ML20129H588
| ML20129H588 | |
| Person / Time | |
|---|---|
| Site: | Maine Yankee |
| Issue date: | 10/28/1996 |
| From: | Dan Dorman NRC (Affiliation Not Assigned) |
| To: | Frizzle C Maine Yankee |
| References | |
| TAC-M94834, NUDOCS 9611010238 | |
| Download: ML20129H588 (5) | |
Text
_
October 28, 1996 i
Mr. Charles D. Frizzle, President i
Maine Yankee Atomic Power Company 329 Bath Road
-Brunswick, ME 04011
SUBJECT:
REQUEST FOR ADDITIONAL INFORMATION ON SBLOCA ANALYSES RELATED TO AN 4
ITEM 0F LOOP-SEAL-INDUCED-CORE-UNC0VERY - MAINE YANKEE ATOMIC POWER STATION (TAC NO. M94834)
Dear Mr. Frizzle:
Enclosed is a request for additional information (RAI) on the small-break loss-of-coolant accident (SBLOCA) analyses performed by Siemens for Maine Yankee full power operation. The request identifies an item of concern regarding the loop-seal-induced-core-uncovery phenomena. This item of concern must be resolved for the NRC staff to complete its review of the SBLOCA analyses for full power operation of Maine Yankee Atomic Power Station.
If you have any questions regarding this matter, please call me at (301) i 415-1429.
Sincerely, (Original Signed By)
Daniel H. Dorman, Project Manager Division of Reactor Projects - I/II Office of Nuclear Reactor Regulation Docket No. 50-309 l
Enclosure:
Request for Additional Information cc w/ encl: See next page Distribution Docket File PUBLIC
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RConte, RI 010033 DOCUMENT NAME: G:\\D0RMAN\\M94834.RAI Ti esce've a copy of thee document, andcate in the box:
"C" = Copy without attachment / enclosure
'E" = Copy with attachmentlenclosure "N" = No copy l0FFICE PM:DRPE
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l lNAME DDorman"h W EPeyton
- T JZwolinskT lDATE 10/25/96 10/MV96 10/#/96 0FFICIAL RECORD COPY l
9611010238 961028 PDR ADOCK 05000309 P
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UNITED STATES g
j NUCLEAR REGULATORY COMMISSION t
WASHINGTON, D.C. 205NM001 October 28, 1996 l
Mr. Charles D. Frizzle, President Maine Yankee Atomic Power Company i
329 Bath Road Brunswick, ME 04011
SUBJECT:
REQUEST FOR ADDITIONAL INFORMATION ON SBLOCA ANALYSES RELATED TO AN ITEM 0F LOOP-SEAL-INDUCED-CORE-UNC0VERY - MAINE YANKEE ATOMIC POWER STATION (TAC NO. M94834)
Dear Mr. Frizzle:
Enclosed is a request for additional information (RAI) on the small-break loss-of-coolant accident (SBLOCA) analyses performed by Siemens for Maine Yankee full power operation. The request identifies an item of concern regarding the loop-seal-induced-core-uncovery phenomena. This item of concern must be resolved for the NRC staff to complete its review of the SBLOCA analyses for full power operation of Maine Yankee Atomic Power Station.
If you have any questions regarding this matter, please call ine at (301) l 415-1429.
)
Sincerely, C1.1-
' ~ ' -
ctu Daniel H. Dorman, Project Manager Division of Reactor Projects - I/II Office of Nuclear Reactor Regulation Docket No. 50-309
Enclosure:
Request for Additional Information cc w/ encl: See next page
i j
Maine Yankee Atomic Power Station Maine Yankee Atomic Power Company l
cc:
Friends of the Coast i
Mr. Charles B. Brinkman P.O. Box 98 Manager - Washington Nuclear Edgecomb, ME 04556 Operations ABB Combustion Engineering Mr. Christopher R. Shaw 12300 Twinbrook Parkway, Suite 330 Plant Manager 4
Rockville, MD 20852 Maine Yankee Atomic Power Station P.O. Box 408 Thomas G. Dignan, Jr., Esquire Wiscasset, ME 04578 Ropes & Gray 4
One International Place Mr. G. D. Whittier, Vice President Boston, MA 02110-2624 Licensing and Engineering 2
Maine Yanke Atomic Power Company Mr. Uldis Vanags 329 Bath Road State Nuclear Safety Advisor Brunswick, ME 04011 State Planning Office i
State House Station #38 Mr. Patrick J. Dostie Augusta, ME 04333 State of Maine Nuclear Safety Inspector Mr. P. L. Anderson, Project Manager Maine Yankee Atomic Power Station Yankee Atomic Electric Company P.O. Box 408 580 Main Street Wiscasset, ME 04578 Bolton, MA 01740-1398 Mr. Graham M. Leitch Regional Administrator, Region I Vice President, Operations U.S. Nuclear Regulatory Commission Maine Yankee Atomic Power Station 475 Allendale Road P.O. Box 408 King of Prussia, PA 19406 Wiscasset, ME 04578 First Selectman of Wiscasset Mary Ann Lynch, Esquire Municipal Building Maine Yankee Atomic Power Company i
U.S. Route 1 329 Bath Road Wiscasset, ME 04578 Brunswick, ME 04578 Mr. J. T. Yerokun Mr. Jonathan M. Block Senior Resident Inspector Attorney at Law Maine Yankee Atomic Power Station P.O. Box 566 U.S. Nuclear Regulatory Commission Putney, VT 05346-0566 P.O. Box E Wiscasset, ME 04578 Mr. James R. Hebert, Manager Nuclear Engineering and Licensing Maine Yankee Atomic Power Company 329 Bath Road Brunswick, ME 04011
i i
i, l
MAINE YANKEE ATOMIC POWER COMPANY i
MAINE YANKEE ATOMIC Pp 9 STATION i
DOCKET NO. 50-s09 REQUEST FO't ADDITIONAL INFORMATION l
}
ON THE SBLOCA ANALYSES TO SUPPORT FULL POWER OPERATION 1
l l
Loon-Seal-Induced-Core-Uncoverv l
The results of small-break loss-of-coolant accident (SBLOCA) analyses provided I
by the industry for pressurized-water reactors (PWRs) show that a loop-seal-3 l
induced-core-uncovery may occur for breaks located at or near the top of the j
reactor coolant pump (RCP) discharge leg. The resulting peak cladding temperature of 1000,F to 1500 *F may last for extended periods of time.
