ML19345E602
| ML19345E602 | |
| Person / Time | |
|---|---|
| Site: | Big Rock Point File:Consumers Energy icon.png |
| Issue date: | 08/24/1977 |
| From: | Bixel D CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.) |
| To: | Desiree Davis Office of Nuclear Reactor Regulation |
| References | |
| NUDOCS 8102050269 | |
| Download: ML19345E602 (27) | |
Text
{{#Wiki_filter:. \\ \\ C ' C0fiSumBIS .r P0CCr j Company o.n.r. ome.. 2 2 w.a we e.n Av.no..s.cw.oa u.cnee.n.on. A,.s coo., si? 7esjg~ ; l _c. August 24, 1977 'I,d")f} of' t' .\\ LE - t, 4 ;.. 6.- y-- r p0G'# aP. H W QQJf v g; e c u Director of Nuclear Reactor Regulation Att: Mr Don K Davis, Acting Branch Chief Operating Reactor Branch tio 2 U3 Huclear Regulatory Commission Washington, DC 20555 DOCKET 50-155, LICEISE DPR, BIG ROCK POIliT PLANT - CYCLE 15 RESPONSE TO LETTER DATED AUGUST 3,1977 By letter dated August 3, 1977, Consu=ers Power Company was requested to provide additional information concerning the Cycle 15 Reload Application for the Big Rock Point Plant, dated April 15, 1977 The requested information is provided as an enclosure to this letter. (- , ( Q p 'jh b SJ 'x% L David A Bixel Nuclear Licensing Administrator CC: JGKeppler, USNRC W/oposCe2&J W#
Ite 1. Trovide core maps of the planned and of the "as loaded" core, including location of the irradiated fuel assembliea with F, modified F, G and G-lU type fuel rods.
Response
A cap of the planned core loading for Cycle 15 is provided in Figure 1. Because no failed fuel was found as a result of fuel sipping and inspection, the actus1 core loading vill be the same as the planned core loading. The following in-for=ation relates the map designations with the various irradiated fuel assembly types: Fuel Type Man Desienations F i Modified F 2 0 3 G-1U h G-3 5 i o 1 3
b 4 l Item 2. Provide the design conservatism factors (DCF's) used for the reference cycle significant reactor kinetics para =eters such as void and Doppler coefficient, scram reactivity, etc. Do the nominal Cycle 15 parameters used in your analyses include the same DCF's? If not, apply acceptable DCF's to the Cycle 15 values so that a comparison on an equivalent basis can be made.
Response
Design conservatism factors, as required by current acceptance criter'a, for the 4 i significant kinetrics parameters were not used in the reference tri tent analysis. However, it is believed thr.t sufficiently conservative kinetics parameters were e= ployed. The two most important kinetics parameters for the Big Rock Point Plant for the "at power" transients (such as turbine trip without bypass) are the Doppler coef-ficient and the void coefficient. The Doppler coefficient used in the reference -5 f transient analysis was -5 k2 x 10 Ak/k/ F, which is 30% less than the Doppler -5 i coefficient predicted for Cycle 15 (-T.06 x 10 Ak/k/ F). Th.'.s is believed to be sufficient design conservatism. The void coefficient used for the reference cycle analysis was approximately .22 ak/k/ unit void, which is' 30", more negative than the most negative void coefficient predicted for Cycle 15 (.166 ak/k/ unit void). A6ain, this is believed to be sufficient design conservatism. The scram reactivity function is a critical kinetics paraceter for fast, ac power, transients (like turbine trip without bypass) for jet pump boiling water reactors. However, it is not a critical parameter for Big Rock Point. Although the pre-dicted "at power" scram function for Cycle 15 does not result in as rapid a I . reactivity insertion rate as compared to the scram function employed in the analysis, sensitivity studies for the turbine trip without bypass