ML17353A505

Nonproprietary Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for FP&L Turkey Point Units 3 & 4.

ML17353A505
Person / Time
Site:	Turkey Point
Issue date:	09/30/1995
From:	Ciocca C WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML17353A502	List: ML17353A501 ML17353A503 ML17353A504 ML17353A505 ML17353A506
References
WCAP-13718, WCAP-13718-R02, WCAP-13718-R2, NUDOCS 9512270314
Download: ML17353A505 (64)

Text

WESTINGHOUSE PROPRIETARY CLASS 3 WCAP-13718 Rev. 2 WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY For Florida Power & Light Company Turkey Point Units 3 & 4 September 1995 C. F. CIOCCA 95122703i4 95i2i8 PDR ADOCK 05000250 P PDR WESTINGHOUSE ELECTRIC CORPORATION Nuclear Technology Division P.O. Box 355 Pittsburgh, Pennsylvania 15230-0355 1995 Westinghouse Electric Corporation All Rights Reserved

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TABLE OF CONTENTS Section Title Pacae I. INTRODUCTION II. METHODOLOGY III. INSTRUMENTATION UNCERTAINTIES IV. CONCLUSIONS 32:

V. REFERENCES 33

LIST OF TABLES Table Number Title Pacae Pressurizer Pressure Control System Accuracy Rod Control System Accuracy Flow Calorimetric Instrumentation Uncertainties 21 Flow Calorimetric Sensitivities 22 Calorimetric RCS Flow Measurement Uncertainties 23 Cold Leg Elbow Tap Flow Uncertainty 26 Power Calorimetric Instrumentation Uncertainties 30 Secondary Side Power Calorimetric Measurement Uncertainties 31 3.3.

LIST OF FIGURES Fi ure Number Title Pacae RCS Flow Calorimetric Schematic 35 Power Calorimetric Schematic 36

WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY 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 (T,~), 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 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 cycle.

The RCS Cold Leg elbow taps are evaluated against the precision calorimetric and used for daily 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 via the temperature input to the rod control system for insertion only, .and the uncertainty reflects this control system. This report is 'based on the elimination of RTD Bypass Loops, that was implemented at Turkey Point in 1991, in the design to measure hot and cold leg reactor coolant system temperatures. The RTDP"'s 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 various DNB correlations. Use of the RTDP methodology requires that variances in the plant operating parameters be justified. The purpose of the following evaluation is to define the specific Turkey Point Units 3 6 4 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 2 and Trojan) used the methodology

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outlined in WCAP-8567 "Improved Thermal Design Procedure",<""

which: is based on the conservative assumption that the uncertainties can be described with uniform probability distributions. Another approach is based on the more realistic assumption that the uncertainties can be described with random, normal, two sided probability distributions.< 'his approach is used to substantiate the acceptability of the protection system setpoints. for many Westinghouse plants, e.g., D. C. Cook 2<'>,

V. C. Summer, Wolf Creek, Millstone Unit 3 and others. The second approach is now utilized for the determination of all instrumentation errors for the RTDP parameters and protection functions.

The RTDP methodology has been previously approved by the NRC in Westinghouse topical report WCAP-11397-P-A<"'. By letter L 131, dated May 5, 1995'"', FPL submitted a proposed revision to the Turkey Point Units 3 and 4 Technical Specifications to include the implementation of Westinghouse NRC approved Revised Thermal Design Procedure. The proposed changes to the instrument uncertainties are based on the use of the RTDP methodology as documented in Revision 1 of this WCAP. In support of the thermal power up-rate project, Westinghouse generated this revision to address only the changes to the instrument uncertainties affected as a result of an increase in the core power level.

METHODOLOGY The methodology used to comb'ine 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 [ 1"".

the range for this parameter is [ ]"'. This technique has been utilized before as noted above, and has been endorsed by the NRC staff" nd various industry standards'"'"'.

