ML17352B151
| ML17352B151 | |
| Person / Time | |
|---|---|
| Site: | Turkey Point  |
| Issue date: | 01/31/1995 |
| From: | Ciocca C WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML17352B147 | List: |
| References | |
| WCAP-13718, WCAP-13718-R01, WCAP-13718-R1, NUDOCS 9505120144 | |
| Download: ML17352B151 (61) | |
Text
WESTINGHOUSE PROPRIETARY CLASS 3 WCAP-13718 Rev.
1 WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY Florida Power 8 Light Company Turkey Point Units 3 E
4
- January, 1995 C.
F. Ciocca WESTINGHOUSE ELECTRIC CORPORATION Nuclear Technology Division P.O.
Box 355 Pittsburgh, Pennsylvania 15230-0355 e 1995 Westinghouse Electric Corp., All Rights Reserved 9505<iOZm V50505 PDR ADOCK 05000250 P,
PDR"-
~
Section Title TABLE OF CONTENTS Page INTRODUCTION II.
METHODOLOGY INSTRUMENTATION UNCERTAINTIES IV.
CONCLUSIONS 26 V.
REFERENCES 27
l'
LIST OF TABLES Table Number
= Title Page Pressurizer Pressure Control System Accuracy Rod Control System Accuracy Flow Calorimetric Instrumentation Uncertainties 16 4
Flow Calorimetric Sensitivities Calorimetric RCS Flow Measurement Uncertainties Cold Leg Elbow Tap Flow Uncertainty Power Calorimetric Instrumentation Uncertainties 17 18 21 24 Secondary Si de Power Cal or imetri c Measurement Uncertainties 25
I
LIST OF FIGURES Figure Number Title Page RCS Flow Calorimetric Schematic 29 Power Cal orimetri c Schemati c 30 111
WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY FOR TURKEY POINT UNITS 3 81 4 I.
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 monthly 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 and the uncertainty reflects this control system.
This report is based on the elimination of RTD Bypass Loops in the design to measure hot and cold leg reactor coolant system temperatures.
The RTDP"" is used to predict the plant's DNBR design limit, The RTDP methodology considers the uncertainties in the system operating plant parameters, fuel fabrication and nuclear and thermal parameters and includes the use of 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
5 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 outlined in WCAP-8567 "Improved Thermal Design Procedure","'
which is based on the conservative assumption that the uncertainties can be described with uniform probability distributions.
Another approach (for McGui re and Catawba) 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 both RTDP parameters and protection functions.
II.
METHODOLOGY The methodology used to combine the error components for a channel is the square root of the sum of the squares of those groups of components which are statistically independent.
Those errors that are dependent are combined arithmetically into independent
- groups, which are then systematically combined.
The uncertainties used are considered to be
- random, two sided distributions.
The sum of both sides is equal to the range for that parameter, e.g.,
Rack Drift is typically j"', the range for this parameter is [
]"'.
This technique has been utilized before as noted
- above, and has been endorsed by the NRC staff"" and 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:
For precision parameter indication using Special Test Equipment or a digital volt meter (DVM) at the input to the racks; CSA
= ((SCA + SMTE + SD)
+ (SPE)
+ (STE) + (RDOUT) )'
BIAS Eq. I 2.
For parameter indication utilizing the plant process computer; CSA
= ((SCA + SMTE + SD)
+ (SPE)
+ (STE)
+ (RCA + RMTE + RD)'
(RTE)
+ (ID)
+ (A/D) }'
BIAS Eq.
2 3.
For parameters which have control systems; CSA
= ((PMA)
+
(PEA)
+(SCA + SMTE + SD)
+
(RCA +
RMTE +
RD + CA)
+
(RTE) ) '
BIAS Eq.
