ML16343A474

WCAP-11595,Rev 2, W Improved Thermal Design Procedure Instrument Uncertainty Methodology Diablo Canyon Units 1 & 2 24 Month Fuel Cycle Evaluation.

ML16343A474
Person / Time
Site:	Diablo Canyon
Issue date:	01/31/1997
From:	Andre S, Rood M, Trozzo R WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML16342D530	List: ML16342D529 ML16342D531 ML16342D532 ML16342D533 ML16342D535 ML16343A474 ML17083C623
References
WCAP-11595, WCAP-11595-R02, WCAP-11595-R2, NUDOCS 9702120283
Download: ML16343A474 (76)

Text

Westinghouse Non-Proprietary Class 3

+ + + + + + + +

Westinghouse Improved Yherrnal Design Procedure Instrument Uncertainty Methodology Diablo Canyon Units 1 8z 2 24 Month jFuel Cycle Evaluation Westinghouse Energy Systems

.i 9702i20283 97013i PDR ADQCK 05000275 P PDR

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WESTINGHOUSE NON-PROPRIETARY CLASS 3 WCAP-11595 .

REV.2 WESTINGHOUSE IMPROVED THERMAL DESIGN PROCEDURE, INSTRUMENT UNCERTAINTY METHODOLOGY DIABLO CANYON UNITS 1 AND 2 24 MONTH FUEL CYCLE EVALUATION II January 1997 S. V. Andre'.

C. Rood R. W. Trozzo WESTINGHOUSE ELECTRIC CORPORATION Nucl ear Servi ces Di vi sion P. 0. Box 355 Pittsburgh, Pennsylvania 15230-0355 1997 Westinghouse Electric Corporation All Rights Reserved

ACKNOWLEDGMENTS The authors of this report wish to acknowledge L. Bates, T. Oonat, C. Ciocca, and P. Kennevan for their cooperation and technical assistance throughout the Diablo Canyon Units I and 2 Extended Surveillance Interval Program.

TABLE OF CONTENTS SECTION TITLE PAGE Introduction II. Methodology III. Instrumentation Uncertainties

1. Pressurizer Pressure Uncertainty 4

2. Tavg Uncertainty 7

3. RCS Flow Measurement Uncertainty 10

4. Reactor Power Measurement Uncertainty 24 IV. Conclusions 28 References 29

LIST OF TABLES TABLE TITLE PAGE NUMBER Pressurizer Pressure Control System Uncertainty Tavg Control System Uncertainty Calorimetric RCS Flow Measurement 18 Instrumentati on Uncertainti es Calorimetric RCS Flow Measurement 19 Sensitivities Calorimetric RCS Flow Measurement Uncertainty 20 Cold Leg Elbow Tap RCS Flow Indicator 23 Uncertainty Daily Power Measurement Instrumentation 26 Uncertainties Daily Power Measurement Uncertainty 27

LIST OF ILLUSTRATIONS FIGURE TITLE PAGE NUMBER

\

Calorimetric RCS Flow Measurement 31 Calorimetric Power Measurement 32

WESTINGHOUSE IMPROVED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY FOR DIABLO CANYON UNITS 1 AND 2 24 MONTH FUEL CYCLE EVALUATION INTRODUCTION Four operating parameter uncertainties are used in the uncertainty analysis of the Improved Thermal Design Procedure (ITDP). These parameters are Pressurizer Pressure, Primary Coolant Temperature (T,,), Reactor Power, and Reactor Coolant System Flow. They are frequently monitored and several are

'sed for control purposes. Reactor power is monitored by the performance of a secondary side heat balance measurement once every 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. RCS flow is monitored by the performance of a calorimetric RCS flow measurement at the beginning of each cycle, The RCS Cold Leg elbow taps are normalized to the calorimetric RCS flow measurement and used for shift surveillances (with a small increase in uncertainty). Pressurizer pressure is a controlled parameter and the uncertainty reflects the verification of the control system through indication. T,, is a controlled parameter via the temperature input to the rod control system and the uncertainty reflects this c'ontrol system.

This work is applicable for 30 month instrumentation surveillance intervals.

Westinghouse has been involved with the development of several techniques to treat instrumentation uncertainties. An early version (for D. C. Cook 2 and Trojan) used the methodology outlined in- WCAP-8567 "Improved Thermal Design Procedure", ' which is based on the conservative assumption that the uncertainties can be described with uniform probability distributions.

Another approach (for McGuire and Catawba) is based on the more realistic assumption that the uncertainties can be described with random, normal, two sided probability distributions. This approach is used to substantiate the acceptability of the protection system setpoints for many Westinghouse plants, e.g., D. C. Cook 2 , V. C. Summer, Wolf Creek, and others. The second approach is now utilized for the determination of all instrumentation uncertainties for both ITDP parameters and protection functions.

