ML15261A061

Forwards Proprietary Topical Rept DPC-NE-2003, Oconee Nuclear Station Core Thermal-Hydraulic Methodology Using VIPRE-01. Rept Withheld (Ref 10CFR2.790)

ML15261A061
Person / Time
Site:	Oconee
Issue date:	08/31/1988
From:	Tucker H DUKE POWER CO.
To:	NRC OFFICE OF ADMINISTRATION & RESOURCES MANAGEMENT (ARM)
Shared Package
ML15261A062	List: ML15261A061
References
NUDOCS 8809060384
Download: ML15261A061 (66)

Text

1 DUKE PROPRIETARY DPC-NE-2003 August 1988 I

DUKE POWER COMPANY OCONEE NUCLEAR STATION I CORE THERMAL-HYDRAULIC METHODOLOGY USING VIPRE-01 I

K. B. Jones D. R. Koontz I

Duke Power Company Design Engineering Department Nuclear Engineering Section Charlotte, North Carolina

'I I

I I

II

-PROPRIETARY INFORMATION NOTICE THE ATTACHED DOCUMENT MAY CONTAIN "PRO PRIETARY INFORMATION" AND SHOULD BE HANDLED AS NRC "OFFICIAL USE ONLY" INFOR MATION. IT SHOULD NOT BE DISCUSSED OR MADE AVAILABLE TO ANY PERSON NOT REQUIR ING SUCH INFORMATION IN THE CONDUCT OF OF FICIAL BUSINESS AND SHOULD BE STORED, TRANSFERRED, AND DISPOSED OF BY EACH RECI PIENT IN A MANNER WHICH WILL ASSURE THAT ITS CONTENTS ARE NOT MADE AVAILABLE TO UNAUTHORIZED PERSONS.

COPY NO.

DOCKET NO.

CONTROL NO.

REPORT NO.

REC'D W/LTR DTD.

NRC Form 190 PROPRIETARY INFORMATION (4-78)

ABSTRACT This report presents Duke Power Company's methodology for using VIPRE-01 for performing thermal-hydraulic analyses in support of Oconee Nuclear Station licensing activities. The VIPRE-01 thermal-hydraulic methodology and models are presented along with the results of sensitivity studies used in determining the acceptability of the various input criteria. This report meets the licensing requirements addressed in the Safety Evaluation Report for EPRI NP-2511-CCM, VIPRE-01, ref. 3.

I I

I I

I I

I I

Ii

Table of Contents 1 Section

1.0 INTRODUCTION

Page 1

2.0 CODE DESCRIPTION 2 3.0 STATION DESCRIPTION 3 4.0 CORE MODELING 3 4.1 STEADY-STATE SINGLE PASS MODEL COMPARISONS 6 4.2 TRANSIENT MODEL COMPARISONS 6 4.3 TRANSITION CORE MODEL COMPARISONS 7 4.4 RESULTS

SUMMARY

8 5.0 VIPRE-01 DATA 8 5.1 AXIAL NODING 9 5.2 ACTIVE FUEL LENGTH 9 5.3 CENTROID DISTANCE 10 5.4 EFFECTIVE CROSSFLOW GAPS 10 5.5 SPACER GRID FORM LOSS COEFFICIENTS 11 5.6 CORE BYPASS FLOW 11 1 5.7 INLET FLOW DISTRIBUTION 12 5.8 VIPRE-01 CORRELATIONS 12 5.8.1 FRICTION PRESSURE LOSS 12 5.8.2 TURBULENT MIXING 14 5.8.3 TWO-PHASE FLOW CORRELATIONS 15 I 5.9 REFERENCE DESIGN POWER DISTRIBUTION 17 5.10 AXIAL POWER DISTRIBUTION 17 5.11 HOT CHANNEL FACTORS 18 5.12 FLOW AREA REDUCTION FACTOR 19 5.13 BWC CRITICAL HEAT FLUX CORRELATION 19 1

1 111

• Table of Contents (Continued)

Section Page 6.0 OCONEE THERMAL-HYDRAULIC ANALYSES 22 6.1

SUMMARY

22 6.2 THERMAL-HYDRAULIC DESIGN CRITERION 22 I 6.3 CORE SAFETY LIMITS 23 6.4 PRESSURE-TEMPERATURE ENVELOPE 23 6.4.1 REFERENCE POWER DISTRIBUTION 24 6.4.2 CORE POWER 24 6.4.3 RCS FLOW 25 6.4.4 CORE INLET TEMPERATURE 25 I 6.5 GENERIC MAXIMUM ALLOWABLE PEAKING LIMIT CURVES 25 6.6 PUMP COASTDOWN TRANSIENT ANALYSES 27

7.0 REFERENCES

30 I

I II I

I I iv

PRiPRIETARY List of Tables DUKE POWER COQ I Table Title Page 3-1 Mark BZ Fuel Assembly Data...................... 31 4-1 Operating Conditions........................... 32 4-2 Comparison of[64 Channel and[8]Channel Model Steady-State Results........................ 33 4-3 Comparison of[64] Channel and[8]Channel Model Transient Results......................... 34 4-4 Comparison of[64] Channel and[ Channel Transition Core Model Steady-State Results.................. 35 4-5 Comparison of[64]Channel and[9]Channel Transition Core Model Transient Results.................... 36 5-1 L8]Channel Model Axial Node Length Sensitivity Study.............................. 37 5-2 [8]Channel Model Inlet Flow Distribution Sensitivity Study.............................. 38 5-3 [8]Channel Model Two-Phase Flow Correlation And Friction Multiplier Sensitivity Study......... 39 5-4 [8]Channel Model Turbulent Momentum Factor Sensitivity Study.............................. 40 6-1 RPS Trip Functions............................. 41 I

I 3 v

PROPRIETARY List of Figures DUKE POWER CO, Figure Title Page 4-1 [64]Channel Model [Eighth]Core Representation....... 42 4-2 [64]Channel Model Hot Assembly Detail.............. 43 4-3 [8]and[9]Channel Model Hot Assembly Detail......... 44 4-4 [8]Channel Model[Eighth core Representation........ 45 4-5 [9]Channel Model[Eighth] Core Representation........ 46 5-1 VIPRE-01 vs. LYNX2 M/P CHF........................ 47 5-2 VIPRE-01 vs. LYNX2 Mass Velocity ................. 48 5-3 VIPRE-01 vs. LYNX2 Quality ....................... 49 5-4 Measured vs. Predicted CHF ....................... 50 5-5 Measured-to-Predicted CHF vs. Quality ............ 51 5-6 Measured-to-Predicted CHF vs. Mass Velocity ...... 52 5-7 Measured-to-Predicted CHF vs. Pressure ........... 53 6-1 RPS Core Protection Safety Limits................. 54 6-2 RPS P-T Core Protection Envelope.................. 55 6-3 High Temperatures MAPS............................ 56 6-4 Low Pressure MAPS................................. 57 6-5 Flux-to-Flow MAPS................................. 58 6-6 Typical Two Pump Coastdown Transient Results...... 59 Iv I

. vi

1.0 INTRODUCTION

Duke Power Company's Oconee Nuclear Station reactor core thermal-hydraulic design and licensing analyses have traditionally used very conservative methods to establish the maximum permissible core power and power distribution for various combinations of core outlet pressure and reactor outlet temper ature to ensure that DNBR criteria are met. Conservative "closed-channel" computer codes have been used for Oconee Nuclear Station thermal-hydraulic analyses using the methodology described in reference 1. Crossflow computer codes which can predict flow redistribution effects within an open lattice reactor core, can more realistically predict the local fluid properties and thus, the departure from nucleate boiling ratio (DNBR) in the hot channels of the core.

This report presents the procedure used to apply the VIPRE-01 computer code for thermal-hydraulic analyses of Oconee reactor cores and fulfills the requirements addressed in the SER for using VIPRE-01 for licensing analyses, ref. 3. The geometric representation of the core is illustrated and discussed along with the models and empirical correlations used to determine friction pressure losses, coolant mixing and subcooled voids. Descriptions of the methodology used to determine the thermal-hydraulic limits which define the regions of safe operation in terms of power level, reactor coolant temperature and pressure, and power distribution are included in this report. The Oconee thermal-hydraulic analyses will continue to treat uncertainties, tolerances, and measurement errors conservatively. The methodology used to perform generic Oconee thermal-hydraulic analyses is discussed in this report.