The relationships between the error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint Methodology<"> and are defined as follows:

e .. For precision parameter indication using Special Test Equipment or a digital voltmeter (DVM) at the input to the racks, and no trending of transmitter calibration and drift; CSA = ((SD + SMTE) + (SRA) + (SPE) + '(STE) +

(RDOUT)z)xa + ((SCA + SMTE)a)xa + BIAS Eq. 1

2. For parameter indication utilizing the plant process computer and no trending of transmitter. calibration and drift; CSA = ((SD + SMTE) + (SRA) + (SPE) + (STE) +

(RCA + RMTE) + (RD + RMTE) + (RTE) + (ID) +

(A/D) ) + ((SCA + SMTE) )x/2 + BIAS Eq. 2

3. For parameters which have control systems and no trending of transmitter calibration and drift;

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

(STE) + (RCA + RMTE) + (RD + RMTE) +

(RTE) ) + ((SCA + SMTE) ) + BIAS Eq. 3 PMA and PEA terms are not included in equations 1 and 2 since the equations are to determine instrumentation uncertainties only.

PMA and PEA terms are included in the determination of control system uncertainties.

where:

CSA Channel Statistical Allowance Process Measurement Accuracy PEA Primary Element Accuracy SCA Sensor Calibration Accuracy Sensor Reference 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 CA Controller Allowance ID Computer Isolator Drift A/D Analog to Digital Conversion Accuracy The parameters above are as defined in references 5 and 12 and are based on SAMA Standard PMC 20.1, 1973'"'. However, for ease in understanding they are paraphrased below:

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 calibration tolerance for a sensor/transmitter.

reference (calibration) accuracy for a sensor/transmitter.

SPE change in input-output relationship due to a change in static pressure for a differential pressure (d/p) cell.

STE change in input-output relationship due to a change in ambient temperature for a sensor or transmitter.

SD change in input-output relationship over a period of time at- reference conditions for a sensor or transmitter.

RCA 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.

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 it's most accurate applicable range for the parameter measured, or 1/2 the smallest increment on an indicator (readability).

CA allowance for the accuracy of the median signal selector and the controller circuitry..

ID change in input-output relationship over a period of time at reference conditions for a control or protection signal isolating devices A/D allowance for conversion accuracy of an analog signal to a digital signal for process computer use.

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

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A more detailed explanation of the Westinghouse methodology noting the interaction of several parameters is provided in references 6 and 13.

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

1. PRESSURIZER PRESSURE 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/indicators and controller. As noted on Table 1, the controller uncertainty for this function is [+ ]"" which corresponds to an accuracy of [ )"" for the average of 3 indicators. 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 [ ]"'as made for this effect. Therefore, a total control system uncertainty of

,[ ]" " is calculated, which results in a standard deviation of [ ]"'assuming a normal, two sided probability distribution).

TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY

+a,c SCA, SRA SMTE STE SD BIAS RCAzNo RMTEzo RTEzNp RDzNo RDOUTzNp RC+Tz.

+a c ELECTRONICS UNCERTAINTY PLUS ELECTRONICS UNCERTAINTY PLUS CONTROLLER UNCERTAINTY (Average of 3 Indicators)

2. ~Tav T, is controlled by a system that compares the median T from the loops with a reference derived from the First Stage Turbine Impulse Chamber Pressure. T, is the average of the narrow range T and Tc values. The median loop T,~ is then used for rod control. Allowances are made (as noted on Table 2) for the RTDs, transmitter and the process racks/indicators and 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 Hot and Cold Legs. Based on the assumption that two T and one Tc cross-calibrated RTDs are used. to calculate T, and the RTDs are located in the hot and cold leg piping, the CSA for the electronics is [ ]"'. Assuming a normal, two sided probability distribution results in an electronics standard deviation (a,) of [

this include the deadband for control i.e, auto

]"'owever, does not or manual of + 1.5 F. For T, the controller accuracy is the combination of the instrumentation accuracy and the deadband.

The probability distribution for the deadband has been determined to be [

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

(a,)' [ ]

+a c

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

(a )' (a,)' (cr,)' [

The controller a~ = [ ]"'or a total random uncertainty of [ ]+a,c

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With the incorporation of T~~ streaming, an additional bias of 0 1 F is included in Table 2. Therefore, the total uncertainty of the controller with the T, ~~ streaming included is [

and [ ]"" bias. ]"'andom ig~

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0 SENSOR/TRANSMITTER Tavg in 'F TABLE 2 ROD CONTROL SYSTEM ACCURACY Turbine Pressure in+age% span PMA =

SCA SRA SMTE=

STE SD BIAS=

PROCESS RACKS Tavg .ERI EAO INDICATOR CONTROLLER TURBINE All values in  % span Span 150 F 75 F 75 F 75 F 630 psig+a,c RCA RMTE RTE RD RDOUT=