II. METHODOLOGY The methodology used to combine the uncertainty components for a channel is the square root of the sum of the squares of those groups of components which are statistically independent. Those uncertainties that are dependent are combined arithmetically into independent groups, which are then systematically combined. The uncertainty components are considered to be random, two sided distributions. This technique has been utilized before as noted above, and

"' and various industry standards (10,11) has been endorsed by the NRC staff (6,7,8,9)

~ ~

The relationships between the uncertainty components and the channel statistical allowance are variations of the basic Westinghouse Setpoint Methodology 'nd are defined as follows:

For precision parameter indication using Special Test Equipment or a DYM at the input to the racks, and tr'ending of transmitter calibration and drift; CSA = ((PMA) + (PEA) + (SMTE+SD) + (SPE) + (STE) + (SRA) +

(SMTE+SCA) + (RDOUT) ) + BIAS (Eq. I)

2. For parameter indication uti 1 i zing the plant process computer, and trending of transmitter calibration and drift;.

CSA = {(PMA) + (PEA) + (SMTE+SD) + (SPE) + (STE) + (SRA) +

(SMTE+SCA) + (RMTE+RD) Np +(RTE) A~p

+ (RMTE+RCA) gp) +

BIAS (Eq. 2)

3. For parameters which have control systems verified through indication, and with trending of transmitter calibration and drift; CSA = {(PMA) + (PEA) + (SMTE+SD) + (SPE) + (STE) + (SRA) +

{SMTE+SCA) + {CA) + (RMTE+RD) wp + (RTE) mp + (RMTE+RCA) mp +

(RDOUT) 1Np) + BIAS (Eq. 3)

0 where:

CSA Channel Statistical Allowance PMA Process Measurement Accuracy PEA Primary Element Accuracy SRA Sensor Reference Accuracy SCA Sensor Calibration Accuracy SMTE Sensor Measuring II Test Equipment Accuracy SPE Sensor Pressure Effects STE Sensor Temperature Effects

'D Sensor Drift CA Controller Accuracy RCA Rack Calibration Accuracy RMTE Rack Measuring P Test Equipment Accuracy RTE Rack Temperature Effects RD Rack Drift RDOUT Readout Device Accuracy.

The parameters above are as defined in references 5 and 12 and are based on ISA S51. 1-1979 (R93) . However, for ease in understanding they are paraphrased below:

PMA non-instrument related measurement uncertainties, e.g.,

temperature stratification of a fluid in a pipe, PEA uncertainties due to a metering device, e.g., elbow, venturi, orifice, SRA reference (calibration) accuracy for a sensor/transmitter, SCA calibration tolerance for a sensor/transmitter based on plant calibration procedures, SMTE- measuring & test equipment used to calibrate a sensor/

transmitter, SPE change in input-output relationship due to a change in static pressure for a d/p cell, STE change in input-output relationship due to a change in ambient temperature for a sensor/transmitter, SD change in input-output relationship over a period of time at reference conditions for a sensor/transmitter,

CA - the accuracy of the controller, RCA - reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated, RMTf - measuring 5 test equipment used to calibrate rack modules RTE - change in input-output relationship due to a change in

, ambient temperature for the rack modules, RD - change in input-output relationship over a period of time at reference conditions for the rack modules, RDOUT - the accuracy of a special, local test gauge, a digital voltmeter or multimeter on it's most accurate applicable range, or 1/2 of the smallest increment on an indicator, BIAS - a non-random uncertainty for a sensor/transmitter or a process parameter.

A/D - the uncertainty component is associated with a computer readout.

IND ,- the uncertainty component is associated with an analog indicator.

A more detailed explanation of the Westinghouse methodology noting the interaction of several parameters is provided in references 5 and 12.

III. INSTRUMENTATION UNCERTAINTIES The instrumentation uncertainties will be discussed first for the two parameters which are controlled by automatic systems, Pressurizer Pressure and T,, (through Rod Control).

1. PRESSURIZER PRESSURE UNCERTAI TY Pressurizer Pressure uses a closed-loop control system with a comparison of the measured vapor space pressure to a reference value. Proper operation of the control system is verified through indication. The control uncertainties established for use in the ITDP analysis are [ ] ". In conjunction with the move to a 24 month fuel cycle (30 month surveillance interval) the control uncertainties were determined consistent with the Diablo Canyon Units 1 and 2 administrative procedures and account for the verification that all three channels indicate above the limit. Uncertainties are from the transmitter and the process racks/indicators as shown in Table 1. Equation 3

I has been modified to allow for a functional dependency between the transmitter/racks and the indicators or controller. That is, the indicators are independent of the controller; however, an error in the transmitter or rack modules will be common to both the indicator and the controller. As shown in Equation 4 below, the statistical combination of the uncertainties for this function results in a total uncertainty of [ ]"'hen one.

of the 4 normal indicators is unavailable.