The need to perform the thermal-hydraulic analyses in conjunction with a reload arises when there is a change in the fuel assembly design, a change in input assumptions of the generic analysis, or a change in the regulatory criteria.

2.0 CODE DESCRIPTION VIPRE-01, ref. 2, is an open channel, homogeneous equilibrium, thermal hydraulic code which features diversion crossflow and turbulent mixing to calculate the departure from nucleate boiling ratios (DNBRs). The code accepts input data which defines the geometric, hydraulic and thermal charac teristics of the core, and permits the user to select correlations and solution methodologies.

Generally, core representation is made by inputting parameters defining and describing the number of channels and subchannels within the model and their individual channel and subchannel characteristics, such as flow area, wetted and heated perimeters, adjacent channel data, and centroidal distances between adjacent channels. Hydraulics of the code are defined by crossflow resistances determined from gap dimensions through which the channels communicate, spacer grid locations and form loss coefficients, mixing coefficients, two-phase flow correlations, friction pressure losses, and inlet flow distributions. Thermal modeling of the reactor core is a function of the core radial and axial power distribution, core power, operating conditions, hot channel factors, heat transfer correlations and correlation limits.

VIPRE-01 was designed to perform steady-state and transient thermal-hydraulic PROPRIETARY CO.

DUKE POWER analyses of nuclear reactor cores for normal operating conditions and several accident conditions. The VIPRE-01 code has been reviewed by the NRC and was found to be acceptable for referencing in licensing applications with the limitations addressed in ref. 3.

3.0 STATION DESCRIPTION Oconee Nuclear Station consists of three, Babcock & Wilcox (B&W) designed, pressurized water reactors with each reactor rated at 2568 Mwt. Each reactor core consists of 177 fuel assemblies with each assembly having 208 fuel rods, 16 control rod guide tubes, and an instrument tube arranged into a 15 x 15 array. Eight Eon-mixing vane spacer grids provide lateral stiffness and fuel rod positioning. Typical dimensions and characteristics of the current Oconee in-reactor fuel assembly designs are given in Table 3-1.

4.0 CORE MODELING Traditionally, core thermal-hydraulic analyses have been performed using multi-pass analyses. In a multi-pass analysis, fuel assemblies and the subchannels of the hot assembly are modeled in separate simulations and sometimes in different computer codes. A more direct approach involves only a single-pass. In a single-pass analysis, the hot subchannel and adjacent subchannels are modeled individually with larger and larger channels modeled toward the periphery of the core; the result is that all thermal-hydraulic DNB calculations can be performed using one code._ VIPRE-01 has the capability to perform single-pass analyses.

L

-3

PROPRIETARY DUKE POWER CO.,

An Oconee Nuclear Station reactor core is geometrically modeled using

[eighth-core symmetry with the center assembly modeled as the "hot" assembly, Figure 4-1. The hot assembly is the assembly in which the minimum DNBR (MDNBR) can be expected to occur. The hot assembly is divided into subchannels with boundaries formed by fuel rods and guide tubes within the assembly, Figure 4-2. The hot assembly contains the "hot" subchannel (i.e.,

the subchannel which yields the MDNBR of the core). To conservatively determine the MDNBR for the core, the models use a[high, relatively flat]

radial pin power distribution along with the application of hot subchannel factors and reduced hot subchannel flow area. The derivation.and application I of these factors will be discussed in more detail in Sections 5.11 and 5.12.

Selection of single-pass models for performing thermal-hydraulic analyses requires the development of different size models and comparisons of the different models at various operating conditions. Three different model sizes were developed and compared for modeling Oconee Nuclear Station fuel:

[64]Channel Model

[9]Channel Model

[8]Channel Model

[All three models were developed assuming eighth-core symmetry. The 64 channel model consists of 36 subchannels making up the hot assembly with the remaining 28 channels individually modeling the rest of the assemblies in the eighth-core segment. The 64 channel model is depicted in Figures 4-1 and 4-2.

The 8 and 9 channel models were formed by including two rows of subchannels]

PROPRIETARY DUKE POWER CO.

around the hot subchannel (Channel 1) and lumping the rest of the hot assembly into one channel (Channel 7), Figure 4-3. The remaining 28 assemblies were either lumped into one large channel, Channel 8, in the case of the 8 channel model, Figure 4-4., or into two large channels, Channels 8 and 9, in the case of the 9 channel model, Figure 4-5. The 64 and 8 channel models were compared to confirm the accuracy of the 8 channel model which will be used for steady-state and two-pump coastdown analyses. The 9 channel model will be used to evaluate potential transition core effects of differing fuel assembly types. As Table 3-1 shows, the different fuel assembly designs currently in-reactor only incorporate minor changes in the basic Mark-BZ fuel assembly design. The use of the 64, 9 and 8 channel models is not limited to the fuel assembly designs listed in Table 3-1; moreover, the 64, 9 and 8 channel models will be used to predict and evaluate the thermal-hydraulic effects of future fuel assembly designs.]

[For illustrative purposes, the number of assemblies lumped together to form Channels 8 and 9 of the 9 channel model, Figure 4-5, was based the on Oconee Unit 1, Cycle 11 core. The number of assemblies lumped together may vary with the cycle specific core configurations being evaluated. The assemblies are arranged in a manner which will give conservative DNBR results.]

To determine the modeling detail required to accurately evaluate the hot channel local coolant conditions and the minimum departure from nucleate boiling ratio (MDNBR), the[64, 9 and 8]channel models were run using the operating conditions stated in Table 4-1. The[RECIRC]numerical solution V

PROPRIETARY DUKE POl^/ER Co.

option was chosen to calculate the results. The VIPRE-01 SER, ref. 3, pg. 17 states that the[RECIRC]numerical solution is acceptable for licensing calculations. The first two operating conditions, Cases 1 and 2, correspond to the high temperature and the low pressure safety limits associated with the Reactor Protection System, ref. 1. -The Case 4 operating conditions correspond to the initial conditions for the two pump coastdown transient. [The Case 3 operating conditions correspond to the operating conditions occuring at the limiting MDNBR during the two pump coastdown transient (i.e., the limiting statepoint in Figure 6-6)] The Case 3 operating conditions are used to develop the maximum allowable pin peaks discussed in Section 6.5. Additional details of the Reactor Protection System will be discussed in Section 6.

4.1 STEADY-STATE SINGLE PASS MODEL COMPARISONS The[64 and 8]channel model results are compared in Table 4-2. Results for the Case 1 operating conditions show that the[8]channel model conservatively predicted the MDNBR by[1.2%]when compared to the[64] channel model MDNBR.

Results for the Case 2 operating conditions showed the[8]channel model exhibited a[O.44%]conservative difference in MDNBR when compared to the[64]

channel model. Likewise, the[8jchannel model exhibited a[2.2%]conservative change in MDNBR for Case 3.

4.2 TRANSIENT MODEL COMPARISONS The two pump coastdown transient is the[most limiting DNB transient therefore, the two pump coastdown transient was chosen to make a comparative study between the[64 and 8] channel models. The development of the transient modeling details are presented in Section 6.6. The transients were performed PROPRIEARY DUKE POWER Co.

using the initial operating conditions from Table 4-1, Case 4. The[64 and 8 channel model transient results are presented in Table 4-3. Throughout the transient, the[8]channel model produced conservative MDNBRs in comparison to the[64lchannel model. The limiting MDNBR observed for the[]channel model occurred at[4.1Jseconds where the MDNBR =[1.216](i.e., which is conservative in comparison to the[64]channel model MDNBR of[1.23 ]also occurring at[4.J seconds).

