ML20079D603

Westinghouse Revised Thermal Design Procedure Instrument Uncertainity Methodology for Alpr,Farley Nuclear Plant,Units 1 & 2 (for RTD Bypass Loops)

ML20079D603
Person / Time
Site:	Farley
Issue date:	05/31/1991
From:	Moomau W WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19302E789	List: ML20079D601 ML20079D602 ML20079D603 ML20079D605 ML20079D607 ML20079D610
References
WCAP-12770, NUDOCS 9107180286
Download: ML20079D603 (37)

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WESTINGHOUSE CLASS 3 $

WCAP 12770 WESTINGHOUSE-REVISED THERMAL DESIGN PROCEDVRE INSTRUMENT UNCERTAINTY METHODOLOGY FOR ALABAMA POWER-FARLEY NUCLEAR PLANT UNITS 1 AND 2 (FOR RTD BYPASS LOOPS)

~MAY, 1991 W.H.Moomau -1

! o l

[

L

~

Westinghouse Electric Corporation Energy Systems P.O. Box 355 Pittsburgh,-Pernsylvania 15230

'(c)1991 Westinghouse- Electric, All- Rights Reserved

,n g , - e- a,,,a w .e -- - - - - - - . - , - . . - - .

T*) 7. < 7 ,

ACKNOWLEDGEMENTS a

The author wishes to recognize the contributions of Mr. Michael G. Eidson of Southern Nuclear Company for his careful review of this report.

W i

TABLE OF CONTENTS SECTION TITLE PAGE I. Introduction 1

!!. Methodology 2 III. Instrumentation Uncertainties S IV. Conclusions 27 References 28 58 i

l LIST OF TABLES i

l

i. TABLE NUMBER TITLE PAGE i 4

1 Pressurizer Pressure Control 6 I System Accuracy 2 Rod Control System Accuracy 8 3 Flow Calorimetric Instrumentation 17 Uncertainties 4 Flow Calorimetric Sensitivities 18 5 Calorimetric RCS Flow Heasurement 19 Uncertainties  ;

6 Cold leg Elbow Tap riow Uncertainty 22 7 Power Calorimetric Instrumentation 25 Uncertainties 8 Secondary Side Power Calorimetric 16 Measurement Uncertainties O

iii

l l

LIST OF ILLUSTRATIONS FIGURE NUMBER TITLE PAGE 1 RCS Flow Calorimetric Schematic 30 2 Power Calorimetric Schematic 3)

+

iv

WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY HETHODOLOGY FOR ALABAMA POWER FARLEY NUCLEAR PLANT UNITS 1 AND 2 ,

(FOR RID BYPASS LOOPS) 1 I. INTRODUCTION l

l l

Four operating parameter uncertainties are used in the uncertainty analysis of the Revised Thermal Design Procedure (RTDP). These parameters j are Pressurizer Pressure Primary Coolant Average Temperature (Tayg),

Reactor Power, and Reactor Coolant System Flow. They are frequently monitored and several are used for control purposes. Reactor power is monitored by the performance of a secondary side heat balance (power j calorimetric) 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 precision flow calorimetric at the beginning of each cyt.l e . Pressurizer pressure is a controlled parameter and the uncertainty reflects the control :ystem. T,yg is a controlled parameter via the tempe;ature input to the rod control system and the uncertainty reflects this control system. This report is based on RTD Bypass Loops included in the design to measure hot and cold leg reactor coolant system temperatures and is applicable for 2660 and 2785 Hwt operation. The RTDP(14) is used to predict the plant's DNBR design limit. The RTDP methodology considers the uncertainties in the system operating plant parameters, fuel fabrication and nuclear and thermal parameters and includes the use of variou ONB correlations. Use of the R'OP methodology requires that variances in the plant operating paramebrs be justified. The purpose of the following evaluatit is to define the specific Farley Nuclear Plant instrument uncertainties for the four primary system operating parameters.

