ML19316B225

Westinghouse Multiple Aperture Small Break ECCS Analyses.

ML19316B225
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Site:	McGuire
Issue date:	05/31/1980
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WESTINGHOUSE MULTIPLE APERTURE SMALL BREAK ECCS ANALYSES May 1980 8006120VIS

INTRODUCTION In the course of extensive plant licensing activities Westinghouse has continued to study the general behavior of small break loss'of coolant accidents (LOCA) to ensure safe operation of a Westinghouse NSSS in case of such an accident. Many design analyses have been performed to under-stand and identify the most limiting small breaks under Appendix K criteria. These analyses have b'en presented in various Westinghouse ,

topical reports and in each plant's FSAR.

This report documents analytical work in a new area of small break work, that being the multiple break domain. Until recently all studies of small break behavior have been limited to the modeling of one rupture in the NSSS. Previous comparative studies of break spectra for small hot and cold leg breaks have shown cold leg breaks are limiting with respect to core uncovery and the resulting peak clad temperature. The purpose of this report is to show that multiple loop breaks are also bounded by an equivalent size cold leg break, and therefore evaluation of the worse case cold leg breaks is sufficient to ensure safety.

All cases presenLd in this report, (whether developed for this report or used from other reports) have utilized the October 1975 version of the Westinghouse ECCS Evaluation Model for small break LOCA, which was previously submitted, reviewed, and approved by the NRC as complying with 10CFR50 Appendix K. This model is documented in Reference 2 and 3.

A three loop Westinghouse NSSS plant was selected for the analyses pre-sented in this report. This plant was selected since, in general three loop plants yield relatively high peak clad temperature results compared ,

to other plant designs. This is true because of the relatively high power rating per loop and lower safety injection rate per loop in this type of Westinghouse plant. This will allow the application of the conclusion reached in this report to other Westinghouse plant designs, specifically the Westinghouse two and four loop plants. The selection of the three loop plant will also enable the use of various cold leg breaks analyzed in Reference 1 for the break spectrum comparison in this report.

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Appendix K assumptions were utilized for all analyses presented in this study. Since all simultaneous breaks analyzed include a cold leg break, it was assumed that one safety injection line was spilling. This'is a conservative assumption since all pumped safety injection flow in the broken loop is assumed to spill to the containment. This would be the case only where the break is on or near the safety injection line. For all cases with a cold leg break the ECCS accumulator on the broken loop is not modeled and is assumed to discharge into the containment also.

Finally, it was assumed that no offsite power was available and that one diesel failed to start, making one train of safeguards equipment ,

inoperable. \

For the simultaneous hot and cold leg break cases presented in this report, hot leg nodes were added to the system nodal network described in Reference 4 and 5. These nodes were added to better represent the phenomena occurring for breaks in the hot leg.

Before this change to the model was used, a comparison was made ,to verify that the addition of the hot leg nodes in the broken and unbroken loops has an insignificant effect on the small break transient. Once this had been done, the multiple breaks were compared to the equivalent cold leg break cases.

In addition to the above mentioned comparisons, a model improvement which gives more accurate results for relatively "large" small cold leg breaks and which has little effect on " smaller" cold leg breaks will be documented. The change has to do with how the flow path connecting the downcomer to the cold leg break node is modeled. This change will be incorporated in the analyses performed for this report.

This report consists of three sections. The first section will contain the change to the flow path connecting the downcomer to the cold leg break node. In this section an eight inch cold leg break was analyzed with and without the model improvement to show the effect it has on that size break. In addition, a three inch cold leg break, (the limiting cold leg break for this type plant) was analyzed to verify that this I.2~

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flow path model change has little effect on " smaller" cold leg break cases as mentioned earlier. The results are documented in Section 1 of the report.

Following Section 1 will be the verification of the new small break model that includes hot leg control volumes. To verify the hot leg nodes, two different size cold leg breaks were studied with and without hot leg nodes. They were a "large"'small cold leg break, (eight inch ,

diameter) and the worst case three loop generic small break, (three inch diameter cold leg break). The results of this portion'of the report can be found in Section 2.

Finally, the multiple break comp;rison study will be documented in the last section, Section 3. A spectrum of break sizes were analyzed to assure that the worse case simultaneous hot and cold leg break was bound and that it was less limiting than its equivalent size cold leg break.

The following multiple break comparisons were made:

1. The simultaneous 2.12 inch hot leg and 2.12 inch cold leg break (Case 5) which was compared to the 3.0 inch cold leg break docu-mented in Reference 1, (Case B).

2. The simultaneous 2.83 inch hot leg and 2.83 inch cold leg break (Case 6) which was compared to the 4.0 inch cold leg break. docu-mented in Reference 1, (Case C).

3. The simultaneous 4.24 inch hot leg and 4.24 cold leg break (Case 7) wh'ch was compared to the 6.0 inch cold leg break documented in Reference 1, (Case 0). ,

4. The simultaneous 8.0 inch hot leg and 8.0 inch cold leg break (Case 8) which was compared to the 11.0 inch cold leg break (Case 9).

All the cold leg break results obtained from Reference 1 utilized the same assumptions that were used in this report. Plots of the important l

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parameter transient response have been attached for Case 1 through 9.

The important input. parameters ~ used for these small break cases can be found in Table A.1. Table A.2 lists the nine cases analyzed for this report and Table A.3 lists the figures included in the attachments.

