Text

Requests Proprietary Responses to NRC Questions Re Optimized Fuel Assembly Design Be Withheld (Ref 10CFR2.790).Affidavit Encl
	ML20095F318
Person / Time
Site:	Point Beach
Issue date:	08/07/1984
From:	Wiesemann R; WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:	Harold Denton; Office of Nuclear Reactor Regulation
Shared Package
ML19269A439	List: ML20095F276; ML20095F318;
References
	CAW-84-78, TAC-49969, TAC-49970, NUDOCS 8408270246
	Download: ML20095F318 (111)
	v • d • e

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Westinghouse Water Reactor Nm Men Electric Corporation Divisions ecx 3912 Pascuri;n Pemsylvania 15230 CAW-84-78 i

August 7, 1984 Mi. Harold R. Denton Director Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Comission Phillips Building 7920 Norfolk Avenue Bethesda, Maryland 20014

Dear Mr. Denton:

APPLICATION FOR WITHHOLDING PROPRIETARY INFORMATION FROM PUBLIC DISCLO5URE

Reference:

Wisconsin Electric Power Company letters, Fay to Denton, dated March 14, 1983 and September 6, 1983 The proprietary material for which withholding is being requested by the Wisconsin Electric Power Company is proprietary to Westinghouse and withholding is requested pursuant to the provisions of Paragraph (b)(1) of Section 2.790 of the Comission's regulations. Withholding from public disclosure is requested with respect to the subject information which is further identified in the affidavit accompanying this application.

The proprietary material transmitted by 'the referenced letter supplements the proprietary material previously submitted. Further, the affidavit submitted to justify the previous material was approved by the Comission on April 17, 1978, and is equally applicable to the subject material.

Accordingly, withholding the subject information from public disclosure is requested in accordance with the previously submitted affidavit, AW-76-60, a copy of which is attached.

Accordingly, this letter authorized the use of the proprietary information and affidavit AW-76-60 by the Wisconsin Electric Power Company for the Point Beach Nuclear Plant.

8408270246 840817 PDR ADOCK 05000266 P

PDR

l Mr. Harold R. Denton August.7, 1984 Correspondence with respect to this application for withholding or the accompanying affidavit should reference CAW-84-78 and be addressed to the undersigned.

Ver truly yours, LL "1 %

Ro ert A. Wiesemann, Manager Regulatory and Legislative Affairs Enclosures (s) cc:

E. C. Shomaker, Esq.

Office of the Executive Legal Director, NRC

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3 ENCLOSURE 4

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AW-76-60 AFFICAVIT r

COMMCfGEALTH OF PEN!iSYLVANIA:

ss CDUtiTY OF ALLEGHEtiY:

Before me, the undersigned authority, personally appeared Robert A. Wiesemann, who, being by me duly sworn ace:rcing to law, de-poses and says that he is authorized to execute this Affidavit on behalf of W'estinghouse Electric Cor;cration ("'.lestinghouse") and that the aver-ments of fact set forth in this Affidavit are true and correct to the i

best of his knowledge, information, and belief:

'b<.L. !!-

.:* 2.%s RoDert A. Wiesemann, Manager Licensing Programs Sworn to and subscribed before,me this 2 day of 80!/.ird21) 1976.

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, AW-76-60 (1) I am Manager, Licensing ' Programs, in the Pressurized Water Reactor Systems Divisien, of Westinghouse Electric Corporation and as such,

.I have been specifically delegated the function of reviewing the proprietary information sought to be withheld frem public dis-closure in connection with nuclear power plant licensing or rule-making proceedings, and am authori ed to apply for its withholding

n behalf of the Westinghouse Water Reacter Divisiens.

o (2), I am making this Affidavit in confornance with the provisions of 10 CFR Section 2.790 of the Comissicn's regulatiens and in con-junction with the Westinghouse application for withholding ac-companying this Affidavit.

(3) I have personal knculadgs of the critaria and precedures utilized by Westinghcuse Nuclear Energy Systems in designating infer =atien h

as a trade secret, privileged er as conficantial exxurcial or financial infcrmation.

(4) Pursuant to the previsions of paragraph (b)(4) of Section 2.790 of the Counission's regulations, the following is furnished for consideration by the Comission in determining whether the in-formation scusht to be withheld frem public disclosure should be wi thhe1,d.'

(1) The informatien sought to be withheld frem public disclesure is owned and has been tield in confidence by 'destinghouse.

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' AW-7G-60 (ii) The infomation is of a type cust:marily held in confidence by Westinghouse and not cust:msrily disclosed to the public.

Westinghouse has a raticnal basis for determining the types of infomation customarily held in c:nfidence by it and, in that connection, utilizes a system to determine when and whether to hold certain types of information in ecnfidence. The ap-plication of that system and the substance of that systam constitutas Westinghouse policy and provides the rational basis required.

Under that systam, information is held in confidence if it falls in one or mor=. of several types, the release of which might result in the loss of an existing or potential ecm-petitive advantage, as follows:

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(a) The infor-nation reveals the distinguishing aspects of a

' process (or component, structure, teol, method, etc.)

where prevention of its use by any of Westinghouse's competitors without license frem Westinghouse constitutas a competitive economic advantage over other companies.

(b) It consists of supporting data, including test data,

, relative to a process (or ccmcenent, stmeture, tecl, method, etc.), the application of which data securts a competitive eccncmic advantage, e.g., by actimi:stien or improved marketability.

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(c) Its use by a con:petitor would reduce his expenditure of resources or improve his cecpetitive position in the design, manufacture, shipment, installation, assurance of quality, or licensfiig a similar product.

(d) It reveals c:st or price infor ation, production cap-acities, budge 1evels or c:rnercial strategies of Westinghouse. Its cust mers or suppliers.

(e) It reveals aspects of past, present, or future tiest-inghouse or customer funded development plans and pro-grams of potential cen=ercial valbe to Westinghet:se.

(f) It c:ntains patentable ideas, for which patent pro-taction may be desirable.

(g) It is not the property of Westinghouse, but must be treated as proprietary by Wastinghouse ac:crding to agreements with the owner.

There are sound policy reasens behind the Westinghouse system.which include the follcwing:

(a) The use of suca infor=ation by Westingn:use gives Westinghouse a cem;etitive advantage over its ccm-petitors. It is,' therefore, withheld frca cisclosure to protect the Westinghouse c:m;etitive positicn.

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5-AW-75 60 (b) It-is information which is marketable in many ways.

The extent to 'which such infomation is available to competitors diminishes the Westinghouse ability to sell products and services involving the use of the information.

(c) Use by our creatiter would put Westinghouse at a competitive disadvantage by reducing his expenditure of resources at our expense.

(d) Each component of proprietary information pertinent to a particular cec:patitive advantace is potantially as valuable as the total competitive advantage. If competitors acquire components of proprietary infor-antion, any one component may be the key to the entire puzzle thereby depriving Westinghouse of a competitive advantage.

(e) Unrestricted disclosure would jeopardi:e the position

' of prominence of Westinghcuse ini the world market, and thereby give a market advantage to the c:mpetition in those countries.

(f) The Westinghouse capacity to invest c:rporate assets in research and development decends upon the success in obtaining and maintaining a ccm:etitive advantage.

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. AW-76-60 (iii)

The infonnation is being transmitted to the Ccanission in j

confidence and, under the previsions of 10 CFR Section 2.750, it is to be received in confidence by the Cc=ission.

(iv)

T1fe information is not available in public sources to the best of our knowledge and bejief.

'(v)

The proprietary information sought to be withheld in this sub-sittal is that which is appropriately marked in the attach-

. ment to Westinghouse letter nunter NS-CE-1298, Eiche1dinger to a

Stolz, dated December 1,1975, ccncerning ini:rmation relating to NRC review of WCAP-C567-P and '. CAP-8563 entitled, " Improved Thennal Design Procedure," defining the sensitivity of ctg ratio to various core parameter:. The letter and attachttent are being submitted in respcase to the NRC request at the October 2S,1976 NRC/Westinghcuse : nesting.

This information enables Westinghouse to:

(a) Justify the Westinghouse design.

(b) Assist its c::st:mers to obtain if censes.

(c Meet warranties.

(d) Provide greater cperational flexibility to cust:: ars assuring them of safe and reliable cceration.

(e) Justify increased ;cwer capability or c;erating margin for plants wnile assuring safe and reliatie c;eratien.

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> AW-76-oO (f) Optimize reactor design and performance while maintaining a high level of fuel integrity.

Further, the information gained fec= the improved thenr.1 design procedure is of significant cc:=:ercial value as foklews:

(a) Westinghouse uses the infomatien to perform and justify analyses which are sold to customers.

(b) Westinghouse sells analysis servicas based upon the experience gained and the methods developed.

. Public disclosure of this ini'omation concerning design pro-cedures is likely to cause substantial harm to the competitive h

position of Westinghouse because competitors could utilize this information to assess and 3ustify their own designs without consnensurate expense.

The parametric analyses performed and their evaluatica represent a considerable amcunt of highly qualified develop =ent effort.

4 This work was contingent upon a design method development pro-gram which has been underaay during the past tro years.

A1,tegether, a substantial amcunt of meney and effort has been expended by Westinghouse which could only be duplicated by a competitor if he were to invest similar su=s of money and pro-vided he had the appropriate talent available.

Further the deponent sayeth not.

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ENCLOSURE 2 I

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-Question l':

The safety analysis for~ Point Beach references WCAP-9500-A which describes a 17x17. optimized fuel assembly (OFA). Provide justification for its application to the 14x14 0FA used in Point. Beach.

Response

The fuel design bases and criteria for Westinghouse 14x14 0FAs are the same as those discussed in Sections 4.2 and 4.4.1.2 of WCAP-9503 for the Westinghouse 17x17 0FA design. Verification that these criteria are met for Westinghouse fuel in the Point Beach Units is performed using the design methodology and models discussed in WCAP-9272, " Westinghouse Reload Safety Evaluation Methodology". These methods and core models used in the reload transition analysis are the same which have been used in the past Point Beach reload cycle designs. No changes in the nuclear design philosophy, methods or models are necessary due' to the transition to 0FA. Based on prototype hydraulic testing of the standard and the OFA assemblies, it was concluded that they are hydraulically compatible, and all of the current thermal and hydraulic design criteria are satisfied.

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1565L/081084 l-l l

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's Ouestion 2:

In our Safety Evaluation Report on WCAP-9500, " Reference Core Report 17x17 Optimized Fuel Assembly," the staff required that those plants using the Westinghouse Improved Thermal Design Procedure (ITDP) supply additional information on the plant specific application of the ITDP. Since the licensee is using the ITDP to perform their thermal-hydraulic analyses, we will require

.the following:

1)

Provide the sensitivity factors (5 ) and their range of 9

applicability; 2)

-If the S values used in the Point Beach analyses are different-9 from those used in WCAP-9500, then the applicant should re-evaluate the use of an uncertainty allowance for application of equation 3-2 of WCAP-8567, " Improved Thermal Design Procedure," and they should validate the linearity assumption; 3)

If there are any changes to the THINC-IV correlations, or parameter values outside of previously demonstrated acceptable ranges, the staff will require a re-evaluation of the sensitivity factors and/or the use of equation 3-2 of WCAP-8567.