Westinghouse has been involved with the development of several techniques to treat instrumentation uncertainties. An early version (for D. C. Cook l

2 and Trojan) used the methodology outlined in WCAP-8S67 " Improved Thermal Design Procedure",(1,2,3) which is based on the conservative assumption 1-

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.(4) This approach is used to substantiate the acceptability of the protection system setpoints for many Westinghouse plants, e.g., D. C. Cook 2(5), V. C. Summer, Wolf Creek, Millstone Unit 3 and others.

The second approach is now utilized for the determination of all instrumentation errors fr- both RTDP parameters and protection functions.

The unceitpinty calculations in this report are based on a detailed review of Farley Nuclear Plant procedures for innrument calibration, heat balance calculations, and RCS flow measurement. The evaluation of heat balance uncertainties includes both the precision heat balance for RCS flow determination as well as the plant process computer heat balance used for the daily nuclear instrumentation alignment surveillance.

II. METHODOLOGY The methodology used to combine the error components for a channel is the square root of the sum of the squares of those groups of components which are statistically independent. Those errors that are dependent are combined arithmetically into independent groups, which are then systematically combined. The uncertainties used are considered to be random, two sided distributions. The sam of both sides is equal to the range for that parameter, e.g., rack drif t is typically

, [ ]+a,c, the range for this parameter is ( )+a,c, This technique has been utilized before as noted above, and has been endorsed by the NRC staff (6,7,8,9) and various industry standards (10,ll),

2-

j i

The relationships between the error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint  !

Methodology (12) and are defined as follows:

, 1. For precision parameter indication using special test equipment or a DVM at the input to the racks, CSA - ((SCA + SMTE + SD)2 + (SPE)2 + (STE)2+ (RDOUT)2)l/2

+ BIAS. Eq. 1

2. For parameter indication utilizing the plant process computer,-

CSA - ((SCA + SMTE + SD)2 + (SPE)2 + (STE)2 + (RCA + RMTE + RD)2

+ (RTE)2 + (ID)2 + (A/D)2)l/2 + BIAS. Eq. 2

3. For parameters which have control systems, CSA - ((PMA)2 + (PEA)2 +(SCA + SMTE + SD)2 + (SPE)2 + (STE)2

+ (RCA + RMTE + RD + CA)2 + (RTE)2)l/2 + BIAS. Eq. 3 where, ,

CSA - Channel Allowance PMA - Process Measurement Accuracy PEA - Primary Element Accuracy SCA - Sensor Calibration Accuracy SMTE - Sensor Measurement and Test Equipment Accuracy SPE - Sensor Pressure Effects STE - Sensor Temperature Effects SD = Sensor Drift-RCA - Rack Calibration Accuracy RMTE - Rack Measurement'and Test Equipment Accuracy

. RTE - Rack Temperature Effects RD - Rack Drift

, . -RDOUT - Readout Device Accuracy (DVM or Gauge)

ID - Computer Isolator Drift l A/D = Analog to Digital Convers & trwracy CA - Controller Accuracy.

I - - . . .-. - -- .

f The parameters above are as defined in references 5 and 12 and are based on  !

SANA Standard PMC 20.1,1973(I3). However, for ease in understanding they are paraphrased below: ,

PMA -

non-instrument related measurement errors, e.g., temperature stratification of a fluid in a pipe, PEA -

errors due to a metering device, e.g., elbow, venturi, orifice, l SCA -

reference (calibration) accuracy for 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, RCA - reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated, or tuned, to this accuracy, i 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 measurement accuracy of a special test local gauge, digital voltmeter or multimeter on its most accurate applicable range for the parameter measured, ID -

change in input-output relationship over a period of time at reference conditions for a control / protection signal isolating device, A/D - allowance for conversion accuracy of an analog signal to a ,

digital signal for process computer use, CA -

allowance for the accuracy of a controller, not including deadband.

BIAS - a non-random uncertainty for a ' sensor / transmitter or a process parameter.

