude, for the reasons stated above, that the improved load follow l
package will continue to prevent the 2.32 peaking factor limit from being ex-g ceeded in nomal operation of the power plant, and therefore is acceptable.
,,,3
.e Two types of instrumentation systems are nomally provided to monitor core l
y power distribution. Excore detectors with two (axial) sections are used to monitor core power, axial offset and azimuthal tilt for the 2.32 F limit, g
and movable incore detectors pemit detailed power distributions to be measured. These systems are used in operating reactors supplied by Westinghouse and we find their use acceptable for Shearon Harris when a 2.32 limit is the minimum requirenent (or possibly lower when cycle specific 9
18 case analyses so indicate).
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4.3 - 4 l
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The 2.32 peaking factor is an acceptable limit for all events considered in Section 15 with the exception of LOCA.
For that event the analysis for Shearon Harris, as described in Section 15.6, uses a 2.11 value to maintain a clad temperature of less than 2200*F. Since this is less than the generic 2.32 value, some additional or altered means of power distribution analysis or surveillance must be used to assure compliance with the 2.11 limit. The generic 2.32 value is essentially never reached in reactor-cycle specific normal operation situation, and lower limiting values can usually be detemined and us'ed in conjunction with cycle specific 18 case analyses or with improved surveillance.
In response to questions in this area Shearon Harris has indicated that it intends to provide an improved surveillance system, the Axial Power Distribution Monitoring System (APDMS). This uses a set of four section (4xial) excore neutron detectors rather than the usual two section detectors.
This provides a better measurement of the axial power distribution than is available from the usual axial offset correlation limit.
The APDMS mode of rurveillance, using an incore instrume:
stem, has been previously approved by the staff and used in other Westi.ua 3e reactors.
The excore four section detector and axial distribution monitoring has been described in a topical report (WCAP-9105) in connection with the Westinghouse 414 reactor.
It has not as yet been used in any domestic operating reactor (otherthantests). The review of WCAP-9105 indicated that it was acceptable for describing the system and techniques, but that further material was needed to provide associated uncertainty levels. Shearon Harris has indicated that a topical report will be submitted on the subject. Acceptance of the system will depend on the results of the review of that report.
Until the APDMS, or some alternate system or analysis methodology and result is accepted, the issue of the peaking factor to be used for LOCA calculations remains open. Until the issue is settled the reactor would be limited to about 91 percent power.
4.3 - 5
m.
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=
Reactivity Coefficients The reactivity coefficients are expressions of the effect on core reactivity of changes in such core conditions as power, fuel and moderator temperature, moderator density, and boron concentration. These coefficients vary with fuel burnup and power level. The applicant has presented values of the coefficients in the FSAR and has evaluated the uncertainties of these values. We have reviewed the calculated values of reactivity coefficients and have concluded that they adequately represent the full range of expected values. We have
[_
reviewed the reactivity coefficients used in the transient and accident analyses and conclude that they conservatively bound the expected values, including uncertainties.
Further, moderator and power Doppler coefficients along with boron worth are measured as part of the startup physics testing to assure that actual values are within those used in these analyses.
Contro1 To allow for changes in reactivity due to reactor heatup, load following, and l
fuel burnup with consequent fission product buildup, a significant amount of excess reactivity is built into the core. The excess reactivity is controlled i
by a combination of full length control rods and soluble boron. Soluble boron is used to control changes due to:
Moderator density and temperature changes from ambient to operating temperatures.
Equilibrium xenon and samarium buildup.
Fuel depletion and fission product buildup - that portion not controled by lumped burnable poison.
Transient xenon resulting from load following.
l b
4.3 - 6 w - - -.
Control rods are used to control reactivity change due to:
Moderator reactivity changes from hot zero to full power.
i Fuel temperature changes (Doppler reactivity changes).
l l
Burnable poison rods placed in some fuel assemblies are used for radial flux shaping and to control part of the reactivity change due to fuel depletion and fission product buildup.
- h The applicant has provided data to show that adequate control exists to satisfy the above requirements with enough additional control rod worth to provide a hot shutdown effective neitiplication factor less than the design basis value of 0.9823 during initial and equilibrim fuel cycles with the most reactive control rod stuck out of the core.
In addition, the chemical and volume control system will be capable of shutting down the reactor by adding soluble boron and maintaining it shut down in the cold, xenon free condition at any time in core life. These two systems satisfy the requirements of Gener31 Design Criterion 26.
Comparisons have been made between calculated ani measured control rod bank worth in operating reactors and in critical experiments. These comparisons lead to the conclusion that bank worths may be calculated to within approxi-mately ten percent.
In addition bank worth measurements are perfomed as oart of the startup test program to assure that conservative values have been used in safety analyses.
Bared on these comparisons, we conclude that the applicant has made suitably conservative assessments of reactivity control requirements and that adequate control rod worths have been provided to assure shutdown capability.
4.3 - 7
ll-i 8-Control Rod Patterns and Reactivity Worths The control rods are divided into two categories - shutdown rods and regulating rods. The shutdown rods are always completely out of the core when the reactor is at operating conditions. Core power changes are made with regulating rods which are nearly out of the core when it is operating at full power. Regulating rod insertion will be controlled by power-dependent insertion limits required in the Technical Specifications to assure that:
There is sufficient negative reactivity available to pemit rapid shutdown of the reactor with adequate margin.
The worth of a control rod that might be ejected is not greater than that which has been shown to have acceptable consequences in the safety analyses.
