LIC-03-0081, Response to Request for Additional Information, Pressure - Temperature Limits Report Amendment Request; Low Temperature Over Pressure

Response to Request for Additional Information, Pressure - Temperature Limits Report Amendment Request; Low Temperature Over Pressure

ML031610723
Person / Time
Site:	Fort Calhoun
Issue date:	06/04/2003
From:	Ridenoure R Omaha Public Power District
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
LIC-03-0081
Download: ML031610723 (49)

Text

upm Omaha Public Power Distnct 444 Souitht 16thi Street fall Onmahla NE 68102-2247 June 4, 2003 LIC-03-0081 U. S. Nuclear Regulatory Commission Attn: Document Control Desk Washington, D.C. 20555

References:

1. Docket No. 50-285

2. Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)

3. Letter from NRC (A. B. Wang) to OPPD (R. T. Ridenhoure) dated May 21, 2003, Request for Additional Information Related to Fort Calhoun Station Pressure-Temperature Limit Report Submittal (TAC No. MB6468)

(NRC-03-103)

SUBJECT:

Response to Request for Additional Information, Pressure-Temperature Limits Report Amendment Request; Low Temperature Over Pressure In support of the license amendment request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (Reference 2), the Omaha Public Power District (OPPD) provides the attached response to the Nuclear Regulatory Commission's (NRC's) Request for Additional Information of Reference 3.

I declare under penalty of perjury that the forgoing is true and correct (Executed on June 4, 2003). No commitments are made to the NRC in this letter.

If you have any questions or require additional information, please contact Dr. R. L. Jaworski of the FCS Licensing staff at (402) 533-6833.

Sincerely, R. T. idenou Divisi nMa ager Nuclar Ope ations

/ RT /rj Ab0\

Employnent vitht Equal Opportunity

U. S. Nuclear Regulatory Commission LIC-03-0081 Page 2

Attachment:

Response to NRC Request for Additional Information Pressure-Temperature Limits Report (PTLR); Low Pressure Over Temperature I c: T. P. Gwynn, Acting Regional Administrator, NRC Region IV A. B. Wang, NRC Project Manager J. G. Kramer, NRC Senior Resident Inspector

LIC-03-0081 Attachment Page 1 Attachment Response to NRC Request for Additional Information Pressure-Temperature Limits Report Low Temperature Over Pressure

LIC-03-008 1 Attachment Page 2 Response to NRC Request for Additional Information Pressure-Temperature Limits Report (PTLR); Low Pressure Over Temperature NRC Question 1:

The LTOP analysis employed RELAP5/MOD3.2 which is not the latest version.

RELAP5/MOD3.3 contains improved water property data at low pressure. Why was not RELAP5/MOD3.3 used and what would have been the impact on the LTOP transients?

OPPD Response:

RELAP5/MOD 3.3 was not used to perform the low temperature overpressure protection (LTOP) analysis due to the analysis being completed prior to the release of the RELAP5/M0D 3.3 code. The impact it would have had on the analysis is described in Appendix 1.

NRC Question 2:

Did ITS Corporation perform the LTOP analysis using the same version RELAP5 as that used by OPPD? If not what were the differences and do they impact the analysis?

OPPD Response:

ITS Corporation did not run RELAP5 to perform their LTOP analysis review. They analyzed the model and performed a series of hand calculations to verify that RELAP5/MOD 3.2 was predicting correct results (Reference A). Please refer to Appendix 2 for ENERCON Services, Inc discussion of Reference A.

NRC Question 3:

Code benchmarking and validation is presented in the attachment to the October 8, 2002 submittal named NEPTUNUS. Did INEEL use the same version as that used by OPPD in the LTOP analysis? (The INEEL RELAP5/3-D version differs from the ISL version used by OPPD). Did OPPD benchmark the version obtained from ISL? Please provide the validation results justifying the use of RELAP5/MOD3.2d for the LTOP analysis.

OPPD Response:

In the report NEPTUNUS, NEEL used RELAP5/MOD 3.2 as noted in the cover page of Reference B. OPPD's benchmark of RELAP5/MOD 3.2 is described in Section 5.1.5, page 15 of Attachment 1 to Reference C and pages 158 - 163 of Reference D.

NRC Question 4.A:

NEPTUNUS simulated pressurization (and subsequent depressurization) with sprays and an initial void in the pressurizer. Many of the LTOP analyses were run for a water solid condition.

What data were used to validate the RELAP5 for water solid conditions?

OPPD Response:

The water solid transient involves only a small flow rate of water at near constant temperatures into a fixed volume. The consequence is a pressure rise until the power operated relief valve (PORV) setpoint is reached and then water flows out of the PORV after a suitable time delay.

LIC-03-0081 Attachment Page 3 The RELAP result consists of the pressure rise rate and the PORV flow rate. The pressure rise rate was verified to be acceptable by a hand-calculation as discussed below in Response 9. The PORV flow rate was verified to be reasonable by a hand calculation and discussed below in Response to NRC Question 5. These were considered sufficient validation since they are the only parameters of real interest.

NRC Question 4.B:

The NEPTUNUS pressurizer nodalization employed 12 cells while the LTOP Fort Calhoun analysis utilized 6 cells. Please provide the sensitivity study justifying the Fort Calhoun study.

OPPD Response:

The use of six nodes was based on a standard pressurizer model obtained from a sample input deck. The noding was not made finer because the transient analyzed did not require it.

Specifically, for cases with a steam bubble in place, the inrush of cold water would be expected to form thermal stratification. This is what is observed. Please refer to Section Pressurizer in Appendix 2. With this hot-water-on-top stratification, buoyancy cells will not form so there is no need for side-by-side flow nodes. The insurge of water is relatively mild so inlet plumes are not expected to be dramatic or affect the temperature of the final layer that is in contact with the steam bubble. For the water solid case, the insurge is slightly warmer due to the conservative assumption of loss of decay heat removal simultaneous with the transient. ITS in its review recommended a single pressurizer node to generate equilibrium mixing. We do see a slight temperature inversion, however, this does not impact either the pressure rise or the PORVs ability to relieve water, and therefore does not impact the peak pressure predictions.

NRC Question 4.C:

What sensitivity studies were performed for time-steps and number-of-cells, which justify the time steps and number of cells in the Fort Calhoun model?

OPPD Response:

The cell nodalization in the Fort Calhoun model (this refers to all cells, not just the pressurizer) was based primarily on the existing CESEC plant model, since this allowed the use of consistent data. Great care was taken during the model construction to avoid any unusually small or large nodes. The minimum time step used is a millionth of a second, and the maximum time step used for model development was on the order of 0.1 seconds. After completing the model, smaller maximum time steps were utilized until the results were not affected. The final runs were performed with a very small maximum time step (0.001 seconds for the period of transient activity after initial equilibrium is reached) to assure that time step choice would not affect the final results.

NRC Question 4.D:

The Massachusetts Institute of Technology (MIT) pressurization test series showed that for pressurizer insurge the peak pressure was controlled by wall heat transfer rather the water-steam interfacial heat transfer. Please show the wall nodalization justifying the OPPD modeling approach.

OPPD Response:

The Fort Calhoun model does not credit heat loss to the walls of the entire Reactor Coolant System (RCS). That is, our pressurizer model is an adiabatic model. This is discussed in more

LIC-03-0081 Attachment Page 4 detail in the Response to NRC Question 8 below. In brief, the water solid transients are mild enough that the temperature rise in the pressurizer is only a few degrees so the adiabatic assumption is conservative and small. The steam bubble cases involve a slow collapse of the bubble that also results in only a few degrees of increase in the steam region. The heat input into the RCS in general in the steam bubble cases is assured to be conservative by the assumption of loss of shutdown cooling simultaneous with a startup of a reactor coolant pump at extremely conservative RCS-secondary side temperature differential.

NRC Question 5:

The power operated relief valves (PORV) discharge coefficient was based on high pressure steam conditions. Was the coefficient also used for liquid conditions at low pressure? If so, justify the use of the discharge coefficient.

OPPD Response:

The ITS Corp report notes that the PORV at Fort Calhoun has a much greater capacity than is required to mitigate these LTOP transients. The PORV is conservatively modeled as providing zero flow until 1.5 seconds when testing shows the PORVs will be fully opened (and even then we model the PORVs as ramping open over an additional 0.5 seconds). Once the PORVs are fully opened, in all cases the flow rate is well above that required to mitigate the transient. Hence even large errors in flow rate will have no effect on peak pressure.

A PORV discharge coefficient was not used to perform the LTOP analyses. Instead the flow rate for liquid conditions is based on a constant area. The flow area of 0.94 square inches was reduced to 0.77 square inches for this analysis as described in the Section entitled "PORV Flow Rate" on page 26 of Reference D. The liquid flow rate is then generated by RELAP based on the pressure drop across a flow area of 0.77 square inches. The resulting RELAP flow rates are further discussed in response to question 14 where the flow rates are seen to be at least 2.5 times greater than the injection flow rates. In summary, conservatism in the peak pressure is assured by a conservatively slow PORV opening time and by the fact that the PORV flow capacity is much greater than required to mitigate these events.

Finally, the PORV flow rate was independently checked by the use of the American Petroleum Institute Standard 520 relief valve flow rate methodology. See Appendix 3 for the comparison calculation.

NRC Question 6:

The benchmarking is insufficient for over-pressurization events. There are relevant data from Shippingport, Connecticut Yankee, and Millstone 2. Also a series of insurge non-equilibrium experiments at Massachusetts Institute of Technology (MIT) by Griffith which covers low pressure. Please justify the adequacy of the benchmarking or show the results with the above data. Also provide a comparison of RELAP5 with data in a water solid condition. Please discuss the data in the literature and your reasons for your choice of separate effects and integral experiments.

LIC-03-008 1 Attachment Page 5 OPPD Response:

The benchmarking was performed to demonstrate accurate RELAP results for sample inputs that are provided with the code, and relevant cases as discussed in Response to NRC Question 3.

Further benchmarking is contained in Appendix 2, pages 87 and 88 to Reference D that verifies the specific model was consistent with expected flow rates and pressure drops. Discussion regarding the MIT data is provided in Response to NRC Question 8 below. Based on all the benchmark results stated previously in Reference C, and per Response to NRC Question 8 below, OPPD considers that the benchmarking is adequate and sufficient in determining RELAP5/MOD3.2's capability to determine the peak pressure following LTOP transients.

Please refer to Section 5.1.5, page 15 of Attachment 1 to Reference C for OPPDs reasoning in determining the verification and validation of using RELAP5/MOD 3.2 for performing LTOP analyses.

NRC Question 7:

Once residual heat removal (RHR) conditions are met, the reactor coolant system (RCS) can develop a bubble in the top of the vessel. Please discuss the effect of the bubble in the reactor vessel. It is anticipated that a bubble in the upper head would not affect the peak pressure but only the timing of pressure increase. Please discuss whether a bubble in the upper head impacts the results and conclusions of the analysis.