3 PMA and PEA terms are not included in equations I 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 PMA PEA SCA SMTE SPE STE SD RCA RMTE RTE RD RDOUT ID A/D CA Channel Allowance Process Measurement Accuracy Primary Element Accuracy Sensor Calibration Accuracy Sensor Measurement and Test Equipment Accuracy Sensor Pressure Effects Sensor Temperature Effects Sensor Drift Rack Calibration Accuracy Rack Measurement and Test Equipment Accuracy Rack Temperature Effects Rack Drift Readout Device Accuracy (DVM or gauge)
Computer Isolator Drift Analog to Digital Conversion Accuracy Controller 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:
PMA PEA SCA SPE STE SD RCA RTE non-instrument related measurement errors, e.g.,
temperature stratification of a fluid in a pipe.
errors due to a metering devi ce, e.g.,
elbow, venturi, orifice.
reference (calibration) accuracy for a sensor or transmitter.
change in input-output relationship due to a change in static pressure for a differential pressure (d/p) cell.
change in input-output relationship due to a change in ambient temperature for a sensor or transmitter.
change in input-output relationship over a period of time at reference conditions for a sensor or transmitter.
reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated, or tuned, to this accuracy.
change in input-output relationship due to a change in ambient temperature for the rack modules.
4-
RD RDOUT ID CA BIAS-change in input-output relationship over a period of time at reference conditions for the rack modules.
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 at reference conditions for a control or protection signal isolating device.
allowance for conversion accuracy of an analog signal to a digital signal for process computer use.
allowance for the accuracy of a controller, not including deadband.
a non-random uncertainty for a sensor or 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.
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).
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 and controller.
As noted on Table 1, the electronics uncertainty for this function is [
]"'hich corresponds to an accuracy of [
]'*'.
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 [
]"'s calculated, which results in a standard deviation of [
]"'assuming a normal, two sided probability distribution).
r
TABLE I PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY SCA =
MS.TE=
=
SD BIAS=
RCA =
MME=
RTE
=
RD CA ELECTRONICS UNCERTAINTY =
PLUS ELECTRONICS UNCERTAINTY =
PLUS CONTROLLER UNCERTAINTY TAYG T,, is controlled by a system that compares the median loop T,, with a reference, derived from the First Stage Turbine Impulse Chamber Pressure.
T,,
is the average of the narrow range T and T, values.
The median loop T,, is then used in the controller.
Allowances are made (as noted on Table 2) for the RTDs, transmitter and the process racks 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 RTD bypass manifold or in the hot and cold legs.
Based on the assumption that two T and one T, cross-calibrated Weed RTDs are used to calculate T,, and the RTDs are located in the hot and cold
- legs, the CSA for the electronics is [
]"'.
Assuming a normal, two sided probability distribution results in an electronics standard deviation
(~~) of [
However, this does not include the controller accuracy is the t
the deadband.
The probability determined to be [
]."'he variance the controller deadband of +1.5 'F.
For T,,
combination of the instrumentation accuracy and distribution for the deadband has been for the deadband uncertainty is then:
]
+4>C Combining the variance for instrumentation and deadband results in a controller variance of:
]
<C>C The controller e,
= [
]"'nd, with.a [
temperature streaming, the total uncertainty is [
]"'ias for cold leg
]
+L,C SENSOR/TRANSMITTER Tavg PMA =
SCA
=
SMTE=
=
SD BIAS=
PROCESS RACKS TABLE 2 ROD CONTROL SYSTEM ACCURACY Turbine Pressure
RMTE=
RTE
=
RD CA EAO ANALOG TURBINE MSS t
8 HOT LEG RTDs
2 ROD CONTROL SYSTEM ACCURACY SENSOR/TRANSMITTER Tavg PMA
SCA =
SMTE=
=
SD BIAS=
8 COLD LEG RTDs
= I Turbine Pressure
+d,c PROCESS RACKS Tavg ERI RCA =
RMTE=
RTE =
RD CA EAO ANALOG TURBINE MSS
+a,c ELECTRONICS CSA ELECTRONICS SIGMA =
CONTROLLER SIGMA CONTROLLER BIAS CONTROLLER CSA
I RCS FLOW RTDP and plant 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 90K 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.,
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:
W-L (A)tQse Qp + (
)](Vg N
(h-hg Eq.
