ML18101B377: Difference between revisions

From kanterella
Jump to navigation Jump to search
(Created page by program invented by StriderTol)
 
(StriderTol Bot change)
 
(4 intermediate revisions by the same user not shown)
Line 3: Line 3:
| issue date = 08/31/1993
| issue date = 08/31/1993
| title = Non-proprietary Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Salem Units 1 & 2.
| title = Non-proprietary Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Salem Units 1 & 2.
| author name = CIOCCA C F
| author name = Ciocca C
| author affiliation = WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
| author affiliation = WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
| addressee name =  
| addressee name =  
Line 17: Line 17:


=Text=
=Text=
{{#Wiki_filter:**
{{#Wiki_filter:WESTINGHOUSE CLASS 3 (NON-PROPRIETARY)
* WESTINGHOUSE CLASS 3 (NON-PROPRIETARY)
WCAP~13652 WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY SALEM UNITS 1 & 2
WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY SALEM UNITS 1 & 2 August, 1993 C. F. Ciocca WESTINGHOUSE ELECTRIC CORPORATION Nuclear & Advanced Technology Division P.O. Box 355 Pittsburgh, Pennsylvania 15230-0355 1993 Westinghouse Electric Corp., All Rights Reserved 96o522oo5s  
* August, 1993 C. F. Ciocca WESTINGHOUSE ELECTRIC CORPORATION Nuclear & Advanced Technology Division P.O. Box 355 Pittsburgh, Pennsylvania 15230-0355
-96082-5-----=--
                ~ 1993 Westinghouse Electric Corp., All Rights Reserved 96o522oo5s - 96082-5-----=-- -----1 PDR   ADOCK 05000272                 I
-----1 PDR ADOCK 05000272 I p. PDR I TABLE OF CONTENTS Section Title Page I. INTRO DU CTI ON 1 I I. METHODOLOGY 3 III. INSTRUMENTATION UNCERTAINTIES 5 IV. CONCLUSIONS 24 v. REFERENCES 27 ;
: p.               PDR                 I
LIST OF TABLES Table Number Title Page 1 Pressurizer Pressure Control System Accuracy 6 2 Rod Control System Accuracy 8 3 Flow Calorimetric Instrumentation Uncertainties 16 4 Flow Calorimetric Sensitivities 17 5 Calorimetric RCS Flow Measurement Uncertainties 18 6 Cold Leg Elbow Tap Flow Uncertainty 21 7 Power Calorimetric Instrumentation Uncertainties 25 8 Secondary Side Power Calorimetric Measurement Uncertainties 26 i i Figure Number 1 2
 
* LIST OF FIGURES RCS Flow Calorimetric Schematic Power Calorimetric Schematic iii 29 30 I. 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 (l'avg), Reactor Power, and Reactor Coolant System Flow. They are frequently monitored and several are used for control purposes:
TABLE OF CONTENTS Section Title                         Page I.     INTRO DU CTI ON                 1 I I. METHODOLOGY                     3 III. INSTRUMENTATION UNCERTAINTIES   5 IV. CONCLUSIONS                     24
Reactor power is monitored by the performance of a secondary side heat balance (power calorimetric) once every 24 hours. RCS fl ow is monitored by the performance of a precision fl ow calorimetric at the beginning of each cycle. The RCS Cold Leg elbow taps are evaluated against the precision calnrimetric and used for monthly surveillance (with a small increase in uncertainty).
: v.     REFERENCES                     27
Pressurizer pressure is a controlled parameter and the uncertainty reflects the control system. Tavg is a controlled parameter via the temperature input to the rod control system and the uncertainty  
 