OPPD Response:

The key in determining the peak pressure is the rate of pressurization. In every scenario that opens a PORV, the analytical question is "What is the peak pressure between the time that the setpoint is reached and the PORV fully opens?", due to once the PORV opens its large capacity provides an immediate depressurization. Anything that could help the elasticity of the RCS will slow the rate of pressurization. It is noted in Section 2.3.3.1 of Reference E that it is conservative to not credit letdown, RCS volume expansion or RCS metal thermal inertia. A reactor head bubble would similarly be a non-conservative assumption since it would be an RCS volume expansion benefit. Therefore, the bubble that could develop in the upper head would act to reduce the peak pressure.

NRC Question 8:

In many of the LTOP events, collapse of the bubble in the pressurizer will occur. Please explain how the bubble collapses during the insurge prior to opening of the PORV. It appears that the nodalization in the pressurizer is too coarse so that artificial mixing of the fluid during the insurge when there is a bubble in the pressurizer will reduce the magnitude of the pressurization.

During such an insurge, the increase in the liquid is expected to compress and superheat the upper steam region. Some heat transfer between the liquid and steam region will occur initially, however, the liquid surface will saturate and a thermal layer will form insulating the steam from the lower cooler liquid region. Under these conditions, the upper steam region would not be expected to totally collapse as the RELAP5 can predict. Please discuss the above comparison with the MIT pressurization tests will show these non-equilibrium effects.

LIC-03-0081 Attachment Page 6 OPPD Response:

This NRC question was discussed further in a telephone conversation between the analysis authors and the reviewer on April 23, 2003. It was decided that the best approach would be to apply the existing Fort Calhoun LTOP pressurizer component to the same type of transient as that run by the MIT researchers. This was done and is presented in Appendix 4. The conclusions are as follows:

• The Fort Calhoun LTOP pressurizer model is adiabatic, so it in fact exaggerates the superheat effect noted in Response to NRC Question 8. At the conditions of the MIT experiment, an adiabatic assumption causes conservatively high pressure predictions, and this is verified by applying our LTOP pressurizer model to the MIT dimensions and insurge flow. Evaluation of the results show that the steam bubble, initially at 303 F, reaches 490 F if the wall heat transfer is ignored, and the pressure rise is 114 psi compared to data showing only 11 psi (this over prediction of the adiabatic approach is consistent with both the MIT paper and a related ICONE paper (Appendix 5) as discussed in Appendix 4). However, the MIT test conditions are not similar to the Fort Calhoun LTOP transient.

• The interfacial area between the liquid and steam phases is much smaller for the MIT test set up, and the insurge is much more dramatic in terms of percent volume decrease. If those two factors are corrected to match the Fort Calhoun conditions (same interfacial area to bubble height, same percentage of reduction in bubble size) the pressure rise predicted by the model is only 14 psi. That is, the adiabatic assumption becomes much less important for the conditions addressed in LTOP transients than for the MIT test conditions.

• The purpose of the LTOP model is to provide a conservatively high peak pressure. The adiabatic assumption increases the peak pressure of the model, and is therefore conservative in that measure. The adiabatic assumption does make it more likely that a PORV will open, but in most transients that does not occur even with the adiabatic assumption (the only case with a bubble that results in a PORV opening is Case 12 of Reference D, which assumes unusual initial conditions, and then the PORV opening is seen to easily handle the transients). In other words, the adiabatic case is also conservative with regards to whether or not the PORV lifts. The only direction of non-conservatism relates to whether the transient will collapse the steam bubble or not, assuming wall heat transfer would decrease the bubble size. However, the transients performed at Fort Calhoun show very little temperature rise (as witnessed by the very small change in bubble pressure). For example, the steam bubble temperature in case HP509S30 of Reference D, the worst case in terms of shrinking the bubble, sees the pressurizer steam temperature rises from 323 F to only 327 F. Thus the heat transfer to the walls would be very small if modeled. Since the transients always demonstrate large remaining bubbles of over 100 cubic feet after 600 seconds, the adiabatic assumption does not affect the conclusion that bubbles remain until the RCS and SGs equilibrate. Please refer to Appendix 4 for more details.

LIC-03-0081 Attachment Page 7 NRC Question 9:

Please explain why the events with injection from a liquid-solid condition do not result in an immediate and faster pressurization.

OPPD Response:

The pressurization rate is verified by checking against hand calculations to be reasonable. The ITS reviewer did this by the use of an Excel spreadsheet model (Please refer to Reference A for more information). In brief, the RCS is a large volume and the water injected takes as long as RELAP predicts to cause the pressure to rise to the PORV setpoint.

NRC Question 10:

Which critical flow model was used in the RELAP5 model and what is the basis for the choice?

OPPD Response:

The choice was to use the default RELAP5 model, and the basis was that the PORV capacity is so much greater than required that the specific flow model has no significance to the final result.

As noted in response to the ITS review (Appendix 2), we did perform an independent verification of the flow rate using the same pressure differentials and a methodology developed by the American Petroleum Institute (API) for flashing flow through a constant area. The API approach agreed well with the RELAP model results (Please refer to Appendix 3 for the specific comparison).

NRC Question 11:

Non-condensables collect in the pressurizer and the upper vessel head. Please describe the impact of non-condensables on the LTOP analysis. Are there any scenarios where non-condensables can affect the calculated peak pressure and the development of the LTOP limits?

OPPD Response:

As noted in Response to NRC Question 7, any increase in the elasticity of the RCS slows the pressurization rate, so the presence of non-condensables in the water solid transients is a benefit because it decrease the pressure rise between the time that the setpoint is reached and the PORV opens. Since the need is to remove volume from the RCS, and gas flow through a PORV has a much higher volumetric flow rate than water flow, any non-condensables present at the time of PORV opening would also be a benefit for the water solid transients. For steam bubble cases, the effect of non-condensables could be to increase pressure, since the non-condensable gas could not condense to a liquid phase. A confirming evaluation was performed and is attached as Appendix 6. However, the final conclusion is that the slightly higher peak pressures are still well below the PORV setpoints. An additional case in Appendix 6 verifies that if the PORV setpoint was reached due to unusual initial conditions, the non-condensable gas provide adequate volumetric relief. This additional case verified the final conclusion of Case 12 in Reference D, i.e., the PORV lifting on a steam bubble cases provides fast pressure relief and a less limiting peak pressure than the water-solid transients.

LIC-03-008 1 Attachment Page 8 NRC Question 12:

Was inadvertent actuation of emergency sprays evaluated for the cases where the pressurizer is water solid?

OPPD Response:

Fort Calhoun Station (FCS) does not have automatic initiation of emergency (i.e., auxiliary) sprays.

NRC Question 13:

What assumptions are made regarding the quench tank? Once the quench tank ruptures, would this result in higher pressurizer pressure due to the additional quench tank resistance? What is the relief area from the quench tank compared to the PORV? Please show that the analysis without a quench tank model is bounding.

OPPD Response:

The FCS LTOP model is only concerned with the peak pressure. The peak pressure occurs as soon as the PORVs open, since from that point on the relief flow rate is much greater than the injection flow rate. OPPD conservatively modeled the backpressure as the relief valve setpoint (of the quench tank) plus 5 psi even though the initial lift (with a non-pressurized quench tank) is the only lift of significance for this analysis. During an actual event, the backpressure at the initial lift would be atmospheric plus line losses. Modeling the higher backpressure assures that the initial lift with its slightly higher decay heat bounds any later reseat and re-lift.

NRC Question 14:

For each case, show the PORV mass flow rates as compared to the injection rate. Also show the void fraction in the top cell and the temperature distribution in the pressurizer for each case.

OPPD Response:

The PORV mass flow rate vs. injection rates are given for Cases 2, 7 and 8 of Reference D.

These are the only significant cases for mass flow rate vs. injection rate because Cases 1 through 6 are all at about the same PORV relief pressure, so the flow rates are about the same. In each case, the PORV flow rate is more than 2.5 times the injection rate.

(Text from Case 2)

The peak mass flow rate out of the PORV is shown in the output file to be 49.5 lbm/s. At 50°F, the density of water is about 62.4 lb1m/ft 3. Therefore the flow rate is:

49.5 lbnjsecond*l/62.4 lbnVft3 *7.481 gallons/ft 3 *60 seconds/min = 356 gpm This is well above the injection flow rate of 132 gpm. As expected, modest errors in PORV flow rate may affect the depressurization rate, but they have no impact on the peak pressure

LIC-03-0081 Attachment Page 9 (Text from Case 7)

This is the first case since Case 2 that the PORV flow rate is significantly different because this is the first case with a different driving pressure. The computer output shows the peak PORV flow rate is 76.4 lb 1 jsecond. At 255°F, the density of water is 58.7 lbm/ft3 . The volumetric flow rate is therefore:

76.4 lb,/second*1/58.7 lb,/ft3 *7.481 gallons/ft3 *60 seconds/min = 584 gpm This is well above the injection flow rate of 215 gpm at this pressure. As expected, modest errors in PORV flow rate may affect the depressurization rate, but they have no impact on the peak pressure.

(Text from Case 8)

The computer output shows the peak PORV flow rate is 90.4 lb,/second. At 305°F, the density of water is 57.1 lb /ft3. The volumetric flow rate is therefore:

90.4 lbm/second*1/57.1 lbm/ft3 *7.481 gallons/ft 3 *60 seconds/min = 710 gpm This is well above that needed to offset the injection and decay heat. As noted, the injection flow rate is limited to the 132 gpm supplied by the charging pumps since the pressure is greater than the HPSI pump shutoff.

NRC Question 15:

How were the quality assurance findings identified in the ITS Corporation letter dated September 9, 2002, addressed relative to their impact on the LTOP analysis? Please discuss each of the findings and their impact on the analysis.

OPPD Response:

Please refer to Appendix 2 of this attachment.

NRC Question 16:

Discuss the pertinence of the INEEL validation presented for the SCDAP/RELAP5 simulation of the TMI-2 accident in view of the fact that the SCDAP/RELAP5 code differs from RELAP5/MOD3.2 used in the OPPD analysis; the nodalization is very different from the OPPD model, and TMI-2 is a different design compared to the CE-designed FCS. The SCDAP/RELAP5 simulation does not validate nor justify the application of a different version RELAP5/MOD3.2d for use in simulating LTOP events in a CE-designed plant. As such the benchmarking is weak. Additional benchmarking using the same version of RELAP5 that was used in LTOP analysis by OPPD needs to be employed in the analysis. Please consider the MIT pressurization test data, as well as RELAP5 simulations of over-pressurization events in CE-designed plants. Benchmarking of the code against water solid relief is also needed.