CSA = ((PMA) + (PEA) + [(SMTE+SD) + (SPE) + (STE) + (SRA) +

(SMTE+SCA) + (RMTE+RD) f/' (RTE) gpss

+ (RMTE+RCA) gpss

+

(RMTE+RD) ppp + (RTE) fgQ + (RMTE+RCA) ppQ] /(N-1) } +

((CA) + [(RMTE+RD) iso + (RTE) iHo + (RMTE+RCA) ron +

(RDOUT) iso]/(N-1) ) + BIAS (Eq. 4)

+a,c CSA =

In addition, an allowance, is made for pressure overshoot or undershoot due to the interaction and thermal inertia of the heaters and spray. Based on an evaluation of plant operation, an allowance of [ ]"'as made for this effect. The high side of the controller also operates with a 10 psi no-action A

deadband between 2250 and 2260 psig. This value has been included in the uncertainty calculations as a 10 psi bias on the high side only. Therefore, a total pressurizer pressure uncertainty of [ ] 's calculated for a 30 month surveillance interval as noted on Table 1 which is bounded by the ITDP analysis.

TABLE 1" PRESSURIZER PRESSURE CONTROL SYSTEM UNCERTAINTY SENSOR/TRANSMITTER

+a,c SRA SCA SMTE STE SD BIAS PROCESS RACKS EAI EAO CONTROLLER IND RCA RMTE RTE RD RDOUT CA All above values in X of instrument span. Span = 1250 psi.

ELECTRONICS UNCERTAINTY =

PLUS =

ELECTRONICS UNCERTAINTY =

PLUS =

CONTROLLER UNCERTAINTY (3 Indicators. Available)

2. ~22 2 Tavg uses a closed-loop control system that compares the auctioneered high Tavg from the loops to a reference derived from the First Stage Turbine Impulse Chamber Pressure. Proper operation of the control system is verified through indication. Tavg is the average of the narrow range Thot and Tcold values, and the highest loop Tavg is used in the controller. The control uncertainties established for use in the ITDP analysis are [ ]"'. In conjunction with the move to a 24 month fuel cycle (30 month surveillance interval) the control uncertainties were determined consistent with the Diablo Canyon Units 1 and 2 administrative procedures. Uncertainties are from hot leg and cold leg streaming, the RTDs, the turbine pressure transmitter, and the process racks/indicators (as noted on Table 2). Based on the assumption that 2 Thot and 1 Tcold cross-calibrated RTDs are used to calculate Tavg (assuming one failed Thot RTD per loop), Equation 3 has been modified to allow for a functional dependency between the transmitter/racks and the indicators or controller and to average the uncertainties associated with the multiple RTDs and the ERI cards. Using Equation 5 below, the electronics uncertainty is calculated to be [ ]"'. Assuming a normal, two sided probability distribution results in an electronics standard deviation (s>) of [ ]"'.

CSA = (((PMATH) + (PMAsErA) + (PEA) +

(((((SMTE+SD) + (SRA) + (SMTE+SCA) ]/Ng] +

(((SMTE+SD) + (SRA) + (SMTE+SCA) ]/Ntj ) /2) +

(((((RMTE+RD) ERt + (RTE) ERg

+ (RMTE+RCA) ERi]/NH] + [L(RMTE+RD) ERI

+

(RTE) pg( + (RMTE+RCA) fR[]/Nc] )/2) * (SPANE<>/SPAN<zpg) +

(RMTE+RD) Ego

+ (RTE) Ego

+ (RMTE+RCA) Ego)/(N-I) +

[(SMTE+SD)>> + (SPE)>> + (STE)>> + (SRA)>> + (SMTE+SCA)>> +

(RMTE+RD)>> + (RTE)>> + (RMTE+RCA)>>](TP Sen) ) + ((CA) +

[(RMTE+RD),, + (RTE),, + (RMTE+RCA),, + (RDOUT),,]/(N-I)) (Eq. 5) 2'

I

+a,c CSA = I However, this does not include the controller deadband of + 1.5 'F. The controller uncertainty is the combination of the electronics uncertainty and the deadband. The probability distribution for the deadband has been determined to be [

] ."'he variance for the deadband uncertainty is then:

]+a,c Combining the variance for the electronics and the variance for the deadband results in a controller variance of:

]+a,c (sc) = (s>) + (s>)' [

The controller standard deviation st = [ ] " results in a total-random uncertainty for a 30 month surveillance interval of [ ]"'nd a cold leg streaming bias of [ ] ", which are bounded by the ITDP analysis.

