ML19290D914
| ML19290D914 | |
| Person / Time | |
|---|---|
| Site: | Fort Calhoun  |
| Issue date: | 02/25/1980 |
| From: | OMAHA PUBLIC POWER DISTRICT |
| To: | |
| Shared Package | |
| ML19290D913 | List: |
| References | |
| NUDOCS 8002290324 | |
| Download: ML19290D914 (24) | |
Text
i ,
.: . 0 : 7T:::G C01 DITIO: 3 FOR OPERATIO:i 2.10 ; tor Core (Continued) 2.10.J 9e w.itity Control k atema and Core Physica Parameters Limits Applicability Appliec to operation of ecntrol element assemblies and ronitor-in m of celected core parameters whenever the reactor is in cold mi hot chutdown, hot standby, or power operaticn conditionc.
Ob?octive To enuare (1) adequate shutdcwn targin following a reacter trip, (2) the ':TC is within the limits of the cafety analysic, and (3) control element ascembl', operation ic within the limits of the setpoint and safety analycic.
recilication (1) Shuticwn tariin with Te _old >21CCF
'. nenever the reactor is in het chutdown, hot ctandby or pcuer operation conditions, the shutdown rarcin shall be
>3.03 Ah/k. '.with the shutdown targin <3.05 Ak/k, initiate and ecntinue 'coraticn until the required shutdcun margin is achieved.
(2) Shutdown I:argin with T olri e _1210CF
'eenever the reactor is in cold chutdcwn conditions, the chutdown cargin chall be >2.C5 Ak/k. With the shutdcva c.argin <2.03 Ak/k, initiate and ccntinue boration until the required shutdown r.arcin is achieved.
( 3) 'cierator Terrerature Coefficient The .oderator temperature coefficient (MTC) shall be:
i
- a. Leon pocitive th'tn +0.2 x 10-" Ap/CF includini; uncer-tainties for power levels at or above 8% of rated power.
i
- b. Lecc pocitive than +0 5 x 10-'4 no/ F including uncer-tainties for power levela below 80% of rated power'.
- c. ftore positive than -2.30 x 10 h ggjoF including uncer-tainties at *tted power.
..ith mlerator te::, orature coef ficient confirmed cut-cide anj cne of the above nr.itc, chan"e reactivit'. centro 2 parr.eter; to bru - the extrapolated ':TC /alue within the abcVe limito within 3 wura or be in at '. e a c t het chutdenen within o heura.
Amend ~.ent .o.I,3I,h3
/
2-50
ITEM 3
"'ha minim:n D:iBR value j'cr C l icic G ic given ao 1. 70. The FSAR givea a value cloce to 1.,3.
E= lain the difforcncco in analysis or accumption ::hich give thi:, di;fcrcnec. Provide the initia2 Dil?R (not thc miniracn) caiculated for Cycia 5. If it io significantly different from the Cycle C value of 1.87, cxr) lain the .: fference.
RESPONSE
The transient analysis (Xft-NF-79-79) has been supplemented to: (1) include protectior. against penetration of core saturation limits by TM/LP; and (2) allow implementation of the Cycle 6 TM/LP on existing Ft. Calhoun RPS hardware. The new minimum DNBR resulting from this analysis is calculated as 1.36 (see Attach-ment 1).
The initial DNBR for Cycle 5 was not calculated as part of the Cycle 6 analysis.
ITEM 4 In rcrfaming the D.7BR calculationa for non LOCA accident analyccc, hou ia the limitin] auluhannel determincd. Plcacc diacuca the ucc of tuo different fuel iccigna in providing your ancocr. (5cc Qucation 1G)
RESPONSE
The thermal-hydraulic models used to calculate the TM/LP and LC0 for DNB monitoring explicitly modeled the hydraulic performance of both fuel types in order to determine the appropriate limiting subchannel flow. The determination of the limiting assembly flow, and subsequent MDNBR, is accomplished in two steps: (1) Core flow distribution to determine the limiting assembly flow rate; and (2) Limiting assembly calculation for evaluation of the core thermal margin (MDNBR).
The core flow distribution calculation directly models the thermal and hydraulic performance of each fuel type as appropriate single hydraulic channels.
The thermal performance is evaluated using ENC neutronics methods to determine the core and assembly peaking distribution while the hydraulic performance is determined using the results of pre 3sure drop testing performed by ENC for both fuel types. The results of the calculations indicate that both limiting fuel assemblies will experience no less than 95" of core average assembly flow rates within the ENC assembly having less flow. Thus, the limiting ENC fuel assembly was selected for TM/LP and LC0 calculations.
The limiting assembly calculations model the limiting ENC fuel assembly into appropriate subchannels with the assembly flow rate as determined above.
The calculation is consistent with the methodology used for the core flow distribution calculations. This calculation results in determining the limiting subchannel flow rate and local fluid conditions used in the ensuing MDNBR calcu-lation for each fuel rod in each subchannel, The limiting subchannel as modeled in the non-LOCA analysis is initialized to be consistent with the results obtained in the analysis listed above.
ITEM 5 E=piain hcu lw:2 tcryerature (or gap conductance) is calculated for each cuent diaausaed in Rafarence 1. Are canaitivities done to shou chcther a high or ico calue of fuct tceparature io conservative? Hou ia fuel bur >wp concidared?