LIC-03-0081 Attachment Page 10 OPPD Response:

The benchmarking was performed to demonstrate that the code is functioning correctly on the OPPD system and could adequately predict the peak pressure following an LTOP transient. The same version of the RELAP code (and, in fact, the same computer) was used to perform both the benchmarking and the LTOP analysis. An attempt was made to find documented benchmarking data for cases as similar as possible. The adequacy of the RELAP program to model a transient like this is discussed in Response to NRC Question 6.

It is stressed that the Fort Calhoun LTOP transients are relatively mild. The mass addition case is water added to a closed system causing that system to pressurize. ITS hand calculation shows the pressurization rate is accurate as mentioned in the Response to NRC Question 9. The water relief aspect is merely calculating a flow rate. The conservatism regarding PORV flow relief comes from a conservative delay time prior to opening, and the peak pressure as noted is insensitive to the flow rate. In any case, calculations in Appendix 1 show good agreement with the RELAP. The MIT pressurization test data is considered in Appendix 4 and hence the heat addition cases were demonstrated to be conservative. Therefore sufficient validation and verification has been performed to ensure that RELAP5/MOD 3.2 is adequate in predicting a conservative peak pressure following an LTOP event.

NRC Question 17:

Please explain why the pressure does not cycle open and close the PORV as steam is initially vented and then remain at the PORV setpoint when the discharge transitions to a pure liquid condition, stabilizing at the condition where injection into the RCS equals the PORV discharge flow. Please show the injection rates compared to PORV mass flow and the quality exiting the pressurizer for those cases.

OPPD Response:

The exact flow rate through the PORV is not critical to the analysis if the flow is sufficient to halt the pressure rise. That is, if as soon as the PORV is full open, the flow out of the PORV is greater than the RCS volumetric increase (due either to mass addition or heat expansion), then the pressure will fall. Small errors in the PORV flow rate will affect the rate of depressurization, but not the peak pressure. This analysis is only concerned with peak pressure.

For transients where the pressure is not immediately relieved at PORV lift, the RCS pressure will continue to rise until an equilibrium condition exists. Should this scenario occur, the accuracy of the PORV equation is important to the peak pressure. Conservatism is assured in this analysis by the use of a high backpressure of 90 psia. As demonstrated in the Results section of Reference D, all PORV lifts immediately reduced the transient pressure; hence modest changes in the PORV flow rates would not affect the peak pressures.

LIC-03-0081 Attachment Page 11 In summary, this model is only concerned with peak pressure. In every case, the initial PORV lift immediately relieved the pressure transient with more than adequate flow. Since the initial lifts are performed at a bounding conservative high backpressure, and subsequent lifts will occur with less decay heat input, it is not necessary to model the transient past the peak pressure point and the model does not attempt to simulate this period. The PORV mass flow rates versus injection rates are included in Response to NRC Question 14 above.

References:

A) Letter from ITS Corporation (K. Ross) to OPPD (F. James Jensen) dated September 9, 2002, "ITS Corporation's Cursory Review of OPPD's LTOP Analysis." [Note: This Reference was included in Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. I License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

B) R5-02-01, Validation Report for NEPTUNUS Pressurizer using RELAP5/MOD 3.2, dated April 12, 2002. [Note: This Reference was included in Letter from OPPD (D. J.

Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS)

Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

C) Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk), dated October 8, 2002, "Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)

D) FC06877, Rev. 0, "Low Temperature Overpressure Protection (LTOP) Analysis, Revision 1." [Note: This Reference was included in Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No.

1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

E) Letter from the NRC (S. A. Richards) to the CEOG (R. Bernier) dated March 16, 2001, "Safety Evaluation of Topical Report CE NPSD-683, Revision 6, "Development of a RCS Pressure and Temperature Limits Report (PTLR) for the Removal of P-T Limits and LTOP Requirements from the Technical Specifications" (TAC No. MA9561)

LIC-03-0081 Attachment Page 12 Appendix 1 Comparison between RELAP5/MOD 3.2 and RELAP5/MOD 3.3 An evaluation was performed to compare the results of RELAP5/MOD 3.2 and RELAP5/MOD 3.3 for a few select low temperature overpressure protection case runs (Reference D). For each overpressure event (i.e. heat and mass addition) two case runs were performed using RELAP5/MOD 3.3. No modeling changes were performed for any of the cases.

For the heat addition event (HA), Cases 11 and 12 were performed. Please refer to Figures 1 and

2. These figures demonstrate that using RELAP5/MOD 3.3, the resultant peak pressure is significantly lower and depicts the two codes having essentially the same trend. The conclusion is that both codes seem capable in determining the peak pressure following a HA event and demonstrate essentially the same trends. For both HA cases, it appears that RELAP5/MOD 3.2 conservatively predicts a higher peak pressure.

For the mass addition (MA) event, Cases 2 and 6 were performed. Please refer to Figures 3 and

4. These figures demonstrate that using RELAP5/MOD 3.3, the resultant peak pressure is essentially identical to that predicted by RELAP5/MOD3.2 and depicts the two codes predicting essentially the exact same trend. The conclusion is that both codes are capable of determining the peak pressure following a MA event and demonstrate essentially the same peak pressure and trend.

The overall conclusion of the comparison between RELAP5/MOD 3.3 and RELAP5/MOD 3.2 is that it appears the improved water property tables provide an improvement (i.e. a lower peak pressure) in determining the peak pressure following a HA event. In the four test cases that were performed, the peak pressure is higher using RELAP5/MOD 3.2 (HA events only) and both codes predict essentially the same trends. Thus it is concluded that RLAP5/MOD 3.2 provides conservative results and thus its application for the Fort Calhoun Station LTOP analysis is also valid and conservative.

Reference D: FC06877, Rev. 0, "Low Temperature Overpressure Protection (LTOP) Analysis, Revision 1." [Note: This Reference was included in Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

LIC-03-008 1 Attachment Page 13 Figure 1 Case 1i (HA Event. PORV does not lift) 460 _

RELAP5Mod 3.3 4l-:: RELAP5/Mod32 0

420

~40 380

360 0 200 400 600 Time (s)

LIC-03-0081 Attachment Page 14 Figure 2 Case 12 (HA Event, PORV opens) 480 460 440 C.

a 420

Rel

- -l  : . f >~~~c-> Rel 400 380 0 inn A i i600 n00

o- ReIap5/Mad 3.3 LIC-03-0081 Attachment Page 15 Figure 3 Case 2 (MA Event, PORIV opens) 500

>-~ Reiap5/Mod 3.3 RelapMod 3.2-400 En

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1001 - .

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0 20 40 60 lme (s):

LIC-03-008 1 Attachment Page 16 Figure 4 Case 6 (MA Event, PORV opens) 800 s S B e _-r r

Relap5IMod 3,3 :t i 600 ' Retap5IM3.2 p, a

(0 0.

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LIC-03-0081 Attachment Page 17 Appendix 2 Letter OPP1-LTR-007

-:OPPI-LTR-007 September 30, 2002 Omaha Public Power District Attn. Mr. F. James Jensen IlIl 444 South 16h Street Malt Omaha, NE 68102-2247

Subject:

ITS Review of the Low Temperature Overressure Protectton (LTOP) Analysis

Reference:

1) ITS Corporation's Cursory Review of OPPD's LTOP Analysis, ITS 01.OPPD-02-004-01 -1, 919/02

2) Fort Calhoun Low,Temperature Overpressure Protection Final Report, Revision 1, 3/15/02, ENERCON Services

Dear Mr. Jensen,

In Reference 1, ITS Corporation descibes its cursory review of the RELAP model developed by ENERCON Services and OPPD for analyzing postulated Fort Calhoun transients having potenUal to exercise the Low Temperature Overpressure Protection (LTOP) System (Reference 2). This letter responds to the key points raised in the ITS review. As noted by ITS, no modeling concems were raised to question the conclusions of the analysis.

SecUons of text taken from the referenoed document are reproduced here in:

italics.

Injectfon Water Temperature

-].The temperature of the injection water in the mass additionscenados was taken to be 250 F. This temperature Isunrealistically high tor safety injecion waterand for makeup (charging) waterunderoold shutdown' conditions.'The reasoningbehind the oiginatorsuse of elevated !niection water temperature seems questionable.

However, the LTOP reportargues convincingly that the elevated temperature s conservativefthe reasonsbein that.-

1. injectinmass ftow rates were specified assumingthat the njection waterwas cold

2. The peak reactorcoolant system (RCS) pressurepredictedby RELAP for a particularscenario was comparedtotfe allowablepressure on the P17curve associatedwith the temperature of the RCS at the beginningof the scenario (as opposed to the higherpressureon the

LIC-03-008 1 Attachment Page 18 PIT curve associatedwith the higher temperatureof te RCS at the time the peakpressureoccur-ed).

The argumentthat the use of elevated inection watertemperatureIs conseative is believable. However, a reiew recommendation is that any future RELAP LTOP calculationsbe..made with realisticnjectonwatertemperatures.

Response: There is a trade-off between density and enthalpy effects associated with the water temperature assumption. Cooler water has greater density increasing mass injection; hotter water has greater enthalpy inreasing energy injection. It is not possible to use a realistic injecion water value without numerous sensitivity analyses for each scenario to determine the worst-case value. We simplified the effort by using extreme non-mechanistic bounds - density just above freezing and enthalpy associated with 2500 F. The 2500 F value admitedly would require unusual scenarios and might only exist for a brief period, but it was based

on the maximum conceivable VCT temperature. Any other temperature would have required more jusUfication. The energy difference between 250 and 120 (if jusstifable) would be about 130 BtIbm. At 132 gpm. that amounts to about 2.4 MWs, calculated as 130 Btu/lbm'132 gallm 8 bmigalV1I0 slmi1055 Ws/Btu*1e46 MWiW. This extra energy of 2.4 MWs Is less than 10% of the 25.7 MW decay heat, so the water injection temperature conservatism Is smal compared to the conservative way we addressed decay heat.

PORV FlowResistnce A hand calculation was made to verify the fiow resistance offered by the PORV in the RELAP model This was done on account of questions that arose in the course of the review,ragardingfh adequacy of the PORV modelingfor subcooled liquid flow. In several of the mass-addition LTOP scenados, RCS temperature remains below the saturationtemperature downstream of the PORV. In these scenarios, the PORV is flowing liquid water. The hand caloulatonis included as Attachment

1. The resufts of the handcalulaionwere comparedto the results of the RELAP calulation for the base mass-addion LTOP scenaro. The RELAP calculatonhad to be extended for a comparison to be made. The specific comparson was of the:

steadystat pssure drp across the PORVgiven a cold water-solid RCS end a xed chargingflow rate. CiticalIn consideringthe mass addition scenaros involving the charging pumps is realizingthat these pumps at FCS are positive displacement pumps (as opposed to centrifuga?pumps). Such pumps develop a certain flow irrespective of head. As such, the pressure oxcursion that would be

expenenced by FCSgiven spunous opertion of al chargingpumps d one, operatingPORIk in a coldshutdown condition would be largelydiffernt (smaller) fhan what would be experienced by a plant having centrifugal chargingpumps. For 132 gpm chargingflow, the hand calculaioOand the RELAP calculaion pedict a pressure drop cm0s the PORV of 54 and57 psld, respecely. The RELAP modeling then of the flow,resistance offered by the PORV to subcoled liquid fow shows to be accurateand on fhe consenatWe side. (A review recommendation is,

LIC-03-0081 Attachment Page 19 however, tha PORV modeling be done differently In future RELAP LTOP calculatibons. The current RELAP modeling of the PORV for subsonic singlephase flow (e.g., cold liquid water flow) is not clean. (Is of the abrupt expansion model sthouldbe replaced by the Inclusion of a physical flow coefficient (Cv) table. The currentPORVmodelingIs conservative because the area of the ofice in the vae has been defined 18% smallr than physial. Were a physclrepresentatve:

oriftce area defined, the fow resistance offered to cold flowing liquid would be too lowandnon-conservatve. A physcalyrepresentative Cv value of 26.99 fore fult-open FCS PORV is cakulatedin Aftachment 1.)