4)

Provide and justify the variances and distributions for the input parameters; 5)

Justify that the nominal conditions used in the analyses bound all permitted modes of plant operation (including future operating cycles);

6)

Provide a discussion of what code uncertainties, including their values, are included in the DNBR analyses; and 7)

Provide a block diagram depicting sensors, processing equipment, computer and readout devices for each parameter channel used in the 1565L/081084 2-1 r

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uncertainty analysis. Within each element of the block diagram identify the accuracy, drift, range, span, operating limits, and setpoints. Identify the overall accuracy of each channel transmitter to final output and specify the minimum acceptable accuracy for use with the new procedure. Also identify the overall accuracy of the finaloutputvalueandmIximumaccuracyrequirementsforeachinput channel for this final output device.

Response

1)

The sensitivity factors and their range of applicability are given in Table 2-1 for 14x14 0FA fuel. These sensitivities have been determined using the WRB-1 DNB correlation.

2)

The S values used in the Point Beach 0FA analyses are different j

from those used in WCAP-9500 because the WCAP-9500 sensitivity values are not applicable to the 14x14 0FA fuel geometry. The uncertainty allowance calculation is shown in Tables 2-2 and 2-3 for typical and thimble cells, respectively.

3)

For the Point Beach units, the THINC IV code and the WRB-1 DNB correlation are the same as that used in WCAP-9500 for Westinghouse OFA fuel. All parameter values are within the ranges of codes and correlations used, and sensitivity factors have been cetermined specific to the fuel type over the range of Point Beach plant parameters.

(4), (5), and (7) - The responses to these questions are given in Appendix 1 and Attachments A and B to Appendix 1.

The uncertainties conservatively bound those associated with Point Beach instrumentation.

(6) Code uncertainty values that have been used in the CNB calculations are shown in Tables 2-2 and 2-3.

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1565L:6/081084

e ATTACHNENT TO QUESTION 2 APPENDIX 1 INSTRUENT UNCERTAINTIES FOR THE POINT BEACH ITDP CALCULATIONS

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Use of the Improved Themal Design Procedure (ITDP) requires the calculation of the standard deviation (a) for RCS pressure, temperature, power, end flow. For plants using Sostman Jor Rosemount RTDs, Westinghouse has determined on 'a generic basis, the ta1 certainty for the T p rtion of the Rod Control System and a av8 precision RCS flow calorimetric. In addition generic calculations have been made to detemine the t$ certainty for the Pressurizer Pressure control system and a daily power calorimetric. R ese generic calculations are outlined in Attachment A.

The calculations for Point, Beach were performed using input-supplied by the plant (Attachment B) and the methodology and equations supplied in Attachment A.

?lant specific variations from the generic calculations will be identified in each of the sections covering the four parameters. The first parameter to be discussed is Pressurizer Pressure.

Pressurizer Pressure he trcertainty in pressure is based on the accuracy of the Pressurizer Pressure control system. Using Equation 3 from Attachment A and the' individual uncertainties noted on Attachment B, it was detemined that the total uncertainty for the control system is [

]+a, c Allowing for the interaction of. the Pressurizer spray and heaters results in an uncertainty of

[

]+a,c Assuming a normal, two sided probability distribution results in-a =('

]+a,c which is the value used in the ITDP analysis.

kvg he uncertainty in T is detemined by looking at the T input to the Rod ava avg Control System. As noted in Attachment A an auctioneered value is compared with a reference as a function of power. We two inputs to the control' system are

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T,yg ({Tg + T }/2) and First Stage Turbine Impulse Chamber Pressure (the c

reference signai). Using Equation 3 of ~ Attachment A and the instrtrnent g

uncertainties.for Point Beach from Attachment B, it was determired that the accuracy of the control system is [

]+8' However, this does not

' include the mcertainty for the centrol system deadband. The deadband uncertainty is noted on page 7b of Attachment A and when combined with the control system tmcertainty, results in a total uncertainty of

[

I+8'

' Asstming a normal two sided probability distribution results in e

=[

]+a,c, the value used in the analysis.

Reactor Pm To determine the uncertainty in the daily power calorimetric is somewhat more ccmplicated than determining the uncertainties in pressure and T However avg.

section III.3.b of Attachment A lists the generic asstaptions and equations used by Westinghouse. A similar calculation was performed for Point Beach using Equation 2 of Attactrnent A and those tacertainties from Attachment B noted as

" Power Calorimetric". Using these and plant specific sensitivity values an equivalent of Table 2b of Attachment A can be constructed. For Point Beach the equivalent is as follows:

F

Commnent Instrtment Error UncertM nty

+a,c Feedwater Flow Venturi.

hermal Expansion Coefficient Temperature Material Density Temperature Pressure op Feedwater Enthalpy Temperature Pressure Steam Enthalpy Pressure Moisture Net Pu::p Heat Addition Dependent Paraneters

[

]+a,c

- Dependent Parameters

[

]+a,c As noted above some of the parameters are statistically dependent. In Attachment A this was ignored based on the conservatism of the assuned values.

However for Point Beach actual plant values were used, therefore the degree of conservatism was less. he Point Beach calculation was performed treating dependent parameters correctly, as noted above. Carrying through the calculations for a single loop, the uncertainty in power is [

]+a,c,

For a two loop plant the uncertainty is [

]+a,c The standard deviation for this parameter, as used in the ITDP calculations is a

=

[

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ai=

-.a a

._,4, 4A a

RCS Flow The tacertainty in RCS Flow is the combination of two mcertainties,1) a precision flow calorimetric (performed at the beginning of each cycle), and 2) the Cold Leg' Elbow Taps (which are normalized to the flow calorimetric). The first uncertainty to be discussed is the flow calorimetric.. A flow calorimetric is essentially a power calorimetric with the additional measurement of'T, T g

e and Pressurizer Pressure. Unlike the daily power calorimetric which assmed that the measurement values were from the plant process computer, the precision 1

flow measurement assmes that the most accurate means reasonably available is used. This implies the use of recently calibrated special test instrmentation and DVMs. It also asstanes that multiple measurements of each loop's parameters are made over an appreciable period of time-(typically once every five minutes over a one hour period). These two assumptions eliminate drift effects and 4

small parameter variations due to power or temperature oscillations. In j.

addition, the measurement should be. performed at the beginning of the cycle (or use an LEFN) to eliminate possible venturi fouling after startup.-

s The basic methodology and equations used are noted in Section III.4.b of Attachment A.

For the Point Beach specific calculations, Equation 1 of i

Attachment A and those uncertainties from Attachment B noted as " Flow Calorimetric" were used. In addition Point Beach has noted that multiple channels will be measured for a given parameter on a loop. An example of this

]

is Steamline Pressure. There are three channels for measuring Steamline Pressure on each of the two steamlines. All three of the channels will be measured and averaged to calculate the average steamline pressure for that loop for the period of the measurement. Those parameters for which multiple channels L

will be measured on each loop are:

Steamline Pressure - 3 channels / loop, T

- 2 channels / loop and H

T,

- 2 channels / loop.

In addition all four Pressurizer Pressure channels will be averaged, thus l'

requiring this tacertainty to be treated as a system error, not a loop error.

It should also be noted that Point Beach has installed a Leading Edge Flew Meter s

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(LEFM) in the common feedwater header. Use of the LEFM would then require treatment of the feedwater flow uncertainty as a systen error.

Based on use of the uncertainties of Attachment B and plant specific sensitivities, it is possible to construct a table equivalent to Table 3b of Attachment A.

For Point Beach the equivalent table asstaning use of the feedwater venturi's for flow measurenent is:

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Comtenant Instrtment Error Uncertainty

+a,c Feedwater Flow Venturi Ihermal Expansion Coefficient Temperature Material Density Temperature Pressure ap 4

Feedwater Enthalpy Temperature Pressure Stem Enthalpy Pressure Moisture Net Pm p Heat Addition Hot Leg Enthalpy Temperature Strem ing Pressure Cold Leg Enthalpy 3

Temperature Pressure

Dependent Par m eters [

]+"'"

6 Dependent Par m eters [

]+a,c O Dependent Farmeters [

]*"'

For a single loop the uncertainty (after treating dependent parameters in a rigorous manner) is [

]**'".

For two loops the system uncertainty is

[

]+a, In this instance the standard deviation (assuning a nonnal, two sided probability distribution) is [

]+a,c,

If the LEFM is used to measure feedwater flow (with a corresponding more accurate measurement of feedwater temperature) the secondary side uncertainties are revised to as follows:

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Comconent Instrunent Error Uncertainty

+a,c Feedwater Flow Feedwater Enthalpy Temperature Pressure Stean Enthalpy Pressure Moisture Net Punp Heat Addition For a single loop the uncertainty is [

]+*'".

For two loops the uncertainty is (af ter correct treatment of system uncertainties) [

]+a,c In this instance the standard deviation is [

]+a,c However, this is only half of the ficw measurement uncertainty. The second half is the uncertainty of the Cold Leg Elbow Tap. Assuning that only one elbow tap per loop is ready by the plant process computer, Attachment B notes the uncertainty for one loop. Combining uncertainties results in the follcwing total uncertainties for RCS Flow:

[

]+a,c assuming use of the feedwater venturi,

[

]+"'" assuning use of the LEFM for feedwater flow measurement. To be conservative, the ITDP calculations assuned the use of the feedwater venturi for the flow calorimetric. This

i would then result in a standard deviatica of (

]+8' In surunary, the following uncertainties and standard deviations were calculated for use in the ITDP analysis.

Uncertainty a

Pressurizer Pressure Control Rod Control (Temperature)

Power Calorimetric RCS Flow 1

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' Question 3:

Please identify the limiting DNBR transient and provide the results of the analysis of the event, including the calculated value of minimum DNBR.

Response

The limiting DN8 transient is Uncontrolled RCCA Withdrawal at Power. Results of this postulated event, including figures of the minimum calculated values of DNBR, are attached.

1565L/C31084 3-1

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. Question 3 Attachment 14.1.2 Uncontrolled RCCA' Withdrawal at Power 4

An uncontrolled RCCA withdrawal at power resuits in an increase in core heat

- flux. Since the heat extraction from the steam generator remains constant,

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there is.a net increase in reactor coolant temperature. Unless terminated by manual or automatic action, this power mismatch and resultant coolant temperature rise would eventually result in DNB. Therefore, to prevent the possibility of damage to-the cladding, the Reactor Protection' System is designed to terminate any such' transient with an adequate margin to DNB.

The automatic features of the Reactor Protection System which prevent core damage 'in a rod withdrawal accident at power include the following:

1.

Nuclear power range instrumentation actuates a reactor trip if two out of the four channels exceed an overpower setpoint.

2.

Reactor trip is actuated if any two out of four AT channels exceed an overtemperature AT setpoint. This-setpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNB.

3.

Reactor trip is actuated if any two out of four AT channels exceed j.

j an overpower AT setpoint. This setpoint is automatically varied with axial power imbalance and coolant temperature to ensure that the j

allowable heat generation rate (kw/ft) is not exceeded.

4.

A high pressure reactor trip, actuated from any two out of three pressure channels, is set at a fixed point. This set pressure will be less than the set pressure for the pressurizer safety valves.