Under these conditions, the metal-water reaction may result in significant l
core wide and peak local oxidations, which must meet the acceptance criteria j
of 10 CFR 50.46(b) for oxidation limits and long term cooling.
l For SBLOCAs with the break location at the top of the RCP discharge leg and a
break cross-sectional areas of about 0.005 to 0.02 ft, reactor coolant system (RCS) pressure holds constant at a relatively high value because the energy l
added to the RCS from decay heat power is matched by the heat removal through the steam generators (SGs) and the energy loss through the break.
For these-j break sizes, which are large enough to prevent the refilling of the RCS with emergency core cooling system (ECCS) injection, but small enough to avoid the l
RCS depressurization following establishment of such a pressure plateau, the 4
j RCS two-phase level eventually decreases to the elevation of the break at the top of the discharge leg and stabilizes at this level. During this stage of a SBLOCA, the steam generated in the core passes through the SGs to the break.
Calculations show that about 90 to 95 percent of the steam generated in the 4
core is condensed in the SGs, and a low steam flow rate exits the SGs into the loop seal region. Since the low steaming rate is insufficient to entrain and remove water from the loop seal region, the hydrostatic head in the loop seal i
remains at high levels as steam bubbles through the loop seal region to the break.
Since the SGs will not condense all of the steam in the SGs for this i
range of small breaks, the low rate of steaming from SG to the loop seal j
causes a pressure buildup in the reactor vessel upper plenum to balance the t
hydrostatic head of water in the loop seal.
If the elevation of the lower i
portion of the loop seal is below the top elevation of the core (as in Maine Yankee), the pressure buildup in the reactor vessel upper plenum may become large enough to cause the core to uncover. Depending upon the break size, break orientation, decay heat level, vertical elevation of the loop seal and ECCS pump head characteristics, the loop-seal-induced core-uncovery may develop and exist for extended periods of time. The peak cladding temperature may be in the range of 1000 to 1500 *F for an extended period of time and raises concerns regarding the resulting metal-water reaction exceeding the limits of 10 CFR 50.46(b).
2 The attached non-proprietary submittal by Framatome Technologies, Incorporated provides additional information to describe the staff concerns and the associated phenomena.
You are requested to provide results cf relevant calculations and to address concerns associated with RCP loop seal clearing and break orientation, and compliance with the requirements of 10 CFR 50.46(b), including concerns regarding metal-water reaction and long term cooling.
Attachment:
Framatone Report l
l
M FRAMATOME j
TECHNOLOGIES Integrated Nuclear Services JHT/96-46 July 15,1996 1
U. S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington, D.C. 20555 l
Subject:
Supplementary Information to FTl's Response to NRC's Request for i
Additional Information on BAW-10168, Volume 11, Revision 2, October 1992; RSG LOCA - BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants.
Reference:
J. H. Taylor to Document Control Desk, " Response to NRC's Request for Additional Information on BAW-10168, Volume ll, Revision 2. October 1992; RSG LOCA - BWNT Loss of-Coolant' Accident Evaluation Model for Recirculating Steam Generator Plants," JHT/94-171, October 28, 1994.
Gentleman:
The reference transmitted FTI's response to an NRC request for additionalinformation en topical report BAW-10168, Revision 2. The attachment provides supplemental information to the referenced response. The material enclosed herein is considered non-proprietary to Framatome Technologies.
Very truly yours, t
. H. T or, M ag r Licensing Services cc:
Frank R. Orr, NRC R. B. Borsum L. W. Ward, INEL - DC C. P. Fineman, INEL - ID 9D U o ogg'- 3315 Old Forest Road, P.O. Box 10935, Lynchburg, VA 24506-0935 e.
Telephone: 804 832-3000 Fax: 804-832 3663
i
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Break Discharge Coefficients:
For SBLOCA, the leak flow requirements of 10CFR50.46 Appendix K have generally been interpreted as use of the Moody discharge of 1.0 for the entire two-phase flow regime. However, In BAW-10168 Revisio Section 4.3.2.4, FTI proposed the use of realistic break discharge coefficients for I
calculations. Comparisons between the Moody discharge correlation and experiment
}
that Moody overpredicts the leak flow rate for void fractions of 70 percent (corre i
quality of 10 percent at a pressure of 1000 psi) or greater. To account for this i
better predict system depressurization, FTI's method used a C, of 0.7 for void frac i
i percent or greater. For subcooled, superheated, and saturated discharges up to void fracti 70 percent, a C, of 1.0 was still used. FTI's break discharge methodology w i
based on our qualitative evaluation of the approach and with a request for a qua
}
before or with its first application. FTI provided the NRC-requested evaluations w l
l response to requests for additional information on Revision 2 of BAW-10168, Volume H.
j After consultation with NRC personnel, it became clear that FTI's discharg
}
a sound technical basis, would be considered as non-standard, requiring a substant licensing effort. We have concluded that the expenditure of such an effort i
be productive. In point of fact, for most SBLOCAs the use of either method
}
comparable trends and results, since little time is spent at leak void fractions whe i
differences are noted between Moody and test data. Therefore, FTI is modifying
}
break flow model to reflect the common interpretation of Appendix K. A dischar
]
of 1.0 will be used regardless of leak flow quality--subcooled, saturated, or sup Discharge correlations-Extended Henry-Fauske (subcooled), Moody (saturated), a l
Baurnan (superheated)--will remain unchanged. This, coupled with a break with the intent and requirements of Appendix K for SBLOCA.