Response: Other than a conservative,: bnef ramp toopen, the PORVs are modeld as onstant area. The modeling is consistent with other sites, the RELAP manual examples, and with Amercan Petroleum Industry relief valve equations. The valve area used was back calculated using trial and error to produce flows consistent with the design condition (this is what resulted in the area being 18% smaller than listed in the spec sheet). The area reduction could be considered as equivalent in, bottom line effect to determining the valve full-open Cv value and helps to explain the good agreement between the reviewers hand cate and the RELAP result.

RELAP addresses subcooled liquld as well as liquid that flashes to two phases. Of course, in general, It Is best to use manufacturers Cv values, but such data was not available. We are confident that the flow equations used are good for this purpose. We also note in the final reportthat the flow:rate at the time of PORV o pening is much greater than injection rate, and so small errors in PORV flow rate w11 have no impact on peak pressure since the PORV is more than adequate at reducing the pressure as soon as It is opened for all scenarios. (Note: the reviewer calculated different pressure drops for the same flow rate. In LTOP analysis, the PORV setpoint and the assumed downstream pressure determine the pressure drop, and the flow rate s calculated.)

ROS PressurizationRote In Mss-Addidon Scenarios A hand calculaton was made to verity the tfme taken for the RCS t pressurize to the PORVsetpoint in the base mass-additionLTOP case. This was done on.:

accountofquestons thatarose In the course of the reviewregarding the seemingly slow pressuizationrate n the RELAP cakulation of further water addition to a water-solidsystem. The calculaton is included as Attachment 2. It simply relates the charging flow rate to the volume of the RCS and t compressibility f iquid betwen te initialRCS pressure and the P4RV set point The handcaculation and theREfLAP calculation predict an elapse of 19.5 and 18.3 sec, respectively, from the time charing flows initate to the tiffme e PORV:

set point Is reached. This fair comparison satisfiedthe review questions regarding pressuzaftion rate.

Response:ENERCON also did hand calculaions to estimate pressurization rate:

and found agreement with the RELAP model.

LIC-03-0081 Attachment Page 20 ReactorCoolanf Pumps, The heating of RCS inventoty associated with irreversible flow losses in the system is accounted for in the RELAP model by depositing energy in the fluid as it flows through the reactorcoolantpumps. This Is appropriatebut there is a conservtismhere thatmayhave been overookedBydefault, RELAP dposits

the irrevesible loss associatedwith wall frcion into the fluid locally as heat.

Typically walfitin accounts forroughly,half of the fowloss in an RCS; the other halfbeing attributedto wminor-toe flow loses through ittngs, abruptexpansions and contractions, etc. Minor4p fnow losses are not deposited in t fluid as heat by RELAP. The heat additions made then to the RELAP calculatfons to account for reactor coolant pump operation are roughly 5% higherthan physical Response: Early runs without RCP heat slabs showed that the RCP heat was not being added through fiction at a conservative rate, perhaps for the reason that the reviewer notes. The heat slabs were added as an after thought to assure model conservatism. The additional frction heat conservatism was not mentoned because it was small compared to other. conservatisms, and ft was difficul to explain and quantify.

SIt was noticed that in cases wher a reactorcoolant pump was not operating, the pump component was removed from the RELAP model and a simplistic control-volume component was substtuted. It is unclearwhy this was done. A substantial effort was clearly made In the modeling the pumps as evidenced by the complete set of homologous curves defined. It would be good to take advantage of he.

thorough pump modeling giWn the reverse lop flows that develop in many of the LTOP scenaros. f the reason forremoving the pumps was robustness-related (e.g., code stops),it woild have been good to state this in the LTOP report. In any case, it would have been good to nclude a description of how ft resistance offered by a stoppedreacorcoolantpuMp. to reverse fowwas capturdinthe surrogate component.

Response: the text of Reference 2 states: Fort Calhoun RCPs have anti-reverse-rotation devices, so all the secured RCPs (which will all have reverse flow) are modeled with a loss coefficient as described in Attachment 2.. In attachment 2, the loss coefficient Is described and the reference given. This is consistent with other Fort Calhoun models.

Volume Control Rags it was noticed n the course of the revew that a handful of controlvouBmeshad the calculation of wall fritkion disabled. It is unclearwhy this was the case. ff this was Inadvertent, it wuld be good to enable 'ction in hese control volumes for.

consistency.

LIC-03-0081 Attachment Page 21 Response: wall friction decisions were described inAttachment 2 to Reference 2.

Cases where wall friction was set to zero were based on consistency:with existng plant models, such as the design basis CESEC model. For example, under..

component 330, CESEC Node 12 neglects friction losses (hydraulic diameter -

.9E99) so this model does likewise. The hydraulic diameter is set to 100 ft and the control flag Isset toO01D to Ignore mcUon losses!-

Pressurizer The pressurizer Ismoderec as a single stack of 6cont volumes. ith respec to interfaciaheat andmass transferconsiderations, it wauldbe betterto useeiter2 ormore adjacent stacks of cells or simply a single ceI to represent the pressunzer.

The reason for this is the tendency for unrealistic stratification to develop. Inan

actualpressudzer, the liquid Invenfory s well mixed by cirulative natural convecton flows. Ina single stack of contro volumes, RELAP has no way to develop such flows, Consequently, stratfed ayers of larely varying temperature can develop. RetatiVely cold layers can unrealisticallysit atop relaf;elyhot layers.

This unphyscl stratificatfon can impact the realism of the interfacialheat and mass transfercaculated by RELAP between the liqud region and vaporspace of the pressurzer.

Response: The OPPD LTOP presunzer model is consistent wlth other sites and RELAP manual examples. The "reverse stratification3 does not occur for cases

with pressurizer bubbles because the pressurizer is the hottest location in the RCS. This means coolerwater is introdued through the surgeline. If everthe surgeline flow were hoter than the saturated water at the top, it would boil. The output Contained Infile hp5O3s30.o, for example, ends up Inthe final edit statement with pressurizer temperature from bottom to top of 2680 F, 314 0 F, 350° F, 3950 Fs 423° F, and 449° F. However, for the mass addition cases that are originally water solid with constant temperature RCS, slightly warmer fluid Is introduced as the decay heat warms up the fluid. The temperatures Inthe final edit, statement In m305p2.o are, from bott to top, 310A°, 310.10. 307.4, 306501 306.4 ° and 306.3 °F. This has no impact on peak pressure because the PORV flow rate easily causes the pressure to fall, that Is,there is no consequene to ai

.small change in PORV flow due to a few degrees difference. Note: use of a single cell as'recommended by ITS would be less accurate for pressunzer bubble cases,:

since it would eliminate the temperature stratification that we expect to be there.

The pressurizerInventoy in e heat addition LTOP scenarios was appropriftely Inifialized saturated.in the mass addton cases, however, the pressunzer inventory was initiazedat t initialtemperatureof the RCS. This seems questionablegiven that 1)before the spurious injection, the pressurzer Inventory twouldhave been saturatedatthe Initialpressure of the RCS, andthat 2) the pressurizerheaters are assumed to be operating as epressurizer fis withliquid.

It might be more defendable to startmass-addiionscenados With a realistic pressurizercondition (.e., saturatedwith levelin the nominalrange)and then

LIC-03-0081 Attachment Page 22 allow the pressurizerto fill with the heaters operating. It could be that the pressure drop across the PORV will differ meaningfully dependent upon the temperature of the liquidin the pressuzer. (This might especially be true if the liquid temperature were greaterthan the saturaiontemperature downstam of the PORV.)

Response: Inthe mass addition cases, there is no pressurizer bubble so there is no reason for te pressurizer to be at saturated temperature. Reference 2 does include one case (Case 4) with saturated pressurizer and bubble to see Ifthe bubble made a difference, and determined that It had no signfficant impact since the bubble collapsed pnor to PORV lift.:As to PORV flow rate differences, as noted

in Reference 2, the PORVfloW rate is much higherthan is needed to start a pressure decrease, so changes inflow due to different inlet temperatures do not affectpeak pressure.

Stedy State A review recommendation Is that in future LTOP analyses documentation, results be presented of an extended steacstate RELAP calculation. Te objCt of Including the steady-state results would be to dentifvyclose correspondence between the RELAP LTOP model and actual FCS monitoredparameters. The calculation should have reactorpowerat t full operating value, andshould indude realistic feedwater temperature and acte steam generator level contro.

The goal here would be to convincingly illustrate the base realism of the RELAP model.

Response: We did compare reatistic steady-state results to other models In, Reference 2,Table A2-3, as recommended.

Steam Generators The secondatyside oftthe steam 9eneratorsand the steam generatortubingmetal gmass were conservatvly excluded from the mass-additonscenarios. In the heat addition scenarios, the generatorswere initiaized entrely full of liquid which 'was hot relative to RCS temperature. ntiaingthe steam generatorsfullofliquid.

seems unrealisticaly conservaUve. A suggesion of the rview s that future heat addition LTOP calculations be Initialized with steam generator level In t nominal range consistent with where the operatorswould maintain it Response: It s agreed that the Reference 2 steam generator treatment is clearly conservative, but it allowed fora simpler boundary of the model. If we had used a realistic amount of water, we would also have to nclude heat transfer from the outer steam generator walls. We would also have had to model steam condensing as the seondary side cooled, including models of the'metal ms and surfacs in the steam region. Note: the model does Include the steam generatortubing metal mass in the heat addton scenarios.

LIC-03-0081 Attachment Page 23 Summary In summary, the model shortcomings identified in the course of the review are not thought tb have the potential meaningfully impact he conclusions of OPPD's current RELAP LTOP calculations.The RELAP model seems well suited to performing LTOP transients and is vey well documented. Modeling uncertinties appear to have been consistenly addressed in a conservative manner.

Response: While we appreciate the conclusion,-we do not agree that any model shortcomings exist. Where the reviewer identifies tnsersrs, we believe the conservatisms ate smatl and justfied for their simplificaUon of the model. Since better PORV flow data is not available, and pressurzer reverse stratification does not affect the peak pressure calculation, we do not see any advisable changes to the model.