We have reviewed the calculated rod worths and the uncertainties in these worths, and conclude that rapid shutdown capability exists at all times in core life assuming the most reactive control rod asseribly is stuckout of the core.
Stability The stability of the Shearon Harris Units 1 and 2 cores to xenon induced spatial oscillations is discussed in the FSAR. The overall negative reactivity (power) coefficient provides assurance that the reactor will be stable against total power oscillation. The applicant also concluded that sustained radial or azimuthal xenon oscillations are not possible. This conclusion is based on measurements on an operating reactor of the same dimensions which showed stability against these oscillations. We concur with this conclusion.
This core is predicted to be unstable with respect to axial xenon oscillations after about 12000 Megawatt days per ton of exposure. The spp11 cant has acceptably shown that axial oscillations may be controlled by the regulating rods to prevent reaching any fuel damage limits.
4.3 - 8
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, i, Criticality of Fuel Assemblies Criticality of fuel assemblies outside the reactor is precluded by adequate j
design of fuel transfer and storage facilities. The applicant presents information on calculational techniques and assumptions used to assure that criticality is avoided. We have reviewed this information and the criteria which will be employed and find them to be acceptable.
Vessel Irradiation Values are presented for the neutron flux in various energy ranges at mid-height of the pressure vessel inner boundary. Core flux shapes calculated by standard design methods are input to a transport theory calculation (Sn) which 10 results in a neutron flux of 2.9 x 10 neutrons per square centimeter per 6
second having energy greater than 10 electron-volts at the inner vessel boundary. This results in a fluence of 2.9 x 1019 neutrons per square centi-meter for a forty year vessel life with an 80 percent use factor. The methods used for these calculations are state of the art, and we conclude that accept-able analytical procedures have been used to calculate the vessel fluence.
The Materials Engineering Branch will review the requirements for surveillance progtams and the pressure-ten.perature limits for operation.
4.3.3 Analytical Methods The applicant has described the computer programs and calculational techniques used to obtain the nuclear characteristics of the reactor design. The calcul-ations consist of three distinct types, which are performed in sequence:
detemination of effective fuel temperatures, generation of macroscopic few-group parameters, and space-dependent few-group diffusion calculations. The programs used (e.g., LASER, TWINKLE, LEOPARD TURTLE and PANDA) have been applied as part of the applications for most earlier Westinghouse designed nuclear plant facilities and the predicted results have been compared with measured characteristics obtained during many startup tests for first cycle and reload cores. These results have validated the ability of these methods to predict experimental results. We, tharefore, conclude that these methods are acceptable for use in calculating the nuclear characteristics of the Shearon Harris Units I and 2.
4.3 - 9
....m.
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4.3.4 Summary of Evaluation Findings The Sheaton Harris nuclear design was reviewed according to Section 4.3 of the Standard Review Plan (NUREG-0800). All areas of raview and review procedures from that section have been followed aither for this reactor or for previous similar reactors (e.g., Sumer) or for Topical Report reviews.
The applicant has described the computer programs and calculational techniques used to predict the nuclear characteristics of the reactor design and has provided examples to demonstrate the ability of the analyses to predict reactivity and physics characteristics of the Shearon Harris Units 1 and 2 plants.
To allow for changes of reactivity due to reactor heatup, changes in operating conditions, fuel burnup, and fission product buildup, a significant amount of excess reactivity is designed into the core. The applicant has provided sub-stantial information relating to core reactivity balances for the first cycle and has shown that means have been incorporated into the design to control excess reactivity at all times. The applicant has shown that sufficient con-trol rod worth is available to make the reactor subcritical with an effective multiplication factor no greater than 0.9823 in the hot condition at any time during the cycle with the most reactive control rod stuck in the fully with-drawn position. On the basis of our review, we conclude that the applicant's assessment of reactivity control requirements over the first core cycle is suitably conservative, and that adequate negative worth has been provided by the control system to assure shutdown capability.
Reactivity control require-ments will be reviewed for additional cycles as this information becomes available. We also conclude that nuclear design bases, features, and limits have been established in confonr.ance with the requirements of General De.'.gn Criteria 10, 11, 12, 13, 20, 25, 26, 27, and 28.
G 4.3 - 10
This conclusion is based on the following:
1.
The applicant has met the requirements of GDC 11 with respect to prompt inherent nuclear feedback characteristics in the power operating range by:
Calculating a negative Doppler coefficient of reactivity, and s.
b.
Using calculational methods that have been found acceptable.
The staff has reviewed the Doppler reactivity coefficients in this case and found them to be suitably conservative.
2.
The applicant has met the requirements of GDC 12 with respect to power oscillations which could result in conditions exceeding specified acceptable fuel design limits by:
l a.
Showing that such power oscillations are not possible and/or I
can be easily detected and thereby remedied, and b.
Using calculational methods that have been found acceptable.
3.
The applicant has met the requirements of GDC 13 with respect to provisions of instrumentation and controls to monitor variables and l
systems that can affect the fission process by:
l a.
Providing instrumentation and systems to monitor the core power distribution, control rod positions and patterns, and other process variables such as tenperature and pressure, and i
b.
Providing suitable alams and/or control room indications for these monitored variables.
e 4.3 - 11 l
- l. -
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The applicant has met the requirements of GDC 26 with respect to provision of two independent reactivity control systems of different designs by:
a.
Having a systeri that can reliably control anticipated operational occurrences.
b.
Having a system that can hold the core suberitical under cold conditions, and c.
Having a system that can control planned, normal power changes.
5.
The applicant has met the requirements of GDC 27 with respect to reactivity control systems that have a combined capability in l
conjunction with poison addition by the emergency core cooling l
s.5 stem of reliably controlling reactivity changes under postulated 4ccident conditions by:
a.
