ML20092A738

Nonproprietary Westinghouse Improved Thermal Design Procedure Instrument Uncertainity Methodology for CP&L Shearon Harris Nuclear Power Station

ML20092A738
Person / Time
Site:	Harris
Issue date:	09/30/1989
From:	Moomau W WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML18010A520	List: ML18010A519 ML20092A737 ML20092A738
References
WCAP-12341, NUDOCS 9202100266
Download: ML20092A738 (35)

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WESTINGHOUSC CLASS 3 WCAP 12341 a

WESTINGHOUSE IMPROVED THERMAL DESIGN PROCEDURE INSTRUMENT UNCER1AINTY HETHODOLOGY FOR CAROLINA POWER AND l.lGHT SHEARON HARRls NUCLEAR POWER STATION SEPTEMBER, 1989 W.H.Hoomau Westinghouse Electric Corporation Energy Systems P.O. Box 355 Pittsburgh, Pennsylvania 15230 Copyright by Westinghouse Electric 1989, C All Rights Reserved

TABLE OF CONTENTS SECTION TITLE PAGE I.

Introduction I

tti

. Methodology 2

Instrumentation Uncertainties 4

IV.

Conclusions 22 References 26 4

l t

l i

t

LIST OF TABLES i

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 Colorimetric Sensitivities 16 5

Calorimetric RCS Flow Hoasurement 17 Uncertainties

'6 Cold leg Elbow Tap Flow Uncertainty 20 7

Power Calorimetric Instrumentation 24 Uncertainties 8

Secondary Side Power Calorimetric 25 Measurement Uncertainties I

s H

l

-M ii

LIS1 Of IllVSTRATIONS FIGURE NUMBER TITLE PAGE i

RCS flow Calorimetric Schematic 28 s

2 Power Calorimetric Schematic 29 i

i I

I F

h I

L t

iii

WESilNGHOUSE IMPROVED THERMAL DE51GN PROCEDURE INSTRUNENT UNCERTAIN 1Y METHODOLDGY FOR CAROLINA POWER AND LIGHT SHEARON HARRIS NUCLEAR POWER STATION

+

1.

IN1RDDUCT10N four operating parameter uncertainties are used in the uncertainty analysis of the Improved Thermal Design Procedure (ITDP). These parametersarePressurizerPressure,PrimaryCoolantTemperature(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 calorimetric)onceevery24 hours.

RCS flow is monitored by the performance of a precision flow calorimetric at the beginning of each cycle. The RCS Cold leg elbow taps are normalized against the precision calorimetric and used for monthy surveillance (with a small increase in uncertainty).

Pressurizer pressure is a controlled parameter and the uncertainty reflects the control system. T is a controlled parameter avg via the temperature input to the rod control system and the uncertainty reflects this control system.

Westinghouse has been involved with_the development of several techniques to treat instrumentation uncertainties. Anearlyversion(ford.C. Cook 2 and Trojan) used the methodology outlined in WCAP 8567 " improved Thermal DesignProcedure",(1,2,3)whichisbasedontheconservativeassumption 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.:twosidedprobabilitydistributions.(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 Enow utilized for the deterniination of all instrumentation errors for both

!TDP parameters and protection functions.

.l.

HE1HODOLOGY 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 arithmetica11y 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 noted above, and has been endorsedbytheNRCstaff(6.7,8,9)andvariousindustry standards (10,ll),

The relationships between the error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoir.t Methodology (12)andaredefinedasfollows:

1.

For precision parameter indication using Special Test Equipment or a DVH at the input to the racks CSA - ((SCA + SMTE + 50)2 + (SPE)2 + (STE)2+ (RD0VT)2)l/2

+ BIAS Eq. 1 2.

