ML20099K268

Proposed Tech Specs Re Boron Concentration in Moderator Water for Standby Liquid Control Sys

ML20099K268
Person / Time
Site:	Pilgrim, 05000000
Issue date:	03/19/1976
From:	BOSTON EDISON CO.
To:
Shared Package
ML20099K024	List: ML20099K053 ML20099K111 ML20099K116 ML20099K155 ML20099K172 ML20099K232 ML20099K268 ML20099K308 ML20099K324 ML20099K349 ML20099K371 ML20099K494 ML20099K508 ML20099K512 ML20099K542 ML20099K556 ML20099K559
References
FOIA-84-105 NUDOCS 8411290366
Download: ML20099K268 (26)

Text

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. BASES:

3.4 STAtmny Lmino cornmot Sumt A. The condi,fAons under which the Standby Liquid I Control' System must provide shutdown capability are identified via the Station Nuclear Garcty .

Operational Analysis (Appendix G). The require-ments of this specification are taken from the Operational Nuclear Safety Sequirements of subsec-tien 3.8.6 of the Final Safety Analysin Report. If no acre than one operable centrol red is withtirawn, the basic shutdown reactivity requirement for the core is satisfied and the Standby Liquid Control

, system is not required. Thus, the basic reactivity requirement for the core is the urinary determinant j of when the liquid centrol system is required. -

The purpose of the liquid control system is to pro-vide the capability of bringing the reacter frca full pcuer to a cold, xenon-free shutdown conditica assumin6 th:.t none of the withdrawn control' rods can be inserted. To =eet this objective, the liquid centrol system is designed to inject a quantity of boren that produces a cencentration of 7CO ppn of -

beren in the reactor core in less than 125 =inutes. 5' e Be 700 ppm concentratien in the reacter core is requirci to bring the reactor frcs full pcuer to a three percent Ak suberitical condition, censidering

] the hot to celd reactivity difference, xenon poisen-

"s ing, etc. "he time requirement for inserting the -

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boren Oclutien was selected to override the rate of reactivity insertien caused by cooldown of the re-f acter follouing the xenen poison peak.

2e ninimum '4 *tatien en the relief valve setting

is intended to prevent the less of liquid centrol

! solutien via the lifting of a relief valve at tco lov a pressure. 2 e upper 11:10 en the relief valve settings provides syste: protection from overpressure, i

3. Only one of the two standby liquid control pu= ping ,

i loops is needed for cparating the system.

Cne inoperable pumping. circuit does not i==ede

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intcly threaten shutdoun capability, and re-actor operation can continue while the circuft

'is being repaired. Assurance that the 4

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. A. h.. i ATTACHMENT A Responses to NRC Questions on Pilgrim I Reload 2 Licensing Submittal (NEDO-20855-01)

Question A1: Does BECo. plan to use an operating limit MCPR of 1.31 for 7 x 7 asse=blies and 1.39 for S x 8 assemblies for the entire cycle?

Response: Yes Question A2: Provide the cross sectional areas assumed for indiv'idual com-ponents in the DBA - ie, jet pumps, cleanup loop, etc.

Response: The component flow areas assumed are:

Suction line vessel nozzle area: 3.56 ft2 Cleanup line area .08 Jet pump (10) nozzle area .71 TOTAL 4.35 ft-Question A3: Does a reactor scram occur as the result of a feedwater transient?

Response: In Paragraph 7. 3. 3.1. 6.1 (Page 7-13) of NED0-20S55-01, the peak neutron flux was incorrectly stated as 120% of initial. The corrected statement should be: " Neutron flux increases to a value of 119.8% of initial at 109 seconds." Since the indicated neutron flu: does not ex-ceed the scram setpoint of 120%, scram does not occur.

A time plot of the transient is shewn in Figure 3-1 (attached).

Question A4: Provide a power-flow map for Reload 2 showing a 92%

pcwer limit.

Response: The Pilgrim Cycle 3 power / flow map with the nominal 92%

flow control line is shewn in Figure 5-1, attached.

Refer to response to Question A10 for a description of how Pilgrim 1 will operate on the power /ficw map to achieve the 92% power, 100% flow endpoint at EOC-3.

Question AS: For the over pressure transient with one relief valve out of service, what steam flow and power level were assumed? .

Are all other assumptions the same as those assumed in the previous analysis with all valves operable?

Response: The one relief valve out of service analysis was performed at 92% power and steam flow and 100% core flow and all other assumptions are the same as for the analysis with all valves operable.

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t Question'A6: The Rod Withdrawal error transient doesn't appear to follow the .

core loading maps provided in-NED020855-01. Clarify the incon-sistencies that relate to bundle identification and location on j the various core maps and figures provided in support of the RWE-transient.

f, Response: The coordinates of Figure 2-1 of NEDO-20855-01 identify the locations of the fuel bundles and those of1 Figure 7-12 identify 1 the location of control rods. Location of the control. rods in Figure 2-1 may be determined by. interpolation.

Erroneous bundle identifications are used in Figures 7-13

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through 7-16. Corrections on these figures as well as for

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Figure 7-12 are given below.

L-Figure No. As Shown Should Be 7-12 "end rod is 35-18" " Error rod is 18-35" 7-13 11, 11 21, 32 11, 10 21, 34 09, 09 17, 36 7-14 12,.12 23, 30 -

10, 10 19, 34 7-15 11, 11 21, 32 9, 9 17, 36 7-16 12, 8 23, 33 8, 10 15, 34 10, 10 . 19, 34 Question A7: Provide an analysis that establishes that a loading error ac-cident does not significantly affect adjacent fuel assemblies.

