ML20037A994
| ML20037A994 | |
| Person / Time | |
|---|---|
| Site: | Dresden  |
| Issue date: | 09/14/1967 |
| From: | Julie Hughes COMMONWEALTH EDISON CO. |
| To: | |
| References | |
| NUDOCS 8008280628 | |
| Download: ML20037A994 (37) | |
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- 9 Dr. Peter A. Morris, Director O nruim Division of Reactor Licensing %
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U. S. Atomic Energy Commission M """"
Washington, D.C. 20545 y
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Subject:
Proposed Change No. 14 to the Operating License DPR-2, as amended Docket 50-10
Dear Dr. Morris :
@egulatog, Suppl File Cf Commonwealth Edison Company requests authorization, pursuant to 10CFR5059 and Paragraph 3.a.(4) of the Dresden Nuclear Power Station Operating License DPR-2, as amended, to make such changes as are itemized hereafter to provide for the operation of Dresden Unit 1 reactor with a replacement fuel batch, designated Type VI fuel.
The items to be changed are as follows:
Item No. 1 Change item "2.
Nuclear Core" of Section "3.
DESIGN FEATURES" of Appendix "A" to DPR-2 to read in its entirety:
"2.
Nucicar Core
" Maximum core diameter (circumscribed circle) 129 in.
Maximum active fuel length - cold 112 in.
Maximum number of fuel assemblics by types:
Type I 65 Type III 190 Type III-F 100 Type V 106 Type VI 96 Maximum total number of fuci assemblies 4SS "The reactor may be operated at any power up to and including rated power with any configuration of the various types of fuel assemblies installed, provided the maximum number is within the limits specified above."
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em Item No. 2 Change the second paragraph of item "3.
Fuel" of Section "B.
DESIGN FEATURES" of Appendix "A" to DPR-2 to read in its entirety:
"3.
Fuci "The nominal fuci pollet density is 93.5% of theore-tical for all Type VI UO, fuel rods.
For Gd,0.-UO 2 fuci rods of Type VI fuct, the nominal fuel pefict density is 91.50 of theoretical.
Pcilot dishing is used for all Type VI fuel pc11 cts which reduces the initial volume of UO2 in given stack height by up to 20 over non-dished pellets.
Previously supplied fuel has a minimum density of 940.
Iten No. 3 Change item "3.
Determination of Maximum Reactor Power" of Section "D.
POWER OPERATION" of Appenalx "A" to DPR-2 to rena in its entirety:
"3.
Determination of Maximum Reactor Power The rated power of the reactor shall be limited to a maximum steady state value of 700 MU(t).
"The maximum allowabic stead / state heat flux limits 2
expressed in units of Btu /(hr)(ft ) shall never ex-coed the following values:
Fuel Type I 350,000 Fuci Type III 360,000 Fuci Type III-F 360,000 Fuci Type V 360,000 Fuel Type VI 360,000 "The reactor shall be operated within the above limits such that a minimum critical heat flux ratio (MCHFR) of at 1 cast 1.5, evaluated at 125% of rated! power, will be maintained in each type of fuct closest. o burnout t
in the hottest channel in the core based on a uniform steam quality over the cross section of the channel.
This burnout ratio shall be based upon the correlation in Edison's " Recommended Curves of Burnout Limit for Design and Operation of Boiling Mater Reactors", dated January 5, 1962.*
The reactor shall be operated always well within the bounds of stability, as evidence by the operation itself and any experimental data prc#uced.
Item No. 4 Change Appendix "A" to DPR-2 by deleting Table 2 (revised 8/25/66) attached thereto, and substituting Tabic 2 (dated 7/28/67) attached herewith.
- APED-3892 ii e
r-2 September 14, 1967 Page 3 Pursuant to 10CFR5 0.59 and 3.2. (4) of DPR-2, a Description. and Safety Evaluation Report in support of Proposed Change No.14 to Appendix "A" is attached hereto as "EXIIIBIT I".
In our opinion Proposed Change No. 14 shall not result in hazards which are greater than or different from, those analy:cd in the
!!a:ards Su'mmary Report, specifically there is (1) no increase in the probability of, or (2) no increase in the consequences of, or (3) the creation of a credibic probability of an accident different from, those accidents previously analyzed in the Hazards Summary J
Report as amended or in connection the amendments and changes to Operating License DPR-2.
Very truly yours, COMMON!IEALTH EDISON COMPANY Q/ f.
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clear LicensinV Administrator j
s Attachments:
Tabic II Exhibit I SUBSCRIBED and sworn to before me this /H6% day of/h.
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i EXHIBIT I Dresden Nuclear Power Station
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Description and Safety Evaluation Report of Type VI Fuel Contents SECTION I Description of Proposed Change to Appendix "A" to DPR-2 Physical Char \\acteristics of the.Pseload Fuel SECTION II SECTION III Nuclear Characteristics of the Fuel and Core SECTION IV Thcroal and Hydraulic Characteristics SECTION V Previous Experience with Gadolinia UO2 and Background Information SECTION VI Safety Considerations This report provides technical information in support of the attached application for change of Dresden Operating License DPR-2, as anended.
It is not intended that the material contained her6in constitutes " Technical Specifications" in the sense of the Licen. sing regulations (10 CFR, Part 50, Section 50.36).
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- e SECTION I - Description of Proposed Change to Appendix "A"
to DPR-2 These four itens are proposed primarily for the purpose of obtaining authority to utilize Type VI fuel assemblies in the Dresden reactor.