0 TABLE 2 TAVG CONTROL SYSTEM UNCERTAINTY SENSOR/TRANSMITTER (All values in X span)

Tavg ~

Turbine Pressure*

Span 100 F 620 psig

+a,c PMA BETA SRA SCA SMTE STE SD BIAS TP Sen PROCESS RACKS (All values in X span)

Tavg ERI EAO INDICATOR CONTROLLER TURBINE Span 150 F 100 F 100 F 100 F 620 psig

+a,c RCA RMTE =

RTE RD RDOUT=

CA NH

= ¹ HOT LEG RTDs = 2 Nc = ¹ COLD LEG RTDs = 1

+a,c ELECTRONICS CSA ELECTRONICS SIGMA =

CONTROLLER SIGMA CONTROLLER BIAS CONTROLLER CSA

• Based on Rosemount uncertainties for Turbine Pressure. This bounds the results for the Barton transmitters.

• • Sensitivity of Tref to Turbine Pressure uncertainties. Based on relationship between turbine pressure and Tref from no load to full power.

3. RCS FLOW MEASUREMENT U CERTAINTY Calorimetric RCS Flow Measurement ITDP and plant Technical Specifications require a calorimetric RCS flow measurement every fuel cycle above 90% of Rated Thermal Power with a high degree of accuracy. It is assumed for this uncertainty analysis that the calorimetric RCS flow measurement is performed before feedwater venturi fouling is significant, therefore, no separate allowances have been made for feedwater venturi fouling. This analysis is based on information supplied by PGEE in References 15 and 17. Any changes in plant configuration or procedures made subsequent to that submittal will be addressed by PGRE.

The flow measurement is performed by determining the steam generator thermal output (corrected for the RCP heat input and the loop's share of primary system heat losses) and the enthalpy rise (Delta-h) of the primary coolant.

Assuming that the primary and secondary sides are in equilibrium, the RCS total vessel flow is the sum of the individual primary loop flows, i.e.,

(Eq. 6)

The individual primary loop volumetric flows are determined by correcting the thermal output of the steam generator for steam generator blowdown (if not secured), subtracting the RCP heat addition, adding the loop's share of the primary side system losses, dividing by the primary side enthalpy rise and multiplying by the cold leg specific volume. The equation for this calculation is:

Wc = ~A~~p~+~N~vt (hH - hc) (Eq. 7) where; w( Loop flow (gpm)

A 0.1247 gpm/(ft /hr)

Qs6 Steam generator thermal output (Btu/hr)

Op RCP heat addition (Btu/hr)

Primary system net heat losses (Btu/hr) 10

>c Specific volume of the cold leg at Tz (ft'/lb)

N Number of primary side loops h Hot leg enthalpy (Btu/lb) hc Cold leg enthalpy (Btu/lb).

The thermal output of the steam generator is determined by a secondary side calorimetric measurement, which is defined as:

gsG

= (h, - hg)Wg - pgp(h, - hgp)VBD/A (Eq. 8) where; h, Steam enthalpy (Btu/lb) hg Feedwater enthalpy (Btu/lb)

Wg Feedwater flow (lb/hr) pen Blowdown density (lb/ft')

Van Volumetric flowrate of blowdown (gpm) hen Blowdown enthalpy (Btu/lb).

The steam enthalpy is based on measurement of steam generator outlet steam pressure, assuming saturated conditions. The feedwater enthalpy is based on the measurement of feedwater temperature and an assumed feedwater pressure.

The feedwater flow is determined by multiple measurements and the following calculation:

where K venturi flow coefficient hp'eedwater Fa Feedwater venturi correction for thermal expansion pr Feedwater density (lb/ft')

Feedwater venturi pressure drop (inches Hq0).

The feedwater venturi flow coefficient is the product of a number of constants including as-built dimensions of the venturi and calibration tests performed by the vendor. The thermal expansion correction is based on the coefficient of expansion of the venturi material and the difference between feedwater temperature and calibration temperature. Feedwater density is based on the 11

measurement of feedwater temperature and an assumed feedwater pressure. The venturi pressure drop is obtained from the output of the differential pressure transmitter connected to the venturi.

t RCP heat addition is determined by calculation, based on the best estimate of coolant flow, pump head, and pump hydraulic efficiency. The primary system net heat losses are determined by calculation, considering the following system heat inputs and heat losses:

Charging flow Letdown flow Seal injectioq flow RCP thermal barrier cooler heat removal Pressurizer spray flow Pressurizer surge line flow Component insulation heat losses Component support heat losses CRDM heat losses.

A single calculated sum for 100K RTP operation is used for these losses or heat inputs.

The hot leg enthalpy is based on the measurement of the hot leg temperature and the pressurizer pressure. The cold leg enthalpy and specific volume are based on measurement of the cold leg temperature and calculation of the cold leg pressure from the pressurizer pressure measurement.