*reflects this control systemL This report is based on the elimination of RTD Bypass Loops in the design to measure hot and cold leg reactor coolant system temperritures.
LIST OF TABLES Table Number Title                             Page 1       Pressurizer Pressure Control System Accuracy                     6 2       Rod Control System Accuracy         8 3       Flow Calorimetric Instrumentation Uncertainties       16 4       Flow Calorimetric Sensitivities     17 5       Calorimetric RCS Flow Measurement Uncertainties           18 6       Cold Leg Elbow Tap Flow Uncertainty                   21 7       Power Calorimetric Instrumentation Uncertainties       25 8       Secondary Side Power Calorimetric Measurement Uncertainties           26 ii
The 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.
LIST OF FIGURES Figure Number 1        RCS Flow Calorimetric Schematic 29 2        Power Calorimetric Schematic   30
The purpose of the following evaluation is to define the specific Salem Nuclear Plant instrument uncertainties for the four primary system operating parameters.*
* iii
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 11 Improved Thermal Design Procedure", 11*2*3 l which is based on the conservative assumption that the uncertainties can be described with uniform probability distributions.
WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY 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 (l'avg), Reactor Power, and Reactor Coolant System Flow. They are frequently monitored and several are used for control purposes: Reactor power is monitored by the performance of a secondary side heat balance (power calorimetric) once every 24 hours. RCS fl ow is monitored by the performance of a precision fl ow calorimetric at the beginning of each cycle. The RCS Cold Leg elbow taps are evaluated against the precision calnrimetric and used for monthly surveillance (with a small increase in uncertainty). Pressurizer pressure is a controlled parameter and the uncertainty reflects the control system. Tavg is a controlled parameter via the temperature input to the rod control system and the uncertainty
Another approach (for McGuire and Catawba) is based on the more realistic assumption that the uncertainties can be described with random, normal, two
*reflects this control systemL This report is based on the elimination of RTD Bypass Loops in the design to measure hot and cold leg reactor coolant system temperritures. The RTDP< 1 ~ is used to predict the plant's DNBR design limit.
* sided probability distributions.
The RTDP methodology considers the uncertainties in the system operating plant.*
141 This approach is used to substantiate the acceptability of the protection system setpoints for many Westinghouse plants, e.g., D. C. Cook 2 15 1, V. C. Surrmer, Wolf Creek, Millstone Unit 3 and others. The second approach is now utilized for the determination of a 11 instrumentation errors for both RTDP parameters and protection functions.
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 Salem Nuclear Plant instrument uncertainties for the four primary system operating parameters.*
* II. METHODOLOGY The methodology used to combine the error components for a channel is the square root of ium of the squares of groups of components which are statistically independent.
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 11 Improved Thermal Design Procedure", 11 *2 *3l which is based on the conservative assumption that the uncertainties can be described with uniform probability distributions.
Those errors that are dependent are combined arithmetically into independen.t groups, which are then systematically combined.
Another approach (for McGuire and Catawba) is based on the more realistic assumption that the uncertainties can be described with random, normal, two sided probability distributions. 141 This approach is used to substantiate the
The uncertainties used are considered to be random, two sided distributions.
* acceptability of the protection system setpoints for many Westinghouse plants, e.g., D. C. Cook 2151, V. C. Surrmer, Wolf Creek, Millstone Unit 3 and others.
The sum of both sides is equal to the range for that parameter, e.g., Rack Drift is typically  
The second approach is now utilized for the determination of a11 instrumentation errors for both RTDP parameters and protection functions.
[ ] +a.c, the range for this parameter is [ J +a,c. This technique has been utilized before as noted above, and has been endorsed.
II. METHODOLOGY The methodology used to combine the error components for a channel is the square root of th~ ium of the squares of tho~e groups of components which are statistically independent. Those errors that are dependent are combined arithmetically into independen.t 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
by the NRC staff 16*7*8*9' and various industry standards 00*11'. The relationships between error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint Methodology!
[             ] +a.c, the range for this parameter is [           J+a,c.
12 l and are defined as follows: 1. For precision parameter indication using Special Test Equipment or a digital volt.meter (DVM) at the input to the racks; CSA = {(SCA + SMTE + SD)2 + (SPE)2 + (STE)2+ (RDOUT)2}112 + BIAS Eq. 1 2. For parameter indication utilizing the plant process computer; CSA = {(SCA + SMTE + SD)2 + (SPE)2 + (STE)2 + (RCA + RMTE + RD)2 + (RTE)2 + (ID)2 + (A/D)2}112 +BIAS 2 3. For parameters which have control systems; CSA= {(PMA)2 + (PEA)2 +(SCA+ SMTE + SD)2 + (SPE)2 + (STE)2 + (RCA + RMTE + RD.+ CA)2 + (RTE)2}112 + 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.
This technique has been utilized before as noted above, and has been endorsed. by the NRC staff 16 *7*8 *9' and various industry standards 00
*
* 11 '.
* 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  
The relationships between th~ error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint Methodology! 12 l and are defined as follows:
= Computer 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(nl. However, for ease in Understanding they ate paraphrased below: PMA -non-instrument related measurement errors, e.g., temperature stratification of a fluid in a pipe. PEA -errors due to a metering device, e.g., elbow, venturi, orifice. SCA -reference (calibration) accuracy for a sensor or transmitter.
: 1.     For precision parameter indication using Special Test Equipment or a digital volt.meter (DVM) at the input to the racks; CSA = {(SCA + SMTE + SD) 2 + (SPE) 2 + (STE)2+ (RDOUT) 2 } 112
SPE -. change in input-output relationship due to a change in static pressure for a differential pressure (d/p)
                    + BIAS                                                         Eq. 1
STE -change in input-output relationship due to a change in ambient temperature for a sensor or transmitter.
: 2.     For parameter indication utilizing the plant process computer; CSA   = {(SCA + SMTE + SD) 2 + (SPE) 2 + (STE) 2 + (RCA + RMTE + RD) 2
SD change in input-output relationship over a period of time at
                + (RTE) 2 + (ID) 2 + (A/D) 2 } 112 +BIAS                         Eq~ 2
* reference conditions for a sensor or transmitter . . RCA -reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated, .or tuned, to this accuracy.
: 3.     For parameters which have control systems; CSA= {(PMA) 2 + (PEA) 2 +(SCA+ SMTE + SD) 2 + (SPE) 2 + (STE) 2
* 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 RDOUT -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.
                + (RCA + RMTE + RD.+ CA) 2 + (RTE) 2 } 112 + 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
ID change in input-output relationship over a period of time at reference conditions for a control or protection signal isolating device. A/D -CA BIAS -allowance for conversion accuracy of an analog signal to a digital signal for process computer use. allowance for the accuracy of a not including deadband:
* included in the determination of control system uncertainties.
a non-random uncertainty for a sensor or transmitter or a process parameter.
where:
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 (through Rod Control).  
* CSA PMA PEA SCA
: 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 and controller.
                    =
As noted on Table l, the electronics uncertainty for this function is [
                    =
which corresponds to an accuracy of [ ra,c .. In addition to the controller accuracy, an allowance is made for pressure overshoot or undershoot due to the interaction and thermal inertia of the heaters and spray. Based on an evaluation of plant operation, an allowance of [
                    =
was made for this effect. An  
                    =
* *
Channel Allowance Process Measurement Accuracy Primary Element Accuracy Sensor Calibration Accuracy SMTE      =    Sensor Measurement and Test Equipment Accuracy SPE      =    Sensor Pressure Effects STE      =    Sensor Temperature Effects SD        =   Sensor Drift RCA      =   Rack Calibration Accuracy RMTE      =   Rack Measurement and Test Equipment Accuracy RTE      =   Rack Temperature Effects RD        =   Rack Drift RDOUT    =   Readout Device Accuracy (DVM or ga~g~)
* additional bias of [ J *a,c was included for a zero span shift during calibration of the transmitter.
ID        =   Computer Isolat~r Drift A/D      =   Analog to Digital Conversion Accuracy CA        =   Controller Accuracy
Therefore, a total control system uncertainty of [ ] *a,c is calculated, which results in a standard deviation of [ ra,c (assuming a normal, two sided probability distribution).
* The parameters above are as defined in references 5 and 12 and are based on SAMA Standard PMC 20.1, 1973(nl. However, for ease in Understanding they ate paraphrased below:
TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY SCA = SMTE= STE = SD = BIAS= RCA = RMTE= RTE = RD CA = +a,c ELECTRONICS UNCERTAINTY  
PMA -     non-instrument related measurement errors, e.g., temperature stratification of a fluid in a pipe.
= PLUS ELECTRONICS UNCERTAINTY  
PEA -     errors due to a metering device, e.g., elbow, venturi, orifice.
= PLUS CONTROLLER UNCERTAINTY
SCA -     reference (calibration) accuracy for a sensor or transmitter.
= +a,c
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.
* Tavg is controlled by a system that compares the auctioneered high Tavg from the loops with a reference, and derived the First Stage Turbine Impulse Chamber Pressure.
SD       change in input-output relationship over a period of time at
Tavg is the average of the narrow range TH and Tc values. The highest loop Tavg is.then used in the controller.
* reference conditions for a sensor or transmitter .
Allowances are made (as noted on Table 2) for the RTDs, transmitter and the process racks and controller.
        . RCA -     reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated,
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 one TH and one Tc Weed (N9004E-2A-SP)
                  .or tuned, to this accuracy.
RTDs are used to calculate and the RTDs are located in the Hot and Cold legs, the CSA for the electronics is [ ]+a,c. Assuming a normal,*two sided probability distribution results in an electronics standard deviation (a 1) of [ ]*a,c. However, this does not include the controller deadband of +/- 1.5 &deg;F. For the controller accuracy is the combination of the instrumentation accuracy and the The probability distribution for the deadband has been determined to be [ ] :a,c The variance for the deadband uncertainty is then: J +a,c. Combining the variance for instrumentation and deadband results in a controller variance of: J +a,c The  
RTE -   change in input-output relationship due to a change in ambient temperature for the rack modules.
[ ] +a,c for a total random uncertainty of [ ra,c. With the incorporation of Tcold streaming, an additional bi as of [ ] +a,c is included in Table 2.
RD       change in input-output relationship over a period of time at reference conditions for the rack modules.
* Therefore, the total uncertainty of the controller with the Tcold streaming included is [ J +a,c random and [ ] +a,c bi as. I '! 
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.
*
ID       change in input-output relationship over a period of time at reference conditions for a control or protection signal isolating device.
* TABLE 2 ROD CONTROL SYSTEM ACCURACY Tavg TURB PRES PMA = SCA = SMTE= STE = SD = BIAS= RCA = RMTE= RMTE= RTE = RD = CA = BIAS= _} +a,c # RTDs USED -TH = 3 TC = 1 ELECTRONICS CSA = ELECTRONICS SIGMA = CONTROLLER SIGMA = CONTROLLER BIAS = CONTROLLER CSA = +a,c *A Tcold bias of [ J+a,c in Tav, due to cold leg streaming, is not applicable when determining the DNBR core limits. A bias for Tcold streaming is accounted for in determining the Thermal Design Flow.  
A/D -     allowance for conversion accuracy of an analog signal to a digital signal for process computer use.
* *
CA        allowance for the accuracy of a controlle~. not including deadband:
* 3. RCS FLOW RTDP and plant Technical Specifications require an RCS flow measurement with a high degree of accuracy.
BIAS -  a non-random uncertainty for a sensor or transmitter or a process parameter.
It is assumed for this error analysis that the flow measurement is perfonned within thirty days of completing the calibration of the hot leg and cold leg narrow range RTDs. Therefore, drift effects are minimized.
A more detailed explanation of the Westinghouse methodology noting the interaction of several parameters is provided in references 5 and 12.
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., 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: where; WL = Loop fl ow ( gpm) A = 0.1247 gpm/(ft 3/hr) QSG = Steam Generator thermal output (Btu/hr) Qp = RCP heat addition (Btu/hr) QL = Primary system net heat losses (Btu/hr)  
III. INSTRUMENTATION UNCERTAINTIES The instrumentation uncertainties will be discussed first for the two parameters which are controlled by automatic systems, Pressurizer Pressure, and T~ 9 (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 and controller. As noted on Table l, the electronics uncertainty for this function is [               ]~.c which corresponds to an accuracy of [             ra,c .. In addition to the controller accuracy, an allowance is made for pressure overshoot or undershoot due to the interaction and thermal inertia of the heaters and spray. Based on an evaluation of plant
* Ve = Specific volume of the Cold Leg at Tc (ft 3/lb). N = Number of primary side loops hH = Hot Leg enthalpy (Btu/lb) he = Cold Leg enthalpy (Btu/lb).
* operation, an allowance of [           ]~.c was made for this effect.
The thermal output of the Steam Generator is determined by a precision secondary side calorimetric measurement, which is defined as: where; hs hf wf = =
An
= Steam enthalpy (Btu/lb) Feedwater enthalpy (Btu/lb) Feedwater flow (lb/hr). Eq. 6 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.
additional bias of [               J*a,c was included for a zero span shift during calibration of the transmitter.           Therefore, a total control system uncertainty of [             ] *a,c is calculated, which results in a standard deviation of
The Feedwater flow is determined by multiple measurements and the following calculation:
[           ra,c   (assuming a normal, two sided probability distribution).
Eq. 7 where; K = Feedwater venturi fl ow coefficient Fa = Feedwater venturi correction for thermal expansion Pt = Feedwater density (1 b/ft 3) d/p = Feedwater venturi pressure drop (inches H 2 0). 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.
TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY
Feedwater density is based on the measurement of Feedwater temperature and Feedwater pressure.
                                  +a,c SCA =
The venturi pressure drop is obtained from the output of the differential pressure cell connected to the venturi .  
SMTE=
*
STE =
* RCP heat addition is detennined by calculation, based .on the best estimate of coolant flow, pump head, and pump hydraulic efficiency.
SD   =
The primary system net heat losses are determined by calculation, considering the following system heat inputs and heat losses: Charging fl ow Letdown fl ow Seal injection flow RCP thennal barrier cooler heat rembval 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.
BIAS=
RCA =
RMTE=
RTE =
RD CA =
* ELECTRONICS UNCERTAINTY =
PLUS ELECTRONICS UNCERTAINTY =
                                                                                    +a,c PLUS CONTROLLER UNCERTAINTY =
* Tavg is controlled by a system that compares the auctioneered high Tavg from the loops with a reference, and derived ~rom the First Stage Turbine Impulse Chamber Pressure. Tavg is the average of the narrow range TH and Tc values.
The highest loop Tavg 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 one TH and one Tc cross~calibrated Weed (N9004E-2A-SP) RTDs are used to calculate T~ 9 and the RTDs are located in the Hot and Cold legs, the CSA for the electronics is
[       ]+a,c. Assuming a normal,*two sided probability distribution results in an electronics standard deviation (a1) of [             ]*a,c.
However, this does not include the controller deadband of +/- 1.5 &deg;F. For Ta~
the controller accuracy is the combination of the instrumentation accuracy and the deadb~nd. The probability distribution for the deadband has been
* determined to be [
                  ] :a,c The variance for the deadband uncertainty is then:
J+a,c.
Combining the variance for instrumentation and deadband results in a controller variance of:
J+a,c The controller~= [               ] +a,c for a total random uncertainty of [               ra,c.
With the incorporation of Tcold streaming, an additional bi as of [           ] +a,c is included in Table 2.
* Therefore, the total uncertainty of the controller with the Tcold streaming included is [           J+a,c random and [     ] +a,c bi as.
I
 