Current plans are to load up to 96 Type VI fuel assemblics in the Dresden reactor during the refueling for the operating Cycle No. 6.
The physical, thermal-hydraulic, and nucicar properties of the Type VI fuel and the effect of using
~this fuel in a mixed core are covered in Sections II through VI of-this report.
These itcas also provide an up-to-date listing of the fuel assemblies availabic for the next refueling.
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e SECTION II - physical Characteristics of the Reload Fuel
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It is proposed to reload the reactor with a maximum of 96 fuel assemblies.
This reload will consist of 83 " normal" assemblies (Type VI) and 13 " instrumented" assemblies Type VI-I.
The Type VI fuel consists of six rows of six Zircaloy-2 clad fuel rods as shown in Figure 1.
The Type VI-I fuel is identical to Type VI except that one fuel rod is replaced by an instrument channel.
The Type VI fuel is physically very similar to the Type III-F fuel.
Each of the 96 assemblies contains one seg-k monted fuel rod which connects and positions five spring spacers as in Type III-F fuel.
Each of the 96 assemblies contains six fuel rods of lower than average enrichment which nre placed in corner locations to reduce local power peaking.
Each of the 83 Type VI assenblies contain one removable fuel rod.
This rod is loaded with Gd 023 (gadolinia) burnable poison incor-porated in the UO2 pellets.
In each of the 13 Type VI-I assemblies the instrument channel is located in the position occupied by the removable, burnable poison bearing fuel rod of a Type VI assembly.
The Type VI-I assemblies contain no burnable poison.
The two main differences between the Type VI fuel and the Type III-F fuel are as follows:
l.
In the Type VI fuel, the Gd 023 burnable poison is carried in a single UO2 (urania) rod rather than in a single inert alumina rod.
Experience with gadolinia-UO2 is summarized in Section V.
2.
In the Type VI fuel, the fuel pellets are dished to reduce axial thermal expansion of the pellet stack and to provide distributed space for pellet swelling.
The tot'al UO2 loading in the Type VI asse-blies is nearly identical to that of the Type III-F assemblies since the loss in loading from the use of dished pellets and somewhat lower pellet density is compensated for by the gain in loading from the use of UO2 as the carrier for the poison rod rather than Al 02 3-
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The purpose of the Gd 023 burnable poison is to aid the control rods in holding down excess reactivity at beginning of Cycle 6.
Dresden operating data from Cycle 4 which contained Gd 02 3 bearing Type III-F fuel, and calculations on Type VI fuel show that the burnable poison will be effectively depleted by the time that approximately 3000 MWD /T are generated in the fuel assemblies loaded at beginning of Cycle 6.
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SECTION III - Nuc1 car Characteristics of the Fuel and Core
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.The key nucicar characteristics of the Type VI frtel and the l
resulting core characteristics when this fuel is - loaded into the l
Dresden reactor are given in the follouing paragraphs.
A repro-l sentative core loading for the 96 assenbly reload is niven in Fig. 2.
l Deviations from this pattern nay result at the time os loading, but l
the loading vill be essentially a uniform "1 in 4" scatter of the new fuel,as employed in previous Dresden fuel cycles.
p The -basic physics lattice data for Type VI fuel are sunnarized in Tabic 1.
Data are presented for the Type VI and-Type VI-I assemblics.
There are no safety probicas introduced uith the use of the gadolinia since the control rods are always availabic as a backup.
The uncertainty in the calculated depletion rate of the gadolinia is expected to be small, based on the perfornance of gadolinia in the Type III-F fuel in Cycle 4.
If the gadolinia depiction rate is faster than anticipated, control rods can be inserted to control reactivity and maintain acceptabic power distributions.
If the gadolinia depiction rate is slower than anticipated, the control rods in the core n.ay be withdraun or in the norst case, after all rods are fully uithdraun, the reactor may have to be dorated sooner than anticipated.
Abnormal or unexpected poucr distributions can be detected with reactor incore instrumentation and periodical flux wire irradiations.
Fuel uill be verified.to be as designed by ex-haustive quality control procedures, including reactivity substitution tests, following completion of fuel fabrication and assembly.
The nuclear calculation nothods used are described and discussed at the end of this Section.
A.
Pouer Distribution The overall gross core power peak is expected to be about the saac in magnitude as it was in Cycle 4.
A degree of radial power flattening will be obtained by the shifting of a number of Type V assemblies to the far periphery (sco Fig. 2).
These assemblics, at BOL Cycic 6, are about at their naximum reactivity, having essentially depicted their gadolinia poison by the end of Cycle 5.
For Cycle 6, the predicted overall gross pouer peak (radial l
x axial x local) at any tine during the cycle will not exceed 3.24, uhich is the value established from the license requirement that the naxinun heat flux = 360,000 Btu /hr.- ft2 at 700 MUg.
In addition, l
the predicted minimum critieni heat flux ratio'MCHrR at 125% power j
at any tiac during the cycle will be Dhl.5.
Thornal-hydraulic char-acteristics are given in Section IV.
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L The maximum local peaks for the Type VI fuel are shown' in Table 2.
The maximum local peaks occur at BOC 6.
The magnitude of the local peak for Type VI is similar to that of Type III-F.
The local peak in Type VI-I is lower than in Type VI due to the absence of the gadolinia rod.
With burnup and depletion of gadolinia (in Type VI), the local peaks will i
decrease (and shift to different rods) from their initial values.