Response: The fuel loading error analysis has been performed for Pilgrim Reload 2 with bypass flow holes plugged. Subsection 7.3.2.5.2 of NEDO-20S55-01 is hereby revised as shown below.

"7.3.2.5.2 Results and Consequences The analysis of the loading error accident is based on operating MPCR's at the limiting point in the cycle where the "B" scram re-

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activity curve is still applicable. This results in a peak linear-heat generation rate (LHCR) of 16.6 KW/ft and a minimum critical

. power ratio-(MCPR) of 0.96 in.the misplaced'b'undle. This linear--

heat generation rate is below the value at which 1% plastic ~ strain of the cladding occurs; Fuel damage is not expected to occur with a LEGR lower than that needed to cause a 1% plastic strain in the cladding (see Section 3.2.1 of Reference 1). Therefore, fuel fail-ure is not expected for this event.

Fuel bundles adjacent to the misplaced bundle are insignificantly.

affected by the presence of the misplaced bundle."

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In the above analysis, the Reload 2 80 262 fuel is the limiting '

bundle. Seven pins are expected to experience boiling transi-tion. Bundles adjacent to the misloaded. bundle are isolated by the water gap so that the thermal neutron flux, hence power, is-not significantly increased; therefore, the effect on these bundles is considerably less than that for-the. misplaced bundle.

The operating MCPR over 'the range of "B" scram curve applicability (1.26 for 7 x 7 fuel and 1.33 for:8 x 8) on which this analysis is' based are significantly lower than the ones to be administered throughout Cycle 3 (1.31 for 7 x 7 fuel and 1.39 for 8 x 8; see response to Question-A1). The difference in CPR between the oper- .

ating MCPR and the MPCR of the misplaced bundle is relatively in--

sensitive to the initial value of operating MCPR. Therefore, the MCPR for the loading error identified above is conservative by '

about 5%; i.e., a MCPR of about 1.01 for the misplaced bundle would be experienced, based on the actual MCPR limits to be admin-istered.

Question AS: Describe th.e extent, if any, of shuffling of the fuel from the initial core loading and Reload i locations. If fuel shuffles are to be made, discuss the applicability of the transient and accident analyses presented for Reload No. 2.

Response: Fuel shuffling assumptions used to design the reference loading pattern for the Pilgrim Reload 2 licensing analysis were as follows:

1. The icwest reactivity bundles were assumed to be discharged.

If in actuality others were discharged,-the final core con-figuration will be lower in reactivity than' the design ref--

erence loading pattern since higher reactivity bundles were discharged.

2. The design of Reload 2 reference loading was based on the. .

lowest projected Cycle 2 shutdown exposure. Hence, extension-of Cycle 2 operation would produce a core with a lower reac-tivity than that presented in the license-submittal.

3. It was assumed that any bundle could Ina shuffled provided the resulting pattern met licensing criteria (e.g., MCPRs. shut-down margin) and fuel cycle criteria (e.g., energy: requirement).

4 Quarter core mirror symmetry was maintained as in-Cycle 2.

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The d'esigd' reference' loading was' based on the maximum number.

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of Reload 2 bundles which could be loaded.at the assumed EOC-2 exposure. Thus, any reduction in the number in 'the Reload - 2 bundles loaded will result in a lower reactivity core than.

oresented in' the license submittal.

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n- r3.y The criteria to be used in establishing the final loading pattern for Cycle 3 are given as follows:

1. The design reference locations of all Reload 1 and Reload 2 bundles remain unchanged.

2. If the leaker bundle is not one of those bundles originally scheduled for discharge per design ~ reference loading pattern, it shall be replaced by a sound bundle with higher exposure; i.e., lower in reactivity.

3. Maintain greater than 1% shutdown margin (desig'). n ,

4 Minimum cycle energy requifement shall be satisfied.

Application of the above criteria for establishing the final loading pattern assures that both the reactivity of the four-bundle cell which contains the replaced bundle and the worth of the control rod in that cell will be less than those of the

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design reference loading pattern. In addition, the extended Cycle 2 operation has made the core average exposure of_ the final loading pattern higher than that of the design ref erence loading pattern. The conbination of these two factors will make the final laeding pattern more conservative than the design reference loading pattern as far as shutdown margin, rod with-drawal error, and other safety relatad analyses are concerned.

Thus, the transient and accident analyses presented for Reload 2 are applicable.

Question A9: What power level is assumed in Table 3 '.? Provide a discussion that this assumption results in'the limiting case.

Response: Tacle 5.4 in NED0-20333-01 was derived frca ?ilgrim's rated conditions. The parameters listed tnerein were used as initial conditions on both the 100% and 92% power level transient. anal-ysis listed in Table 5.3. The CETA3 evaluations for these tran-sients are not significantly affected by total core pcwer level.

The sane initial MCPR for different total core power levels can be obtained by adjusting the radial power peaking factor. The transient aCPR is evaluated using the relative ' change in core conditiens and is not very sensitive to initial factors such as

, the value of the radial power peaking factor, Thus a 10% in-t crease in core power. for example, can e the same change in i

CPR in a bondie at a part icular :!C1'R whet her the core power i .<

l 90; or 1002 of rateel. Therefore, the initial comiit ions in Table 5-4 are applicable for CEI AB transient analyses at both 100% and.92% power levels. '

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F Question A10: Are the analyses at the endpoints of 100% power with "B" scram curve and 92% power with "EOC-3" scram curve bounding for all MCPR and pressure transients for all power and burnup combina-tions shown on Figure 7-11. Provide a description of how Pilgrim I will operate on the power / flow cap to achieve the

[ 92% power, 100% flow endpoint of ECC-3.