Questons on this response can be addressed to Ralph Berger or mysef at 510-,

632-1734.

Sincerely, Tien Lee Engineeing Manager Enercon Services, Inc.

TPLTjtn

LIC-03-008 1 Attachment Page 24 Appendix 3 Verification of PORV Flow Rates Independent Check Of RELAP PORV Water Flow Rates Based On Inlet Conditions, Exit Pressure And 0.77 Square Inches Relief Area Three cases from Reference D were checked: Case 2, Case 7 and Case 8. The methodology used is taken from the American Petroleum Industry (API) Standard 520. It is important to note that this methodology requires a relief valve coefficient kd. API recommends, in the absence of other data, to assume kd = 0.85 (the sensitivity is such that higher values of kd result in higher flow rates, since flow is proportional to kd). Trial and error found that a value of kd of 0.62 provided a close match, which means that the RELAP water flow rate is lower by about 15% relative to the default API approach for a generic relief valve. In the below cases, the value of the three coefficient product kdkbkc is set to 0.62, however evaluation of the methodology identifies that kb and kC should be 1.0 under these conditions so this term really represents just kd..

Results:

Case Upstream P Upstream T Downstream RELAP API Flow Rate (psia) (OF) P (psia) Flow Rate kd=0.62 (lbn/s) (Ibm/s) 2 484 50 90 49.5 50.1 7 1071 255 90 76.4 76.6 8 1522 305 90 90.4 91.3 CASE 2 API SOLUTION CALCULATION OF TWO-PHASE FLOW RATE This calculation is based on the specification of an inlet state, an outlet pressure, and a relief path, including area and loss terms. The flow rate through this path is calculated based on the following references:

• The American Petroleum Institute (API) Recommended Practice 520, Sizing Selection and Installation of Pressure Relieving Devices in Refineries, Appendix D, 7th edition, January 2000.

• The Crane Manual, also known as Flow of Fluids through Valves, Fittings, and Pipe, Technical Paper No. 410, The Crane Company, Twentieth printing, 1981.

• Easily Size Relief Devices and Piping for Two-Phase Flow, Joseph Leung, Chemical Engineering Progress, December 1996.

These references are referred to below as the API, Crane, and Leung, respectively.

LIC-03-0081 Attachment Page 25 INLET STATE The water inlet state is subcooled water. The pressure is 484.0 psia and temperature is 50.0°F.

There is no non-condensable gas present.

Water/steam state properties are as follows, where f indicates fluid, g indicates gas, and o indicates inlet state. Water/steam saturation properties are given for a temperature of 50.0°F and pressure of 0.18 psia.

Enthalpies: ho=19.43, hfo =18.05, hgo =1083.40 Btu/ lbm Specific Vols: vo=0.01586, vfo =0.01602, vgo =1704.80000 ft3/ lbm Densities: po=63.0673, pfo =62.41, pgo =5.8657E-4 lb,,,/ft3 Entropies: so=0.04729, sfo =0.03610, sgo =2.12620 Btu/ lbm.F Spec. heats: cpo=1.002, cpfo =1.002, cpgo =0.444 Btu/ lbm0 F Based on the backpressure of P2 = 90.0 psia, flashing will not occur.

CALCULATION Step 3: Calculation and Final Result The specific heat ratio cp/cv is calculated using Figure A-9 from Crane. A value of 1.2751 is interpolated based on a temperature of 50.0°F and pressure of 484.0 psia.

This is subcooled water with no non-condensables present and Po<1604 psia and To<634.5 0F.

The appropriate formula for the omega factor is D.8 from the API:

omega=0. 185/vo*cpfo*(To+460)*Ps*(vgo_vfo) 2 /(hgo-hfo) 2 Ps is the saturation pressure associated with To. With To = 50.0°F, the saturation pressure is 0.2 psia. Here the specific volume and enthalpy changes are evaluated at Ps and are: vgo=1704.800 ft3/ lbm, vfo=0.01602 ft3/ lbm, hgo=1 083.4 Btu/ lbm, and hfo=18.05 Btu/ lbm. Putting this into the equation for omega gives an omega value of 2717.02.

To determine whether this is a low subcooling or high subcooling region, we calculate the parameter nst = 2*omega/(1+2*omega) to be 0.9998. Since the saturation pressure at a To of 50.0F, calculated to be 0.178 psia, is less than nst*Po = 483.911 psia, this is a high subcooling region. Since the saturation pressure is less than the downstream pressure P2, which is 90.0 psia, critical flow is not achieved. The mass flux G is given by the API Equation D. 11:

G=96.3*SQRT[(Po-P2)/vfo]

G is calculated to be 15100.888 lbm/ft2 .

The value of 0.62 was provided for KdKbKc.

LIC-03-0081 Attachment Page 26 The flow rate through an area of 0.7700 square inches is given by the formula W =

kdKbKc*A*G/0.04 and is 180229.10 lbm/hr or 50.0636 Ibm/s. This is equivalent at a specific volume of 0.0159 ft3 / lbm to 47.629 cubic feet/minute, or 371.981 gpm.

The exit state is subcooled water at 90.0 psia and 49.02°F. Properties are:

Enthalpy: h2=18.45 Btu/ bm Specific Vol: v2=0.01585 ft3 / lbm Density: p2=63.0868 lbm/ft 3 Entropy: s2=0.04568 Btu/ lbm0F Spec. heat: cp2=1.002 Btu/ lbm0 F CASE 7 API SOLUTION CALCULATION OF TWO-PHASE FLOW RATE This calculation is based on the specification of an inlet state, an outlet pressure, and a relief path, including area and loss terms. The flow rate through this path is calculated based on the following references:

. The American Petroleum Institute (API) Recommended Practice 520, Sizing Selection and Installation of Pressure Relieving Devices in Refineries, Appendix D, 7th edition, January 2000.

• The Crane Manual, also known as Flow of Fluids through Valves, Fittings, and Pipe, Technical Paper No. 410, The Crane Company, Twentieth printing, 1981.

• Easily Size Relief Devices and Piping for Two-Phase Flow, Joseph Leung, Chemical Engineering Progress, December 1996.

These references are referred to below as the API, Crane, and Leung, respectively.

INLET STATE The water inlet state is subcooled water. The pressure is 1071.0 psia and temperature is 255.0°F.

There is no non-condensable gas present.

Water/steam state properties are as follows, where f indicates fluid, g indicates gas, and o indicates inlet state. Water/steam saturation properties are given for a temperature of 255.0°F and pressure of 32.53 psia.

Enthalpies: ho=225.97, hfo =223.67, hgo =1165.75 Btu/ lbm Specific Vols: vo=0.01702, vfo =0.01705, vgo =12.74300 ft3 / lbm Densities: po=58.7698, pfo =58.66, pgo =0.0785 lbm/ft3 Entropies: so=0.37099, sfo =0.37485, sgo =1.69305 Btu/ lbmnF Spec. heats: cpo=1.014, cpfo =1.014, cpgo =0.511 Btu/ lbmnF Based on the backpressure of P2 = 90.0 psia, flashing will not occur.

LIC-03-0081 Attachment Page 27 CALCULATION Step 3: Calculation and Final Result The specific heat ratio cp/cv is calculated using Figure A-9 from Crane. A value of 1.2564 is interpolated based on a temperature of 255.0°F and pressure of 1071.0 psia.

This is subcooled water with no non-condensables present and Po<1604 psia and To<634.5°F.

The appropriate formula for the omega factor is D.8 from the API:

omega=0.1 85/vo*cpfo*(To+460)*Ps* (vgo-vfo) 2 /(hgo-hfo) 2 Ps is the saturation pressure associated with To. With To = 255.0°F, the saturation pressure is 32.5 psia. Here the specific volume and enthalpy changes are evaluated at Ps and are:

vgo=12.743 ft3 /lbm, vfo=0.01705 ft3 /lbm, hgo=1165.8 Btu/lbm, and hfo=223.67 Btu/lbm. Putting this into the equation for omega gives an omega value of 46.78.

To determine whether this is a low subcooling or high subcooling region, we calculate the parameter nst = 2*omega/(1+2*omega) to be 0.9894. Since the saturation pressure at a To of 255.OF, calculated to be 32.532 psia, is less than nst*Po = 1059.674 psia, this is a high subcooling region. Since the saturation pressure is less than the downstream pressure P2, which is 90.0 psia, critical flow is not achieved. The mass flux G is given by the API Equation D. 11:

G=96.3*SQRT[(Po-P2)/vfo]

G is calculated to be 23101.321 lbd/ft2 .

The value of 0.62 was provided for KdKbKc.

The flow rate through an area of 0.7700 square inches is given by the formula W =

kdKbKc*A*G/0.04 and is 275714.26 lbrjhr or 76.5873 lbm/.s This is equivalent at a specific volume of 0.0170 ft3/ lbmto 78.190 cubic feet/minute, or 610.667 gpm.

The exit state is subcooled water at 90.0 psia and 255.56°F. Properties are:

Enthalpy: h2=226.54 Btu/lbm Specific Vol: v2=0.01702 ft 3 /lbm Density: p2=58.7543 lbm/ft 3 Entropy: s2=0.37179 Btu/lbm0 F Spec. heat: cp2=1.014 Btu/ lbmnF CASE 8 API SOLUTION CALCULATION OF TWO-PHASE FLOW RATE This calculation is based on the specification of an inlet state, an outlet pressure, and a relief path, including area and loss terms. The flow rate through this path is calculated based on the following references:

LIC-03-0081 Attachment Page 28

• The American Petroleum Institute (API) Recommended Practice 520, Sizing Selection and Installation of Pressure Relieving Devices in Refineries, Appendix D, 7th edition, January 2000.

• The Crane Manual, also known as Flow of Fluids through Valves, Fittings, and Pipe, Technical Paper No. 410, The Crane Company, Twentieth printing, 1981.

• Easily Size Relief Devices and Piping for Two-Phase Flow, Joseph Leung, Chemical Engineering Progress, December 1996.

These references are referred to below as the API, Crane, and Leung, respectively.

INLET STATE The water inlet state is subcooled water. The pressure is 1522.0 psia and temperature is 305.0°F.

There is no non-condensable gas present.

Water/steam state properties are as follows, where f indicates fluid, g indicates gas, and o indicates inlet state. Water/steam saturation properties are given for a temperature of 305.0°F and pressure of 72.19 psia.

Enthalpies: ho=277.54, hfo =274.85, hgo =1181.15 Btu/ bm Specific Vols: vo=0.01740, vfo =0.01750, vgo =6.02910 ft3 / lbm Densities: po=57.4676, pfo =57.14, pgo =0.1659 lbm/ft3 Entropies: so=0.44080, sfo =0.44395, sgo =1.62910 Btu/ lbm0 F Spec. heats: cpo=1.028, cpfo =1.028, cpgo =0.556 Btu/ lbm0 F Based on the backpressure of P2 = 90.0 psia, flashing will not occur.

