ML031830503

Cold Leg Elbow Tap Flow Coefficients

ML031830503
Person / Time
Site:	Catawba
Issue date:	07/01/2003
From:	Martin R NRC/NRR/DLPM/LPD2
To:	Gordon Peterson Duke Energy Corp
Martin R, NRR/DLPM, 415-1493
References
Download: ML031830503 (15)

Text

July 1, 2003 Mr. G. R. Peterson Vice President, Catawba Site Duke Energy Corporation 4800 Concord Road York, SC 29710

SUBJECT:

CATAWBA NUCLEAR STATION, UNIT 2 RE: COLD LEG ELBOW TAP FLOW COEFFICIENTS

Dear Mr. Peterson:

By letters dated October 2, 2001, and March 19, 2003, the U. S. Nuclear Regulatory Commission (NRC) staff transmitted amendments to the Facility Operating License for the Catawba Nuclear Station, Unit 2, concerning the determination of reactor coolant system (RCS) flow rate. In the Safety Evaluations transmitted with those letters, the NRC staff stated that reactor coolant pump energy was incorrectly addressed in the Duke Power Company (Duke) application and that a correct calculation would provide a conservatism that Duke had not elected to take credit for. The NRC staffs conclusion, in this regard, was incorrect. The existing Duke calculations that supported the information provided to the NRC staff in the associated Duke submittals is correct. Since the error relates to an additional conservatism that was not incorporated into Dukes submittals to the NRC staff, the error does not affect the NRC staffs prior findings that the license amendment applications submitted by Duke are acceptable.

The error is related to the modeling of the thermodynamics of RCS flow. The enclosure discusses the validity of various modeling approaches and confirms the NRC staffs prior findings, issued on October 2, 2001, and March 19, 2003, that the license amendment applications submitted by Duke, are acceptable.

This information is provided to ensure that the record is correct on this issue. Please contact me at (301) 415-1493, if you have any other questions on these issues.

Sincerely,

/RA/

Robert E. Martin, Senior Project Manager, Section 1 Project Directorate II Division of Licensing Project Management Office of Nuclear Reactor Regulation Docket No. 50-414

Enclosure:

Supplement to Safety Evaluation cc w/encl: See next page

July 1, 2003 Mr. G. R. Peterson Vice President, Catawba Site Duke Energy Corporation 4800 Concord Road York, SC 29710

SUBJECT:

CATAWBA NUCLEAR STATION, UNIT 2 RE: COLD LEG ELBOW TAP FLOW COEFFICIENTS

Dear Mr. Peterson:

By letters dated October 2, 2001, and March 19, 2003, the U. S. Nuclear Regulatory Commission (NRC) staff transmitted amendments to the Facility Operating License for the Catawba Nuclear Station, Unit 2, concerning the determination of reactor coolant system (RCS) flow rate. In the Safety Evaluations transmitted with those letters, the NRC staff stated that reactor coolant pump energy was incorrectly addressed in the Duke Power Company (Duke) application and that a correct calculation would provide a conservatism that Duke had not elected to take credit for. The NRC staffs conclusion, in this regard, was incorrect. The existing Duke calculations that supported the information provided to the NRC staff in the associated Duke submittals is correct. Since the error relates to an additional conservatism that was not incorporated into Dukes submittals to the NRC staff, the error does not affect the NRC staffs prior findings that the license amendment applications submitted by Duke are acceptable.

The error is related to the modeling of the thermodynamics of RCS flow. The enclosure discusses the validity of various modeling approaches and confirms the NRC staffs prior findings, issued on October 2, 2001, and March 19, 2003, that the license amendment applications submitted by Duke, are acceptable.

This information is provided to ensure that the record is correct on this issue. Please contact me at (301) 415-1493, if you have any other questions on these issues.