4.3 TRANSITION CORE MODEL COMPARISONS As mentioned earlier in Section 1.0, a thermal-hydraulic analysis must be performed whenever there is a change in the fuel assembly design, a change in input assumptions of the generic analysis, or a change in the regulatory criteria. Combinations of different fuel assembly designs in a reactor core constitutes a mixed (transition) core which must be evaluated to determine its effect on thermal-hydraulic performance. [Transition core effects are determined by comparing results of a thermal-hydraulic (T-H) analysis explicitly modeling the mixed core with that of a T-H analysis for a non-mixed core. If the comparison shows the MDNBR is adversely affected, then a penalty must be assigned to that particular operating cycle]

The[64]channel and[9]channel transition core models were compared on a steady-state and transient basis to ascertain the accuracy of the[9]channel model. Table 4-4 presents a comparison of the[64]and[9]channel models steady-state results. In all cases, the[9]channel model produced conservative results. The steady-state runs using the Case 3 operating conditions produced the lowest MDNBRs, with the[9]channel model predicting a MDNBR[1.9]more conservative than the[64]channel model. A comparison of theL64]channel and[9]

t PROPRIETARY DUKE POWER CO.

channel model two pump coastdown transient results using the Case 4 initial operating conditions revealed that the[9]channel model again produced conservative MDNBRs (see Table 4-5); therefore, the[9]channel model will be used to assess any future Oconee Nuclear Station reloads involving transition cores.

4.4 RESULTS

SUMMARY

In all of the studies and comparisons performed between the[64, 9, and 8]

channel models, the[9 and 8 channel models consistently produced conservative results. Duke Power Company will use the smaller channel models to perform thermal-hydraulic analyses since the[64]channel model requires an extensive amount of computer processing time. [The 64 channel model will be used if a situation arises which requires the 1-2% conservatism currently available with the 8 and 9 channel models. The larger model would only be used for cycle specific evaluations requiring the additional margin]

5.0 VIPRE-01 DATA The fuel assembly data used to develop each of the input parameters, such as flow area, wetted and heated perimeters, centroid distances, and gap widths are given in Table 3-1. Other important VIPRE-01 input is discussed in detail in the subsections which follow.

II

PROPRIETARY DUKE POWER CO.

5.1 AXIAL NODING Given the axial power shape and a specified heated rod length, VIPRE-01 determines the axial power factor for each axial node, ref. 2. The node length determines how well the code approximates the axial power shape, the shorter the node length, the better the approximation of the curve. Volume 4 of the VIPRE-01 manual states as a general rule that nodes on the order of 2 or 3 inches long are recommended in the region where MDNBR is likely to occur, ref.2. Calculations involving node sizes smaller than 2 or 3 inches require more computer processing time without gaining significant increases in the accuracy of the results.

Results of an axial node length sensitivity study performed with the[8]channel steady-state model are presented in Table 5-1. A comparison was made[between a three-inch node length, uniformly applied to the axial length of the rod from 4.125 to 142.125 inches, and two ranges of two-inch axial node lengths applied to the rod at elevations ranging from 32.125 to 94.125 inches and 81.125 to 143.125 inches] As Table 5-1 shows, the bree-inch node lengths produced slightly conservative MDNBRs; therefore, the[three-inc node length will be used for all Oconee Nuclear Station thermal-hydraulic analyses.

5.2 ACTIVE FUEL LENGTH

[Uranium fuel both densifies and swells when irradiated. Densification effects are predominant at low burnup and swelling effects are predominant at higher burnup. Fuel densification decreases the active fuel length while fuel swelling tends to increase the active length. For low densification fuel as used in Oconee cores, the active fuel length at reactor conditions is greater]

PROPRIETARY DUKE POWER CO.

[than the nominal cold stack height. Since the fuel thermally expands in the reactor more than it shrinks due to densification, it is conservative to use the cold nominal active fuel length]

5.3 CENTROID DISTANCE The location of each subchannel or channel is defined by numbering all the channels, inputting connecting channel numbers, and defining the distance between centroids of adjacent channels. The centroidal distance in a normal square array, is the subchannel pitch. The centroidal distance determines the length over which the crossflow exists and defines the lateral pressure gradient in the crossflow momentum equation. The centroidal distance for a channel cut by a line of symmetry is the same as the centroidal distance for the complete channel, ref. 2. For the lumped subchannels, the centroidal distance is increased from its individual subchannel value in proportion to the number of rod rows between channel centroids. Likewise the centroidal distances between lumped assemblies is increased in proportion to the rows of assemblies between the lumped channel centroids.

5.4 EFFECTIVE CROSSFLOW GAPS g Crossflow resistances are calculated by inputting connecting channel information and crossflow gap widths. The product of the gap width and the axial node length defines the lateral flow area between channels. The gap widths are easily calculated given the rod pitches and diameters. The gap width for a fuel assembly or any lumped channel is the sum of the subchannel gaps through which the two assemblies communicate.

8 PROPRIETARY 5.5 SPACER GRID FORM LOSS COEFFICIENTS DUKE POWER CO.

Form loss coefficients are used-to account for the unrecoverable pressure losses caused by the abrupt variation in flow area and turbulence at a spacer grid. The Mark-BZ fuel assemblies have six ihtermediate zircaloy spacer grids and two inconel end grids. Form loss coefficients determined for the different types of subchannels (i.e. unit, thimble tube, peripheral, instrument guide tube, and corner channels) and for the overall grid are used in the thermal-hydraulic analyses. [Spacer grid form loss coefficients are developed from full size fuel assembly flow tests performed by the vendor.

Individual subchannel form loss coefficients are determined analytically by the vendor from the overall grid form loss coefficients]

5.6 CORE BYPASS FLOW Core flow is equal to the total reactor coolant system flow less the bypass flow, which is defined as that part of the flow which does not contact the effective heat transfer surface area. The bypass flow paths are the 1) core shroud, 2) core barrel annulus, 3) control rod guide tubes and instrument tubes, and 4) all interfaces separating the inlet and outlet regions of the reactor vessel. A typical value of the design bypass flow is[9.0%] however, the bypass flow rate is dependent on the number of control rod and burnable poison rod assemblies in the core since they act as guide tube plugging devices. The actual core bypass flow must be verified each cycle to assure that it is less than that used in the generic analysis.

II

PROPRIETARY 5.7 INLET FLOW DISTRIBUTION DUKE POWER CO.

VIPRE-01 allows the user to specify the core inlet flow maldistribution. The Oconee core thermal-hydraulic analyses include a[5% reduction in inlet flow to the hot assembl to conservatively represent the results obtained in B&W's 1/6-scale Vessel Model Flow Test, ref 1. More restrictive flow maldistri bution factors are used for operation with less than four reactor coolant pumps. Table 5-2 shows that the use of a[5] inlet flow maldistribution produces slightly conservative results compared with a uniform inlet flow distribution.

5.8 VIPRE-01 CORRELATIONS Empirical correlations are used in the VIPRE-01 code to model turbulent mixing and the effects of two-phase flow on friction pressure losses, non-equilibrium subcooled boiling, and the relationship between the quality and void fraction.

The correlations which have been selected for use in the Oconee thermal-hydraulic analyses are discussed in the subsections which follow.

5.8.1 FRICTION PRESSURE LOSS Pressure losses due to frictional drag are calculated for flow in both the 1 axial and lateral directions. In the axial direction the friction pressure loss is calculated by dP = f G2 V' dZ 2gc Dh 3

3 PROPRIETARY DUKE POWER CO.

where f = friction factor determined from an empirical correlation defined by user input G = Mass flux, lbm/sec-ft 2 v'= specific volume for momentum, ft. 3/lbm g = force-to-mass units conversion factor, 32.2 1bm-ft/lbf-sec 2 Dh= hydraulic diameter based on wetted perimeter, ft.

[Based on the recommendation in ref. 2, vol. 4 of the VIPRE-01 manual, the default Blasius smooth tube friction factor expression f = 0.32 Re- 0.25 + 0.0 will be used to calculate the friction pressure loss for turbulent flow.]

Based on sensitivity study results given in Table 5-3, the friction pressure loss for two-phase flow will be calculated using the[EPRI two-phase friction]

multiplier.

In the lateral direction the pressure loss is treated as a form drag loss that iscalculated by A\P = KG w'w v' 2S2gc where KG = loss coefficient in the gap between adjacent channels w = crossflow through a gap, lbm/sec-ft v' = specific volume for momentum, ft 3/lbm S = gap width, ft gc = force-to-mass units conversion factor, 32.2 1bm-ft lbf-sec 2 PROPRIETARY DUKE POWER CO.

When rod arrays are modeled as lumped channels the effective crossflow resis tance is the sum of the resistance of the rod rows between the lumped channel centroids. The lateral loss coefficient becomes K. = N KG where N is the number of rod rows between lumped channels and KG is the nominal drag coefficient for a single gap. Crossflow resistance coefficients are not precisely known, but sensitivity study results discussed in Volume 4 of ref. 2 show that for applications where the axial flow is predominant relative to crossflow, crossflow resistance has an insignificant effect on mass flux and DNBR. A subchannel drag coefficient, K , of[0.5]will be used with the coefficient for lumped channels calculated internally by the code based on the input centroid distances between lumped channels and the standard subchannel fuel rod pitch.