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1.0 DOWNCOMER TO COLD LEG FLOW PATH MODELING TECHNIQUE In this section of the report a 3 inch and an 8 inch cold leg break were analyzed utilizing two different modeling techniques for the flow path that connects the downcomer control volume to the cold leg break node.

The two techniques considered for flow path No.15 were the " continuous" flow path and the " point contact" flow path (the latter has been used in the past). The major difference between these two modeling techniques ,

is sumarized below. A point contact flow path will only allow either all steam or all two-phase mixture into or out of a control volume.

Therefore, until the mixture level in the control volume drops below the point of contact of the flow path, only two-phase liquid can leave the control volume. Contrary to this, a continuous flow path will allow both steam and two-phase liquid to enter or leave a control volume at a rate proportional to where the mixture level is between the top and bottom elevation of the flow path. Therefore, there is a range of mix-ture levels where the control volume can purge a varying ratio of steam and two-phase mixture together. This range is equivalent to the diameter of the continuous flow path considered.

For "sufficiently large" small cold leg break analyses, a situation may arise where the continuous flow path techniques would be required to obtain accurate results. Here "sufficiently large" is defined as large enough for the steam flowing through the intact loop to significantly ,

affect the downcomer :ixture level. This can be seen by observing the 8 inch cold leg break with flow path No.15 as a point contact path, l, Case 1. Refer to the attached plot for this case which are Figures 1.1  !

through 1.26. .

Notice the downcomer mixture level plot, Figure 1.5. At approximately 120 seconds and again more noticeable at approximately 335 seconds, the downcomer mixture level drops quite suddenly. This is an unrealistic phenomenon caused by the point contact flow path connecting the downcomer to the cold leg control volume and the method the WFLASH computer code utilizes to distribute steam coming into a control volume.

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The WFLASH code takes all steam coming into a two-phase mixture and distributes it as bubbles, uniformly, throughout the two-phase mixture in the contrcl volume. Keeping this in mind let's look at what.is occurring just p-ior to 335 sec into the transient. The downcomer con-trol volume was modeled as sketched below:

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During this part of the transient the RCS is slowly depressurizing after the first accumulator injection. The downcomer mixture level has stab-ilized at the exit elevation of flow path No.15. This would be expec-ted since the break in the cold leg node causes flow to occur from the downcomer to the break via path No. 15. Flow path No. 15 is a point contact flow path located at the cold leg centerline in this case, so the downcomer mixture level will be at the exact point of contact of -

that path. This enables flow path No. 6 to bring steam into the two phase mixture in the downcomer, since the intact loop seal has drained.

Because flow path No. 6 is a continuous flow path which is half covered at the entrance to the downcomer, the steam coming into the downcomer from No. 6 is split. Part of it is put into the two-phase mixture and part into the saturated steam phase above the downcomer mixture level.

The proportion is a function of the area fraction in the cross section l

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1 of the continuous flow path as determined by the location of ths mixture i level in the downcomer node. For this case, there is approximately a  !

50-50 split.

As mentioned before, the steam put in the two-phase mixture is uniformly distributed as bubbles throughout the two-phase mixture in the control volume. Therefore, an average steam bubble which actually comes into the downcomer mixture very near the. mixture surf ace elevation js unreal-istically put in the middle of the downcomer two-phase mixture in the WFLASH calculation. Thus, it has to travel approximately thirteen feet to the surf ace to escape the two-phase mixture, instead of less than two feet as it would in a realistic situation. This phenomena will keep more bubble mass in the mixture than it should, which overpredicts the void fraction in the downcomer mixture.

Because of this excess bubble mass, the downcomer mixture level is T

l incorrectly kept at the flow path No.15 point of contact, spilling two-phase mixture out to the cold leg break each time the level froths up over that elevation. This leads to an unrealistically high break

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flow for this part of the transient which causes the RCS mass depletion to be overpredicted. Eventually, due to downcomer liquid mass deple-tion, the mixture level can no longer be maintained at the nozzle eleva-

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tion and it f alls below the bottom of the cold legs.

Steam bubbles are no longer put into the downcomer's two-phase mixture, but rather are assumed to go into the upper saturated steam phase directly to the broke 1 loop nozzle. Now, because of the high void frac-tion, the bubble escape rate in the downcomer is very large causing the buoble mass and the void fraction to decrease very rapidly, (see Figure .

A.1). This causes the downcomer mixture level to fall. Since steam is no longer being held up in the downcomer, the pressure balance between the core and the downcomer changes. The core begins to empty into the downcomer, until the hydrostatic pressure balance is once more obtained. While the flow between the downcomer and core is reversed, the core's mass depletion rate increases, adversely affecting I

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insignificant effect on the core uncovery and any other important tran-sient characteristic which is a function of the peak clad temperature.

The 3 inch cold leg break was chosen because it was the limiting break for the Westinghouse three Loop NSSS plant, as was dccumented in Refer-ence 1. Refer to the attached plots for Case 3, (i.e. the 3 inch cold leg break with flow path No. 15 continuous), Figures 3.1 through 3.27.

When compared to the results of the 3. inch cold leg braak in Reference 1, (Case B) it is obvious there would be little effect on the peak clad temperature calculated for Westinghouse plants. The similarity between the core mixture level transients for the two cases reinforces this conclusion. A comparison of the important small break transient char-acteristics for both cases can be found in Table 1.1.