TABLE 2
* PMA =
ROD CONTROL SYSTEM ACCURACY Tavg TURB PRES
                                      +a,c SCA =
SMTE=
STE =
SD =
BIAS=
RCA =
RMTE=
RMTE=
RTE =
RD =
CA =
BIAS=                 _}
              # RTDs USED -   TH = 3     TC = 1
                                                      +a,c ELECTRONICS CSA =
ELECTRONICS SIGMA =
CONTROLLER SIGMA =
CONTROLLER BIAS =
CONTROLLER CSA     =
  *A Tcold bias of [     J+a,c in Tav, due to cold leg streaming, is not applicable when determining the DNBR core limits. A bias for Tcold streaming is accounted for in determining the Thermal Design Flow.
: 3. RCS FLOW
* RTDP and plant Technical Specifications require an RCS flow measurement with a high degree of accuracy. It is assumed for this error analysis that the flow measurement is perfonned within thirty days of completing the cross-calibration of the hot leg and cold leg narrow range RTDs. Therefore, drift effects are minimized. 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.,
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:
where; WL       =   Loop fl ow (gpm)
A         =   0.1247 gpm/(ft 3 /hr)
QSG       =   Steam Generator thermal output (Btu/hr)
Qp       =   RCP heat addition (Btu/hr)
QL       =   Primary system net heat losses (Btu/hr)
Ve       =     Specific volume of the Cold Leg at Tc (ft 3 /lb).
* N hH he
                    =
                    =
                    =
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:
Eq. 6 where;   hs   =   Steam enthalpy (Btu/lb) hf  =  Feedwater enthalpy (Btu/lb) wf  =  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:
Eq. 7 where;   K   =   Feedwater venturi fl ow coefficient Fa   =   Feedwater venturi correction for thermal expansion Pt   =   Feedwater density (1 b/ft 3) d/p =   Feedwater venturi pressure drop (inches H20).
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 detennined 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 fl ow Letdown fl ow Seal injection flow RCP thennal barrier cooler heat rembval 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 pressure.
The Cold Leg specific volume is based on measurement of the Cold Leg temperature and nominal Pressurizer pressure.
The RCS flow measurement is thus based on the following plant measurements:
The RCS flow measurement is thus based on the following plant measurements:
Steaml i ne pressure (P 5) Feedwater temperature (Tt) Feedwater venturi differential pressure (d/p) Hot Leg temperature (TH) Cold Leg temperature (Tc) Steam Generator blowdown (if not secured) and on the following calculated values: Pressurizer pressure (PP)  
Steaml i ne pressure (P 5 )
*
Feedwater temperature (Tt)
* Feedwater pressure (Pf) Feedwater venturi flow coefficients (K) Feedwater venturi thermal expansion correction (FJ Feedwater density (pf) Feedwater enthalpy (ht) Steam entha l PY (hs) Moisture carryover (impacts h 5) Primary system net heat losses (QL) RCP. heat addition (QP) Hot Leg entha l PY (hH) Cold Leg enthalpy (he)* These measurements and calculations are presented schematically on Figure 1: The derivation of the measurement errors and flow uncertainties on Table 5 are noted .below. Secondary Side The secondary side uncertainties are in four principal areas, Feedwater flow, Feedwater enthalpy, Steam enthalpy and RCP heat additi6n.
Feedwater venturi differential pressure (d/p)
These four areas are specifically identified on Table 5. For the measurement of Feedwater flow, each Feedwater venturi is calibrated by the vendor in a hydraulics laboratory under controlled conditions to an accuracy of [ ]+a,b,c. The calibration data which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of [ ra,c is included for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of [ ra,c. Si nee RCS loop fl ow is proportional to Steam Generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [ ra,c. It should be noted that no allowance is made for venturi fouling. The venturis should be inspected, and cleaned .if necessary.
Hot Leg temperature (TH)
prior to performance of the precision measurement.
Cold Leg temperature (Tc)
If fouling is present but not removed, its effects must be treated as a flow bias.  
Steam Generator blowdown (if not secured) and on the following calculated values:
* *
* Pressurizer pressure (PP)
* The uncertainty applied to the Feedwater venturi thermal expansion correction (Fa) 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&deg;F in the nominal Feedwater temperature range changes Fa by +/- 0.002 % and the Steam Generator thermal output by the same amount. An uncertainty in Fa of +/- 5 % for 304 stainless steel is used in this analysis.
Feedwater pressure (Pf)
This results in an additional uncertainty of [ ]ff,c in Feedwater flow. Westinghouse uses the conservative value of [ ]ff*c. Using the 1967 ASME Steam Tables it is possible to determine the sensitivities of various parameters to changes in Feedwater temperature and pressure. .Table 3 notes the instrument uncertainties for the hardware used to perform the measurements.
* Feedwater venturi flow coefficients (K)
Table 4 lists the various sensitivities.
Feedwater venturi thermal expansion correction (FJ Feedwater density (pf)
As can be seen on Table 4, Feedwater temperature uncertainties have an impact on venturi Feedwater .density and Feedwater enthalpy.
Feedwater enthalpy (ht)
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)(l/2)(transmitter span/100)2 Feedwater flow transmitter span is [ J +a,c of 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 [ Table 4. fa,c. This value is noted on The net pump heat uncertainty is derived from the combination of the primary system net heat losses and pump heat addition and are surrmarized for a four loop plant as follows:
Steam entha l PY (hs)
* System heat losses Component conduction and convection losses Pump heat adder Net Heat input to RCS -2.0 MWt -1.4 +15.4 +12.0 MWt The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [
Moisture carryover (impacts h5 )
of the *calculated value. Since direct measurements are not possible, the uncertainty
Primary system net heat losses (QL)
* on component conduction and convection losses has been assumed to be [
RCP. heat addition (QP)
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 II and by input power measurements from several plants, therefore, the uncertainty for the pump heat addition is estimated to be [ J*a,c of the best estimate value. Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are less than the value used in the analysis which is [ ] +a,c of core power. Primary Side The primary side uncertainties are in three principal areas, Hot Leg enthalpy, Cold Leg enthalpy and Cold Leg specific volume. These are specifically noted on Table 5. Three primary side parameters are actually measured, TH, Tc and* Pressurizer pressure.
Hot Leg entha l PY (hH)
Hot Leg enthalpy is influenced by TH, Pressurizer pressure and Hot Leg temperature streaming.
Cold Leg enthalpy (he)*
The uricertainties 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.
These measurements and calculations are presented schematically on Figure 1:
For the plants with direct ilTITlersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the streaming uncertainty is [ ra,c for both random and systematic For the Salem Units with RTDs located in thermowells placed in the scoops (bypass manifolds eliminated), the streaming uncertainty is, [ J +a,c random and [ ra,c systematic.
The derivation of the measurement errors and flow uncertainties on Table 5 are noted .below.
* *
Secondary Side
* The Cold Leg enthalpy and specific volume uncertainties are impacted by Tc and Pressurizer pressure.
* The secondary side uncertainties are in four principal areas, Feedwater flow, Feedwater enthalpy, Steam enthalpy and RCP heat additi6n. These four areas are specifically identified on Table 5.
Table 3 notes the Tc instrument uncertainty and Table 4 provides the sensitivities.
For the measurement of Feedwater flow, each Feedwater venturi is calibrated by the vendor in a hydraulics laboratory under controlled conditions to an accuracy of [                 ]+a,b,c. The calibration data which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of [               ra,c is included for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of
Noted on Table 5 is the plant specific RTD cross-calibration systematic allowance.
[             ra,c. Si nee RCS loop fl ow is proportional to Steam Generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [               ra,c. It should be noted that no allowance is made for venturi fouling. The venturis should be inspected, and cleaned .if necessary. prior to performance of the precision measurement. If fouling is present but not removed, its effects must be treated as a flow bias.
When necessary, an allowanc.e is made for a systematic temperature error due to the RTD cross-calibration procedure.
The uncertainty applied to the Feedwater venturi thermal expansion correction
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 4 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 is: # of loops flow uncertainty  
*    (Fa) 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&deg;F in the nominal Feedwater temperature range changes Fa by +/- 0.002 % and the Steam Generator thermal output by the same amount.
(% flow) 4 [ J +a,c
An uncertainty in Fa of +/- 5 % for 304 stainless steel is used in this analysis. This results in an additional uncertainty of [             ]ff,c in Feedwater flow. Westinghouse uses the conservative value of [               ]ff*c.
* TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION  
Using the 1967 ASME Steam Tables it is possible to determine the sensitivities of various parameters to changes in Feedwater temperature and pressure. .Table 3 notes the instrument uncertainties for the hardware used to perform the measurements. Table 4 lists the various sensitivities. As can be seen on Table 4, Feedwater temperature uncertainties have an impact on venturi F~