The RCS flow measurement is thus based on the following plant measurements:

Steamline pressure (P,)

Feedwater temperature (Tq)

Feedwater pressure (Pq assumed value)

Feedwater venturi differential pressure (hp)

Hot leg temperature (TH)

Cold leg temperature (Tc)

Pressurizer pressure (Pp) 12

Steam generator blowdown (V8p) and on the following calculated values:

Feedwater venturi flow'coefficients (K)

Feedwater venturi thermal expansion correction (F,)

Feedwater density (pq)

Feedwater enthalpy (h<)

Steam enthalpy (h,)

Moisture carryover (impacts h,)

Primary system net heat losses (gL)

RCP heat addition (Pp)

Hot leg enthalpy (hz)

Cold leg enthalpy (h<)

Cold leg pressure (P~L)

Cold leg specific volume (vz)

Blowdown density (p8p)

Blowdown enthalpy (h>>).

These measurements and calculations are presented schematically on Figure 1.

The deri,vation of the measurement uncertainties and flow uncertainties on Table 5 is noted below.

Secondar Side The secondary side uncertainties are in four principal areas, feedwater flow, feedwater enthalpy, steam enthalpy and RCP heat addition. These four areas are specifically identified on Table 5.

For the measurement of feedwater flow, each feedwater venturi is calibrated by accuracy of [ j '.

the vendor in a hydraulics laboratory under controlled conditions to an The calibration data which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of [ ]"'s included for installation effects, resulting in an overall flow c'oefficient (K) uncertainty of [

Since RCS loop flow is proportional to steam generator thermal output which is 13

proportional to feedwater flow, the flow coefficient uncertainty is expressed as + 0.5 X flow. It should be noted that Westinghouse makes no explicit allowance for venturi fouling.

The uncertainty applied to the feedwater venturi thermal expansion correction (F,) is based on the uncertainties of the measured feedwater temperature and the coefficient of thermal expansion for the venturi material, usually 304 stainless steel. For this material, a change of + 1 'F in the nominal feedwater temperature range changes F, by + 0.002 X and the steam generator thermal output by the same amount.

uncertainty in F, of + 5 X for 304 stainless steel is used in this I

An analysis. This results in an additional uncertainty of [ ] 'n

~+a,c feedwater flow. Westinghouse uses the conservative value of f Using the Steam Tables it is possible to determine the sensitivities of various parameters to changes in feedwater temperature and pressure. Table 3 notes the instrument uncertainties for the hardware used to perform the measurements. Table 4 lists 'the various sensitivities. As can be seen on Table 4, feedwater temperature uncertainties have an impact on venturi F feedwater density and feedwater enthalpy. Feedwater pressure uncertainties impact feedwater density and feedwater enthalpy.

Feedwater venturi hp uncertainties are converted to X feedwater flow using the following conversion factor as defined in Table 3-28 of Reference 14:

X flow = (hp uncertainty in Xdp span)(1/2)(transmitter span/100)

The feedwater flow transmitter span is 115 X of nominal flow.

Using the Steam Tables again, it is possible to determine the sensitivity of steam enthalpy to changes in steam pressure and steam quality. Table 3 notes the uncertainty in steam pressure and Table 4 provides the sensitivity. For steam quality, the Steam Tables were used to determine the sensitivity at a moisture content of [ j ". This value is noted on Table 4.

14

The net pump heat uncertainty is derived from the combination of the primary system net heat losses and pump heat addition and are suranarized for Diablo Canyon as follows:

System heat losses -2.3 MWt Component conduction and convection losses Pump heat adder ~+1 .8 Net Heat input to RCS +14.4 MWt An additional component of 1.3 MWt has been identified by PG&E to account for heat losses on the core side between the hot and cold leg RTDs. The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [ ]"'f the calculated value. Since direct measurements are not possible, the uncertainty on to be

[ ]"'f component conduction and convection losses has been assumed the calculated value. Reactor coolant pump hydraulics are known to a relatively high confidence level, and are supported by system hydraulics tests performed at Prairie Island II and by input power measurements from several plants. Therefore, the uncertainty for the pump heat addition is estimated to be [ ]

" of the best estimate value.

Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are approximately [ ]"'f the total net heat input. For this calculation, a conservative value of [

] " is used in Table 5.

Primar Side The primary side uncertainties are in three principal areas, hot leg enthalpy,'old leg entlialpy and cold leg specific volume. These are specifically noted on Table 5. Three primary side parameters are actually measured, T<, Tc and pressurizer pressure. Hot leg enthalpy is influenced by T>, pressurizer pressure and hot leg temperature streaming. The uncertainties for the instrumentation are noted on Table 3 and the sensitivities are provided on Table 4. The errors associated with measuring hot leg streaming are split into two components, one random and the other systematic. For Diablo Canyon Units 1 and 2, the streaming measurement uncertainty is [ ]"'. random and 15

[ ] " systematic. The Hot Leg and Cold Leg Streaming Biases due to fluid stratification are not considered when determining RCS Flow Measurement Uncertainty because these biases add conservatism to the determination of total RCS flow.

The cold leg enthalpy and specific'olume uncertainties are impacted by T{ and the pressure in the cold leg (which is determined by PGRE by adding a correction factor to the pressurizer pressure). Table 3 notes the Tt instrument uncertainty and Table 4 provides the sensitivities.