The 3 inch cold leg break with flow path No. 15 continuous also has hot leg control volumes added to its small break model. This, however, is shown to have an insignificant effect on small break transients in Section 2 and can therefore be ignored.

Conclusions The information documented in this portion of the report, has shown that for a "sufficiently large" (approximately 6 to 8 inch diameter or larger) cold leg break, it is necessary to model both the flow' paths connected to the downcomer at the cold leg elevation (i.e. flow path No.

6 and No. 15) as continuous flow paths.

In addition, it has been shown that for "sufficiently small" (less than 6 to 8 inch diameter) cold leg breaks, this modeling change has little ,

effect on core uncovery or any other important small break transient characteristic which is a function of the peak clad temperature.

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the core mixture level transient as can be seen in Figure 1.3. This can lead to larger core uncoveries and unrealistically high peak clad temperature results.

Therefore, for sufficiently large cold leg breaks, a situation arises where the point contact flow path connecting the downcomer to the cold leg node is no longer suitable. "Sufficiently large" can now be defined ~

as approximatel; ' ' 8 inch diameter cold leg breaks because they are 1arge enough to drain the intact loop seal and artifically hold the downcomer mixture level at the point of contact of flow path No.15.

This problem was solved by using the continuous flow path modeling tech-nique for flow path No. 15. This can be seen by comparing the 8 inch cold leg break plots with flow path No.15 continuous to those with flow path No. 15 point contact. Refer to the attached plots for Case 2 (Fig-ures 2.1 through 2.26) and case 1 (Figures 1.1 through 1.26) respectively.

From the core and downcomer mixture level plots, (Figures 1.3 vs. 2.3 and 1.5 vs 2.5 respectively) the effect that the flow path modeling technique has is best shown. Notice with flow path No. 15 continuous, J the downcomer mixture level (Figure 2.5) is no longer artificially held at the cold leg centerline elevation from 100 to 250 seconds anynore, but tends to seek a more realistic level determined by the upper downcomer steam flowrates.

This can be explained by the fact that when the intact loop seal drains, the downcomer mixture level is predicted to drop to the bottom elevation of the cold leg and the steam coming into the downcomer from flow path No. 6 will be put into the saturated steam phase above the two-phase mixture. This enables the downcomer mixture level and void fraction to be accurately calc'ulated and shows, in Figure 2.3, that the second major core uncovery should not have been predicted.

In the beginning of this section, it was mentioned that a 3 inch cold leg break was also analyzed utilizing the above two modeling techniques for flow path No.15. This was done to verif' that for " smaller" cold leg breaks the modeling technique used for this flow path has an 1.4 5106A

2.0 HOT LEG N00E VERIFICATION In this portion of the analysis, the effect that the addition of hot leg nodes have on a small cold leg break transient will be inve.tigated.

The purpose of this comparison is to show that the addition of hot leg nodes to the system nodal network for a small cold leg break will have i

very little, if any,.significant effect on the results.

This verification was required to alleviate any concerns about the effect that hot leg nodes could have on the multiple break comparison discussed in Section 3.

Two cases were analyzed for this portion of the report. They consist of a "large" small cold leg break and the worst cold leg break for a Westinghouse three loop generic plant. The two cases are:

1. 8 inch cold leg break with and without hot leg nodes, Case 4 and 2 respectively.

2. 3 inch cold leg break with and without hot leg nodes, Case 3 and B respectively. Case B was taken from Reference 1.

As mentioned in Section 1, Cases 2 through 9 utilized the continuous flow path model for the flow path connecting the downcomer to the cold leg node, see Table A.2. The 3 inch cold leg break case without hot leg nodes taken from Reference 1, does not have this new flow path model but this has no effect on the results for "sufficiently small" cold leg breaks as was shown in Section 1. The following discussions are provided te point out similarities in the important characteristics and transient responses for each comparison.

8 Inch Cold Leg Break Plots of the transient response for the 8 inch cold leg break without hot leg nodes can be found in the attached plots for Case 2, Figures 2.1 through 2.27. Plots of the 8 inch cold leg break with hot leg nodes are 2.1 5106A.

Figures 4.1 through 4.27. As expected, the two transients show very good agreement as can be seen from the core mixture level plot (Figure 2.3 and 4.3). The only discrepancy significant enough to discuss is the difference in the height that the core mixture level stabilizes at after an accumulator injection. Notice for the case with hot leg nodes, the core mixture level is lower; 15.83 feet verses 17.04 feet for the case without hot leg nodes. This 1.208 foot difference can be attributed to .

the type of flow path model connected to the core at the hot leg elevation because, for this size break the RCS refills to that level af ter an accumulator injection.

For the case without hot leg nodes there is a " point contact" flow path connected to the core at the 17.04 foot elevation. For the case with hot leg nodes a " continuous" flow path is used instead so that the heterogenous hot leg control volumes will completely drain during the transient. The bottom of the continuous flow path is 15.83 feet.

Therefore, the RCS is simply filling up to the hot leg flow path's bottom elevation and any additional mass is spilling into the loops.

Thus, the core mixture level difference is caused by the 1.208 foot difference in the flow path model connections.

This difference, however, has an insignificant effect on the core uncovery or any of the other important transient characteristics that is a function of the peak clad temperature. This can be seen in Table 2.1 which gives the chronology of events and peak clad temperature, axial location, and time for the case listed above.