5 where; W
A QsG Qp Loop flow (gpm) 0.1247 gpm/(ft'/hr)
Steam Generator thermal output (Btu/hr)
RCP heat addition (Btu/hr)
0
Qp Qg
~c N
h hc RCP heat addition (Btu/hr)
Primary system net heat losses (Btu/hr)
Specific volume of the cold, leg at T, (ft'/lb)
Number of primary side loops Hot leg enthalpy (Btu/lb)
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:
QG= (h,-h)W, Eq.
6 where; h,
=
Steam enthalpy (Btu/lb) hf Feedwater enthal py (Btu/1 b)
W,
=
Feedwater flow (lb/hr).
The Steam enthalpy is based on the measurement of steam generator outlet Steam t
- 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:
Eq.
-7 where; Fa Pr hP Feedwater venturi flow coefficient Feedwater venturi correction for thermal expansion Feedwater density (lb/ft')
Feedwater venturi pressure drop (inches H,O).
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 t
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 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:
Steamline pressure (P,)
Feedwater temperature (Tf)
Feedwater venturi differential pressure (hP)
Hot leg temperature (T)
Cold Leg temperature (Tq)
Pressurizer pressure (Pp)
Steam Generator blowdown (if not secured) and on the following calculated values:
Feedwater venturi flow coefficients (K)
Feedwater venturi thermal expansion correction (F,)
Feedwater density (p,)
Feedwater enthalpy (hf)
Feedwater pressure (P,)
Steam enthalpy (h,)
Moisture carryover (impacts h,)
Primary system net heat losses (g,)
RCP heat addition
((}p)
Hot leg enthalpy (h)
Cold leg enthalpy (h,).
These measurements and calculations are presented schematically on Figure l.
The derivation of the measurement errors and flow uncertainties on Table 5 are noted below.
Secondary Side e
The secondary side uncertainties are in four principal areas, Feedwater flow; Feedwater
- enthalpy, Steam enthalpy and RCP heat addition.
These four areas are specifically identified on Table 5.
For the measurement of Feedwater flow, each Feedwater venturi was calibrated by the vendor in a hydraulics laboratory under controlled conditions to an accuracy of [
].""
The cal ibrati on data whi ch substanti ates thi s accuracy i s pr ovi ded to the plant by the vendor.
An'dditional uncertainty factor of [
]"'s included -for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of
[
]."'ince RCS loop flow is proportional to steam generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [
]."'t should be noted that no allowance is made for venturi fouling.
The venturis are 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 The uncertainty applied to the Feedwater venturi thermal expansion correction
(F,) is based on the uncertainties of the measured Feedwater temperature and the coefficient of thermal expansion for the venturi material, usually 304 stainless steel.
For this material, a change of ~l 'F in the nominal Feedwater temperature range changes F, by a0.002%
and the steam generator thermal output by the same amount.
An uncertainty in F, of t5X 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 [
]
."'sing the 1967 ASME Steam Tables it is possible to determine the sensitivities of various parameters to changes in Feedwater temperature and pressure.
Table 3 notes the instrument uncertainties for the hardware used to perform the measurements.
Table 4 lists the various sensitivities.
As can be seen on Table 4, Feedwater temperature uncertainties have an impact on venturi F Feedwater density and Feedwater enthalpy.
Feedwater pressure uncertainties impact Feedwater density and Feedwater enthalpy.
Feedwater venturi hP uncertainties are converted to X Feedwater flow using the following conversion factor:
X flow =
(hP uncertainty)(1/2)(transmitter span/100)'he Feedwater flow transmitter span is [ ]"'fnominal 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 [
]."'his 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.
These are summarized.for a t
three loop plant as follows:
System heat losses Component conduction and convection losses Pump heat adder Net Heat input to RCS
- 2.0 MWt
- 1.4
+11.4
+ 8 MWt The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [ ]"'fthe calculated value.
Since direct measurements are not possible, the uncertainty on component conduction and convection losses has been assumed to be
[ ]"'fthe calculated value.