Sincerely,

/RA/

Robert E. Martin, Senior Project Manager, Section 1 Project Directorate II Division of Licensing Project Management Office of Nuclear Reactor Regulation Docket No. 50-414

Enclosure:

Supplement to Safety Evaluation cc w/encl: See next page DISTRIBUTION:

PUBLIC OGC CHawes WLyon PDII-1 R/F ACRS RMartin JUhle JNakoski LPlisco, RII RHaag, RII ADAMS Accession: ML031830503 OFFICE PDII-1/PM PDII-1/LA PDII-1/SC NAME RMartin CHawes JNakoski DATE 06/25/03 07/1/03 07/1/03 OFFICIAL RECORD COPY

SUPPLEMENT TO SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION DETERMINATION OF REACTOR COOLANT SYSTEM FLOW RATE FROM A CALORIMETRIC HEAT BALANCE ON THE SECONDARY SYSTEM DUKE ENERGY CORPORATION CATAWBA NUCLEAR STATION, UNIT 2 DOCKET NO. 50-414 1.0

SUMMARY

Reactor coolant system (RCS) flow rate is often determined by first performing a calorimetric heat balance on the secondary side to determine the heat transfer rate from the RCS into the steam generators (SGs). This heat transfer rate is then used in conjunction with RCS hot leg and cold leg temperature measurements to determine flow rate and to calibrate cold leg elbow tap flow meters. The Nuclear Regulatory Commission (NRC) staff recently completed two reviews of the determination of RCS flow rate for Duke Power Companys Catawba, Unit 2 (References 1 and 2). During further investigation of RCS flow rate, the NRC staff found an error in those reviews. Although the error does not change the NRC staffs finding of acceptability for the Duke Power amendment requests, it does involve modeling of the thermodynamics of RCS flow rate that should be clarified. This supplement to the previous safety evaluations, issued on October 2, 2001, and March 19, 2003, provides that clarification.

2.0 BACKGROUND

The RCS configuration illustrated in Figure 1 is typically used as a basis for flow rate analyses. An overall heat FIGURE 1. RCS LOOP CONFIGURATION balance is taken over the control volume defined by the surface of the RCS to obtain Equation 1: STEAM GENERATOR Qcore + QRCP - Qloss = Qcal (1)

HOT LEG where: Qcore = nuclear heat generation rate in the core, REACTOR VESSEL QRCP = rate of energy addition to RCS by reactor coolant pumps (RCPs), REACTOR COLD LEG Qloss = net rate of RCS heat loss exclusive COOLANT PUMP of SGs and RCPs, and Qcal = rate of heat removal by SGs.

This formulation assumes no mass transfer through the control volume surface1 and the control volume is constant (the boundaries are rigid). Since Qcal is determined from the calorimetric, and Qloss and QRCP can be estimated, this equation will provide Qcore.

With Qcore determined, one may assume that mass flow rate through the core could be determined by dividing Qcore by the enthalpy difference across the core if the enthalpy difference were known. However, these enthalpies are not known because core inlet temperatures are not measured and there is a wide variation in core outlet temperatures with position across the top of the core. Consequently, it is necessary to make the determination where meaningful temperatures are measured. In practice, the hot and cold leg resistance temperature device (RTD) temperatures, Th and Tc respectively, are used.

The NRC staff considers three different approaches to determine mass flow rate, M. The first is used in many typical fluid flow calculations. It simply assumes a constant pressure process with the following expression:

M = (Qcore - Qloss T) / [Cp (Th - Tc)] (2) where: Qloss T = heat loss rate associated with all RCS components located between the cold and hot leg RTDs.

Cp = heat capacity at constant pressure.

Although it is not immediately apparent, Equation 2 does not correctly account for the effects of water flowing within the RCS between the measurement locations of Tc and Th. This understanding is developed in Section 3, below.

The second approach depends upon the calculation of RCS pressures from an assumed RCS flow rate. With pressures determined, one can determine the hot and cold leg enthalpies, hh and hc respectively.2 Then RCS flow rate in the hot and cold legs can be determined by the following relationship:

M = (Qcore - Qloss T) / (hh - hc) (3)

This is the approach used by the Duke Power Company that the NRC staff discussed in References 1 and 2. Although not immediately obvious, it is shown in Section 3, below, that this is a thermodynamically correct representation that follows, in part, from the difference in hot 1

A complete formulation must consider mass flow and all meaningful heat flows through the control volume boundary. This includes charging flow (+), letdown flow (-), seal injection flow (+), RCP thermal barrier cooler heat removal (-), pressurizer spray flow (-), pressurizer surge line flow (+),

component insulation heat loss (-), component support heat loss (-), and control rod drive mechanism heat loss (-). The effect of these items is often incorporated into Qloss a step that must be considered carefully if mass flow rates are involved.