3-2 15GSL:G/081084

B 5.

A high pressurizer water level. reactor trip, actuated from any two out of three level-channels, is actuated at a fixed setpoint. This affords additional protection for RCCA withdrawal accidents.

The manner in which the combination of overpower and overtemperature AT trips provides protection over the full range of reactor coolant system

- conditions is illustrated in Figure 14-1.

Figure 14-1 presents allowable reactor loop average temperature and AT for the design power distribution and flow as a function of primary coolant pressure. The boundaries of operation defined by the overpower AT trip and the overtemperature AT trip are represented as " protection lines" on 'this diagram. These protection lines are drawn to include all adverse instrumentation and setpoint errors, so that under nominal conditions trip would occur well within the ' area bounded by.

these lines. A maximum steady state operating condition for the reactor is

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also shown on the Figure.

The utility of the diagram just described is in the fact that the operating limit imposed by any given DNB ratio can-be represented as a line on this coordinate system. The ONB lines represent the locus of conditions for which the DNBR equals the limit value (1.65 for the thimble cell and 1.66 for the typical cell). All points below and to the left of this line have a ONB ratio greater than this value. The diagram shows that DNB is prevented for all

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cases if the area enclosed within the maximum protection lines is not traversed by the applicable ONB ratio line at any point.

The region of permissible operation (power, pressure, and temperature) is completely bounded by the combination of reactor trips: nuclear overpower (fixed setpoint); high pressure (fixed setpoint); low pressure 3-3 1565L:6/081084

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FIGURE 14-1 4

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. (fixed setpoint); overpower and overtemperature AT (variable setpoints).

These trips are designed to prevent overpower and a ONB ratio of less than the limit value.

Method of Analysis Uncontrolled rod cluster control assembly bank withdrawal is analyzed by the LOFTRAN code. This code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables, including temperatures, pressures, and power level. The core limits, as illustrated in Figure 14-1, are used as input to LOFTRAN to determine the minimum departure from nucleate boiling ratio during the transient. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567.

In order to obtain conservative values of departure from nucleate boiling ratio, the following assumptions are made:

1.

Initial Conditions - Initial reactor power, reactor coolant average temperatures, and reduced reactor coolant pressure (2000 psia) are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAp-8567.

2.

Reactivity Coefficients - Two cases are analyzed.

3-5 1565L/081484

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a.

Minimum Reactivity Feedback - A positive (5 pcm/ F) moderator coefficient of reactivity is assumed,-corresponding to the beginning of. core life. A variable Doppler power coefficient with

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core power is used in the analysis. A conservatively small (in absolute magnitude) value is assumed.

b.

Maximum Reactivity Feedback - A conservatively large positive moderator density coefficient and a large (in absolute magnitude)

-negative Doppler power coefficient are assumed.

3.

The rod cluster control assembly trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position.

4.

The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118". of cominal full power. The overtemperature AT trip includes all adverse instrumentation and setpoint errors; the delays for trip actuation are assumed to be the maximum values.

No credit was taken for the other expected trip functions.

l 5.- The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the combination of the two control banks having the maximum combined worth at maximum speed.

l The effect of rod cluster control assembly movement on the axial core power distribution is accounted for by causing a decrease in overtemperature and overpower AT trip setpoints proportional to a decrease in margin to ONB.

Results Figures 14.1.2-1 and 14.1.2-2 show the response of neutron flux, pressure, average coolant temperature, and departure from nucleate boiling i

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l 3-6 l

1555L:6/081084

t ratio to a rapid rod cluster control assembly withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the.

start of the accident. Since this is rapid with respect to the thermal time constants of the plant, small changes in T,yg and pressure result, and a large margin to DNB is maintained.

The response of neutron flux, pressure, average coolant temperature, and DNBR for a slow control rod assembly withdrawal from 10% power is shown in Figures.

14.1.2-3 and'14.1.2-4.

Reactor trip on overtemperature AT occurs after a longer period, and the rise in temperature and. pressure is consequently larger than for rapid rod cluster control assembly. withdrawal. Again, the minimum DNBR is greater than the limit value.

Figure 14.1.2-5 shows the minimum departure from nucleate boiling ratio as a-function of reactivity insertion rate'from initial full power operation for the minimum and maximum reactivity feedback cases.

It can be seen that two reactor trip channels provide protection over.the whole range of reactivity insertion rates. These are the high neutron flux and overtemperature AT trip channels. The minimum ONBR is never less than the limit value.

l Figures 14.L2-6 and 14.1.2-7 show the minimum departure from nucleate boiling ratio as a function of reactivity insertion rate for rod cluster control assembly withdrawal incidents starting at 60% and 10% power respectively. The results are similar to the 100% power case, except that as the initial power is decreased, the range over which the overtemperature AT trip is effective is increased. In neither case does the departure from nucleate boiling ratio fall below the DNBR limit value.

In the referenced figures, the shape of the curves of minimum departure from nucleate boiling ratio versus reactivity insertion rate is due both to reactor core and coolant system transient response and to protection system action in initiating a reactor trip.

3-7 156SL:6/G81084

Referring to Figure -14.1.2-6, for example, it is noted that:

1. ' For' high reactivity insertion rates (i.e., between ~100 pcm/second and ~5 pcm/second), reactor trip is initiated by the

.high neutron. flux trip.for the minimum reactivity feedback cases.

The neutron flux level in the core rises rapidly for these insertion rates, while core heat fiux and coolant system temperature. lag behind due to the. thermal capacity of the fuel and coolant system fluid. Thus,.the reactor is tripped prior to significant increase in~ heat flux or water temperature with resultant high minimum departure from nucleate boiling ratios during the transient. Within this range, as the reactivity insertion rate decreases, core heat

. flux'and coolant temperatures can remain more nearly in equilibrium with the neutron flux; minimum DNBR during the transient thus decreases with decreasing insertion rate.

2.

With further decrease in reactivity insertion rate, the overtempera-ture AT and high neutron flux trips become equally effective in terminating the transient (e.g., at ~4 pcm/second reactivity insertion rate).

The overtemperature AT reactor trip circuit initiates a reactor trip when measured coolant loop AT exceeds a setpoint based on I

measured reactor coolant system average temperature and pressure.

It is important in this context to note, however, that the average temperature contribution to the circuit is lead-lag compensated in order to decrease the effect of the thermal capacity of the reactor coolant system in response to power increases.

1 For reactivity insertion ra' es between ~4 pcm/second and t

~.6 pcm/second, the effectiveness of the overtemperature AT trip increases (in terms of increased minimum departure from nucleate

'i 3-8 1565L:6/081084

-. a

boiling ratio) due to the fact that, with lower insertion rates, the power increase rate is slower, the rate of rise of average coolant temperature is slower, and the lead-lag compensation provided can increasingly account for the coolant system thermai capacity lag.

3.

For maximum reactivity feedback cases reactivity insertion rates -less than ~60 pcm/second, the rise in reactor coolant temperature is sufficiently_ high so that the steam generator safety valve setpoint is reached prior to trip. Opening these valves,.which act:as an additional heat load on the reactor coolant system, sharply decreases the rate of rise of reactor coolant system average temperature. This decrease in rate of rise of the average coolant system temperature during the transient is accentuated by the lead-lag compensation, causing the overtemperature'AT trip setpoint to be reached later with resulting lower minimum departure from nucleate boiling ratios.

Figures 14.1.2-5, 14.1.2-6, and 14.1.2-7 illustrate minimum departure from nucleate boiling ratio calculated for minimum and maximum reactivity feedback. The calculated sequence of events for this accident is shown in Table 14.1.2-1.

Conclusions In the unlikely event of an at power (either from full power or lower power levels) control rod bank withdrawal incident, the core and reactor coolant system are not adversely affected since the minimum value of DNB ratio reached is in excess of the ONB limit value for all rod reactivity rates. Protection is provided by nuclear flux overpower and overtemperature AT.

Additional protection would be provided by the high pressurizer level, overpower AT, and the high pressure reactor trip. The preceding sections have described the effectiveness of these protection channels.

1565L:6/081084 a

TABLE 14.1.2-1 TIME SEQUENCE OF EVENTS FOR UNCONTROLLED RCCA WITHDRAWAL AT POWER Time of Each Event Event (Seconds)

Case A:

Initiation of uncontrolled rod cluster 0

control assembly withdrawal at full power and maximum reactivity insertion rate (100 pcm/sec)

Power range high neutron flux high trip point reached 1.061 Rods begin to fall into core 1.561 Minimum departure from nucleate boiling ratio occurs 2.5 Case B:

Initiation of uncontrolled rod cluster control 0

assembly withdrawal at 107. power and at a small reactivity insertion rate (4 pcm/sec)

Overtemperature AT reactor trip signal initiated 125.80 Rods begin to fall into core 127.80 Minimum departure from nucleate boiling ratio occurs 128.2 3-10 1565L:6/031084

4000 3

5 1.2500 - -

5 2

8.1.000J -

i::y

.75000 - -

E Er

=

.50000 - -

5 e

L

=

.25000 - -

Ud

.a 2-30 0

2600.0 2500.0 - -

~

g v.

S 2400.0 - -

y 2300.0 - -

S W

2200.0 - -

=

5 2100.0 -

~

=

S 2000.0 C

E 1900.0 -

1800.0 l

0 2.5 5.0 7.5 10.0 12.5 15.0' TIME (SECONOS)

FIGURE 14.1.2-1 R00 WITH0RAWAL AT POWER, FULL POWER, MINIMUM FEEDBACX,100 pen /sec WITH0RAWAL RATE NUCLEAR POWER & PRESSURIZER PRESSURE VS. TIME l

1 3-11

620.00 9

3 610.00 - -

e 600.00 - -

m 37 590.00 - -

E I

580.00 -

O 570.00 - -

E 5

560.00 -

y 550.00 - -

v 540.00 4.0000 f

3.5000 - -

5.0000 - -

=

i 2.5000 -

C 2.0000 - -

1.5000 - -

1.0000 0

2.5 5.0 7.5 10.0 12.5 15.0 1

TIME (SECONOS)

FIGURE 14.1.2-2 R00 WITH0RAWAL AT POWER, FL'LL POWER, MINIMUM FEECBACK,100 pcm/sec WITH0RAWAL RATE CORE AVERAGE TEMPERATURE & ONBR VS. TIME 3-12 i

9 1.4000 0

5 1.2500 - -

z S

8 1.0000 - -

_i Q

. 75000 - -

5

~

=

.50000 - -

5 2

=.<.25000 - -

i.