RESPONSE
The fuel temperatures calculated for the Cycle 6 analysis are calculated consistent with the methodology as given in XN-74-5, Rev. 1 For the Cycle 6 analysis, a value for the gap conductance of 500 Btu /hr-ft was used in the analysis of all plant transients.
The results of XN-NF-79-79 indicate that the limiting DNB trar ient is the rod drop transient and, as such, establishes the limiting conditions for steady-state operation. The MDNBR calculated for this transient occurs late in the transient at a time when the heat flux is essentially at steady state. Hence, the resulting MDNBR, including peaking augmentation, is established via a steady-state calculation which assumes a one-to-one correspondence between the power level and the heat flux at the time of MDNBP, and is virtually independent of the value of the gap conductance used in the analysis. Thus, the analysis performed to establish the plant operating thermal margins for the plant is judged to be adequate.
ITEM 6 It ic hypothetical:y poccible for a trancient to be iniciated f wm anycharc within the accepmbic opcmting range ac defined in the Tachnical Srceificationa.
Decer;be the mti:cd uced to accura the conditiona chocen for initiation of a tem tent ara the n:oct concervativa. Includa poccible pouar dictributionc chich 751l1 bd Cahadti bl} .T' anon tv: MciCnta. =
RESPONSE
The initial core conditions assumed in the analysis of anticipated operational occurrences assurae a simultaneous occurrence of variables to be at their most limiting values as defined by the proposed Cycle 6 Technical Specifications as indicated in Table 2.1 of XN-NF-79-79. Specifically, the initial reactor power is 1025 of rated power, system pressure is reduced by 47 psia to a value of 2053 psia, core inlet temperature is increased by 2 F to 547 F, and a minimum anticipated total core flow of 71.7 x 106 lb/hr is used. In addition, the limiting fuel rod power is also assumed to be consistent with the maximum value allowed by the Technical Specifications plus augmentation due to peaking uncertainties. The assumption of the above values is considered to be conservative to evaluate the operating thermal margins (MDNBR) in that the probability associated with this simultaneous occurrence is considered to be very small and the analysis protects even this low probability event.
The impact upon the initial power distribution during normal operation due to xenon transients is considered to be an axial effect and, as such, is considered in establishing the limiting conditions for operation (LCO's). The axial power profile used in the transient analysis exceeds that allowed by the limiting con-ditions for operation and hence results in an initial MDNBR lower than actually anticipated. Thus, the initial MDNBR is set at a conservatively low value providing additional conservatism in the initial plant conditions to evaluate thermal margins.
ITEM 7 (a) Fig. 3.13 of X:l-liF-70-79 chouc that the TlULP trip ic the first trip to or wata far the CEA Withdrawal at Full Po:xr over the entire range of ra.rtivity incertionc. Explain uhy thic ic co. Thace resulto do not agvce uith thoca of the FSAR analycec. Fig:a'ec 3. 7 to 3.10, uhich shou the ve;'ulic of the Fact CEA Withdrawal, chou only a clight inercace in coolant t cr~'cra t ur- or decreacc in praccurc until after the reactor trip.
(b) -C The Cacir 5 m:i=c reactivitu incertion for this trancient uac <310:10
~ .
ForChoicCthevalueis100:10~. E=viain the reduction in the rcactivity tncertion 2utc. !!cu vac cach of theca reactivity incarcion rates calculated?
(3) l,h t o ~ 4:~ 1 cuar cizapa and radial pouar dictributionc are used in the analycia of the Rod withdrao22 Event?
(1) In Figurc 3.13, c= plain uhy the rod uithdrawal cuant giucc the c=c D:l9R at ECC an,i ECC.
(e) Provile corn of PTS?WR and XCOBR-1 input data used to calculate the fact rod uithdraull event. This info 2=ation chould be providad ac coon ac roccibic to .t 2 th 3:aff thne to perfo2m indcpendent calculationc of thic cuent, if naduccar .
RESPONSE
(a) The TM/LP trip function applied in Xfi-f1F-79-79 at full power is P
var
=
31.61
- B + 20.77
- T.in - 12483.4 where P
ar
=
trip pressure (psia)
B =
high auctioneered thermal (AT) nuclear power (* of rated)
T.
in
=
core coolant inlet temperature ( F)
This TM/LP was conservatively generated through the use of a single limiting axial power profile for values of B > 100% which is calculated to be more limiting than required by APD. Hence, a higher than required sensitivity of P with respect to power (B) results. This increased sensitivity results in a large change in P with respect to a small change in power to the extent of initiating a reactor trip via Tli/LP before the overpower or high pressure trip set points are reached.
(b) The maximum reactivity insertion rate is determined by the maximum differential CEA group worth and the maximum speed of withdrawal (46 in/ min). The maximum reactivity insertion rate calculated for
-4 Cycle 6 is 1.725 x 10 Ap/sec. An additional CEA withdrawal analysis using the modified TM/LP was performed with a withdrawal rate of 1.725 x 10 an/sec. The results are not significantly different from
-4 the 1.0x 10 Ap/sec withdrawal rate. Thus, the impact of the CEA withdrawal upon plant thermal margins (MDf BR) are judged to be adequately represented by the range of withdrawal rates analyzed.