Response: Figure 7-11 of NEDO-20S55-01 shows the resultant maximum power level profile as a function of cycle exposure. The derate schedule is shown as a linear function of fuel ex-posure connecting the specified calculational operating power limit points. The use of the linear relationship to connect the two calculational points in Figure 7-11 is .

conservative because the scram reactivity degrades gradually and would thus be a smooth function of core exposure. Thus, for the pressure transients connecting the actual calcula-tional points with a straight line will conservatively main-tain a minimun pressure margin of 25 psi since the actual allcwable power level would be expected to lie somewhere above this operating limit line. Conservatism is incorporated into the operating MCFR's by imposing the limiting operating

.dCPRs of 1.31 for 7 x 7 fuel and 1.39 for the S x 8 fuel, .

calculated for the worst degraded condition (and-of-cycle),

over the entire cycle. Thus, the analyses at the endpoints of 100% power with "B" scran curve, and 92% power with the ECC-3 scran curve, are bounding for all MPCR and pressure transients for all power and burnup combinations shown on Figure 7-11.

Operation at 1000 pcuer level is permissable over the range of "3" scran curve applicability. Seyond that point, 2600

>SD/T into the cycle, the power will be reduced f rom 100%

power to 92% power at ECC-3, limited by the maximua pcwer profile shewn on Figure 7-11. This limit will be administered by imposing small step darates, each of which is valid for sc e incremental exposure. For each derate there will be ad-ministered a corresponding nominal power-ficw line, interpolated between the nominal 100% flow control line and the nominal 92%

flow control line shown on Figure 5-1.

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ATTACHMENT S Responses to NRC Questions on Pilgrim Single Loco coeration (NEDO-20999) m Question 31: Provide the details of the calculations which are used in the evaluation and comparisons of one and two pump operation for 37R/3 and SWR /4 type plants. The information required includes the input MAPLHGRS for one and two pump operation and the ECCS type information such as transient core flow, core pressure, lower plenum enthalpy, MCPR, convective heat transfer coefficient, water level, vessel pressure, peak cladding temperature, and break .

spectrum curve. -

Response: The requested details of the calculations for one and two pump operation are compared in Table.1-1. It is noted that the cal-culation of the two pump MAPLEGR and peak cladding temperature (PCT) values presented in NEDO-20999 are entirely consistent with the procedures specified in NEDE-20566. The evaluation of MAPLHCR and PCT values for one-pump operation are also consistent with NEDE-20566 with :he exception that boiling transition is assumed at 0.1 seconds after the LCCA. The time to boiling transition is provided as input to the one-pump analysis and is not calculated as part of the analysis as is the case in the two-pump analysis.

For cwo-pump operation, the transient core flow, core pressure, lower plenum enthalpy, and MCPR are calculated, as always, using the LAM 3 and SCAT computer codes. From these evaluations the duration of' nucleate boiling after the LOCA is calculated. For aingle-loop operation, the transition from nucleate boiling is conservatively assumed at 0.1 seconds. Therefore, the LAM 3 and SCAT conputer codes are net required and are not used in the eval-uations for one pump MAPLHGR's and PCT's. With this exception, the heat transfer coefficients used for the evaluation of one-pump and two-pump MAPLHGR's and PCT's are identical. Water level and vessel pressure are calculated by the SAFE /REFLOOD Code for both one-pump and two-pump operation. PCT and break spectrum curves are calculated by the CHASTE (with SAFE /REFLCOD) computer code for both one-pump and two-pump operation.

Question BlA: In particular how are differences in reflood time, spray time and uncovery time used to predict a change in MAPLHCR. For example,  ;

a one second reflood delay is equivalent to a one-half percent reduction in MAPLHCR for late reflooders; while changes in the above parameters is equivalent tc a one percent reduction in MAPLGHR for early reflooders.

Response: Sensitivity studies with Appendix K ECCS evaluation models provide the basis for estimating the effect of small changes in reflooding time (TFLOCD), uncovery time (TUNC), and core spray cooling initia- i tion time on the MAPLHGR and PCT for early and late reflooders. For example, an increase of I second in reflooding time increases the PCT by about 3of for a late reflooder and by about 30F for an early reflooder. The one second increase in reflooding time decreases the MAPLHGR by about 0.10% for a late reflooder and by about 0.25% for an early reflooder.

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TABLE l-1 i

COMPARISON OF FAPLHGR AND TCT CALCULATION CETAILS .e0R ONE-PUFP VERLUS TWO-PUMP OPERATICN r

TWO-PUMP ONE-PUMP OPERATION OPERATIO" MAPLNGR For Table 2-3 of NECO- 15.1 kw/ft 12.Z kw/f t .

20999 Core Pressure Calculation LAMB Not Applicabie*

Transient Core Flow Calculation LAMS Not Applicable-Lower Plenum Enthaley LAM 8 Not Applicable *

' Calculation .

MCPR SCAT Not Applicable' Convective Feat Transfer NEDE-20565 NECE-20S66**

C: efficient Water Level Calculation SAFE /REFLOOD SAFE /REFLCCO Vessel Pressure Calcula!.icn SAFE /PEFLOOD . SAFE /REFLC00 Peak Cladding Tec:erature CHASTE CHASTE Calculation Break Sp'ctrum Calcula:icns CHASTE plus CHASTE plus SAFE /REFLC00 SAFE /REFLC00 .