5.8.2 TURBULENT MIXING The VIPRE-01 transverse momentum equation includes terms to calculate the exchange of momentum between adjacent channels due to turbulent mixing. Two parameters must be input to include turbulent mixing: a turbulent momentum factor (FTM) and a turbulent mixing coefficient (B).

The turbulent momentum factor (FTM) defines how efficiently the turbulent crossflow mixes momentum. FTM can be input on a scale from 0.0 to 1.0, where 0.0 indicates that the crossflow mixes enthalpy only and 1.0 indicates that it mixes enthalpy and momentum with the same strength. In actuality, some proportion of enthalpy and momentum mixing does take place; therefore, PROPRIETARY DUKE POWER CO.

turbulent momentum factors of 0.8 and 1.0 are probably more representative of actual crossflow effects. Sensitivity studies discussed in Vol. 4 of ref. 2 show that changes in the fraction of momentum mixing have negligible impact on the flow field; therefore, FTM = 0.8 is recommended, ref. 2. Sensitivity studies using theE8]channel model were performed for the Case I and 2 operating conditions given in Table 4-1. The runs were made using FTM = 0.0, 0.8 and 1.0. The results of the analyses are presented in Table 5-4.[Since the results show that MDNBRs for an FTM = 0.8 lie between the MDNBRs for FTM =

0.0 and 1.0, and since FTM = 0.8 more realistically assumes some momentum mixing, an FTM = 0.8 will be used in all future Oconee thermal-hydraulic analyses]

Turbulent crossflow between adjacent channels is calculated by I w'=8SG where w' is the turbulent flow per axial length, B is the turbulent mixing coefficient, S is the gap width, and G is the average mass flux of the adjacent channels. [Based upon vendor predictions of mixing test results, a mixing coefficient of 0.01 will be used for all Oconee Nuclear Station core thermal-hydraulic analyses.]

I 5.8.3 TWO-PHASE FLOW CORRELATIONS I

Two correlations are used in VIPRE-01 to make two-phase flow predictions. The first correlation is referred to as the subcooled void correlation. It uses a g PROPRIETARY DUKE POWER CO.

quality model to calculate the flowing vapor mass fraction including the effects of subcooled boiling. Once the flowing vapor mass fraction is calcu lated, the bulk void correlation is applied to calculate the void fraction including any effects due to slip, ref. 2, Vol. 1.

Sensitivity studies were performed using three different combinations of subcooled void and bulk void correlations to evaluate their effects on the hot channel local coolant conditions and MDNBR.

Subcooled Void Bulk Void Levy Zuber-Findlay Levy Smith EPRI EPRI The hot channel local coolant conditions and MDNBRs are given in Table 5-3 for the Case 1 and 2 operating conditions. As Table 5-3 shows, the combination of the[Levy subcooled void correlation and the Zuber-Findlay]bulk void correlation yields slightly conservative results. Section 3.3 of Vol. 4 of the VIPRE-01 manual, ref. 2, presents the results of VIPRE-01 predictions of the Martin void fraction tests at high pressure (1565 and 1991 psia).[Of the two-phase flow correlations evaluated, the Levy/Zuber-Findlay combination I compared most favorably with the test results. The Levy subcooled void correlation and the Zuber-Findlay bulk void correlation will be used for Oconee thermal-hydraulic analyses.]

__________________PROPRIETARY(

POR 5.9 REFERENCE DESIGN POWER DISTRIBUTION DUKE POWER Co.

Co.

The reference design power distributions are shown in Figures 4-1 through 4-5.

The power distributions were designed to be conservatively[high and relatively flat]in the vicinity of the hot subchannel. The pin power peaking gradient within the area of the hot subchannel is approximately[1%] The pin power distribution was verified to be conservative by comparison with predicted physics power distributions. The reference design power distribution was developed using a radial-local hot pin peak, FNH, of [1.714]

and an assembly power of L1. 6 147 ] The FNH = L1 .7 14 Jis the same reference pin peak used in the methodology discussed in reference 1.

5.10 AXIAL POWER DISTRIBUTION I The axial power shape, used to develop the results presented in this report, was a[1.65 chopped cosine axial power shape. Predicted and actual axial power shapes vary for cycle specific reloads and transients since they are functions of control rod positions, xenon transients, etc. The effect on DNB of different axial flux shapes is taken into account as discussed in Section 6.5.

EA routine has been added to the VIPRE-01 code to generate axial power shapes with inlet, symmetric, or outlet peaks. The routine is based on the following constraints on an axial power shape]

3 PROPRIETARY DUKE POWER CO.,

F(x) is continuous from ME)

F'(x) is continuous from (B,E)

E -B F(x) dx = 1.0 B

where F(x) = axial power shape as a function of the axial location, x B,E = beginning and ending normalized location of the active fuel length II The reference 1.65 axial flux shape is generated using the new axial shape routine.

5.11 HOT CHANNEL FACTOR The local heat flux factor, F ", and the power factor, Fq are conservatively applied to the hot subchannel[(i.e., the instrument guide tube subchannel)]of the hot assembly to compensate for possible deviations of several parameters from their design values. The local heat flux factor F " = 1.0137, ref. 1.,

incorporates variations in pellet density, pellet cross-sectional area, weight per unit length, local enrichment, and local clad outer diameter. [This factor is only used in the computation of the surface heat flux of the hot pin when calculating the DNBR for the hot subchannel, ref. 1 The power factor, F =

1.0107] ref. 1., accounts for variations in average pin power caused by differences in the absolute number of grams of U-235 per rod. The loading tolerance on U-235 per fuel stack and variation on the powder lot mean 3 PROPRIETARY DUKE POWER CO.

II enrichment are considered in determining the factor, ref. 1.[Fq is applied to the heat generation rate of the hot pin of the hot subchannel. Both factors are applied to the hot subchannel during steady-state and transient thermal hydraulic analyses. However, in the determination of maximum allowable peaking limits (as will be discussed in Section 6.5), the local heat flux factor, F ", is increased by applying two additional penalities] [First, a 1.007 penalty is applied to account for power spikes occurring as a result of the flux depressions at the spacer grids. Secondly, a 1.016 penalty is applied to account for axial nuclear uncertainty, ref. 1. The Fq" for calculating maximum allowable peaking limits for Mark-BZ fuel is 1.0137 x 1.007 x 1.016 = 1.0371, ref.1]

5.12 FLOW AREA REDUCTION FACTOR lI The hot subchannel flow area is reduced by 3%jto account for variations in I as-built subchannel coolant flow areas.

5.13 BWC CRITICAL HEAT FLUX CORRELATION I

The BWC critical heat flux (CHF) correlation, ref. 4, will be used for Oconee thermal-hydraulic analyses. The BWC correlation was originally developed for B&W 17 x 17 Mark-C fuel. Subsequently, as discussed in ref. 4, B&W showed that the BWC correlation can be used for 15 x 15 Zircaloy grid Mark-BZ fuel.

The BWC correlation was developed by B&W using the LYNX2 crossflow computer code, ref. 5. To justify use of the BWC correlation with the VIPRE-01 code the Zircaloy grid CHF test results given in ref. 4 were predicted using PROPRIETARY DUKE POWER CO.,

3 VIPRE-01 and compared with B&W's LYNX2 results. The VIPRE-O1/BWC results for all[211]data points were used to determine a DNBR limit which provides a 95%

probability of precluding DNB at a 95% confidence level.

I Figures 5-1, 5-2, and 5-3 show the B&W LYNX2 versus VIPRE-01 calculations for II the Measured-to-Predicted (M/P) CHF, mass velocity, and quality at the CHF 3 location, respectively. These figures show that the VIPRE-01 coolant conditions and BWC CHF predictions are essentially the same as B&W's LYNX2 predictions. Figure 5-4 shows the measured CHF versus the VIPRE-01 predicted CHF for all[211]data points demonstrating that the overall prediction of the correlation is correct. The ratio of measured-to-predicted CHF is plotted versus local quality, mass velocity, and pressure in Figures 5-5 through 5-7, respectively. These figures show that there is no bias in the correlation relative to the important fluid parameters. Calculation of the design DNBR limit is based on the assumption that the M/P CHF values are normally distributed. This was verified statistically using the D-prime test.