(c) The most limiting allowable axial power profile at full power is used in the CEA withdrawal transient. This axial power profile is selected in accordance with the axial power profiles allowed by APD at full power.
The radial peaking factor used in CEA withdrawal is 1.65, which is the nominal maximum peaking of 1.57 allowed by the proposed Cycle 6 Technical Specifications augmented by the radial distortion factor calculated at full power (see Response to Question 14).
(d) There are variations in transient MDf;BR between BOC and EOC conditions.
The report XN-NF-79-79 shows the lower fiDf;BR value of the two points in the Cycle. The results of the analysis of the CEA withdrawal with the revised TM/LP are provided fer both BOC and E0C in Attachment 1.
(e) The input data used in the analysis of the fast CEA withdrawal has been previously provided to the flRC Staf f.
ITEM 8 (Figurc 3.1) E.rplain uhy ti:a coolant flou incroacco during the CEA vithch'aval event.
RESPONSE
The increase of total coolant flow during the analysis of the CEA withdrawal event is a direct result of the calculated decrease in primary coolant temperature (coolant density increases).
ITEM 9 (Section 3.2) Explain the gradual decay of main oteam flou over a pcriod of foie ccconda for a Luca of External Land transient. If the timbine otop valoc h2e closed, and no credit is bcing taken for bypaca valvec or relief valves, hco can the alcam flou perciot for 4 caconda?
RESPONSE
The steam flow represented in the results of the analysis of the Loss of External Load transient represents the steam flow in a volume between the steam dome and the turbine stop valve. The analysis of this transient does assume the closure of the turbine stop valve and consequently the steam flow entering the turbine does go to zero. The steamline volume in question is up-stream of the pressure relief / safety valves and hence the steam flow in this volume is governed by the compression of the steam due to pressure increases during the first four seconds of the transient.
ITEM 10 (Section 3.7)
(a) For th ' lar; c at.mlina break, Figura ?. C3 choua a return tO pcccr reashir.]
a maxime: at ar;:rcrimatalil 12.5 ceconin. At this tir:c the praccura has dareanad 1:00 paia, rukir:g :::e RC:l preac:we appr::i":ately 300 raia. Danon-O tV::ta Ui Il G critical IlCat [iu: <*0rralatiCrt aproj'riatC to thia prGUCurC LI:ab thCr2 ia 1:0 dCYarture frCm nu2?CatC bCiliq. Or that ita C((Cata arc adCQ;c;OC[] C01:0idCrCd.
(b) i:: plain 1:w ncn-unifon in!ct tcepcraturc ellcata are inclwicd in the etcar:
1inc bv.:ak calcula lan.
RESPONSE
(a) A minimum critical heat flux ratio (q"DNBI9" ACTUAL) equal to 1.14 was calculated for the large steamline break at hot standby via a modified MacBeth critical heat flux correlation. Local coolant conditions for the limiting subchannel were obtained from XCOBRA-IIIC calculations performed at core conditions pro-vided by the PTSPWR2 transient simulation. The calculated minimum CHF ratio of 1.14 indicates that there is no departure from nucleate boiling in the limiting subchannel during this transient.
The modified MacBeth CHF correlation employed in this calculation is described in XN-74-21, Rev. 2, and is ar,~opriate to reactor conditions which encompass those in question:
- 1) Rod dianeter between 0 20 and 0.550 in.
- 2) Pressure between 600 and 1450 psia.
- 3) Coolant mass flux between 0.20 and 4.0 M lb/ hr-f t .
- 4) Inlet subcooling between 0 and 283 Btu /lb.
(b) The system model used in the analysis of plant t ansients for Cycle 6 is described in XN-NF-79-79 and separately models 'oth the " intact" loop as well as the loop modeled to have the steamline break and, as such, considers any difference in primary coolant from each loop during the transient. The fluid
in each loop is assumed to mix prior to determining the core inlet temperature during the transient and, as such, no inlet temperature maldistribution is assumed to occur.
ITEM 11 (Scotion .3. C. 2, pay? 74) For the roi drop it is otated th:t cycic' pracctwa ic ana:c ed conatant, but tha preca:wa actually decreaccc 105 paia. E.: plain uhy the decreace in ECS precoure vac not included in the DNBR analynca.
RESPONSE
The PTSPWR simulation of the Rod Drop event indicates that at 70 seconds:
(1) the reactor power tends to return to its initial value; (2) the core inlet temperature decreases approximately 7 F; (3) the system pressure decreases about 103 psia; and (4) the MDNBR value occurs at this point.
Sensitivity studies performed with XCOBRA-IIIC indicate that a conservative evaluation of the MDNBR for the rod droo event can be obtained by assuming initial core conditions plus the peaking augmentation anticipated for this event. In other words, no credit is taken fer the decrease in core inlet temperature and no penalty taken for the pressure decrease.
Since the credit associated with the decrease in inlet temperature exceeds the penalty associated with the pressure decrease, the analysis of the MDNBR for the rod drop event is found to be more conservatively calculated using the initial conditions (7 F higher core inlet temperature, and 108 psia high pressure) than the asymptolic conditions which exist in the transient.