incident in-dicate that even a two-second delay in the time of reactor trip results in only a 3% increase in peak heat flux (a reduction in MCHFR of about 0.1) and a 60 psi increase in peak system pressure, when assuming identical scram insertion characteristics. Considering the fact that following a scram initiation signal the control rods are 905 inserted in at most 2.5 seconds by Technical Specifica-tions and nominally fully inserted in less than 2 seconds, the sensitivities in-peak heat flux and pressure noted above are a very conservative indication of I u 2
i r I the relative insignificance of the differences in the scram insertion curves. Further, it is felt that the proposed conservative MCEFR limit to be submitted more then compensates for this effect. e 9 7 l l \\ t 3
9 Item 3 The scra= reactivity strongly effects the plant response for fast transients such as a loss of load. Provide a comparison of the reference cycle scram reactivity function and the Cycle 15 scram reactivity function for B0C and EOC. The comparison should be on the same basis (ie, either nominal or including DCF's). Res ponse Hot standby scram reactivity functions at beginning and end of Cycle 15 are compared to the reference hot standby scram reactivity function in Table 1. A similar ecmparison is =ade for the hot full power scram reactivity function in Table 2. The calculated scram functions for Cycle 15 are based on the design scram insertion time of 2 5 seconds for 905 insertion, and do not include any additional design conservatisms. The reference hot standby scram function in table 1 is the scram function actually used in the rod drop transient analyses, and includes no other design conservatism. It should be noted that measured times for full control rod insertion on, scram have historically been less than 2.0 seconds based on surveillance test results. It is concluded, however, that this approach is both accurate and consistent since both the reference cycle and Cycle 15 curves are developed under the same assumptions. The reference hot full power scram function given in Table 2 is the scram fune-tion actually used in the plant transient analysis presented in the Final Hazards Su==ary Report for Big Rock Point. i h
c s TABLE'1 4-Hot Standby- _Scras Reactivity Insertion Rate Reactivity Inserted (Fraction of Total) ' Time BOC 15 EOC 15 Reference Analysis 0 0.0 0.0' O.0 57 .115 .002 .007 78 .253 .021 .021 99 .382 .045 .0h1 1.19 .459 .076 .063. 1.ho .49h .119 .100 1.61 .535 .181 -.125 .1.82. .589- .266 .170 2.03 .660 .367 .2h5 2.28- .771 .527 .h50 2.43 .899 .765 .625 ~! 2 73 -1.0 -1.0 -1.0 Note: These scram curves were calculated on a consistent basis assuming 90% of full. control rod insertion in 2.5 seconds; the nominal time required for full insertion is less than 2.0 seconds based on historic surveillance test results. i I f o 5 1 . ~.
1' f~ TABLE 2 l Hot Full Power Scram Reactivity Insertion Rate Reactivity Inserted (Fraction of Total) ' Time BOC 15 ECC 15 Reference Analysis _0 0.0 0.0 0.0 .03 30-57 .05h .025 .3h 78 .110 .oh9 .h9 99 .163 .070 .59 1.19 .212 .089 .66 ~ 1.40 .280 .121 .715-1.61 .365 .169 .765 1~ 82 .h69 .237- .81 2.03 .592 .362 .85 2.28 .777 .500 .895 s 2.h3 .979 .807 .9h -2.73 -1.0 -1.0 -1.0 Note: These scram curves were calculated on a consistent basis assu=ing 90% of full control rod insertion in 2.5 seconds; the nominal time required for full insertion is less than 2.0 seconds based on historic surveillance test results. ) 6 ,= 6 f a
Ite= 4. Provide the results of the analysis of the loss of recirculation pump that you discuss in Paragraph 7.1.h of your April 15, 1977 submittal.