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• Soiling transition for LOCA from one-loca oceration is assumed at 0.1 seconds; therefore, LAMS and SCAT calculations not recuired.

• • For one-purp 0;eration, loss of nucleate boiling assumed at 0.1 seconds af ter the LOCA.

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Question 313: For the two-pump MAPLHGR is a constant heat transfer coefficient (HTC = 30) used until lower plenum flashing as in previous sub-

< mittals or was the calculation done as described in NEDE-20566, (Dougall-Rohsenow with Ellion used to calculate steam generation)?

Response: The calculation was done as described in NEDE-20566 (Dougall-Rohsenow with Ellion).

Question B1C: Are hot node uncovery and core uncovery as used on Page 2-5 and 1-7 synomonous?

Response: Yes.

Question B2: Provide the data that shows that boiling transition occurs earlier .

for discharge than for suction breaks of the recirculation line, as discussed in Section 2.2.4 on Page 2-7 (NED0-20999).

Response: This is a mistake in the text. The seccad sentence in the first paragraph of Section 2.2.4 should read: " Curves for both suction and discharge breaks are presented because the onset of boiling transition occurs significantly later for discharge breaks" (cor-tected in Supplement 1, attached). Examples of boiling transition times for suction break versus discharge break for several plants are listed below. Note that the enset of boiling transition occurs much later for discharge breaks.

ONSET OF BOILING TRANSITION (sec)

Suction Break Discharge Break PLANT A S.3 15.9 PLANT 3 6.0 12.4 PLANT C 1.4 3.S Question 33: Pr: ride a discussion en the effect of core plugging on MAPLHCR reductien~for one pump operation.

. Response: Plugging the core plate holes, ,chich substantially reduces re-flooding capability after a LCCA, results in longer calculated reflooding times. The FMPLHGR reduction factor (a multiplication factor on the two pump MAPLHCR) which is a function of reflooding time, increases as the reflooding time increases (i.e., for plugged cores) as shown in Figure 1 of NEDC-20999. Therefore, the effect of core plugging is to increase the reflooding time and thereby increase the MAPLHGR reduction factor. Although the SMPLHGR Re- +

duction Factor increases for pluggeu core, the resultant MAPLHCR for single loop operation is less fer a plugged core than for the unplugged core. The assumed time to boiling transition is inde-pendent of whether or not the core is plugged.

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Question B4: Provide a curve of transient PCT of the lowest axial plain to experience CPR = 1.0 prior to jet pump uncovery versus time with one pump operation. .

Response: The cladding heatup analysis of the LOCA from one pump operation assumes that boiling transition occurs over the entire length of the fuel bundle at 0.1 seconds after the LOCA. In other words, all axial planes are assumed to experience CPR = 1.0 prior to jet pump uncovery. Therefore, the high-power axial plane experiences the most severe cladding heatup. Plots of calculated peak clad-ding temperature versus time for the high-power axial plane are provided in Figures 4-1 and 4-2 for early and late reflooding 3'a'R/4 plants.

Question 35: Provide a discussion on the one pump vessel blowdown and reflooding calculations relating to the assumption of 102" of rated power and flow as being conservative when compared to operating at a reduced pcwer level.

Response: The Appendix K ECCS blowdown-reflooding calculations (SAFE /REFLOOD computer codes) are performed assuming that the reactor is operating at 102" rated power with corresponding core flow, steam flow, etc.

at the time of the postulated LOCA. For single loop operation, the reactor will be operating at considerably less than 100% rated pcwer.

The parameters input to the cladding heatup analysis are somewhat sensitive to the reactor operating conditions (core power, core ficw, etc.) assumed in the blowdown-refloeding calculations.

Table 5-1 (attached) presents a comparison of calculations for the high-pcwer axial plane uncovery time (TU::C) , the core spray cooling initiation time (TSPRAY), and the high-power axial plane reflooding time (TFLOOD) for the following two cases:

CASE 1: LOCA from two-loop operation assuming the reactor is operating at 102% rated power with corresponding core flow, steam flow, etc.

CASE 2: LOCA assuming the reactor it operating at the ap-proximate reduced power level for single loop operation (approximately 75% rated power) with corresponding core flow, steam flow, etc.

The comparisons are made for the DBA, 80% DBA, and 60% DBA break sizes for the example plant. Table 5-1 indicates longer uncovery time, longer time for the single-loop LOCA (reduced power and flow).

The longer uncovery time and shorter reflooding time decrease the calculated peak cladding temperature (PCI) for the full power case, while the longer time for the initiation of spray cooling tends to increase the PCT by a small amount. The net effect of using the parameter values for the full power case rather than for the reduced power case is to increase calculated PCT for the DBA by approximately 1GoF. Thus, since the PCT is slightly increased by this assumption, it is convenient and conservative to use the parameter values cor-responding to the two-pump LOCA from full power for the cladding heatup calculations for the single-loop LOCA.

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, , m, TABLE i-1 COMPARISON OF BLO*.iD0iP4-REFLOODING PArWETERS FOR FULL PC'.!ER (TWO-LOOP OPERATION) CASE VERSUS RECUCEO POWER CASE (SINGLE-LOOP OPERATION)

BREAK SIZE CASE 1* CASE 2**

d TUNC 25.4 25.1 DBA TSPRAYD 33.5 32.5 - '

TFLOODc 106.3 104.7 TUNC 28.1 29.2 80% DBA TSPRAY 39.6 40.9 TFLOOD 103.6 99.4 TUNC 33.0 37.6

-.,,n...

o u, n a.n, -

. an r t ,i . 7 - , -

aa.o TFL000 105.4 10a.3 "TUSC =

H:: nede unccvery time (sec) b TSPRAY = time ac v.hicn credit is assumed for core spray heat transfer (sec) c TFLCOD = Reficcding tire for hot ncce (sec)

" CASE 1: LOCA frca cao-icco c:eration as:umina the reacter is cceratino at 102; rated ;:wer witn correspccc'5g core fic'.1, steam fica,'e:c.