A DNBR limit is calculated so that cores can be designed to operate below the CHF. The DNBR limit is the lowest DNBR that can be calculated (for any core condition) for the limiting pins in the core and ensure with 95% confidence that 95% of the limiting pins are not in film boiling. The design DNBR limit was calculated using the following expression developed in ref. 4:

I I

I 5

PROPRIETARY 3.

1.0 DUKE POWER CO, DNBR Limit M/P - K N,6,P where M/P = mean measured-to-predicted CHF ratio KN,,P = one-sided tolerance factor based on Ii degrees of freedom (N), confidence level (6) and portion of population protected (P).

a = standard deviation of measured-to-predicted CHF values For the VIPRE-01/BWC combination the design DNBR limit is 1.161] The DNBR limit is calculated as shown in the following.

II N = 211 I M/P = 1.0076 K211 , 0.95, 0.95Y = 1.832 a = 0.0797 1.0 DNBR Limit = 1.0076 - 1.832 (0.0797) = 1.161 For all Oconee thermal-hydraulic analyses using VIPRE-01 and the BWC correlation, a design DNBR limit of 1.161 + margin will be used.

The applicable range of variables for the BWC correlation are:

Pressure 1600 < P < 2600 psia 5 Mass Velocity 0.43 < G < 3.8 Mlbm/hr-ft 2 Quality -0.20 < X < + 0.26 I

I I 6.0 OCONEE THERMAL-HYDRAULIC ANALYSES I 6.1

SUMMARY

I A thermal-hydraulic analysis of the Oconee reactor cores is necessary to I define the core thermal margin and acceptable operating limits. The crossflow

• code thermal-hydraulic analysis methods used to derive the core safety and operating limits are the same as the previously approved methods in ref. 1.

The safety and operating limits are used to ensure core protection against anticipated transients and steady-state operation. Some of the Reactor Protection System (RPS) trip functions incorporate these safety limits as I setpoints which would trip the reactor prior to exceeding the thermal design limits. A list of RPS trip functions is given in Table 6-1. The safety limits are derived from generic thermal-hydraulic analyses based upon various combinations of power, pressure, temperature and flux-to-flow limits. A new analysis is performed for a reload core whenever there is a significant change in the fuel assembly or core design, a change in the input assumptions of the generic analysis, or a change in the regulatory criteria.

I 6-2 THERMAL-HYDRAULIC DESIGN CRITERION I

The thermal-hydraulic design criterion is that no core damage due to DNB occur 1 during steady-state operation or anticipated transients. DNB is defined as the point where bubble generation on the clad heat transfer surface forms an insulating blanket over .the surface heating area, thus, causing a large clad surface temperature rise. The departure from nucleate boiling ratio (DNBR) is defined as the ratio of the critical heat flux at a point on the rod to the I actual heat flux at the same point. DNBR is calculated using Babcock 3

PROPRIETARY I DUKE POWER CO.

and Wilcox's BWC correlation. The minimum DNBR (MDNBR) is limited to 1.161 +

margin] as previously explained in Section 5.13.

6.3 CORE SAFETY LIMITS Core safety limits are determined to protect the core during steady-state I operation and anticipated transients. The core safety limits prevent overheating and possible rupture of the cladding which would release fission products to the coolant. Fuel clad overheating is prevented by restricting operation to within the nucleate boiling regime where clad temperature is only slightly above the coolant temperature. Two core safety limits directly provide DNB protection:

1. Pressure - Temperature Envelope

2. Power - Power Imbalance Limits II 6.4 PRESSURE-TEMPERATURE ENVELOPE The Pressure-Temperature (P-T) envelope defines a region of allowable operation in terms of reactor coolant system (RCS) pressure and vessel outlet temperature. The P-T envelope provides DNB protection as well as protection for the RCS. The three reactor trips that define the region of allowable operation as shown in Fig. 6-3 are:

II

1. High temperature trip

2. Low pressure trip

3. Variable low pressure trip I

3 PROPRIETARY DUKE POWER CO.

I To ensure that the P-T envelope provides DNB protection, P-T curves are determined for[4, 3, and 2]reactor coolant (RC) pump operation. The P-T curves are the combinations ofRCS pressure and vessel outlet temperature that yield the design DNBR limit[(BWC correlation limit plus marginJ or the BWC correlation quality limit. The P-T envelope must be more restrictive than the I most limiting P-T curve as shown in Fig. 6-3.

The P-T curves are calculated using the[8]channel model discussed in Section 4.0. The VIPRE-01 input that is used to calculate the generic P-T curves is discussed in subsections 6.4.1 through 6.4.4 which follow.

II 3 6.4.1 REFERENCE POWER DISTRIBUTION The reference power distribution discussed in Section 5.9 and shown in Fig.

4-3 and 4-4 is used to calculate the P-T curves. The reference axial power profile used to calculate the P-T curves is a symmetric chopped cosine with a peak to average value of 1.65j The axial power shape can change as a result of rod motion, power change, or due to a xenon transient. Power - power imbalance limits, ref. 1, provide protection for the core from the effects of skewed axial power distributions. To determine the power-power imbalance limits maximum allowable peaking (MAP) limits are calculated as discussed in Section 6.5.

6.4.2 CORE POWER The maximum power level for 4 pump operation, 112% FPJ is set by the high flux trip setpoint with adjustment made for uncertainties and margin. [The maximum]

3 PROPRIETARY DUKE POWER CO.

Ipower level for 3 and 2 pump operation are based on the flux/flow trip setpoint including the appropriate power and flow uncertainties] The PT curves are calculated for the maximum power levels forF4, 3, and 2 pump operation]

6.4.3 RCS FLOW The generic Oconee thermal-hydraulic analyses will be based on an RCS flow of

[366,080] gpm [(104%]of the design flow of [88,000] gpm/pump) which is lower than the measured flow for any of the three Oconee units. This value could be increased for a cycle specific analysis to take credit for the flow margin at I a particular unit.

6.4.4 CORE INLET TEMPERATURE For a given core power, flow (number of operating RC pumps), and pressure the vessel outlet temperature at which the MDNBR equals the design DNBR limit defines a point along a PT curve. VIPRE-01 is run at several pressures to determine the core inlet temperatures that yield the design DNBR limit.

6.5 GENERIC MAXIMUM ALLOWABLE PEAKING LIMIT CURVES I In order to provide DNB protection for axially assymetric and symmetric power distributions, a series of maximum allowable pin peaks are calculated such that the MDNBR limit is obtained. [Maximum allowable peaking (MAP) limits are calculated in the form of lines of constant MDNBR for a range of axial peaks with the location of the peak varied from the bottom to the top of the core.

This is performed for axial peaks of 1.1 to 1.9 for distances up the channel, 3

PROPRIETARY DUKE POWER CO.

[X/L, of 0.1 to 0.9. The axial peaks were generated using the new axial shape routine discussed in Section 5.10. The maximum allowable peaks are multiplied by their respective axial peaks to obtain Total Maximum Allowable Peaking Limits (i.e., MAP limits). The MAP Limits are plotted for each axial peak and X/L to form a set of MAP limit curves. The MAP limits provide a basis for equating the symmetric and asymmetric power distributions. MAP limits are I compared in a maneuvering analysis with peaks resulting from design power transients as discussed in ref. 1. Two sets of generic MAP Limit curves are determined. One set is used to determine the DNB operational offset limits, and the other set is used to determine the Reactor Protection System (RPS) DNB offset limits. The RPS MAP Limit curves.are developed using the extreme conditions of the P-T core protection envelope, Figure 6-2. The extreme conditions are the high temperature and low pressure safety limits based on the limiting P-T curve as shown in Figure 6-2 and defined in Table 4-1, Cases 1 and 2. A typical set of RPS MAP Limit curves generated with the 8 channel model are shown in Figures 6-3 and 6-4. The final set of generic RPS MAP limit curves ia a conservative overlay of both the high temperature and the low pressure RPS MAP Limit curves I

[Operational MAP Limit curves are developed in the same manner as the RPS MAP limits based on the two pump coastdown transient as explained in the following section. A typical set of Operational MAP limit curves generated with the 8 channel model is shown in Figure 6-5.