Response
Listed below are the results of the analysis of the two-pump trip from rated power for the limiting fuel bundle at the " worst" time in core life: Core Average Time Heat Flux Core Flow MCHER 0.0 1.0 1.0 h.50 0 50 988 .807 h.21 1.00 948 .676 h.02 1 50 .888 582 3 96 2.00 .8h8 511 3 74 k.00 .680 343 3.13 6.00 580 .290 3.6h 8.00 5ho .290 h.89 10.00 510 .290 5 96 The worst time in core life was found to be end of life when the highest bundle peaking factor (at any time during Cycle 15) occurred. ( 7
~l s Item 5 Describe the fuel loading error calculation performed for your transient analyses. Provide the lowest MCPR resulting from the misloading and justify that the worst case has been identified by stating other cases -. calculated, and demonstrating qualitatively that no worse case should exist. Resconse ' Eight independent fuel loading error cases were studied for Cycle 15 The results of this study are sum =arized in Table 3 The bundles referred to in Table 3 are situated in the planned loading pattern as-shown in Figure 1. The-G999 bundle is a bundle from new fuel storage which is not intended for use in Cycle 15 and is used to' emulate the beginning of life characteristics of a G-3 assenbly. The peaking factor and overpever MCHFR quoted for each case on Table 3 are for the vorst time in life for Cycle 15_ assuming a full cycle de-pletion with the misplaced bundle (s). The study was conducted to determine the worst assembly misloading error and the resultant core MCHFR. The study assumed the insertion of new assemblies (either from the core outer row or from new fuel storage) into limiting MAPLHGR and/or MCHCR locations on the "as planned" core. There is reasonable confidence that this approach bounds the worst case misplacement. The vorst case misplacement (Case 8) was found to occur when a new bundle (G303) from the outside of the core was interchanged with G207, a mixed oxide bundle in its second cycle of depletion. Even at 122% overpower, this worst case mis-loading would result in an MCHFR of only 2.4T7 Other less limiting cases which were considered are as follows: Case Description 1 A new assembly (G302) from the outside row was interchanged with G05, one of the limiting MAPLHGR assemblies in Cycle 15 4 2 A new assembly from new fuel storage (G999) was inserted for G05 3 A new assembly from new fuel storage (G999) was inserted for Gil such that four new assemblies would be adjacent to each other. ( 8 + e
as y
.. _ _.=. .=, .f. ~ Case Description h . A new assembly from new fuel storage (G999) was inserted for G03 which is near two control rods in the center.of the core. 5 A new assembly (G303) from the outside of the core.was interchanged with a rodded twice burnt assembly (G13) in the middle of the core. 6 ~ A new assembly (G304) from the outside was interchanged with a rodded assembly (G08) near a high radial peaked assembly (G219). 7. A new assembly (G999) was inserted into the limiting MCHFR location for the properly loaded core (G207). E. 4 a 2 4 9
TABLE 3 Fuel leading Error 122% overpower Case Assembly Interchanges Radial X Axini Peak
- MCHFR*
1.81 2.663 Planned Gore Loading 1 GOS ',G302 1 91 2.627 2 G999** ",GOS 1.88 2.617 3 G999"' ',Gil 1.86 2 525 h 0999** ' G03 1 96 2.489 5 G303 " G13 1 93 2.548 6 G304 ",G08 1.86 2 521 '( G999** ",G207 1 94 2.h84 8 Gr07 e-- G303 1 98 2.477 "At worst time in cure life.
- G999 is a G-3 bundle from new fuel storage that is not intended for use in Cycle 15
P l Ites-6. Paragraph T.. 6 of your reload submittal discusses the " start-up event" t
- and the relative values of the key nuclear parametern. The staff vill.not accept a "linearization" of the peak fuel temperature variation with Doppler coefficient change without further substantiation. Accordingly, justiff your asau=ption of a linear approximation, or reanalyze this event using the relevant nuclear parameters.