Providing a movable control rod system and a liquid poison system, and b.
Performing calculations to demonstrate that the core has sufficient shutdown margin with the highest-worth stuck rod.
6.
The applicant has met the requirements of GDC 28 with respect to postulatedreactivityaccidentsby(reviewedunderSection15.4.8):
a.
Meeting tfie regulatory position in Regulatory Guide 1.77, b.
Meeting the criteria on the capability to cool the core, and c.
Using calculational methods that have been found acceptable for reactivity insertion accidents.
l l
4.3 - 12
i 7.
The applicant has met the requirements of GDC 10, 20 and 25 with respect to specified acceptable fuel design limits by providing analyses demonstrating:
a.
That nonnal operation, including the effects of anticipated operational cccurrences, have met fuel design criteria, 1
b.
That the autonatic initiation of the reactivity control system assures that fuel design criteria are not exceeded as a result of anticipated operational occurrences and assures the automatic operation of systems and components important to safety under accident conditions, and c.
That no single malfunction of the reactivity control system causes violation of the fuel design limits.
e 4.3 - 13
- s REFERENCES WCAP-8385 T. Morita, et al.. " Power Distribution Control and Load Follow Procedure. Septaber 1974.
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4.4 Themal-Hydraulic Design
~ 4.4.1 Themal-Hydraulic DesIan Bases The principal thennal-hydraulic design basis for the Shearon Harris Units 1 and 2 cores is the avoidance of themal-hydraulic induced fuel damage during nomal steady-state operation and anticipated operational transients.
In order to satisfy the design basis, design analysis is perfomed and design limits are established based on the criteria in the subsections which follow.
.' 4.4 I.1 Departure From Nucleata Boiling The margin to departure from nucleate boiling at any point in the core is expressed in tems of the departure from nucleate boiling ratio (DNBR). The DNBR is defined as the ratio of the heat flux required to produce departure from nucleate boiling at the calculated local conditions to the actual local heat flux.
The themal-hydraulic design basis in the SHNPP FSAR Section 4.4.1.1 for the DNBR is as follows:
"There will be at least a 95 percent probability that departure from nuclear boiling (DNB) will not occur on the limiting fuel rods during noraml operation, operational transients, or during transient conditions arising from faults of moderate frequency (ANS Condition I and II events), at a 95 percent confidence level. Historically this has been conservatively met by adhering to the following themal design basis: there must be at least a 95 percent probability that-the minimum departure from nucleate boiling ratio (DNBR) of the limiting power rod during ANS Condition I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The DNBR limit for the correlation is established based on the variance of the correlation such that there is a 95 percent probability with 95 percent confidence that DNB will not occur when the calcalated DNBR is at the DNBR limit. For SHNPP. a minimum DNBR of 1.30 was used."
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During modes of operation associated with ANS Condition I and ANS Condition II events, there is at least a 95 percent probability at. the 95 percent confidence melting temperature.
level that the peak kW/ft fuel rods will not exceed the UO2 istakenas5080*F,(Ref.4.4.1-1),
The melting temperature of unirradiated U02 melting, the fuel and decreasing 58'F per 10,000 MWD /MTU. By precluding UO2 geometry is preserved and possible adverse effects of molten UO on the cladding 2
are eliminated. To preclude center melting and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700*F has been selected as the overpower limit. This provides sufficient margin for uncertainties in the themal evaluations as described in i
Section 4.4.2.9.1."
This design basis is evaluated in Section 4.2.3.2(4) of this Safety Evaluation Report 4.4.1.3 Core Flow Design Basis A minimum of 93.9 percent of the themal flow rate passes through the fuel rod Coolant flow through region of the core and is effective for fuel rod cooling.
the thimble tubes as well as the leakage from the core barrel-baffle region into the core are not considered effective for heat removal.
Ccre cooling evaluations are based on the themal flow rate (minimum flow) l entering the reactor vessel. A maximum of 6.1 percent of this value is allowed as bypass flow. This includes rod cluster control guide thimble cooling flow, head cooling flow, baffle leakage, and leakage to the vessel outlet nozzle.
4.4.1.2 Hydrodynamic Stability The hydrodynamic stability design basis in the SHNPP FSAR Section 4.4.1.4 is as follows:
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" Modes of operation assoicated with ANS Condition I and II events shall not lead to hydrodynamic instability."
4.4.2 Themal-Hydraulic Desian Methodology 4.4.2.1,Themal-H.ydraulic Comparisen The themal-hydraulic design of SHNPP is similar to that of V. C. Summer Units I and 2, which have been approved for licensing, and to North Anna Units 1 and 2, which are operating units.
Values of critical design para-meters for the three reactor designs are compared in Table 1. ' There is no change in design criteria for SHNPP, the reactor is designed to a minimum DNBR of 1.30 using W-3, as well as no fuel centerline melting during nomal oper-ation, operational transients and faults of moderate frequency.
The themal-hydraulic parameters for SHNPP are very similar to V. C. Summer which supports the acceptability of the SHNPP themal-hydraulic design. The staff has also reviewed the differences in flow and inlet tenperature between SHNPP and 'the operating North Anna units and have found these differences in flow and inlet temperatures are consistent with the difference in minimum DNBR.
4.4.2.2 Departure fiom Nucleate Boilina l
DNBRs are calculated by using the W-3 critical heat flux correlation. The THINC-IV computer code (Ref.1) is used to detemine the flow distribution in the core and the local conditions in the hot channel for use in the DNB correlation.