CA I HOT LEG RTDs = 2 4 COLD LEG RTDs = 1 ROD CONTROL SYSTEM ACCURACY SENSOR/TRANSMITTER Tavg Turbine Pressure All values in  % span Span 754F 630 psig +a,c PMA SRA =

SCA =

SMTE=

STE SD BIAS=

TABLE 2 (Continued)

ROD CONTROL SYSTEM ACCURACY PROCESS RACKS Tavg ERI EAO INDICATION TURBINE All values in  % span

.Span 75 F 75cF 75 F 630 psig RCA RMTE=

RTE RD

+&g C ELECTRONICS CSA ELECTRONICS SIGMA CONTROLLER SIGMA CONTROLLER BIAS CONTROLLER CSA i+i 0

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3,. RCS FLOW RTDP and Turkey Point's Technical Specifications require an RCS flow measurement with a high degree of accuracy. Six month drift effects have been included for feedwater temperature, feedwater flow, steam pressure, and pressurizer pressure. It is assumed for this error analysis that the flow measurement is performed within ninety days of completing the cross-calibration of the hot leg and cold leg narrow range RTDs. Therefore, partial drift effects are 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 venturi fouling, and the calorimetric is performed above 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 equilibrium, the RCS total vessel flow is the sum of the individual primary loop flows, i.e.,

WRqq

= N(W~) 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:

Eq. 5 (hH hg

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where; W~ Loop flow (gpm)

A Constant conversion factor 0.1247 gpm/(ft'/hr)

QSG Steam Generator thermal output (Btu/hr)

Qp RCP heat addition (Btu/hr)

Qz, Primary system net heat losses (Btu/hr)

Vc Specific volume of the Cold Leg at Tc (ft'/lb)

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

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

QsG (hs hf)Wf Eq. 6 t where; h, hf W,

=

=

=

Steam enthalpy (Btu/lb)

Feedwater enthalpy (Btu/lb)

Feedwater flow (lb/hr).

The Steam enthalpy is based on the measurement of Steam Generator outlet Steam pressure, assuming saturated conditions. The Feedwater enthalpy is based on the, measurement of Feedwater temperature and nominal Feedwater pressure. The Feedwater flow is determined by multiple measurements and the following calculation:

Wf (K) (F ) { (pf ) (d/p) )'" Eq. 7 where; K = Feedwater venturi flow coefficient F, = Feedwater venturi correction for thermal expansion pf = Feedwater density (lb/ft3) d/p = Feedwater venturi pressure drop (inches H,O).

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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.

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 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.

The Hot Leg and Cold Leg enthalpies are based on the measurement of the Hot Leg temperature, Cold Leg temperature and the nominal Pressurizer pressure. The Cold Leg specific volume is based on measurement of the Cold Leg temperature and nominal Pressurizer The RCS flow measurement is thus based on the following plant measurements:

Steamline pressure (P,)

Feedwater temperature (T~)

Feedwater venturi differential pressure (d/p)

Hot Leg temperature (T)

Cold Leg temperature (T~)

Steam Generator blowdown flow (if not secured)

Pressurizer pressure (P )

and on the following calculated values:

Feedwater venturi flow coefficients (K)

Feedwater venturi thermal expansion correction (F )

Feedwater density (p,)

Feedwater enthalpy (h~)

Feedwater pressure (P,)

Steam enthalpy (h,)

Moisture carryover (impacts h,)

Primary system net heat losses (Q~)

RCP hea't addition (Q )

Hot Leg enthalpy (h)

Cold Leg enthalpy (h,).

These measurements and calculations are presented schematically in Figure l.

The derivation of the measurement errors and flow uncertainties on Table 5 are noted below.

Secondar Side The secondary side uncertainties are in four principal areas, Feedwater flow, Feedwater enthalpy, Steam enthalpy and RCP heat addition. These four areas are specifically identified on 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 [ ]" b". The calibration data which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of

[ ]"" is included for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of [ ]"'. Since RCS loop flow is proportional to Steam Generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [ ]"'. 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.