Relative power distributions within the Type VI assemblies are presented in Figs. 3 and 4.
B.
Temperature and Void Coefficients The temperature and void coefficients become less negative with fuel exposure in the Dresden reactor.
Thus, the limiting
--moderator coefficients occur at the end of the fuel cycle, at cold (680F) conditions.
Moderator coefficients were determined at ECC 6 conditions for an average core exposure of ^/9,900 MWD / Ton from 2-dimensional
.multigroup calculations.
Temperature coefficients were deter-mined as a function of core temperature at the EOC 6 average core exposure.
The effect of the presence of control rods wac con-servatively neglected in the temperature coefficient calculations.
The core effective moderator coefficients for Cycle 6 comply with the license restrictions and are given in Table 3 and Fig. 5.
C.
Reactivity Control The infinite multiplication factors under controlled and uncontrolled conditions are presented in Table 1 for each type of fuel to be loaded into Cycle 6.
Inspection of the table l
shows that the fuel and control rod characteristics' combine to produce the most reactive core at ambient temperature.
- Also, in Cycle 6, the effective reactivity 1cvel of the mixed core containing these fuel assemblies will decrease with exposure.
Thus, the most reactive state of the core will be at ambient temperature at the beginning of Cycle 6.
The license requirement for cold shutdown margin is that the core must be at least 1% suberitical with the worst control rod stuck out.when the core is in its most reactive state.
l Cold shutdown calculations were performed for BOL Cycle 5 and for Cycle 6 with a 96 assembly. relohl.
The Cfcle 5 data'. ___ _ _
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L provided a basis for checking and normalization of the results since experimentally derived cold shutdown data for BOL Cycle 5 are available.
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L The calculational results showed that with the core. reload configuration displayed in Fig. 2,.the license requirement for cold shutdown margin will' be met.
That is, with any peripheral rod stuck out of the core, the shutdown margin is in excess of 1%.
Greater shutdown margins are obtained with any of the
- non-peripheral control rods stuck in the out position.
D.
Burnable poison Behavior Type VI fuel-utilizes gadolinia as the burnable poison.
The gadolinia which acts as a highly self-shielded poison is I
contained in a single, removable rod in each Type VI assembly, having the same dimensions as other fuel rods.
The gadolinia is carried in UO2 (2.34 w/o enriched) pellets and has a con-
-centration within the UO per rod or.0.183 g/cm )*2 matrix of 1.74 w/o (55.0 gm gadolinia 3
The concentration and geometric con-figuration of gadolinia is such that it is effectively depleted in the high cross section isotopes, Gd-155 and Gd-157, after
~3000 MWD /T exposure.
As shown in Table 1, the reactivity worth of the gadolinia is ~6% 6k at 680F.
Because the proposed loading pat, tern con-tains only 83 Type VI.gadolinia bearing assenblics.on a scatter -
load, the effective poisoning of the core is-only ~1.1% Ak.
Evaluation of the gadolinia burnup in Type VI fuel indicates that with the proposed loading the core will have maximum-reactivity'at the beginning of the cycle.
That is, the positive contribution to core reactivity due to gadolinium burnup is always less than the negative contribution of fuel burnup throughout the core.
Thus, the net change in reactivity with
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time will always be negative.
E.
Nuclear Calculation Methods 1.
Fuel 'dssenb"1v liucic5r ~C5a~rheterNtINs The nuclear constants of the individual fuel assemblies are calculated using zero-dimensional reactivity and burnup formulations.
Neutron spectrum is recalculated before each discrete time step.
The multigroup Amouyal-Benoist method is used to calculate thermal flux self-shielding over the thermal neutron spectrum.
Effective burnable poison thermal cross sections are obtained from three dimensional Monte Carlo calculations performed at several burnable poison depletion states.
A non-lattice peaking factor is applied as a cor-rection to the homogenized model to account for different thermal. flux levels between the water outside the shroud and the water in intimate contact with the fuel'r5ds.
The non-lattice peaking factor is obtained from two dimensional dif-fusion theory calculations.
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- Experience with UO -Gadolinia is given in Section V.
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Nucicar constants derived from the above calculations include M2, k and number densities of all important isotopes g3,
as a function of burnup.
In addition, the power-dependent xenon worth, required in the three dimensional burnup calcu-lations, is computed as a function of burnup.
This calculation method was checked extensively against light water reactor cold critical and operating data, and against Monte Carlo calculations, with excellent agreement.
2.
Three-Dimensional Power and Burnup Distributions f
Three dimensional power and burnup distributions were cal-culated for Cycle 6 starting with burnup distributions projected for the remaining fuel at the end of Cycle 5.
In these cal-2 M,
culations nuclear constants such as koo and migration area, for each fuel assembly type were allowed to. vary as a function of the local burnup and localcoolant void-fraction, for all fuel assemblies, in coupled nuclear-thermal hydraulic reiterative calculations.
The local reactivity was also made to vary with local power generation reflecting the effect of fuel Doppler temperature coefficient and equilibrium Xenon concentration which vary with power generation.
In this calculation model, neutrons are absorbed within the node boundaries in which they are thermalized, since node dim-ensions ar'e in excess of 5 in.,which is several times the mean free path of thermal neutrons.
The fast neutrons are allowed to migrate to adjacent nodes and beyond until they are ther-malized.
Allowance is made for the presence or absence of a control rod at each node.
Reflectors are treated by calculating an albedo at the core boundaries.