-Response Thestart-upeventwasreanalyzedforG-3typefuelusingEheRELAPhcomputer code. A two-heat conductor, one control volu=e and one flow junction model was esployed. One heat conductor was used to represent the average core, while the core hot spot was represented by the second heat conductor. RELAPh-requires that at least one control volume and one flow junction be modeled in order.to perfom-a transient calculation. However, because this event is terminated before any significant amount of energy can be transferred to the coolant (it is conservatively assu=ed in the analysis that the fuel heats up adiabatically), an accurate representation of the reactor coolant system is not necessary. However, the core coolant temperature must be properly defined to ensure proper definition of the initial fuel stored energy. Important input para =eters for this analysis are listed below: i (1 ^ K W Naore, W H Rett N, "RELAPh - A Computer Program for Transient Thermal .Hydreulic Analysis," ANCR-ll27, Rev 1, March 1975 11
t. +
- (.'
Cold Hot -Nominal Start-Up -Start-Up Cycle 15 Value Maximum Reactivity Insertion, %Ao 39 k.2 2.03 i p Doppler Coefficient, Ak/k/*F -0.8 x 10 ' -0.8 x 10 ' 0 95 x 10 ' ~ ~ ~ Local Peaking Factor, F h.0 4.0 39 q ~1 - B/1* ,s 183 183 183 Init'ial Fuel / Coolant Temperature, F 130 560 Fuel Rod Outside Diameter, Inches hh9 ~ Cladding Thickness, Inches 3k Fuel Pellet Outside Dia=eter, Inches 3715 Pellet-to-Clad Heat Transfer 2 Coefficient, Btu /h/ft /oF (h ) <<1 A ver/ lov h was chosen (adiabatic assumption) in order to maximize the fuel pellet tenperatue rise at. the hot spot. Transient results ror both the cold start-up and hot start-up events are dis-cussed below. For the cold start-up event, nuclear flux and hot spot fuel -temperature are plotted as a function of time on Figure 2. The peak nuclear fluy is about 125 times rated, and the hot spot fuel temperature '.ncreases to l 3,270 F at 0.25 second after the flux peak. At this,same time, control rod insertion due to reactor trip on high flux (125% of rated) begins. Heat trans- ?i fer to the primar/ _ coolant would then begin to cool the core. 1 Nuclear flux and hot spot fuel temperatures are plotted as a function of time on Figure 3 for the hot start-up event. In this case, peak nuclear flux is about 176 times rated, and the hot spot fuel temperature at 0.25 second after the flux i peak is 3,950 F. It should be noted that in both cases the transient has been essentially terminated when reactor trip is initiated and, hence, scram has little effect on the transient. In both cases, the hot spot fuel temperature is significantly below the melting temperature for UO and, as such, results in 2 acceptable consequences. (, 12 t* +,r y .-m y y.,g w g --*? =r=+ erflerw---r w +-*'"=s f* =>w
i l ( t Ite= s 7 Propose changes to the Technical Specifications which will have the effec
- of limiting the reactor core stability decay ratio X2/Xo to no greater than 0 50. Provide a thermal-hydraulic stability analysis which demonstrates that during limiting reactor core stability conditions, as proposed by the specifications, the decay ratio does not exceed 0 50.
Or, provide analyscs to show recirculation pu=p start-up from natural recirculation doesn't cause too severe a reactivity transient.