0.0 2600.0 2500.0 - -

2 2400.0 -

I y

2300.0 - -

N W

2200.0 -

l 2100.0 -

6

~

ne S

2000.0 m

W E

1900.0 -

1800.0 0

25 50 75 100 125 150 175 200 TIME (SECON05)

FIGURE 14.1. 2-3 R00 WITH0RAWAL AT POWER,10% PCWER, MINIMUM FEEO8ACX, a pcn/sec WITHDRAWAL RATE NUCLEAR POWER & PRES $URIZER PRESSURE VS. TIME 3-13 i

y

_m,- - - ~ - - - -

u--

---,,~_-m,

,---w

-cy.-

620.00 C.

610.00 - -

8 600.00 - -

m 7

590.00 -

I 5

I 580.00 - -

O 570.00 - -

E 5

560.00 - -

y 550.00 v

540.00,

5.0000 4.5000 - -

4.0000 - -

3.5000 - -

a9 3.0000 - -

2.5000'- -

2.0000 - -

1.5000 - -

1.0000 O

25 50 75 100 125 150 175 200 TIME (SECONOS)

FIGURE 14.1.2-4 A00 WITH0RAWAL AT POWER,10% POWER, MINIMUM FEECBACX, 4 pen /sec WITH0RAWAL RATE CORE AVERAGE TEMPERATURE & ONBR VS. TIME 3-14 l

4 2.2 1

Minimum feedback Maximum feedback

2. 0 _

i 3

e j

4' e"

j g

1. 8 _

f E

-~m/

nr a

1 1.6

.1

.2

.4

.6

.8 1.0 2.0 4.0 6.0 8.0 10 20 40 60 80 100 j

a REACTIVITY INSERTION RATE (PEH/SEC) j FIGURE 14.1.2-5 i

R00 WITH0RAWAL AT POWER FULL POWER i

HINIHUM DNBR VS REACTIVITY INSERTION RATE i

(

4 I

i 1

O g

O l

i e

b i

e i

4 a

w i

p#

..C p

C 4,%

g N

W N

..C 2

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L<W I

>=

=2

/

..C w

vC v

w N

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w "-

j s

5

/

m Cs C

e

- I

C g 4

-W y*

g I

g g"

M M l

u v

. G3

=

w

.a: as:

.. C.

5_

, g" 88 m

w

..=w

..g g

k. m m W g

w w w

(

KK

..F W

C M

x a>

- w gg Rg C -

= -

K 5

h*

>=

U 3

C 2."

m

=K m

'.L 4

.se L

.. N.

I I

9 l

8 W

i m.

A.

0%.

D.

N N

N W

H8NG WOWINIW i

l 3-16 nn-.

e--

9 9

0 0

/

1

/

0

8

0 6

0 4

R EE WT OA

)

PR C

0 E

TN 2

S N O

/

E I

M C T C

R R

~

P E E S

0

(

P q

1 7

N K K E

0I C C

0 T

2 1

Y A A A

T B 8

8 R

1 D D R

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4 E

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1 W T F F

6 I

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P H

0 R

R A

l l

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T t

R M M 4

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l i

T TM C

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W M E

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RM 6

4 2

]

7 5

3 1

9 5

7 2

2 2

2 1

1 1

%soEEEx e

Ysu iiii l!,!

}

,i;

Question 4:

Please provide a table of kinetics parameter ranges for the standard core, mixed cores and 0FA core.

Response

Table 4-1 provides the required kinetic parameters.

4-1 1565L:6/081084

f l-

.g.

I

, r

~

y Y

c 1ASLL 4-1 e

l LIMliS USED IN THE IMANSIENT ANALYSIS TyriCAL CYCLE' j<.

-" ~

d' 4

. P rdocLgf,

Statula rd Fe.dl Core QF A Co re*

Jyoical Valise L imi t t rd f t:mi lygg.. _

f,..

-9

' L' : F.;g#,

Nht,husitive MIC (pca/*F) 0.0 o 's. u

' +3.0

~ Of A Hust' Ncgative MIC (pce/*f)

-35.0

-31.0 SID ~

f:

Nst skigative DIC (pca/*F)

-1.6

-2.9

-2.4 SID 4

l e:ast-seegacita DIC (pcm/*f)

-1.0

-0.91

-l.4 OfA

-13.1 SID Wa t beg a t i ve DPC ( pcm/* f ) -

-8.0 of A -

c s aast Negative DPC (pce/% power)

-13.3 Of A Mamanue 80ross Wrtta (pce/ ppm) s 4

-6. 5 -

SID

. M,nimase tsorose Wrth (pca/ppe) 7 Oes700

.0.0072

,0.00584 -

. OfA M.swimase Seta off 0.

Manimien'> ta est 0.00453.

0.0043 0.amM15 51D

-6 e

4 Mamimus Fromp houtrca Lifetime (10 6cc) it

/

26

[ 21 '

j'

- OtA

e. j O

J

.ll m.

is.

.y, '

y a

4 M

(*

2*

ig v.

a ltou limats for a transition caso arv ttae sames as for an all OfA coru.

e 1%uL:6/uS4044 L

Question 5:

Provide more information on the results of the reanalysis for the Uncontrolled Rod Withdrawal at Power. In the updated FSAR this event is apparently the Ilmiting DNBR transient. Provide quantitative discussion of the effect of 0FA and positive MTC on this event.

Response

The safety evaluation for Pt. Beach 0FA transition core has demonstrated that the ON8 design basis has been met for this transient for both optimized and standard fuel consistent with the. Thermal / Hydraulic methodology used to evaluate each fuel type as specified in the Safety Evaluation for Point Beach Units 1 and 2 Transition to Westinghouse 14x14 Optimized Fuel Assemblies (Reference Attachment B of September 6,1983 letter from Fay (WEPCO) to Denton (NRC)]. The moderator temperature coefficient has been assumed at its most conservative value within the bounds of Technical Specification 15.3.1 regarding MTC.

\\

1540L:6/081084 5-1, Li

y s-1 l

-Question 6:

Provide a qualitative ' discussion of the effect of the various changes (OFA, positivEMTC,RAOC,'etc.)onallothereventsthatwerereanalyzed.

'q Response:

General discussions of the impacts due to 0FA, positive MTC, F

. AH multiplier, RAOC and other changes are discussed in Section 6.1 of Attachment

' B to letter of Septem6er 6,1983 from Fay (WEPCO) to Denton (NRC);

discussions :pecific to the variousi ransients are included in Section 6.2 of t

that document.

i 4

\\

1

<i s

1540L:6/081084 6-1

,.--.....I

- ~ - -

Question 7:

. Provide a discussion of the Primary System Pipe Rupture (small break LOCA)

- event.

Response

The small-break LOCA analysis for Point Beach applicable to transition and full 0FA core. cycles was reanalyzed dua to the differences between-Westinghouse standard and 0FA designs. The currently approved October 1975 small break ECCS evaluation model was utilized for a spectrum of cold-leg breaks.

When assessing the transition core impact on small break LOCA, the only mechanism available-to cause a transition core to have a greater calculated PCT than a full core of either fuel is the possibility of flow redistribution due to fuel assembly hydraulic resistance mismatch.

4 -

The W-FLASH computer code was used to model the core hydraulics during a small break LOCA event. Only one core flow channel was modeled in W-FLASH, since the core flowrate during a small break LOCA is relatively low, and this provides enough time to maintain flow equilibrium between fuel assemblies 4

(i.e., crossflow). Therefore, hydraulic resistance mismatch is not a factor for small break LOCA. Thus it was not necessary to perform a small break 1

evaluation for transition cores and it was sufficient to reference the small break LOCA for the' full core of the OFA design.

The small break 0FA LOCA analysis for Point Beach utilizing the currently approved 1975 Small Break Evaluation model resulted in a PCT of 992*F for the 6-inch' diameter cold-leg break. The analysis assumed the worst small break power. shape consistent with a LOCA F envelope of 2.32 at core midplane g

elevation and 1.5 at the top of the core.

Analyses showed that the high'and low head portions of the ECCS, together with the. accumulators, provided sufficient core flooding to keep the calculated PCT 1

7-1

. 1565L:6/081084

~

well below the recuired limits of 10 CFR 50.46. Adequate protection is therefore afforded by the ECCS in the event of a small break LOCA in the Point Beach Units.

7-2 1565L:6/081084

Question 8:

c Westinghouse has recognized that, under certain circumstances, the FSAR analysis of the cropped rod event for turbine runback plants may not be complete. Specifically dropping a very low worth rod would lead to a turbine runback to 78 percent power but would not reduce core power by this amount.

Thus a core turbine mismatch would be created with possible violation of DNBR limits. Please confirm that an analysis of this scenario has been performed, describe the analysis procedure, and provide the results for the limiting core configuration.

- Response:

A reanalysis of the dropped rod event for Point Beach has been performed to address this issue. A description, including methods, assumptions, results and conclusions is attached. The DN8 design basis has been confirmed to be met for this event.

l 8-1 1565L:6/081084 I

Question 8 Attachment Rod Cluster Control Assembly (RCCA) Drop Dropping of a full-length RCCA occurs when the drive mechanism is deenergized. This would cause a power reduction and an increase in the hot channel factor. If no protective ac'tfon occurred, the Reactor Control System would restore the power to the level which existed before the incident. This would lead to a reduced safety margin or possibly DNB, depending upon the magnitude of the resultant hot channel factor.

If an RCCA drops into the core during power operation, it would be detected by-either a rod bottom signal, by an out-of-core chamber, or both. The rod bottom signal device provides an indication signal for each RCCA. The other independent indication of a dropped RCCA is obtained by using the out-of-core power range channel signals. This rod drop detection circuit is actuated upon sensing a rapid decrease in local flux and is designed such that normal load variations do not cause it to be actuated.

A rod drop signal from any rod position indication channel, or from one or more of the four power range channels, initiates the following protective action: reduction of the turbine load by a preset adjustable amount and blocking of further automatic rod withdrawal. The turbine runback is achieved by acting upon the turbine load limit and/or on the turbine load reference.

The rod withdrawal block is redundantly achieved.

Method of Analysis The transient following a dropped RCCA accident is determined by a detailed digital simulation of the plant. The dropped rod causes a step decrease in reactivity and the core power generation is determined using the LOFTRAN code. The overall response is calculated by simulating the turbine load runback and preventing rod withdrawal. The analysis is pre:ented for the case in which the load cutback is greater than that required to match the worth of the dropped rod (75pcm). The load is assumed to be cut back from 100 to 76%

of full load at a conservatively slow rate of approximately 1% per second.

8-2 1565L/081484

...s e

a t

i

~

~The least negative values of moderator and Doppler-temperature coefficients of.

reactivity are used in this analysis resulting lin the: highest heat flux during

~

the transient. These are moderator density coefficient of reactivity of 0.0 ap/gm/cc and a Doppler temperature coefficient of reactivity of

-1 pcm/*F.'

This accident'is analyzed with the Improved Thermal Des'ign Procedure as described in WCAP-8567.

Results l

Figures 14.1.3-4 through 14.1.3-6 illustrate the transient response following a dropped rod of worth 75 pcm. The reactor coolant average temperature decreases initially, due to the decrease in reactor core power. Since the drop in power is less than the drop in load, with no reactivity feedback, coolant temperature then increases. The higher' primary power level (as opposed to secondary) is eventually matched by opening of the steam generator safety valves. Steady-state conditions are then achieved.

a Y

d I

4 i

5 8-3 t'

..... -....~. -

....- - ~

1.2000 1.1000 - -

-1 E

1.0000 -

w C

I

-g

.90000 -

=

.w

.80000 - -

a.