{% SPAN) FW TEMP FW PRES FW d/p STM PRESS TH SCA = SMTE= SPE = STE = SD = R/E = RDOUT= BIAS= CSA = # OF INST USED OF INST SPAN = 400 psi 1200 3 % d/p psi OF 120% Flow 1200 120 1 OF 120 PRZ PRESS +a,c 1 psi 800 INST UNC. :[ J +a,c {RANDOM) INST UNC. (BIAS) NOMINAL = 433 860 psi a 760 psi a 612.6 543.2 2250 psi a
Feedwater .density and Feedwater enthalpy. Feedwater pressure uncertainties impact Feedwater density and Feedwater enthalpy .
* The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures.  
* Feedwater venturi d/p uncertainties are converted to % Feedwater flow using
# The pressurizer pressure is not a measured parameter, therefore the control uncertainty is utilized and conservatively rounded from [ J +a,c
  *the following conversion factor:
* TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW Fa TEMPERATURE MATERIAL = DENSITY TEMPERATURE  
          % flow= (d/p uncertainty)(l/2)(transmitter span/100) 2 Th~  Feedwater flow transmitter span is [     J+a,c of nominal fl ow.
= PRESSURE DELTA P = FEEDWATER ENTHALPY TEMPERATURE  
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.
= PRESSURE = hs = hf = Dh(SG) = STEAM ENTHALPY PRESSURE = MOISTURE = HOT LEG ENTHALPY TEMPERATURE  
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 [       fa,c. This value is noted on Table 4.
= PRESSURE = hH = he = Dh(VESS) = Cp (TH) = COLD LEG ENTHALPY TEMPERATURE  
The net pump heat uncertainty is derived from the combination of the primary system net heat losses and pump heat addition and are surrmarized for a four
:[ PRESSURE Cp (Tc) = COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE :[ +a,c 1200.4 BTU/LBM 411.4 BTU/LBM 789.0 BTU/LBM 631.7 BTU/LBM 538.8 BTU/LBM 92.9 BTU/LBM 1. 504 BTU/ LBM-&deg; F 1.227 BTU/LBM-&deg;F  +a,c J +a,c J +a,c
* loop plant as follows:
* TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT FEEDWATER FLOW VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE A 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
* System heat losses Component conduction and convection losses Pump heat adder
* TEMPERATURE PRESSURE COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE RTD CROSS-CAL SYSTEMATIC ALLOWANCE INSTRUMENT ERROR * **, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS FLOW UNCERTAINTY
                                                                        - 2.0 MWt
+a,c 
                                                                        - 1.4
*
                                                                        +15.4 Net Heat input to RCS                              +12.0 MWt The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [               ]~.c of the
* TABLE 5 (CONTINUED)
  *calculated value. Since direct measurements are not possible, the uncertainty
CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT 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 .; SINGLE LOOP (WITHOUT BIAS VALUES) [ N LOOP UNCERTAINTY  
* on component conduction and convection losses has been assumed to be
.(WITHOUT BIAS VALUES) N LOOP UNCERTAINTY (WITH BIAS VALUES) FLOW UNCERTAINTY
[     ]~.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 II and by input power measurements from several plants, therefore, the uncertainty for the pump heat addition is estimated to be [         J*a,c of the best estimate value. Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are less than the value used in the analysis which is [             ] +a,c of core power.
+a,c J +a,c 
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, TH, Tc and*
*
Pressurizer pressure. Hot Leg enthalpy is influenced by TH, Pressurizer pressure and Hot Leg temperature streaming. The uricertainties 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 the plants with direct ilTITlersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the streaming uncertainty is [                 ra,c for both random and systematic ~omponents. For the Salem Units with RTDs located in thermowells placed in the scoops (bypass manifolds eliminated), the streaming uncertainty is, [           J+a,c random and [       ra,c systematic.
* As noted earlier, the precision flow calorimetric is used as the reference for determining the accuracy of the Cold Leg elbow taps. To perform the Technical Specification required surveillance Salem utilizes board mounted instrumentation.
The Cold Leg enthalpy and specific volume uncertainties are impacted by Tc and
Table 6 notes the instrument uncertainties for determining the accuracy of the elbow taps, assuming one elbow tap per loop. The d/p transmitter uncertainties are converteq to % flow on the same basis as the Feedwater venturi d/p. The elbow tap uncertainty is then combined with the prec1s1on flow calorimetric uncertainty.
* Pressurizer pressure. Table 3 notes the Tc instrument uncertainty and Table 4 provides the sensitivities.
This combination of uncertainties results in the following total flow uncertainty:  
Noted on Table 5 is the plant specific RTD cross-calibration systematic allowance. When necessary, an allowanc.e is made for a systematic temperature error due to the RTD cross-calibration procedure. No allowance was necessary for this plant.
.{ # of loops flow uncertainty  
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).
(% flow) 4 +/- 2 .8 The corresponding values used in RTDP are: # of loops standard deviation  
Using Table 5, the 4 loop uncertainty equation (with biases) is as follows:
(% flow) 4 [ J +a,c  
*                                                                          +a,c*
*
Based on the number of loops, number, type and measurement method of RTDs, and the vessel Delta-T, the flow is:
* TABLE 6 COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES  
        # of loops         flow uncertainty (% flow) 4             [     J +a,c TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION   UNCERTAINTI~S
% d/p SPAN % FLOW PMA = PEA = SCA = SMTE = SPE = STE = SD = RCA = RMTE = RTE = RD = ID = A/D = RDOUT= BIAS = FLOW CALORIM. BIAS = FLOW CALORIMETRIC  
{% SPAN)     FW TEMP   FW PRES   FW d/p STM PRESS     TH             PRZ PRESS
= INSTRUMENT SPAN = SINGLE LOOP ELBOW TAP FLOW UNC = N LOOP ELBOW TAP FLOW UNC = N LOOP RCS FLOW UNCERTAINTY (WITHOUT BIAS VALUES) = N LOOP RCS FLOW UNCERTAINTY (WITH BIAS VALUES) = +a,c +a,c 
                                                                                  +a,c SCA =
** 4. REACTOR POWER Generally a plant performs a primary/secondary side heat balance once every 24 hours when power is above 15% Rated Thermal Power. This heat balance is used to verify that the plant is operating within the limits of the Operating License and to adjust the Power Range Flux channels when the difference between the NIS and the heat balance is greater than that required by the plant Technical Specifications.
SMTE=
Assuming that the primary and secondary sides are in equilibrium; the core power is determined by sulTlTling 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 equatton for this calculation is: where; RP = N = OsG = Qp = QL = H = { (N) [QSG -Op + ( ( 100) RP= H Core power (% RTP) Number of primary side loops Eq.8 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 anaiysis (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.
SPE =
*
STE =
* 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.
SD =
The sensitivities calculated are the same as those noted for the secondary side on T_ab 1 e 4. As noted on Table 8, Westinghouse has determined the dependent sets in the calculation and the direction of interaction.
R/E =
This is the same as that performed for the RCS fl ow 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.
RDOUT=
Using the power uncertainty values noted on Table 8, the 4 loop uncertainty (with bias values) equation is as follows: +a,c Based on the number of loops and the instrument uncertainties for the four parameters, the power measurement uncertainty for the secondary side power calorimetric is: # of loops power uncertainty  
BIAS=
(% RTP) 4 [ J +a,c  
CSA =
*
  # OF INST USED                                           3        1        1 OF       psi     % d/p     psi       OF       OF      psi INST SPAN = 400          1200  120% Flow 1200         120       120     800
* IV. CONCLUSIONS The preceding sections provide the methodology to account for pressure, temperature, power and RCS flow uncertainties for the RTDP analysis.
* INST UNC.
The plant specific instrumentation data and procedures have been reviewed for Salem Units 1 & 2 and the uncertainty completed based on the Salem Units RTD Bypass Loop Elimination design. These uncertainty values or more conservative values are used in the RTDP analysis .  
{RANDOM)
* *
INST UNC.
* TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES  
(BIAS)   :[                                                                J
(% SPAN) FW TEMP FW PRES FW d/p STM PRESS SCA = SMTE= SPE = STE = SD = BIAS= RCA = RMTE= RTE = RD. = A/D = CSA = INST SPAN = INST UNC :[ (RANDOM) INST UNC (BIAS) NOMINAL = OF psi 400 1200 433 860 +a,c % d/p psi 120% Flow 1200 J +a,c psi a 760 psi a The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures .
                                                                                  +a,c NOMINAL     = 433       860 psi a         760 psi a 612.6   543.2     2250 psi a
* 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 BIAS VALUES FEEDWATER DELTA P FEEDWATER PRESSURE STEAM PRESSURE POWER BIAS TOTAL VALUE DENSITY ENTHALPY ENTHALPY INSTRUMENT ERROR *, ** INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES) N LOOP UNCERTAINTY (WITHOUT BIAS VALUES) N LOOP UNCERTAINTY (WITH BIAS VALUES) POWER UNCERTAINTY