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

!!!. INSTRUMENTATION UNCERTAINTIES The instrumentation uncertainties will be discussed first for the two

, parameters which are controlled by automatic systems -- Pressurizer Pressure and RCS average temperature (Tayg). Then the development of the uncertainties for the RCS flow and the secondary side power calorimetric measurements will be discussed .

1. PRESSURIZER FRESSURE Pressurizer pressure is normally controlled automatically to simplify plant operation and to maintain pressure within the normal steady state envelope of operation assumed in the safety analysis. To ensure that pressure is restored within its limit following load changes and other expected transient operation, a 12 hour1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> surveillance of pressurizer pressure through instrument readout is included in the Technical Specifications.

Pressurizer pressure is controlled by comparison of the measured vapor space pressure and a reference value. Allowances are made for the transmitter and the process rat.ks/ controller. As noted on Table 1, the electronics uncertainty for this function is [ ]+a,c which corresponds to an accuracy of ( )+a.c. In addition to the controller

. accuracy, 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 allowence of ( ]+a,c was made for this effect. Therefore, a total control system uncertainty of (

]+a,c is calculated, which results in a standard deviation of (

]+a,c (assuming a normal, two sided probability distribution).

-S-

TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY

+a,c

- SCA -

SMTE-STE -

• SD -

BIAS-RCA -

RMTE-RTE -

RD -

CA -

ta C ELECTRONICS UNCERTAINTY -

PLUS ELECTRONICS UNCERTAINTY -

PLUS CONTROLLER UNCERTAINTY -

This calculation is performed assuming that:

+a,C M

b e

6-

2. TAVG T,yg is normally controlled automatically through the rod control

, system to simplify plant operation and to maintain T,yg within the normal steady state envelope of operation assumed in the safety analysis. To ensure that temperature is restored within its limit following load changes and other expected transient operation, a 12 hour1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> periodic surveillance of RCS T,yg through instrument readout is included in the Technical Specifications.

T,yg is controlled by a system that compares the auctioneered high T,yg from the loops with a reference that is derived from the first stage turbine impulso chamber pressure. T avg is the average of the narrow range Tg and TC values. The highest loop T,yg is then used in the controller. Allowances are made (as noted on Table 2) for the RTDs, transmitter and the process racks / controller. The CSA for this function is dependent on the type of RTD, pressure transmitter, and the location of the RTDs, i.e., in the RTD bypass manifold or in the hot and cold legs. Based on the assumption that 1 TH and 1 TC cross-calibrated RTDs are used to calculate Tavg and the RTDs are located in the RTD bypass manifold, the CSA for the electronics is

( )+a,c. Assuming a normal, two sided probability distribution results in an electronics standard deviation (si) of

[ j+a,c, However, this does not include the controller deadband of 1.5 0F.

The controller accuracy is the combination of the instrumentation accuracy and the deadband. The probability distribution for the deadband has been determined to be (

l, ).+a,c The variance for the deadband uncertainty is then:

(s2) "[ ]+a,c, 1

l

. 7

. .- .~ . - . - = . - - - . - - - - - - - . _ - _ - - - . - . - -

Combining the variance for instrumentation and deadband results in a controller variance of:

(sT) = (si)2 + (s2) "I 3+"'"'

O With a controller si = [ ]+ ' , the total uncertainty is

( )+a.c,  :

TABLE 2 R00 CONTROL SYSTEM ACCURACY Tavg TURB PRESS

- +a.c .

PMA =

SCA -

SMTE-STE -

SD -

BIAS- "

RCA =

RMTE-RMTE-  !