For parameter indication utillring the plant process computer; CSA 2 ((SCA + SMTE + 50)2 + (SPE)2 + (STE)2 + (RCA + RH1E + RD)2

+ (RTE)2 + (10)2 + (A/D)2)1/2 + BIAS Eq. 2 3.

for parameters which have control systems 6 CSA = ((PKA)2 + (PEA)2 + (SCA + SMTE + SD)2 + (SPE)2 + (STE)2

+ (RCA + RHTE 4 RD + CA)3 + (RTE)2)1/2 + BIAS Eq. 3 2-

where Channel Allowance CSA Process Heasurement Accuracy PHA Primary Element Accuracy PEA Sensor Calibration Accuracy SCA Sensor Hessurement and Test Equipment Accuracy SHIC Sensor Pressure Effects SPE STE Sensor Temperature Effects

$D Sensor Drift Rack Calibration Accuracy RCA Rack Heasurement and lest Equipment Accuracy RHTE Rack Temperature [ffects RTE RD Rack Drift Readout Device Accuracy (DVH or gauge)

RDOUT Computer Isolator Drift 10 Analog to Digital Conversion Accuracy A/D Controller Accuracy CA The parameters above are as defined in references 5 and 12 and are based on SAKA Standard PHC 20.1,1973(13). However, for case in understanding they are paraphrased below:

PHA -

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 (calibration) accuracy for a sensor / transmitter, SPE -

change in ir.put-output relationship due to a change in O

static pressure for a d/p cell, STE -

change in input output relationship due to a change in p

ambient temperature for a sensor / transmitter, change in input output relationship over a period of time SD at reference ennditions for a senscr/ transmitter, RCA -

reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string g

calibrated, or tuned, to this accuracy.

RTE -

change in input output relationship due to a change in ambient temperature for the rack modules, 3-l

change in input output relationship over a period of time RD at reference conditions for the rack modules.

RDOUT -

the measurement accuracy of a special test local gauge, digital voltmeter or multimeter on it's most accurate applicable range for the parameter measured, change in' input output relationship over a period of time ID at reference conditions for 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, na' including CA 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.

111.

Instrumentation Uncertainties The instrumentation uncertainties will be discussed first for the two parameters which are controlled by automatic systems, Pressurizer Pressure, and Tavg (through Rod Control).

1.

PRESSURIZER PRESSUE 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 [

~

)+ac.

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 allowance of

(

)+a,c was made for this effect. Therefore, a total control system uncertainty of [

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

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

4

_ _ ___-_- _ _______ ___ _ ______u

TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY

+a.c SCA.

SHTE-STE -

$D BIAS.

RCA -

e RHTE.

RTE.

RD CA

~

~

48.C ELECTRONICS UNCERTAINTY -

PLUS

[LECTRONICS UNCERTAINTY -

PLUS CONTROLLER UNCERTAINTY 9

9 w

5-

2.

TAVG T,yg is controlled by a system that compares the auctioneered high T,yg from the loops with a reference, usualli derived from the first stage Turbine impulse Chamber Pressure. T is the average of the ayg 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 RIDS, 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 RID bypass manifold or in the Hot and Cold legs. Based on the assumption that 1 Ty and 1 TC cross + calibrated Rdf R1Ds are used to calculate T and the RTDs are ayg located in the RID bypass manifold, the CSA for the electronics is

[

)+a,c.

Assuming a normal, two sided probability distribution results in an electronics standard deviation (sg) of

(

)+a.c, O.f However, this does not include the controller deadband of 1 1.5 The controller accuracy is the combination of the instrt. mentation accuracy and the deadband. The probability distribution for the deadband hasbeendeterminedtobe[

).4a,c The va tance for the deadband uncertainty is then:

(s2) *I 3+"'c '

Combining the variance for instrumentation and deadband results in a contro11er variance of:

4 (sT) * (5 ) +(s2) "I 3 "'

1 Thecentro11ers7-[

)+a c for c total uncertain *.y of

[

)+a,c,

.s.

TABLE 2 ROD CONTROL SYSTEM ACCURACY lavg TURD PRES 44,C PHA =

SCA -

SMTE-

$1E -

50 BIAS =

RCA -

RM1E.

RM1E-R1E -

RD CA BIAS.

1. R1Ds USE0 - TH - 1 TC - 1 ELECTROMICS CSA

=

ELECTRONICS SIGMA -

CONTROLLER SIGMA CON 1 ROLLER BIAS CONTROLLER CSA e

9

~7-

3.

8';S it0W ITDP, and some plant Technical Specifications, requires an RCS flow measurement with a high degree of accuracy, it is assumed for this error analysis that the flow measurement is 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 been made for feedwater venturt fouling. The minimum power level t

assumed for the measurement is 90% 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 in ev, librium, the RCS total vessel flow is the sum of the individual primary loop flows, i.e.,

e WRCS = N(W ).