Westinghouse has perfomed a test program (Ref. 2) on the 17x17 fuel assembly.
A correction factor was developed to adopt the W-3 correlation to 17x17 assemblies with top split mixing vane grids referred to as "R" grids. The l
SHNPP design includes a conservative multiplier.("R" grid DNB correlation) of
.865 for all DNB analyses.
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The test results indicate that a reactor core using this geometry may operate with a minimum DNBR of 1.28 and satisfy the design criterion. However, a j
minimum DNBR of 1.30 is used for SHNPP.
The design value of 6.1% bypass flow is based on calculations performed using drawing tolerances in the worst direction and accounting for uncertainties in pressure drops. These calculations show the core bypass flow to be no greater
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than the design value.
i i
l These calculations use experimental data from a series of 1/7 scale hydraulic reactor model tests where core flow distributions and pressure drops were j
measured.
y The staff has reviewed the 6.1% bypass flow design basis and has found it to be' consistent and conservative relative to previously approved designs and to have been properly included in' determining the minimum DNBR and is therefore acceptable to the staff.
30 minimum DNBR as described in Section 4.4.2 The staff finds the use c!
of the SHNPP to be acceptn.a.
4.4.3 Design Abnormalities 4.4.3.1 Fuel Rod Bowing A significant parameter that influences the thermal-hydraulic design is rod-to-rod bowing within fuel assemblies.
The staff has developed interim criteria for evaluating the effects of rod bow on DNB for application to the Westinghouse standard 17x17 fuel assembly.
The resultant reduction in DNBR due to rod bow is given by:
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Burnup DNBR Reduction (MWD /MTU)
(5) 0 0
3500 0
5000 0
10000 2.15 15000 4.64 20000 6.74 25000 8.59 30000 10.27 35000 13.07 40000 19.09 The appropriate provisions should be incorporated into the Technical Specifications. The applicant should also insert into the basis of the Technical Specifications any generic or plant specific margin that may be used to offset the reduction in DN8R due to rod bowing, and reference the source and staff approval of each generic margin. With these requirmients satisified by the applicant, the staff concludes that they have adequately accommodated the reductions listed above.
4.4.3.2-Crud Deposition Crud deposition in the core and an associated change in core pressure drop ha've been observed in some PWRs, In response to a staff question concerning crud deposition the licensee responded:
"There has been no case reported to Westinghouse of significant flow redgetion
(
in a relatively short period of time due to buildup of crud on the fuel rods l
of any Westinghouse plant. Additionally, there has been no report to Westing-house of a significant flow reduction in a relatively short period of time at any Westinghouse plant."
The staff does not agree with implicatic ~ b ( crud deposition can not happen at a Westinghouse plant.
However, tu s.ar' finds the flow measurement w
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technique as described for SHNPP is sufficient to detect a crud buildup if this should occur. Therefore, SHNPP technical specifications-should require
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that the RCS flow is monitored every 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.
In describing the flow measurement technique in Amendment 5 to the FSAR, the applicant presented a set of uncertainty values indicating a total loop flow uncertainty of 3.05 and a reactor flow uncertainty of 1.75%.
If the applicant mvises the Westinghouse Standard Tet.hnical Specification value f
for flow uncertainty which is 3.55, such revision must be reviewed by the staff.
4.4.4.6 Hydrod_ynamic Stability In steady-state, two phase, heated flow in parallel channels, the potential for hydrodynamic instability exists. The applicant provided the following infomation in the FSAR to support the contention that the SHNPP core is themal hydraulically stable.
" Boiling flows may be susceptible to themohydrodynam,1c instabilities. These instabilities are undesirable in reactors since they may cause a change in
- themohydraulic conditions that may lead to a reduction in the DMB heat flux relative to that ob
- erved during a steady flow condition or to undesired forced vibrations of core components. Therefore, a themohydraulic design criterion as developed which states that modes of operation under ANS Condition I and II events shall not lead to thermohydrodynamic instabilities.
Two specific types of flow instabilities are considered for Westinghouse PilR operation. These are the Ledinegg, or flow excursion type of static instability, and the density wave type of dynamic instability.
A Ledinegg instability involves a sudden change in flow rate from one steady state to another. This instability occurs when the slope of the reactor coolant system pressure drop-flow rate curve $a PA G internal) becomes algebraically smaller than the loop supply (piang head) pressure drop-flow rate curve 6a P/ G external). The criterion for stability is thus 6a PA G internal
>6P/dG external. The Westinghouse reactor coofant pump head curve has a negative slope ( 8F/ G external 4) whereas the reactor coolant system pressure drop flow curve has a positive slope $a P/dG internal:0) over the ANS Condition I and ANS Condition II operational ranges. Thus, the Ledinegg instability will not occur.
u.;
A simple method has been developed by Ishii (Reference 3) for parallel closed channel systems to evaluate whether a given condition is stable with respect to the density wave type of dynamic instability. This method has been used to l
assess the stability of typical Westinghouse reactor design including Virgil C.
Summer, under ANS Condition I and II operation. The results indicate that a large margin to density wave instability exists. e.g., increases on the order of 200 percent of rated reactor power would be required for the predicted inception of this type of instability.
, The application of the method of Ishii to Westinghouse reactor designs is conservative due to the parallel open channel feature of Westinghouse PWR cores. For such cores, there is little resistance to lateral flow leaving the flow channels of high power density to lower power density channels.
This coupling with cooler channels has led to the opinion that an open channel configuration is more stable than the above closed channel analysis under the same boundary conditions.
