ML20070H301

Nonproprietary Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Vogtle 1 & 2 Nuclear Power Station (for RTD Bypass Loops)

ML20070H301
Person / Time
Site:	Vogtle
Issue date:	12/31/1989
From:	Moomau W WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19310D315	List: ELV-02597, Forwards Proprietary WCAP-12460 & Nonproprietary WCAP-12461, Westinghouse Revised Thermal...(For RTD Bypass Loops) & Proprietary WCAP-12462 & Nonproprietary WCAP-12463, Westinghouse Revised Thermal Design Procedure.. ML20070H271 ML20070H289 ML20070H301 ML20070H305
References
WCAP-12461, NUDOCS 9103140030
Download: ML20070H301 (33)

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-WESTINGHOUSE CLASS'3-WCAP 12461 F y WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY -FOR GEORGIA POWER ^ V0GTLE l'AND 2 NUCLEAR POWER STATION (FOR RTO BYPASS. LOOPS) l w p. av . DECEMBER,11989 W.H.Moomau l \\ - i e a ifb - _ 1 -:o Westinghouse Electric Corporation-Energy Systems P.O. Box 355 Pittsburgh,-Pennsylvania 15230' Copyright.tur Westinghouse Electric '1989,_ c-All Rights Reserved: . I i u "ai n i s' __mm.-.___m. ..m.__--__._m__ __o

TABLE OF CONTENTS SECTION TITLE PAGE I. Introduction 1 II. Methodology 2 III. Instrumentation Uncertainties 4 IV. Conclusions 22 References 26 o. k i

a LIST OF TABLES TABLE NUMBER TITLE PAGE 1 Pressurizer Pressure Control 5 System Accuracy 2 Rod Control System Accuracy 7 3 Flow Calorimetric Instrumentation 15 Uncertainties 4 Flow Calorimetric Sensitivities 16 5 Calorimetric RCS Flow Measurement 17 Uncertainties '6 Cold leg Elbow Tap Flow Uncertainty 20 7-Power Calorimetric Instrumentation 24 Uncertainties 8 Secondary Side Power Calorimetric 25 Measurement Uncertainties 1 ii l

LIST OF ILLUSTRATIONS . FIGURE NVMBER TITLE PAGE i 1 RCS Flow Calorimetric Schematic 28 2 Power Calorimetric Schematic 29 T 111

WESTINGHOUSE REVISED THERMAL DESIGN PROCEDVRE INSTRUMENT UNCERTAINTY METHODOLOGY FOR GEORGIA POWER V0GTLE 1 AND 2 NVCLEAR POWER STATION (FOR RTD BYPASS LOOPS) 1. INTRODUCTION Four operating parameter uncertainties are used in the uncertainty analysis of the Revised Thermal Design Procedure (RTDP). These parameters are Pressurizer Pressure, Primary Coolant Temperature (Tavg), 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 calorimetric) once every 24 hours. RCS flow is monitored by the performance of a precision flow calorimetric at the beginning of each cycle. Pressurizer pressure is a controlled parameter and the uncertainty reflects the control system. T is a controlled parameter via the temperature input to the rod avg 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. 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",(l'2'3) 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 core 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 for both .RTDP parameters and protection functions. The uncertainty calculations in this report are based on a detailed review 1

of P_lant Vogtle procedures for instrument 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 PROTEUS

. plant-computer heat balance used for the daily nuclear instrumentation alignment surveillance. 11. 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 sum of both sides is equal to the - range for that parameter, e.g., Rack Drift is typically [' ]+a,c, the range for this parameter is [ ]+a,c, -This technique has been utilized before as ne ed above, and has been -endorsed by the NRC staff (6,7,8,9) and various industry standards (10,ll), Ths relationships between the error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint Meth'odology(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+ (RD0VT)2)l/2 Eq.1 + BIAS 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)1/2 + BIAS Eq. 3.