ITEM 12 (Section 3.6.1)
(a) L'.:pl zin th.? una of a lov fico trip at K.E'; for tha tuo p:e p coactd:un an ! 07; for the four p:c"p coutdour2.
(b) 'iaa :: the act, flou va. tirnc cur:Ja uacd in the analyaic obtaincd?
(:-) :Sa iu a 0.3 inultiplinv uccd far Dorpler for ti:c Loca of Ccolant Flow thn a 0. 3 zu uced for other trv:aicnte
(d) H':u are unecriaintica in the lov flou trip cet point a ecuntad for both in the analyciu of the Lw of Flcu and Sci;wd Rotor cucnto and in the plant ana tr:encntation?
RESPONSE
(a) The FSAR analysis for Ft. Calhoun indicated that of the loss of flow transients the two-pump coastdown was the more limiting. In order to provide a conservative estimate of the MDNBR for this transient, the analysis was performed consistent with the FSAR analysis with a low flow trip at 92.5% of rated flow.
(b) The core flow versus time curves used in the analysis of the pump coastdown transient were consistent with the curves given in Section 14.6 of the Ft. Calhoun FSAR. Since the pump performance is independent of fuel type, these curves are judged to be adequate for the Cycle 6 analysis.
(c) A multiplier of 0.8 on the Doppler coefficient was used in the analysis of the Loss of Coolant Flow transients. The reported value of 0.9 is incorrect.
(d) The low flow trip set point is 95% of rated flow. To account for the 25 measurement uncertainty, the low flow trip point was set at 93% in the analysis.
ITEM 13 '
(5 cation 3. 7) Dica:cc accwytione mia about the location of the brcah for a cicam linc break. Alco, diccuca the aca w tionc chich acra made ac to the quality of the etca": and the dwcharja coaffte 'ent. ,
RESPONSE
The location of the steam line break was assumed to be at the outlet nozzle of the steam dome in the analysis. The fastest cooldown of the primary is thus achieved. The discharge coefficient was assumed to be one so that maximun possible discharge rate could be realized. Break flow was cal-culated each time step based on a choke flow model proportional to the steam generator pressure. The steam was assumed saturated; computations during the transient indicated the quality was essentially unity. Break flow was conputed based on an ideal gas flow model and results in a greater flow than calculated using Moody's results. Thus, the above model is judged to result in a more rapid cooldown and hence an increased likelihood of return to power.
ITEM 14 E.mlain hou chan3co in pouer a:atribution during an accident arc included in the acetdcnt analycia. Such cim :a ocuid be cauccd, for examic, by moucmant of contral roda or cooldoon or hcatup of the raactor coolant. (Sec Qucation 7c).
RESPONSE
The analysis of each transient is initiated from conditions which assume simultaneous occurrence of system conditions and nuclear peaking in a fashion so as to minimize the calculated MDNBR. The PTSPWR code calculates the system conditions (core flow, pressure, temperature, etc.), and these changes are utilized in deternining the MDNBR for each transient event.
The change in assembly radial peaking (assembly power) anticipated in those transients which result as a consequence of CEA movement, e.g. , CEA withdrawal,CEA drop is accounted for in the determination of the MDNBR. Steady-state, or equilibrium, neutronics calculations were performed covering the range of calculated coolant temperatures during the transient analysis. The results showed no significant increases in radial peaking and, in the case of reduced coolant temperatures, some decrease. No credit was taken for any decrease in radial peaking for those transients calculated to result in a cooldown of the primary coolant such as CEA drop, loss of feedwater heating, and CEA withdrawal .
The changes in axial power profile which may exist due to reactor scram and/or coolant temperature changes were conservatively neglected in the plant transient analysis. This a ,. roach is substantiated by the following arguments.
The critical parameter which most strongly affects the calculated MDNBR during the transient is the rod surface heat flux which is calculated to occur during the transient. The heat flux is predominately determined by the pellet stored energy distribution which is assumed to exist at the initiation of the transient. In principle, a fission product distribution would exist in the fuel
in proportion to the initial axial power distribution virtually independent of any transient occurrence. Since the half-life of those fission products which would continue the power generation process (and hence heat flux) .
is as long as s 50 sec., it is anticipated that the axial power distribution (axial fission density) would remain in proportion to the initial distribution for a time interval far in excess of that necessary to establish the MDNBR limits. For most of the DNB transients, e.g. , CEA withdrawal, loss of load, loss of feedwater flow, loss of coolant flow, the MDNBR occurs early in the transient and is well within the 50 second value listed above. Thus, based upon this phenomenon above, the use of a conservative initial axial power distribution is judged to be adequate in determining the MDNBP, during the plant transients.
In a fashion similar to that above, the energy created in the fuel pellets does not instantly appear as heat flux on the fuel rod surface. This is due to the time delay associated with the heat transfer of the fission energy through the fuel pellet and fuel cladding. Thus, the axial heat flux distribution actually depends upon the fuel pellet axial stored energy distribution which is proportionate to an axial power distribution which existed at some previous time in the transient and is anticipated to exist for several fuel time constants into the transient (on the order of several seconds). Thus, the initial assumption of a conservative axial power profile is adequate to determine the MDNBR during the plant transient analysis.