RTE -

RD -

CA =

BIAS =

1. RTDs USED - TH = 1 TC = 1 ta.c ELECTRONICS CSA -

l ELECTRONICS SIGMA -

CONTROLLER SIGMA -

CONTROLLER BIAS -

CONTROLLER CSA -

1 ~e ,

• 0 Includes the controller deadband of 1.5 F This calculation was performed assuming that: +a,c 6 WWEERBI i

. - , . _ . _ _ .._c_.._._ . . . . . . ~ . _ _ _ _ _ . _ _ _ _ . _ _ _ _ _ , _ _ . . . _ _ _ _

3. RCS FLOW  !

RTDP ated the plant Technical Specifications require three RCS flow i surveillances: a total RCS flow measurement every fuel cycle (every

, 18 months) that is used to calibrate the RCS flow instrument channels; a monthly total RCS flow measurement; and a qualitative RCS

, flow verification every 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> to maintain RCS flow within the assumed safety analysis value. The 18 month RCS flow surveillance is j satisfied by a precision RCS flow measurement; the monthly RCS flow surveillance is satisfied by a process computer measurement from the RCS flow instrument channels whose calibration is based on the 18 month precision RCS flow measurement; and the 12 hour1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> RCS flow surveillance is satisfied by control board RCS flow indicator readings. No drift is assumed in this error analysis for hot and cold leg RTDs and the feedwater flow transmitters used in the RCS flow measurement procedure. Six (6) month drift effects are included for feedwater temperature, steam pressure and pressurizer pressure measurements. It is also assumed that the calorimetric flow measurement is performed at the beginning of a cycle, i.e., no allowances have been made for mid cycle feedwater venturi fouling, and above 90% RTP.

The RCS 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 tne 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.,

l. WRCS-N(W). L Eq. 4

..- .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:

.g.

WL = 16.110SC - Op._ _1Q L /N)H Vc l (hg - hC) Eq. 5 where, W

L

= Loop flow (gpm)

A = 0.1247 gpm/(ft3 /hr) 0g3

- Steam generator thermal output (Btu /hr)

Qp -

RCP heat addition (Btu /hr)

Qt

- Primary system net heat losses (Btu /hr)

VC

- Specific volume of the cold leg at TC (ft /lb)

N - Number of primary side loops hg - Hot leg enthalpy (Btu /lb) h C

- Cold leg enthalpy (Btu /lb).

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

Osa - (hs - h )Wff Eq. 6 where, h

s

- Steam enthalpy (Btu /lb) hr

- Feedwater enthalpy (Btu /lb)

Wf -

Feedwater flow (1b/hr).

The steam enthalpy is based on measurement of steam generator outlet steam pressure, assuming saturated conditions. The feedwater enthalp based on the measurement of feedwater temperature and steam pressure The feedwater flow is determined by multiple measurements and the following calculation:

W f - (K)(F,)((pf)(d/p))l/2 Eq. 7 where, Wf

- Feedwater loop flow K - Feedwater venturi flow coefficient F, - Feedwater venturi correction for thermal expansion pf - Feedwater density (1b/ft3 )

d/p - Feedwater venturi pressure drop (inches H2 0).

l The feedwater venturi flow coefficient is the product of a number of l

l constants including as built dimensions of the venturi and calibration j tests performed by the vendor. The thermal expansion correction is based l on the coefficient of expansion of the venturi material and the i g difference between feedwater temperature and calibration temperature. j Feedwater density is based on the measurement of feedwater temperature

{

and steam pressure. The venturi pressure drop is obtained from the  ;

output of the differential pressure cell connected to the venturi. I 1

RCP heat addition is determined by calculation, based on the best estimate of coolant flow, pump head, and pump hydraulic efficiency.

The prireary system net heat losses are determined by calculation cor;idering the following system heat inputs and heat losses:

Charging flow Letdown flow Seal injection 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 100% RTP operation is used for these losses or heat inputs.