CALCULATION Step 3: Calculation and Final Result The specific heat ratio cp/cv is calculated using Figure A-9 from Crane. A value of 1.2529 is interpolated based on a temperature of 305.0°F and pressure of 1522.0 psia.

This is subcooled water with no non-condensables present and Po<1604 psia and To<634.5 0 F.

The appropriate formula for the omega factor is D.8 from the API:

omega=0. 185/vo*cpfo*(To+460)*Ps*(vgo_vfo) 2 /(hgo-hfo) 2 Ps is the saturation pressure associated with To. With To = 305.0°F, the saturation pressure is 72.2 psia. Here the specific volume and enthalpy changes are evaluated at Ps and are: vgo=6.029 ft3/lbm, vfo=0.01750 ft 3 /lbm, hgo=1181.2 Btu/lbm, and hfo=274.85 Btu/lbm. Putting this into the equation for omega gives an omega value of 26.55.

LIC-03-0081 Attachment Page 29 To determine whether this is a low subcooling or high subcooling region, we calculate the parameter nst = 2*omega/(1+2*omega) to be 0.9815. Since the saturation pressure at a To of 305.0°F, calculated to be 72.185 psia, is less than nst*Po = 1493.866 psia, this is a high subcooling region. Since the saturation pressure is less than the downstream pressure P2, which is 90.0 psia, critical flow is not achieved. The mass flux G is given by the API Equation D. 11:

G=96.3*SQRT[(Po-P2)/vfo]

G is calculated to be 27547.283 lbm/ft2 .

The value of 0.62 was provided for KdKbKc.

The flow rate through an area of 0.7700 square inches is given by the formula W =

kdKbKc*A*G/0.04 and is 328776.82 lbm/r or 91.3269 lbnJs. This is equivalent at a specific volume of 0.0174 ft 3 /lbm to 95.351 cubic feet/minute, or 744.694 gpm.

The exit state is subcooled water at 90.0 psia and 307.12°F. Properties are:

Enthalpy: h2=279.75 Btu/lbm Specific Vol: v2=0.01743 f/lbm Density: p2=57.3869 lbJ/ft3 Entropy: s2=0.44353 Btu/ lbm.F Spec. heat: cp2=1.029 Btu/ lbm0 F Reference D: FC06877, Rev. 0, "Low Temperature Overpressure Protection (LTOP) Analysis, Revision 1." [Note: This Reference was included in Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

LIC-03-0081 Attachment Page 30 Appendix 4 Evaluation of the Massachusetts Institute of Technology (MIT) Test Results A concern was raised in the review of Reference 4-3 (below) that the code used, RELAP5/

MOD3.2, had been shown to inaccurately predict the pressure behavior discovered in Reference 4-1. Two reasons for this difference are apparent at first view.

The first difference is that the MIT experiment (Reference 4-1) is more dramatic than the Fort Calhoun LTOP transients analyzed in Reference 4-3. The level in the pressurizer model rises from 17 inches to 34 inches in 31 seconds (Experiment BB4) which compresses the steam volume vertical distance from an initial 28 inches to 11 inches (from 100% volume to 36%

volume) in 31 seconds. By comparison, the Reference 4-3 transient Case 10 is from 350 ft3 to 200 ft3 in 100 seconds (100% to 57%), or roughly six times slower.

The second difference is that the test setup had a very small interface surface area compared to volume (8 inch diameter versus 28 inch height, whereas the Fort Calhoun steam bubbles had a diameter = 6.86 feet and a height 9.75 feet).

The Reference 4-3 model was adiabatic, in that no wall heat transfer was assumed in the pressurizer (this was identified as a known conservatism). Reference 4-2, Figure 1 implies that adiabatic models greatly over predict pressure rise when trying to model a small, skinny tank with high rate of bubble compression. The MIT paper also concludes adiabatic models will greatly over predict pressure when modeling this transient setup.

The experiments in the MIT tests that are applicable to Fort Calhoun's LTOP analyses are the insurge to partially filled tanks (cases ST4, BB4, and TR8). This can be simulated with the FCS pressurizer model assuming the following changes:

Existing Model New Model Vol = 900 ft3 Vol = pi*(8/24) 2 *45/12 = 1.309 ft3 Length = 24.364 ft Length = 45/12 = 3.75 ft Area = 36.94 ft2 Area = 0.394 ft2 Input for Experiment BB4 Initial P = 70.1 psia Initial T = 303°F Initial water level = 17 inch Insurge T = 70°F Insurge flow rate = level change

• area/time = 16/12*0.394/31 = 0.017 ft3 /s = 7.6 gpm = 1.06 lbr/s at a density of 62.3 lb1 Jft3 Insurge flow time = 31 seconds Data for pressure history: Figure A. 1.1 pg 67

LIC-03-0081 Attachment Page 31 The Fort Calhoun pressurizer model from Reference 4-3 is as follows:

• PRESSURIZER 4100000 pres pipe 4100001 6 4100101 36.94,6 4100201 36.94,5 4100301 2.9232, 5 4100302 9.744, 6 4100601 90. 6 4100801 0.00015 0.0 6 4100901 0.0 0.0 5 4101001 0,6 4101101 0,5

• Manually set Przr pressure and water level 4101201 2 95. 0.0 0. 00,5 4101202 2 95. 1.0 0. 00,6 4101300 1 4101301 0.0 0.0 0.0,5 Experiment BB4 The Fort Calhoun Pressurizer model is revised to match the experiment case, for an initial water level of 17 inch (1.42 ft) with 5 volumes of height 1.42/5=0.283 and the bubble volume of height (45-17)/12 = 2.333 ft, and initial temperature of 70.1°F.

• PRESSURIZER 4100000 pres pipe 4100001 6 4100101 0.394,6 4100201 0.394,5 4100301 0.283, 5 4100302 2.333, 6 4100601 90. 6 4100801 0.00015 0.0 6 4100901 0.0 0.0 5 4101001 0,6 4101101 0,5

• Manually set Przr pressure and water level 4101201 2 70.1 0.0 0. 00,5 4101202 2 70.1 1.0 0. 00,6 4101300 1 4101301 0.0 0.0 0.0,5

LIC-03-0081 Attachment Page 32 The boundary condition is setup with a forced flow rate 1.06 lbm/s for 31 seconds. This is simulated with two components, a 70°F reservoir and a time dependent junction. The complete input file is as follows:

=MIT Experiment Model 100 new transnt 102 british british 105

• time step control 201 90.0 1.0-60.0015 100250 1000

• Output control 301 p 410060000

• trip cards

• 501 run stop time

• 502 insurge start time 501 time 0 ge null 0 60.0 1 502 time 0 It null 0 31.0 n 600 501

• Source of insurge to set P & T 2510000 si2 tmdpvol

• flow area, length, volume, horiz angle, vert angle 2510101 50.0 10.0 0.0 0.0 0.0

• elev change, roughness, hydraulic diameter, flags 2510102 0.0 0.0 0.0 00

• 3 makes 201 card P&T, trip number 2510200 3 2510201 0.0 90.0 70.0 2510202 1000.0 90.0 70.0

• Insurge flow rate 2520000 insurge tmdpjun

• from, to, area 2520101 251000000 410000000 0.1

• 1 means mass flows/0=velocity, trip number, table var & location 2520200 1 502 p 410010000

• pressure, liq bm/s, vapor velocity, interface (=0) 1 Pump Curve 2520201 0.0 0.0 0.0 0.0 2520202 10.0 1.06 0.0 0.0 2520203 2000.0 1.06 0.0 0.0

LIC-03-0081 Attachment Page 33

• PRESSURIZER 4100000 pres pipe 4100001 6 4100101 0.394,6 4100201 0.394,5 4100301 0.283, 5 4100302 2.333, 6 4100601 90. 6 4100801 0.00015 0.0 6 4100901 0.0 0.0 5 4101001 0,6 4101101 0,5

• Manually set Przr pressure and water level 4101201 2 70.1 0.0 0. 00,5 4101202 2 70.1 1.0 0. 00,6 4101300 1 4101301 0.0 0.0 0.0,5

• end of cases What happens is similar to what is shown on page 62 of Reference 4-1, or in Figure 1 of Reference 4-2. The FCS adiabatic model shows a much higher pressure rise. The peak pressure calculated was 184 psia, and peak temperature was 490°F.

Experiment BB4 simulated with RELAP adiabatic model L

200 00)

U, 2Q- 50 a-0 20 40 60 80 Time (s)

However, this is not similar to the Fort Calhoun LTOP transient. As noted, the Reference 4-3 flow rate is much slower and the relative surface area to volume is much greater. To do a better comparison, one should use a consistent area to height ratio and a consistent level rise rate. To

LIC-03-0081 Attachment Page 34 get a consistent area to height ratio for the test bubble height of 2.333 ft, the interfacial area should be:

=ht

• LTOP Area/LTOP bubble height = 2.33

• 36.94/9.75 = 8.83 ft2 instead of 0.394 ft2 The insurge rate to get a decrease in volume from 100% to 57% in 100 seconds would be:

= volume/time = 8.83*.43*2.33/100 = 0.0885 ft 3/s = 5.51 bm/s Making these two changes by adding the following lines to the input file:

• Changes for a better comparison 4100101 8.83,6 4100201 8.83,5 2520202 10.0 5.51 0.0 0.0 2520203 2000.0 5.51 0.0 0.0 Gives the following pressure trace Experiment BB4 simulated with RELAP adiabatic model Similar Area and Flow Rate to R. Calhoun LTOP CD, 85-CT 1 ~ -1 0 ~ 2 804 6 I.-

a-0 20 40 60 80 Timne (s)

This is the pressure predicted by RELAP for a similar bubble shrinkage rate (as occurs at Fort Calhoun during a heat addition case) for the MIT test case if the MIT test case had a surface area in proportion to the bubble height. Here the effects of the rapid shrinkage is reduced and the heat transfer to the liquid phase is increased. The pressure rise is still probably higher than actual (MIT's case BB4 had a peak pressure of 81 psia) but the effect of neglecting heat transfer to the pressurizer wall are obviously much reduced.

Conclusions The RELAP model used at Fort Calhoun is an adiabatic model. References 4-1, 4-2 and our simulation agree that the adiabatic model predicts extremely high pressures for rapid insurge, small area tanks. The reason is that the steam space gets extremely hot due to compression and

LIC-03-0081 Attachment Page 35 the water/steam surface area is insufficient to remove the heat. Our simulation of the test BB4 predicted a steam temperature of 490°F, while the test measured temperature was only about 310°F. Obviously, in the test case, heat has been transferred from the steam to the pressurizer vessel.