Any changes in axial power profile due to CEA movement are conservatively neglected in the plant transient analysis since any CEA movement would result in shif ting the axial power lower in the reactor. A CEA withdrawal would create not only an instantaneous but also a prolonged increase in fission density in the lower portion of the core due to the addition of positive reactivity to the
lower portion of the core prior to addition in the upper part of the core.
Due to the time dependent characteristics of the fission product inventory and heat transfer as discussed above, this shift in axial heat flux distribution would tend to exist for a period of time in excess of that required to establish the plant operating thermal margins.
The effect of reactor scram, which occurs in all transients with the exception of the rod drop transient, would have the same effect of redistributing the heat flux to the bottom of the core since negative reactivity is added to the top of the core, resulting in a similar increase in fission product density at the bottom Of the core.
Any reduction in coolant temperature such as exists in the more limiting transients, CEA witFdrawal, and rod drop has the effect of increasing the fission product density in the bottom portion of the core which in turn would result in an axial heat flux distribution tending to peak more in the bottom of the core than the axial power distribution used in the plant transient analysis. Thus, the use of a conservative axial power distribution in the plant transient analysis is adequate and envelopes any changes in the axial power distribution which may exist during the transient.
ITEM 15
& plain ai:3 tiinra ia no amlyaic of a dropped part inngti: control rod or lcca of AC pomr.
RESP 0lSE (a) The part length control rods are non-operable in the Ft. Calhoun reactor.
The worto associated with a part length CEA drop is anticipated to be less than the full length rod. Hence the consequences of a part length CEA drop is enveloped by the full length CEA drop analysis. The full length CEAs are allowed partial insertion in accordance with Cycle 6 PDIL require-ments. However, the peaking augmentation associated with the latter CEA drop is anticipated to be enveloped by the full length drop. Thus, the analysis presented in XV-fiF-79-79 adequately protects penetration of MDilBR limits for all anticipated rod drop events for Cycle 6.
(b) The most severe case of loss of AC power leads to the coastdown of the feedwater pumps and the core coolant pumps. The plant is protected by the low primary flow trip signal, the auxiliary feedwater system and the steam dump system.
This transient is not analyzed because it is judged to be bounded by the four-pump coastdown event, fhe reasoning is given as follows:
During the loss of AC power, the primary flow decreases at the same rate as in the four-pump coastdown event. The reactor will trip on low primary flow signal at approximately three seconds after the initiation of the transient. The heatup in the core would be the same provided the heat removal capability is the same. This is justified by the following rationale.
As the feedwater pumps trip, the steam generator level starts to decrease.
This may result in a decrease in secondary coolant flow from the downcomer to the active heat transfer region. However, the heat removal capability
will not be reduced, because the heat transfer action vill be enhanced by the increased boiling region on the secondary side, which is a con-sequence of the reduced flow. The secondary pressure is kept approximately the same as for the four-pump coastdown case. Hence, the secondary tempera-ture will be approximately the same. Therefore, the heatup in the core would be equal to or less than that of the four-pump coastdown event.
Af ter the reactor scrams, the reactor power decreases rapidly. It may take a longer time for this case to remove the residual heat, but since the MDNBR occurs immediately after scram (s?.0 sec for four-pump coastdown),
this has little effect on the thermal margin. Therefore, this transient is not analyzed.
ITEM 16 Dcc!rica Um nthod uccd to calculata the reculto of the CIM Dro; Event. :,'ha t coolant candition.' are acc:enad? What acc:cmtion: arc mada about the peak radial inar and th. lo:: tion of tha fact roi uith the peak rahal poucr before and after ine dro;>? l Rat vac the pacition of the CEAc ieat vac the valuc of th3 anyantation factor oithou t w:ccrtaintioc? Specify the unwrtainty factora and their valua,. Wint valua of buenup yielda the larycat aupentation factor?
Wh3 u the corth of the drof vcd rad that recultad in tha 15 au;~:antation?
Da.wil a the ma :al (includin.; road arial and ralial racer diatributicno) uced for the ai l> rop Event D:.2it calculaticnc. Idct any hot ch.v:nel petorc includcd in thaw rol pc~:cr dictributionc.
RESPONSE
The CEA Drop Event was simulated with PTSPWR to determine the core conditions as functions of time throughout the transient. The analysis simulates a drop of the most reactive CEA (3.4 x 10-3 @) which, taken with the neutronics parameters reported on Page 73 of XN-NF-79-79, represents the most limiting CEA drop event.
Initial core conditions are as described on Page 8 of XN-NF-79-79. To envelope the spectrum of possible CEA drop incidents, it is postulated that the limiting bundle undergoes the maximum calculated peaking increase which exists at any point in the core. The maximum augmentation factor of 1.21 (without uncer-tainties) was calculated for EOC conditinns. A factor of 1.16 was calculated at BOC but was not used in any ensuing calculations. In addition, all calcula-tions were performed with a 37 uncertainty factor. The pin power distribution within the hot assembly is assumed to remain unchanged through the transient.
The MDNBR for the CEA drop incident was obtained from steady-state XCOBRA-IIIC calculations. This approach was shown to be conservative in the Response to Question 11.
The XCOBRA-IIIC sukhannel model is illustrated in Figure 16.1 and assumes 1/8 assembly symmetry. The axial power distributions used to establish the MDNBR for the Rod Drop Event are consistent with the methodology used to establish the LCO for DNB monitoring. This methodology was described in detail in the response to NRC Question #23 on ENC Set Point Methodology (XN-NF-507).