2

".i

.70000 -

M=

.80000 - -

.50000

!.2000 l

1. 1000 -

1o 2

1.0000 - -

L e

E

.90000 -

u

d

.80000 -L r<w=

w

.70000 - -

=cw

.60000

-L

.50000 L

l 0

50 100 150 200 250 300 TDE (SEC::NOS)

FIGURE 4.1.M 1

l i

RE!FONSE M A OROPFED AC":A 0F WORTH - 75 ::c:n 8-4 l

NUC:.!AR 90VER i. CORE HEAT FLUX '15. TU1E

700.00

~

675.00 - -

c 650.00 - -

a wc

~

625.00 - -

c.Z

\\

w

~

600.00 --

z<

a 8

575.00 -

~

v o

550.00 -

525.00 -

500.00 2400.0 2300.0 -

w 5

2200.0 - -

M w

CE C.

5 2100.0 -

~

3 m

W W

g 2000.0 -

1900.0 -

1800.0 i

0 50 100 150 200 250 300 i

l l

TIME (SEC'.,N05)

FIGURE 14.1.3-5 REIFONSE TO A CROPPED AC'.A 0F WGRT*ri - 75 ::c:t 8-5 PRE 33URIIIR FREISLRE & AVG C'0LANT 79.P. 'IS. TIME

600.00 580.00 - -

C ew 560.00 - -

w=3 E

w Q.

I 540.00 - -

w

>=

w 5-520.00 - -

500.00 1.0 s

s 0.9

-a<z 0.8 c5

<=

C~5 5,.

0.7

~

m-v as:

Et 0.6 O.5 0

50 100 150 200 250 300 1 JE (5It'.N05)

FIG'RE 14.t.2-4 4

AEIFONSE TO A CROFeED RC'A 0F '4Rin - e : :c-.

8-6 int.ET TEMP.1 STE;M LCAO 't!. TIP.E i

f v_ i d -

k Question 9:

Technical Specification 15.3.10 is not complete. The measu' red Fg value must be multiplied by a. factor which accounts for xenon transient effects before it.is'ccmpared to the F (z) limit..This Specification n

should be altered to be consistent with the sample Specification given in Appendix A to Part B of NS-EPR-2649, "The F Surveillance Technical q

Specification." The W(z) report required by that Specification should be. submitted in a timely manner.

Response

The revision to Technical Specification 15.3.10'is complete without the W(z) factor. Although the Westinghouse submittal (WCAP-10216-P-A) to the NRC deals with both RAOC and F Surveillance Technical g

~ Specifications, the implementation of RAOC does not require the use of F Surveillance. The revised Point Beach Technical Specification g

(15.3.10) is consistent with the present Point Beach specification which requires a measurement of F every 30 days. Adding the W(z) factor g

results in an F surveillance Technical Specification which is not g

related to RAOC operation.

1565L:6/081084 9-I

.i

A y

i ATTACHENT A - GENERIC CALCULATIONS (T0 APPENDIX 1) 4)

Provide and justify the variances and distributions for input parameters.

5)

Justify that the nominal conditions used in the analyses bound all permitted modes of plant operation.

7)

Provide a block diagram depicting sNsor, processing equipment, computer, and readout devices for each paraneter channel used in the uncertainty analysis. Within each element of the block dia-gram identify ths accuracy, drift, range, span, operating limits, and setpoints. Identify the overall accuracy of each channel transmitter to final output and specify the minimum acceptable accuracy for use with the new proceddre. Also identify the over-all accuracy of the final output value and maximum accuracy requimments fo: each input channel for this final output device.

Resoonse :

Rosemount RTDs I.

INTRODUCTION Four operating parameter uncertainties art used irr the uncertainty ana-lysis of the Imoroved Thermal Desigrr Procedure (ITDP). These operating parameters are pressurizer pressure. primary coolant temperature (T,yg), reactor power, and reactor coolant systent flow. These para-notars are monitored on a-regular basis and several are used for control purposes. The reactor power is monitored by the perfomance of a secon-dary side heat balance (power calorimetric measurtment) at least once every 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months . The RCS flow ~ir monitored by the perfomance of a pre-cision flow calorimetric measuremen't at the beginning of each cycle.

The RCS 1cco elbow tags, can then be nomalized against :ne precisi,on calorimetMe and used for eenthly survettance (wit.t's small increase ia total uncertainty) or a.precisforr flow calorimet-ic can be perfomec on 1b

l the same surveillance schedule. Pressurizer pressure is a controlled parameter and the uncertainty for the Improved Thermal Design Procedure reflects the use of the control system. T is a controlled para-ayg

. meter thmugh the use of the temperature input to the Control Rod con-trol system; the uncertainty presented here reflects the use of this control system.

Since 1978 Westinghouse has been deeply involyed with.the development of several techniques to treat instrumentation uncertainties, errors, and allowances. The earlier versions of these techniques have been docu-mented for several plants; one approach uses the methodology outlined in WCAP-8567 " Improved Themal Design Procedure.0,2,0 which,is based on the conservative assumption that the uncertainties can be described with unifom probability distributions. The other approach is based on the

~

more realistic assumption that the uncertainties can be described with normal probability distributions. This asstaption is also conservative in that the " tails" of the normal distribution are in reality " chopped" at the extremes of the range, i.e., the ranges.,r uncertainties are

~

finite and thus, allowing for some probability in excess of the range limits is a conservative assumption. This approach has been used to substantiate the acceptability of the protection system setpoints for several plants with a Westinghouse NSSS, e.g., D. C. Cook II

, No rth Anna Unit 1, Salem Unit 2, Sequoyah Unit 1, Y. C. Succer, and McGuire Unit 1.

Westinghouse now believes that the latter approach can be used for the deteminatiorr of the instrumentation errors and allowances for the ITDP parameters. The total instrumentation erttrs presented in this response are based on this appmach.

II.

METHODOLOGY

.~,e methodology used to combine the error components for a channel is basically the appropriate statistical cocDination of those groups of components which are statistically independent, i.e., not interactive.

1 Those errors which are not independent are combined arithmetically to fem independent groups, which can then be systematically coccined. The

. statistical combination technique used by Westingnouse is the C.

25

~

]+a,c.e of the instrumentation uncer-tainties. The instrumentation uncertainties are two sided distribu-tions. The sum of both sides is equal to the range for that parameter, e.g., Rack Drift is typically [

3+a,c, the range for this parameter is [

]+a,c.

This technique has been utilized before as noted 'above and has been endorsed by the staff (5,6,0 and various I'

industry standards The relationship between the error components and the statistical instrumentation error allowance for a ch' nnel is defined as follows:

a 1.

For parameter indication in the racks using a DVM;

- a,c

~

Eq. 1 2.

For parameter indication utilizing the plant process computer;

+. a, c Eq. 2 3.

For parameters which have control systems;

- a.c

~

Eq. 3 where:

CSA Channel Statistical Allowance

=

PMA Process Measurement Accuracy

=

PEA Primary Element Accuracy

=

SCA Sensor Calibration Accuracy

=

50 Sensor Drift

=

Ob

l i

I l

l STE Sensor Temperature Effects

=

SPE Sensor Pressure Effects

=

l RCA Rack. Calibration Accuracy

=

RD Rack Drift

=

RTE Rack Temperature Effects

=

DYM Digital Voltmeter Accuracy

=

ID Computer Isolator Drift

=

'A/D Analog to Digital Conversion Accuracy

=

CA Controller Accuracy

=

The parameters above are as defined in reference 4 and a're based on SAMA standard PMC-20-1973(10). However, for ease in understanding they are paraphrased below:

~

PMA non-instrument related measurement errors, e.g., tempera-ture stratification of a fluid in a pipe, PEA errors due to metering devices, e.g., elbows, venturis,

orifices, SCA reference (calibration) accuracy for a sensor / transmitter, SD change,in input-output relationship over a period of time at reference conditions for a sensor / transmitter, 1

STE change in input-output relationship due to a change in ambient temperature for a sensor / transmitter, SPE change in input-output relationship due to a change in static pressure for a. a.s cell,.

reference (calibration) accuracy for all rack modules in RCA 1,oop or channel assuming the loop or channel is tuned to this accuracy. This assumption eliminates any bias that could be set up through calibration of individual mocules in the loop or channel.

RD change in input-output relationship over a period of time at reference conditionc for the rack modules, RTE -

change in input-output relationship due to a change in ambient temperature for the rack modules, DVM the measurement accuracy of a digital voltmeter or multi-meter on it's most accurate applicable range for the parameter measured, ft

s i

ID change in input-output relationship over a period of time at reference conditions for a control / protection signal isolating device, allowance for conversion accuracy of an analog signal to A/D a digital signal for process computer use, allowance for the accuracy of a controller, not including CA deadband.

A more detailed explanation of the Westingh'ousa met' odology noting the h

interaction of several parameters is provided in reference 4.

III. Instrumentation Uncertainties The instrumentation uncertainties will be discussed first for the two parameters which are controlled by automatic systems, Pressurizer pres-sure. and T,yg (thro' ugh Rod Control).

The uncertainties for both of these parameters are listed on Table Ib, Typical Instrumentation Uncer-tainties.

~

1.b.

Pressurizer Pressure Pressurizer. pressure is controlled by a system that compares the mea-sured pressure against a reference value. The pressure is measured by a pressure cell connected to the vapor space of the pressurizer. All ow-ances are made as indicated on Table Ib for the sensor / transmitter and the process racks / controller. As noted, the CSA for this function is

[

]+a,c which corresponds to a control accuracy of [

]+a,c,

^

~

The accuracy assumed in the ITDP analysis is [

]+"' C, thus,

margin exists between analysis and the plant. Being a controlled para-meter, the nominal value of 2235 psig is reasonable and bounded by ITDP error analysis assumptions, i.e., assuming a normal, two sided distribu-tien for CSA and a 95+t probability distribution (which will be docu-mented later in this response), e for the noted CSA ecuals

[

]+a,c.

Assuming a normal, two sided distribution for the ITDP assumotion of [

]+a,c and a 95+t probability distribution results in a = = [

]+a,c.

Thus, Sb

e O

I I

ss e

U-a b

o E

I 8

-.a 8 -

o *

.b 15 8

3-E.

5.

g g

b I a

2. %.. :

l$

. -.i. t a

e t

2

.t.t 3 0

.. =.

- g i.e E

a E

w de Id b2

= 2 8

I 2

5

.S u

.= 0

==

m.--

Ir a-

=

e-

~

g

.a

2. t =.

4*2 I

=

: 2..

G.D

2. ). :.2 IT 4 )1 g 8

w ---

It

-g

O. 2

. 6

- =

5:

A S

8 w a v

~

.m.

=

g l

-3

.s J-

.t. :

I e.

2 3 *. ". 3 e

o a.

eee.

Ot

,I 3

. sae a*-

5*

e 8

7,.,*".",.

....e OeCaO A.. ~E3g 6

=

s.

l 33 3

-02223

: 3.:) j l

-~

w.

-~~--

=

g*<a 2=E o

.eS

= S *

= e>5e

~

margin exists between the expected and assumed standard deviations for Pressurizer pressure.,

2.b.

T AVG T

is controlled by a system that compares the auctioneered high ayg T

from. the loops with a reference derived from the First Stage ayg Turbine Impulse Pressure. Tayg is derived from the average of the narrow range TH and TC from the bypass manifolds. The highest loop T

is then used in the controller. Allowances are mace as noted on ayg Table b for the sensor / transmitter and the process racks / controller. As noted, the CSA for this function is [

]+a,c.which corre-spends to an instrumentation accuracy of [

]+a,c.