Tho hot leg and cold leg enthalpies are based on the measurement of the hot leg temperature, cold leg temperature and pressurizer pressure. The cold leg specific volume is based on measurement of the cold leg temperature and pressurizer pressure.

i The RCS flow measurement is thus based on the following plant measurements: 1 1

Steamline pressure (Ps )

o feedwater temperature (Tf )

Feedwater venturi differential pressure (d/p) l

. Hot leg temperature (Tg)

Cold leg temperature (TC)

Pressurizer pressure (Pp )

Steam generator blowdown (if not secured) and on the following calculated values:

Feedwater venturi flow coefficients (K) feedwater venturi thermal expansion correction (F,)

Feedwater density (pf)

Feedwater pressure (Pf)

.Feedwater enthalpy (hr)

Steam enthalpy (h,)

Moisture carryover (impacts hs )

Primary system net heat losses (QL )

'RCP heat addition (Qp)

Het leg enthalpy (hH)

Cold leg enthalpy (hC )*

l These measurements and calculations are presented schematically on Figure 1.

The derivation of the measurement errors and flow uncertainties on Table

. 5 are noted below.

. Secondary 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.

pu-ww .- yr .-.w-& .- m--cs- it- -:tw w a- -w-yev- --

w - -M w- 9 *-r*-

l l

For the measurement of feedwater flow, each feedwater venturi is calibrated by the vendor in a hydraulics laboratory under controlled conditions to an accuracy of [ J+a c. The calibration data which substantiates this accuracy is provided to the plant by the vendor. An edditional uncertainty factor of ( )+a,cis included for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of [ )+a,c. Since the calculated RCS loop flow is related to the calculated steam generator thermal. output which in turn is proportional to feedwater flow, the flow coefficient uncertainty is expressed as [ ]+a,c. It should be noted that no allowance is made for venturi fouling. The venturis should be inspected, and cleaned if necessary, prior to performance of the precision ~ measurement. If fouling is present but not removed, its  ;

- effects must be treated as a flow bias. l l

The uncertaint/ applied to the feedwater venturi thermal expansion i correction (F,) is based on the uncertainties of the measured feedwater ,

temperature and the coefficient of thermal expansion for the venturi i material, usually 304 stainless steel. For this material, a change of i 13.0 F0in the nominal feedwater temperature range changes F, by l 1 0.026 % and the steam generator thermal output by the same amount.

Based on data introduced into the ASME Code, the uncertainty in F, for 304 stainless steel is i 5 %. This results-in an additional uncertainty of [ ]+a c in feedwater flow.

Using the 1967 ASME-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 l- feedwater enthalpy. Feedwater pressure uncertainties impact feedwater

- density and feedwater enthalpy.

i i

. n.v', ,,',.,, -

feedwater venturi d/p uncertainties are converted to % feedwater flow using the following conversion factor:

% flow - (d/p uncertainty)(1/2)(transmitter span /100)2

. The feedwater flow transmitter span is ( )+a,c of nominal flow.

Using the 1967 ASME 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 ( ]+a,c. This value is noted on Table 4.

The net pump heat uncertainty is derived from the combination of the primary system net heat losses and pump heat addition and are summarized for a three loop plant as follows:

System heat losses ,

-2.0 MWt Component conduction and convection losses 1.4 Pump heat adder 110J Net Heat input to RCS + 6.6 MWt.

The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [ ]+a,c of the calculated value. Since direct measurements are not possible, the uncertainty on component cnnduction and convection losses has been assumed to be [ ]+a,c of the calculated value. Reactor coolant pump hydraulics are known to a relatively high confidence level, supported by

. system hydraulics tests performed at Drairie Island II and by input power measurements from several plants; therefore, the uncertainty for the pump heat addition is estimated to be [ ]+a,c of the best estimate value.

Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are less than

( )+a,c of the total, which is [ ]+a,c of core power.

Primary Side The primary side uncertainties are in three principal areas -- hot leg enthalpy, cold leg enthalpy and cold leg specific volume. These are specifically noted on Table 5. Three primary side parameters are actually f measured, Ty and TC and pressurizer pressure. Hot leg enthalpy is influenced by TH, pressurizer pressure and hot leg temperature streaming.