Response
Reports concerning the stability and transient perfomance of the Big Rock Point reactor are: 1. GEAP-3795, " Consumers Big Rock Point Nuclear Power Reactor Stability Analysis," 1961. 2. GEAP-h567, "Preoperational Power Stability Analysis of the Consumers Big Rock Point Plant," 196h. 3 GEAP-hh48, " Control Rod Oscillation and Transient Pressure Tests - Big Rock Point Boiling Water Reactor," 1966. k. APID-4230, " Big Rock Point Plant Control and Transient Perfor=ance Tests," 1963. 5 GEAP Lh96, " Core Perfor=ance and Transient Flow Testing - Big Rock i Point Boiling Water Reactor," 1965 References 1 and 2 address reactor core stability using analytical techniques. Reference 3 addresses core stability based on a number of control rod oscillation and transient pressure tests.which were run at Big Rock Point between 1963 and 1965 References h and 5 address a nu=ber of transient conditions, including recirculation pu=p trip and recirculation pump start-up from natural circula-tion, based on actual plant tests htween 1963 and 1965 The documentation above (in particular Reference 3) represents significant proof of the stability of the Big Rock Point reactor and is further justified by 15 years of operation with no evidence of core instability. In general, the reports discuss the stability of the reactor as a whole and =ake the conclusion that the reactor is stable if operated within technical specifi-cation limits. It should be noted that this also included operation on natural circulation which is no longer permitted at Big Rock Point. Therefore, the { reactor should continue to be stable with the Exxen Nuclear Company (ENC) reload 13 f 4 -~ -- - , + ~, - - - -. ,--.,n--
b 1 fuel, provided t'2at hydraulic compatibility between fuel types exists. As such, a generalized '. ydraulic stability analysis of.the ENC reload fuel is provided in Attachment A as part of the justification. This study demonstrates that the fuel design does not promote parallel-channel flow instability within tech-nical specification operating limits. References 1 and 2 evaluated the stability of the Big Rock Point reactor using analytical techniques and concluded that for normal modes of operation '(ie, both recirculation pu=ps operating) the reactor would be stable with' both 56-and 8h-bundle cores. Reference.3, on the other hand, provides test results docu=enting the stability of the reactor during operation with natural circulation as well as with one and two recirculation pu=ps operating. Since all work in the area of boiling water reactor stability has indicated that the natural circulation mode is_the operating mode most prone to instability, it is prudent to discuss those tests and results. The natural circulation 1,ntrol rod oscillation tests (Test No 28) were run for the 8k-bundle core at s ther=al power of 128 MW', a primary pressure of 1,370 psia, an inlet subcooling of 419 Btu /lbm, and a void coefficient of about t .125 ok/k/ unit void. The reactor was shown to be stable for this adverse-operating condition over a control rod oscillation frequency range from 0.05 to 3.0 cycles /second. The measured gain and phase margins for this test indi-cated even greater stability margin than did the pretest predictions. Referring to Table 4-1 on Page k-72 of Reference 3, it is noted that both pretest predic-tions and the test results indicatei that natural circulation is the most severe operating mode with respect to core stability. However, the reactor was found i to be stable in this mode of operation (ie, Gain Margin = k.7 dB, Phase Margin = .50 degrees). Again, it should be noted that the natural circulation mode of operation'is no longer permitted at Big Rock Point. The tests discussed in Reference 3 vere run for operating conditions (ie, pri-mary pressure, inlet subcooling, void coefficient core flow, core power) which are' not appreciably different from those presently existing at Big Rock Point. 4 Probably the only significant changes that have taken place since the time that the tests were run are changes in fuel design. ( 1h ,m
4 ^ \\ - l-The applicable tests were run when the 8h-bundle fuel loading was primarily ec=- . posed of 12 x 12 rod array UO fuel with a rod outer diameter of 0 388 inch, a 2 . nominal pellet-to-clad gap vidth or v 9025 inch and a clad thickness of 0.019 - L 2h. Present-day fuel is primarily composed of 11 x 11 rod array fuel with a rod OD of 0.kh9 inch, a nominal pellet-to-clad gap vidth of 0.03h5 inch, and a clad thickness of 0.03h inch. The sum total of these differences indicate a - lower pellet-to-clad gap conductance and a longer fuel time constant for present-day fuel as compared to-the fuel type tested. These differences, however, have been shown to result'in improved stability margins in recent work" by General Electric (
Reference:
NED0-21506,' " Stability and Dyna =ic Performance of the General Electric Beiling Water Reactor," 1977). Based on the above discussion of total tystem stability and on the reload fuel channel hydraulic stability study provided in Attachment A, it is concluded that the -Big Rock Point reactor can be operated safely within present technical. spec-a ification limits with no danger of reactor core instability. t' 1 4 1 I 4 i t 15 i d ne n , +. m ,=< a -,r. -n m, ,,s -. ~ -e-------,-- s'. s
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i Item 8. The Hench-Levy correlation used in your transient analyses has not been accepted by the staff for further license amendment applications since it was based on data generated utilizing a uniform axial heat generation rate. Such data has been shown to be nonconservative for application to rods having c nonuniform axial heat generation rate, such as Big Rock Point fuel rods. It is our position that you must provide results for the transients analyzed in terms of an acceptable correlation. The XN-2 correlation (as described in XN-75-34, Rev 1, "The XH-2 Critical Power Correlation," August 1,1975) is an acceptable correlation for use on Big Rock Point. Respor 1, A study has been conducted to determine a synthesized Hench-Levy minimum critical heat flux ratio (MCHFR) limit which ensures that the minimu= critical power rate (MCPR) limit of 1.32 for the XH-2 correlation vill not be violated at Eig Rcck Point. The results of this study and applicable Technical Specifications changes vill be forwarded under separate cover. l ( 16
~. /> -- 4 J Item:- 9 State your schedule _ for submitting to NRC a brief summary report of physics start-up tests. This report should include'both' measurement and predicted -values. If the difference between the measured and predicted value exceeded the acceptance criterion, the report should discuss the actions that were taken and justify the adequacy of these actions.