Assuming a normal, two sided distribution for CSA and a 95+% probability distribu-tion results in a standard oeviation, c = [

]+3*C, However, this does not include the controller deadband of + 1.5'F.

To determine the controller accuracy the instrumentation accuracy must be combined with the deadband. Westinghouse has determined that the proba-bility distribution for the c;eadband is [

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

[

3+aic and the standard deviation, e = [

]+a,c,

~

Combining statistically the stancard deviations for instrumentation anc deadband results in a controller standarc deviation of:

=fc)2+c b

j*

o 2

7b

+

Therefore, the controller uncertainty for ~a 95+% normal probability distribution is % [

] +a,c This is the uncertainty assumed for the ITDP error analysis and reasonably bounds the ncminal value corresponding to the full power T,yg.

3.b.

Reactor Power Generally a plant performs a primary / secondary side heat balance once every 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months when power is above 15% Rated Thermal Power. This heat balance is used to verify that the' plant is operating within the limits of the Operating License and to adjust the-Power Range Neutron Flux channels when the difference between the NIS and the heat balance is greater than that allowed by the plant Technical Specifications.

Assuming that the primary and secondary sides are in equilibrium; the core power is detemined by suming the themal output of the steam gene'rators, correcting the total secondary power for steam generator blowdown (if not secured), subtracting the RCP heat addition, accing the primary side system losses, and dividing by the core ratec Btu /hr at full power. The equation for this calculation is:

f

[N

)

RP =

E[03g /0 ) +0L Eq. 4 100 3

(

/

n where; RP Core power ( % RTP)

=

N

=

Number of primary side loops 03g Steam Generator thermal output (Stu/nr)-

=

Op RCP heat adder (Btu /hr)

=

Ot Primary system net heat losses (8tu/hr)

=

H Core rated Stu/hr at f ull power.

=

For the purecses of this uncertainty analysis (ana basec on H notea above) it is assumec that the plant is at 10C% RTP when the measurement is taken. Measurements per#crmed at lower power levels will result in Ob 4.-

~

1 O

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 themal output of the steam generator is detamined by a calorime-tric measurement defined as:

Ogg (h, - h ) Wf Eq. 5

=

f where; h,

Steam enthalpy (Btu /lb)

=

Feedwater er.;.halpy (Btu /lb) h

=

f Feedwater flow (1b/hr).

W

=

.f The steam enthalpy is based on the measurement of steam generator outlet steam pressure, assuming saturated conditions. The feedwater ent,halpy is based on the measurtment of feedwater temperature and an assumed feedwater pressure based on steamline pressure, plus 100 psi. The feed-water flow is detemined by multiple measurements and a calculation based on the following:

f (X)(F,) (V Er te)

Eq. 6 W

=

where:

Feedwater venturi flow coefficient K

=

F, Feedwater venturi correction for themal expansion

=

2 Feedwater density (!b/ft )

ef

=

Feedwater venturi pressure drop (inches H O).

ap

=

2 The feecwater venturi flow coefficient is the product of a nu=ber of constants including as-built dimensions of the venturi and calibration tests perfomed by the vendor. The themal expansion correction is based on the coefficient of expansion of the venturi material and the l

l l

cb l

l

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 differtntial pressure call connected to the venturi.

The RCP heat adder is detamined by calculation, based on the best esti-mates of coolant flow, pump head, and pump hydraulic efficiency.

The primary system net heat losses an deterniined by calculation, con-sidering the following system heat inputs and heat losses:

Charging flow Letdown flow Seal injection flow RCP themal barrier cooler heat removat Pressurizer spray flow Pressurizer surge line flow Component insulation heat losses Component support heat losses CRDM heat losses A single calcuated sum for full power operation is used for these los-ses/ heat inputs.

The core power measurement is based on the following plant measurements:

Steamline pressure (P )

s Feedwater temperature (T )

f Feedwater pressurt (P )

f Feedwater venturi differential pressure (ap)

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

Feedwater venturi flow coefficient (K)

Feedwater venturi themal expansion correction (F,)

Feedwater density (:f)

ICb r

s--

+e e--e,--

y v-

Feedwater enthalpy (h )

f Steam enthalpy. (h,)

Moisture carryover (impacts h )

s i

Primary system net heat losses (Qg)

RCP, heat adder (Q )

p Tnese measurements and calculati ns are presented schematically on Figure 1.

Starting off with the Equation 6 parameters, the detailed derivation of the measurement errors is noted below.

Feedwater Flow Each of the feedwater venturis is calibrate'd by the vendor in a hydrau-lic laboratory under controlled conditions to an accuracy of

[

]+a,b,c % of span. The calibration data which substantiates this accuracy is provided for all of the plant venturis by the respective vendors. An additional uncertainty factor of [

]**'C %

is included for installation effects, resulting in an overall flow coef-ficient (K) uncertainty of [

]**'C %.

Since steam generator themal_

output is proportional to feedwater flow, the flow coefficient uncertainty is expressed as [

] ?"'O %.powe r.

The uncertainty applied to the feedwater venturi themal expansion correc-tion (F,) is based on the uncertainties of the measured feedwater tem-perature and the coefficient of themal expansion for the venturi material, usually 304 stainless steel. For this material, a change of + 2*F in the feedwater temperature range changes F, by [

3a,b,c : and the steam generator themal output by the same amount. For this deriva-l tion, an uncertainty of [

]+a,c in feedwater temperature was l

assumed (detailed breakdown for this assumption is provided in the feed-water enthalpy section). This results in a total uncertainty in F and a

steam generator output of [

]'"'C 5.

lib v-e r.

-c--_v-----,,---m-,

-.,,,v--

-,<e-,--w------

- + - -.. -

Based on data introduced into the ASE code, the uncertainty in F, for 304 stainless steel is _+5 percent. This results in an additional uncer-tainty of [

]+a,c % in feedwater flow. A conservative value of

[

3a c 5 is used in this analysis.

Using the ASME Steam Tables (1967) for compressed water, the effect of a

[

3+a c error in feedwater temperature on the %fs

[

3+8'C 5'in steam generator themal output. An error of

[

]+a c in feedwater pressure is assumed in the analysis (detailed breakdown of this value is provided in the steam enthalphy section). This results in an uncertainty in Gof [

]+a,c ;

in steam generator themal output. The combined effect of the two results in a total

/o uncertainty of [

]**'C % in steam f

generator themal output.

Table Ib provides a listing of the instrumentation errors for feedwater ap (inciuding an allowance for the venturi as defined above) assuming display on the process computer. With the exception of the computer readout error, the electronics errors are in percent AP span and must be translated into pen:ent feedwater flow at' full power conditions.

This is accomplished by multiplying the error in percent ap span by the conversion factor noted below:

II Isoan of feedwater flow transmitter in of ncminal flow ) 2 1

100 For a feedwater flow transmitter span of [

]+8'C % nominal flow, the' conversion factor is [

]C (which is the value used for this analysi s).

As noted in Table 2b, the statistical sum of the errors for feecwater flow is [

]+8'C % of steam generator themal output.

l i

12b l

l

t

~ -

Feecwater Enthalpy The next major. error cogonent is the feedwater enthalpy used in Equa-tion 5.

For this parameter the major contributor to the error is the uncertainty in the feedwater temperature. Table Ib provides the' detailed error. breakdown for this tegerature measurement assuming indication on the process computer. Statistically summing these errors (utilizing

' Eq. 2) results in a total tegerature error of [

]+a c % span.

4 Assuming a span of [

]+a,c results in a tegerature error of

[

.].a,c A conservative, bounding.v.alue of [

]+a,c was

+

assumed for this analysis. Assuming smaller spans results in smaller temperature errors.

Using the ASME steam tables (1967) for compressed water, the effect of a

[

]+a,c error in feedwater temperature on the feedwater enthalpy (h ) is [

]+a.c % in steam generator thermal output.

f I

Assuming a [

]+a,c error in feedwater pressure (detailed break-down provided in the steam enthalpy section) results in a

[

]+a,c % effect in hf and s' team generator thermal output.

The combined effect of the two results in a total hf uncertainty of

[

]+a.c%.

A conservative value (based on rouno-off effects of individual instrumentation errors) of [

]+a,c % for hf uncer-tainty is used in this analysis (as notea on Table 2b).

Steam Enthaloy The steam enthalpy has two contributors to the calorimetric error, steamline pressure and the moisture content. For steamline pressure the errors are as rioted en Table 16, assuming display on the process compu-ter. This results in a total instrumentation error (utilizing Ec. 2) of

[

]+a,c % span. Based on a 1200 psig span this equals

[-

.].a,c A conservative value of [

]+a,c is assumec

+

in this analysis. The feedwater pressure is assumed to be 100 asi higher than the steamline pressure with a conservatively high measure-ment error of [

.]. a.c 7aole Ib provices a breakoown of expected errors if feecaater pressure is measure: cirectly anc ois;layec 13b

e on the process computer. The results indicate an expected error of

[

]+a,c, well within the assumed value.

Using the ASME Steam Tables (1967) for saturated water and steam, the effect of a [

]+a,c (g 3+a,c) error in steamline pressure on the steam enthalpy (h ) is [

]**.' C % in steam generator s

themal output. Thus a total instrumentation error of [

]**'C in steamline pressum results in an uncertainty of [

]+a,c 5 in steam generator themal output.

The major contributor to h uncertainty is moisture content. The s

nominal or best estimate perfomance level is assumed to be [

]+a,c g, which is the design Ifmit to protect the high pressure turbine. The most conservative assumption that can be made in regards to maximizing steam generator themal output is a steam moisture' content of zero. This conser-vatism is introduced by assigning an uncertainty of [-

]+C to the moisture content, which is equivalent through enthalpy change to

[

.]**'C % of themal output. The combined effect of the steamline pressure and moisture content on the total h uncertainty is s

[

]+C % in steam generator thermal output.

Looo Power The loop power uncertainty is obtained by statistically combining all of the error components noted for the steam generator thermal output (Q3g) in tems of loop power. Within each loop these components are inoependent effects (or fonned into independent. quantities) since they are independent measurements. Technically, the feedwater temperature and pressure uncer-tainties are common to several of the error components. However, they are treated as independent quantities because of the conservatism assumed and the arithmetic sumation of their uncertainties before squaring them has no significant effect on the final result.

i 14b

~,

y

.e,+,.-,,,-.--y--,-

..-...,..-...,..-.--...--r-

---..--,--,,-e

The only effect which tends to be dependent, affecting all loops, is the accumulation of crud on the feedwater ventuMs, which can effect the ap for a specified flow. Although it is conceivable that the crud

~

accumulation could affect the static pressure distribution at the ven-turi throat pressure tap in a manner that would result in a higher flow for a specified ap, the reduction in thmat area resulting in a lower flow at the specified ap is the stronger effect. All reported cases of venturi fouling have been associated with a significant loss in elec-tMcal output, indicating that the actual themal power has been below the measured power rather than above it. Losses in net power generation which have been correlated with venturi fouling have occurred in about half of the more than 20 Westinghouse pressurized water reactors oper-ating in the United States. These power losses have been generally in the range of two to three percent. Power losses have also occurred in at least three, and possibly five plants out of the more than ten West-inghouse plants operating abroad. In no case has venturi fouling been reported which resulted in a non-conservative feedwater flow measure-ment. Because the venturi crud fomations have resulted in a conserva-tive, reduced power condition, no uncertainty has been included in the analysis of power measurement error for-this phenomenon.