Flow stability tests (P.:f. 4) have been conduct d t.srs the clossd channel systens were shown to be less stable than when the same channels were cross connected at severa1 locations. The cross connections were such that the resistance to channel to channel cross flow and enthalpy perturbations would be greater than that which would exist in a PWR core which has a relatively low resistance to cross flow.
Flow instabilties which have been observed have occurred almost exclusively in closed channel systems operating at low pressure relative to the Westinghouse PWR operating pressures. Kao, Morgan, and Parker (Ref. 5) analyzed parallel
. closed channel stablity experiments simulating a reactor core flow. These experiments were conducted at pressures up to 2200 psia. The results showed that for flow and power levels typical of power reactor conditions, no flow
. oscillations could be induced above 1200 psia.
Additional evidence that flow instabilities do not adversely affect themal margin is provided by the data from the rod bundle DNB tests. Many Westing-house rod bundles have been tested over wide ranges of operating conditions with no evidence or premature DNB or of inconsistent data which might be indicative of flow instabilities in the rod bundle.
.. _ ~.. _...
In suinnary, it is concluded that thermohydrodynamic instabilities will not occur under ANS Condition I and II modes of operation for Westinghouse PWR reactor designs.
A large power margin, greater than doubling rated power, exists to predicted i
inception of such instabilities. Analysis has been perfomed which shows that minor plant to plant differences in Westinghouse reactor designs such as fuel i
assenbly arrays, core power to flow ratios, fuel assembly length, etc. will l
not result in gross deterioration of the above power margins."
l l
l The staff is presently conducting a generic study of the hydrodynamic stability-characteristics of pressurized water reactors. Limitations to the themal-l hydraulic design resulting from the staff study will be compensated by appropr-l iate operating restrictions if necessary; however, no operating restrictions are anticipated.
In the interim, the staff concludes that past operating experience, flow stability experiments, and the inherent themal-hydraulic characteristics of Westinghouse pressurized water reactors provide a basis for accepting the SHNPP stability evaluation for issuance of an operating license.
4.4.6.4 Loose Parts Monitoring Systems l
The applicant has provided documentation of the loose parts monitoring system in Amer.dment 5 to the FSAR. The applicant has not done a complete evaluation of the system for confomance to Regulatory Guide 1.133 or committed to confom to the Regulatory Guide 1.133.
i The applicant must complete the evaluation of the system for confomance to Regulatory Guide 1.133 providing justifications for any deviations and the staff must review and approve these deviations.
N-1 Pump Operation In response to staff question 492.3 CP&L stated that it does not intend to pursue N-1 loop operation as a licensing basis for SHNPP at this time.
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Hence, operation at power with less than four pu;nps operating will not be pemitted. The staff will require that the technical specifications include appropriate provisions to ensure that these types of operation are prohibited.
TMI-Action Plan NUREG-0737 Iten II.F.2 The applicant provided in reference 6.a partial response to the item-by-item documentaton required by NUREG-0737.
Reference 6 promises that additional infomation to support the acceptability of the ICC monitoring system will be provided by August 30, 1983 in the applicants response to Supplement 1 of NUREG-0737,(item Regulatory Guide 1.97 Rev. 2). The applicant must submit the remainder of the documentation required by NUREG-0737 and the complete documentation must be received, reviewed and approved prior to the licensing of SHNPP.
Conclusion and Sum:nar_y The themal-hydraulic design of the core for SHNPP has been reviewed. The scope of our review included the design basis and the steady-state analysis of the core themal-hydraulic perfomance. The acceptance criteria used as the basis of our evaluation are set forth "in the Standard Review Plan (SRP),
NUREG-0800, in Section 4.4 "Themal and Hydraulic Design." The review con-centrated on the difference between the proposed design and those designs that have been previously reviewed and found acceptable by the staff.
Based on our review, the staff concludes that the themal-hydraulic design of the initial core of SHNPP is acceptable, provided satisfactory resolution is obtained for the following items identified in this SER:
1.
The applicant has provided documentation of the loose parts monitoring system in Amendment 5 to the FSAR. The applicant has not done a complete evaluation of the system for confomance to Regulatory Guide 1.133 or cormnitted to confom to Regulatory Guide 1.133.
The applicant must complete the evaluation of the systen for confor-mance to Regulatory Guide 1.133 providing justifications for any deviations and the staff must review and approve these deviations.
_-m 2.
The applicant must submit the remainder of the documentation required by Item II.F.2 of NUREG4737 and the entire documentation package must be reviewed and approved prior to the licensing of SHNPP.
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TABLE 1.0 THERMAL AND HYDRAULIC CGtP!4RISON TABLE North Anna Design Parameters Shearon Harris V. C. Summer Units 1 & 2 Reactor core heat output (Nwt) 2775 7775 2775 System pressure, minimum steadystate(psia) 2220 2220 2220 Miminum DNBR at nominal design conditions Typical flow channel 1.98 1.98 2.15 Thimble (cold wall) flow channel 1.68 1.68 1.77 Minimum DNBR for design transients 1.30 1.30 1.30 DNB correlation "R" (W-3 with modified Spacer Factors)
Coolant Flow l
Total thermal flow rate l
'(10 l b,hr.)
109.1 109.6 105.2 6
Coolant Temperatures Nominal inlet (F) 556.0 556.0 546.8 Heat Transfer _
Active heat transfer, surface area (ft.2) 43,600 48.600 48,600
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TABLE 1.0 Con't THERMAL AND HYDRAULIC COMPARISON TABLE l
l North Anna w
_ Design Parameters Shearon Harris V. C Summer thits 18 2 Heat Transfer 1
Average heat flux
. (Btu /hr.-ft.2) 189.800 189,800 189.800 Average linear power (kW/ft) 5.44 5.44 5.45 Peak linear power for nomal operation (kW/ft) 12.6 12.6 13.6 4
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SECTION 4.4 REFEF.D j.ES,
{
1.