i This switch in methodology will not invalidate the studies and benchmarks pe
}
of Revision 2 nor will Fr! totally abandon the use of its more accurate modeling
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will reanalyze SBLOCA cases having clad ten.perdures in excess of 1800 F model. Reductions in the rate of system depressurization occurritig during the " core b
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(or high void phase of the transient), resulting from the use of the variable C, metho j
adversely impact ECC injection, core inventory, and possible lead to clad t r.geisure above those predicted using the normal App =k K techrique. Analyzing high t--r SBLOCA transients using both Appendix K and our variable C, methods will assure th ereture is not underpredicted. SBIhCA transients below 1800 F are not highly ni prible to large clad i
j temperature changes resulting from items such as the incidence of rupture and its a inskle/outside metal-water energy addition; the reverse bewiring true as i.edures cl i
l 1800 F. At and above 1800 F, the energy contribution from the metal-water rea increasingly significant. For those cases just below 1800 F, a reasonable safety j
400 F to the PCT criterion is provided. Hence,1800 F is a logical transition point b i
analyzing a SBLOCA transient using only the Appendix K method and analyzing
]
both methods.
4 i
I i
/
i
-., ~. - - - -
i 5
In summary, FTI will use a discharge coefficient of 1.0 for the entire two-phase lea All other aspects of our break modeling wiB remain unchanged. This methodology com j
the intent and recuirements of Appendix K for SBLOCA. For SBLOCA transients pred i
temperatures above 1800 F using the Appendix K technique, FTI will also analyze such cases using its variable C, method. 1800 F will be the established transition point. Analyzing i
l cases with both techniques assures that the PCT will be conservatively predicted.
I PartialImop Seal Clearing: In response to questions regarding panial loop seal clearing, l
additional SBLOCA cases were run using the plant model shown in Figure 1. Break sizes wer varied from 1.6 to 2.0-inch to study the transition from no loop seal clearing to the clea l
the broken loop. It was found that RELAP5/ MOD 2-B&W predicts this transition for breaks between 1.9 and 2.0 inches. The liquid levels in the broken loop pump suction piping for j
j two cases are shown in Figures 2 and 3. Figure 4 shows the core liquid levels for the two cases.
From Figure 4 it can be observed that the minimum core liquid levels of about 9.0-ft occur at) i about 1600 seconds and increase thereafter. For the 2.0-inch break, the core liquid level is i
10.0-ft at the time ofloop seal clearing. The loop seal spillunder elevation corresponds to 3 j
height from the bottom of the core.
I 1he steam velocity in the upside pump suction piping for the 2-inch break is shown in Fig j
j Once the steam venting process initiates, the head imbalance in the loop seal accelerates the s
}
flow and can be expected to reach a termmal velocity sufficient to clear the loop seal. For the i
inch beak the terminal steam velocity in the upside pump suction piping reaches about 10.0 ft/s at the time of loop seal clearing as shown in Figure 5. Tuomisto and Kajanto'show that th will clear completely for steam velocity greater than 6.2 ft/s (1.9 m/s) at 870 psia (60 bar). Thi is based on the floodmg criterion for large diameter vertical pipes, Kutateladze Number Ku (S Equation 5 in Reference 1) equals 3.2. This flooding criterion is defined as a zero downward flow of falling film on the tube surfaces. They also show that, at pressures above about 145 psia (
bar), vertical flooding is the limiting mechanism for loop seal clearing rather than the drople entrainment from the stratified liquid in the horizontal section of the loop seal. For the 2.0-inch break case, the system pressure is abou: 1000 psia and therefore the loop will clear for steam velocities lower than 6.2 ft/s. The 1.9-inch break case in ROSA (see response to Question 14) demonstrates the loop seal clearing mechanism diaen==~i above. For these break sizes, it is possible to am== taw some of the liquid in the loop seal once the initial acceleration of steam is complete as observed in the test. This liquid fall back is also observed in the RELAPS simulation cf the 1.95-inch break case which is discussed at the end of this section.
Figure 3 shows that the liquid level in the upside of the loop seal section starts to ds.a after about 1700 seconds. The void fractions in Nodes 255, 260, and 265 are shown in Figures 6 through 8, respectively. From these figures it can be seen that the liquid level decrease in the loop seal upside aareian is caused by the increase in void fraction in the pump vohane (Node 26 Steam venting from the loop seal occurs only after about 2200 seconds as shown in Figure 6. The pump discharEC P ping on the other hand is highly voided after about 750 seconds due to the steam i
flow from the upper head spray nozzles into the downcomer. At about 1600 seconds the break junction void fraction increases rapidly from zero to a highly vcided state and the flow in the cold 2
(_-__-_.-
1 i
l i:
leg starts to oscillate. Injection of the cold ECC water into the highly voided cold leg a l
break node amplify these oscillations. This results in a flow of steam from the cold le i
pump volume. Note that in the broken loop, up until loop seal clearing, the HPI water is in to the Node 276 (a vertical node), and the CCI water is injected into the cold leg j
equilibrium option is selected in Node 276, making Node 276 a major source of oscillation j
Stratified flow is expected in the pump discharge piping and RELAP5 allows only sm condensation when the flow is stratified. The voiding of the pump node prior to loop is discussed further in the next section.
1 i
To further study the possibility of predicting partial loop seal clearing, a 1.95-inch break c j
run. The broken loop also cleared for this case. However, some liquid remained in the section and in the pump node, possibly as a liquid film on the pipe walls that fell back after t i
high steam flow period ended. This water eventually accumulated in Node 255 as show
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Figure 9. All other nodes in the loop seal were almost completely voided. The liquid d into node 250, which represents the lowermost portion of the U-bend. This is consistent w j
discussion in Reference 1, 4
1 Pumn Nodine Sensitivity Study i
The broken loop pump suction noding for the base model is shown in Figure 10. T loop seal clearing, Node 248, representing the lower portion of the downside piping, w a small node height, I foot. The bottom of Node 248 coincides with the spill under elevation of the loop seal. Node 250 represents the horizontal portion of the U-bend and the height of j
is the radius of the pipe. Node 260 represents the pump. The height of Node 260 is 5.81 ft
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is the actual height of the pump up to the centerline of the discharge piping. In RELAP5,
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pump volume also uses the high mixing flow regime, and, therefore, slug flow (Wilson drag not used in this node, even though it is a vertical node.