Mininua critical heat flux ratios (MCHFR) ucre determined for all nodes froa power generation and quality data generated in the above calculations, taking into account the pouer-dependent coolant flou distribution, and the dependence of miniaua critical heat flux on local quality and nass velocity.
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The validity of the above calculation =cthod has been checked bycomparisonuithfuelcycle{gngghsandpouerdistributionsfron.
Fuel cycle 1cngths calculated Dresden I and an operating PUR by this acthod are uithin 5% of those actually achieved in Cycles 1 and 4 of Dresden I and also within 50 of the cycle icngth expected for Cycle 1 of the Trino Vercellesc (SELNI) PUR reactor.
Radial pouer distributions calculated for Cycle 1 of Trino at about 3000 MUD /MTU show excclient agreenent with experinental pouer distribu-tions derived fron Mnd6 activations.
The average deviation between calculation and experiment for each fuel assenbly is icss than 20.
In addition, calculated axial power distributions-for Dresden I, Cycle 4, are in.very good agreenent with flux wire neasurements.
3.
Moderator Coefficients Moderator coefficients at EOC conditions uero deternined for a generalized fuel assenbly representing Types III-F, V and VI fuel assenblics follouing burnout of burnable poison, at EOC core average burnup.
These calculations ucre performed using a tuo dimensional diffusion theory progran with cross section input derived from the
- cro dimensional reactivity and burnup calculations.
The effect of the presence of control rods was conservatively negiccted in the calculation of the coefficient.
The justifications for using the generalized fuel assenbly representing Types III-F, V, and VI fuel, and for neglectin3 the l
control contribution are as follous:
1.
At ECC 6, essentially all gadclinia is depleted fron all assemblics, and 650 of the core is comprised of Types III-F, V, and VI assenblics.
These three assenbly types are geonctrically sinilar and have the sane initial enrichacnt.
As shown in. Fig. 6, the calculated tenperature coefficient for the generalized c1ccent is a nearly linear function df exposure in the range of interest.
A scatter-loading refueling pattern is onployed which tends to average out Type I fuel'(1.45% initial enrichnent) given in Fig. 6{py the nuclear properties of the fuel in the core.
Data show the temperature coefficients for this fuel to be nearly the saac as a function of exposure as for the general-i:cd fuel assembly.
The~tenperature coefficients expected for Type III-B fuel assenblies should be no more positive than those shown in Fig. 6 for Type I or the generalized fuel assenbly.
This is because Type III-D (uhich has an l
initial enrichnent internediate betueen Type I and the generalized fuel assenbly) contains orbiun, a slow burning l
poisonul((htendstonakethecoderatorcoefficientsless positive Therefore, the entire core is reasonably I
represented as fueled entirely uith the generalized fuel type, for the purpose of calculating coderator coefficients.
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The contribution of controlled nodes to the overall coefficient is sm' ll, particularly at EOC conditions, a
because the fraction of the core controlled at EOC conditions is relatively small; also, the flux square importance weighting required.to obtain the overall coefficient would further reduce the contribution from the controlled nodes since the flux is depressed in these nodes by virtue of the presence of control.
Inclusion of the control contribution would tend to make the computed coefficients more negative (or less positive).
The temperature coefficient computed in the above manner at 680F as a function of exposure is presented in Fig.
6.
The slope of the line drawn through the calculated points is +0.40 x 10-3
.(ok/k)eff/oF per 1000 MWD /T advance in average core exposure.
This slope is in good agreement with that obtained from EOQ 4 and EOC 5 coefficients reported in previous submittals(3,41which yield a slope of +0.45 x-10-5(ak/k)eff/oF.
Also shown on Fig. 6 are the EOC average core exposures for Cycles 4, 5, and 6.
Since the advance in average core exposure from EOC 4 and EOC 5 to EOC 6 is small because the equilibrium cycle is being approached, only a small increcsc of +.3 x 10-0 in moderator coefficients from EOC 5 to EOC 6 would be expected.
Based upon the above calculations the temperature coefficient on a keff basis at 680F at EOC 6 is +4.1 x 10-3/oF, compared with
+5.5 x 10-5 I.subnittals(foF obtained from an extrapolation of data in previous p
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e References - Section III 1.
TID-7672, ANS Topical Meeting, Nuclear Performance of Power Reactor Cores, pg. 80, Sept. 1963.
2 Dresden Nuclear Power Station Description and Hazards Evaluation Report of Type III Fuel, June 15, 1963.
3 Dresden Nuclear Power Station Description and Hazards Evaluation Report of Type III-F Fuel, November 17, 1964.
4.
Dresden Nuclear Power Station Description and Hazards
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Evaluation Report of Type V Fuel, August 25, 1966.
5.
" Calculation of Fuel-Cycle Burnup and Power Distribution of Dresden I Reactor with the TRILUX Fuel Management Program",
L. Goldstein et.
al.,
Presented at 1967 Annual Meeting of ANS, San Diego, Calif., June 11-15,.1967.
6.
" Application of Reactor Operating Data to the Design and Fuel Management of Replacement Fuel", G.
Sofer et al.,
Presented at Conference on Reactor Operating Experience, Atlantic City, N.J.,
July 23-26, 1967 e
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SECTION IV - Thermal and Hydraulic Characteristics The thermal and hydraulic characteristics of a complete
-core consisting of different fuel types for Cycle 6 have been examined.