Shefcheck, J., " Application of the THINC Program to PWR Design,"
WCAP-7359-L (Froprietary), August 1969 and WCAP-7838, January 1972.
2.
Motley, F. E., Wenzel, A. H. and Cadek, F. F., " Critical Heat Flux Testing of 17x17 Fuel Assembly Geometry with 22-Inch Grid Spacing "
WCAP-8536 (Proprietary), May 1975 and WCAP-8537 May 1975.
A s 3.
Saha, P., Ishii, M., and Zube N., "An Experimental Investigation of the Thermally Induced Flow Oscillations in Two-Phase Systems,"
J.
Heat Transfer, November 1976.
' 4.
S. Kakec T. N. Veziroglu, K. Akyuzlu, O. Berkol
" Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parra11el-Channel Upflow System " Proc. of 5th International Heat Transfer Conference, Tokyo, September 3-7, 1974.
5.
H. S. Kao, C. D. Morgar., and W. B. Parker, " Prediction of Flow Oscillation in Reactor Core Channel." Trans. ANS, Vol. 16, 1973, pp. 212-213.
6.
Letter, M. A. McDuffie (CFSL) to H. R. Denton (NRC), dated August 11, 1983.
4 9
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15.4.1
_tmcontrolled Rod Cluster Control Assembly (Rod) Bank )Hthdrawal From Zero power Conditions Discussion 3
The consequences of an uncontrolled rod cluster control assembly bank withdrawal at zero power have been analyzed. Such a transient can be caused by a failure of the reactor control or rod control systems. The analysis assumed a conservatively small (in absolute magnitude) negative Doppler coefficient and a positive moderator coefficient.
Further, hot 2ero power initial conditiens with the reactor just critical are chosen because they are known to maximize the calculated consequences. The reactivity insertion rate is assumed to be equivalent to the simultaneous withdrawal of the two highest worth banks at maximum speed (45 inches per minutes).
Reactor trip is assumed to occur on the low setting of the power range neutron flux channel at 35 percent of full power (a ten percent uncertainty has been addedtothesetpointvalue). The maxirium heat flux.is much less than the full power value and average fuel tamperature increases to a value lower than the nominal full power value. The minimum DNBR at all times remains above the limiting value of 1.30.
Evaluation Findings We have reviewed this event according to the Standard Review plan (NUREG-0800)
Section 15.4.1.
The possibilities for single failures of the reactor control system which could result in uncontrolled withdrawal of control rods under low pcwer start-up conditions have been reviewed. The scope of the review has included investi-l gations of initial conditions and control rod reactivity worths, the course l
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of the resulting transients or steady-state conditions, and the instrument response to the transient or power maldistribution. The methods used to determine the peak fuel rod response, and the input into the analysis, such as power distributions and reactivity feedback effects due to moderator and fuel temperature changes, have been examined.
We conclude that the requirements of General Design Criteria 10, 20, and 25 have been met. This conclusion is based on the following:
The applicant has met the requirement of GDC 10 that the specified acceptable fuel design limits are not exceeded GDC 20 that the reactivity control systems are automatically initiated so that specified acceptable fuel design limits are not exceeded, and GDC 25 that single malfunctions in the reactivity control system will not cause the specified acceptable fuel design limits to be exceeded.
These requirements have been met by comparing the resulting extreme operating conditions and response for the fuel (i.e., fuel duty) with the acceptance criteria for fuel damage (e.g., critical heat flux, fuel t::nperatures, and clad strain limits should not be exceeded), to assure, that fuel rod failure will be precluded for this event. The basis for accept:
in the staff review is that the applicant's analyses of the maximum
- tents for single error control rod withdrawal from a suberitical or low-pwer condition have been confimed, that the analytical methodt and input data are reasonably conservative and that specified acceptable fuel design limits will not be exceeded.
15.4.2 Uncontrolled Rod Cluster Control Assembly (Rod) Bank Withdrawal a_t Power l
Discussion The consequences of uncontrolled withdrawal of a rod bank in the power operating range have been analyzed. The effect of such an event is an increase in coolant temperature (due to the core-turbine power mismatch) which must be teminated prior 'to exceeding fuel design limits.
15.4 - 2
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The analysis is perfomed as a function of reactivity insertion rates, reactivity feedback coefficients and core power level. Protection is provided by the high neutron flux trip, the overtemperature AT and over-power AT trips, and pressurizer pressure and pressurizer utter level trips. In no case does the departure from nucleate boiling ratio fall below 1.30.
Adequate fuel cooling is therefore maintained. The maximum heat flux reached
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including uncertainties does not exceed 118 percent of full power, thus pre-cluding fuel centerline melting.
Evaluation Findings
~
We have reviewed this event according to Section 15.4.2 of the Standard Review Plan (NUREG-0800).
The possibilities for single failures of the reactor control system which
, could result in uncontrolled withdrawal of control rods beyond nomal limits under power operation conditions have been reviewed. The' scope of the review has included investigations of possible initial conditions and the range of reactivity insertions, the course of the resulting transients and the l
instrunentation response to the transient. The met' hods used to detemine the peak fuel rod response, and the input into the analysis, such as power distri-butions, rod reactivities, and reactivity feedback effects of moderator and fuel temperature changes, have been examined.