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The early voiding of the pump node for the 2.0-inch break case, as discussed in the p j
section, may have been caused by the height of Node 260. To study the sensitivity of p size, the base input model was modified by dividing the pump volume into three nodes (2 259-2, and 260) as shown in Figure 11. Node 260 still represents the pump. In this case the inch break case did not clear the loop seal. For a 2.1-inch break case, the loop seal clear thout 3300 seconds. Collapsed liquid levels in the loop seal and core and the void fractions in the loop seal nodes, pump node, and the pump discharge node of the broken loop are sho
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Figures 12 through 22. From these figures the following observations can be made. Steam venting through the loop seal starts after Node 245 is highly voided. This occurs at about 1400
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anana. The void fr.cdos in Node 259-1, which is part of the actual pump, is close to the void fraction in Node 258. Node 260 is highly voided and the void fraction in node 259-2 is
~
somewhere b&~ the values for Nodes 259-1 and 260. The void distribution in the upside U bend, including the pump volume, is improved over that in the base calculation. The v i
l steam causes the liquid level in the upside of the U-bend to decrease, reducing the core level
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depression. Figures 12,14, and 15 show liquid level oscillations on the order of 1.0 foot in the down side of the U-bend from about 1500 seconds until the time ofloop seal clearing, ab seconds. The oscillations are mainly caused by the condensation of steam on the cold ECC water i
i
injected into the cold legs Rothe, Wallis, and Thrall 2 discussed the pressure due to the condensation of steam on ECC water in the cold legs. CE) and Wes 2
scale tests (See Table X in Reference 2) both show condensation induced press the order of 10 to 20 psi. Therefore, the RELAPS calculated 1.0 foot oscillations Conclucinn From this study the follo, wing conclusions are made. The transition fro to clearmg of one loop occurs within a narrow range of break sizes. Condensatio oscillations causes steam venting through the loop seal before the liquid level in section of the loop seal reaches the spillunder elevation. This substantiall of core uncovery at the time ofloop seal clearing for these break sizes. The core for the break sizes studied here.
The revised pump noding will be used in SBLOCA EM. However, this model impact previous EM studies and benchmarks.
Referene#<
1.
H. Tuomisto and P. Kajanto, "Two-Phase Flow in a Full $: ale Facility," Nu Design 107, pp 295-305,1988.
2.
P. H. Rothe, G. B. Wallis, and D. E. Thrall, Cold I2e ECC Flow Oscillations, EPRI 282, November 1976.
3.
W. E. Burchill, P. A. IAwe, and J. R. Brodnk, Steam-Water Mixine Test proeram T D: Formal Renort for Task B and Final Renort for the Reaam Relier Ph Program, CENPD-101, AEC-C00-2244-1, October 1973.
4
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Break Orientation: The break orientation, for SBLOCA studies, is placed at the bottom of the cold leg piping, between the ECCS injection location and the reactor vessel, since this
}
configuration poses the greatest challenge to the ECCS in providing sufficient coolant flow to mainemin core cooling. With the break so situated, ECCS entering the RCS through the inj i
noale in the broken cold leg must pass over the break prior to penetrating the reactor vessel.
i Unless the pump discharge piping is already full, the emergency coolant will be passed out of t i
break, unable to provide core cooling. This limits the effective ECCS flow, during critical cooling times, to that injected into the remaining ! oops (intact loops). For that reason, mos l
have limits on the amount of injection that can be delivered to any_ one loop or leg dur SBLOCA. A typical limit is that no more than 70 percent of the total ECCS flow can be delivered i
to any one injection nozzle.
t l
The issues involved with the evolution of SBLOCA transients having alternate break orientation j
are primarily concerned with the longer term management of the accident than with the measurement of the capability of the ECCS system to provide sufficient and timely injection. The j
investigation of an SBLOCA scenario with the break at the top of the pump discharge i
illustrative. For the first period of the transient-reactor trip, ECCS initiation, and loop through loop seal clearing-the LOCA is essentially the same irrespective of the break orien top, side, or bottom. The pump discharge piping is essentially full of water. Plant pressure I controlled by a balance bemeen the volumetric discharge through the break, the vapor genera in the core, and condensation in the steam generator, if that is needed. Plant inventor lost rapidly and a liquid level imbalance is being setup between the downcomer and the core in order to achieve loop seal clearing. Loop seal clearing, when it occurs, is self advancing and rapid. At the end ofloop seal clearmg, one or more loops have been cleared ofliquid; the li is retained in the core and downcomer. The downcomer core level imbalance necessary to drive steam to the break. This process, though dependent on break size, is independent of break orientation; it occurs in essentially the way same for bottom, top, and breaks. Some arguments exist that side and top breaks offer less potential for liquid diversion t the break during loop seal clearmg and, thus, arrive at a stable cleared configuration with vessel inventories than do bottom breaks. That effect, however, is difficult to demonstrate.
Following loop seal clearing, the ECCS system is challenged as to its ability to supply water at a rate sufficient to replace the water that is being boiled offin the core. In the critical cases, with i
a single failure of one of the high pressure injection systems (HPIs) and the break located at the bottom of the dieharge piping, the ECCS cannnt immediately keep pace with core boiling. Th system is then in a boildown mode. The inventory in the reactor vessel continuously decreases until the decay heat drops or the ECCS flow increases (ha~ of system depressurizadon) to the point of achieving a match with the core boiling. If the imhalance is sufficient, the core may uncover, exposing its upper regions to steam cooling before the match occurs. Modeling this phase of the transient with a bottom break is limiting because top or side breaks have effective ECC flows, that are up to 40 percent higher. Thus, for the initial system response and the determination of the adequacy of the ECCS, the bottom break is clearly the conservative choice.
27 i
w - - - -..