The fresh Type VI fuel assemblies attain the highest powers and therefore are the most limiting assemblies during Cycle 6.-
These assemblies meet the existing. license requirements for the Type III-F fuel which are:
1.
The ' maximum heat flux at design power shall not exceed 360,000 Btu /hr-ft2, 2
The minimum critical heat flux ratio shall be greater than 1.5 when calculated at 125% design power.*
The hydraulic characteristics'of the Type,VI fuel are essen-tially the same as those of the Type III-F assemblies.
There-fore, the flow through the Type VI assemblies will not appre-ciably change the existing flow patterns in the reactor nor adversely affect the operating stability.
Maximum allowable thermal-hydraulic characteristics for the Type VI fuel assemblies are listed in Table 4..
Hydraulic Characteristics Total recirculation flow through the reactor is a function of the pump supply characteristics and the total pressure drop around the recirculation loop.
Data obtained during Cycles 3
'and 4 show marked increase in core pressure drop from the beginning to the end of each cycle.
However, this increase in core pressure drop resulted in only a small reduction in total recirculation flow during the cycle.
Based on operating data for Cycles 3 and 4, the total flow used in the calculations 5 x 106 lb/hr and that used of.the beginning of the cycle is 25 6 at the end of the cycle is 24.5 x 10 lb/hr.
Only 92% of this total flow passes through the. fuel assemblies.
The core is divided into two flow zones to provide higher than average flow for the higher power central fuel assemblies.
The outer 208 assemblies have relatively small inlet orifices which result in flow rates less than the core average.
The contral 256 assemblies have relatively large orifices which provide flow rates higher than the average.
Actual flow through a given assembly depends on the pressure difference across the core, the power production in the assembly, and the assembly hydraull'c characteristics as a function of residence time.
The flow rates for the different fuel assemblies under various operating conditions have been determined by a computer program which calculates the pressure drop along a BWR fuel assembly for different flow and power conditions.
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- Based on " Burnout Limit Curves", APED-3892 and average quality of the cross section of the assembly.
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for friction losses, losses due to local restrictions (e.g.,
orifices, end plates, and spacers), hydrostatic head, and momentum changes.
pressure loss coefficients for various com-ponents of the Type VI fuel assemblies were measured experi-mentally and used as input for the computer program calcula-l tions.
The calculation method used has been checked against two phase flow high temperatura pressure drop data as shown in Fig. 7.
The flow rates of Types I, III-B, III-F, and VI fuel assemblies for a typical beginning of cycle operating condition p
are shown in Fig. 8 as a function of relative assembly power, f
These assembly flow rates are based on clean inlet orifices and a total flow rate of' 25.5 x 106 lb/hr.
The outer region assemblies have very little dependence on assembly power be-cause of their large inlet orifice pressure drops.
The Type VI fuel assembly has the same flow rates as the Type III-F fuel l
assemblics at the beginning of the cycle.
All of the fuel assemblies were examiued and cleaned at the end of Cycle 4.
It was noted that there were large crud
- deposits on the inlet sections of many assemblies which reduced
.the. orifice flow area and thereby increased the hydraulic re-
..sistance of the assemblies.
Examination of the pressure. drop.-
I data obtained before and after cleaning the fuel assemblies show an effective diametral decrease of 0.275 inches for the single hole orifices.
This decrease in flow area increases the inlet pressure' loss for a single hole orifice by a f actor of
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2.5 times at constant flow.
Similar analysis of the pressure drop data for the seven hole orifices used in the outer flow region show an increase in pressure drop by a factor of 9 times at constant flow due to crud buildup.
The crud effect is more pronounced in the seven hole orifices because of the large circumferential area available for crud buildup.
Using orifice loss coe,fficients calculated for end of cycle crud. conditions, the assembly pressure drops were recalculated for a number of flow rates and assembly powers.
The flow rates for Types I, III-B, III-F, and VI fuel assemblies at typical end of cycle conditions are shown in
. Fig. 9.
The a total flow rate of 24.5 x 10ge flows are based on 100% power, lb/hr, and a core pressure drop increase of 7 psi over the BOC value.
Table 5 gives a comparison of flow rates in different assembly types at BOC and EOC condition for a relative assembly power of 1.0 Type III-B and III-F fuel assemblies in the outer flow region have a reduction in flow rate at the end of cycle because of the large effect of crud buildup' on the seve,n hole orifices.
Type VI fuel assembly in the outer flow region shows a slight increase in flow rate be-cause the effect of crud buildup on the single hole orifice of Type VI fuel is less than the effect on the sev~n~ hole orifices.
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Fuel assemblics in the central flow region show an increase in flow rate over the BOC value which is a consequence of the flow rate reduction in the outer region assemblies.
The Type VI fuel assembly has a somewhat lower flow rate than the Type III-F at the end of cpcle because its clean orifice diameter is smaller than the III-F's and therefore the crud buildup produces a slightly larger hydraulic resistance.
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Thermal Characteristics The maximum allowable hot spot f actor for a Type VI fuel assembly is 3.44 which is the ratio of allowable to average heat flux (360,000/104,444).
This hot spot factor is composed of engineering and nuclear factors.
The engineering hot spot factor for Type VI fuel assemblies is 1.061 which is based on pellet enrichment (1.047) and rod diameter (1.013) tolerances.
Based on this engineering factor, the allowable nuclear hot spot factor is 3.24 (3.44/1.061).
The nuclear factor consists of radial, j
axial, and local power f actors that are functions oi' core location, control ro,d pattern, ' and fuel burnup.