3 Inch Cold Leg Break .

Plots of the transient response for the 3 inch cold leg break with hot leg nodes, Case 3, can be found in Figures 3.1 through 3.27. The 3 inch cold leg break without hot leg nodes was analyzed in Reference 1 and its plots can be found there. Again, the transients are very similar even though there is the additional difference of how flow path No. 15 is modeled.

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The only differences in the transient response for this cast can again be attributed to the difference in the flow paths connected to the core at the hot leg elevation. Notice for the case with hot leg !1 odes t.he core mixture level stabalizes at 15.83 feet (the bottom of the core hot

leg exit elevation) because of the continuous flow path model. While for the case without hot leg nodes the core mixture level stays at 17.04 feet'because of the point contact flow path connected to the core. As for the eight inch cold leg break case, these differences have little -

effect on the small break transient which can be seen in Table 2.1.

Conclusions With the information documented in this portion of the report, it can be concluded that the incorporation of hot leg nodes into the WFLASH computer model will have an insignificant effect on the resultant peak clad temperature for Westinghouse snall break analyses. This was shown to be true for the worst and also for a relatively "large" small cold leg break on the three loop generic plant.

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3.0 MULTIPLI

APERTURE BREAK COMPARISON In this section of the report, a study will be made of the simultaneous hot and cold leg small break transient. A spectrum of simultaneous break sizes will be considered and compared to their equivalent size cold leg break. As mentioned in the introduction, a spectrum of cold leg breaks are currently analyzed for'each Westinghouse NSSS design, to ,

ensure their safe operation in case of such an accident. The purpose of this analysis is to show that the spectrum of cold leg breaks, (men-tioned above) are limiting for peak clad temperature considerations, when compared to the equivalent size multiple aperture break case.

The break comparisons are:

1. 2.12" diameter hot leg plus 2.12" diameter cold leg break (Case 5) which was compared to a 3" cold leg break (Case B in Reference 1).

2. 2.83" diameter hot leg plus 2.83" ciameter cold leg break (Case 6) which was compared to a 4" cold leg break (Case C in Reference 1).

3. 4.24" diameter hot leg plus 4.24" diameter cold leg break (Case 7) which was compared to a 6" cold leg break (Case D in reference 1).

4. 8.0" diameter hot leg plus 8.0" diameter cold leg break (Case 8) f which was compared to an 11" cold leg break (Case 9).

As mentioned in Section 1, Case 5 through Case 9 all utilized the con-tinuous flow path model for the path connecting the downcomer control ,

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! volume to the cold leg break node (i.e. flow path No. 15). The small break transients taken from Reference 1 do not use this modeling tech- 1 nique for flow path No. 15. They are, however, still applicable for i this application. This 'was proven in Section 1 where it was shown that l for "sufficiently small" cold leg breaks the modeling technique used for I flow path No. 15 is insignificant. "Significant small" was defined as small enough for the steam flowing through the intact loop to not sig- l nificantly affect the downcomer mixture level. The 6 inch cold leg 3.1

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t break from Reference 1 is large enough te clear the intact loop seal but l before the transient can be affected to any significant degree the accumulators inject and alter the transient. Therefore, the difference in the modeling technique used for flow path No. 15 can be neglected.

All the cases considered, including those taken from Reference 1, ,

assumed minimurn safeguards safety injection. Loss of offsite power is also assumed to occur at the reactor trip time, therefore, the only .

means of venting steam on the secondary side is through the steam generator safety valves. The analytical model and all other analysis assumptions are in confonnance with Appendix K criteria.

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The following discussions are provided to point out important character-istics and transient responses for the simultaneous hot and cold leg small break cases. Similar discussions for the cold leg break tran-sients can be found in Reference 1. Table 3.1 gives the chronology of events and peak clad temperature, axial location, and elapse time for each comparison listed about.

Simultaneous 2.12 Inch Diameter Hot and Cold Leg Break Refer to the attached plots for Case 5 for the important transient responses, Figures 5.1 through 5.27. For this postulated break combi-nation, the Reactor Coolant System (RCS) will depressurize rapidly early in the transient, and an automatic reactor trip and safety injection signal will be generated based on low pressurizer pressure. During the early stages of the depressurization, when the system is still full of two-phase liquid, the break flow, which also will be mostly liquid, is not capable of removing all of the decay heat generated in the core. .

Therefore, the early depressurization is limited by steam generator energy removal considerations, and the RCS pressure will temporarily hang up above the steam generator safety valve setpoints, assuming no steam dump is available. The system pressure stays at this level in order to provide a temperature difference from the primary to the secondary system so that core heat may be removed by the steam ,

generators.

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During this early portion of the transient, the pumped safety injection flow is less then the combined break flows, and there is a net loss of mass from the system. This causes the RCS to drain, starting from the top of the steam generator tubes which leads to voiding throughout the primary system. The rate of system drain is determined by the net loss of liquid inventory, a function of both high pressure injection (HPI) and the combined break size. Prior to the occurrence of draining, heat is removed mainly from the steam generator through continuous two-phase natural circulation, with two-phase mixture flowing over the top of the steam generator tubes. This is also typical for a small cold leg break transient as can be seen in Reference 1. The presence of a hot leg break does not affect two phase natural circulation during this part of the accident.