Parameter dependent effects are identified on Table 5. Westinghouse has determined the dependent sets in the calculation and the direction of interaction, i.e., whether components in a dependent set are additive or subtractive with respect to a conservative calculation of RCS flow. The same work was performed for the instrument bias values. As a result, the calculation explicitly accounts for dependent effects and biases with credit taken for sign (or direction of impact).

Using Table 5, the 4 loop uncertainty equation (with biases) is as .follows:

Flow = ((FSv) + {(ps + h~, - F~){I3 + (F<<) + (pz - hip + (SGBD)

+ (h, )+ (NFM) +.{(W) + (h,P + (hii,) + (his~) + (hc, + vg,) )/N

+(hH, - hg, - vga + (hHQ ) + (pp Bias) - (hf Bias)

- (hB, Bias) + (hH Bias) - (h(- Bias) - (v~ Bias) 16

Based on four (4) loops, the number, type and measurement method of RTDs, and the vessel Oelta-T, the uncertainty for the precision RCS flow measurement is:

8 of loops flow uncertainty (X flow) 4

).a,c 17

TABLE 3 CALORIMETRIC RCS FLOW MEASUREMENT INSTRUMENTATION UNCERTAINTIES (X SPAN) FW TEMP FW PRES FW d/p STM PRESS TH T< PRZ PRESS

+a,c SRA SCA SMTE SPE STE SD SDBIAS BIAS RCA RMTE RTE RD RCAEAI(ERI)

RT EEAI (ERI)

EAI (ERI)

RCAEAQ RTEEAO RDEAO NUMBER OF INSTRUMENTS USED 3 1 2/LOOP 1/LOOP 3/LOOP 1/LOOP 1 SPAN = 500 'F 2000 psi 115%FLOW 1200 psi 100'F 100'F 1250 psi OF OF psi Mp span psi OF psl

+a,c INST UNC.

(RANDOM) =

INST UNC.

(BIAS)

NOMINAL = 430 855 PSIA 100'%FLOW 820 PSIA 600 540 2250 PSIA

• Uses uncertainties associated with Barton transmitters. These values bound the Rosemount values for -the RCS Flow Calorimetric function.

• • These values are in Tavg Span. l

• • " Bias due to additional temperature effect on Barton transmitters identified in Reference 16 and location of transmitters per Reference 17.

Feedwater pressure is assumed. TH and TC are calculated as read from the MMI.

All other parameters are calculated as read by the process computer. A 35'F temperature rise is assumed for the EAGLE-21 racks and a 50'F temperature rise is assumed for the process computer. The static pressure span effect for the feedwater flow dp transmitter is based on a feedwater pressure of 905PSIA and maximum transmitter output.

18

TABLE 4 CALORIMETRIC RCS FLOW MEASUREMENT SENSITIVITIES FEEDWATER FLOW ta,c Fa TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE h, 1198.5 BTU/LBM hf 408.3 BTU/LBM Dh(SG) = 790.2 BTU/LBM SG BLOWDOWN 0.009246 'X/gpm NET PUMP/SYSTEM HEAT = 0. 1202 'X/MWt

+a,c STEAM ENTHALPY PRESSURE MOISTURE HOT LEG ENTHALPY TEMPERATURE PRESSURE hH 612.7 BTU/LBM hc 534.5 BTU/LBM Dh(VESS) 78.3 BTU/LBM Cp(T) 1.419 BTU/LBM-'F

+a,c COLD LEG ENTHALPY TEMPERATURE PRESSURE Cp(Tc) 1. 217- BTU/LBM-'F

+a,c COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE 19

TAB,LE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTY COMPONENT INSTRUMENT ERROR FLOW UNCERTAINTY

+a,c FEEDWATER FLOW VENTURI, THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM'NTHALPY PRESSURE MOISTURE SG BLOWDOWN NET PUMP HEAT ADDITION HOT LEG ENTHALPY TEMPERATURE STREAMING, RANDOM STREAMING, SYSTEMATIC PRESSURE COLD LEG ENTHALPY TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE

• , **, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS 20

TABLE 5 (CONTINUED)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTY COMPONENT FLOW UNCERTAINTY'a,c BIAS VALUES FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRESSURE ENTHALPY - HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME - COLD LEG FLOW BIAS TOTAL VALUE

+a,c SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES) 4 LOOP UNCERTAINTY (WITHOUT BIAS VALUES) 4 LOOP UNCERTAINTY (WITH BIAS VALUES) 21

Cold Leg Elbow Tap RCS Flow Indicator Uncertainty As noted earlier, the calorimetric RCS flow measurement is used as the reference for the normalization of the cold leg elbow tap RCS flow indicators at the beginning of each fuel cycle. The Technical Specification surveillance requirement can be met either through the process computer or by reading the control board indicators. This calculation assumes the use of the board indicators which bounds the computer readout. Table 6 notes the instrument uncertainties for normalization of the elbow taps, assuming three elbow tap RCS flow indicators per reactor coolant loop. The dp transmitter uncertainties are converted to X flow in a similar manner to the feedwater venturi hp. The elbow tap RCS flow indicator uncertainties are combined with the calorimetric RCS flow measurement uncertainty for determination of the total elbow tap RCS flow indicator uncertainty.