• • CASE 2: LOCA frcm single-lcoc oceration assuming the reactor is operating at reduced pcxer wicn corresponding ccre flow, steam ficw, etc.

l

J Question 36: Discuss the significance of the PCT term in the MAPLHGR Reduction Factor (F) equation and its tendency to lower the "F" f actor when the 2 pumps FCT is below 2200 F. Clarify the statement in Para-graph 3 of Page 2-9 (NEDC-20999) that no credit for PCT margin is taken to calculate the one pump MPALHGR.

Response: This question is not relevant to Pilgrim operation in Cycle 3 because MAPLHCR limits are derived from the 22000F Appendix K limits throughout the cycle. Ecwever, in the interest of for-warding information to the NRC, the following information is forwarded from G.E. Since this informatioa is generic in nature and is not applicable to Pilgrim, it is expected that it will be used for information purposes and will not effect the NRC review ..

~

of the Pilgrim single loop submittal.

The equation:

F=1- )

MAPLHGR (2 Punn)(2-Pump)

MAPLHGR MAPLHCR (1 Pump)J

~

~2200 F - PCT (2 Pump)

,0o.y x 0.01 -

is used to calculate tha generic MA?LHOR Reduction Factor (F) curve in Figure 1 of NE00-20999. The standard procedure followed in calculating the data points in Figure 1 is as follows:

(1) One-pump P_;FLHGR and PCT values for selected SWR /3-BWR/4 plants are conservatively calculated acccrding to the as-sumptions of Section 2.2.3 (two-pump MAPLHGR and PCT values are already available for previous Appendix ~ ECCS analysis submittals);

(2) These MAPLHGR and ?CT values for ene and two pump operation are then used in the above equation to calculate F. If the two-pump PC~ is equal to 22000F, then the PCT term in the abcVe equation is zero and the equation to calculate the MAPLEGR Reduction Factor is simply.

_ p=1 _ 'MAPLHGR (2 ourn) - MAPLGHR (1 Puno)

MAPLHGR (2 Pump)

If the two pump PCT is less than 22000F then the ?CT term is employed to ensure that a conservatively low MAPLEGR Reduction Factor for the generic curve in Figure 1 is calculated. For example, suppose that results of specific heatup calculations for a particular 3ZR/3-BWR/4 are as follows:

MAPLHGR (2 Pump) = 15.00 kw/ft; PCT (2 Pump) = 2000 F.

! MAPLHGR (1 Pump) = 14.25 kw/ft; PCT (1 Pump) = 22000F.

i These results show that the one-pump MAPLHGR is reduced to 14.25/

15.00 = .95 of the two-pump MAPLHGR. It is noted for this example that the MAPLHGR for single-loop operation is reduced by a compar-stively s=all amount (5*) because there is 2000F (22000F - 20000F) of margin in the two-pump PCT before the MAPLHCR is limited by the

' W e A:.p,n.u.. e 11M r.

m; Q ne The results cited in this example are now used to calculate a generic data point for Fi:;ure 1. Using the above equation, the calculated MAPLHCR Reduction Factor for Figure 1 is F=1- Il5.00 - 14.25' -

~2200 - 2000 x .01 =

L 15.00 . . 20 1.0 - 0.05 - 0.10 = .85-Therefore, for this example plant, the effect of the PCT term in the equation for F is to reduce the calculated generic MAPLHCR Reduction Factor for Figure 1 by 0.10 (f rom 0.95 to 0.35) . There-fore, 0.85 is a conservatively low MAPLHCR Reduction Factor that may be applied to a different plant with an equal reflooding time  :

and with a two-pump PCT equal to 22000F.

^

To answer the second part of this question, consider the following example. A single-loop operation MAPLHCR is required for a' plant with the same reflooding time as the plant in the above example, but the two pump PCT is 2100 F. The generic MAPLHGR Reduction Factor (F) from Figure 1 is 0.55, and therefore the single-loop MA?LEGR is: .

MAPLHCR (1 Pump) =  ? x MAPLHGR (2 Pump)

= 0.35 x MAPLHCR (2 Pump) 4 It is noted in this example that credit for the 100 0F margin (2200U F - 210G U F) in the two pump PCT is conservatively i;nored in the calculation of the single-loop MAPLEGR frcm Figure 1.

Taking credit for this margin w0uld increase the MAPLHOR reduc-tion factor to 0.?0 (instead of 0.55) and thereby increase the one-pump operation MAPLEGR by approximately 31.

In summary, it is conservative to account for the two puup PCT margin in the calculation of the generic MAPLHGR Reductica Factor Curve (Figure 1) because this results in conservatively low values of F to be applied generically to all SWR /3-BWR/4 plants with com-parable reflooding times. On the other hand, it is conservativa to ignore the two pump PCT nargin when using the generic MAPLCHR Rcduction Factor (Figure 1) curve to calculate the one pu=p opera-tion MAPLGHR because, by so doing, conservatively low MAPLHGP.'s are obtained.