As the draining phenomenon progresses, the mixture level in the primary system eventually drops below the steam generator tubes. The hot leg break uncovers, turning to steam flow and the natural circulation mode of heat removal as just defined, ceases. This occurs at approximately 350 seconds into the transient. Now that the hot leg break flow has gone all steam, it is capable of removing a much greater percentage of the core decay heat, but for this size hot leg break it is not capable of removing all of it. Therefore, the steam generator is still relied upon for heat removal by the condensation mode. However, now a smaller amount of heat removal is necessary, and with minimum auxiliary feed-water available, the steam generator secondary side wi begin to slowly depressurize below the steam generator secondary side safety valve set-points. The primary side will also slowly depressurize along with the secondary side, but will remain slightly above the secondary side in

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pressure to maintain the necessary temperature difference to remove the portion of core heat that the breaks and subcooled safety injection (SI) can't remove. This can be seen in the attached RCS pressure and steam generator secondary pressure plots, Figures 5.1 and 5.2 respectively.

Almost immediately after the primary pressure stabilizes above the steam generator safety *.alve setpoints, the addition of the hot leg break 3.3 5106A -

begins to have a noticeable effect on the small break transient. This can be seen by the early decrease in the steam generators secondary pressure, Figures 5.2. The early decrease iri the steam generator secondary pressure is caused by the additional vent path for steam gen-erated in the core, created through the hot leg break, as mentioned above. Now there is no need for the broken loop seal to clear before this phenomenon can occur as was required for the 3 inch diameter cold leg break. Also, since the hot leg break has been uncovered, less liquid break flow is occurring than for the equivalent cold leg break.

Therefore, the system' drains slower. The loop seal can be expected to drain later for. the multiple aperture breaks relative to the equivalent cold leg break. Also, the predicted core uncovery will be later and the core will be at a lower decay heat level when uncovery occurs. Thus, the heat up due to core uncovery is less than for the equivalent cold leg break.

With the mixture level in the primary system below the stecm generator tubes at 350 seconds the loop seals can begin to drain and will continue to drain until the flow path at the bottom of the loop seal uncovers at about 700 seconds. (For the equivalent cold leg break, this occurred at about 550 seccnds.)

Because there is also a cold leg break in this scenerio, the core mix-ture level decreases in conjunction with the loop seal draining. The core level will decrease to the poin*. where the loop seal has cleared.

At this point, approximately 700 seconds into the transient, the core pressure decreases relative to the downcomer pressure, and the hydro-static head in the downcomer recovers the core. The minimum core level is dependent on the relative height and fluid density between the core .

and the loop seal, as well as the loop pressure drop between the core and the loop seal. This loop pressure drop is sma..er for the multiple aperture break relative to the equivalent cold leg break since most of the steam is being taken out of the system via the hot leg break.

Therefore, the amount of mass forced out of the vessel by the blowing of the loop seal is smaller for the multiple aperture break. The important l

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effect is that upon draining the loop seal, the core will recover to a higher level and the subsequent core uncovery will be less savere. The core mixture height plot (Figure 5.3) illustrates this effect.

Soon af ter the loop seal clears, the downcomer mixture level has dropped below the bottom elevation of the cold leg, see Figure 5.5. Now steam, instead of two-phase liquid is being fed to the cold leg break node which by 720 seconds has uncovered the cold leg break. This can be seen by a rapid decrease in break flow on the cold leg break flow rate plot (Figure 5.14). This can also be seen, on the break flow quality plot, (see Figure 5.20).

When the cold leg break has gone t'o all steam, the total break flow is much greater than the incoming safety injection and there still is a net loss of liquid from the primary system. Because of this net loss of mass, the core mixture level will once more decrease, eventually uncovering the core by 920 seconds. Again, the ;'resence of a hot leg break will lead to less core uncovery than for the equivalent cold leg break. As mentioned before, with a hot leg break there is less steam binding and thus less driving force is necessary to push steam through the loops. This driving force is normally provided by a difference in static head between the core and downcomer. Thus, with less steam binding, the core mixture level can be higher at any point in time.

This leads to less heating up of the fuel cladding as can be seen by the peak clad temperature plot, Figure 5.27.

With the cold leg break all steam, nearly all the core decay heat is now being removed fran the system via the break flow paths. This, combine <4 with less heat transfer from the pa-tially covered core to the primary fluid, results in an increase in the RCS depressurization, which shortly results in an accumulators injection at 1450 seconds of the transient.

As the accumulators injects additional water. into the primary system, the core mixture level eventually recovers and the peak clad temperature transient turns around. This added mass increases the RCS pressure since the core is recovered and more steam is being generated again, and 3.5 5106A

eventually the accumulators shut off. Since the total break flow is still larger than the pumped safety injection flow, the core mixture level will decrease once again due to primary system mass depletion.

Finally at 1900 seconds, the RCS has depressurized to the point where safety injection flow and total break flow are equal and there is no further net loss of mass from the system. The core mixture level will gradually rise and no further core uncoveries will occur. The RCS .

pressure will eventually stabilize at approximately 475 psi wh'ch is the pressure at which the safety injection flow matches the break flow.

This is defined as the end of a small break transient.