* The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures.
+a,c 
  #     The pressurizer pressure is not a measured parameter, therefore the control uncertainty is utilized and conservatively rounded from
*-*
[                         J+a,c TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW Fa                                             +a,c TEMPERATURE MATERIAL         =
* V. REFERENCES  
DENSITY TEMPERATURE     =
: 1. Westinghouse letter NS-CE-1583, C. Eicheldinger to J. F. Stolz, NRC, dated 10/25/77.  
PRESSURE DELTA P               =
: 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.
FEEDWATER ENTHALPY TEMPERATURE     =
* 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. Su1T111er Nuclear Station Unit No. l, Docket 50-395, August, 1982. 8. Regulatory Gui'de 1.105 Rev. 2, 11 Instrument Set points for Safety-Related Systems 11 , dated 2/86. 9. NUREG/CR-3659 (PNL-4973), 11 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, 11 Criteria for Technical Specifications for Nuclear Power Stations 11* 11. ISA Standard S67.04, 1987, "Setpoints for Nuclear Safety-Related Instrumentation Used in Nuclear Power Plants 11*
PRESSURE         =
* 12. Tuley, C. R., Miller, R. B., "Westinghouse Setpoint Methodology for Control and Protection Systems 11 , IEEE Transactions on Nuclear Science, February, 1986, Vol. NS-33 No. l, pp. 684-687. 13. Scientific Apparatus Manufacturers Association, Standard PMC 20.1, 1973, 11 Process Measurement and Control Terminology 11* 14. Westinghouse WCAP-11397-P-A, 11 Revised Thermal Design Procedure 11 , dated April, 1989 .
hs           =   1200.4 BTU/LBM hf           =   411.4 BTU/LBM Dh(SG)       =   789.0 BTU/LBM STEAM ENTHALPY
SECONDARY SIDE PRIMARY SIDE
                                                          +a,c PRESSURE         =
* P, I I I Pt K I r w, ilh I I I -calculated D QSG [OJ -measured
MOISTURE         =
* I QJN L Qp ***1 I . I WL I L Other Loops I RCS FLOW Figure 1 RCS Flow Calorimetric Schematic
HOT LEG ENTHALPY TEMPERATURE     =
* SECONDARY SIDE [] pf [] G -I I I I hs hf Pt Fa K I .. . wf , I OsG D -calculat ed D -measured . ' " L Other Loops ' + + -QL L I Qp Core Power Figure 2 Power Calorimetric Schematic .}}
PRESSURE         =
hH           =   631.7  BTU/LBM he           =   538.8  BTU/LBM Dh(VESS)     =     92.9  BTU/LBM Cp (TH)     =   1. 504 BTU/ LBM- &deg; F COLD LEG ENTHALPY TEMPERATURE PRESSURE        :[                             J+a,c Cp (Tc)     =   1.227 BTU/LBM-&deg;F COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE         :[                             J +a,c TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT                        INSTRUMENT ERROR  FLOW UNCERTAINTY FEEDWATER FLOW
                                                                              +a,c VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE                A 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 RTD CROSS-CAL SYSTEMATIC ALLOWANCE
        * **, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS TABLE 5 (CONTINUED)
* BIAS VALUES CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT FEEDWATER PRESSURE         DENSITY FLOW UNCERTAINTY
                                                                              +a,c ENTHALPY STEAM PRESSURE             ENTHALPY PRESSURIZER PRESSURE       ENTHALPY - HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME - COLD LEG FLOW BIAS TOTAL VALUE J
SINGLE LOOP UNCERTAI~TY (WITHOUT BIAS VALUES)     [             +a,c N LOOP UNCERTAINTY     .(WITHOUT BIAS VALUES)
N LOOP UNCERTAINTY       (WITH BIAS VALUES)
As noted earlier, the precision flow calorimetric is used as the reference for determining the accuracy of the Cold Leg elbow taps. To perform the Technical Specification required surveillance Salem utilizes board mounted instrumentation. Table 6 notes the instrument uncertainties for determining the accuracy of the elbow taps, assuming one elbow tap per loop. The d/p transmitter uncertainties are converteq to % flow on the same basis as the Feedwater venturi d/p. The elbow tap uncertainty is then combined with the prec1s1on flow calorimetric uncertainty. This combination of uncertainties results in the following total flow uncertainty:
                                            .{
      # of loops   flow uncertainty (% flow) 4         +/- 2 .8 The corresponding values used in RTDP are:
      # of loops     standard deviation (% flow) 4     [       J   +a,c TABLE 6
* COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES
                % d/p SPAN     % FLOW
                                        +a,c PMA =
PEA =
SCA =
SMTE =
SPE =
STE =
SD =
RCA =
RMTE =
RTE =
RD =
ID   =
A/D =
RDOUT=
BIAS =
FLOW CALORIM. BIAS =
FLOW CALORIMETRIC =
INSTRUMENT SPAN       =
                                                      +a,c SINGLE LOOP ELBOW TAP FLOW UNC   =
N LOOP ELBOW TAP FLOW UNC       =
N LOOP RCS FLOW UNCERTAINTY (WITHOUT BIAS VALUES)       =
N LOOP RCS FLOW UNCERTAINTY (WITH BIAS VALUES)           =
: 4. REACTOR POWER
** Generally a plant performs a primary/secondary side heat balance once every 24 hours when power is above 15% Rated Thermal Power. This heat balance is used to verify that the plant is operating within the limits of the Operating License and to adjust the Power Range ~eutron 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 sulTlTling 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 equatton for this calculation is:
{ (N) [QSG - Op + ( ~)]} ( 100)
RP=                                       Eq.8 H
where; RP            =    Core power (% RTP)
N            =    Number of primary side loops OsG          =    Steam Generator thermal output (BTU/hr) as defined in Eq. 6 Qp            =    RCP heat adder (Btu/hr) as defined in Eq. 5 QL            =    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 anaiysis (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. The sensitivities calculated are the same as those noted for the secondary side on T_ab 1e 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 fl ow 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 4 loop uncertainty (with bias values) equation is as follows:
                                                                    +a,c Based on the number of loops and the instrument uncertainties for the four parameters, the power measurement uncertainty for the secondary side power calorimetric is:
        # of loops             power uncertainty (% RTP) 4                   [       J   +a,c 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 have been reviewed for Salem Units 1 & 2 and the uncertainty c~lculations completed based on the Salem Units RTD Bypass Loop Elimination design. These uncertainty values or more conservative values are used in the RTDP analysis .
TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES
(% SPAN)   FW TEMP FW PRES   FW d/p STM PRESS
                                                          +a,c SCA =
SMTE=
SPE =
STE =
SD =
BIAS=
RCA =
RMTE=
RTE =
RD. =
A/D =
CSA =
OF      psi      % d/p      psi
* INST SPAN INST UNC (RANDOM)
                =
:[
400     1200       120% Flow     1200
                                                            +a,c INST UNC (BIAS)
NOMINAL    =  433    860 psi a J
760 psi a
* The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures .
TABLE 8 SECONDARY SIDE POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES COMPONENT                           INSTRUMENT ERROR  POWER UNCERTAINTY
                                                                                +a,c FEEDWATER FLOW VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION BIAS VALUES FEEDWATER DELTA P FEEDWATER PRESSURE       DENSITY ENTHALPY STEAM PRESSURE           ENTHALPY POWER BIAS TOTAL VALUE
        *, **   INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)
N LOOP UNCERTAINTY     (WITHOUT BIAS VALUES)
N LOOP UNCERTAINTY     (WITH BIAS VALUES)
V. REFERENCES
*- 1. 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
* 7.
Company, dated 2/12/81.
NUREG-0717 Supplement No. 4, Safety Evaluation Report related to the operation of Virgil C. Su1T111er Nuclear Station Unit No. l, Docket 50-395, August, 1982.
: 8. Regulatory Gui'de 1.105 Rev. 2,     11 Instrument Set points for Safety-Related Systems 11 , dated 2/86.
: 9. NUREG/CR-3659 (PNL-4973), 11 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,   11 Criteria for Technical Specifications for Nuclear Power Stations 11 *
: 11. ISA Standard S67.04, 1987, "Setpoints for Nuclear Safety-Related
* Instrumentation Used in Nuclear Power Plants 11 *
: 12. Tuley, C. R., Miller, R. B., "Westinghouse Setpoint Methodology for Control and Protection Systems 11 , IEEE Transactions on Nuclear Science, February, 1986, Vol. NS-33 No. l, pp. 684-687.
: 13. Scientific Apparatus Manufacturers Association, Standard PMC 20.1, 1973, 11 Process Measurement and Control Terminology 11 *
: 14. Westinghouse WCAP-11397-P-A, 11 Revised Thermal Design Procedure 11 , dated April, 1989 .
SECONDARY SIDE                             PRIMARY SIDE
* P, I
I           I Pt                     K I
r w,                       ilh I   II                      - calculated D
QSG                       - measured
[OJ
* QJN             L I
Qp
                                                                      ***1 I                                                 . I WL I
L             Other Loops I
RCS FLOW Figure 1 RCS Flow Calorimetric Schematic SECONDARY SIDE
[] pf I
[]           G     -
I I                   I hs hf                 Pt             Fa         K
                                          ..       I wf ,
I OsG D - calculat ed D - measured   .
L           Other Loops
                          +
              +                 -
QL                   L           I Qp Core Power Figure 2 Power Calorimetric Schematic
                        .}}