Question B7: Provide assurance that all BUR /3 and SWR /4 plants vithout the LPCI modification will be limited by the suction line break. For example, BSEP :!2 with a LPCI modification is limited by the discharge line break, it is not obvious that without the LPCI modification that the limiting break will rivert to the suction line. Plugging of.

the bypass flow holes in the core support plate may also have an effect on the limiting break location.

y - - - - - - v- -.

,. m '

l l

Respoase: In the heatup analysis for sin).le loop operation, boiling transi- '

tion is assumed at 0.1 seconds for both suction and discharge line breaks. Therefore, the reflooc.ing time is the primary parameter determining the peak cladding temperature. For BWR/3-BUR /4 plants without the LPCI modification, the reflooding time for the suction line break is always longer than for the discharge break. This results because inventory losses during vessel blowdown are always less for the discharge line break due to smaller break area (re-stricted by the limiting flow area through the recirculation pump).

Since the suction break has a longer reflooding time, it will always be more limiting on MAPLHGR than the discharge line break for BWR/3-BWR/4 plants without the LPCI modification.

Plugging the bypass leakage holes significantly retards core re-flooding for either suction break or discharge break. 'The suction break reflooding time is longer for plugged bypass holeg (for the .

same reasons cited above) and therefore limits the >MPLHCR in this case also.

Question 58: Section 2.2.5 needs clarification on the following items.

33A: The last paragraph on Page 2-10 needs clarification with regard to boiling transitica for large breaks being maximum relative to the ,

DBA. Will a smaller break relative to the DBA have a longer time -

to 3.T. and therefore a greater MAPLHCR reduction for one pump operati:n.

Response: In Section 2.2.5 it is stated that for the plant selected for tne calculations in Table 2-3, "the tire to boilina transition is maximum relative to the u c, a ,, .

.:o unders tand tne meaning or tnis, considerrac. le S-l . ane ,iast column sgcus tne boiling transition times for large breaks (8G3 CBA, 601 ;BA, and I ft-) minus -he :SA toiling transitico tire for two plants. Rela:ive to Plant B, i- is seen far Plant A tnat the dura:icn of nucleate boiling fcr tne large brea<s is longer relative to tne CEA boiling transition time. Therefore, the assuraticn cf early boiling transition for one-pur; operatico increases the large creak PCT's relative to tne OEA PCT to a larger extent f;r P1 ant A than for Plant 3. Inus, Plant A was selected for tne calculations sccwn in Table 2-3 t0 illustrate tnat the PCT for the large break portion of the creak spactrum decreases witn decreasing break area.

Wita regard to the seccnd part cf questicn #2A, it is true that the ceiling transiticn times increase witn decreasing break areas. I t i s r.a t true ,

however, that a break smaller tnan the DBA will limit tne one-puna lGPLEGR for a BWR/3-C'..'R/4 plant. In Section 2.2.5, the evaluations presented in Table 2-3 are applied generically to all CUR /3-CWR/4 " lead plant" analyses.

This conclusively demonstrates tnat the large break PCT's for all " lead clants" (and thus for all CWR/3-BWR/4 plants) are always less than the CBA PCT for one-pump operation.

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. l

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}

For smaller breaks (i.e., break neea <l.0 ft 2) Section 2.2.5 presents a comparison (for break area = 0.0/ f2t ) of the effect of single-loop operation on the PCT. For this one case, the PCT increased from 1725'F to 1760*F in going from two-pump to one-pump operation.* The slight increase in PCT for this example is explained in the second paragraph of page 2-12. The last paragraph on this page summarizes the arguments that the small break PCTs remain well below DBA PCT.

There is a tendency to draw an analogy between single-loop operation in jet pump BWR plants and non-jet pump BWR plants because the duration of nucleate boiling for both cases is quite short (less than 1-2 seconds). Since the MAPLHGR's for non-jet pump SWR plants are currently limited by breaks smaller than the DBA, it has been suggested by the f'RC staff that this may be the case for single loop operation in jet pump BWR's. However, such is not the case because for jet pump BXR's the PCT transient is terminated; by reflooding, .

whereas the non-jet pumo BWR's rely on PCT turnover by core spray cooling to terminate tne heatup transient. The reflooding phenomencn occurs much earlier, particularly for smaller breaks, than does PCT turnover as illustrated in 1 Table 8-2, which compares the reflooding time for a late reflooding SWR versus PCT turnover time for a non-jet pump SWR.

For single loop operation, immediate (0.1 sec.) loss of nucleate boiling is assumed independent of break si:e. T'a us , the initial temperature re-sponse is identical for breaks of different sizes. The larger break un-covers earlier and therefore it has a higher temperature after the tine of uncovery for the large break. ery late in the transient, the later spray initiation for the case of the smaller break causes the temperature dif ference between the large and small to be reduced. Ecwever, reflooding cccurs at early enough times such that the larger break has the higher terperature. Specific detailed calculations have sacwn this to be the case (see SEDo-22999, Section 2.2.5) l 253: Are the two pump PCT's in Table 2-3 calculated using the original MA?LEGR

and the one purp ?CT's calculated with a two pump MAPLHGR recuced by the reduction factor with no credit for "22000F - PCT" nargin?

"It is noted that tnese small break PCT's were calculated with the CHASTE computer code. For two-pump operation (PCT : 1725'F) toe small break assumptions are emoloyed (i.e., nucleate boiling until uncovery).. For one-pump operation (PCT = 1760'F) Ellion cool toiling (htc = 20 Btu /hr-f 2t *F) is assuned until uncovery. The PAPLHGR used in the one-pump PCT calculation is reduced by 151 from the two-pump MAPLHGR.

. . . ~.

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P L' fit A "* ESA 8.3 --

e- re.n eu. .