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

An uncertainty in F, of + 5 for 304 stainless steel is used in this analysis. This results in an additional uncertainty of

[ ]"'n Feedwater flow. Westinghouse uses the conservative value of [ ]+a,c Using the NBS/NRC Steam Tables it is possible to determine the sensitivities 'of various parameters to changes in Feedwater temperature and pressure. Table 4 notes the instrument uncertainties for the hardware used to perform the measurements.

Table 5 lists the various sensitivities. As can be seen on Table 5, Feedwater temperature uncertainties have an impact on venturi if'>

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F Feedwater density and 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)'he Feedwater flow transmitter span is 120 % of nominal flow.

Using the NBS/NRC Steam Tables, 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 [ ]"'. This value is noted on Table 4.

t 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 +11.4 Net Heat input to RCS + 8.0 MWt The uncertainty on system heat losses, which is essentially all to charging and letdown flows, has been estimated to be due

[ ]"'f the calculated value. Since direct measurements convection losses has been assumed to be [ ]"'f are not possible, the uncertainty on component conduction and calculated value. Reactor coolant pump hydraulics are known to a the relatively high confidence level, supported by system hydraulics i+i 0,

tests performed at Prairie Island Unit 2 and by input power measurements from several other plants, therefore, the uncertainty for the pump heat addition is estimated to be

[ ]"" of the best estimate value. Considering these parameters as one quantity, which is designated the,net pump heat uncertainty, the combined uncertainties are less than' of the total, which is .[ ]' of core power.

Primar 'ide 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 measured, T, T~ and Pressurizer prcssure'ot Leg enthalpy is influenced by T, Pressurizer pressure and Hot Leg temperature streaming. The uncertainties for the instrumentation are noted on Table 3 and the sensitivities are provided on Table 4. The hot leg streaming is split into random and bias (systematic) components. For Turkey Point Units 3 a 4, the RTDs are located in thermowells placed in the loops (bypass manifolds eliminated). A plant specific evaluation has'een performed which resulted in a streaming uncertainty of [ ]"" for random and [. ]"'or systematic components.

The cold leg enthalpy and specific volume uncertainties are impacted by T, and Pressurizer pressure. Table 3 notes the T, 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 procedure. No allowance was necessary for Turkey Point.

Parameter dependent effects are identified on Table 5.

Westinghouse has determined the dependent sets in the calculation 4>.

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).

4 Using Table 5, the 3 loop uncertainty equation (with biases) 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 is:

of loops flow uncertainty (% flow)

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TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION UNCERTAINTIES

(% Span)

FW TEMP FW PRES FW d/p STM PRESS T Tc PRZ PRESS

+4 C SCA SRA SMTE SPE STE SD RDOUT=

RCA RMTE RTE RD BIAS CSA =

OF INST USED oF psi d/p 'psi oF oF psi 120% Flow INST SPAN = 500 1500 120 1200 120 120 1000 INST UNC. +&gC (RANDOM)=

INST UNC.

(BIAS)

NOMINAL = 432'F 879psia 779psia 607.8'F 546.6'F 2250psia 4>

TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW

+AgC TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE h, 1199.'7 BTU/LBM h~ 410.5 BTU/L'BM Dh(SG) 789.2 BTU/LBM STEAM ENTHALPY

+4 C PRESSURE MOISTURE HOT LEG ENTHALPY TEMPERATURE PRESSURE h 624.0 BTU/LBM hc 542.6 BTU/LBM Dh(VESS) 81.4 BTU/LBM CV(T.) 1.465 BTU/LBM- F COLD LEG ENTHALPY

+Ay C TEMPERATURE PRESSURE'P(

c) 1.233 BTU/LBM- F COLD LEG SPECIFIC VOLUME

+ABC TEMPERATURE PRESSURE igt 4i 0

TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT INSTRUMENT ERROR FLOW UNCERTAINTY FEEDWATER FLOW VENTURI THERMAL EXPANSION COEFFICIENT.

TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA' FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE HOT LEG .ENTHALPY TEMPERATURE STREAMING, RANDOM STREAMING, SYSTEMATIC PRESSURE COLD LEG ENTHALPY TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE

• , **, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS if'>

'.TABLE 5 (CONTINUED)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT FLOW UNCERTAINTY

+4 BIAS VALUES C FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRESSURE ENTHALPY HOT LEG ENTHALPY COLD LEG SPECIFIC VOLUME COLD LEG FLOW BIAS TOTAL VALUE

+4 C SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES) igt II

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

Since the elbow tap bP transmitters feed the control board indicators, it is a simple matter to perform Technical Specification required surveillance. Table 6 notes the instrument uncertainties for determining flow by using the elbow taps, assuming three elbow taps per loop. The hP transmitter uncertainties are converted to percent flow on the same basis as the Feedwater venturi hP. The elbow tap uncertainty is then combined with the precision flow calorimetric uncertainty. 'This combination of uncertainties results in the following total flow uncertainty:

I of loops flow uncertainty (% flow)

+ 3.6 t The corresponding I of loops values used in standard deviation RTDP are:

(% flow)

• Based on Florida Power a Light procedures and information.

TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES PEA ALL VALUES IN d/p SPAN L  %

SCA 0.50 SMTE 0.34 SPE 1.12 STE'D 1.44 0.51 BIAS* 0.00 +& C RTE A/D

+&pC RCA ALL, VALUES IN '%LOW RD RMTE RDOUT OF LOOP = 3 +&gC FLOW CALORIMETRIC FLOW CAL. BIAS PRESS. CONTROL TEMP. CONTROL FLOW SPAN ACCURACY OF INDICATED RCS FLOW FROM CONTROL BOARD INDICATION

+& C PMA ALL VALUES IN  % FLOW PMA PEA SCA 0.36 SMTE 0.24 SPE 0.37 STE 1.04 SD 0.37 BIAS 0.00

+&gC RCA RMTE RTE RD A/D RDOUT LOOP ELBOW TAP 2.6 FLOW t

N LOOP ELBOW TAP 1.5  %

'%LOW FLOW'.6 N LOOP RCS FLOW (NO BIAS)

• Based on Florida Power & Light procedures and information.

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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 of the Operating License and to adjust the Power Range Neutron Flux channels when the difference between the NES 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:

((Nj[Q+G Q + ( )])(100)

Eq. 8 where; RP Core power (% RTP)

N Number of primary side loops Qsc Steam Generator thermal output (BTU/hr) as defined in Eq. 6 Qp RCP heat addition (Btu/hr) as defined in Eq. 5 Q~ Primary system net 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 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 main control board indicators will be used for the calculations. The sensitivities calculated are the same as those noted for the secondary side in Table 4. As noted in 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 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 results 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:

+ABC 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:

I of loops power uncertainty (% RTP) ill 41

TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES

(% SPAN) FW TEMP FW PRES FW d/p STM PRESS

+4 C SCA SRA SMTE=

SPE STE SD BIAS=

RCA =

RMTE=

RTE RD ID A/D CSA oF psi  % d/p psi 120% Flow INST SPAN = 500 1500 120 1200

+4 INST UNC C (RANDOM)

INST UNC (BIAS)

NOMINAL = 432'F 879 .psia 779 psia i+i

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TABLE 8 SECONDARY SIDE POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES COMPONENT INSTRUMENT ERROR POWER UNCERTAINTY 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 P2QVLMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES)

ZV. CONCLUSIONS The preceding sections provide the methodology to account for pressure, temperature, power and RCS flow uncertainties for the RTDP analysis. The plant specific instrumentation data and procedures have been reviewed for Turkey Point Units 3 & 4 and the uncertainty calculations. completed. These uncertainty values or more conservative values are used in the RTDP analysis.

V REFERENCES

1. Westinghouse WCAP-11397-P-A, "Revised Thermal Design Procedure", dated April 1989.

2. Westinghouse letter NS-CE-1583, C. Eicheldinger to J. F.

Stolz, NRC, dated 10/25/77.

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

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

5. Westinghouse letter NS-EPR-2577, E. P. Rahe Jr. to C. H.

Berlinger, NRC, dated 3/31/82.

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

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

8. 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.

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

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

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

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

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

14 'cientific Apparatus Manufacturers Association, Standard PMC 20.1, 1973, "Process Measurement and Control. Terminology".

15. WCAP-11397-P-A, "Revised Thermal Design Procedure", dated April 1989.

16. Letter T. F, Plunket (FPL) to USNRC, "Proposed License Amendments Implementation of the Revised Thermal Procedure and Steam Generator Water Level Low-Low Setpoint", L-95-131, dated May 5, 1995.

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SECONDARY SIDE PRIMARY SIDE Pc Tx Pp Tc F K h hc Wr 0

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