At the beginning of the cycle, the local peaking f actor for a Type VI fuel assembly with i
gadolinia is 1.28 which allows a radial x axial f actor of 2.53 i
At the end of cycle, when the gadolinia is burned up, the local f actor becomes 1.21 which allows an radial x axial f actor of 2.68 Detailed nuclear analysis using three-dimensional core representation, allowing for residual fuel burnup distribution, non-uniform void distribution and planned control rod withdrawal i
I patterns, has shown that the nuclear peaking f actors can be main-tained below the designated limits through Cycle 6.
l Critical heat fluxes were calculated for all assemblies using the Janssen-Levy critical heat flux correlation (APED-3892),
the average quality over the cross section of the assembly and the assembly flow. rate.
These calculated critical heat fluxes were compared with the predicted operating heat fluxes based on the worst power distributions calculated for Cycle 6.
The mini-t i
mum critical heat flux ratios calculated.at 125% design were all
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above 1.5 which is the license limit.
i The actual critical -heat flux ratios to-be obtained during-Cycle 6 will be recalculated for all assemblies during the cycle fuel management and the control rod patterns will be re-adjusted where necessary to maintain ratios greater than the license limit of 1.5.
j Cri ti c al heat flux data were measured for a 16 rod 6 f t.
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tive of the Dresden Type VI fuel assemblies.
The minimuti crTtiu"
~
cal heat flux ratio at 125% overpower based on these data was also in excess of 1.5 The effect of the spacers was to increase the critical heat flux by several percent.
No credit was taken for this increase in critical heat flux in satisfying the 1.5 MCHFR at 125% overpower.
The centerline ' fuel temperature is less than the melting point of UO2 at the maximum allowabic heat flux of 450,000 Btu /
l hr-ft2 (19.4 kw/f t).
The thermal conductivity of fuel used in l
this calculation is based on data' presented in GEAP-4624.
The l
fuel-clad gap conductance was conservatively taken as 1000 Btu /
2 hr-ft
- F.
The peak fuel temperature for design conditions and l
125% design power are 3940*F and 4930*F respectively.
~
1 e-13 l
e SECTION V - Previous Experience with Gadolinia-UO9 and Background Information The fuel tube and pellet dimensions of the,Gd20 -UO2 fuel-3 poison rods are identical to those of the regular fuel rods of the Type VI assemblies.
The enrichment of the gadolinia bearing UO2 fuel is 2.34 w/o U235 The fuel-poison pellets are cold pressed and sintered and dished to the same extent as the UO2 pellets.
The gadolinia content'of 55 grams Gd 0 is equivalent 2
to 0.183 grams Gd 0 /cm3 of fuel or 1.74 w/o Go2 3 The Gd 023 23 g'
loading is uniform axially.
The nominal density of these pellets is 91.5% allowing a somewhat greater margin for swelling due to burnup than in the UO2 rods.
The use of gadolinia-urania fuel poison rods in the Type VI fuel is not the first use of this material in a reactor.
The following is a summary of experience with gadolinia-urania, including some background information.- All indications are that the gadolinia-urania rods will behave essentially the same as UO2 rods:
1.
Gd 02 3 and UO2 form a single phase, face-centered, cubic solid solution up to 40 mole percent Gd 023 in UO2-
~
Additions of Gd 023 to UOo result in a small contraction lattice (l).
However, the theoretical density of the UO2 of UO -Gd 02 3 solutions is lower than that of UO2 For 2
the low Gd u3 concentrations being used in Type VI fuel, 2
the temperature and time used to sinter UO2 pellet is sufficient to form a uniform solid solution of Gd203 in UO2 starting with a blend of Gd 02 3 and UO2 powders (4 4).
has been developed (4,5) 2.
A phase diagram of Nd 0 -UO2 23 and shows a decrease in the melting point of UO2 Of
.about 500C for a 5 mole percent Nd 02 3 composition.
Due to the similarity in chemical behavior between Nd 02 3 and Gd 023, the expected melting point change would be similar and not limit the use of UO2 with Gd 02 3 additions at the composition of interest.
Heating of Gd 0 -UO pellets with 1.65 w/o Gd 0 to 45000F 23 9
2 3 indicated no melting or significant densification of these pellets.
3.
The nuciear behavior of gadolinia has been well pre-dicted.
Design methods for calculating gadolinia effects have been developed.
These methods have been verified by the proper prediction of the gadolinia depletion rate in Type III-F fuel during Cycle 4, as observed from control rod movement in Cycle 4.
.Furthermore, minimum critical tests of fabricated Type VI fuel assemblics will be performed prior to loading'the Type VI fuel in the Dresden reactor.
These tests will include critical tests with and
' without Gd 02 3 and should yield a verification of
.~
.Gd 0 worth in the "as fabricated" assemblies.
23 14
7s s
4.
Gadolinia-urania rods are currently being irradiated in the-Dresden and Big Rock Point reactors.
In the Dresden reactor, four special Gd 0 -UO2 test rods in 23 Type III-F assemblies were placed into the reactor in Cycle 4.
These rods have received exposures of about 5000+ MWD /T as of the EOC 4 and have been retained in Cycle 5.-
A.large number of Gd 0 -UO2 23 rods are currently undergoing irradiation at Dresden as part of the Type V fuel assemblies loaded in Cycle 5.
k Six rods containing axially distributed gadolinia in the urania are being irradiated in the Big Rock Point e
reactor.