I To verify that the MAP limits are acceptable for a specific reload design, the MDNBR is determined for the most limiting predicted power distribution. The predicted radial power distribution and axial flux shape is input directly WI into the VIPRE-01 code.

PROPRIETARY DUKE PO W ER CO.

I _ _ _ _ _ _ _ _ _ _

6.6 PUMP COASTDOWN TRANSIENT ANALYSES I

The flux/flow trip prevents the core from violating the DNBR criterion during a loss of one or more reactor coolant (RC) pumps. DNB protection is also provided by pump monitors which provide an immediate trip signal on loss of electrical power to the pump motors. The pump monitors at Oconee will trip the reactor[for a loss of two or more reactor coolant pumps from above 55%

power. Thus, it is conservative to determine the flux/flow trip setpoint assuming the loss of two RC pumps]

The two pump coastdown transient is analyzed using VIPRE-01 to assure that the

[1.161 + margin] design DNBR limit is not violated after the loss of one or more RC pumps. The VIPRE-01,[8]channel model was used with rods1-4] (see Fig. 4-3) modeled using the conduction model available in the VIPRE-01 code. For steady-state analyses "dummy" rods are used with the power (heat flux) applied directly to the coolant. During a transient; however, once the reactor is tripped, the neutron power generated in the fuel decreases rapidly, but the thermal power reaches the coolant with some time delay through conduction and convection from each fuel rod. To model the conduction and stored energy effects the VIPRE-01 conduction model is used.

I I Conduction through the gap between the fuel pellet and the clad is determined using the gap conductance model in VIPRE-01. The NRC concluded in the VIPRE-01 SER, ref. 3, that the fuel rod heat conduction model is acceptable for licensing analyses.

U To select the input for the conduction model sensitivity studies were performed varying the following input parameters:

PROPRIETARY DUKE POWER CO.

Pellet/clad gap Gas composition I Pellet radial power profile Sensitivity studies discussed in the VIPRE-01 manual, ref. 2, show that the I gap conductance model is most sensitive to the specified gap width. Maximum and minimum gaps, based on predicted pellet densification and clad creepdown, were evaluated. Studies were also performed to determine the effect of the fill gas composition and fission gases released into the gap on the transient DNBR results. Cases were also run to determine the effect that the pellet radial power profile has on the transient DNBR results.

Based on the sensitivity study results, the generic VIPRE-01 two-pump coastdown analyses will be performed using the following conservative conduction model input:

Nominal Rod OD Nominal Clad Thickness Maximum Pellet/Clad Gap Minimum Pellet Diameter Minimum Prepressure Nominal Plenum Volume I Helium and Nitrogen Fill Gas

• Uniform Pellet Radial Power Profile Heat transfer correlations are used by the VIPRE-01 code only when the conduction model is specified. Convection and nucleate boiling correlations are selected since only conditions up to the point of DNB are normally PROPRIETARY DUKE POWER CO.

analyzed. [The default single-phase forced convection correlation, the Dittus-Boelter correlation with the leading coefficient compatible with the I' EPRI void model, will be used for Oconee pump coastdown analyses. The Thom subcooled and saturated nucleate boiling correlation will be used.] A sensitivity study showed that the choice of nucleate boiling correlations made very little difference in the pump coastdown DNBR results.

I Generic pump coastdown transient analyses are performed for each unit to verify that the flux/flow trip setpoint provides DNB protection for the loss of one or more RC pumps. The flux/flow trip setpoint also provides overpower protection for three and two pump steady-state operation. The generic analyses are performed using the reference power distribution (FNH =1.714])

shown in Fig. 4-3 and 4-4 along witht the[1.65 cosine] reference axial power shape. To ensure that the DNBR criterion is met during a pump coastdown transient with any possible axial flux shape, Operational MAP limits are calculated as previously discussed in Section 6.5. [The Operational MAP limits are based on the limiting two pump coastdown statepoint (power, pressure, temperature, etc.) as shown in Fig. 6-6. The operational MAP limits are used to determine operational power-power imbalance limits as discussed in ref. 1]

I I

I I

I 1

7.0 REFERENCES

1. Oconee Nuclear Station Reload Design Methodology II, DPC-NE-1002, Duke Power Company, Charlotte, NC, March 1985.

2. J. M. Cuta, et. al., "VIPRE-01: A Thermal-Hydraulic Code for Reactor Cores," EPRI-NP-2511-CCM, Vol. 1-5, Battelle Pacific Northwest Laboratories, July 1985.

I

3. Letter from C. E. Rossi (NRC) to J. A. Blaisdell (UGRA), "Acceptance for Referencing of Licensing Topical Report, VIPRE-01: A Thermal Hydraulic Analysis Code for Reactor Cores," EPRI-NP-2511-CCM, Vol. 1-5, May 1, II 1986.

I 4. BWC Correlation of Critical Heat Flux, BAW-10143-A, Babcock and Wilcox, Lynchburg, VA, April 1985.

I 5. LYNX2: Subchannel Thermal-Hydraulic Analysis Program, BAW-10130-A, Babcock and Wilcox, Lynchburg, VA, July 1985.

I I

I I

PROPRIETARY DUKE POWER CO.

TABLE 3-1. MARK-BZ FUEL ASSEMBLY DATA (TYPICAL)

GENERAL FUEL SPECIFICATIONS Fuel rod diameter, in. (Nom.) 0.430 Thimble tube diameter, in. (Nom.) 0.530 Instrument tube diameter, in. (Nom.) 0.493 Fuel rod pitch, in. (Nom.) 0.579 Fuel assembly pitch, in. (Nom.) 8.587 Fuel rod length, in. (Nom.) 153.7 GENERAL FUEL CHARACTERISTICS Grids: Material Quantity Location Type Inconel 2 Upper and Lower Non-mixin I Zircaloy 6Intermediate Non-mixing I'Fuel rods: Materi.al Quantity[Vn Zircaloy-4 208

• Fuel Cycle Design Assembly Features Fuel Assy. Mark Designation: B4Z Mark B5Z Mark B6 0.3 Mark B7 IIIFeatures: Requires Mechanical Same as Mark B4Z Same as Mark B5Z except Recon-Same as Mark B6 except Thinner Iretain Retainer to BPRA except Anti-straddle stitutable Upper End Fitting Lower End Fitting and Longer Fuel Rod Lower End Fitting and no separate mechani cal retai ner for BPRA

• -31

PROPRIETARY TABLE 4-1. OPERATING CONDITIONS DUKE POWER CO.

Inlet CASE* Power Flow Pressure Temperature I_%  % PSIA OF 1 112 104 2085 568.6 I 2 112 104 1800 539.8 3 103.3 78.2 2135 555.3 4 108 104 2135 555.3

• All cases were performed using a[1.65]axial peak unless otherwise noted.

I I

I I

I I

I I

I I

• -32 I

TABLE 4-2. COMPARISON OFE64ICHANNEL ANDE8]CHANNEL MODEL STEADY-STATE RESULTS (TYPICAL)

MDNBR MASS VELOCITY EXIT QUALITY (MLBM/HR-FT2 CASE la Ch. 1 Ch. 2 Ch. 3 Ch.1 Ch.2 Ch.3 Ch. 1 Ch. 2 Ch.3

[64] Channel Model 1.412 1.490 1.569 1.86 2.06 2.03 0.121 0.114 0.1021 C8Channel Model L1.395 1.476 1.558 1.80 2.00 1.99 0.124 0.115 0.102]

CASE 2a

[64]Channel Model 1.601 1.674 1.750 1.93 2.15 2.19 0.087 0.081 0.070 L8]Channel Model 1.594 1.670 1.749 1.90 2.13 2.17 0.088 0.081 0.069]

CASE 3 a

[64]Channel Model 1.243 1.338 1.454 1.37 1.54 1.53 0.163 0.154 0.138

[8]Channel Model 11.216 1.315 1.433 1.32 1.48 1.47 0.166 0.156 0.138 a) Cases 1, 2, 3 are in reference to the operating conditions given in Table 4-1.

o:OO? t,

PROPRIETARY DUKE POWER CO, TABLE 4-3. COMPARISON OF[64]CHANNEL AND[8]CHANNEL MODEL TRANSIENT RESULTS (TYPICAL) 1 Time 64 Channel Model