Response
A start-up report vill be submitted in accordance with Big Rock Point Technical Specification 6.9.1. .5 j i t a b J -I i w i i l 1 4 17 I I 1 i
ATTACICENT A M CREEL IfCRAULIC STABILITY i In order to estabiish channel hydraulic stability for the ENC reload fuel in the Big Rock point boiling water reactor, a stability analysis was perfonned to cover the expected operating range of assemblies in the reactor. The occurrance of parallel channel flow instability is characterized by a pressure drop versus mass velocity curve which demonstrates a nonpositive slope over a portion of *.he mass velocity range.U) An analysis was perfonned for the ENC reload fuel using the XCOBRA com-puter code ( ) to detennine pressure drop versus mass velocity for the range of expected operatirg conditions. The operating conditions for the hot assembly are given in Table I. The analysis considered the effect of assembly power, mass velocity, inlet enthalpy, operating pressure, and orifice zoning. The results are shown in Figures 1 to 3. Figure 1 shows the flow stability calculation covering the anticipated operating range of pressure. As indicated in the figure, the slope of all pressure curves remains positive over a wide range of assembly flow rates and indicates no point of flow instability. The results of Figure 2 indicate that no flow instability occurs for assembly power up to 6 megawatts. As seen from Table 1, this is well above the expected power of the hot assembly at 122% overpower. Figure 3 shows the results of the flow stability analysis considering the effects of inlet enthalpy. This analysis was performed assuming a change of inlet enthalpy which conservatively envelopes any expected variations as shown in Table 1. Again there is no indication of flow instability for the ENC reload fuel. (
( t i It is therefore concluded that for the operating range of the Big Rock Point reactor, sufficient margin exists to protect against the occurrence of parallel channel flow instability for the ENC reload fuel. As indicated in the reload G design report,(3) the ENC fuel is hydraulically ccmpatible with the existing fuel. 4 f 5 a 4 4 i 2 I ~ 3 m er e e- -r
m e t -t, --? --t-N7*- V V
// 35 PRESSURE (psia) // //
- 1200-INilER ORIFICE 20tlE
/ f A1350 // 15
- 30
/ / 012001 / / 61350 OUTER ORIFICE 20flE // l !/ 25 01500 // // // 20 // // // 3 // E // {l5 // /- 7 / /l // 10 // s /j' / 5 0 0 0.25 0.50 0.75 1.00 1.25 1.50 6 2 G (10 lb/br-f t ) FIGURE 1 BIG ROCK P0lflT HYDRAULIC STABILITY At!ALYSIS AT 3.5 MW, Hin = 575 Btu /lb [