Eq. 4 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 enthalpy rise and multiplying by the Cold leg specific volume. The equation for this calculation is:

WL - IALLQso_:._Qpd / Nil (Vg1 t

(hg-h)

Eq. 5 C

where t

Loop flow (gpm)

W 3

0.1247 gpm/(ft /hr)

A Steam Generator thermal output (Btu /hr)

QSG RCP heat addition (Btu /hr)

Op Qt Primary system net heat losses (Btu /hr) 3 VC Specific volume of the Cold leg at TC (ft /lb) 8

Number of primary side loops N

Hot leg enthalpy (Btu /lb) ha

,old Leg enthalpy (Btu /lb).

hc The thermal output of the Steam Generator is determined by precision secondary side calorimetric measurement, which is defined as:

Q c - (h - hr)Wr Eq. 6 S

s where; h

Steam enthalpy (Btu /lb) 3 hr Feedwater enthalpy (Btu /lb)

Feedwater flow (1b/hr).

Wr 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 Feedwater pressure. The Feedwater flow is determined by multiple measurements and the following calculation:

Wr - (K)(F,)((pt)(d/p))l/2 Eq. 7 where; K Feedwater venturi flow coefficient Feedwater venturi correction for thermal expansion F

a 3

Feedwater density (ib/ft )

pr d/p Feedwater venturi pressure drop (inches H O).

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 the

~

difference between Feedwater temperature and calibration temperature.

Feedwater density is based on the measurement of Feedwater temperature and Feedwater pressure. The venturi pressure drop is obtained from the output of the differential pressure 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.

.g.

The primary system net heat losses are determined by calculation, considering the following system heat inputs and heat losses:

Charging flow Letdown flow Seal injection flow RCP thermal barrier cooler heat remcval Pressurizer spray flow Pressurizor surge line flor Component insulation heat losse; L

Component support heat losses CRDM heat losses.

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

Tha Hot leg and Cold Leg enthalpies are based on the measurement of the Hot leg temperature, Cold Leg temperature and the Pressurizer pressure, The Cold Leg specific volume is based on measurement of the Cold Leg temperature and Pressurizer pressure.

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

Stoamline pressure (P )

3 feedwater temperature (Tr)

Fee % Sressure (P )

f FeNaa B renturi differential pressura (d/p)

Hot Leg ttaperature (Tg)

Cold Leg temperature (T )

C Pressurizer pressure (P )

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

Feedwater venturi flow coefficients (K)

Feedwater venturi thermal expansion correction (F )

a f adwater density (pf). _ _ _ _ _ _ _ _ __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

Feedser enthalpy (hr)

Steam enthalpy (h )

s Hoisture carryover (impacts h )

s Primary system net heat losses (Q )t RCP heat addition (Q )

p Hot Leg enthalpy (h )

H Cold Leg enthalpy th )*

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

The derivation of the messurement errors and flow uncertainties on Table 5 6.re noted below.

Secondary Side The secondary side uncertainties are in four principal areas, feedwater flow, Feedwater enthalpy, Steam enthalpy and RCP heat addition, These 4

four areas are specifically identified on Table 5.

,($

k.3 For the measurement of Feedwater flow, each Feedwater venturi is Iy calibrated by the vendor in a hydraulics laboratory under controlled 1

conditions to an accuracy of (

)+a,b,c.

The eslibration dat which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of [

]+a,c js included for installato effects, resulting in a conservative overall flow coefficient (K) uncertainty of [

]+a,c.

Since RCS loop flow is proportional to Steam Generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [

]+a,t.

It should be noted that no allowance is mada 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.

1he uncertainty applied to the Feedwater venturi thermal expansion correction (F ) is based on the uncertainties of the measured Feedwater a

temperature and the coefficient of thermal expansion for the venturi _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

material, usually 304 stainless steel, for this material, a change of O

11 f in the nominal Feedwater temperature range changes F by a

0.002 % and the Steam Generator thermal output by the same amount.

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. Westinghouse uses the conservative value of [

]+a,c, Using the 1967 ASME Steam Tables it is p9ssible to determine the sensitivities of various parameters to c'ianges in Feedwater temperature and pressure.

Table 3 notes the instrument uncertainties for the hardware used te 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 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:

% flow - (d/p uncertainty)(1/2)(transmitter span /100)2 Typically, the Feedwater flow transmitter span is [

]+a,c nominal fl ow.