-~. nere: Channel Allowance CSA = 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 = Rack Calibration Accuracy RCA = RMTE Rack Measurement and Test Equipment Accuracy = Rack Temperature Effects . RTE = RD ' Rack Drift = Readout Device Accuracy (DVM or gauge) RDOUT. = ID. Computer-Isolator Drift . Analog-to Digital Conversion Accuracy -A/D = Controller Accuracy CA- -= The_ parameters above are as defined in references 5 and-12 and are b a d on SAMA Standard PMC 20.1,-1973(13). 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, SCA---

reference (calibfation) 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, lSD: 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. RTE - change in input-output _ relationship due to a change in ambient temperature for the rack modules, -

i changefin~ input-output relationship _over a period of time RD. at' reference conditions for the rack modules, RD0VT. the measurement accuracy of. a special test local gauge, digital-voltmeter or multimeter on-it's most accurate i applicable range for the parameter measu-ed, change.in input-output relationship over a period of timt ID at-reference conditions fer a control / protection signal isolating device, allowance for conversion accuracy of an analog signal to a A/D digital signal for process-computer use, allowance for the accuracy of a controller, not including CA= .deadband. BIAS - a non-random uncertainty for a sensor / transmitter or a process parameter. A more detailed explanation of the West'inghouse methodology noting the 'interaction of several para;neters 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 Tayg (through Rod Control).- 1. ' PRE 3SUR17ER PRESSURE ' 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 racks / controller. As noted on Table 1, the electronics uncertainty for _this function is-[ ]+a,c which corresponds to an accuracy of [ ]+a,c. Inl 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 allowance of l[ ']+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)..

I 1 s TABLE 1= PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY +a,c. 'SCA = SMTE= STE = SD ;- BIAS =~ RCA. RMTE=- . RTE = l RD '- CA ' - +a,c . ELECTRONICS UNCERTAINTY = PLUS- . ELECTRONICS UNCERTAINTY =. PLUS ECONTROLLER UNCERTAINTY- = I + 4 q

bi 32.

TAVG-f T' avg is controlled by a system that compares the auctioneered high

'T from the loops with a reference, usually derived from the First avg Stage Turbine Impulse Chamber Pressure. T is the average of the avg narrow' range _Ty and TC values. The highest loop T is then used avg 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 Tg and 1 TC-cross-calibrated RdF RTDs are used to calculate T and the RTDs are avg . located in the RTD bypass manifold, the CSA for the electronics is [ ]+a,c. Assuming a normal, two sided probability l - distribution results in an electronics standard deviation (si) of [_ -]+a,c, i 'However,'th'is does not include the controller deadband of i 1.5_O. F The controller accuracy is the combination of the instrumentation

accuracy.and the.deadband. The probability distribution for the deadband has been' determined to be [

].+a,c The variance 'for the deadband uncertainty is then: (s2)

• I 3+ ' -

Combining-the variance for instrumentation and 'deadband results in a controller variance of: (sT) --(s))2 + (s2) " (- l "' ll p. The controller sT - [ ]+ ' for a total. uncertainty of l [- ]+a,c, l l

i..

1 Y 1 'k. i 1 TABLE 2' R0D CONTROL SYSTEM ACCURACY + Tavg TURB' PRES j u +a,c-PMA = --SCA = LSMTE-0 STE'-'- l' SD =- --BIAS =: RCA-='- RMTE-- ,a' RMTE-RTE = ? RD = CA = BIAS =

1. RTDs USE0 -:'TH = 1 TC = 1

+a,c ELECTRONICS CSA = ELECTRONICS SIGMA = CONTROLLER. SIGMA = -- . CONTROLLER-BIAS = CONTROLLER CSA = '{ s [. :' J 4' ,f.. 5 I L 4 i V' l r l a.

g-1 i t3. -RCS FLOW-RTDP and some plant Technical Specifications require an RCS-flow measurement with a high degree of accuracy. It is assumed for this error analysis that the flow measurement ;s performed within thirty days of calibrating the measurement instrumentation. Therefore, -except -where necessary due to sensor location, drift effects are not = included. It is also assumed that the calorimetric flow measurement is performed at the beginning of a cycle, i.e., no allowances have j -been made for Feecwater venturi fouling, and above 70% RTP. 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 inLequilibrium,= the RCS total vessel flow is the sum of the l individual primary loop flows, i.e., WRCS - N(W ). Eq. 4 1 L '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 ~ ent.halpy rise and multiplying by the Cold Leg speci fic volume. The equation for this calculation'is: (d /EjliyCI-Wt = (A)(OSC - O 4 p t (hH-h) Eq. 5 C where; p L ' Loop flow (gpm) -W' 3 -A 0.1247gpm/(ft/hr) Q33 ' Steam Generator thermal output (Btu /hr) Qp RCP heat addition (8tu/hr) Qt Primary system net heat losses (Btu /hr) 3 V-C Spec (fic volume of the Cold Leg at TC (ft /lb)