Channel 1 8 Channel Model Channel 1 (sec) MDNBR MDNBR 0.0 1.833 1.830 0.5 1.813 1.807 1.0 1.774 1.767 1.5 1.715 1.708 2.0 1.645 1.636 2.5 1.569 1.558 2.7 1.528 1.517 2.9 1.501 1.489 3.1 1.471 1.454 3.3 1.436 1.420 3.4 1.396 1.376 3.5 1.356 1.333 3.6 1.331 1.308 3.7 1.306 1.285 3.8 1.280 1.260 3.9 1.261 1.238 4.0 1.241 1.221 4.1 1.234 1.216 4.2 1.245 1.221 5 4.3 1.281 1.250 I

I I

I II -34 I

TABLE 4-4. COMPARISON OFE64 CHANNEL AND[9]CHANNEL TRANSITION CORE MODEL STEADY-STATE RESULTS (TYPICAL)

MDNBR MASS VELOCITY EXIT QUALITY (MLBM/HR-FT2 )

CASE 1a Ch. 1 Ch. 2 Ch. 3 Ch.1 Ch.2 Ch.3 Ch. 1 Ch. 2 Ch.3

[64lChannel Model 1.412 1.490 1.568 1.86 2.06 2.03 0.121 0.114 0.102

[9]Channel Model L1.398 1.479 1.560 1.80 2.00 1.99 0.123 0.115 0.102]

CASE 2a

[64 Channel Model 1.600 1.674 1.750 1.93 2.15 2.19 0.087 0.081 0.0701

[9]Channel Model L1.597 1.673 1.751 1.90 2.13 2.18 0.087 0.081 0.0691 CASE 3a

[64]Channel Model 1.243 1.338 1.454 1.37 1.54 1.53 0.163 0.154 0.138 L9 ]Channel Model L1.220 1.318 1.436 1.32 1.48 1.47 0.165 0.155 0.138]

a) Denotes Cases 1, 2 and 3 are operating conditions from Table 4-1.

PROPRIETARY DUKE POWER CO.

TABLE 4-5. COMPARISONS OF[64] CHANNEL ANDL9]CHANNEL TRANSITION CORE MODEL TRANSIENT RESULTS (TYPICAL)

[64IChannel Model [9]Channel Model Time Channel 1 Channel 1 (sec) MDNBR MDNBR 0.0 1.834 1.831 0.5 1.814 1.808 1.0 1.775 1.769 1.5 1.715 1.710 2.0 1.646 1.639 2.5 1.571 1.562 2.7 1.529 1.519 2.9 1.502 1.491 3.1 1.472 1.456 3.3 1.438 1.422 3.4 1.396 1.379 3.5 1.359 1.337 3.6 1.333 1.301 3.7 1.309 1.283 3.8 1.283 1.262 3.9 1.263 1.243 4.0 1.244 1.229 4.1 1.238 1.224 4.2 1.249 1.230 4.3 ----- 1.256 4.4 ----- 1.302 I

I I

I

• -36 I

TABLE 5-1. [81CHANNEL MODEL AXIAL NODE LENGTH SENSITIVITY STUDY (TYPICAL)

Node Operating Axial X Size Node Study Channel 1 MDNBR @

Conditions Peak L (in.) Elevation (in.) MDNBR Elevation (in.)

(CASE 1) 1.65 0.5 3 4 .125-142 .12 5b 1.395 97.1-100.1 L1.65 0.5 2 81.125-143.125 1.398 99.1-101.11 (CASE 2) 1.65 0.5 3 4.125-142.125b 1.594 94.1-97.1 1.65 0.5 2 81 .125 -143 .125 1.598 95.1-97.1 (CASE 1) 1.70 0.1 3 4.125-142.125 1.731 64.1-67.1 L1.70 0.1 2 32 .125- 94 .125 1.736 66.1-68.1 (CASE 2) 1.70 0.1 3 4.125-142.125 1.932 58.1-61.1 L1.70 0.1 2 32.125-94.125 1.941 58.1-60.1 Notes a) Operating conditions from Table 4-1.

b) [4.125-81.125 in. range modeled with three-inch nodes.]

c) L4.125-32.125 and 94.125-143.125 in. ranges modeled with three-inch nodes.]

-0 0i

g PROPRIETARY DUKE POWER CO.

TABLE 5-2. [1CHANNEL MODEL INLET FLOW I DISTRIBUTION SENSITIVITY STUDY (TYPICAL)

Operating Condition Case 1 from Table 4-1.

U Percent Flow MDNBR to Hot Assy. Channel 1 Channel 2

[W Channel 3 95 1.395 1.476 1.558 100 1.409 1.492 1.572]

MASS VELOCITY (MLBM/HR-FT2) 95 1.8042 1.9977 1.9874 100 1.8148 2.0127 2.00351 3 EXIT QUALITY 95 0.1238 0.1154 0.1016 L100 0.1201 0.1118 0.0981]

U I

II I

• -38 I

TABLE 5-3. [8]CHANNEL MODEL TWO-PHASE FLOW CORRELATION AND FRICTION MULTIPLIER SENSITIVITY STUDY (TYPICAL)

Sub- Two-Phase Operating Cooled Bulk Friction Exit Void Exit Conditionsa Void Void Multiplier MDNBR Mass-Velocity Fraction Quality (MLBM/HR-FT2 )

CASE 1:

Ch. 1 Ch. 2 Ch. 3 Ch. 1 Ch. 2 Ch. 3 Ch. 1 Ch. 2 Ch. 3 Ch. 1 Ch. 2 Ch. 3 LEVY ZUBR EPRI 1.395 1.476 1.558 1.804 1.998 1.987 0.428 0.412 0.380 0.124 0.115 0.102 LEVY ZUBR HOMO 1.413 1.489 1.569 1.850 2.032 2.018 0.420 0.404 0.373 0.120 0.112 0.098 LEVY SMIT HOMO 1.413 1.488 1.566 1.855 2.048 2.013 0.408 0.393 0.365 0.120 0.112 0.099 EPRI EPRI EPRI 1.437 1.515 1.591 1.876 2.071 2.071 0.220 0.206 0.138 0.117 0.110 0.096 EPRI EPRI HOMO 1.446 1.521 1.596 1.902 2.114 2.085 0.218 0.180 0.136 0.115 0.108 0.095 CASE 2:

LEVY ZUBR EPRI 1.594 1.670 1.749 1.904 2.129 2.174 0.402 0.382 0.346 0.088 0.081 0.069 LEVY ZUBR HOMO 1.606 1.678 1.754 1.943 2.158 2.196 0.393 0.375 0.337 0.085 0.078 0.066 LEVY SMIT HOMO 1.604 1.675 1.752 1.945 2.155 2.187 0.379 0.362 0.330 0.085 0.078 0.067 EPRI EPRI EPRI 1.625 1.697 1.767 1.993 2.220 2.242 0.156 0.143 0.084 0.083 0.077 0.066 EPRI EPRI HOMO 1.629 1.699 1.768 2.014 2.276 2.247 0.155 0.120 0.084 0.082 0.076 0.064 a) Denotes operating conditions from Table 4-1.

Table 5-4. [8]CHANNEL MODEL TURBULENT MOMENTUM FACTOR SENSITIVITY STUDY (TYPICAL)

Operating MASS VELOCITY Conditions FTM MDNBR @ MDNBR LOCATION EXIT QUALITY (MLBM/HR-FT2 Ch. 1 Ch. 2 Ch. 3 Ch. 1 Ch. 2 Ch. 3 Ch. 1 Ch. 2 Ch. 3 CASE 1:

0.0 1.37 1.49 1.55 1.75 2.03 1.96 0.125 0.114 0.1021 0.8 1.39 1.48 1.56 1.80 2.00 1.99 0.124 -,0.115 0.102 1.0 1.40 1.47 1.56 1.81 1.99 1.99 0.124 0.116 0.102

. CASE 2:

I 0.0 1.57 1.68 1.74 1.85 2.17 2.15 0.089 0.080 0.069 0.8 1.59 1.0 L1.60 1.67 1.67 1.75 1.75 1.90 1.91 2.13 2.12 2.17 2.18 0.088 0.088 0.080 0.081 0.069 0.0691

-4c

m -n - un -meam - -MmMM TABLE 6-1. RPS TRIP FUNCTIONS Reactor Trip Honitored Parameter Trip Setpoint During Purpose of Trip 4-Pump Operation I. Overpower Neutron flux 105.51 FP To provide core protection during transients trip involving uncontrolled power increase.