Sixty-three percent of the Gd 0 -UO2 pellets 23 used in these rods have gadolinia concentrations equal to or greater than that of the gadolinia in the Type VI assemblies.
The exposure of the Big Rock Point assemblies containing the gadolinia was over 9000 MWD /T as of August 1, 1967.
Thus, the use of gadolinia in orania in the Type VI fuel is technically justified as gadolinia has performed as expected in irradiations to date.
5.
The corrosion resistance of pellets fabricated from.
Gd O2 3 mixtures has been tested in deoxygenated and oxygenated water at 6800F and found to be somewhat superior {o)UO2 pollets fabricated without the Gd 023 additions 6 6.
The fuel supplier is well experienced in the fabrication of UO -Gd 02 3 pellets, having fabricated the UO2-Gd 023-2 pellets for the Pathfinder Core boiler fuel.
O e.
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/
e References - Section V
.S q,
-s 1.
ANL-6868, " Annual Report for 1963 - Metallurgy Division"[
. ',, \\
e page 158-161 m
r.
,i 2
Letter from Consumers Power Company, Jackson, Michigan to
- h\\
Dr.
R..L.
Doan, Director, Division of Reactor,I/lcensing,-
USAEC, Washington, D.C.,
dated Sept. 2, 1964.
(Re:
Amendment No. 14 to the Application for Reactor Construc-tion Permit and Operating License for the Big Rock Point-p Nuclear Plant (Docket No. 50-155).
7 '
'.l"W 3..
Exhibit 1 of Amendment of Appendix A, DPR-2 to permit.<
N
~
1966. 'rq(_ '
operation of Dresden 1 with Type V Fuel, August 25, tt
' y' U 4.
"High-Temperature Reactions of UO2 with Various Metal s
Oxides", NBS Circular 568, February 20, 1956.
v*
5.
D. Kolar, J. H. Handwerk, and R.
J.
Beak, "Investigatiorts '
in the System Urania-Neodymia", ANL-6631, December 1962.
k uw 6.
ANL-6868, " Annual Report for 1963 - Metallurgy Division",
page 195-196.
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g SECTIdifVI Safety Considerations N s-s
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p;There are no foreseen problems in the general area of 4
A safety associated 'with Cycle 6 operation with UNC Type VI ~
e-
- fue,l "e.,The Type-VI fuel' is, - in most respect,s, similar to
-i.
4 prov,ious"Dresden fuel, particularly Type III-F and Type V, E.
Groupkli andi thurefore does not result in significant changes in thd hazardGJevaluation:previously performed.-
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Table 2 Local Peaking Factors for Dresden Type VI and Type III-F. Fuel Basis -
No. Rods f
VI-I 35 1.17 III-F 36 1.31 35 1.27 6
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4 Table 3 Predicted Limiting Void Reactivity Coefficier.t During Cycle 6 t
Void Fraction Moderator Void Coefficient (Interior to Flow Channel)
Temp. (OF)
(ak/koff)/% Void EOC 6 EOd5 EOC 4 0
68
-1.0x10-4
-1.3x10-4
-1.6x10-4
' 10 68
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Table 4 Allowable Thermal-Hydraulic Characteristics Type VI Fuel-Assemblies Design Power 125?o Design Power Reactor Thermal Power, MW 700 875 Maximum Heat Flux, Btu /hr-ft2 360,000 450,000 Average Linear Power, kw/ft.
4.62 5.78 Maximum Linear Power, kw/ft.
15.5 19.4 Minimum Critical Heat Flux Ratio
') 1.5 Maximum Fuel Temperature 3940
'4830*
Corresponding to Maximum Li.near Power Generation, CF
- Fuel. melting point:
50800F s
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Table'5 Cycle 6 l
Comparison of Assembly Flow Rates Relative Assembly Power = 1.0 BOC EOC Ib/hr.
Ib/hr.
Inner Region Type III-B 61,000 69,000 Type III-F 60,500 69,000 Type VI 60,500 66,000 Outer Region Type I 26,800 19,500 Type III-B 42,500 24,500 Type III-F 42,500 24,500 Type VI 42,500 45,500 i
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6' 65 6E 67 66 d,E 70 h
t Fig. 2 -Dresden Nuclear Power 8: troa C' cle C Loading 1
6 J
I
-c
,~
~
Se 4
Control rod position (water) j.
(
l i
I I
I I
I l 1.08
__ __._ _ _] 1.27 I
l 1.27 1.08 1.23 1.28 l
T - - d -' - -
PPC I
I l
l l0 1.08 l 1.14_ _ i_ __ __ L _.9 6_.i_ __ _ _l. _
1.00 l 1.00 1.16 l
i I
l l 0.88 0.83 0.87 1.'01 1.27
.l 1.00__._ p _ __1._ _._ll_ _ _ lp _ __
l
- l I
I I
4 i
1.23 l 0.96 l 0.83 l 0.77 l 0.78 l 0.94 r--
p-p-
r-F0 8
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i i
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1.27 l 1.00 l 0.87 l 0.78 in vo2 7 _ _ 7 _ _ 7 _ _ y _ _ 1 0.9 5 PPC PPC I
1 1.