A maximum clad temperature of 14520F was reached 1478.5 seconds into the transient at the 10.75 foot elevation. As expected, the equivalent size cold leg break gave worse results, as can be seen in Table 3.1, which compares the iTiportant small break transient parameters for these two cases.

Simultaneous 2.83 Inch Diameter Hot and Cold Leg Break The characteristic behavior made for this multiple break case is very similar to the 2.12 inch diameter multiple aperture break case described earlier, with a few important deviations. This portion of the report will not go into as detailed an explanation as the previous case, but will instead make reference to the previous description and explain in detail only the significant deviations in their behavior. Refer to the attached olots for the 2.83 inch diameter multiple break case, (Figures 6.1 through 6.27) for the important transient responses in Case 6.

The early depressurization and RCS stabilization above the steam gener-ator safety valve setpoints as described previously, also applies for the simultaneous 2.83" diameter hot and cold leg break. This phenome-non, however, occurs slightly earlier in the transient due to the greater total liquid break discharge rate for this case. The RCS draining phenomenon also occurs similarily.

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s However, when both breaks have turned to all steam all the core decay heat is being removed from the primary system without the help of the steam generators. This eventually leads to the RCS primary pressure dropping below the secondary side pressure which was not observed for the smaller simultaneous break case. Refer to Figures 6.t and 6.2 to observe this phenomenon.

~* For this case, a maximum clai temperatiure of 1470.50F was reached .

802.5 seconds into the transient at the 11.0 foot elevation. Again, the 4 inch diameter equivalent cold leg break was more limiting as can be seen in Table 3.1.

Simultaneous 4.24" Diameter Hot and Cold Leg Break The characteristic behavior mode for this size multiple break has some significant deviations from the two cases described earlier. Again, this discussion will not repeat all the previous discussion, but will make reference to the previous behavior description and will explain in detail the significant deviations in behavior. Refer to the attached plots for the 4.24 inch diameter multiple break case (Figures 7.1 through 7.27) for the important transient responses in Case 7.

As observed before, the larger break a'ca leads to a greater total liquid break discharge rates, which causes an even earlier depressuri-zati in and stabilzation of the RCS pressure above the steam generator safety valve setpoints. The RCS draining phenomenon also occurs very similar to the smaller break cases. For multiple break cases in this size range, the hot leg break is capable of removing all the decay heat from the core the instant that the break flow goes to all steam. This ,

enables the RCS primary pressure to rapidly drop below the secondary side pressure at 80 seconds such that the heat transfer is from the secondary to the primary system, see Figure 7.1 and 7.2.

Again, as the loop seal clears, the core undergoes a rapid recovery, followed i:nnediately by another phase of core uncovery due to the net

' icss of mass from the system. Refer to Figure 7.3 in the attached plots 3.7 5106A

I to see this core mixture level transient. The cold leg break turns to ,

all steam by 180 seconds, which is soon after the loop seal clears. l This leads to an even more rapid depressurization of the primary system now that both breaks are steam. This rapid depressurization causes the accumulators to inject by 290 seconds, which is before the core has had time to uncover to any significant extent. As the accumulators sub-cooled water comes into the downcomer, control volume steam above the two-phase mixture is condensed. Condensing this steam lowers the down- -

comer pressure which disrupts the hydrostatic pressure balance between the downcomer and the core.

This leads to a temporary condition where the core is emptying fluid into the downcomer which cause the core mixture level to drop. Event-ually the accumulator fills the downcomer, the pressures balance, and the core mixture level recovers (see Figure 7.3). This is only temporary, however, since the total break flow is still much larger than the pumped safety injection flow so the core mixture level will decrease again. This phenomenon repeats until finally at 1405 seconds the RCS has depressurized below the shutoff haad of the low head safety injection pumps and this provides enough of an addition to the pumped safety injection flow for it to exceed the total break flow. No further core uncoveries will occur which can be seen in Figure 7.3.

For this multiple break case, a maximum clad temperature of 868.60F was reached 312.5 seconds into the transient at the 10.75 foot eleva-tion. Refer to Figure 7.27 for the 4.24 inch diameter multiple aperture break cases peak clad temperature plot. The 6 inch diameter equivalent cold leg break was more limiting as can be seen in Table 3.1.

Simultaneous 8.0" Diameter Hot and Cold 1.eg Break Refer to the attached plots for Case 8, (Figure 8.1 through 8.27) for its important transient responses. As shown, the RCS pressure rapidly decreases to the pressure at the steam generator safety valve setpoint, but instead of stabilizing there it continues to drop until the accumu-lator discharge at 69.0 seconds into the transient (refer to Figure 3.8 1

5106A i l

l

8.1). Because of the size of these breaks, (equivalent to an 11 inch diameter cold leg break) the total mass discharge rate is large enough to drain the primary system very rapidly. Since the breaks are so large, the RCS pressure is below 600 psi by 70 seconds and the inirtion of subcooled accumulator water leads to the first core uncovery at that time, (see Figure 8.3). .

A large volume of water is injected into the primary system before the pressure rises enough to shut off the accumulators. This large influx of water quickly recovers the core to its hot and cold leg elevations and it maintains that level until the second accumulator injection at 213.0 seconds occurs. As seen in previous cases, this lowers the pres-sure in the downcomer which leads to a flow reversal between the core and the downcomer control volumes, causing the core to uncover again.

This time the accumulators empty, but before this can happen the RCS must depressurize below the shutoff head of the low head safety pumps which provides sufficient safety injection flow to recover the core.