We conclude that the requirements of General Design Criteria 10,10, and 25 have been met. This conclusion is based on the following:
The applicant has met the requirements of GDC 10 that the specified acceptable fuel design limits are not exceeded, GDC 20 that the reactivity control systems are automatically initiated so that specified acceptable fuel design limits are not exceeded, and GDC 25 the single malfunctions in the reactivity control system will not cause the spec *.1ad acceptable fuel design limits to be exceeded.
These requirements have been met by comparing,the resulting extreme operating conditions and response for the fuel (i.e., fuel duty) with the acceptance 15.4 - 3
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criteria for fuel damage -(e.g., critical heat flux, fuel temperatures and clad strain limits should not be exceeded), to assure that fuel rod failure will be precluded for this event. The basis for acceptance in the staff review is that the applicant's analysis of maximum transients for single error control rod malfunctions have been confimed, the analytical nethods and input data are reasonably conservative and that specified acceptable fuel design limits will not be exceeded.
15.4.3 Rod Cluster Control Assembly Malfunctions Discussion Rod cluster control assembly misalignment incidents including a dropped full length assembly, a dropped full length bank, a misaligned full length assembly and the withdrawal of a single assembly while operating at power have been a'nalyzed by the applicant. Misaligned rods are detectable by: (1) asymmetric power distributions sensed by excore nuclear instrumentation or core exit
- thermocouples, (2) rod deviation alata, and (3) rod position indicators. A deviation of a rod from its bank by about 15 inches or twice the resolution of the rod position indicator will not cause power distribution to exceed design limits.
Additional surveillance will be required to assure rod alignment if one or more rod position channels are out of service.
In the event of a dropped assembly or group of assemblies the reactor will typically scram on a neutron flux negative rate trip, and analysis indicates that themal limits will not be exceeded for the event.
If the rod locations are such that the reactor does not scram, however, the automatic controller i
may return the reactor to full power and the control could result in a power overshoot. An analysis methodology for this event has been developed by Westinghouse, and reported in WCAP-10297-P " Dropped Rod Methodology for Negative Flux Rate Trip Plants," January 1982. This methodology has been reviewed and approved by the NRC staff. The review is in a memorandum for F. Miraglia from L. Rubenstein, " Review of the Westinghouse Report ' Dropped Rod Methodology for Negative Flux kate Trip Plants'", December 1983.
Generally, 15.4 - 4
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detailed analyses for most reactors, for most cycles, show that if this event occurs thermal limits will not be exceeded. However, the analysis is reactor and cycle specific, and the analyses for Shearon Harris for Cycle 1 has not been completed as yet. The staff has also accepted an interim position for operating reactors which consists of a restriction on operations above ninety percent power such that either the reactor is in manual control or rods are required to be out greater than 215 steps. This restriction will be applied to Shearon Harris thits I and 2 in the event that calculations for Cycle 1 operation are not completed in time for initial operations. With this restriction thermal limits will not be exceeded. Approval of the analysis specific to Shearon Harris for Cycle 1 will result in removing the restriction.
Similar analysis will also be needed for each subsequent reload cycle.
For cases where a group is inserted to its insertion limit with a single rod in the group stuck in the fully withdrawn position analysis indicates that departure from nucleate boiling will not occur. We have reviewed the calculated estimates of the expected reactivity and power distribution changes that accompany postulated misalignments of representative assemblies.
We have concluded that the values used in this analysis conservatively bound e
the expected values including calculational uncertainties.
The inadvertent withdrawal of a single assembly requires multiple failures in the control system, multiple operator errors or deliberate operator actions combined with a single failure of the control system. As a result the single assembly withdrawal is classifid as an infrequent occurrence.
The resulting transient is similar to that due to a bank withdrawal but the increased peaking factor may cause departure from nucleate boiling to occur in the regfon surrounding the withdrawal assembly.
Less than five percent of the rods in the core experience departure from nucleate boiling for such a transient.
i 15.4 - 5
Evaluation Findinas We have reviewed this event according to Section 15.4.3 of the Standard Review Plan (NUREG-0800).
The possibilities for single failures of the reactor control system which could result in a movement' or malposition of control rods beyond nomal limits have been reviewed. The scope of the review has included investig-ations of possible rod malposition configurations, the course of the resulting transients or steady-state conditions, and the instrumentation response to the transient or power maldistribution. The methods used to detemine the peak fuel rod response, and the input to that analysis, such as power distribution changes, rod reactivities, and reactivity feedback effects due to moderator and fuel temperature changes, have been examined.
We conclude that the requirements of General Design Criteria 10, 20 and 25 have been met. This conclusion is based on the following:
The applicant has met the requirements of GDC 10 that the specified acceptable fuel design limits are not exceeded, GDC 20 that the mactivity control systems are automatically initiated so that specified acceptable fuel design limits are not exceeded, and GDC 25 that single ' malfunctions in the reactivity control systen will not cause the specified acceptable fuel design limits to be exceeded.
These requirements have been met by comparin2 the resulting extreme operating l
conditions and response for the fuel (i.e., fuel duty) with the acceptance criteria for fuel damage (e.g., critical heat flux, fuel temperatures and clad i
strain limits should not be exceeded), to assure that fuel rod failure will be l
precludet For this event. The basis for acceptance in the staff review is that maximum co.ifigurations and transients for single error control rod malfunctions have been analyzed, that the analysis methods and input data are resonably conservative and that specified acceptable fuel design limits will not be exceeded.