The net pump heat uncertainty is derived in the following manner. Th'e ~

primary system net heat losses and pump heat adder for a four loop plan are suir.tarized as follows:

Systems heat losses

- 2.0 MWt l

Component conduction and l

convection losses

- 1. 4 Pump heat adder

+18.0 l

Net Heat input to RCS

+14.6 MWt 1Cb

I i

The uncertainties for these quantities are as follows: The uncertainty on system heat losses, which are essentially all due to charging and letdown flows, has been estimated to be [

]+a.c % of the calculated value. Since direct measunements are not possible, the uncertainty on component conduction and convection losses has been assumed to be

[-

]+a,c 5 of the calculated value. Reactor coolant pump hydraulics are known to a relatively high confidence level, supported by the system hydraulics tests perfomed at Prairie Island II and by input power mea-surements from several plants, so the uncert'a'inty for the pump heat adder is estimated to be [

.]+"'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

[

]+"'C % of the total, which is equivalent to [

]*"'C 5 of core power.

The Total Loop Power uncertainty (noted in Table 2 as [

]+"'C %)

SG ' I' 3*

I' I

is the statistical sum of the Loop Power uncertainty (Q and the Net Pump Heat Addition, [

]+"'C %.

The Total Secondary Powe: uncertainty is the statistical combination of the Loop Power uncertainty and the number of primary side loops in the plant. As noted in Table 2b, the Secondary Power uncertainty for N loops is as follows:

4 uncertainty

+ 1.2 power N

=

3

+ 1.4 i power 2

+ 1.7 power In all cases the total secondary Power uncertainty is less than or equal to the historically used value of + 2 % power. For ITDP, credit is taken for the increased knowledge of reactor power and the values noted above are used in the ITDP error analysis, i.e., the standard deviation for reactor power, at the 95+% probability. level is:

1Gb

s

,x T

FIGURE 1 POWER CALORIMETRIC SCHEMATIC 7

P, P

j T

AP y

f

4. s

\\.-

,g h

h s

f, cf F

E a

U

=

0 - measured Q g O - calculated 1

Y l

b l

T Core.: ewer 1

s 175 s

n

,.4.,,,.

i l.

l TABLE 2b l.

SECONDARY POWER CALORIETRIC EASUREENT UNCERTAINTIES i

Power Comoonent Instrument Error Unce rtainty Feedwater Flow. '

+a c VentuM, K Thermal Expansion Coefficient Temperature MateM al Density Temperature Pressure Electronics ao Cell Calibration Sensor Pressure Effects Sensor Temperature Effects Sensor Drift Rack Calibration Rack Temperature Effects Rack Drift Ccmputer Isolator Drift Computer Readout Total Electronics Error fr(e)2 (e)2 Total Feedwater Flow Error w

.n

. f t'_

[ y' l

a-

. i ?

TABLE 2b (Cont)-

t SECONDARY POWER CALORIMETRIC MEASUREMENT UNCERTAINTIES t

Power

' ' ' Conconent Instrument Error Uncertainty s

s

-(

n Feedwa'ter Enthalpy Tegerature (Electrer.ics)

+a,c-RTD Calibration R/I Converter

(

Rack Accuracy Rack ~ Tegerature Eff ects

~

Rack Drif t Coguter Isolator Drff t

k. '

Cog uter Reacout

. Total-Electronics Error fI(e)2

Fenewater Temperature Error Assumed Prr.ssure y

'Tokal Feecwater Enthalpy Error fI(e)2 Steam Enthalpy Steamline Pressure (Electronics)

Pressure Cell Calitr stion Sensor Temperature Eff ects Sensor Drif t Rack Calibration Rack Temperature Eff ects l

3Ct v

se g-m w,--

e.-

y

+

4 m

TABLE 2 b(Cont)

SECONDARY POER CALORIETRIC EASUREENT UNCERTAINTIES Power Comoonent Instrument Error Uncertainty Steam Enthalpy (Cont)

+a,c

+a,c Rack Drif t Computer Isolator Drift Computer Readout Total Electronics Error fI(el 2

Steamline Pressure Error Assumed Moisture Carryover Total Steam Enthalpy Error 'f fil)2 LoopPowerUncertaintyfz(e)2 Net Pump Heat Addition Uncertainty

~

Total Loop Power Uncertainty (8)

Total Secondary Power Uncertainty

[I(e)2]/N where N = 4 loops

+ 1.2t 3 loops

-- 1.at 1

2 loops

+ 1.7t

~

.---.--...s.

NOTES FOR TABLE 2 b 1.

Temperature effect on Themal Expansion Coefficient is assumed to be linear with an uncertainty of [

]+a,b,e per 2*F change.

2.

Conservative assumption for value, 'particularly if steamline pressure

~

+ 100 psi is assumed value. Uncertainty for steamline pre'ssure noted in Steam Enthalpy.

3.

To transfom error in percent ap span to percent of feedwater flow at 100% of nominal feedwater flow; multiply the instrument error by:

2

[1/2 Sean of feedwater flow transmitter in cercent of nominal flow

(

/\\

100

/

In this analysis the feedwater flow transmitter span is assumed to be

[

]+a,c % of nominal flow.

4 In this analy' sis assumed an error of [

]*"'C and a maximum swing in feedwater pressure from no load to full power of [

].C 5.

[

3+a,c 6.

[

]+a,c span of [

]*"'C equals [

]+a,c wnich equals

[

1+a c power.

7.

Conservative. assumption for instrumentation error for this analysis.

8.

Statistical sum of Loop Power Uncertainty and Net Pump Heat Addition Uncertai nty.

i l

l 21b

~

4 1

+a.c N

=4 a.

power 3

power 2

powe r J

44,RCS FLOW The Improved Thermal Design Procedure (ITDP) and some plant Tech-nical Specifications require an RCS flow measurement with a high degree of accuracy. It is assumed for this error analysis, that this flow measurement is perfomed within seven days of calibrating the measurement instrumentation therefore, drift effects are not included (except where necessary due to sensor location). It is also assumed that the calorimetric flow measurement is performed at the beginning of a cycle, so no allowances have been made for feed-water venturi crud buildup.

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 ( Ah) of the primary coolant. Assuming that the primary and secondary sides are in equilibrium; the RCS total vessel fimv is the sum of the individual primary loop flows, i.e.,

RCS " 23 W (Eq. 7)

W L

The individual primary loop 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 scecific volume of the RCS cold leg. The equation for this calculation is:

L=(T)h l0 \\ Y L

/033 - Ce + ( ;-) ( (V )

e (Ec. a) nuH ~ "c2 7.~.-

i

'0CP II'" IIP"I where; W L

3 y

0.1247 gpm/(ft /hr)

=

Steam Generator themal output (Stu/hr)

Q

=

3g RCP heat adder (Btu /hr)

Q,

=

QL Primary system net heat losses (Btu /hr)

=

3 e

C (ft /lb)

Y Specific volume of the cold leg at T

=

Number of primary side loops N

=

H Hot leg enthalpy (Btu /lb) h

=

Cold leg enthalpy -(Btu /lb).

h

=

e The themal output of the steam generator is detemined by the same calorimetric measurement as for reactor power, which is defined as:

Qgg = (h -h)W (Eq. 5) s f

f Steam enthalpy (Btu /lb) where; h

=

s f

Feedwater enthalpy (Stu/lb) h

=

Feedwater flow (1b/hr).

W

=

f The steam enthalpy is based on measurement of steam cenerator outlet steam pressure, assuming saturated conditions. The feedwater enthalpy _

is based on the-measurement of feedwater temperature and an assumed feedwater pressure based on steamline pressun plus 100 psi. The feed-water flow is detemined by multiple measurements and the same calcula-tion as used for reactor power measumments, which is based on the fol-lowing:

f = (X) (F,){ V cf as}

(Eq. 6)

W Feedwater venturi flow factor where; K

=

F, Feedwater venturi correction for themal expansion

=

3 O

Feedwater density (1b/ft )

f Feedwater venturi pressure drop (inches H O).

.sp

=

2

?.;b

,n

The feedwater venturi flow coefficient is the product of a number of constants including as-built dimensions of the venturi and calibration tests perfomed by the vendor. The themal expansion correction is.

l based on the coefficient of expansion of the venturi mateHal and the difference between feedwater temperature and calibration temperature.

Feedwater density is based on the measurement of feedwater temperata:t and feedwater pressure. The ventuH pressure drop is obtained from the h

output of t' e differential pressurt cell connected to the ventuM.

The RCP heat adder is detemined by calculation, based on the best esti-mates of coolant flow, pump head, and pump hydraulic efficiency.

The primary system net heat losses are detemined by calculation, con-sideMng the following system heat inputs and heat losses:

Charging flow Letdown flow Seal injection flow RCP themal barrier cooler heat removal Pressurizer spray flow Pressurf ree surge l'ine flow Component insulation heat losses Component support heat losses CRDM heat losses.

A single calculated sum for full power operation is used for these los-ses/ heat inputs.

The hot leg and cold leg enthalpfes are based on the measurement of the hot leg temperature, cold leg temperature and the pressurizer pressure.

The cold leg specific volume is based on measurement of the cold leg temperature and pressurizer pressure.

The RCS flow measurement is thus based on the following plant measure-ments:

2no

l Steamline pressure (P )

3 Feedwater temperature (T )

f Feedwater pressure (P )

f Feedwater ventud differential pressure (Ap)

Hot leg temperature (T I H

Cold leg temperature (T )

C Pressurizer pressure (P )

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

Feedwater venturi flow coefficients (K)

Feedwater venturi themal expansion correction (F,)

Feedwater density (of)

Feedwater enthalpy (h )

f Steam enthalpy (h )

s Moisture carryover (impacts.h )

s Primary system net heat lossss (O )

t RCP heat acMer (0 )

p Hot leg enthalpy (h I

~

M Cold leg enthalpy (h ).

e These measurements and calculations are presented schematically on Figure 2.

Starting off with the Equation 6 parameters,. the detailed derivation of the measurement errors is noted below.

Feedwater Flow Each of the feedwater ventuM s is calibrated by the vendor in a hydrau-lics laboratory under controlled concitions to an accuracy of

[

]+a,b,c : of span. The calibration data which substantiates this accuracy is provided for all of the plant venturis by the respec-tive vendors. An additional uncertainty factor of [

]+a,c ; 9,

Gb

~

included for installation effects, resulting in an overall flow coef-ficient (K) uncertainty of [

]+a,c %.