The uncertainties for the instruneentation are noted on Table 3; the sensitivities are provided on Table 4. The hot leg streaniing is split into random and bias (systematic) components. For the Farley units with direct immersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the hot leg streaming uncertainty is ( )+a,c for both random and systematic components.

The cold leg enthalpy and specific volume uncertainties are impacted by Tc and pressurizer pressure. Table 3 notes the TC instrument uncertainty and Table 4 provides the sensitivities.

Noted on Table 5 is the plant specific RTD cross calibration systematic allowance. When necessary, an allowance is made for a systematic temperature error due to the RTD cross-calibration precedure. No allowance was necessary for this plant.

Parameter dependent effects are identified on Table 5. Westinghouse has detertnined 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 3 loop uncertainty equation (with biases) is as follows:

+a,c i l

3 Based on the number of loops, number, type and measurement method of RTDs, and the vessel delta-T, the flow uncertainty for the precision flow calorimetric is:

1. of loops flow uncertainty (% flow)

+a,c 3

e 0

TABLE 3 FLOW CALORlHETRIC INSTRUMENTATION UNCERTAINTIES

(% SPAN) FW TEMP FW PRESS FW d/p STM PRESS TH TC PZR PRESS

- -4ac SCA =

SMTE-SPE -

STE -

SD -

R/E -

RDOUT-BIAS-CSA =

1. OF INST USED 1 1 2 0F (I) psig(2) % d/p(3) psig(l) 0F(4) 0F (4) psig(5)

INST SPAN - 500. 2000. 120% 1200. 120. 120. 800.

INST UNC. -- -

+a,c (RANDOM) -

INST UNC.

(DIAS)

NOMINAL 4370 F 893 psia 100% Flow 793 psia 610.90F 543.50F 2250 psia Notes:(1) Based on permanently installed plant instrumentation and read from the plant computer.

(2) A steam pressure measurement is read from the plant computer and is substituted for a feedwater pressure measurement. A conservative uncertainty value is used.

(3) Heasured with a Barton transmitter. This does not include the venturi uncertainty.

(4) Temperature measured with DVH at the input to Westinghouse process instrumentation using an RTD test rig.

(5) Based on permanently installed plant instrumentation and read from the plant computer. The calculation is based on control board indicators.

These calculationt. were performed assuming that: +a,c 0

1 TABLE 4 FLOW CALOR! METRIC SENSITIVITIES FEEDWATER FLOW TEMPERATURE -

o MATERIAL - ,

DENSITY i TEMPERATURE - J PRESSURE -

l DELTA P -

FEEDWATER ENTHALPY TEMPERATURE -

PRESSURE - .

h -

s 1199.6 BTU /LBM i

- 416.4 BTU /LBM  :

hr Dh(SG) 783.2 BTU /LBM i

STEAM ENTHALPY 43C PRESSURE -

MOISTURE -

HOT LEG ENTHALPY TEMPERATURE -

' PRESSURE -

h = i 629.1 BTU /LBM

'h H -

539.1 BTU /LBM Ok(VESS) - 90.0 BTU /LBM ,

Cp(T) H 1.492 BTU /LBM Of COLD LEG ENTHALPY

+a,c

. TEMPERATURE -

PRESSURE -

Cp(TC ) -

1.228 BTU /LBM Of COLD LEG SPECIFIC YOLUME TEMPERATURE - 1 PRESSURE -

TABLE 5 CALORIMETRIC PCS FLOW MEASUREMENT UNCERTAINTIES (Page 1 of 2)

COMPONENT INSTRUMENT ERROR FLOW UNCERTAlHTY FEEDWATER FLOW +a,c VENTURI THERKAL EXPANSION COEFFICIENT TEMPERATURE KATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION HOT LEG ENTHALPY TEMPERATURE STREAMING, RANDOM STREAMING, SYSTEKATIC PRESSURE COLD LEG ENTHALPY TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPELATURE PRESSURE RTD CROSS-CAL SYSTEKATIC ALLOWANCE