The RCS pressure oscillates during this portion of the transient causing a fluctuation in safety injection flow rate but the break flow is less than SI flow at 800 seconds and no further core uncovery will occur.

For the 8" diameter multiple break case, no clad heat up abeve the steady-state operating conditions occurs. That is, the maximum clad temperature is 707.750F at the 10.0 foot elevation at 0.0 seconds of the transient. The 11" diameter equivalent cold leg break, (Case 9) had a maximum clad temperature of 715.50F at the 10.75 foot elevation at 78.5 second into the transient and is therefore more limiting. Plots of ,

the 11" cold leg break transient response can be found in the attached ff % res, see Figure 9.1 through 9.27. Table 3.1 gives the chronology of j events and peak clad temperatures, axial locations, and time for all four of the comparisons made in this secton of the report. i I

l j 3.9 5106A

Conclusions .

Multiple break analyses on a three loop Westinghouse NSSS plant were perfonned in this section for various break sizes and they were compared to their equivalent size cold leg breaks. From this comparison, it was shown that simultaneous hot and cold leg breaks of equal area are less limiting than their equivalent size cold leg break with respect to core uncovery and any other small break transient characteristic that affects .

the peak clad temperature calculation.

As mentioned above, a three loop plant was utilized for this analysis.

This plant was selected since in general three loop plants yield rela-tively high peak clad temperature results compared to other plant designs. This is true because of the relatively high power rating per loop and lower safety injection rate per loop in these plants. Since all Westinghcuse plants would experience the same transient character-istics which aake the equivalent cold leg break limiting, even with some differences in the plant design, it is concluded that the equivalent size cold leg break is limiting for all Westinghouse plants. This includes all two loop, three loop, and four loop, UHI and non-UHI Westinghouse plants.

Cold leg breaks tend to be the worst postulated break location for Westinghouse plants because of three major reasons.

1. It is necessary to clear the loop seal of water in order to vent steam generated in the core, out the break for a cold leg break analysis.

2. The cold leg break cases have the greatest amount of steam binding l because the flow path from the core to the break has a larger i resistance.

3. The cold leg break analysis assumptions also include the spillage of safety injection from the line of least resistance.

3.10 5106A j

TABLE A.1 INPUT PARAMETERS FOR SMALL BREAK ANALYSES Licensed Core Power MWt 102% of 2775 i

Nunber of Reactor Coolant Loops 3 12.2 at 10 ft Peak Linear Power kw/ft Peaking Factor (At License Rating) 2.32 Fuel Array 17 x 17 Number of Fuel Assemblies 157 Accumulator Water Volume (ft ) 1.000 Break Initiation Time (sec.) 0.0 4

1 e

s' i

4

-5106A

. Considering the above, the results reached in this report can be applied to a more general class of multiple aperture break cases for the follow-ing reasons. ,

All multiple aperture breaks studied in this report had a simultaneous hot and cold leg break and these breaks were of equal areas. Since, it is the additional vent path, (i.e. the hot leg break) for steam gerer-ated in the core, which results in a l'ess limiting peak clad tempera- ,

ture, the addition of any size hot leg break will reduce the peak clad temperature results. As this additional vent path's area is reduced, the multiple aperture small break transient will approach the equivalent cold leg break transient and the peak clad temperature will change appropriately. Therefore, any multiple aperture break area combination will be bounded by its equivalent size cold leg break case.

Ia addition, because the cold leg break analysis assumptions also include the spillage of safety injection from the line of least resis-tance, any multiple aperture break case without a cold leg break, will be less limiting than an equivalent size cold leg break. Therefore, any simultaneous hot and crossover leg break case will definitely be bounded by their equivalent size cold leg break cases.

Finally, even with a simultaneous cold and crossover leg break, some additional steam venting is realized. Depending on the location of the break in the crossover leg, the venting may not be as beneficial as a hot leg break but any additional venting will decrease the amount of steam binding and the necessity of the loop seal to clear which in turn lowers the peak clad temperature. Therefore, any simultaneous cold and crossover leg break cases will be bounded by their equivalent size cold ,

leg break also.

Therefore, the infonnation documented in this report has shown any post-ulated multiple loop break scenario will be bounded by its equivalent size cold leg break case. This will include any simultaneous hot and cold leg break, any simultaneous hot and crossover leg break, and .any simultaneous cold and crossover leg break _ cases of any break area combination.

5106A 3.11

i Finally, the conclusion in each plant's FSAR is that the Acceptance-Criteria denoted in 10CFR50.46 are met. Since a full range of small breaks are evaluated in the FSAR of every plant, this conclusion is valid even in light of multiple aperture breaks. The peak clad temper-ature, percent Zircaloy-water reaction, etc., for a multiple aperture break are bounded by the cases evaluated in the FSAR.

e r

3.12 5106A

- - - -TABLE A.2- - - -

TRANSIENT ANALYSES CONSIDERED Case 1 - 8" cold leg break; flow path No. 15 point contact Case 2 - 8" cold leg break; flow path No. 15 continuous Case 3 - 3" cold leg break with hot leg nodes; flow path No.15 continuous Case 4 - 8" cold leg break with hot leg nodes; flow path No.15 continuous Case 5 - 2.12" hot leg plus 2.12" cold leg break; flow path No. 15 continuous Case 6 - 2.83" hot leg plus 2.83" cold leg break; flow path No. 15 continuous Case 7 - 4.24" hot leg plus 4.24" cold leg break; flow path No. 15 continuous Case 8 - 8.0" hot leg plus 8.0" cold leg break; flow path No. 15 continuous Case 9 - 11" cold leg break; flow path No. 15' continuous Case B - 3" cold leg break; flow path No. 15 point contact (documented in Section 3.2 of Reference 1)