The total elbow tap RCS flow indicator uncertainty is:

¹ of loops flow uncertainty (% flow) 4 + 2.24 The corresponding ITDP value would be:

¹ of loops standard deviation (% flow) 4 22

TABLE 6 COLD LEG ELBOW TAP RCS FLOW INDICATOR UNCERTAINTY INSTRUMENT UNCERTAINTIES X d,p SPAN X FLOW ta,c PMA PEA SRA SCA SMTE SPE STE SD BIAS RCAggo RMTEioo RTErNn RDreo RDOUT RCAfA'f RMTEEA~

RTEpAi RDEAy RCAfAQ RMTEfAo RTEEAo RDEAo FLOW CALORIM. BIAS FLOW CALORIMETRIC INSTRUMENT SPAN

+a>c SINGLE LOOP ELBOW TAP FLOW UNC (USE OF 3 RCS FLOW INDICATORS) 4 LOOP RCS FLOW UNCERTAINTY (WITHOUT BIAS VALUES) 4 LOOP RCS FLOW UNCERTAINTY (WITH BIAS VALUES) 23

'II

4. REACTOR POWER MEASUREMENT UNCERTAINTY Diablo Canyon Units 1 and 2 perform a primary/secondary side heat balance once every 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> when power is above 15% Rated Thermal Power. This heat balance is used to verify that the plant is operating within the limits of the Operating License and to adjust the Power Range Neutron Flux channels when the difference between the NIS and the heat balance is greater than that required by the plant Technical Specifications.

Assuming that the primary and secondary sides are in equi librium; the core power is determined by sunning the thermal output of the steam generators, correcting the total secondary power for steam generator blowdown (if not secured), subtracting the RCP heat addition, adding the primary side system losses, and dividing by the core rated Btu/hr at full power. The equation for this calculation is:

RP = QN'~~p~+ 100 (Eq. 10)

H where; RP Core power (X RTP)

N Number of primary side loops Qsg Steam generator thermal output (BTU/hr) as defined in Eq. 8 Qp RCP heat adder (Btu/hr) as defined in Eq. 7 Primary system net heat losses (Btu/hr) as defined in Eq. 7 H Core rated Btu/hr at full power.

For the purposes of this uncertainty analysis (and based on H noted above) it is assumed that the plant is at 100% RTP when the measurement is taken.

Measurements performed at lower power levels will result in different uncertainty values. However, operation at lower power levels results in increased margin to DNB far in excess of any margin losses due to increased measurement uncertainty.

The secondary side power measurement equations and effects are the same as those noted for the precision RCS flow calorimetric measurement (secondary side portion), equations 8 and 9, except for the feedwater venturi flow coefficient (K). An uncertainty of 1.42 X on K has been determined by PGEE to account for the use of an ultrasonic flowmeter to benchmark and trend feedwater flow. The measurements and calculations are presented schematically on Figure 2. Table 7 provides the instrument uncertainties for those measurements performed. Since it is necessary to make this determination daily, the plant process computer is used for the measurements. The sensitivities calculated are the same as those noted for the secondary side on Table 4. As noted on Table 8, Westinghouse has determined the dependent sets in the calculation'nd the direction of interaction. This is the same as that performed for the RCS flow measurement, but applicable only to power.

I Using the power uncertainty values noted on Table 8, the 4 loop uncertainty (with bias values) equation is as follows:

+a,c Based on four (4) loops and the instrument uncertainties for the four parameters, the uncertainty for the secondary side power calorimetric measurement is:

k'f loops power uncertainty

~+a,c (X RTP) t:

The actual value used in the ITDP analysis was based on a power uncertainty of 2.0% RTP and is therefore bounding with respect to this calculation.

25

TABLE 7 DAILY POWER MEASUREMENT INSTRUMENTATION UNCERTAINTIES (X SPAN) FW TEMP FW PRES FW d/p STM PRESS*

+a,c SRA =

SCA =

SMTE=

SPE =

STE =

SD BIAS=

RCA =

RMTE=

RTE =

RD NUMBER OF INSTRUMENTS USED 2/LOOP 1/LOOP INST SPAN = 500 'F 2000 psi 115'AFLOW 1200 psi OF psi Mp span psi

+a,c INST UNC (RANDOM)

INST UNC (BIAS)

NOMINAL = 430 855PSIA 100%FLOW 820PSIA

• Uses uncertainties associated with Rosemount transmitters. These values bound the Barton values for the Power Calorimetric function. EAI and EAO uncertainties apply only to the Steam Pressure signal.