I i
1.08
- 1.16 l 1.01 l 0.94 l 0.95 l 0.85 l
l
~. -. ---
. Fig. 3 -- Relative Power Distribution in a Type VI Fuel Assembly (Void fraction = 0.20; temperature = 546 F; exposure = 0 Mwd /T)
I p*
a h
e
n Control rod position (water)
[
t lPPC PPC PPC I
I I
I l 1.01 i
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1.16
. T - -- d - - '-
PPC l
1 l
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l 1.10
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l l
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,']
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0.82 l 0.87 l 1.01
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l l
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l 0.82 l 0.83 l 1.03 r -
r - - F -
g 0.91 1.13
@e-F--
l l
1 l hole l
__. _ _l(inst.
pband ! _1.13 1.17 0.95 0.87 0.91 4
I l
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~
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1 I~ 1.03 l 1.13 l 0.96 1.01 1.10 l
l 1.01 l
I
~l Fig. 4 - Relative Power Distribution in a Type VI-I Fuel Assembly (Void fraction = 0.20; temperature = 546*F; exposure = 0 Aiwd/T)
....~.
O
\\
1 m
+ 6. 0
+ 5. 0 Q,
'\\
+4.0 N
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s
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EOC 5 "o
(previous submittal) x
+ 2.0 N
\\
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(previous submittal)
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s
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- 5. 0
\\'
- 6. 0
- 7. 0
- 8. 0 0
100 200 300 400 500 600 Temperature, F Fig. 5 - Dresden I - Calculated Temperature Coefficient vs Tem-perature - EOC 6 e
Calculated mixed core EOC 5 Calculated mixed core EOC4 Calculated Type I l
/
Generalized ass'y.^I ph?[f
/
lyl/
Inr, v, v1(
s 4'g l
Measured Type,I y
l l
+ 2.0 l
pr b
O=
0 E
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l l
2 D
i
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l I
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I
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I l o o::
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5
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l g j
- 8.0 A Data from previous hazards j
!__I y g
l l j g
submittals R
O Experimental data, BOC4 and BOC 5 l
l 2
[-0.0 (Refs.: DNPS 80-5-G5; DNPS 106-5-67) l g
j P
A Data for Type I fuel (TID-7072, p.80) l
-14.0 O
2 4
6 8
10 12 14 Exposure, Mwd / ton X 103 Fig. 6 -Dresden I Temperature Coefficient vs Exposure at 68 F
m,
,~
.~
~
4 20
~6 G = 1.8 x 10 6
Calculated data O Experimental data 16 8
2 G = 1.4 x 10 lb/ft -hr 3
a 12 S'
u O
'o 8
o G.= 1.0 x 10 f
h U
o 8
fo6 4~
o A
o G = 0.8 x 108 o
c o
8 4
V
=,0.5 x 10 _
MO 'O
~
0 0
0.1 0.2 0.3 0.4 0.5 Exit Quality l x
._ Jig. 7 - Pressure Drop Across 6-ft Long Heated Burnout Test Section - with Spacers e
4
l 3
-3 Inner Flow Region 70 Type I - none Type IIIB
't Type IIIF
%g g
60 B
"o 50 e7%Z
,S 40
\\
Outer Flow Region N
b j
Type I g
Type IIIB J
_ Type IIIF 30
' Ty]e VI 20+
g 0
0.6 0.8 1.0 1.2 1.4 1.6 Relative Assembly Power
.. Fig. 8 - Assembly Flow Rates - Cycle 6 - Beginning of Cycle -
8 100% Power (Total flow.= 25.5 x 10 lb/hr) e#
e u
em
'N l Inner Flow Region Type I - none Twe mB Type IIIF 70 n
/ Type W w
60
- S "o
50 ci d
c:
4
.'S 40 N
b Outer Flow Region
.o TheI E!
Type IIIB 30 f
Type IIIF
~
Type VI
/
/
20 o
0.6 0.8 1.0 1.2 1.4 1.6 Relative Assembly Power Fig. 9 - Assembly Flow Rates - Cycle 6 - End of Cycle -
0 100?o Power (Total flow = 24.5 x 10 lb/hr)
=
a e
.B m
O.
- l m
m,
~~
n rs p
^
1829 D ATE OF DOCUu(NT:
DATE RECElvtD NO.:
~~Feou:
c - ith Edisen Co.
9-14-67
%1M7 3076 M W. h t.,
eq0, M.
LTR.
utwo:
nEeonT:
oTHER:
(Joimi H. Hughes)
I To:
ORIG.;
CC:
OTHER:
Dr. Peer A. Morrie-DEL 5
ACTION NECE$MAY
]
CONCLIRE**CE
]
DATE ANSWERED:
NO ACTION NECESSARY Q
COMMENT C
BY:
CtAssir.:
nosT orriCE eit! code:
50-10 703 U
.Ec so.
DESCRIPTION: (Mute Be Unciotnefiedl ggFggp?D TO DATE RECEIVED av DATE Ltr req. Proposed Change No. 14 to CL 9 ' 13I DPH to make changes to Items Nei, 2 Yl**FT (J & 4.d cys ree'd-notarized 9-14-67) w/9 qs- - FM AC'T'M ENCLOSURES:
INFO CTS 70:
H. Pries at 3 Laff Morris /Doyd skevkolt
.. n,n 1-.=ppi rue #
cube /Levine 4 ;...
~Q s.J RE=Aans:
i;ist:1-rorsEI Tile s c-1-ac Poa 2-Complianee DO Nqt I?c, move s3 76 m
1-oce u.s. uoMIC ENERGY COMMISStoN MAIL CONTROL FORM roau AEC szes.
- S.6 01 W U. S. GoggenwEnf Pei%fiNG OFFICE: 1966-235 $13 h
l l.
Iw
.