Case C - 4" cold leg break; flow path No.15 point contact (documented in Section 3.2 of Reference 1)

Case D - 6" cold leg break; flow path No. 15 point contact (documented in Section 3.2 of Reference 1) 1 I

5106A

2 TABLE A.3 PLOT FIGURE NUMBERS ,

Case No. Plot No. Plot Description 1 RCS pressure 2 Steam generator secondary pressure 3 Core mixture level .

4 Upper head mixture level 5 Downcomer mixture level 6 Pressurizer mixture level 7 Steam generator coldside mixture level 8 Steam generator secondary mixture level 9 Loop seal mixture level 10 Hot leg mass flow rate 11 Cold leg mass ficw rate 12 M ety injection mass flow rate 13 SG Jafety valve mass flow rate 14 Break mass flow rate 15 Hot leg fluid temperature 16 Cold leg fluid temperature 17 Steam generator secondary temperature 18 Hot leg flow quality 19 Cold leg' flow quality 20 Break flow quality 21 Core mixture quality 22 Downcomer mixture quality 23 Steam generator hot side mixttee quality 24 Steam generator cold side mixture quality .

25 Integral of break mass flow rate 26 Integral of break energy flow rate 27 Peak clad temperature 5106A

.BLE 1.1 Reactor SI Acc. Peak Clad Temperature Trip Signal Injection Value Lccation Time No. Case (sec) (sec) (sec) (OF ) (ft) (sec) 1 8" Cold Leg Break 10.06 8.65 133.85 901.2 10.75 4 36.1 15 point contact 2 8" Cold Leg Break 10.06 8.65 133.59 707.75 10.0 0.00 15 continuous B 3" Cold Leg Break 26.04 32.34 1315.4 1708.4 11.25 1420.8 15 point contact 3 3" Cold Leg Break 27.0 33.47 1308.5 1710.8 11.50 1380.0 15 continuous 5106A ,

.tE 2.1 Reactor SI Acc. Peak Clad Temperature Trip Signal Injection Value Location Time No. Case (sec) (sec) (sec) (cF) (ft) (sec) 2 8" Cold Leg Break 10.06- 8.65 133.59 707.75 10.0 0.00 no hot leg nodes 4 8" Cold Leg Break 10.32 9.35 132.19 727.43 11.0 89.4

.with hot leg nodes B 3" Cold Leg Break 26.04 32.34 1315.4 1708.4 11.25 1420.8 no hot leg nodes 3 3" Cold Leg Break 27.0 33.47 1308.5 1710.8 11.50 1380.0 with hot leg nodes 5106A , .

TABLE J.1 Reactor SI Acc. Peak Clad Temperature Signal Injection Value Location Time Trip No. Case (sec) (sec) (sec) (F) (ft) (sec) 36.25 1449.15 1451.9 10.75 1478.5 5 2.12" Multiple Bk 29.78 t

32.34 1315.40 1708.4 11.25 1420.8 B 3" Cold Leg Break 26.04 18,90 24.70 750.03 1470.5 11.0 802.5 6 2.83" Mutiple Bk 22.00 670.00 1541.1 11.0 682.1 C 4" Cold Leg Break 17.10 16.13 291.97 868.6 10.75 315.5 7 4.24" Multiple Bk 12.19-a 14.30 265.50 1004.3 10.75 286.1 D 6" Cold Leg Break 11.40 5106A

l TABLE 3.1 (Continued) i Reactor SI Acc. Peak Ciad Temperature Trip Signal Injection Value Location Time No. Case (sec) (sec) (sec) ( F) (ft) (sec) 8 8" Multiple Bk 9.15 7.08 68.37 707.75 10.0 0.0 9 11" Cold Leg Break 8.82 6.53 64.09 715.5 10.75 78.5 J

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REFERENCES

1. Anderson, T. M., " Report on Small Break Accidents for Westinghouse NSSS System", WCAP-9600, Valume II, Section 3.2, June 1979 (Proprietary).

2. Esposito, V. J., Kesavan, K., and Maul, B. A., "WFLASH - A FORTRAN-IV Computer Program for Simulation of Transients in a Multi-Loop PWR", WCAP-8261, Revision 1, July 1974 (Non-Proprietary), WCAP-8200, Revision 2, July 1974 (Proprietary).

3. Skwarek, R., Johnson, W., and Meyer, P., " Westinghouse Emergency Core Cooling System Small Break October 1975 Model", WCAP-8970, April 1977 (Proprietary).

4. Salvatori, R., " Westinghouse ECCS-Plant Sensitivity Studies",

WCAP-8340 (Proprietary) and WCAP-8356, July 1974 (Non-Proprietary).

_ 5. Bordelon, F. M., et. al., " Westinghouse ECCS Evaluation Model -

Supplementary Information", WCAP-8471 (Proprietary) and WCAP-8472, April 1975 (Non-Proprietary).

l 1