• • Bias accounts for location of the transmitter per Reference 17.

All parameters are read by the process computer, except feedwater pressure which is an assumed value. A 35'F temperature rise is assumed for the EAGLE-21 racks and a 50'F temperature rise is assumed for the process computer. The static pressure span effect for th'e feedwater flow hp transmitter is based on a feedwater pressure of 905PSIA and maximum transmitter output.

26

TABLE 8 DAILY POWER MEASUREMENT UNCERTAINTY COMPONENT INSTRUMENT.,ERROR POWER UNCERTAINTY

+a,c FEEDWATER FLOW VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION SG BLOWDOWN BIAS VALUES FEEDWATER DELTA P FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY POWER BIAS TOTAL VALUE

• , ** INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS. VALUES) 4 LOOP UNCERTAINTY (WITHOUT BIAS VALUES) 4 LOOP UNCERTAINTY (WITH POSITIVE BIAS VALUES) 27

IV. CONCLUSIONS The preceding sections provide the methodology to account for pressure, temperature, power and RCS flow uncertainties for the ITDP analysis. The plant-specific instrumentation and procedures have been reviewed for Diablo Canyon Units 1 and 2 and the uncertainty calculations are completed for 30 month surveillance intervals. The values used in the ITDP analysis are more conservative than those listed below.

+a,c Pressurizer pressure uncertainty Temperature (Tavg) uncertainty Power measurement uncertainty RCS flow uncertainty - cold leg elbow tap RCS flow indicators normalized to a calorimetric RCS ilow measurement 28

REFERENCES

1. Westinghouse letter NS-CE-1583, C. Eicheldinger to J. F. Stolz, NRC, 10/25/77.

l

2. Westinghouse letter NS-PLC-5111, T. M. Anderson to E. Case, NRC, 5/30/78.

3. Westinghouse letter NS-TMA-1837, T. M. Anderson to S. Varga, NRC, 6/23/78.

4. Westinghouse letter NS-EPR-2577, E. P. Rahe Jr. to C. H. Berlinger, NRC, 3/31/82.

5. Westinghouse Letter NS-TMA-1835, T. M. Anderson to E. Case, NRC, 6/22/78.

6. NRC letter, S. A. Varga to J. Dolan, Indiana and Michigan Electric Company, 2/12/81.

7. NUREG-0717 Supplement No. 4, Safety Evaluation Report related to the operation of Virgil C. Summer Nuclear Station Unit No. 1, Docket 50-395, August, 1982..

8. Regulatory Guide 1. 105 Rev. 2, "Instrument Setpoints for Safety-Related Systems", 2/86.

9. NUREG/CR-3659 (PNL-4973), "A Mathematical Model for Assessing the Uncertainties of Instrumentation Measurements for Power and Flow of PWR Reactors", 2/85.

10. ANSI/ANS Standard 58.4-1979, "Criteria for Technical Specifications for Nuclear Power Stations".

. 11. 'ISA Standard S67.04, 1994, "Setpoints for Nuclear Safety-Related Instrumentation Used in Nuclear Power Plants" 29

12. Tuley, C. R., Williams, T.P., "The Significance of Verifying the SAMA PMC 20.1-1973 Defined Reference Accuracy for the Westinghouse Setpoint Methodology", Instrumentation, Controls, and Automation in the Power Industry, June, 1992, Vol.35, pp.497-508.

13. Instrument Society of America Standard S51.1-1979 (Reaffirmed 1993),

"Process Instrumentation Terminology."

14. WCAP,-11082, Revision 3, "Westinghouse Setpoint Methodology for Protection Systems, Diablo Canyon Units 1 and 2, 24 Month Fuel Cycle Evaluation",

June'1996.

15. PG&E letter Chron. k'30223, S. D. Kamdar to J. Hoebel, 5/23/96.

16. ITT Barton Letter dated September 14, 1995, "Industry Notification of Temperature Errors in Model 763 and 763A Gauge Pressure Transmitters.

17. PG&E letter Chron. 8 231739, S. D. Kamdar to J. Hoebel, 12/2/96.

30

SECONDARY SIDE PRIMARY SIDE Veo P, Pg TH Pp Tc assumed h~ s heos hg Pr F, hH hc PcL Peo Wf >c calculated Qse

- measured/

D assumed QL/N Qp Mass W

Volume' Z Other Loops RCS VOLUMETRIC FLOW Figure 1 CALORIMETRIC RCS FLOW MEASUREMENT 31

SECONDARY SIDE Fou ing assumed Correction

~BD P, Pf from AMAG or ASME Nozzles hz, hBD hf Pf F, PBD Wf QSB a - measured/

assumed Other Loops ZQ, Core Power FIGURE 2 CALORIMETRIC POWER MEASUREMENT 32

4