5 the acceptable match of decay heat and ECCS i
4 ECCS and the reactor vessel will start to refill ecrease downcorner will be refilled with ECCS water i
j the behavior of the bottom, and side and top breaks starts to differ For bo n,
break discharge is sufficient to remove exces s, the liquid i
being hydre d"Ily halsamt against the downcomer, is well covered i
the rising system water level and the excess ECCS nce j
many factors including operator action to manage the j
e pump suction
'Ihat the plant is safe and can be managed acceptably during recovery l
for the plant Emergency Operating Procedures (EOPs) or other dev concern i
recovery of the plant. The initial response of the ECCS, its Mau i
of long-term cooling have, by this phase of the accident, been es sizing, and the establishment
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of 10CFR50.46. Tbc eventual recovery from the accident, the ev i
operator actions, and their affect on the RCS and core are operational matte probable plant behavior is described; aberrant, s j
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remains for LOCA analysis past the initial ECCS respo
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breaks. Actually, even a bottom break will event
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j and top breaks since the break flow cannot be adjusted infinitely, but i
break. The main factors are:an extremely long time period. For our purpos i
I The amount of steam flow possible through the upper head spray noz c.
i This vent path, if it supports the core camia:rate, can eliminate the need for steam venting through the loops. hanne core steaming is dependent on decay hea increase in significance as time progresses j
b.
'Ihe amount of steam or water that can be passed through the reacto l
Hot side to cold side leakage is another vent path capable of eliminat need for loop venting. 'Ihis mechanism responds with time in two ways.
decreases with time reducing the amount of steam to be vented and, se nominal temperature also decreases with time, increasing the fitting ga vent capability. Care should be exercised in applying leakage credit during pa i
A uncovery since the steam in the upper head will be superheated, tending to b t
structures and reduce the gaps.
Whether the mechanism for filling the suction lines evolves gradually or it is c.
spontaneous development.
If the means for spilling water into the suction piping is the decrease in dec build up of excess injection will occur slowly and the accumulation of water in piping will be gradual. The potential for blockage will be imamed gradually a beyond which loop venting may not be needed. If, however, the increase i
rapid, u may occur because of the return to service of a failed injection syste potential for blockage can occur with reasonable rapidity.
d.
The amount of steam flowing through the loops that is not condensed in the st generators.
l This of course is the most direct factor of concern in evaluating the effect of r the loop seals. An important consideration is the degree of management credit l
steam generator pressure control is conducted as intended by the EOPs, the plan evolve to a reflux mode with no need for loop venting except where spontaneous in injection flow occur (item c).
Depending on the plant, the UHSNs can eliminate any concem over a seco process. All Westinghouse plants, classified as T upper head plants, have reasonably large UHSNs. McGuicc/ Catawba and Sequoyah are examples of such plants. An exam Sequoyah calculations for a 1.9-inch break shows that the process of loop se i
interrupted at about 2,000 seconds by the development of a head imbalance betw downcamer and the core that is large enough to support sufficient steam flow thro to eliminate the need for loop venting. For this break and breaks of smaller cross-se the loops never clear and, after achieving a minimum suction piping downside leve piping will gradually refill. Because the core swen factor (mixture level divided level) is approximately proportional to core steam generation and the differen for flow through the UHSNs is proportional to the square of the rate of steam gen elevation head difference between the core and the downcomer will decrease m the swell height dLw as decay hea' t drops. The core mixture level actually increase time, assuring contim=d core cooling. 'Iherefore, for breaks that do not requir during the initial system response, no need for clearing will develop later in the accide Further, for larger breaks that do require loop seal clearing, the ability to flow sufficien through the UHSNs will develop with time, also eliminating the need for loop s Thus. for T upper head plants, because the UHSNs have substantial capability for steam venting, no concern over the refilling of the loop seals with time exists.
i For T upper head plants, the UHSNs are not sufficient to vent a meaningful amount of Such plants can be bm= dad by consideririg the results of excess ECCS for a th I
absent UHSNs and internals leakage. To this end, an evaluation has been condu 29
without UHSNs or internals leakage and for which no operator actions the accident.
The analysis comprises an exammation of the potential condition of the following a 2-inch diameter break in the side or top of the cold leg just af I% hours into the accident, and at six hours into the accident. In each case, su elapsed for the suction piping to have been refilled to the extent predicted.
The plant is considered to be in a transient mode for the evaluation of the conditions p and in a quasi-steady-state for the evaluations at 1 % and six hours. The spectru considered are one and four loops ventmg and one or two HPIs providing mak is arbitrarily lost or spilled from the system. The tuning ofloop seal clearing was available spectrum calculations performed with the evaluation model. The timin slightly for a top break with two HPIs, but that,is not a significant simplification.
One key in understanding the analysis is to realize that a transport mechan must exist. Either the core is boiling and steam is being used to transport energy to the RCS is basically water solid and experiencing natural circulation. A water s at six hours is possible, if the operator has followed the EOPs and depressurized th generators. However, there is no concern for loop seal blockage in a circulating system I
case will not be considered further. Because steam is the transport mechanis i
and the flow rate of water to the core can be determined by balancing the heads betw suction riser section and the core given that the inlet enthalpy is specified. For thi the core inlet enthalpy was assumed to be the injection enthalpy and a level credit wa the difference in the downcomer liquid density and the core average liquid density assumption, that the core inlet is saturated, can be made with no density difference i
between the core and the downcomer. Either approach achieves essentially the same c i
1, vel. One depresses the core collapsed level less, while the other generates a highe swell. Steam generated in the core passes through one or four loops and is mixed w the pump suction riser section at the spill urrier. Here, excess ECCS subcooling enaha=
to the extent possible and any remaining non-condensed steam is bubbled up through t section to the break. For the post-loop seal clearing analysis, the pressure is taken from the reference RELAPS calculation. For the extended time evaluations, the pressure is determi from the break model (Moody or Extended Henry-Fauske) and the consideration of mass a energy equilibrium for the RCS. For the single HPI cases, the break requires peam and water to be in equilibrium and only that steam flow (the break steam) was used to lighten (decrea density) the riser section. For the two HPI cases, the HPI sensible heat was sufficient to abs all of the core heat and no break steam flow occurred. In these cases, the condensat in the bouom of the riser section was== mad o take place in an exponential pattern over t
bottom four feet of the riser section. Forty eight percent of the steam was condensed in the f one-half foot, eighty pswui was condensed by 1% feet, and all the steam was condensed feet.