Latest revision as of 09:00, 23 February 2020

Non-proprietary Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Salem Units 1 & 2.
ML18101B377
Person / Time
Site: Salem  PSEG icon.png
Issue date: 08/31/1993
From: Ciocca C
WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML18101B376 List:
References
WCAP-13652, NUDOCS 9605220058
Download: ML18101B377 (34)
v • d • e


Text

WESTINGHOUSE CLASS 3 (NON-PROPRIETARY)

WCAP~13652 WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY SALEM UNITS 1 & 2

  • August, 1993 C. F. Ciocca WESTINGHOUSE ELECTRIC CORPORATION Nuclear & Advanced Technology Division P.O. Box 355 Pittsburgh, Pennsylvania 15230-0355

~ 1993 Westinghouse Electric Corp., All Rights Reserved 96o522oo5s - 96082-5-----=-- -----1 PDR ADOCK 05000272 I

p. PDR I

TABLE OF CONTENTS Section Title Page I. INTRO DU CTI ON 1 I I. METHODOLOGY 3 III. INSTRUMENTATION UNCERTAINTIES 5 IV. CONCLUSIONS 24

v. REFERENCES 27

LIST OF TABLES Table Number Title Page 1 Pressurizer Pressure Control System Accuracy 6 2 Rod Control System Accuracy 8 3 Flow Calorimetric Instrumentation Uncertainties 16 4 Flow Calorimetric Sensitivities 17 5 Calorimetric RCS Flow Measurement Uncertainties 18 6 Cold Leg Elbow Tap Flow Uncertainty 21 7 Power Calorimetric Instrumentation Uncertainties 25 8 Secondary Side Power Calorimetric Measurement Uncertainties 26 ii

LIST OF FIGURES Figure Number 1 RCS Flow Calorimetric Schematic 29 2 Power Calorimetric Schematic 30

  • iii

WESTINGHOUSE REVISED THERMAL DESIGN PROCEDURE INSTRUMENT UNCERTAINTY METHODOLOGY 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 (l'avg), 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 fl ow is monitored by the performance of a precision fl ow calorimetric at the beginning of each cycle. The RCS Cold Leg elbow taps are evaluated against the precision calnrimetric and used for monthly surveillance (with a small increase in uncertainty). Pressurizer pressure is a controlled parameter and the uncertainty reflects the control system. Tavg is a controlled parameter via the temperature input to the rod control system and the uncertainty

  • reflects this control systemL This report is based on the elimination of RTD Bypass Loops in the design to measure hot and cold leg reactor coolant system temperritures. The RTDP< 1 ~ 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 Salem 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 11 Improved Thermal Design Procedure", 11 *2 *3l which is based on the conservative assumption that the uncertainties can be described with uniform probability distributions.

Another approach (for McGuire and Catawba) is based on the more realistic assumption that the uncertainties can be described with random, normal, two sided probability distributions. 141 This approach is used to substantiate the

  • acceptability of the protection system setpoints for many Westinghouse plants, e.g., D. C. Cook 2151, V. C. Surrmer, Wolf Creek, Millstone Unit 3 and others.

The second approach is now utilized for the determination of a11 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 th~ ium of the squares of tho~e groups of components which are statistically independent. Those errors that are dependent are combined arithmetically into independen.t 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 [ J+a,c.

This technique has been utilized before as noted above, and has been endorsed. by the NRC staff 16 *7*8 *9' and various industry standards 00

  • 11 '.

The relationships between th~ error components and the channel instrument error allowance are variations of the basic Westinghouse Setpoint Methodology! 12 l and are defined as follows:

1. For precision parameter indication using Special Test Equipment or a digital volt.meter (DVM) at the input to the racks; CSA = {(SCA + SMTE + SD) 2 + (SPE) 2 + (STE)2+ (RDOUT) 2 } 112

+ BIAS Eq. 1

2. For parameter indication utilizing the plant process computer; CSA = {(SCA + SMTE + SD) 2 + (SPE) 2 + (STE) 2 + (RCA + RMTE + RD) 2

+ (RTE) 2 + (ID) 2 + (A/D) 2 } 112 +BIAS Eq~ 2

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

+ (RCA + RMTE + RD.+ CA) 2 + (RTE) 2 } 112 + 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 PMA PEA SCA

=

=

=

=

Channel Allowance Process Measurement Accuracy Primary Element Accuracy Sensor Calibration Accuracy SMTE = Sensor Measurement and Test Equipment Accuracy SPE = Sensor Pressure Effects STE = Sensor Temperature Effects SD = Sensor Drift RCA = Rack Calibration Accuracy RMTE = Rack Measurement and Test Equipment Accuracy RTE = Rack Temperature Effects RD = Rack Drift RDOUT = Readout Device Accuracy (DVM or ga~g~)

ID = Computer Isolat~r Drift A/D = Analog to Digital Conversion Accuracy CA = Controller Accuracy

  • The parameters above are as defined in references 5 and 12 and are based on SAMA Standard PMC 20.1, 1973(nl. However, for ease in Understanding they ate paraphrased below:

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

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

SCA - reference (calibration) accuracy for a sensor or 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 - reference (calibration) accuracy for all rack modules in loop or channel assuming the loop or channel is string calibrated,

.or tuned, to this accuracy.

RTE - change in input-output relationship due to a change in ambient temperature for the rack modules.

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.

ID change in input-output relationship over a period of time at reference conditions for a control or protection signal isolating device.

A/D - allowance for conversion accuracy of an analog signal to a digital signal for process computer use.