There are two things to note for the LTOP transient. The first is that the adiabatic assumption has a much less effect for our slower transients and much larger interfacial area. Had the MIT experiment been performed with a surface area to height ratio and bubble compression rate similar to the Fort Calhoun transient, the heat transfer to tank walls would have been a much less significant factor. The adiabatic pressure rise predicted by RELAP for this proposed test is only 14 psi.

The second thing to note is that the Fort Calhoun pressurizer model is conservative in terms of predicting peak pressure. Crediting the heat transfer to the pressurizer walls would provide a mechanism for removing energy from the primary system and keeping the pressure lower. It is true that additional heat transfer to the pressurizer walls might collapse the steam bubble faster; however, this is not a significant effect because even with the adiabatic assumption the steam temperature does not rise very much (in Case 10 of Reference 4-3, the temperature rises only 40 F), so the heat transfer would be small. In any case, the remaining bubbles (for all but one case where the PORV lifted) stay well above 100 ft3 .

Overall, the effects noted in the MIT tests (i.e. the adiabatic assumption) are insignificant for a real pressurizer LTOP event geometry and insurge rates, so it is concluded that the adiabatic assumption is both of minor consequences and conservative in terms of calculating peak pressure.

References:

4-1.1. Insurge Pressure Response and Heat Transfer for PWR Pressurizer, Hamid Reza Saedi, Masters Thesis MIT, 11/82.

4-1.2. Prediction of MIT Pressurizer Data using RELAP5 and TRAC-M, Shumway, Bolander, and Aktas, ICONE-10 paper 22580, 10th International Conference on Nuclear Engineering, 4/14-18/02, Alexandria, VA (Located in Appendix 5).

4-1.3. FC06877, Rev. 0, "Low Temperature Overpressure Protection (LTOP) Analysis, Revision ."[Note: This Reference was included in LIC-02-0109.]

LIC-03-0081 Attachment Page 36 Appendix 5 Prediction of MIT Pressurizer Data using RELAP5 and TRAC-M

LIC-03-0081 Attachment Page 37 Proceeding of ICONE10 l o4 Internatlonal Cotference on Nuclear Engineering AprU 14-1 ,2002, Alexandria, VA ICONE-10 22580 DRAFT Prediction of MIIT Pressurizer Dati using RELAP5 snd TRAC-M Rex Sbumway, MiarkBolander and Birol Aktas lnformadon Systems Libortory, Inc.

Suit 260, 2235 E,25th Street Idaho Falls, Idaho, 83404 208-552-2000, 208-552-6I7, rshunway@islinc.com ABSTRACT INTRODUCTION Tests simulating Pressurized Water Reactorpressunzers under Experimental and analytical work on presurizcd water inflow and outflow Wnditions have been perfrme atlMIT. rcactor prsurizcs was pcrfomoed itt MIT (Saedi and Griffith, Prediction of preurizerpreSSure requires accumte models of 1983). This pape contratces on Test number ST4 by wall heat tansfer as well as interfacil liquid-steam liat comparing the TRAC-M (Aktas and Uhle, 2000) and RELAPS transfer. The US NRC has two computer programs used for (ISL Inc., 2001a) computer codes to the data' TRAC-M began predicting dmTnal hydraulic behavior in reactors; RELAPS and as a modernized version of TRAC-P (Spore eta!., 1993). The TRAC-M. TRAC-M is the Consolidated Thcrmal-hydraulics RELAP5 and TRAC-M thermal-hydraulic codes arc uscd by the Code developcd by combining odels from TRAC-B and US NRC to aid reactr safcty decision making.

RELAPS into a modernizedversion ofTRAC-P. Thc component Important phenomena includi: wall condensation. mixing of models from RELAPS and TRAC-B have been ported to incoming cold water with already present hot watcr inthe vcssel TRAC-M but not the constirtivc models. A suite of asscssmcnt titsndfrec rfchcattiansfcr.

eases arc being developed to guide the corstitutive model improvement process. Assessment ayainst data will determinet Predictions from RELAP5A1OD3.3 version al and TRAC-M Which constitutive relations need to be potted to TRAC-M. This version 3927 are shown. Code options exanined includc:e paper compares the RELAP5 and TRAC.M codes against MIT thcrmal front tracking. kvel tracking, numerical inplicitness prcssurizer daia. As watcr is injected into the bottom of he. ad timestcp si scnsitivity -

pressurizer the steam pressure in the top orthe pressurizer rises.

The pressure increasc ratc is controlled by wall and interfacc TESTDESCRWIPON condcnsation ratcs. Both codes predict this o6mpkx Test ST4 was an insurge experiment Water, submoced by compression process reasonably welL The effect of time step 130 K.was ityecteO into a stinless stecl vessel which wQS sizc and code options ar explored in this paper. The benerits of. partially filled with saturated water at a prcssurc of 0.49 MPa.

using two codes to analyze tbrmal-hydraulic prooesics are :The steam in the uppcr part of the vessel was compressed. As eVident from the rsults. the saturation tmperature rose, the vesscl walls became subcooled and film condensation eccurred. The condens4te ran NOMENCLATURE dowi the walls to ni ct a rising water levcl. A balance between h; ebattransfcrcocfficient interfacial and wall steam cdcstion and steam comprcssion k liquid themal conductiviy determined the pfessur response.

S film thickness The test vessel was 114.3 cm high had anl.D. of 203 cm, and a wall thkness of 0.818 cm.

Water injection into the bottom of the vessel varied over the I Copyright C 1999 by ASME

LIC-03-008 1 Attachment Page 38 first 40.6 seconids at which time it was stopped. The injection after the water insuge was stopped.

rate translated into a vesscl water level rise rate of about I is, Wall and Interface The vessel was insulated t diminish energy losses.

Calibration tests werc uscd to estimate the loss" at 1I1 kUW During the compression process, wall condesation is the (Kim and Griffith, 1987). controlling phenomea Figu I shows thc pressure rise rate when the vcsl wall heat transfer is removed from te RELAP5 CODE MODEL modeL This demonstrates wall heat transfcr is y important.

The cssel was molled using 10 fluid celIs. A more Both codes usc the flmwise condensation coefficient accuratc prcdiction could bc obtaied with mnore cells, howe-mer correlation developed byNusselt (1916).

models of reactor pressurizers usually have les than 10 cells. RELAPS predicts a liquidto interfacc heat transfer cocfficient The water level was initially in cell 4 (he void fration was timcs area valuc of about 3000. When the liquid and var 0.22) and reached its maximum valuc in cell $ (thc void fraction, interf4cisl beat tmnsfcr coefficieits wcrc set to 1.0 internally uws 0.69). the effect on the peak preure uws negligibli. This implies interfacial hcattransfcr is not important.

The experimnters did not report on the type and thickness of the insulation eovering the vcssel. The code model uscd 8.9 cm A study of the rcason for the pressure drop rate di(fereces of fiber glass insulation. Steady itte calculations were betwccn the two codcs, after thc water insuTgc stopped, showed performed to adjust the insulation conductivity so that the model deficicncies in both codes in the cell with the water levcl.

steady state heat loss agreed with the reported value. TRAC-M switchcs from wAll condensation heat trfr node to liquid convection mode when a water level enters a cell.

CODE RESULTS Shutting ofF wait condensaion when Ihe water level tcached ccil 8 caused the noticcable prssure incrcase change at 36 Measured pressure inthe top of the vessel peaked at about secords in Fig. I.

0.59 MPa as shown inFig 1.After the subcoted water insurge stops, thc pressure falls due to further steam condcnsation. The When a water level entes a RELAP5 cell, the code partitions complcx physical processe occuirng ar: wall heat transfer, the wall energy transfer into the regions above and below the Steam-water interfacial heat transfer, and thiermal mixinig water level, However, the cordensation film thickness is bsedl between te cold and hot watcr. tupon he avcrage liquid flow across both the cell inlet and outlet 0.6 cJtions. When watcr enters from below, only the flow from above should be used to determine the film thicknes used in the N1usselt (1916) condcation halt transfercoefiint:

h* (I)

D55 Using the average liquid how rate results in a large tickness and a small heat transfcr coefficimt when war is floing into the cell from below. When the water flow stops the calculated tdickness is small and the hcat trinsfr coefficient suddenly a increase resulting the larie pressure decrease shown in Fig. 1.

I Another problem is the liquid temperature assumed for thc 0.5 water in the film. Nhen a watcr level enters a ccll, the liquid temperaturc used to evaluate the wall fitm condensation heat

'flux should be the abovc ccll liquid tmpraurc instcad of ihe cell liquid ixte temperature.

RELAP5 was altered to perrorm like TRAC-M; e tum off vall condcnsation when awar kvcl nte a ccil. Rcsults are 0 .4 5 '. . . . . . .

shown in Fig 2. The pred;cted pressure is imprved with the 0 21) 40 6° X' f ltered code.lowever thephysics p is still not modeled correctly.

Tume (a)

The co'idel&imi ihiniiin eors in aemh eo would not FIgure 1. Brecasepredktionof prsress. likely have been uncoyered ifeach cod were assessed independent ofthc other code.:

The two codes show fairly accure results during the ;

compression period but RELAPS hi too much condensatic :n :

2 2CopyTight 199 O by ASME

LIC-03-0081 Attachment Page 39 0.6 035 I 1-1 0.

05 j

0.4s 0 20 40 60 Time(s)  : Tiue (s)

Figure 2. REIAPS pressure pretiklion with no will Figure 3. Pressure predction with thermal condensation In vertcal stratficadton. stratiflicaton miodel Therml Front 450 A thermal stratification modcl is included in RELAP to inprove the accuracy of solutions when there is warm fluid above cold fluid in a vertical stick of celts (ISL Inc., 2001a4 42Sf Bccausc dcfiultRELAPS uses a first-otder semI-implicit 0 upwind difTerencing scheme, axial numcrical diffusion of cold water occurs. This bs an unfavorable effect on the accacy of 400 the tcmpcraturc profile. Thc modecl is ctivated when thesr is a density difference bctween upstteam and downstream cclls. Tbe e 375 model achieves a sharp temperature profile by spccifying the liquid enrg cassing a cell junction to be the downstream cell I:1.

erngy. The model Istumed off when the cell liquid cncrgy I_ 0 OUqTr4tUrC-R5b&W equals the upsuram cell liquid c.rgTU 1! DLqT=ma=nr-R T F'rvt The calculated pressure in the top of Ihe vsscl using the 325 1 themal stratification model is shown in Fig. 3. Improved resutts arc obscrved during and after the liquid insurge into the vessel. . .I .

300 0 k-iq1 VOA pitkR" A reason for de Improved pressure behavior is demonstrated in the fluid tempeature response shown in Fig. 4. The thermat front model shatpcns the axial tempratureprofile. The pressure :2t . . .

is higher with the model active becausc the liquid at the vapor- I :0: : , 0.7 IA liquid interface is hottcr; limitingcondrnisation. Axial Ekvation (m)

Level Tracking Figure 4. Axial flid temperature profile vt 35 swonds.