The table presents the results obtained for liquid collapsed levels in the riser sections of the venting suction piping and the reactor core. The table also indicates whether or not the core is covered by the boiling mixture. As can be seen from the table, the core is essentially coveted with a boiling mixture for all cases. The one HPI, four-loop venting case has a core mixture height of 11.9 feet at six hours, which is considered essentially covered. Extending these res
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1 l
to greater times will eventually demonstrate core uncovery. However, operator action in conjunction with the EOPs has been delayed for over 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> for these analyses. Becau action will mitigate the consequences of these transients, it is not necessary to consider t response of the system for longer times.
l j
The evaluations provided are appropriate if me processes described and credited are no l
That may not be true for the condensation process in the riser sections. At that location i
3 steam being forced into subcooled water, water cannon or water hammer effects may In that event, the system can be expected to vary about the nominal conditions derived j
Core mixture levels will be both higher and lower than tl ose indicated, but, because the c t
beating at these times is not rapid, the core overall should be well cooled. Again, if follows the EOPs, the potential for these conditions will be removed early in the event.
i In summary, FTI maintains that the decision to mn 10CFR50.46 calculations for breaks bottom of the piping is appropriate.
These breaks clearly offer the greatest challenge to the emergency core cooling systems. SBLOCA transients may evolve differently for top and :ide j
breaks than for bottom breaks, but the evolution is essentially independent of the EC t
the differences occur during the period of accident management that is the purview of
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Emergency Operating Procedures and they should not be evaluated with the requ conservatisms. Not withstanding these considerations, FTI has considered the evo and side breaks. For T, upper head plants, the evolution of the transient has been shown to produce a smooth increase of core coolant level with sustained and continuous core c a possible initial uncovery. For Tw upper head plants, inter-vessel leakage around the hot leg j
nozzles serves the same purpose as UHSNs for the T plants, making long-term cooling a smooth process with no core uncovery.
j 1
Additionally, top breaks were evaluated out to 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> for a plant without UHSNs or inter-ve
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leakage. It was shown that, at least on the average, the core will be continuously covered demonstrated that the transient can progress past six hours without experiencing serious
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uncovery, requiring many additional hours to produce significant core uncovery. Because the i
potential to require loop venting in the long term is limited (UHSNs and inter-vessel le effects) and hmme the EOPs typically recommend operations to deptussurize the p j
the transient, thereby refilling the plant and mitigating any need for loop venting, FTI belie j
that any consideration of times beyond those presented to be the proper subject of op j
procedures and not suited for consideration under 10CFR50.46.
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Analysis Results for a 2-inch Diameter Pump Discharge Break at the Top o
.i Time Decay HPIs Loops Pump Core Core l
Heat Operating Venting Suction Collapsed Mixture s
Riser IAvel Level Collapsed Ierel i
bours feet feet feet 0.5 2.0 1
1 2.4 14.6
' 12 +
~
4 5.5 11.5 12+
l 2
1 3.7 13.3 12+
4 6.2 10.8 12+
1.5 1.5 1
1 4.4 12.6 12+
4 6.6 10.5 12+
2 1
7.4 11 12 +
1 4
8.2 10.2 12+
6.0 1.0 1
1 5.9 11.1 12 +
4 7.3 9.7 11.9 2
1 7.6 10.4 12 +
4 8.2 9.8 12 +
1 l
l 52
Cross Flow Resistance and Core Modeling: In our 3/28/% telecon, questi to the basis for the crossflow modeling used within the core. The modeling l
4.3.2.5 of volume D of the RSG evaluation model repon, BAW-10168 the model is a 20 axial region core, radially divided into a single assembly rernainder of the core. Each volume in the core model is connected ve Vertical resistance is based on core design factors which in turn are bas fuel assemblies. Correlations for the prediction of lateral resistances vary factor value of 2, based on the interface area between adjacent fuel assem for the evaluation model. This value produces reasonable results that. ag expectations for SBLOCA. The value, however, does not appear to be unique or larger values would also appear to produce valid results. The B&W-desig small break evaluation model uses a value of 200 for the base produce substantially differing predictions.
resistance used in the B&W-designed plant SBLOCA model m the calculation.)
Two adjustments arc imposed on the basic resistance in order to assure co predictions. For the top half of the core, the flow resistance from the average channel is increased by a factor of 10 (flow resistance from the ho( channe is not increased). This has little effect on the behavior of the below the mixture level. However, above the mixture in the steam cooling re core has uncovered, the increased resistance limits any tendency to flow st to the hot channel. It is expected that steam will flow from the hot channel to the of the higher vapor generation in the hot channel. Because flow diversion out o is a conservatism, that flow is not impeded. However, flow reversion back to the would have the effect of reducing the hot channel vapor temperature and in Although some flow reversion is expected, the resistance within the model is incr limit the effect. The factor is only applied to the upper half of the core becau buis, it is not possible to predict acceptable cladding temperatures if the top half uncovered for an extended period. This modeling adjustment, then, is taken to hel conservative evaluation.
For reasons similar to the increased crossflow resistance, the hot channel outlet r resistance was increased to a k-factor of 200 based on the assembly flow area. It that this would reduce the tendency for liquid fall brck into the hot channel b to flow into the average channel and then crossflow to the hot channel. The effective high reverse flow resistance, however, is mitigated by the need for the hydraulic s achieve a pressure hat = ace between the inlet and outlet plenums. As the flows an develop axially within the core, the hot channel innintains a slightly increased voi its higher vapor generation rates. This leads to an apparent pressure imbalance columns (hot and average channels) as the core exit is approached. To adjust the solution allows negative liquid flow into the uppermost volume of the hot i
lower void fraction for that volume. The reduced voiding in the upper volume balanc l
channel pressures.