1.034
<> <> <> <> <.958
- l. 985 8 7 68
- 1. ; O 0 1 86
<0J Rf 1.0 J < 001f Control Rod Guide Tube ( 1.092 1.00 _.
- 1. 0
. O FIGURE 16.1 FT. Call 10Vil SUSCliAfitlEL MODEL FOR MDiiBR EVALUATI0ft
ITEM 17 Decaribe method used to calculate the Lacc of Flav Event and the Sicccd Rotor Event. Decariba hou the recultc of PTSPWR are input into XCO3?A. Eco is the axial poucr dictribution detemined? Arc reactivity fcadback effecto on the axial pcuer dictribution concidered? Ecu are voido in the hot cinnnel considered for both 1: cat trancfor and reactivity effecto. Io collapcing of the voido due to tize incroacc in precoured modeled? The ancuero to qucationc 1, 4, 5, 12 and 14 arc relevant to thic qucation.
RESPONSE
Both the Loss of Flow and Siezed Rotor events are analyzed with the PTSPWR code and the reactor system model is the same as is used in the analysis of the other plant transients. The difference between the analysis of the above two events is essentially the transient initiation. In the analysis of the loss of Flow Events, the transient is initiated by assuming a loss of electrical power to the primary pumps resulting in coastdown of the primary pumps. In the Cycle 6 analysis, the use of the FSAR pump coastdown curves were judged to De adequate (see Response to Question 12). In the analysis of the Siezed Rotor Event, the transient is initiated by assuming loss of one primary pump leaving only three pumps operating.
The MDNBR calculated by PTSPWR at the initiation of the transient is verified by a detailed XCOBRA analysis which determines the limiting subchannel (see Response to Question 4).
The axial power profile used in the analysis of these events is identical to that used in analyzing the other plant transients. This profile envelopes those axial power profiles allowed by the proposed Cycle 6 LCOs and is judged to be conservative with respect to the allowed power profiles for Cycle 6 (see Response to Question 8).
The effects of reactivity feedback calculated to occur during the transient upon the axial power distribution are discussed in detail in the Response to Question 14.
The point reactor kinetics model used in the PTSPWR code represents six delayed neutron groups and includes reactivity feedback due to changes in core average fuel and coolant conditions which are calculated to occur during the transient. As the system transient response is governed by tha core average kinetics response, this modeling is considered adequate and any small contribu-tions due to reactivity feedback from the hot channel are judged not to affect the core average response.
The local coolant conditions which exist in the hot channel during the transient are based upon the assumed initial peaking, which has been shown to be conservative, and are determined with respect to the calculated core average transient responses. These coolant conditions are then used in determining the hot channel MDNBR during the transient.
Any change in system conditions during the transient, such as changes in pressure and temperature, is reflected in the thermohydraulic properties of the coolant in the system. Thus, a pressure increase would be reflected as a coolant density increase and an appropriate change in coolant quality (voids).
ITEM 18 Diccuca in detail hou E=on analycco include the Ccmbuction Engineering and E=on fuct roda. State ento are made in the topical rcporto cubmitted by CPPD uhich ctata that the analyces are done only for E=on fuct. For c ampic, Page 13 of Xi-liF-73-77 (Staa.~: Line Break), Page 13 of Xll-?;F-73-77 (LOCA biculoun analycic).
For LOCA hcatup analycic the core ic analyccd ac if it contained all CE and all E=on fuel. What poucr hictorica arc uccd for the c=cocura analycas? Do they conservatively bound fuel movementa in uhich the CE fuel u a in a lou pouer peri-phaval position and is nou poccibly in a higher pouer position chara it might pcccibly still be limiting cvan though it nau hac more burnup?
RESPONSE
Detailed core flow distribution calculations were performed prior to the analysis of the anticipated operational occurrences. The methodology is consistent with that used to establish the reactor set points and DNBR initialization in the transient. analyses. These calculations explicitly model the hydraulic character-istics of each fuel type in the core, the anticipated core loading, and anticipated nuclear peaking. In order to provide analysis results consistent with the proposed Technical Specifications, it was assumed that each fuel type was on peaking limits.
The results of these analyses indicated that the limiting ENC fuel assembly would rec;' approximately 35 less flow in comparison to the limiting CE fuel assembly.
Thus, the limiting ENC fuel assembly results in an MDNBR value lower than anticipated for a comparable CE fuel assembly and the ensuing thermal margins (MDNBR) analysis was performed for Cycle 6 using the ENC fuel as the limiting fuel assembly. This analysis approach is considered not only adequate for Cycle 6 when the peaking on the ENC fuel in anticipated to be well below limits, but also for future cycles as the peaking on the ENC fuel is anticipated to increase.
Exxon Nuclear Company has performed LOCA ECCS analyses for Fort Calhoun to establish ECCS allowable peaking limits, which assure compliance with 10 CFR 50.46 criteria for both ENC and CE fuel types. These analyses have included a determina-tion of the worst case LOCA (i.e. , limiting break).
These break spectrum calculations which determine the worst case are governed by the NSSS design and the assumed break size, location, and break con fi gu ration . The limiting break does not depend on fuel design or fuel exposure.