10.0 1. 7, .

60% CBA 12.6 4.3 o

1.0 ft' 27.0 18.7 l

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, TABLE ?-2 COMPARIS0t 0F REFLC00I7:3 TD'E FOR A LATE REFLOODIt;G JET PUr7 B':a '/EF. sus PCT .

TURt:0'.'ER TD'E FOR A ?:0N-JET PtFP 8'JR i F

LATE REFLOODIt!G tlCil-JET PUMP JET PLW 80 BWR BREAK SIZE i REFL000!tlG TIME PCT TURI;0VER Tit'E (sec) (sec) t 08A 280 '300 1.0 ft 2 220 340 2 '

0.1 ft 360 550 O.07 ft 2 480 650 6

4 WP a

G v

n l

• l

. fra

~: e'3 i 3: l P

i Response: The. calculations presented in Table 2-3 were performed for the BWR/3-BWR/4 with the earliest reflooding time (106 seconds) as explained in Section 2.2.5.

The MAPLHGR's used for the PCT calculations are 16.1 kw/ft (the original MAPLHGR) and 12.7 kw/ft for two-pung and one-pump operation, respectively.

From Figure 1 of NE00-20999 the IMPLHGR Reduction Factor is 0.74 for this plant. Therefore, the one-pump MAPLHGR for this plant would be

.0.74 x 16.1 = 11.9 kw/ft if no credit is assumed for the "2200*F - PCT" margin. However, taking credit for this margin increases the single loop MAPLHGR to 12.'/ kw/f t. Although the value of 11.9 kw/ft is the recommended MAPLHGR for single loop operation, the 12.7 kw/ft value is used for the one-pump PCT break spectrum calculations in Table 2-3 because this results in higher calculated PCT's for the breaks smaller than the DBA, and therefore is conservative.

It is furtner n for the C.07 ftgted for clarification break was calculatedthat for theaPCT comparison different plant forinwhich S ection the2.2.5 MAPLHGR Reduction Factor for single-loop operation is 0.85 (i.e., a 15%

reduction from the two pump MAPLHGR).

BSC: . Describe how the MAPLHCR's for one and eaa pump operation are derive.d in Table 2-3. -If Figure 1 is used for the shortest reflood time plant,' the reduction factor is 0.74, and the one pump MAPLHCR would be 16.1 x .74 or 11.9. The MAPLHCR reduction factor derived from the MAPLHGR's shown in Table 2-3'are F = 1 F16.1 - 12.7 7 = .79.

A 16.1 J Response: The answer to this part of Question 48 is included in Part B immediately above.

Question 39: For the discussion of small break PCT's in S2ction 2.2.5 the following needs to be clarified.

39A: Are the two pump MAPLEGR's calculated using the SAFE code with the small break assumptions (i.e., nucleate boiling till core uncovery zero heat transfer till core spray, and core spray heat transfer until reflood)?

Response: This question refers to the two pump PCT's, rather than MAPLHCR's, for-small breaks. For single-loop operation, the small break PCT's are cal-culated with the CHASTE Code assuming loss of nucleate boiling at 0.1 C

seconds, folicwed by Ellion pool boiling until core uncovery, zero heat transfer until core spray, and core spray heat transfer until reflooding.

This method is identical to that for the large break PCT calculations for single loop operation. For two-loop operation, the small break PCT is

. calculated with the CHASTE code with the small break assumptions, i.e. ,

nucleate boiling until core uncovery, zero heat. transfer until core spray,

, and core spray heat transfer until reflooding.

B9B: It appears that the small break model (SBM) should yield a greater reduc-tion in MAPLHGR because the core uncovers slowly (no B.T. occurs) and MAPLHGR reduction is greater for one loop operation at longer times to boiling transition (B.T.). 2 Compare the 1.0 ft using the LBM and SBM.

Provide the MAPLHGR's PCT's for two pump and one pump operation.

- . , .- ,s p )

u -

Response: The first part of this question is answered in the response to question 8A.

Table 9-1 presents the 1.0 ft2 PCT for the Large Break Model (LBM) and Small Break Model (SBM) for both two-loop and single-loop operation. The MAPLHGR's used in the calculations are also specified in Table 9-l. It is noted that both the LBM and SBM PCT calculations for single loop operation were performed assuming the Ellion correlation until uncovery. Furthermore, if nucleate boiling (rather than Ellion) until uncovery is assumed for the single-loop operatien SBM, the PCT decreases from 1900*F to 1480*F as ,

shown in Table 9-1.

~

Qusacion 310: Provide the delay to boiling transition based on the GE correlation if calculated. If the delayed boiling transition was not calculated give justification for not considering in one pump operation.  ;

)

Response: i For single-loop operation the delay to boiling transition for all BWR/3-BWR/4 plants is not calculated but rather conservatively assumed to be 0.1 )

seconds after the LOCA. For non-jet pump BWR/2 plants (not included in the scope of this report) the calculated delay to boiling transition is typically 1~.0 - 1.5, seconds. The calculated delay time to boiling transition-(using the non-jet the same for BWR/2, 3, pump plant boiling transition correlation) is essentially and 4 plants because the major determining parameter in t he calculation is bunale power, which is approximately equal for these plants.*

Therefore, the justification for using the assumed value (0.1 seconds) rather than a calet.latec value (approximately 1.0-1.5 seconds) is that this assumption results in conservatively high calculated ceak cladding tecceratures (because the longer duration of nucleate boiling is more effective in removing the stored energy from the fuel before the transition to film boiling, and therefore, if assumed, reduces the cal-culated PCT).