Note should be taken that the upper two volumes of the hot or average Y
j channels do not represent nuclearly heated regions of the plant I
unpowered segments of the fuel pins (the fuel pin upper plenum and inte upper nozzle of the fuel assembly. Thus, the flow and the void reduction core active region. The resultant negative flow from the upper plenum t volume only occurs when the upper plenum contains some mixture. M are not created because once the inner vessel mixture Icvel falls into the c ha m is maintained by a slightly increased mixture level in the hot channel level in the hot channel is physically real and well modeled. Observatio level predictions for the hot and average channel discussed below j
the solution.
possible core reverse flow. The resistance does conditions but is likely to divert flow away from the hot channel under flow co would be a maiagful conservatism if SBLOCA were to involve any substa
. This core flow. Although no such period can be identified, the only reverse cor
}
resistance factor has been kept as a precaution.occurrmg during i
That the hot channel and average channel mixture heights evolve rea uncovery can be observed in the attached figures.
These figures display the axial void in a Westinghouse designed 4-loop plant over the l void fraction versus axial core elevation from the lower plenum to the upp i
elevation of the outlet nozzles. Each void fraction is displayed axially at th j
from which it is taken and is connected to the void fraction of the adj j
line. If not recognized, this technique can introduce some confusion, as o i
cf these graphs, but the linear connection to the i
}
produces a visual irnpression that the lower plenum contains steam as the bo approached. In truth there is a step change in void content beras the lower plen
}
core. The same recognition should be made in reviewing the upperplenum void in part, is the reason that the channels, except for the lower pleman to th connecting lines while the upper plenum volumes are displayed as points. The t j
figure is captured is displayed just above the figure border. Within the u i
most value is at the elevation of the center of the core outlet nozzles. This voh
}
height of the outlet nozzles. The next lower volume is entirely below th j
PIP 888-j Loop seal clearing for the case shown in the figures occurs at approximately 7 i
graphs display the core elevation head / mixture height as the necessary he i
develops on the approach to loop seal clearing and as the core refills after i
provided at 640, 660, 680, 690, 700, 710, 715, 720, and 800 seconds. By 640 i
clearing process has initiated and the core mixture level has fallan telow the no
}
indicated by the void fraction in the upper most volumes.
(The upper volume represents the i
portion of the upper plenum adjacent to the outlet nozzles.) The core is still covere and the depressed void fraction at the exit to the hot channel can be observed. It 1
CQ [
observed that the correspondence in void content between the average channel and the is quite good. Deviations occur, but the general trend is a slightly higher void content in channel. There is no indication that the lower void content of the hot chan propagated downward. By 660 seconds, more of the upper plenum is voided, but the core is stil covered and the core void distributions remain reasonable. At 680 seconds, the columns representing the hot and average channels are starting to void. The upper plenum is 100 percent voided. The core beated regions are still covered since the high voiding h penetrated below the non-heated regions of the fuel assemblies. By this time, before any heatup, the void fraction for the hot assembly upper region has evolved into agreement w of the average channel. At 690 seconds, the heated regions of the hot and average chan started to uncover. Loop seal clearing is now about 25 seconds away. Because the core outle void fraction is at 90 percent, the cladding temperatures remain near saturation.
At 700 seconds, the two upper volumes of the heated core are showing substantial the very top heated node may be experiencing some beatup. For the limited uncov here, mist entrainment from the mixture may be sufficient to prevent core beatup. The average mixture levels are in agreement as the uncovery proceeds. At 710 seconds, the mixt has fallen to its lowest level during loop seal clearing. The hot and average channel mix levels remain in agreement with the hot channel slightly more voided. At 715 seci seal for the broken loop has cleared and the downcomer and core levels are starting t creating a core refill. By 720 seconds, the refill has progressed into the upper plenum. The v fraction at the very outlet of the hot channel is again depressed but that was not observed in t partial refill at 715 seconds.
Thus, the predictions of the hot channel exit void fraction are consistent with the needs of the transient prediction, attaining the required degree of a under conditions when core uncovery is occurring or eminent. By 800 seconds, the refill is complete and the core boil down phase has been entered. As shown, the refill did not co fill the vessel. The region just below the out norzle remains at an elevated void content and the upper plenum at the outlet nozzle elevation is completely voided.
In conclusion, core modeling has been arranged to provide for hot and average channel effe Specific provisions have been incorporated into the EM to achieve conservative predictio cladding temperature (crossflow resistance for the upper half of the core). The modeling wo weh during core uncovery as evidenced by the agreement h,.e the hot and average chann mixture levels. Although a modeling factor does lead to an apparently inconsistent void fraction in an upper unheated volume of the hot channel during those phases of the SBLOCA transient l
when the upper plenum contains mixture, this difficulty is resolved as the core uncovers and is l
not present at any time that the calculation is predicting core uncovery or calculating cladding l
temperature excursions. Therefort, the core modeling approach employed is appropriate for th I
calculation of small break LOCA simulations.
i l
l 35
CORE VOID DISTRIBUTION - 3 in Break 640 seconds 1
X X
0.9 0.8 x
0.7
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CORE VOID DISTRIBUTION - 3 in Break
~
680 seconds 1
x ^ x x
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CORE VOID DISTRIBUTION - 3 in Break 690 seconds 1
^ x x
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CORE VOID DISTRIBUTION - 3 in Break 700 seconds 1
~*"x
^ X x
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.9 0.6 Um LE 0.5
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CORE VOID DISTRIBUTION - 3 in Break 715 seconds 1
r x ^
x x
0.9 0.8 i
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CORE VOID DISTRIBUTION - 3 in Break 720 seconds 1
x x
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0.8 0.7 5
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10 12 14 16 18 Elevation, ft Hot Channel x
Ave Channel
CORE VOID DISTRIBUTION - 3 in Break 800 seconds 1
x 0.9 0.8 0.7 5
0.6
/
2 0.5 x
g E
0.4 -
x 0.3 x
1 0.2 -
0.1 -
0
-2 0
2 4
6 8
10 12 14 16 18 Elevation, ft Hot Channel x
Ave Channel
Supplementary Break Orientation Information:
Range of Upper Head Spray Nozzle Areas:
T-hot Plant
=>
= 0.02 ft: (Trojan, No-3 Anna, Surry, etc.)