For Fort Calhoun, the limiting break was determined to be the largest double-ended guillotine break of a cold leg pipe. In the limiting break cal-culations, the fuel-related parameters correspond to ENC fuel at beginning-of-life.
A complete.LOCA ECCS calculation for the identified limiting break was also performed for CE fuel . Comparison of the limiting break LOCA response for the two fuel types shows that, other than in the fuel and the core itself, the , cal-culated results are essentially identical for the two fuel types. The limiting break primary system LOCA response results for ENC and CE fuels were used to provide core boundary conditions for the detailed core heatup analysis for the.
two respective fuel types.
The heatup analyses provide detailed calculations of core and fuel parameters during the limiting break LOCA transient. The calculations consider detailed fuel design, exposure conditions, and power profile conditions. The power history for exposure analysis assumes full power operation at the ECCS allowable peaking limit over the entirc hurnun period. This maximizes fission gas release and bounds all ECCS allowable power histories. Power peaking limits are computed as a function of burnup for each fuel type. Operation within these limits assures conformance to 10 CFR 50.46 criteria regardless of power history or fuel element position.
ITEM 19 E.rplain the basic and purpose of the fourth critaricn for thermal hydraulic perfor-n : in Mtion G. i doahng uith the claldtn3 temeraturca.
RESPONSE
The cladding temperature limits as defined in Section 6.1 of XN-NF-79-77 provide protection of cladding integrity during irradiation in the reactor environment. Specifically, the inside surface temperature limit of 850 F provides adequate margin to protect against fuel-cladding chemical interaction. The outside surface temperature limit of 675 F retains the expected hydrogen absorption within ENC limits and reduces the potential of corrosion-induced hydrogen absorption of the fuel rods. The volumetric average temperature limit of 750 F represents a reasonable limit between the time-temperature annealing effects and irradiation hardening of the cladding material and provides adequate material strength for in-reactor service.
ITEM 20 (Table 7.1) ihplain the factora (1.07) and (1.08) uced ao multiplicro on the final crorgy deposition at !!FP and the (1. 33) and (1.10) used at li:P.
RESP 0flSE The above question refers to the rod ejection accident in the Fort Calhoun Cycle 6 core. The specific factors in question were taken from the Generic Analysis of the Control Rod Ejection Iransient for Pressurized Water Reactors (I) .
The Generic Analysis discusses these factors in detail .
The total enthaley in a rod ejection accident depends on the initial enthalpy, the control red worth, the power peaking factors, the doppler coefficient and the delayed neutron fraction. In Figures 4.3 (HFP) and 4.4 (HZP) of Reference 1 the deposited enthalpy, corrected for differences in initial enthalpy between the generic case and Fort Calhoun, must be augmented to account for the worst anticipated Doppler coefficient and delayed neutron fraction of the Fort Calhoun Cycle 6 core.
The augmentation factors 1.07 (Figure 4.i of Reference 1) and 1.08 (Figure 4.5 of Reference 1) are used to augment the total enthalpy at HFP in a pellet to account for the worst anticipated doppler coefficient of 0.95 x 10- Ap/ F and the delayed neutron fraction of 0.0045. Correspondingly, the augmentation factors 1.10 (Figure 4.2 of Reference 1) and 1.33 (Figure 4.6 of Reference 1) are used to augment the total enthalpy at HZP in the pellet. The utilization of the augmentation factors is described in more detail in Reference 1. Reference 1 has been submitted to the flRC as the ENC generic analysis of the control rod ejection transient for pressurized water reactors.
- 1) RJ Burnside, TL Krysinski, DW Pruitt, "A Generic Analysis of the Control Rod Ejection Transient for Pressurized Water Reactors," Xft-flF-78-44
ITEM 21 (Scotion A. 2. 3) For the limiting large break, the F value nuat bc reduccd above a car . height of 707a. This is ahoun in Fig:we 9A. 2. Explain uhy th's curva doca not hava to appcar in the Ft. Calhoun Technical Specifications. Is the concluaion likely to change from cycle to cycle?
RESPONSE
The curve of Figure A.2 shows the ECCS allowable peaking as a function at axial position (which conforms to 10 CFR 50.46 and Appendix K to 10 CFR 50 based on ENC analysis). The reduction in allowable peaking at the top of the core shown by Figure A.2 is less than the reduction in peaking which occurs at the top of the core due to the inherent neutronics of the reactor. Therefore, the Technical Specification limits for the allowed peaking, or LHGR, at the 701 elevation or below, is sufficient to also protect against exceeding the limits of Figure A.2. This conclusion is unlikely to change from cycle to cycle unless a design change is implemented which alters the inherent neutronics of the reactor.
_uestion Q 22 It is our understanding that the TM/LP equation to be used is not the equation given on Page 30 (Xii-flF-79-77). Justify the validity of the analysis presented in References 1 and 2 in view of the changed equation.
Response
The impact of the final TM/LP on the analyses reported in Xft-f!F-79-79 has been carefully considered. The transients listed in Category A of Table 1.1 (Xfi-t1F-79-79) which were reported to have tripped on TM/LP have been reanalyzed using the revised TM/LP.