2. Power-flow- Neutron flux, RC flow Flux/Flow = 1.08 To provide core protection during transients Imbalance and power imbalance involving a flow reduction and during core trip conditions involving excessive power peaking

3. RCS pressure- RCS pressure and RC Function of RC outlet To provide core protection during transients temperature outlet temperature temperature involving a reduction in pressure or a trip reduction in core heat removal and to ensure reactor shut down during a LOCA.

4. Low CS RCS pressure 1800 poll To provide core protection during

,pressure transients Involving a pressure trip reduction

5. PC Pump Neutron flux and pump Loss of two pumps To provide core protection during Monitor trip contact monitor voltage above 551 FP loss of RC pumps

6. High RCS RCS pressure 2300 paig To provide protection of RCS pressure pressure boundary from excessive pressures trip no 0

7. High RCS PC outlet temp. 619oF To prevent excessive temperature in the rn temperature RCS trip

8. Nigh PC Rs pressure 4 pal To ensure reactor shutdown during a LOCA pressure and 8L1 inside containment.

trip

g PROPRIETARY DUKE POWER CO.

I FIGURE 4-1. [64 CHANNEL MODEL [EIGHTH-CORE] REPRESENTATION HOT ASSEMBLY 1.6147 I I S1.590 1

1.549 1.2 1,495 1.314 1.350 1.342 1.344 1.138 1 .184 0.946 11.321 1.210 II I\ 0.9 42 0.952 0.961 0.962 0.921 0.750

• b 0.7 69 0.721 0.789 0.510 0.450 I 0.4 47 0.400 0.425 I

I PROPRIETARY DUKE POWER CO.

I FIGURE 4-2. [64] CHANNEL MODEL HOT ASSEMBLY DETAIL I HOT SUBCHANNEL I

I1.714 1.705 I\

3 9 10 11 12 1.680 1.660 1.650 1.630 I\ I13 1.645 14 1.635 15 16 17 1,620 . 1.615 1.610 g18 19 - 20 21 22 23 24 I1.635 1.625 1.620 1.615 1.605 1.600 1.590 25 26 27 28 29 30 31 32 1.580 1.570 1.565 1.525 1.510 1.507 1.500 1495 I

PROPRIETARY DUKE POWER CO.

I FIGURE 4-3. [8 AND 9]CHANNEL MODEL HOT ASSEMBLY DETAIL I

1 2 I 1.714 1.705

\

I\

3 4 1.700 1.698 5 6 7 8 1.695 1.690 1.6 80 1.675 I 7 I-1 1.5941 1.9

£ I

I -PROPRIETARY DUKE POWER CO.

FIGURE 4-4. [8]CHANNEL MODEL [EIGHTH] CORE REPRESENTATION I

HOT ASSEMBLY I

I\

3 I I DZDDD U8 I

0.9961 I -I+

I PROPRIETARY DUKE POWER CO.

I FIGURE 4-5. [9]CHANNEL MODEL[EIGHTH]CORE REPRESENTATION I CHOT ASSY. MK 16/7 I

I I

GA1 LEGEND:

Mk-B6/87 ASSYS I ~Mk-B5 ASSYS PROPRIETARY DUKE POWER CO.

I I I

I I

I 0 I K@o 0 0o 0 I .

• o o I 0 0

-o I

3 LJH3 d/N ZX Url

£ FIGURE 5-2 VIPRE-01 vs. LYNX2 MASS VELOCITY AT CHF BWC CHF CORRELATION 4

u O>D I

2 o~

c --o o --:

0.5 1 1.5 2 2.5 3 3.5 4 VIPRE-01 MASS VELOCITY AT CHF -

FIGURE 5-3 VIPRE-01 vs. LYNX2 QUALITY AT CHF BWC CHF CORRELATION 0.3 0.2 0.1 0.0

-0.

-0.30

-0.3 -o -o.U oT.o 0.1 0.2 oa3 VIPIRE-01 QUALITY AT Hff

U PROPRIETARY DUKE POWER CO.

I I I U0 1

I 08 00-0q 0 oo a0 0 0 I o oo o0 o a0 s I o O0 O

LLJa a 00 0 d' 0 00a

>6 0 Q I 10 MO co 0

c-c I o I -- 4 3dH3 -50flSV 1

ma -m ~aem - mm m m FIGURE 5-5 MEASURED/PREDICTED CHF vs. QUALITY VIPRE-01, BWC CORRELATION 1.3 12 x

xA x xx x xx xx x x x )OK W )( x 1.1 - gg xx x xx x x xx x x xX x Kxx xxxxx xx x Xx x

. x x x x XX 0.9 - x x~x x A xxX x) zX x Xx Xx x ~ Xx t xv xxx a O

0.8 - x xWc 08 0 c) 0.7 I I I I

-0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 . :

QUALITY

FIGURE 5-6 MEASURED/PREDICTED CHF vs MASS VELDCITY VIPRE-01, BWC CORRELATION 1.3 4 1.2 xx x xx A x

. - . x XX x x x xt XX x* xx 1 f x x x)) xx X xx x 09 x xx )x x

0.n <1 )Oxx W X0 x" c**

0.8 x x x 0.7 -

O0 0 0.5 1 1.5 2 Z5 3 3.5 4 MAS VELOCITY, Mlbm/hr-ft

FIGURE 5-7 MEASURED/PREDICTED CHF vs. PRESSURE VIPRE-01, BWC CORRELATION 1.23 x

x x xx 1.1 - x x XW )Jc x x

0.9- x xx xi x x c x x x m --

0.8 -

O0 0.7 n 1400 1800 1800 2000 2200 2400 28600 8800 PRESSURF, (psia)

IROIE DUKE POWER CO.

FIGURE 6-1 RPS CORE PROTECTION SAFETY LIMITS I

Thermaa Power Level. S USACCSPTABLE I OPERATION I ACCEPTABLE OPERATION

.100 UUM ~ACCEPTABLE OPERATION I .3 so I ACCEPTABLE so & 0 PUMP 2.3

~OPERATION S40

-EU -4 -200 0 40 60

• SO *40 *20 0 20 40 60 Reactor Power laalance. S II I

I PROPRIETARY DUKE POWER CO.

I High Outlet I --.-emoerature Tri p II Acceptable Operation Variable Low Pressure Set point/

I 4A /1

' %Unacceptable I/ / / Operation I

Low Pressure /-- Limiting Trip /P-T Curve I

Coolant Outlet Temperature 1 -FIGURE 6-2. RPS P-T CORE PROTECTION ENVELOPE I

I PROPRIETARY I- FIGURE 6-3. HIGH TEMPERATURE TRIP MAPS DUKE POWER CO.

I 4.00 3.80 I b 3.60 I 0 z 3.40 3.20- .

I hIi I J 1 1.80 I . El1 AXIAL PEAKS:

F 2.2 0 ------- 1.9 0 .204 .6-0.

2 ~~-A 3-----__1.3 0 0.2 0.4 0.6: 0.81 I

-56 I

I PROPRIETARY DUKE POWER CO.

FIGURE 6-4. LOW PRESSURE TRIP MAPS 4.00 3.80 160.

w 3.40 I 06. -.

3.20 - '

I j

<(

3.00 3.2 2 0

II J .402.

4 2.0-* - - ---

I AXIAL PEAKS:

I 0....

N 1.7 101.5 I--------1.1 ~-----1.3 1.80 1 1111 11 0 0.2 0.4 0:4 0.8 7xA PROPRIETARY DUKE POWER CO.

FIGURE 6-5. FLUX-TO-FLOW MAPS 3.50 3.40 3.20 I Z 3.10 - - .

3.00 2.90 M2.80 I 0 0

2.70 2.60 I 4 S2.50 I 2.40 X 230 S2.20 AXIAL PEAKS:

1.91.5 2.00- ------- 1.3 1.80 1 f11 11 0 0.2 0.4 0.6 0.8 I -XA I

PROPRIETARY DUKE POWER CO.

I I FIGURE 6-6 TYPICAL 2 PUMP COASTDOWN TRANSIENT RESULTS I m I

I I

I I

I IZ I

Limiting Pump Coastdown Statepoint I

II TIME, SEC.

I I

-59 I