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

l TABLES (CONTINUED)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES l (Page 2 of 2)

' COMPONENT FLOW UNCERTAINTY BIAS VALUES -- --

+a c FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRESSURE ENTHALPY HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME - COLD LEG l FLOW BIAS TOTAL VALUE l

+a,C I SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITHOUT BIAS VALUES) )

N LOOP UNCERTAINTY (WITH BIAS VALUES)

J l

I l

l I

l 1

h 4

n -

The precision flow calorimetric may be used as a reference for the normalization of the cold leg elbow taps. Table 6 notes the instrument uncertainties for normalization of the elbow taps, assuming one elbow tap per loop. The d/p transmitter uncertainties are converted to % flow on the same basis as the feedwater venturi d/p. The elbow tap uncertainty is then combined with the precision flow calorimetric uncertainty. This combination of uncertainties results in the following total flow uncertainty :

1. of loops flow uncertainty (% flow) 3 2.3.

The corresponding value used in R10P is:

1. of loops standard deviation (% flow)

+a,c 3

l l

TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES

% d/p SPAN  % FLOW

+a,c PMA =

PEA -

SCA -

SPE -

STE -

SD -

RCA -

RMTE-RTE -

RD =

ID =

A/D -

RDOUT-BIAS =

FLOW CALORIM. BIAS FLOW CALORIMETRIC -

..NSTRUMENT SPAN -

+ac SINGLE LOOP ELBOW TAP FLOW UNC -

N LOOP ELBOW TAP FLOW UNC N LOOP RCS FLOW UNCERTAINTY (WITHOUT BIAS VALUES) --

N LOOP RCS FLOW UNCERIAINTY (WITH BIAS VALUES)

- 2.3

. 22-

b

4. REACTOR POWER Generally a plant performs 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 I of the Operating License and to adjust the Power Range Neutron Flux  !

3 channels when the difference between the NIS and the heat balance is l greater than that required by the plant Technical Specifications.

Assuming that the primary and secondary sides are in equilibriu.n, the core power is determined by summing 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 j the primary side system losses, and' dividing by the core rated Btu /hr i at full power. The equation for this calculation is:

RP = (fN)(O SC -O p M / LNill (100)

H Eq. 8 where, RP - Core power (% RTP)

N - Number of primary side loops ,

Q3g- - Steam generator thermal output (BTU /hr) as defined in Eq. 6 Op - RCP heat adder (Btu /hr) as defined in Eq. 5 Qt

-- Primary system r,et heat losses (Btu /hr) as defined in Eq. 5 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 calorimetric equations and effects are the same as those noted for' the precision flow calorimetric (secondary side ,

23

a u ~ c.npw n .

portion), equations 6 and 7. Table 7 provides the lastrument uncertainties for those measurements performed. Since it is necessary to make this determination daily, it has been assumed that the plant process computer will be used for the measurements. The sensitivities calculated are the same as those noted for the secondary side on Tabic

4. As noted on Table 8, Westinghouse has determined the dependent sets

, in the calculation and the direction of interaction. This is the same as that performed for the RCS flow calorimetric, but .pplicable only to power. The same was performed for the bias values noted. It should be noted that Westinghouse does not include any allowance for feedwater venturi fouling. The effect of fouling is to result in an indicated power higher than actual, which is conservative.

Using the power uncertainty values noted on Table 8, the 3 loop uncertainty (with bias values) equation is as follows:

-- +a,C Based on the number of loops and the instrument uncertainties for the four parameters of feedwater temperature, feedwater pressure, feedwater flow and steam pressure, the power measurement uncertainty for the secondary side power calorimetric is:

1. of loops power uncerthinty (% RTP)

+a,c 3

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TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES 2 (% SP?N) FW TEMP FW PRES 3 FW d/p STM PRESS

+a,c SCA =

SMTE-SPE -

STE -

SD -

BIAS-RCA -

RMTE-RTE =

RD =

ID -

A/D -

CSA - j OF (I) psig(2) % d/p(3) psig(I)

INST SPAN - 500. 2000. 120% Flow 1200.