• Actually the time to boiling transition would be longer for SWR /3 and 4's

- because, even for a LOCA from single loop operation, there would be signi -

ficant core coastdown flow induced by natural circulation through the jet pumps.

This induced core flow would result in lator boiling transition than that predicted by the correlation used for BWR/2 plants which assumes zero core coastd'o wn flow.

O .

. 1

_ :. .u.. . . -- . . . . . . . '

.s.

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.)

.m .

TABLE 9-1

' COMPARISON OF PCT FOR 1.0 FT2BREAK USING LARGE BREAK MODEL AND SMALL BREAX lCDEL FOR TKO-LOOP f.50 SINGLE-LOOP OPERATION PEAK CLADOING TEMPERATURE (*F) 2 FOR 1.0 FT BREAK TWO-LOOP OPERATION

• ONE-LCOP OPERATION **

Small Break Mcdel" 1715 1900***

Large Break Mcdel b 1730 1925

1. Small Break Model: SAFE /REFLC00 b

large Break Model: CHASTE

• MAPLHGR = 16.1 kw/ft for two-loop operation.

• • MAPLHGR = 12.7 kw/ft for single-loco operation. For single-loca operation, the Ellion correlation is assumec until uncovery for both LBM and SEM.

• • • If nucleate boiling until uncovery is assumed for the SEM single-loop PCT, the result is PCT : 1480*F.

i 1

-w

- m3. .. ,

i

. . 7 O. s

Qusscion Bil: P'rovide the MAPLHCR curves for one loop operation for the plugged.

reload core.

Response: NEDO-20999 is applicable for the bypass flow holes plugged case only.

This clarification is made in Supplement 1 (attached)..

' Quastion 312: Provide assurance that the Kg factors that are derived from the cold water increase transient (recirculation pump speed up, both loops operating) will be bounding for one loop operation.

- Response: The Kr factors are derived assuming that both recirculation loops increase

- speed to the maximum permitted by the M-G Set scoop tube position set screws.

i. This condition produces the maximum possible power increase and hence .

maximum aMCPR for transients initiated from less than rated power and flow.

When operating with only one recirculation loop the flow and power increase

, associated with the increased speed on only one M-G Set will be less than that associated with both pumps increasing speed, and, therefore, the Kf factors derived with the two pump assumption are conservative -for single loop operation.

! B12A: Also provide a discussion of the cold water increase (positive reactivity insertion) transients and how they are bounded by the two-loop full power analysis, s

1 Response: The loss of feedwater heater event is generally the most severe cold water increase event with respect to increase in core power. This event is caused by positi e reactivity insertion from core flow inlet subcooling (see i

Reference 5 of llECO-20999); therefore, the event is independent of two-pump or one-pumo operation. The severity of the event is primarily dependent on the initial pcwer level. The higher tne initial power level, the greater the CPR change during the transient. Since the initial power level during one- ,

q pumo operaticn will be significantly lower, the one-pump cold water increase '

1 case is conservatively bounded by the full power (two-pump) analysis.

i Question B13: P.rovide details on how the curves relating core flow to drive flow, as described in Section 4.2, are obtained.

{ ,

Response: See revised pages to NEDO-20999 (attached).

3' l Question 314: The derivation of the rod block equation appears to have an inconsistency.

Should the first " equation on Page 4-4 read, "RB100 = m(100 + AW) + RB'", f

, rather than, RB100 = m(100 + AW) + RB"?

i- Response: - See revised pages to NEDO-20999 (attached),

I i

1 I

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• p

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l'

• 4 o O q- .

tQuestion B15: Provide the 100% point on the drive flow axis on Figure 4. There i: appears to be an inconsistency on the location of the 100% point.

The point Fc = 100%, W = 100% should be che ,two pump curve, while

[. the abscissa label indicates the one pump curve.

t Response: See revised pages to NEDO-20999 (attached) .

Question B16: Justify using the two-loop uncertainty factors for calculating the Safety Limit MCPR for single loop operation. The reverse flow through the idle jet pumps may result in a higher flow uncertainty factor.

Response: Most of the uncertainties used in statistical analysis (Table 5-llof NEDO-20855-1) are independent of whether flow is provided by two loops or a single loop. The only exception is the-total coreflow which, for. ~

two pump operation, has a standard deviation (% of point) of 2.5. For single loop operation, this value would increase to about 6% of rated core flow. 12ue 3.5% increase'in core flow uncertainty corresponds to an increase in the safety limit of approximately 0.004 which can be

, neglected.

' It should be noted that the steady state operating MCPR with single loop operation will be conservatively established by multiplying the rated flow MCPR limit.by the Kf factore This assures that the 99.9%

statistical limit requirement 1 is always satisfied.

Quastion B17: Provide a technical basis for the proposed changes in the intercept and slope for both the flow biased APRM flux scram and the rod block setting. . Provide a quantitative assessment of margins to the MCPR safety limit at the lower permissible valuesrof core . flows for 100%

control . rod pattern. Include in the assessment such local power in-crease transients where thermal / hydraulic and 2udden effects are in

~

phase. Examples are rod withdrawal errors and xenon redistribution caused by normal operations.

Response: A more detailed justification for the' proposed changes la the intercept and ' slope for the flow bia; sad APRM scram and the rod block setting than has already been provided is r.ot available at this time. General Electric ha's advised us that it will be several months before this information be-comes available. We will'cantinue to review the benefits of-this proposed Technical'Specificaticu revision and will forward the information as ap-

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propriate as it becomes available and following completion of our review of this information.

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