Reactor coolant pump hydraulics are known to a relatively high confidence level, supported by system hydraulics tests performed at Prairie Island II and by input power measurements from several plants, therefore, the uncertainty for the pump heat addition is estimated to be [ ]"'fthe best estimate value.
Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are less than the value used in the analysis, which is
[
]"'fcore power.
Primary Side The primary side uncertainties are in three principal areas, hot leg enthalpy, cold leg enthalpy and cold leg specific volume.
These are specifically noted on Table 5.
Three primary side parameters are actually measured, T, T, and Pressurizer pressure.
Hot leg enthalpy is influenced by T, Pressurizer pressure and hot leg temperature streaming.
The uncertainties for the instrumentation are noted on Table 3 and the sensitivities are provided on Table 4.
The hot leg streaming is split into random and bias (systematic) components.
For Turkey Point Units 3 5 4, the RTDs are located in thermowells placed in the loops (bypass manifolds eliminated).
A plant specific evaluation has been performed which resulted in a streaming uncertainty of
[
]"'or random and [
]"'or systematic components.
1 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 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:
0 Based on the number of loops,
- number, type and measurement method of RTDs, and the vessel Delta-T, the flow is:
8 of loops flow uncertainty (X flow)
[
]<t,c
TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION UNCERTAINTIES (X SPAN)
FW TEMP FW PRES FW hP STM PRESS TH SCA SMTE =
RTE RD RDOUT=
BIAS =
CSA TC PRZ PRESS
¹ OF INST USED INST SPAN
= [
F PSI X hP PSI oF oF PSI jul C INST UNC.
(RANDOM) =
INST UNC.
(BIAS)
NOMINAL
=
437'F 885 PS IA 785 PSIA 602.3'F 546.2'F 2250 PSIA TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW FA TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE hS 1199.8 BTU/LBM hF 415.5 BTU/LBM Dh(SG)
=
784.3 BTU/LBM STEAM ENTHALPY PRESSURE MOISTURE HOT LEG ENTHALPY TEMPERATURE PRESSURE hH 616.5 hC 542.5 Dh(VESS)
=
74.1 Cp (TH)
=
- 1. 440 COLD LEG ENTHALPY TEMPERATURE PRESSURE BTU/LBM BTU/LBM BTU/LBM BTU/LBM-DEGF Cp(TC)
=
1.230 BTU/LBM-DEGF COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE
TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT 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 HOT LEG ENTHALPY TEMPERATURE STREAMING, RANDOM STREAMING, SYSTEMATIC PRESSURE COLD LEG ENTHALPY TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE INSTRUMENT ERROR FLOW UNCERTAINTY
- ,**,+,++ INDICATE SETS OF DEPENDENT PARAMETERS 0
TABLE 5 (CONTINUED)
CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES BIAS VALUES FEEDWATER PRESSURE STEAM PRESSURE PRESSURIZER PRESSURE FLOW BIAS TOTAL VALUE DENSITY ENTHALPY ENTHALPY ENTHALPY -
HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME -
COLD LEG FLOW UNCERTAINTY SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES) 3 LOOP UNCERTAINTY (WITHOUT BIAS VALUES)
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 hP transmitters feed the plant process computer, 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 one elbow tap per loop.
The hP transmitter uncertainties are converted to percent flow on the same basis as the Feedwater venturi bP.
The elbow tap uncertainty is then combined with the precision flow calorimetric uncertainty.
This combination of uncertainties results in the following total flow uncertainty:
k of loops flow uncertainty
(% flow) 3
+3;4 The corresponding values used in RTDP are:
0 of loops standard deviation (X flow) 3
[
j+4C 0
r 0
TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INDICATED RCS FLOW INPUT VALUES PEA SCA SMTE =
RCA RMTE =
RTE RD A/0 ROUT =
~ C ALL VALUES IN X d/p SPAN 8
OF LOOP
=
3 FLOW CALORIMETRIC =
FLOW CAL. BIAS PRESS.