INST UNC -r --

+a,c (RANDOM) -

INST UNC (BIAS)

NOMINAL 541F 0 844 psia 100% 744 psia Notes:

(1) Based on permanently installed plant instrumentation and read from the plant computer.

(2) A steam pressure measurement is substituted for a feedwater pressure measurement. A conservative uncertainty value is used.

(3) Based on permanently installed plant instrumentation (protection channel feedwater flow) and read from the plant computer.

Normally the daily power calorimetric measurement is performed by reading the local Barton d/p transmitter. The feedwater flow channel represents a conservative uncertainty calculation.

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TA3LE 8 SECONDARY SIDE POWER CALORIMET'IC MEASUREMENT UNCERTAINTIES COMPONENT h.sTRUMENT ERROR POWER UNCERTAINTY

+a,c FEEDWATER FLOW VENTURI THERMAL EXPANSION C0FFFICIENT TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESFJRE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION BIAS VALUES -

FEEDWATER DELTA P FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY POWER BIAS TOTAL VALUE

• ** INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOU1 BIAS VALUES)

N LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES)

S 4

IV. CONCLUSIONS The preceding sections provide the methodology for what Westinghouse believes is a reasonable means of accounting for instrument uncertainties for pressure, temperature, power and flow. The plant-specific instrumentation has been reviewed for Farley Nuclear

- Plant Units 1 and 2, and the uncertainty calculations are completed.

These uncertainty values or more conservative values are used in the RTDP analysis.

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REFERENCES

1. Westinghouss letter NS CE-1583, C. Eicheidinger to J. F. Stolz, NRC, dated 10/25/77.

a

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

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

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

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

6. NRC letter, S. A. Varga to J. Dolan, Indiana and Michigan Electric Company, dated ?c12/81.-

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", dated 2/86.

L

, 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.

V

10. ANSI /ANS_ Standard 59.4-1979, " Criteria for Technical Specifications for Nuclear Power Station:3".

11. ISA Standard S67.04, 1982, "Setpoints for Nuclear Safety-Related Instrumentation Used in Nuclear Power Plants"

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i

12. Tuley, C. R., Miller, R. B., " Westinghouse Setpoint Methodology for Control and Protection Systems", IEEE Transactions on Nuclear Science, February,

.i

-1986, Vol NS-33 No. 1, pp. 684-687.-

13. Scientific Apparatus Manufacturers Association, Standard PMC 20.-1, 1973,

" Process Measurement and Control Terminology".

14. Westinghouse.WCAP-11397-P A, " Revised Thermal Design Procedure", dated "

April,.1989.

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- .. .-_ _ ._. - - . . - . - . _ . .. ~ . _ - - _ - ,

FIGURE 1 RCS FLOW CALORIMETRIC SCHDMTIC

(- l OTC *l. I PpOl l OTH *l {Ol Ps Pf {O Tf I

{%

Ap hC hH hs hf pf Fa K

/

P

% %G  :

4-4 -

.E -  % m.s no a C.lculated O H

_ Mass 4.

. qr

-- I'

• vc  :

4 Other Loops

• One temperature per loop is measured ~

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. RCS Volumetric Flow l

FIGURE 2 POWER CALOR! METRIC SCHEMATIC l l P, I P T j AP 7

Yl i l f

l ~

(  ; '

l h h 3 f pf F, K 1r W

7

'r O - measured 0 - calculated sa c

e Other Loops 4+

OL Op s

, Core Power

~

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Attachment 6 l Joseph M. Farley Nuclear Plant Units 1 and 2 i Request for Technical Specifications Changes l VANTAGE-5 Fuel Design l

" Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Alabama Power Farley Nuclear Plant Units 1 & 2 (for RTD Bvoass Elimination)"

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