The level tracting model implemented in RELAPS includes a detection of t: lcvl appearance, calculation of mixture level

~3 Copyright 0 1999 byASME

LIC-03-0081 Attachment Page 40 parameters such as position end velocity of the level and void numerical scheme to solve the consrvation equations as fiations abov and below the Ievel, mixwre level movcment recommended in the users guide (ISL Inc., 2001b). TRAC.M from cell to ce]l1,mass and energy equation modifications, and base rcsults use the implicit advancement scheme since it is the beat transfer calculation modifications. default The base maximum timc step size was set at 0.01 RELAP5 level trackng model applied to this problem seconds for both codes.'

showed only slight improvementis in the predicted pressure. Figwe 6 shows that th RELAPS predicted pressure rise using Duing the presure rise portion of the tr4nsft th predicted tc ncarly-implicit solution bas sore problems when the watcr pressure laid nearly on top of the RELAPS bse rm. Th more levelcrossscellboundarisataboutl2,24and36seconds-notabie improwvcent in the prcdicted pressure was after the TRAC-M des not encounter the sane type orcell boundazy waer insurge was stopped. The predieted pressure after the crossing probei as docs RELAP5 (cr Fig. 1). The TRAC-M watcr insurgc was stopped, laid between the base run and the semi-implicit rmsults ovcrlaythe implicit results at the base run case where the thermal front model was activated. time step size of.0I second..

TRAC-M has new level tracking logic as discussed in Aktas, 0.6 (2002). Whcn the level tracking logic is applied to this problem.

the rcsults arc greatly improved at shOwn in Fi. 5. The peak pressu rises when the level tracking model is active. The logic uses the semi-implicit nuridcal sch C and assuIes thc interfacial heat transfer is zero in the cell with the water level.

05 I

0.55 V's 0.5 0.4S L.

0 20: 40 i X Times)(a Flgure 6. RELAPS pressure using mpliit numerks.

OAS Time Step Size nme (s)

A key to obtainng satisfactory predictions is controlling the Figure S. TRAC-M pressurepredletlen using theleel timc step size. The iplicit numerical solution method allows trncking nedet for the time step size to be larger than the material Courant limit while the semi-impblicit method does not.

RELAPS, with level traking on, results in the same small Figure 7 shows time step size vetsus time using the base value of intetfacial heat transfer as wilh it off. This is because codes with hc maximum time step size raised from 0.01 to O.5 the dcfault RELAPS model, known as the "vertieal seconds. TRAC-M rached the naximu time stepsiz for four stratification" model, already sets the interfacial area to be the inteals dring tt transient. The decrcases in the time step are cross-sectional area. i4: :f- X-trelated to the liquid level crossing cell boundaries. RELAP5, howevcr, was limited by the Courant time step size (dtcrnt)

Numerical Implicitness since the basc case runs in sCMi-implicit modc.

RELAPS base calculations employ the semi-implicit Figure S demonstrates that time step size sensitivity

4 copyright 01999 b ASME

LIC-03-008 1 Attachment Page 41

0. 0.6 0.6 0.53 I I

.4 04

05

[  :.4

0 Ij

. 20 -40 60 0 I :::  : 40 60 TicK(s) Time (s)

Figure 7. Time tep comparlson for a naximam fiume Flgure &Effect or time step sie on TRAC41 step size of 0.5 seconds, predl . -

calculations sluld be performed before accepting a rdiction.

Shown arc TRAC-M implicit pressure predictions using time step sizes of 0.01 and 0.5 seconds Onc possible reason for the lower pressure in the large time stcp size casc is the toiduction .6 solution is not implicitly coupled to the hydraulic solution. This would allow the wall heat flux from vapor to be too large when thc saturationtrnrnaturc Is rising.

0.55 Figurc 9 shovs that RELAP5 poorly ptedicts the pressure afetr 12 seconds when using nearly-implicit numerics and allowing large tirne steps, RELAPS has the ability to perform a conduction solution implicitly coupled tothc bydraulicsolution.

Itowevcr. checks showed the nearly-implicit hydraulic solution was so bad that implicit condcon coupling did not yikld improved results, CONCLUSIONS Both RELAP5 and TRAC-Mi prcdict the MIT pressuizer data teasonably well providcd the time step sizes are not large". The effccts of choosing arious codc options havc been 0.45 demonstrated. The bencfits operfronning calculations with two codes makes codc crrors morc obvious and the codes can be I:ree (s) more rapidly i.proved.

XEFERENCES FIgure 9. RtlspS prediction using Implicit numerics Iwith a large nuximum time step size.

Aktas. D. and Uhc. J 2000, "USNRC Code Consolidation and Development Eon, Procedirof he OECD-C I 5 Copyright I 1999 by ASME

LIC-03-008 1 Attachment Page 42 brkshop om Advnced Tenual4-l t andNeutronk Codes CurrenandFuWeApiatios, Barceiona,i Spain.

Aktas, B., 2002, "Tracldng Interacs in Vertical Two-phase flown"! ThnihInieraioamlConfrnmee on AlctarEnergy,the American Society of Mechanical Engineers, Virginia.

Kim, S. N., andfGrith,P, 1987, -P%VR Pressurizer Modeling",

NuclearEngleerngandDesign10 2,pp. 199-209.

Nusselt, W, 1916, "Die Oberflachenkondensaticn des Wasscrdampfes7" Z Ver. Deutsch.ling., FbI.60, pp. 541-569 TSL Inc.,2001a, "RELAP5MOD3.3Beta Code Manual Nbluno 1: Code Structure' Systam Models, and Solution Methods",

NUREG/CR-5535Rev 1.

IS1. Inc., 2001b, RELAP5iMOD33Befa Code Manual VolumC 5: Uscr Guidelines. NUREG/CR-553SRcv 1.pp 83.

Sacdi, H.R, and Griffith P,1983, 'Thc pressu Responsc ofa PWR Prssurizer During an lnsurv Tsient, Transctions of AAS. 1983 Annuat Meeting, Dcroit, Michigan, June 12-16, pp.

606-607.

Spore, J. W, ct a., 1993, "TRAC-PF1MOD2 Theory Ntanual" LA-1203 1-M, Vol 1,NUREG/CR-5673.

6 ICopyright 0 1999 by ASME :

LIC-03-0081 Attachment Page 43 Appendix 6 Effect of Non-Condensables in Pressurizer A confirming evaluation was performed which concludes that assuming non-condensable gases present within the steam in a pressurizer bubble has no significant impact. The worst case in terms of shrinking the bubble, Case 10 HP509S30 of Reference D, was rerun with an assumed 10% non-condensable gas. The bubble was slightly smaller and the pressure slightly higher, but the change does not impact the analysis conclusions. Case 12 of Reference D, where the input is adjusted to result in a PORV opening, was also re-performed with 10% non-condensables.

Model changes (only two cards change):

1) Added default non-condensable gas (air was used but nitrogen could have been used without a significant difference since air is 80% nitrogen).

2) Changed bubble card to be same pressure of 95 psia, but slightly lower temperature to keep steam saturated. That is, the steam temperature was set to the saturation temperature at 85.5 psia (316.7°F) since that is the partial pressure of steam, i.e., 0.9*95, at 90% quality where here quality is defined as fraction of steam/(steam + air).

Old card:

4101202 2 95. 1.0 0. 00,5 New cards:

110 air 4101202 4 95. 316.7 0.9 00,6 For Case 12 the change is to:

4101202 4 400. 434.4 0.9 00,6 Results:

For Case 10, there is still a significant bubble left at 600 seconds, but it is slightly reduced from 139.0 ft3 to 130.1 ft3 . Without non-condensables, the pressure only went up to 100.3 psia and then came back down. For this analysis it rose to 112 psia, and then returned slowly to 122 psia at 600 seconds. Please refer to the figures below for comparison.

The conclusion is that the non-condensable gases have no significant impact. The bubble is very slightly smaller, but there is still adequate bubble remaining after 10 minutes. The pressure is slightly higher, but that is unimportant since the PORV setpoint is still far away from being reached. Case 12 was run to show that even if the PORV does lift (due to an assumed high initial pressure) the transient is successfully mitigated. Re-performing Case 12 showed a tiny increase in the peak pressure (from 476 to 478 psia), but more rapid depressurization after the valve opens and a growing bubble at 600 seconds.

It was determined that these cases were sufficient in determining the effect of non-condensables on the LTOP analysis and were not needed on the mass addition cases. The resultant conclusion

LIC-03-0081 Attachment Page 44 would be unchanged since none of the heat addition cases are limiting when compared to the mass addition events.

Reference D: FC06877, Rev. 0, "Low Temperature Overpressure Protection (LTOP) Analysis, Revision 1." [Note: This Reference was included in Letter from OPPD (D. J. Bannister) to NRC (Document Control Desk) dated October 8, 2002, Fort Calhoun Station Unit No. 1 License Amendment Request, "Reactor Coolant System (RCS) Pressure and Temperature Limits Report (PTLR)" (LIC-02-0109)]

Old case, no non-condensable gas Figure 13: Case 10 H509S30 Results Pressurizer Bubble E

6 200 -

T M 100-3 0- I o 200 400 600 Time (seconds)

New case, 10% non-condensable gas Figure S2: HP509air Results Pressurizer Bubble

_400 -

0 300 -

E 6 200-n0

.0 100-0 200 400 600 Time (seconds)

LIC-03-0081 Attachment Page 45 Old case, no non-condensable gas Figure 14: Case 10 HP509S30 Pressurizer Pressure 102 101 -

, 100 -

Ts7 99.-

-. 98 -

297-96 95 95 ao94 -

93.

92-,,

0 200 400 600 Time (seconds)

New case, 10% non-condensable gas Figure SI: Case HP5O9air Figure S: Case HP509air Pressurizer Pressure 130 -

.0 120-0.

2 110-0 to E 100-90 I I I 0 200 400 600 Time (seconds)

LIC-03-0081 Attachment Page 46 These are from Case 12 with the same changes made to HP504S30.txt:

Old Pressure curve:

Figure 20: Case 12 HP504S30 Pressurizer Pressure 480 470

- 440 Q 430 j 420 410

& 400 390 380 0 200 400 600 Time (secondE New Pressure curve:

Figure S3: Case HP5O4air Figure S3: Case HP504air Pressurizer Pressure 500 -

ei 450-0L

, 400-C)

" 350 0.

300 0 0 60 0 200 400 60 Time (seconds)

LIC-03-008 1 Attachment Page 47 Old Bubble curve, Case 12:

Figure 19: Case 12 H504S30 Results Pressurizer Bubble

_ 300-c- 250 -:)

E 200- \

6

.0 150 -- \

4) 100-0 500 0 200 400 600 Time (seconds)

New Bubble:

Figure S4: HP504air Results Pressurizer Bubble

_ 300-

• ! 250-E 200-6 150 0 100

.0 0 50 km 0 4 6 0 200 400 600 Time (seconds)