T-cold Plant
=>
= 0.45 ft2 (McGuire/ Catawba, Sequoia, etc.)
Some plants sit in between these limits with areas of 0.2 or 0.3 ftz, The inclosed plots are for the 2.1 inch case that was provided in an earlier communicati i
that with them being part of a larger set they would be more useful. If the specific 2 in is important we can reconstruct it and send the same plots. Some of the definiti are:
UP Upper Plenum V
Volume or Node AC Average Channel HC Hot Channel CVAR Control Variable For the case of AC CVAR and HC CVAR the display is a collapsed water level for the core region with 0.0 taken at the bottom of the active region.
The reason that the values exceed 12 feet is the inclusion to the t unheated volumes of the fuel assemblies that model the fuel pins above the l
uranium pellets and the upper nozzle of the fuel assembly.
Jun Junction or Flow Path J
Junction or Flow Path UH SPRAY Upper Head Spray Nozzle IL ECC Intact loops ECCS flow BL ECC Broken Loop ECCS flow For this case IL ECC CVAR and BL ECC CVAR are simply the high pressure injections. Had the plant depressurized these control variables would have picked up the accumulators and the low head systems.
HOT CH Hot channel, HOT CH, CVAR is a control variable that approximates the mixture level in the core hot channel. For the purpose of this CVAR mixture is defined as a < 0.9. The control variable samples the a from l
l the bonom to the top in each node of the channel. If a is less than 0.9 the l
height of the volume is considered mixture once a is greater than 0.9 the control volume is considered as above the mixture and the search stopped.
l
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0 5
4 n
3 6
V r
P D
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RS/2 2.1 INCH PD BREAK Split BL Pump Vol 260 RELAP5/ MOD 2 Ver 20.0HP I6 IEf3END
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- AC CVAR 301 14-
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RS/2 2.1 INCH PD BREAK Split BL Pump Vol 260 RELAPS/ MOD 2 Ver 20.0HP 400 l
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_ LEGEND IL ECC CVAR 17 350-O BL ECC CVAR 18 300
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RS/2 2.1 INCH PD BREAK Split BL Pump Vol 260 RELAP5/ MOD 2 Ver 20.0HP 32 i
}
__ LEGEND 7
-- HOT CH, CVAR 928 a-
+
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........... -. - + + ~. - -.. + -.. - - -. *
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y
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0:
0 800 1600 2400 3200 4000 4800 5600 6400 Time (sec) t TATH VWun <vrais n en
1 SBLOCA Imag-Term Cooling: In our 3/28/96 telecon, Bob Jones raised an issue as t j
sumciency of FTI's SBLOCA long-term cooling write up on page 8-1 of BAW-10168, R j
2, Volume II. He indicated that the appropriateness of the methodology was diffi relative to the criterion of 10CFR50.46.
As stated on page 8-1, FTI's SBLOCA long-term cooling methodology is basically the same as that tud for LBLOCA and discussed in deta page 8-1 of Volume I. It is repeated below.
FTI contmues its transient small break LOCA computer analysis until the' core is 4
mixture and the clad i.uyeretures have decreased to the coolant saturation temperature.
long-term, the clad will be maintainad within several degrees of the coolant saturation i
by a continuous flow of ECC water. Each plant has established NRC-approved procedu e
1 an orderly transition to long-term cooling, assurmg a continuous flow of ECC water to the re 1
vessel and preventing the crystallization of boric acid in the core. The plant procedu
}
the operator actions necessary to switch to sump recirculation-providing for a continuous j
flow-and to assure a throughput of water to the core-maintaining boric acid concentrations j
below previously-established acceptable levels.
FTI plant applications performed under BAW-10168 will validate the appropriateness of previously-established operator action times, assuring the effect,ive establishment of l cooling. If the need for new operator action times is demonstrated, analyses necessary to l
will be performed for and reported in the plant-specific LOCA application. For SBLOCA, su 4
calculations are usually unnecessary, since, in general, it is bounded by LBLOCA predictions that analysis is used to satisfy the long-term cooling criterion. In FTI's approach, the LO j
procedure interface is properly addressed and in cornbination with as-designed plant i
systems requirements the long-term cooling criterion of 10CFR50.46 is satisfied.
Equilibrium Core Heat Transfer Calculations: FTI's original NRC-approved evaluation model (for both large and small breaks)- BAW-10168, Revision 1-used equilibrium conditions for the RELAPS computation of core heat transfer; this issue was thoroughly explored by the INEL reviewers and it was approved by the NRC. In Revision 2 of the EM, FRAP-T6 was deleted from the large break LOCA calculational technique. No changes were made to the core heat transfe package other than the calculations for the hot channel were now performed in RELAP5. The modeling was still based on equilibrium and it was found to be =ceaptahle for limnaing use NRC. In Revision 3, FRAP-T6 was deleted from SBLOCA. Again, no changes, other than code location, were made to the equilibrium core heat transfer package.
When the RSG evaluation model was originally assembled, FTI installed in RELAPS core heat transfer correlations, covering most of the boiling curve, that were fornmlatad based on
~=uihnum condidoes. The RELAP5 core heat transfer package, designed after that in FRAP-T was used and approved for both large and small break applications. The EM was hanchmarked, most recently against ROSA IV, and shown to produce conservative PCTs. FTI understands that it could upgrade RELAPS to a nonequilibrium core heat transfer calculation, but it would a substantial investment (code revisions, benchmarks, topical report revisions, and licensing) a there is no identified calculational or safety benefit to such a modification. Therefore, FTI has 55
e O
decided to continue to use the equilibrium option. The T-H role of RELAPS is an equilibrium core heat transfer calculation, previously found acceptaNe in FRAP-T being used and has already been approved for LBLOCA calculations. The RE I
approach is NRC-approved and the removal of FRAP-T6 from the SBLOCA EM on its continued validity.
l
)
5+
-