Number of primary side loops N =

h-Hot Leg enthalpy' (Btu /lb) g hC 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: Qsc = (hs - h )W7 Eq. 6 -f where; h Steam enthalpy-(Btu /lb) 3 ' hr. Feedwaterenthopy(Btu /lb) f Feedwater flow ob/hr). W = The' Steam enthalpy is based on measurement of Steam Generater outlet Steam pressure, assuming saturated-coaditions. The Feedwater enthalpy is based on the measurement of feedwater temperature and a Feedwater pressure of 1000 psia. The Feedwater flow is determinec by multiple measurements and the following calculation: Wf = (K)(F )((P )(d/p))l/2 Eq. 7 a f Feedwater venturi' flow coefficient

where; :K

= .Feedwater venturi correction for thermal expansion-F = a 3 _ pf: Feedwater density (1b/ft ) Feedwater venturi pressure drop (inches H 0). d/p =- 2 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 tb difference 'between Feedwater temperature and calibration ~ temperature. Feedwater density _is based on the measurement of Feedwater temperature and 'a Feedwater pressure of 1000 psia. The venttri pressure drop is obtained from the output of'the differcatial prc.>sure _ cell connected to 'the venturi, RCP heat addition is determined by calculation, based on the best . estimate of coolant flow, pump head, and pump hydraulic efficiency. J \\

L j The primary system net heat losses are determined by calculation, considering the following s.vthe heat inputs and heat losses: Charging flow Letdown flow Seal injection flow RCP thermal barrier cooler heat removal 4 Pressurizer spray flow i Pressurizer surge 1tne-flow Coronent insulation heat losses - l Corronent support heat los:ce-CRDM heat losses. ~ A single ca::,: lated sum for 100% RTP operation is t: 2d for these losses [ ar heat inputs. Tie Hot Leg and Cold leg enthalples are based on the measurement of the Hot Leg temperature,, Cold Leg temperature and a Pressurizer pressure of 2250 psia. The Cold Leg specific volume is based on measurement of the Cold leg temperature and a Pressurizer pressure of 2250 psia. The RCS flow measurement is thus based on the followinn plant meat,urements: Steamlinepressure-(P) s Feedwatertemperature(T) f Feedwater venturi differential pressure (d/p)- Hot Leg temperature (T ) H ColdLegtcmperature(Tc) ' Steam Generator blowdown (if not-secured)- h and.on the following calculated values: Feedwater venturi flow coefficients (K) -l Feedwater venturi thermal expansion correction (F ) a Feedwater density (pf) Feedwater pressure (P )- f_ Pressurizer pressura (P ) p, l - .-.a _,_ e

[ i Feedwater enthalpy (hr) Steam enthalpy (h ) s Moitturecarryover(impactsh) i s Pr m y system net heat losses (Q ) L RCPheat-addition (Q) p Hot Leg ent!.alpy (h ) H Cold Leg enthalpy (h )* C These measurements and calculations are presented schematically on Figure l. i The derivation of the measurement errors and flow uncertainties on Table 5 are noted below.- Secondarv Side i ~The secondary side uncertainties are in ur 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 the vendor in a hydraulics laboratory under controlled conditions-to an accuracy.of ( ]+a,b,c. The calibration data which subriantiates this accuracy is provided to the plant by the vendor. An additional 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, it's effects must' be treated as a flow bias. The uncertainty gplied to the feedwater venturi thermal expansion correction (F ) is based on the uncertainties of the measured Feedwater a L temperature.and the coefficient of thermal expansion for the venturi 11 ,.r-, _s .,...r.m,m4 ,m._, m,,n., ..-..-,._myn ,,,,n. m w