I 1: / 35 - / /! l // POWER / MEGAWATTS f l/l I A3.5 INilER ORIFICE Z0t!E / f ~ g6.0 ,1 /l p, l // l l/ l o2.01 / / 25 A 3. 5 0 UTTER ORIFICE ZONE j j 06.0 I /; / // / // /,I/ l I 20 I lj i l // //f! / // 15 /// S II' / // g /l/ l'l //,l l/ 10 pp,/ 'tl '/ A 5 i t 0 0 0.25 0.50 0.75 1.00 1.25 1.50 6 2 G (10 lb/hr-f t ) FIGURE 2 BIG ROCK POINT HYDRAULIC STABILITY A'lALYSIS AT Hin = 575 Stu/lb,1350 psia e e
i 35-IriLET ENTHALPY (Btu /lb) / [/ e546] / // / // i A575 > INNER ORIFICE ZONE / // 3606j / / // 3 f /[ I / / 30 / 05467 / /j f
- /
A575 h OUTER ORIFICE ZONE / / 'f 0606j ,/,/ j ,I 25 f',/ ,/ /! i,! l/ // 20 / / /- f l// 15 j g / E l' 10 / / / / / 5 h 0 0 0.25 0.50 0.75 1.00 1.25 1.50 6 _g (10 lb/hr-ft ) FIGURE 3 BIG ROCK P0Illi HYDRAULIC STABILITY AflALYSIS AT 3.5 MW, 1350 psia
L'. i i TABLE 1 BIG ROCK POINT OPERATING CONDIT0NS(3) Pressure, psia 1350 Inlet Enthaly, Btu /lb 569.5 Hot dssembly Power @ 1.45 radial and 122% reactor overpower, Megawatts 5.05 Hot Assembly Flow 0122% overpower, lb/hr 123,100 Hot Assembly mass velocity @ 122% overpower, 106 lb/hr-ft-0.756 2 Inlet Subcooling, Btu /lb 22.8 - Subcooling considered, Btu /lb 0 to 46.8 i i s e
5 - {. ~ REFERENCES 1. " Study on Distribution of Flow Rates and Flow Stabilities in Parallel Long Evaporators", K. Akagio, et al., Japan Society of Mechanical Engineers, Vol. 14, No. 74, pp. 837-846 (1971). 2. Patten, T. W., "XCOBRA Code User's Manual", XN-NF-77-43, (1977). 3. XN-72-17, " Design Report Big Rock Point Reactor Reload G Fuel", (1972). 1 1 4 I
a..:.. n.n u.... a.. n. m t... si v. .a...au -1_.- l U.S. NUCLEAR GEGULATORY J MISSION DOCKET NUM1Eil, 2RC P ::M 195 6d-/[6
- 2 78)
"""""I" NRC DISTRIBUTION nn PART 50 DOCKET MATERIAL TO: FROM: DATE OF DOCUMENT Consumers Power Company 8/24/77 ackson, McMgan DATE RECEIVED Mr. Don K. Davis David A. Bixel 8/29/77 i dLETTER ONOTORIZED PROP INPUT FORM NUMBER OF COPtES RECEIVED 700tGINAL MNCLASSIFIED CCOPY /g g(( .g',,/' p3-77 /Dy Consists of requested additional infor,ation concerning the Cycle 15 Reload Application.. tA I CfJG_JLED3ED r ' sU lk-.ati.*v hd (1-?) (26-P) PLANT NAME: Big Rock Point Plant RJL 8/29/,77
- EMM SAFETY FOR ACTION /INFORMATION I BRANCH CHIEF: (7)
[d(//S l l I i I I INTERNAL DISTRIBUTION L REG, FILE J l L 'NRC PDR l l IEE (2) OELD t l HMiAUER CHECK j l STELLO EISENHUT SHAO 3AER j BUTLER GRIMES J. COLLINS l 1 i i l 1 EXTERNAL DISTRIBUTION CONTROL NUMBER I 1 O LPDR: 0//4f1ff/dJy) mig /f. TIC II,/1 7 "I 'f j M MsIC 16 C'IS ACRS SENT CATEGORYN ' ) -tiesa so a cA, t L ,}}