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f 15.4.7 Inadvertent Loading of a Fuel Assembly into Improper position Discussion Strict administrative controls in the form of previously approved established procedures and startup testing are followed during fuel loadings to prevent operation with a fuel assembly in an improper location or a misloaded burnable poison assembly.
Nevertheless, an analysis of the consequences of a loading error has been performed.
Comparisons of power distributions calculated for the nominal fuel loading pattern and those calculated for five loadings with misplaced fuel assemblies or burnable poison assemblies are present
'v the applicant. The selected non-normal loadings represent the spectrun:.
Stential inadvertent fuel misplacement. Calculations included, in part.
'r, the power in assemblies which contain provisions for monitoring with inewe detectors.
As part of the required startup testing, the incore detector system is used to detect misloaded fuel prior to operating at po'wer. The analysis described above shows that all but one on the above misloading svents would be detected by this test.
In the excepted case, an interchange of Region 1 and 2 assemblies near the center of the core, the increase in the power peaking is approximately-equal to the uncertainty in the measurement of this quantity
-( 5 percent). This uncertainty is allowed for in analyses so that this misloading event does not result in unacceptable consequences.
Evaluation Findings We have reviewed this event according to Section 15.4.7 of the Standard ReviewPlan(NUREG-0800).
We have evaluated the consequences of a spectrum of postulated fuel loading We conclude that the analyses provided by the applicant have shown errors.
for each case considered that either the error is detectable by the available instrumentation (and hence remediable) or the error is undetectable but the 15.4 - 7
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I offsite consequences of any fuel rod failures are a small fraction of 10 CFR Part 100 guidelines. The applicant affirms that the available incore instrumentation will be used before the start of a fuel cycle to search for fuel loading errors.
We conclude that the requirements of General Design Criterion 13 and 10 CFR Part 100 have been met. This conclusion is based on the following:
The applicant has met the requirements of GDC13 with respect to providing adequate provisions to minimize the potential of a misloaded fuel assembly going undetected and meets Part 100 with respect to mitigating the con-sequences of reactor operations with a misloaded fuel assenbly. These requirements have been met by providing acceptable procedures and design
. features that will minimize the likelihood of loading fuel in a location other than its designated place.
15.4.8 Rurture of a Control Rod Drive Mechanism Housing (Dod Cluster u
Control Assembly E.iection)
Discussion The mechanical failure of a control rod mechanism pressure housing would result in the ejection of a rod cluster control assembly.
For assemblies initially inserted, the consequences would be a rapid reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage. Although mechanical provisions have been made to make this accident extremely unlikely, the applicant has analyzed the consequences of such an event.
Methods used in the analysis are reported in WCAP-7588, Revision 2. "An Evaluation of the Rod Ejection Accident in Westinghouse Reactors Using Spatial Kinetics Methods." which has been reviewed and accepted by the staff. This report demonstrated that the model used in the accident analysis is conservative relative to a three dimensional kinetics calculation.
15.4 - 8
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o, ' ' t 'a The applicant's criteria for gross damage of fuel are a maximum clad temperature of 2700 degress Fahrenheit and an energy deposition of 200 calories per gram in the hottest pellet. These criteria are more con-servative* than those proposed in Regulatory Guide 1.77. Therefore, they are acceptable.
Four cases were analyzed: beginning-of-cycle at 102 percent and zero power and end-of-cycle at 102 percent and zero power. The highest clad temperatures.
2504 degrees Fahrenheit, and the highest fuel enthalpy,180 calories per gram, i-was reached in the beginning-of-cycle full power case. The analysis also shows that less than 10 percent of the fuel experiences departure from nucleate boiling and less than 10 percent of the hot pellet melts. Analyses have been perfomed to show that the pressure surge prodsced by_the rod ejection is mild and will not approach the Reactor Coolant System emergency limits.
Futh_er analyses have shown that a cascade effect, i.e., the ejection of a further rod due to the ejection of the first one, is not credible.
~
Evaluation Findinas We have reviewed this event according to Section 15.4.8 of the Standard Review Plan (NUREG-0800).
We conclude that the analysis of the rod ejection accident is acceptable and meets the requirements of General Design Criterion 28. This conclusion is based on the following:
- Regulatory Guide 1,77 has an acceptance criterion of 280 calories per gram energy deposition and no criterion for clad temperature other than that implicit in requirements for fuel and pressure vessel damage.
15.4 - 9
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The applicant met the requirements of GDC 28 with respect to preventing postulated reactivity accidents that could result in damage to the reactor coolant pressure boundary greater than limited lo:a1 yielding, or cause sufficient damage that would signficantly impair the capability to cool the core. The requirements have been met by demonstrating that the i
regulatory positions of Regulatory Guide 1.77, " Assumptions Used for Evaluating a Control Rod Ejection Accident for PWR's" are complied with.
The staff has evaluated the applicant's analysis of the assumed control rod ejection accident and finds the assumptions, calculation techniques.
and consequences acceptable. Since the calculations resulted in peak fuel enthalpies less than 280 cal /ge, prompt fuel rupture with consequent rapid heat transfer to the coolant from finely aispersed molten UO tes 2
assumed not to occur. The pressure surge was, therefore, calculated on the basis of conventional heat transfer fran the fuel and resulted in a pressure increase below " Service Limit C" (as defined in Section III,
" Nuclear Power Plant Components " of the ASME Boiler and Pressure Vessel Code) for the maximum control rod worths assumed. The staff believes that the cal:ulations contain sufficient conservatism, bo'th in the initial assumptions and in the analytical models, to ensure that primary system integrity will be maintaine.i.
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