Results of the reanalysis are reported in the discussion of the Application for Amendment of Facility Operating License forwarded herewith. Those results indicate that modification of the RPS by inclusion of the revised TM/LP does not alter the protection afforded by the RPS.
Question 23 Describe how delays in the TM/LP trip circuitry (including the initial delays in power and inlet temperature) are modeled in PTSPWR for the Fort Calhoun safety analyses.
Response
The overall scram delay time reported in Xfi-flF-79-79 for the low pressurizer pressure subsumes' signal acquisition and processing delays inherent in the TM/LP trip circuitry. The overall scran time for the low pressurizer pressure used in Fort Calhoun safety analysis PTSPWR calculations is consistent with the Fort Calhoun FSAR.
ATTACib!ENT 1 Sll.'e!ARY The'following letter reports the final TM/LP trip equation for Ft. Calhoun Cycle 6 stretch power operation.
The transient events which were reported in Reference I to have tripped on TM/LP have been reanalyzed to verify the adequacy of the final T11/LP equation.
Results of the reanalysis are summarized.
o No event considered causes a TM/LP trip.
o Adequate margin to DNB is maintained for the duration of the CEA withdrawal events by the high power and high pressure set points.
o The most limiting transient considered, the fast CEA withdrawal event, trips on high power with MDNBR equal to 1.36.
o Excessive load increase incidents trip on high power or establish a new steady-state with substantial margin to DNB.
o 'Ita ef fectiveness of the RPS is not impaired by inclusion of the final
'I M/ LP .
Reference:
XN-NF-79-79, " Fort Calhoun Cycle 6 Reload Plant Transient Analysis Report," October 1979 s
es .
THE 'Df/LP TRIP FUNCTIi The following function is the final version of the Ft. Calhoun TM/LP LSSS:
P var
= 5.54 PF(BjB + 22.48 T.
in
- 10S01. (1) 1.0 B > 100 PF(B) = -0.01B+2.0 50 < B < 100 1.5 B < 50
/
This function is based on final Cycle 6 neutronics calculations. It protects core saturation limits and the SAFDL on DNB during those anticipated operational occurrences listed in category A of Table 1.1, Reference (1). In addition, it shculd prove amenable to existing Ft. Calhoun analog control systems. Isobars from equation (1) are plotted in Figure 1.
Those category A tr: 'ients (see above) for which TM/LP trips are reported have been reanalyzed with the new TM/LP function to assure adherence to the SAFDL on DNB. Analytical methodology for the reanalysis is as described and referenced in (1). The MDNBRs and core conditions calculated in the reanalysis are summari:ed in the attached Table 1, and indicate that the RPS as modified by inclusion of the new TM/LP trip equation adequately protects SAFDLs. The individual analyses are summarized below.
CEA WITHDRAWAL The CEA withdrawal event is described in Section 3.1 of (1) . The analysis was performed for a wide range of reactivity addition rates at both BOC and EOC conditions. All cases resulted in reactor trips on variabic high power except the three BOC cases at the lower end of the reactivity insertion spectrum, which resulted in reactor trips on high pressure. Calculated MDNBRs appear in Figure 2 as a function of reactivity insertion rate, and indicate that adequate thermal margin is maintained for all the CEA withdrawal incidents considered.
EXCESSIVE LOAD INCREASE INCIDENT Excessive load increase incidents and analyses are as described in Section 3.4 of reference (1) . The cases considered are: 1) rapid opening of the turbine control valves, and 2) sudden opening of steam dump and steant bypass valves.
A new steady-state is attained in Case 1) without initiating a reactor trip. Marginal cooldown and pressure decrease occur, accompanied by a small increase in power. The MDNBR is 1.53.
For Case 2), the reactor scrams on high power at about 9 seconds with a MDNDR of 1.47. Peak power level is 1702 M'i. Cooldown and depressurization are more rapid and extensive than for Case 1).
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TABLE 1 SU5BIARY OF RESULTS 9
bl\XIMU51 MAXISRJM MAXIMUM CORE AVERAGE PRESSURIZER POWER LEVEL llEAT FLUX., PRESSURE MDNBR TRANSIENT (PEltCE.NT) (BTU /IIR-FT') (PSIA) (W-3)
Initial Steady-State 102 176,213. 2053. 1.69 Fast CEA Nithdrawal 112.4 188,721. 2270. 1.36 Slow CEA Withdrawal 120.1 183,401 210S. 1.37 Loss of Feedwater lleating 112.S 187,612. 2066. 1.43 Excessive Load Case 1 104.5 180,428. 2056. 1.53 Excessive Load Case 2 113.5 184,256. 2057. 1.47 e
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LOSS OF FEEINATER HEATING The loss of feedwater heating event and analysis are described in Section 3.5 of Reference (1). Reanalysis results in a variable high power trip at about 25 seconds. Calculated MDNBR is 1.43 at about 27 seconds.
Adequate margin to DNB is maintained for the duration of the transient.
g CONCLUSION The results of the transient analyses reported above indicate that the fast CEA withdrawal transient is the most limiting of those previously reported to have tripped a TM/LP. More than adequate margin to DNB is main-tained for the duration of that limiting transient. It may be concluded that modification of the RPS by substitution of the final TM/LP equation does not impinge on the integrity of the RPS.