2 The initially determined RCS pressure distribution will be for an assumed RCS flow rate.

Iteration to obtain a converged solution may not be necessary since enthalpy is a relatively weak function of pressure. However, there are situations where it may be necessary to iterate to obtain a converged solution.

and cold leg pipe diameters and the equal mass flow rates entering and leaving the control volume.

One can also assume RCP potential energy associated with the pressure increase at the RCP is converted to frictional heat throughout the RCS in direct proportion to the pressure drop.

Therefore, since the RCS pressure distribution is known as a function of M, the portion of RCP heat that is distributed between Tc and Th, QRCP T, can be calculated and:

M = (Qcore + QRCP T - Qloss T) / (hh - hc) (4)

In Reference 1, the NRC staff estimated that the difference in M due to accounting for QRCP T was 1450 gpm and the NRC staff concluded that the model represented by Equation 4 should be used. In Reference 3, the licensee agreed in principle with the NRC staffs determination, and calculated that failure to incorporate QRCP T into the model under-predicted M by about 1045 gpm. The licensee elected to continue using Equation 3, the model without QRCP T, since it provided a conservative value of M. The NRC staff accepted this justification in Reference 2.

The NRC staff now shows in Section 3, below, that Equation 4 is incorrect.

3.0 DISCUSSION In the previous section, three typical approaches for determining RCS flow rate, M, were presented in Equations 2, 3 and 4. Each case is an implementation of Equation 1 using parameters measured in the plant or parameters that can be derived, such as the pressure distribution. Only Equation 3 is valid as is discussed below.

3.1 Thermodynamics of Fluid Flow Equation 3 is a simplified expression of conservation of energy of a flowing system. This is shown to be the case by considering a general control volume and summing the forms of energy entering and leaving the volume through the control volume boundary:

heat + [mass flow rate]in{ internal energy + flow energy + kinetic energy + potential energy}in =

work + [mass flow rate]out{ internal energy + flow energy + kinetic energy + potential energy}out, or:

Q + Min { u + [ 144 P V + v2 / (2 g) + Z ] / f }in = W + Mout { u + [ 144 P V + v2 / (2 g) + Z ] / f }out (5) where: Q = heat addition rate, Btu/sec M = weight flow rate, lbs/sec u = internal energy per unit weight, Btu/lb P = pressure, lbs/in2 absolute V = volume per unit weight, ft3/lb v = velocity, ft/sec g = gravitation constant, 32.2 ft/sec2 Z = elevation, ft f = conversion factor = 778 ft-lbs/Btu W = work performed by the fluid, Btu/sec

Since, by definition, enthalpy is:

h = u + 144 P V / f (Btu/lb) (6)

Equation 5 may also be written as:

Q + Min { h + [ v2 / (2 g) + Z ] / f }in = W + Mout { h + [ v2 / (2 g) + Z ] / f }out (7)

Equation 7 may be applied to the RCS by selecting a control volume that encloses the reactor vessel and the pipes between the reactor vessel and the locations of Th and Tc. There is no work done by the system within this control volume and W = 0. The hot and cold leg pipe elevations are identical and Zin = Zout. There is no accumulation of mass and Min = Mout = M.

We will also assume that vin = vout. This velocity assumption is, of course, a simplification but because the heat addition is large, and the differences in hot and cold leg velocity are small (because of the increase in hot leg diameter that accommodates the decrease in fluid density),

the error introduced by this simplification is negligible. Equation 7 therefore reduces to:

Q + M hc = M hh (8)

But Q is Qcore minus that portion of Qloss associated with the reactor vessel and the pipes between locations of Th and Tc, Qloss T. Thus,:

M = (Qcore - Qloss T) / (hh - hc) (9)

Since Th and Tc are measured, and Ph and Pc are known as a function of M, h can be obtained from known water properties. Thus, Equation 9 will provide M.