CA allowance for the accuracy of a controlle~. not including deadband:

BIAS - 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~ 9 (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 and controller. As noted on Table l, the electronics uncertainty for this function is [ ]~.c which corresponds to an accuracy of [ ra,c .. In addition to the controller accuracy, an allowance is made for pressure overshoot or undershoot due to the interaction and thermal inertia of the heaters and spray. Based on an evaluation of plant
  • operation, an allowance of [ ]~.c was made for this effect.

An

additional bias of [ J*a,c was included for a zero span shift during calibration of the transmitter. Therefore, a total control system uncertainty of [ ] *a,c is calculated, which results in a standard deviation of

[ ra,c (assuming a normal, two sided probability distribution).

TABLE 1 PRESSURIZER PRESSURE CONTROL SYSTEM ACCURACY

+a,c SCA =

SMTE=

STE =

SD =

BIAS=

RCA =

RMTE=

RTE =

RD CA =

  • ELECTRONICS UNCERTAINTY =

PLUS ELECTRONICS UNCERTAINTY =

+a,c PLUS CONTROLLER UNCERTAINTY =

  • Tavg is controlled by a system that compares the auctioneered high Tavg from the loops with a reference, and derived ~rom the First Stage Turbine Impulse Chamber Pressure. Tavg is the average of the narrow range TH and Tc values.

The highest loop Tavg 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 one TH and one Tc cross~calibrated Weed (N9004E-2A-SP) RTDs are used to calculate T~ 9 and the RTDs are located in the Hot and Cold legs, the CSA for the electronics is

[ ]+a,c. Assuming a normal,*two sided probability distribution results in an electronics standard deviation (a1) of [ ]*a,c.

However, this does not include the controller deadband of +/- 1.5 °F. For Ta~

the controller accuracy is the combination of the instrumentation accuracy and the deadb~nd. The probability distribution for the deadband has been

  • determined to be [

] :a,c The variance for the deadband uncertainty is then:

J+a,c.

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

J+a,c The controller~= [ ] +a,c for a total random uncertainty of [ ra,c.

With the incorporation of Tcold streaming, an additional bi as of [ ] +a,c is included in Table 2.

  • Therefore, the total uncertainty of the controller with the Tcold streaming included is [ J+a,c random and [ ] +a,c bi as.

I

TABLE 2

  • PMA =

ROD CONTROL SYSTEM ACCURACY Tavg TURB PRES

+a,c SCA =

SMTE=

STE =

SD =

BIAS=

RCA =

RMTE=

RMTE=

RTE =

RD =

CA =

BIAS= _}

  1. RTDs USED - TH = 3 TC = 1

+a,c ELECTRONICS CSA =

ELECTRONICS SIGMA =

CONTROLLER SIGMA =

CONTROLLER BIAS =

CONTROLLER CSA =

  • A Tcold bias of [ J+a,c in Tav, due to cold leg streaming, is not applicable when determining the DNBR core limits. A bias for Tcold streaming is accounted for in determining the Thermal Design Flow.
3. RCS FLOW
  • RTDP and plant Technical Specifications require an RCS flow measurement with a high degree of accuracy. It is assumed for this error analysis that the flow measurement is perfonned within thirty days of completing the cross-calibration of the hot leg and cold leg narrow range RTDs. Therefore, drift effects are minimized. 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.,

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:

where; WL = Loop fl ow (gpm)

A = 0.1247 gpm/(ft 3 /hr)

QSG = Steam Generator thermal output (Btu/hr)

Qp = RCP heat addition (Btu/hr)

QL = Primary system net heat losses (Btu/hr)

Ve = Specific volume of the Cold Leg at Tc (ft 3 /lb).

  • N hH he

=

=

=

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:

Eq. 6 where; hs = Steam enthalpy (Btu/lb) hf = Feedwater enthalpy (Btu/lb) wf = 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:

Eq. 7 where; K = Feedwater venturi fl ow coefficient Fa = Feedwater venturi correction for thermal expansion Pt = Feedwater density (1 b/ft 3) d/p = Feedwater venturi pressure drop (inches H20).

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 detennined 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 fl ow Letdown fl ow Seal injection flow RCP thennal barrier cooler heat rembval 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 pressure.

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

Steaml i ne pressure (P 5 )

Feedwater temperature (Tt)

Feedwater venturi differential pressure (d/p)

Hot Leg temperature (TH)

Cold Leg temperature (Tc)

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

  • Pressurizer pressure (PP)

Feedwater pressure (Pf)

Feedwater venturi thermal expansion correction (FJ Feedwater density (pf)

Feedwater enthalpy (ht)

Steam entha l PY (hs)

Moisture carryover (impacts h5 )

Primary system net heat losses (QL)

RCP. heat addition (QP)

Hot Leg entha l PY (hH)

Cold Leg enthalpy (he)*

These measurements and calculations are presented schematically on Figure 1:

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

Secondary Side

  • The secondary side uncertainties are in four principal areas, Feedwater flow, Feedwater enthalpy, Steam enthalpy and RCP heat additi6n. These four areas are specifically identified on Table 5.

For the measurement of Feedwater flow, each Feedwater venturi is calibrated by the vendor in a hydraulics laboratory under controlled conditions to an accuracy of [ ]+a,b,c. The calibration data which substantiates this accuracy is provided to the plant by the vendor. An additional uncertainty factor of [ ra,c is included for installation effects, resulting in a conservative overall flow coefficient (K) uncertainty of

[ ra,c. Si nee RCS loop fl ow is proportional to Steam Generator thermal output which is proportional to Feedwater flow, the flow coefficient uncertainty is expressed as [ ra,c. It should be noted that no allowance is made for venturi fouling. The venturis should be inspected, and cleaned .if necessary. prior to performance of the precision measurement. If fouling is present but not removed, its effects must be treated as a flow bias.

The uncertainty applied to the Feedwater venturi thermal expansion correction

  • (Fa) 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 Fa by +/- 0.002 % and the Steam Generator thermal output by the same amount.

An uncertainty in Fa of +/- 5 % for 304 stainless steel is used in this analysis. This results in an additional uncertainty of [ ]ff,c in Feedwater flow. Westinghouse uses the conservative value of [ ]ff*c.

Using the 1967 ASME Steam Tables it is possible to determine the sensitivities of various parameters to changes in Feedwater temperature and pressure. .Table 3 notes the instrument uncertainties for the hardware used to perform the measurements. Table 4 lists the various sensitivities. As can be seen on Table 4, Feedwater temperature uncertainties have an impact on venturi F~

Feedwater .density and Feedwater enthalpy. Feedwater pressure uncertainties impact Feedwater density and Feedwater enthalpy .

  • the following conversion factor:

% flow= (d/p uncertainty)(l/2)(transmitter span/100) 2 Th~ Feedwater flow transmitter span is [ J+a,c of 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 [ fa,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 surrmarized for a four

  • loop plant as follows:
  • System heat losses Component conduction and convection losses Pump heat adder

- 2.0 MWt

- 1.4

+15.4 Net Heat input to RCS +12.0 MWt The uncertainty on system heat losses, which is essentially all due to charging and letdown flows, has been estimated to be [ ]~.c of the

  • calculated value. Since direct measurements are not possible, the uncertainty
  • on component conduction and convection losses has been assumed to be

[ ]~.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 II and by input power measurements from several plants, therefore, the uncertainty for the pump heat addition is estimated to be [ J*a,c of the best estimate value. Considering these parameters as one quantity, which is designated the net pump heat uncertainty, the combined uncertainties are less than the value used in the analysis which is [ ] +a,c of core power.

Primary Side The primary side uncertainties are in three principal areas, Hot Leg enthalpy, Cold Leg enthalpy and Cold Leg specific volume. These are specifically noted on Table 5. Three primary side parameters are actually measured, TH, Tc and*

Pressurizer pressure. Hot Leg enthalpy is influenced by TH, Pressurizer pressure and Hot Leg temperature streaming. The uricertainties 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 the plants with direct ilTITlersion RTDs located in RTD bypass manifolds fed by scoops in the legs, the streaming uncertainty is [ ra,c for both random and systematic ~omponents. For the Salem Units with RTDs located in thermowells placed in the scoops (bypass manifolds eliminated), the streaming uncertainty is, [ J+a,c random and [ ra,c systematic.

The Cold Leg enthalpy and specific volume uncertainties are impacted by Tc and

  • Pressurizer pressure. Table 3 notes the Tc instrument uncertainty and Table 4 provides the sensitivities.