CONTROL TEMP.
CONTROL FLOW SPAN ACCURACY OF INDICATED RCS FLOW FROM PROCESS COMPUTER PMA PEA SCA SMTE =
RCA RMTE =
RTE RD A/0 ROUT =
ALL VALUES IN
'X FLOW 1
LOOP ELBOW TAP N
LOOP ELBOW TAP N
LOOP RCS FLOW (NO BIAS) 3.4 X FLOW
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 15Ã 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:
(Njf@SG Q~ + (
)]~(100)
Eq.
8 where; RP N
Qsa Qp H
Core power (X RTP)
Number of primary side loops Steam Generator thermal output (BTU/hr) as defined in Eq.
6 RCP heat adder (Btu/hr) as defined in Eq.
5 Primary system net heat losses (Btu/hr) as defined in Eq.
5 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 e
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 plant computer will be used for the calculations.
The sensitivities calculated are the same as those noted for the secondary side on Table 4.
As noted on Table 8, Westinghouse has determined the dependent sets in the calculation 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 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:
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:
8 of loops 3
power uncertainty (X RTP)
[
]
+ITIC
TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES (X SPAN)
FW TEMP FW PRES FW DP STM PRESS SCA =
SMTE=
=
=
SD BIAS=
RCA =
RMTE=
RTE
=
RD ID A/D =
CSA =
DEG F
PSI PSI INST SPAN
=
r INST UNC (RANDOM) =
INST UNC (BIAS)
NOMINAL
= 437'F 885 PSIA 785 PSIA
TABLE 8 SECONDARY SIDE POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES COMPONENT
.,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 INSTRUMENT ERROR POWER UNCERTAINTY BIAS VALUES FEEDWATER DELTA P FEEDWATER PRESSURE STEAM PRESSURE POWER BIAS TOTAL VALUE DENSITY ENTHALPY ENTHALPY
- ,** INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES) 3 LOOP UNCERTAINTY (WITHOUT BIAS VALUES)
IV.
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 supplied by Florida Power 8 Light Company have been reviewed and the uncertainty calculations completed using this data.
i
v.
REFERENCES Westinghouse letter NS-CE-1583, C. Eicheldinger to J.
F. Stolz,
- NRC, dated 10/25/77.
2.
Westinghouse letter NS-PLC-5111, T.
M. Anderson to E. Case, NRC, dated 5/30/78.
3.
Westinghouse letter NS-TMA-1837, T.
M. Anderson to S. Varga,
- NRC, dated 6/23/78.
4.
Westinghouse letter NS-EPR-2577, E.
P.
Rahe Jr. to C.
H. Berlinger, NRC, dated 3/31/82.
5.
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.
NUREG-0717 Supplement No. 4, Safety Evaluation Report related to the operation of Virgil C.
Summer Nuclear Station Unit No.
1, Docket 50-395,
- August, 1982.
8.
Regulatory Guide 1.105 Rev. 2, "Instrument Setpoints for Safety-Related Systems",
dated 2/86.
9 NUREG/CR-3659 (PNL-4973),
"A Mathematical Model for Assessing the Uncertainties of Instrumentation Measurements for Power and Flow of PWR Reactors",
2/85.
10.
ANSI/ANS Standard 58.4-1979, "Criteria for Technical Specifications for Nuclear Power Stations".
ll.
ISA Standard S67.04, 1987, "Setpoints for Nuclear Safety-Related Instrumentation Used in'Nuclear Power Plants".
n.
- 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".
14 'estinghouse WCAP-11397-P-A, "Revised Thermal Design Procedure",
dated April, 1989.
V
SECONDARY SIDE PRIMARY SIDE P,
Pf Pp hs h,
Pf F,
W, QsG
~
- calculated g
- measured w
Other Loops RCS FLOW Figure I RCS Flow Calorimetric Schematic 1
I 1
~
J
SECONDARY SIDE F,
Wq Qsa Other Loops Core Power figure 2
Power Calorimetric Schematic A