Equations 9 and 3 are identical. Consequently, the NRC staffs conclusion, reached in Reference 1 that Equation 4 was a better representation of the behavior than Equation 3 is incorrect.

Thermodynamically, the enthalpy terms in Equation 9 (and Equation 3) address the energy transition between potential energy and thermal energy. QRCP is not modeled because the pump energy is added to the system outside the control volume by accounting for the pressure rise, inefficiency, and velocity changes across the pump. Since the process is not modeled as isobaric, enthalpy is a function of both pressure and temperature. Therefore, the enthalpy addition stemming from the increase in pressure across the pump is accounted for within the control volume and does not have to be incorrectly modeled as a heat addition term as is done in Equation 4.

3.2 Hot Leg and Cold Leg Flow Velocities Now the assumption that vin = vout is examined by calculating RCS behavior using the flow model described in Reference 4 to provide the pressure distribution throughout the RCS for a flow rate of 100,476 gpm (10337 lbs/sec) in one loop with Th = 620 °F. All control volumes are selected so that Z is constant and the assumption is made that Qloss = 0 to simplify the calculation. The calculation path is summarized in Table 1. The path is initiated at the hot leg where temperature and pressure are known so that fluid properties can be calculated.

(Knowledge of two properties is both necessary and sufficient for determination of all properties

in a single component, single phase fluid.) It is now assumed that cold leg pressure is known and an assumption is made of an isenthalpic compression from the hot leg conditions to obtain the cold leg conditions. The process of working backward is continued to obtain RCP discharge conditions. Then RCP suction conditions can be calculated by considering the pump in steps as an isentropic process, heat addition to account for inefficiency, and a kinetic energy change due to velocity change. The flow is assumed to be isenthalpic to obtain properties in the pipes at the exit and entrance of the steam generator. The kinetic energy change at the RCP and at the SG is essentially the same; 0.25 MW. This is added to the fluid energy at the RCP to obtain the total RCS energy, and subtracted during the energy balance at the SG to obtain the heat removed from the RCS at the SG.

Table 1. Calculation Path Assumptions RCS Location Known P, Process and/or comments psia Hot leg at Th 2225 Physical properties, including h, known from known P and 620 °F temperature via steam tables Isenthalpic compression to 2273.61 psia, decrease h by 862.75 MW core heat Cold leg at Tc 2273.61 Properties from known h and 2273.61 psia Isenthalpic compression to 2276.88 RCP exit 2276.88 Properties determined by above known h and 2276.88 psia Assume pump is represented by of heat addition, isentropic compression, and incorporation of kinetic energy Decrease h by assumed RCP inefficiency of 0.7 MW at constant P to get new h; known P determines new properties Decompress via isentropic expansion to 2186.43 psia RCP entrance 2186.43 Properties determined by known entropy and 2186.43 psia Isenthalpic compression to 2190.54 psia Steam generator 2190.54 (SG) exit Hot leg at Th 2225 Isenthalpic decompression to 2223.94 psia SG entrance 2223.94 Properties determined by known h and 2223.94 psia The results are summarized in Figure 2. The results demonstrate that vc and vh are essentially equal, confirming the assumption used to obtain Equation 8.3 Note that this is not true for flow 3

The velocities could have been determined by a simple calculation based on known P, T, and flow area at Tc and Th. We used the above approach because it introduces additional characteristics that are of interest.

through the RCP and through the SG, where the velocity increases from 42.9523 ft/sec to 54.5389 ft/sec and decreases from 54.5754 ft/sec to 42.9498 ft/sec, respectively.