.. -. - - _ ~.. ~ material, usually 304 stainless steel. For this material, a change of 0 1 2.0 F in the nominal feedwater temperature range changes F, by i 0.004 % and the Steam Generator thermal output by the same amount. l Based on data introduced into the ASME Code, the uncertainty in F for a 304 stainless steel is i 5 %. This results in an additional uncertainty of ( )+a,c in feedwater flow. Westinghou e uses the conservativevalueof( ]+a,c, Using the 1967 ASME Steam Tables it is possible to determine the sensitivities of various parameters to changes in Feedwater temperature ano + 7ssure. Table 3 notes the instrument uncertainties for the I hardware used to perform the measurements. Table 4 lists the various sensitivitie. As can be seen on Table 4, feedwater temperature uncertainties have n impact on venturi f, Feedwater density and a feedwater enthalpy. Feedwater pressure uncertainties impact feedwater density and Feedwater enthalpy. Feedwater venturi d/p uncertainties are converted to % Feedwater flow using the following conversion factor: l % flow - (d/p uncertainty)(1/2)(transmitter span /100)2 Typically, the Feednter flow transmitter span is [ ]+a,c or 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 addi; ion and are summarized for a four loop plant as follows: i t l

System heat losses -2.0 MWt Component conduction and convection losses -1.4 Pump heat adder +14.0 Net Heat input to RCS +10.6 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 conduction and convection losses has been assumed tobe[ J+a,c of the calculated value. Reactor coolant pump hydraulics are known to a relatively high confidence level, supported by

ystem hydraulics tests performed at Prairie Island 11 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 ona quantity, which is designatt.1 the net pump heat uncertainty, the combined uncertainties are less than [ ]+a,c of the total, which is [ )+ac of core power. Primary Side The primary side uncertainties are in three principal areas, Wot Leg enthalpy, Cold leg enthalpy and Cold Leg specific volume. These are specifically noted on Table 5. Two primary side parameters are actually maasured TH and T, and a Pressurizt.r pressure of 2250 psia is used to C determine enthalpies. Hot leg enthalpy is influenced by Tg, Pressurizer pressure and Hot leg temperature streamtog. The uncertainties for the instrumentation are noted on Table 3, the sensitivities are provided on Table 4. The Hot Leg streaming is split into random and bias (systematic) components. For the Vogtle units with direct immersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the streaming uncertainty is [ )+a,c for both random and systematic components. -11 ~

r 1 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.

I 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 procedure. No allowance was necessary for this plant. 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 Tabit 5, the 4 loop uncertainty equation (with bietes) is as follows: +a,c . 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 ist l ( of loops: flow uncertainty (% flow) +a,e 4 f.

TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION UNCERTAINTIES e PRZ PRESS (% SPAN) FW TEMP FW PRES FW d/p STM PRESS TH TC +a,c SCA = SMTE-SPE = STE = SD = R/E = RDOUT= 1 BIAS = CSA - 1

1. OF INST USED 1

1 1 0 (5) op(5) psia (6) 0 (l) psia (E) %d/p(3) psia (4) F F INST SPAN = 500. 1500. 120. 1500. 120. 120. 800. INST UNC. +a,c (RANDOM) = INST UNC. (BIAS) = NOMINAL: -440..

1100, 1000.

618.2 558.9 2250. Notes:- (1) Final feedwater temperature from plant _ computer. A conservative 0F is used. uncertainty value of.2.0 (2) Assumed constant in precision heat balance (not measured). A conservative value is used. 3 . (3) Measured with test d/p gauge. Does not include venturi uncertainty. l (4) Measured with' test' pressure-gauge. (5) Uncertainty assumes temperature measured with DVM st output of R/E-converters. (6) Assumed constant in precision heat balance. The values are based on permanently installed plant instrumentation. 1