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: s

System heat losses

-2.0 MWt Component conduction and

~

convection losses

-1.4 Pump heat adder JJ.J Net Heat input to RCS

+10.1 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 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 Prairie Island 11 and by input power measurements from several plants, therefore, the uncertainty for the pump heat addition is estimated to be [

]4a,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.

c.

Primary Side The primary side uncertainties are in three principal areas. Hot leg enthalpy, Colri Leg enthalpy and Cold Leg specific volume. These are specifically noted on Table 5.

Three primary side parameters are actual y measured, Tg, TC and Pressurizer pressure. Hot leg enthalpy is influenced by T, Pressurizer pressure and Hot Leg temperature streaming.

H 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 systematic components. For plants with direct imersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the streaming uncertainty is (

)+a,c for both random and systematic components.

_The Cold leg enthalpy and specific volume uncertainties are impacted by T_C and--Pressurizer pressure. --Table 3 notes the TC instrument uncertainty and

_ Table 4,provides the sensitivities.

Noted on_ Table _5 is the plant spacific RID cross-calibration systematic allowance. _ When necessary, an allowance is made for a systematic temperaturo error due to the RTD cross-calibrati.on 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 Table-5, the 3. loop-uncertainty equation (with biases) is as follows:

+a,c L-Based on the number of loops, number, type and mea.turement method of RTDs, and the vessel' Delta-T, the flow uncertainty is:

c

_.# of loops flow uncertainty (% flow)

+a,c 3

TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION UNCERTAINTIES

(% SPAN)

FW TEMP FW PRES FW d/p STM PRESS TH TC PRZ PRESS

+a,c SCA =

SMTE-SPE -

STE -

SD R/E -

ROOT =

BIAS-CSA =

1. OF 1/ loop 1/ steam 1/ loop 1/ loop 3 **

INSTRUMENTS USED line OF psia

% d/p psia F

Of psia INST SPAN - 568.

1500.

120.

2000.

100.

100.

800.

INST UNC.

-+a,c (RANDOM) -

INST UNC.

(BIAS)

NOMINAL

- 435.

1064.

964.

620.2 557.4 2250.

+a,c

[

]

Number of Hot Leg and Cold leg RTDs used for measurement in each loop and the number of Pressurizer Pressure transmitters used overall, i.e., one per loop. Measuring and averaging more than one RTD per loop will provide greater accuracy on hot and cold leg temperature measurements. _-_____

TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW F a TEMPERATURE

+a,c HATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE h s 1194.2 BTV/LBM hf 414.0 BTV/LBM Dh(SG) 780.2 BTV/LBM STEAM ENTHALPY PRESSURE 4a,c MOISTURE HOT LEG ENTHALPY TEMPERATURE PRESSURE hH 643.3 BTV/LBM

=

bc(VESS) 556.4 BTV/LBM Dh 86.9 BTV/LBM Cp(Tg) 1.565 BTV/LBM OF COLD LEG ENTHALPY TEMPERATURE

+a,c PRESSURE Cp(T)

C 1.262 BTU /LBM Of

=

COLD LEG SPECIFIC VOLUME TEMPERATURE

+a,c PRESSURE 4. _

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

PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION HOT LEG ENTHALPY lEMPERATURE STREAMING, RANDOM STREAMING, SYSTEMATIC PRESSURE COLD LEG ENTHALPY TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE RTD CROSS-CAL SYSTEMATIC ALLOWANCE

, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS e

TABLE 5 (CONTINUED)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT FLOW UNCERTAINTY BIAS VALUES

+a,c FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRESSURE ENTHALPY - HOT LEG ENTHALPY - COLD LEG 5PECIFIC VOLUME - COLD LEG FLOW BIAS TOTAL VALUE SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

+a,c N LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES)

T, 45 9

4.....

As noted earlier, the precision flow calorimetric is used as the reference for the normalization of the Cold Leg elbow taps.

Assuming that the elbow tap d/p transmitters are used to feed the plant process computer, it is a simple matter to perform Technical Specification required surveillance.