FIGURE 2. THERMODYNAMIC ANALYSIS OF RCS T = temperature, °F Flow Rate =

STEAM GENERATOR P = pressure, psia 100,476 gpm h = enthalpy, Btu/lb (10337 lbs/sec)

S = entropy, Btu/lb-°F

- 857.72 MW = density, lbs/ft3 v = velocity, ft/sec 29 in ID HOT LEG T = 563.587 Th = 620 P = 2190.54 P = 2225 h = 564.427 h = 642.997 REACTOR S = 0.76037 S = 0.83491 VESSEL

= 45.9199 = 41.2946 v = 42.9498 v = 54.5754 T = 564.069 Tc = 564.063 + 852.75 MW P = 2276.88 P = 2273.61 31 in ID h = 564.859 h = 564.859 CROSS-OVER S = 0.76045 S = 0.76046 PIPE = 45.9530 = 45.9509 v =54.5389 v = 54.5414 REACTOR 27.5 in ID COLD LEG COOLANT PUMP

+ 4.97 MW T = 563.581 P = 2186.43 h = 564.427 S = 0.76039

= 45.9172 v = 42.9523

3.3 An Anomaly: RCS Water Cools Due To Friction The Figure 2 results refute the generally accepted conclusion that in a flowing system with a viscous fluid, frictional pressure drop causes the fluid temperature to increase, as assumed in Equation 4. The reverse behavior is illustrated in Figure 2 where, for example, temperature decreases from 563.587 °F at the SG outlet to 563.581 °F at the RCP inlet.

To address this apparent anomaly, consider the calculation results summarized in Figure 3 for conditions at roughly 100 psia and 100 °F with the additional assumption of no core heat addition. In this case, the temperature increases from Tc = 100.098 °F to Th = 100.183 °F, as intuitively expected.

FIGURE 3. RCS BEHAVIOR AT LOW PRESSURE T = 100.183 °F STEAM GENERATOR P = 165 psia h = 68.67921 Btu/lb S = 0.12988 Btu/lb-°F HOT LEG REACTOR COLD LEG COOLANT 100.098 °F REACTOR PUMP 197.07 psia VESSEL 68.67921 Btu/lb 0.12971 Btu/lb-°F 100 °F 108.68 psia 68.3475 Btu/lb 0.12959 Btu/lb-°F The NRC staff concludes that a thermodynamically based calculation yields the expected temperature behavior when temperature and pressure are relatively low, but the reverse occurs at typical RCS operating conditions. Such behavior is typical of the Joule-Thomson throttling process where h = 0. The behavior is described by the Joule-Thompson coefficient:

µ = (MT /MP)h (10)

The condition where µ = 0 represents the inversion temperature. The throttling process will result in cooling if µ > 0 and in heating if µ < 0. Therefore, the sign of the Joule-Thompson coefficient has reversed when changing the process from water near room temperature conditions to typical RCS operating conditions. Although not expected from our experience, in part because the effect is small, the results are consistent with theory.

3.4 RCS Heat Loss Rate Next, the effect of heat loss is considered. It would be expected that the RCS heat loss rate is about 25 percent of the RCP heat, or about 1.25 MWt per loop (including 1/4 of the reactor vessel per loop). As an estimate, assume about half of this heat is lost as the water flows from the location of Tc to the location of Th. When compared to the adiabatic case, this will reduce the Figure 2 Th - Tc by about 0.066 °F and will increase the calculated flow rate for four loops by about 340 gpm.

Failure to include heat loss will decrease calculated core heat generation rate by about 0.16 percent (see Equation 1) and will decrease predicted flow rate by about 0.08 percent (see Equation 3). Hence, failure to include heat loss rate will under-predict the RCS flow rate, a conservatism.

3.5 The Effect of Thermal Power Some calorimetric determinations are done at less than 100 percent core power. To assess this situation, Equation 7 is applied with the assumptions of no change in M, Z, and W, to obtain:

Q + M { h + v2 / (2 g f) }in = M { h + v2 / (2 g f) }out (11) or:

M = (Qcore - Qloss T) / [ hh - hc +(vh2 - vc2) / (2 g f) ] (12)

Taking Qloss T = 0, assuming an isenthalpic expansion from Tc to Th, and adding Qcore, hh can be obtained if hc is known, and hc can be determined if Tc and Pc is known. Consequently, it is assumed as a first step that the pressure distribution is the same as used in development of Figure 2 and behavior as a function of Qcore is examined with the average of Th and Tc approximately constant. The calculations and results are provided in Table 2. The results show that power level has no practical influence on the RCS flow rate if the RCS pressure distribution is assumed constant. (Note, this does not address the accuracy of the calorimetric determination. In general, accuracy will diminish as power level approaches a small value.)