Noted on Table 5 is the plant specific RTD cross-calibration systematic allowance. When necessary, an allowanc.e 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 4 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 is:

  1. of loops flow uncertainty (% flow) 4 [ J +a,c TABLE 3 FLOW CALORIMETRIC INSTRUMENTATION UNCERTAINTI~S

{% SPAN) FW TEMP FW PRES FW d/p STM PRESS TH PRZ PRESS

+a,c SCA =

SMTE=

SPE =

STE =

SD =

R/E =

RDOUT=

BIAS=

CSA =

  1. OF INST USED 3 1 1 OF psi  % d/p psi OF OF psi INST SPAN = 400 1200 120% Flow 1200 120 120 800
  • INST UNC.

{RANDOM)

INST UNC.

(BIAS)  :[ J

+a,c NOMINAL = 433 860 psi a 760 psi a 612.6 543.2 2250 psi a

  • The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures.
  1. The pressurizer pressure is not a measured parameter, therefore the control uncertainty is utilized and conservatively rounded from

[ J+a,c TABLE 4 FLOW CALORIMETRIC SENSITIVITIES FEEDWATER FLOW Fa +a,c TEMPERATURE MATERIAL =

DENSITY TEMPERATURE =

PRESSURE DELTA P =

FEEDWATER ENTHALPY TEMPERATURE =

PRESSURE =

hs = 1200.4 BTU/LBM hf = 411.4 BTU/LBM Dh(SG) = 789.0 BTU/LBM STEAM ENTHALPY

+a,c PRESSURE =

MOISTURE =

HOT LEG ENTHALPY TEMPERATURE =

PRESSURE =

hH = 631.7 BTU/LBM he = 538.8 BTU/LBM Dh(VESS) = 92.9 BTU/LBM Cp (TH) = 1. 504 BTU/ LBM- ° F COLD LEG ENTHALPY TEMPERATURE PRESSURE  :[ J+a,c Cp (Tc) = 1.227 BTU/LBM-°F COLD LEG SPECIFIC VOLUME TEMPERATURE PRESSURE  :[ J +a,c TABLE 5 CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT INSTRUMENT ERROR FLOW UNCERTAINTY FEEDWATER FLOW

+a,c VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE A 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 RTD CROSS-CAL SYSTEMATIC ALLOWANCE
  • **, +, ++ INDICATE SETS OF DEPENDENT PARAMETERS TABLE 5 (CONTINUED)
  • BIAS VALUES CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES COMPONENT FEEDWATER PRESSURE DENSITY FLOW UNCERTAINTY

+a,c ENTHALPY STEAM PRESSURE ENTHALPY PRESSURIZER PRESSURE ENTHALPY - HOT LEG ENTHALPY - COLD LEG SPECIFIC VOLUME - COLD LEG FLOW BIAS TOTAL VALUE J

SINGLE LOOP UNCERTAI~TY (WITHOUT BIAS VALUES) [ +a,c N LOOP UNCERTAINTY .(WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES)

As noted earlier, the precision flow calorimetric is used as the reference for determining the accuracy of the Cold Leg elbow taps. To perform the Technical Specification required surveillance Salem utilizes board mounted instrumentation. Table 6 notes the instrument uncertainties for determining the accuracy of the elbow taps, assuming one elbow tap per loop. The d/p transmitter uncertainties are converteq to % flow on the same basis as the Feedwater venturi d/p. The elbow tap uncertainty is then combined with the prec1s1on flow calorimetric uncertainty. This combination of uncertainties results in the following total flow uncertainty:

.{

  1. of loops flow uncertainty (% flow) 4 +/- 2 .8 The corresponding values used in RTDP are:
  1. of loops standard deviation (% flow) 4 [ J +a,c TABLE 6
  • COLD LEG ELBOW TAP FLOW UNCERTAINTY INSTRUMENT UNCERTAINTIES

% d/p SPAN  % FLOW

+a,c PMA =

PEA =

SCA =

SMTE =

SPE =

STE =

SD =

RCA =

RMTE =

RTE =

RD =

ID =

A/D =

RDOUT=

BIAS =

FLOW CALORIM. BIAS =

FLOW CALORIMETRIC =

INSTRUMENT SPAN =

+a,c SINGLE LOOP ELBOW TAP FLOW UNC =

N LOOP ELBOW TAP FLOW UNC =

N LOOP RCS FLOW UNCERTAINTY (WITHOUT BIAS VALUES) =

N LOOP RCS FLOW UNCERTAINTY (WITH BIAS VALUES) =

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 ~eutron 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 sulTlTling 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 equatton for this calculation is:

{ (N) [QSG - Op + ( ~)]} ( 100)

RP= Eq.8 H

where; RP = Core power (% RTP)

N = Number of primary side loops OsG = Steam Generator thermal output (BTU/hr) as defined in Eq. 6 Qp = RCP heat adder (Btu/hr) as defined in Eq. 5 QL = 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 anaiysis (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. The sensitivities calculated are the same as those noted for the secondary side on T_ab 1e 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 fl ow 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 4 loop uncertainty (with bias values) equation is as follows:

+a,c Based on the number of loops and the instrument uncertainties for the four parameters, the power measurement uncertainty for the secondary side power calorimetric is:

  1. of loops power uncertainty (% RTP) 4 [ J +a,c 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 have been reviewed for Salem Units 1 & 2 and the uncertainty c~lculations completed based on the Salem Units RTD Bypass Loop Elimination design. These uncertainty values or more conservative values are used in the RTDP analysis .

TABLE 7 POWER CALORIMETRIC INSTRUMENTATION UNCERTAINTIES

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

+a,c SCA =

SMTE=

SPE =

STE =

SD =

BIAS=

RCA =

RMTE=

RTE =

RD. =

A/D =

CSA =

OF psi  % d/p psi

  • INST SPAN INST UNC (RANDOM)

=

[

400 1200 120% Flow 1200

+a,c INST UNC (BIAS)

NOMINAL = 433 860 psi a J

760 psi a

  • The feedwater pressure is not a measured parameter, therefore the uncertainty is conservatively assumed to be approximately the difference between the nominal steam and nominal feedwater pressures .

TABLE 8 SECONDARY SIDE POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES COMPONENT INSTRUMENT ERROR POWER UNCERTAINTY

+a,c FEEDWATER FLOW VENTURI THERMAL EXPANSION COEFFICIENT TEMPERATURE MATERIAL DENSITY TEMPERATURE PRESSURE DELTA P FEEDWATER ENTHALPY TEMPERATURE PRESSURE STEAM ENTHALPY PRESSURE MOISTURE NET PUMP HEAT ADDITION BIAS VALUES FEEDWATER DELTA P FEEDWATER PRESSURE DENSITY ENTHALPY STEAM PRESSURE ENTHALPY POWER BIAS TOTAL VALUE

  • , ** INDICATE SETS OF DEPENDENT PARAMETERS SINGLE LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITHOUT BIAS VALUES)

N LOOP UNCERTAINTY (WITH BIAS VALUES)

V. REFERENCES

  • - 1. 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
  • 7.

Company, dated 2/12/81.

NUREG-0717 Supplement No. 4, Safety Evaluation Report related to the operation of Virgil C. Su1T111er Nuclear Station Unit No. l, Docket 50-395, August, 1982.

8. Regulatory Gui'de 1.105 Rev. 2, 11 Instrument Set points for Safety-Related Systems 11 , dated 2/86.
9. NUREG/CR-3659 (PNL-4973), 11 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, 11 Criteria for Technical Specifications for Nuclear Power Stations 11 *
11. ISA Standard S67.04, 1987, "Setpoints for Nuclear Safety-Related
  • Instrumentation Used in Nuclear Power Plants 11 *
12. Tuley, C. R., Miller, R. B., "Westinghouse Setpoint Methodology for Control and Protection Systems 11 , IEEE Transactions on Nuclear Science, February, 1986, Vol. NS-33 No. l, pp. 684-687.
13. Scientific Apparatus Manufacturers Association, Standard PMC 20.1, 1973, 11 Process Measurement and Control Terminology 11 *
14. Westinghouse WCAP-11397-P-A, 11 Revised Thermal Design Procedure 11 , dated April, 1989 .

SECONDARY SIDE PRIMARY SIDE

  • P, I

I I Pt K I

r w, ilh I II - calculated D

QSG - measured

[OJ

  • QJN L I

Qp

      • 1 I . I WL I

L Other Loops I

RCS FLOW Figure 1 RCS Flow Calorimetric Schematic SECONDARY SIDE

[] pf I

[] G -

I I I hs hf Pt Fa K

.. I wf ,

I OsG D - calculat ed D - measured .

L Other Loops

+

+ -

QL L I Qp Core Power Figure 2 Power Calorimetric Schematic

.

Retrieved from "https://kanterella.com/w/index.php?title=ML18101B377&oldid=2505509"