i - ( TABLE 4 FLOW CALORIMETRIC SENSITIVITIES -- FEEDWATER FLOW F, +a c TEMPERATURE l = MATERIAL-DENSITY - TEMPERATURE PRESSURE = DELTA P- = - FEEDWATER ENTHALPY TEMPERATURE = PRESSURE = h 1192.9 BTV/LBM = s hr-419.5 BTU /LBM- = Dh(SG) T/3.4 D10/LBM c - STEAM ENTHALPY +a,c PRESSURE = MOISTURE- - HOT LEG ENTHALPY TEMPERATURE PRESSURE e hH 640.2 BTV/LBM = bc 558.3 BTV/LBM = DH(VESS) '81.9 BTV/LBM = 1.548. BTU /LBM 0F Cp(Tg): = COLD' LEG ENTHALPY ~ +a,c TEMPERATURE = . PRESSURE Cp(T) - l.266 BTV/LBM OF C . COLD LEG SPECIFIC VOLUME +a,c TEMPERATURE PRESSURE l' q 5. 1 (- 's w + w, g-w+ my,- w-we -v y, w-v=*w-.ve---<frw-1-+cr=vw- .g ' e w-meseree =w n.- -mafww---+-w-o%vb,*- er v u- -w +w,-

TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT INSTRUMENT ERROR FLOW UNCERTAINTY FEEDWATER' FLOW 4a,c VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL-DENSITY-TEMPERATURE PRESSURE DELTA P 1 FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE

MOISTURE i

NET PUMP HEAT ADDITION HOT LEG ENTHALPY-TEMPERATURE i STREAMING, RANDOM-STREAMING, SYSTEMATIC -PRESSURE COLD LEG ENTHALPY TEMPERATURE' PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE- -1 RTO CROSS-CAL' SYSTEMATIC ALLOWANCE , +, - ++ INDICATE SETS OF DEPENDENT PARAMETERS. L. L l-I._.._._.._.2..

z..

TABLE S (CONTINUED) CALORIMETRIC RCS FLOW HEASUREMENT UNCERTAINTIES COMPONENT FLOW UNCERTAINTY BIAS VALU;S +a,c FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRES $URE ENTHALPY - HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME COLD LEG FLOW BIAS TOTAL VALUE + 8, C SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES! N LOOP UNCERTAINTY (WITHOUT BIAS VALUES) N LOOP UNCERTAINTY (WITH BIAS VALUES) 9 l l* --

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 i 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 loups flow uncertainty (% flow) 4 2.2 The corresponding value used in RTDP is:
2. of loops standard deviation (% flow)

+ a, c 4 l l 19 l

TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES j l % d/p SPAN % FLOW +a c i= PHA = i PEA = SCA = +- SPE = i STE = 50 = RCA a RMTE= RTE = RD = ID = A/D = RDOUT= BI AS= -- FLOW CALORIM' 81AS: FLOW CALORIM'TRIC a INSTRUMENT SPAN. +a,c SINGLE LOOF ELBOW TAP FLOW UNC = N LOOP ELBOW TAP FLOW UNC .a N LOOP'RCS FLOW UNCERTt.INTY~ (WITHOUTBIASVALUE5) = N' LOOP RCS FLOW-UNC'cRTAINTY 2.2-(WITH BIAS VALUES)- = L i r _

1 ~4 Reactor Power Generally'a plant performs a primary / secondary side heat balance once every 24 hours 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 equilibrium; 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 the p'rimary side system losses, and dividing by the core rated 6tu/hr at full power. The equation for this calculation is: f RP = (fN)(OSC - O M /N))1(100) p t H Eq. 8 t [ where; Core power (% RTP) RP = N ,= Number of primary side loops 3 Steam Generator thermal output (BTU /hr) as defined in Qg = Eq. 6 RCP heat adder (Btu /hr) as defined in Eq. 5 Q = p -QL Primary system net heat losses (Btu /hr) as defined in = Eq. 5 Core rated Btu /hr at full power. H-For the: purposes of this uncertainty analysis (and barea on H noted 'above) it 1s assumed that the plant is at 100% RTP wi.an the measurement is taken.-- lHeasurements 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. l The secondary sido power calorimetric equations and effects are the l. same-as' those noted'for the precision flow calorimetric (secondary side l ~

portion), equations 6 and 7. Table 7 provides the instrument uncertainties for those measurements performed. Since it is necessary .to make this determination daily, it has been assumed that the plant j -process computer will be 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 1 in the calculation and the direction of interaction. This is the s me 'l _as that performed for the RCS flow calorimetric, but applicable only to 4 power. The same was performed for the bias values..sted, it s'ould be 1 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 it, conservative, c 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 the number of loops and the instrument uncertainties for the -four parameters, the power measurement uncertainty for the secondary side power calorinietric_is:

1. of loops power uncertainty (% RTP)

+a,e 4 IV. (.QRCLUSIONS The preceding sections provide the methodology for what Westinghouse believes is _ a reasonable means of accounting for instrument h I l. -.

uncertainties for pressure, temperature, power and flow. The plant-specific instrumentation has been reviewed for Vogtle Units 1 and 2 and the uncertainty calculations are completed. These uncertainty values or more conservative values are used in the RTDP analysis. l 1.

i TABLE 7 POWER CALORIMCTRIC INSTRUMENTATION UNCERTAINTIES (f. SPAN) FW TEMP FW PRES FW d/p STM PRESS -+a,c SMTE-SPE - STE - SD BIAS = RCA - RMTE-RTE - RD ID A/D - CSA.- 0F(I) psia (2) 7. d/p(3) psla(3) INST SPAN - 500. 1500. 120. 1200. INST UNC +a,c (RANDOM) - INST UNC (BIAS) NOMINAL - 440. 1100. 1000. Notes: (1) Final feedwater tempr ature from plant computer. A conservative uncertainty value o' 20 0F is used. (2) Assumed constant in,econdary side heat balance (not measured). A conservative value is used. (3) Based on perranently installed plant instrumentation. l .-.. ~ _ TABLE 8 SECONDARY SIDE POWER CAL 0kIMETRIC MEASUREMENT UNCERTAINTIES 1 COMPONENT INSTRUMENT ERROR POWER UNCERTAINTY +8,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 : BIAS. VALUES FEEDWATER DELTA P FEEDWATER PRESSURE DENSITY -ENTHALPY-STEAM PRESSURE-ENTHALPY POWER BIAS TOTAL VALUE

• ** : INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP llNCERTAINTY.(WITHOUT BIAS VALUES)

Ni LOOP UNCERTAINTY (WITHOUT BIAS VALUES)_ N. LOOP UNCERTAINTY (WITH BIAS VALUES) l t E .. _... -.. ~., _ _. i

I l REFERENCES 1. Westinghouse letter NS CE-1583, C. Eicheldinger to J. F. Stolz, NRC, dated 10/25/77. 2. Westinghouse 'e NS-PLC-5111 T. M. Anderson to E. Case, NRC, dated 5/30/78. 3. Westinghouse letter NS-TMA-1837, T. M. Anaerson to S. Varga, NRC, dated 6/23/78. 4. Westinghouse letter NS-EPR-2577, E. P. Rahe Jr. to C. H. Berlincer, NRC, dated 3/31/82. 5. Westinghouse 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 2/12/81. 7. NVREG-0717 Supplement No. 4, Safety tvaluation 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. 9. NUREG/CR-3659 (PNL-4973), "A Mathematical Model for Assessing the Uncertainties of Instrumentation Measurements for Power and Flow of PWR Reactort",2/85.

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

11. ISA Standard S67.04,1982. "Setpoints for Nuclear Safety-Related Instrumentation Used in Nuclear Power Plants" 26-

12. Tuley, C. R.,

Miller, R. B., ' Westinghouse Setpoint Methodology for Control and Protection Systems", IEEE Transactions on Nuclear Science, February, 1986, Vol. NS 33 No. 1, pp. 684-687.

13. Scientific Apparatus Manufacturers Association, Standard PMC 20.1, 1973,

" Process Measurement and Control Terminology". l I.

FIGURE 1 RCS FLOW CALOR!t1ETRIC SCHEMATIC T P T P P T C p g 3 f f 3p // h h h, h pf F K C H f a Ah z Wf U 0g 4 3 V tieasured Q g, g { L p F ~'"" i Calculated u Mass m. W t 1 P y Vol. W L 1 l V =c Other Loons + l 0 One hot leo temperature per loop. V RCS Volumetric Flow 28

f1GURE 2 POWER CALOR! METRIC SCHEMATIC i l P T l { j f f / h h P F K 3 f p a U W f calculated 03g O measured V Other Loops = i + I I l l+ l E c- [0, g = l i Core Power n L i ,}}