In Reference 5, Duke Power showed that a variation in power from zero to 100 percent changed flow rate by about 4000 gpm (1000 gpm per loop) as measured by elbow tap flow meters. Although this will have some effect on pressure distribution throughout the RCS, the NRC staff does not believe it would have a significant effect on the tabulated comparison conclusions.

Table 2. Effect of Power Variation on Determination of RCS Flow Rate Item 100% Power 90% Power 80% Power Tc 564.063 566.8 569.6 hc 564.859 568.336 571.997 c 45.9509 45.7625 45.5669 vc 54.5414 54.7656 55.0010 Th 620 617.202 614.484 hh 642.997 638.660 634.508 h 41.2946 41.5779 41.8457 vh 54.5754 54.2035 53.8567

( vh2 - vc2 ) / (2 g f) 0.00007 -0.00122 -0.00249 M 10337.4 10337.6 10337.7 M@ v=0 10337.0 10337.4 10337.3 3.6 Comparison of Equations 2 and 3 Two equations were previously provided that differed in the calculation of heat from a temperature determination:

M = (Qcore - Qloss T) / [Cp (Th - Tc)] (2) and:

M = (Qcore - Qloss T) / (hh - hc) (3)

The NRC staff further concluded that Equation 3 is correct if values of h are determined from T and P. As seen from the above Section 3 discussion, thermodynamic considerations yield Equation 3 as a solution. For Equation 2 to be correct, it would be necessary for Cp (Th - Tc) to equal hh - hc. The effect of pressure can be illustrated by considering the typical engineering calculation process of using tabulated values corresponding to saturation temperature. At 610 °F and the saturation pressure of 1665 psia, h = 631.568 Btu/lb. At 2225 psia and 610 °F, h = 627.782 Btu/lb, a significant difference. This is an indication that enthalpy is a function of pressure and that assuming a constant pressure process is inaccurate. Consequently, use of Cp is questionable.

3.8 Assessment of Reference 1 and 2 Reviews References 1 and 2, and the Duke Power Company documentation reviewed in support of those references, did not identify whether Equation 2 or Equation 3 was used to calculate RCS flow rate. This was covered in previous documentation, such as Reference 6, and established that the correct Equation 3 was used.

As discussed in Reference 2, Duke Power Company elected not to credit the additional flow rate that would have been provided via application of Equation 4. The NRC staff noted this

several times in its review by crediting this perceived margin. The most specific reference was the Section 3.3.4 statement that not correcting for the RCP thermal energy error of about 1045 gpm ... clearly compensates for any NRC staff concerns regarding the effect of impeller smoothing. In the same section, the NRC staff also stated that there was a change of 1000 gpm over the time span applicable to the licensees selection of the September 1986, November 1986, and March 1988 calorimetrics for determination of elbow tap coefficients that introduces a conservatism. In Section 3.3.1.1, the NRC staff referenced the licensees estimate that the selected calorimetrics introduced a 1500 gpm conservatism and, although the NRC staff did not specifically prove the licensees value to be correct, the NRC staff provided additional insights to support the conservatism. The licensee additionally selected cold leg RTDs that introduced close to 2000 gpm additional conservatism when compared to selection of normal cold leg RTDs. These licensee selections establish that there is sufficient conservatism to address uncertainties associated with flow rate determination. Consequently, with the exception of the statement that failure to correct for RCP heat was an error, the Reference 2 findings remain valid.

4.0 CONCLUSION

S The Reference 1 and NRC reviews of Duke Power Companys determination of RCS flow rate contain an error. The NRC staff stated that RCP energy was incorrectly addressed and that a correct calculation would provide a conservatism that Duke elected not to credit. The NRC staffs conclusion was incorrect. The existing Duke calculation is correct. Other conservatisms remain and the NRC staffs finding of acceptability of the Duke Power amendment requests remain valid.

A thermodynamically-based assessment of RCS flow rate yielded several behavioral insights that were not immediately evident. These include:

1. Use of enthalpies in the RCS heat balance equations provides a correct accounting for RCP heat input and no corrections are necessary.