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 2.0% with 0.06% flow bias The corresponding value used in ITDP is:

1. of loops standard deviation (% flow)

+a,c 3

4 -_

TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES

-. o

% d/p-SPAN

% FLOW PMA =

+a,c PEA --

SCA -

- SPE -

- STE -

SD RCA =

- RMTE=-

- RTE -

RD ID ' =

- A/D -

RDOT-BIAS =

FLOW CALORIN. BIAS. -

FLOW CALORIMETRIC INSTRUMENT SPAN

. SINGLE LOOP ELBOW TAP FLOW UNC -

+a,c N: LOOP ELBOW TAP FLOW UNC N LOOP RCS FLOW UNCERTAINTY

-(WITHOUT BIAS VALUES)

N LOOP RCS FLOW UNCERTAINTY-

'(WITH BIAS _ VALUES):

2.06

-G l

l -

t

1 4.

Reactor Powet 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 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 primary side system losses, and dividing by the core rated Btu /hr at full power. The equation for this calculation is:

RP - (fN)(O c - Op M / Nill (1001 S

L H

Eq. 8 where; Core power (% RTP)

RP Number of primary side loops N

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 Qt 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 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 marnin 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 _ _ _ _ _ _ _ _ _ _ _ _ _

1 portion), equations 6 and 7.

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, 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 Table 4.

As noted on Table 8. Westinghouse has determined the dependent sets in the calc ulation and the direction of interaction.

This is the same as that performed for the flow calorimetric, but applicable 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, the power measurement uncertainty for the secondary side power calorimetric is:

1. of loops power uncertainty (% RTP)

+a,c 3

IV.

CONCLUSIO!$

The preceding sections provide the methodology to account for instrument uncertainties for pressure, temperature, power and flow.

l 22-1

The plant-specific instrumentation has been reviewed for Shearon Harris

-and the uncertainty calculations are completed.

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

i h

I 1

TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES

(% SPAN)

FW TEMP FW PRES FW d/p STM PRESS SCA -

+a.c SMTE-SPE -

STE -

SD

=

BIAS =

RCA -

RMTE-RTE =

RD ID A/D --

CSA -

_J Of psia

% d/p psia INST SPAN = 430.

1200.

122.

1300.

INST UNC (RANDOM)-

+a,c INST UNC (BIAS)

NOMINAL- - 435.

1064.

964.

Since Feedwater Pressure is calculated, this is an assumed, conservative value.

?

a TABLE 8-

-SECONDARY SIDE POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES t

>~

COMPONENT:

INSTRUMENT ERROR-POWER UNCERTAINTY i

+R,C-FEEDWATER FLOW VENTURI 2 THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL DENSITYL

-TEMPERATURE PRESSURE-'

DELTA P -

-FEEDWATER ENTHALPY TEMPERATURE 1 PRESSURE 7

STEAM ENTHALPY=

PRESSURE-M013TURE 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 1

SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N; LOOP' UNCERTAINTY ~

(WITHOUTBIASVALUES)

-NoLOOP,0NCERTAINTY1 (WITH.BIASVALUES) 5

-l

+

4-lN' "

I I' *

.i iu t-E

{

REFERENCES l.

Westingh'ouse letter NS-CE-1583, C. Eiche1dinger to J. F. Stolz, NRC, dated 10/25/77.

2.

Westinghouse letter.NS-PLC-Sill, T. M. Anderson to.E. Case, NRC, dated l

'5/30/76.

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.

Westinghouse ~ Letter NS-1MA-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 2/12/01.-

7.'

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

August, 19824

? 8. - Regulatory Guide,1.105 Rev. 2, " Instrument Setpoints for Safety-Related Systems", dated 2/86.

9.- NUREG/CRe3659 (PNL-4973), "A Mathematica1'Model for Assessing the t

Uncertainties of-Instrumentation Measurements for Power and Flow of PWR Reactors",2/85.

g L

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"

i

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".

FIGURE 1 RCS FLOW CALOR! METRIC SCHEM TIC j

j {

j (

}

l P,

P j

(

l I

6P g

o h

h h,

h of Fa K

C H

f Ah I

f y

i 95G O

Q Measured t

g I

Calculated O d

Mass W

L y

I y

Vol.

"C M

7 L

I Other Loops I t RCS Volumetric Flow

-~. - -.

FIGURE 2 POWER CALOR! METRIC $CHE ETIC Pf i

q j

)l i

=

N h,

h P

F, K

f f

i I P g

N i

i t

[

calculated-gG S

O -asuma 1 r t

I C

Other Loops l.

4

'9 f

Core Power

~29-

.