2. RCS design results in essentially equal hot and cold leg velocities so that there is no change in kinetic energy when comparing behavior at the hot and cold leg temperature measurement locations.

3. With the exception of the RCPs, the intuitive perception that RCS water temperature will increase due to friction during flow is incorrect. There is actually a slight decrease in temperature, an effect due to a positive Joule-Thompson coefficient for water at full power operating conditions. This condition does not exist at low temperature conditions typical of Mode 4 operation.

5.0 REFERENCES

1. Patel, Chandu P., Catawba Nuclear Station, Unit 2 Re: Issuance of Amendment (TAC No. MB1498), NRC letter to Duke Energy Corporation, October 2, 2001.

2. Martin, Robert E., Catawba Nuclear Station, Unit 2 Re: Issuance of Amendment, Cold Leg Elbow Tap Flow Coefficients (TAC No. MB6529), NRC letter to Duke Energy Corporation, March 19, 2003.

3. Peterson, Gary R., Duke Energy Corporation, Catawba Nuclear Station, Unit 2, Docket Number 50-414, Proposed License Amendment for Unit 2 Reactor Coolant System Cold Leg Elbow Tap Flow Coefficients, Duke letter to NRC, October 10, 2002.

4. Peterson, Gary R., Duke Energy Corporation, Catawba Nuclear Station, Unit 2, Docket Number 50-414, Response to Request for Additional Information for Unit 2 Reactor Coolant System Cold Leg Elbow Tap Flow Coefficients, Duke letter to NRC, February 7, 2003.

5. Peterson, Gary R., Duke Energy Corporation, Catawba Nuclear Station, Unit 2, Docket Number 50-414, Response to Request for Additional Information for Unit 2 Reactor Coolant System Cold Leg Elbow Tap Flow Coefficients, Duke letter to NRC, February 26, 2003.

6. Martin, Robert E., Summary of February 10, 1994, Meeting with Duke Power Company on RCS Flow Measurement Methodology, NRC, April 13, 1994.

Catawba Nuclear Station cc:

Mr. Gary Gilbert Senior Resident Inspector Regulatory Compliance Manager U.S. Nuclear Regulatory Commission Duke Energy Corporation 4830 Concord Road 4800 Concord Road York, South Carolina 29745 York, South Carolina 29745 Virgil R. Autry, Director Ms. Lisa F. Vaughn Division of Radioactive Waste Management Legal Department (PB05E) Bureau of Land and Waste Management Duke Energy Corporation Department of Health and Environmental 422 South Church Street Control Charlotte, North Carolina 28201-1006 2600 Bull Street Columbia, South Carolina 29201-1708 Anne Cottingham, Esquire Winston and Strawn Mr. C. Jeffrey Thomas 1400 L Street, NW Manager - Nuclear Regulatory Washington, DC 20005 Licensing Duke Energy Corporation North Carolina Municipal Power 526 South Church Street Agency Number 1 Charlotte, North Carolina 28201-1006 1427 Meadowwood Boulevard P. O. Box 29513 Saluda River Electric Raleigh, North Carolina 27626 P. O. Box 929 Laurens, South Carolina 29360 County Manager of York County York County Courthouse Mr. Peter R. Harden, IV York, South Carolina 29745 VP-Customer Relations and Sales Westinghouse Electric Company Piedmont Municipal Power Agency 6000 Fairview Road 121 Village Drive 12th Floor Greer, South Carolina 29651 Charlotte, North Carolina 28210 Ms. Karen E. Long Mr. T. Richard Puryear Assistant Attorney General Owners Group (NCEMC)

North Carolina Department of Justice Duke Energy Corporation P. O. Box 629 4800 Concord Road Raleigh, North Carolina 27602 York, South Carolina 29745 Elaine Wathen, Lead REP Planner Richard M. Fry, Director Division of Emergency Management Division of Radiation Protection 116 West Jones Street North Carolina Department of Raleigh, North Carolina 27603-1335 Environment, Health, and Natural Resources North Carolina Electric Membership 3825 Barrett Drive Corporation Raleigh, North Carolina 27609-7721 P. O. Box 27306 Raleigh, North Carolina 27611