Since RCS loop flow is proportional to steam generator themal output which is proportional to feedwater flow, the flow coefficient uncertainty is expressed as

[

1+a,c g fjoy, The uncertainty applied t6 the feedwater venturi themal exp'ansion cor-rection (F,) is based on the uncertainties of the measured feedwater temperature and the coefficient of themal expansion for the venturi material, usually 304 stainless steel. For this material, a change of

+ 2*F in the feedwater temperature range changes F, by

[

3+a,b,c $ and the steam generator themal output by the same amount. For this derivation, an uncertainty of [

j'"'C in feedwater temperature was assumed (detailed breakdown for this assump-tiori is provided in the feedwater enthalpy section). This results in a negligible impact in F, and steam. generator output.

Based on data introduced into the ASE Code, the uncertainty in F, for 304 stainless steel is + 5 %.

This results in an additional uncertainty of [

]**'C : in feedwater-flow. A conservative value of

[

3+a,c % is used'in this analysis.

Using the ASE Steam Tables (1967) for compressed water, the effect of a

[

']+a c error in feedwater-temperature on the /e is f

[

.3+a,c 5 irr steam generator themal output. An error of

[

]+a,c in feedwater pressure is assumed in this analysis (detailed breakdown of this value is provided in the steam enthalpy secti on). This results in an uncertainty in

cf of [

]+a,c in steam generator themal output. The combined effect of the two results in a total / o f uncertainty of [

]+a,c : in steam generator themal output.

It is assumed that the ap cell (usually a Barton or Rosemount) is read locally and soon after the. ap cell and local meter are calibratec (within 7 days of calibration). This allows the elimination of process 25b

rack and sensor drif t errors from consideration. Therefore, the op cell errors noted in,this. analysis are [

3+a,c % for calibration and[

]+a,c % for reading error of the special high accuracy, local gauge. These two errors are in % ap span. In order to be

. useable in this analysis they must be translated into % feedwater flow at full power conditions. This is accomplished by multiplying the error in % ap span by the conversion f actor noted below:

I\\I l

span of feedwater flow transmitter in percent of nominal flowh2

'2 100,,

For a feedwater flow transmitter span of [

]+a,c %. nominal flow, the conversion f actor is [

]+a,c (which is the value used in this analysis).

As noted in Table 3b, the statistical sum of the errors for feedwater flow is [-

]+a,c % of steam generator thermal output.

Feedwater Enthaley The next major error corrponent is the feedwater enthalpy used in Equa-tion 5.

For this parameter the major contributor to the error is the uncertainty in the feedwater temperature. It is assumed that the feec-water tempersepre is determined through the use of an RTD or thermo-couple' whose output is read by a digital voltmeter (DVM) or digital multimeter (DMM) (at the output of the RTD or by a Wheatstone Bricge f or RTD's, or at the reference junction for thermocouples).

It is also assumed that the process components of the above are calibrated within 7 days prior to the measurement allowing the elimination of drif t effects. Therefore, the error breakdown for f eedwater temperature is as noted on Table Ib.

The statis'ical combination of these errors results in a total feedwater temperature error of [

3+a,c, 27b I

____-______-_-_-_-_-__-------A

Using the ASME St.eam Table (1967) for compr'essed water, the effect of a

[

]+a,c error in f eedwater temperature on the feedwater enthalpy (h ) is [

j+a,c%insteamgeneratorthermaloutput.

f Assuming a [

]+a,c error in feedwater pressure (detaileo break-down provided in the steam enthalpy section) results in a

[

]+a,c % effect in hf and steam generator thermal output.

The combined effect of the two results in a total hf uncertainty of

]+a.c % steam genera' or thermal output, as noted on Table 3b.

[

t Steam Enthalor The steam enthalpy has two contributors to the calorimetric error, steamline pressure and the moisture content. For steamline pressure the error breakdown is as noted on Table Ib. This results in a total instru-mentation error of [

]+a c %, which equals [

]+a,cfora 1200 psi span. For this analysis a conservative value of [

]+a'c is assumed for the steamline pressure. The feedwater pressure is assumed to be 100 psi higher than the steamline pressure with a conser-vatively high measurement error of [

]+ac.

If feedwater pres-sure is measuredon the same basis as the steamline pressure (with a DVM) theerroris[

]+a,c % span, which equals [

]+a,c f or a 1500 psi span. Tnus,.an assumption of an error of [

}-a,cis very conservative.

Using the ASME Steam Tables (1967) for saturated water and steam, the effect of a f

]+a,c ([

]+a,c) error in steamline pressure on the steam enthalpy is [

]+a,c % in steam generator thermal output. Thus, a total instrumentation error of [

]+a,c resul ts in an uncertainty of [

]+a.c % in steam generator thermal output, as noted on Table 3b.

The major contributor to h uncertainty is moisture content. The s

nominal or best estimate performance level is assumed to be [

]+a.c; which is the design limit to protect the high pressure turoine. The most ccnservative assum: tion that can be ma:e in regarcs to maximizing stea-

^

20b l

generator themal output is a steam moisture content of zero. This conser-vatism is introduced by assigning an uncertainty of [

["'"%tothe moisture content, which is equivalent through enthalpy change to C

7" ' C 1 of themal output. The combined effect of the steamline pressure and moisture content on the total h uncertainty is s

[

pa c 5 in steam generator themal output.

Secondary Side Loco Power The loop power uncertainty is obtained by statistically combtning all of the error components noted for the steam generator themal output (Q3g) in tems of Stu/hr. Within each loop these components are independent effects since they are independent measurements. Technically, the feed.

water temperature and pressure uncertainties are connon to several of the error components. However, they am treated as independent quantities because of the conservatism assumed and the arithmetic succation of their uncertainties before squaring them has no significant effect on the final resul t.

The only effect whictr tends to be dependent, affecting all loops, would be the accumulation of crud on the feedwater venturis, which can affect the ap for a specified flow. Although it is conceivable that the crud accu-mulation could affect the static pressure distribution at the venturi throat pressure tap in a manner that would result in a higher flow for a specified ap, the reduction in throat area resulting in a lower flow at 4

the specified ap is the stronger effect No uncertainty has been included in the analysis for this effect. If venturi fouling is detected l

by the plant, the venturi should be cleaned, prior to perfomance of the measurement. If the venturi is not cleaned, the effect of the fouling on the detemination of the feedwater flow, and thus, the steam generator

)

power and RCS flow, should be measured and treated as a bias, i.e., the j

error due to venturi fouling should be added to the statistical succation of the rest of the ceasurement errors.

2Cb 1

-...--.o

...-,v,,

,,-._.,,,.--,,,ym.,,,,.-.-.,,.,_rm.,m,--

y,, _,,,

-m,y,-,,.-..-,,-

The net pump heat uncertainty is derived in the following manner. The primary system net heat losses and pump heat adder for a four loop plant are summarized as follows:

System heat losses

-2.0 MWt Component conduction and convection losses

-1. 4 Pump heat adder

+18.0 Net Heat input to RCS

.+14.6 MWt The uncertainties for these quantities are as follows: The uncertainty on systems heat losses, which is essentially all due to charging and letcown flows, has been estimated to be [

]+a,c % of the. calculated value. Since direct measurements are not possible, the uncertainty on component conduction and convection losses has been assumed to be

[

]+a c % of the calculated value. Reactor coolant pump hydraulics are known to a relatively high confidence level, supported by the system hydraulics tests performed at Prairie Island II and b'y inout power mea-surements f rom several plants, so the uncertainty for the puna heat adder is estimated to be [

']+a,c % of the best estimate value.

Considering these parameters as one quantity which is cesignatec the net puen heat uncertainty, the combinec uncertainties are less than

[

]+a,c % of the total, which is [

]+a,c % of core power.

The Total Secondary Sice Locp Power Uncertainty (noted in Tadie 35 a's

[

]+a,c %) is the statistical sum of the secondary sice locp power uncertainty (0 g), [

]+a.c %, and the net pump heat addi-3 tion,[

]+a,c%.

Primary Side Enthaley The primary side enthalpy error contributors are Tg and TC measure-ment errors and the uncertainty in pressurizer pressure. The instrumen-tation errces for Th are as notec on Table 10.

These errers are :asec 30b

l-t on the. assumption that the DVM has been recently calibrated (within 7 days prior to the measurement)~ and the DVM is used to read the output of the RTD, or a bridge, thus allowing the elimination of drif t effects in the racks. The statistical combination of the above errors results in a total TH uncertainty of [

]+a,c, Table Ib also provides the instrumentation error breakdowr! for T. The C

errors are based on the same assumptions as for T, resulting in a H

total TC uncertainty of [

]+a,c, Pressurizer pressure instrumentation errors are noted on Table Ib.

A, sensor drift allowance of [

]+a,c % is included due to the dif-ficulty in calibrating while at power. It is assumed calibration is performed only as required by plant Technical Specifications.

Statistically combining these errors resu,1ts in the total pressurizer pressure uncertainty equaling [

']+a,c % of span, which equals

[

]+a,c for an [

]+a,c span. In this analysis a conservative value of [

]+a,c is used for the. instrumentation error for pressurizer pressure.

The effect of an uncertainty of [

]+a,c in TH on hg is

[

]+a,c % of loop flow. Thus, an error of [

]+a,cin Tg introduces an uncertainty of [

]+a,c percent in h. An g

error. of [

]+a,c in TC isworth[

]+a,c % in h.

e Therefore, an error of [

]+a c in TC results in an uncer-tainty of [

}+a,c % in h and loop flow. An uncertainty of e

[

']+a,c in pressurizer pressure introduces an error of

[

]+a,c%inhgand[

]+a.C % in h. Statistically e

combining the hot leg and cold leg temoerature and pressure uncertain-ties results in an hH uncertainty of [.

]+a,c %, an he uncer-tainty of [

3+a,c%, and a total uncertainty in th of

[

]+a.c % in icop flow.

Statistically comcining tne Total Seconcary Sice !.ocp Fower uncertainu (in Stu/hr) with the primary side enthalpy uncertainty (in Stu/lb),

Olb

FIGURE 2 RCS FLOW CALORIMETRIC SCHEMATIC T

P T

P(-

P T

ap i

H p

g f

f h

h h,

h p

K H

C p

of a

1r

'Ah w

1r 1

Oss

'I Measured O Q

Calcula ed D

~

L 4

e Q Y

9 L,.

m Other Lecos T

RCS Flow

'.2b

~

4 TABLE 3b CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES Flow Comoonent Instrument Error (1)

Uncertainty Feedwater Flow

+a,c Venturi, K Thermal Expansion Coefficient Tegerature Material Density Teg erature Pressure Instrumentation ap Cell Calibration ap Cell Gauge Readout Total Instrumentation Errorf;(e)2 Total Feedwater Flow Error f!(eJ 2

i Feedwater Enthalpy Temperature (Electronics)

RTD Calibration DVM Accuracy Total Temperature Error ft(e)2 Pressure Total Feedwater Enthalpy Error fI(e)2 6

33t

~

TABLE 3b (Cont)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES Flow Conoonent Instrument Error (1)

Uncertainty

+a.c Steam Enthalpy Steamline Pressure (Electronics)

Pressure Cell Calibration Sensor Temperature Effects Rack Calibration Rack Temperature Effects DVM Accuracy Total Electronics Error [I(e)2 Steamline Pressure Error Assumed Meisture Carryover I

Total Steam Enthalpy Error jI(e)2 l

Seconcary Sice Loop Power Uncertainty j I(e)2 Net Pug Heat Addition Uncertainty

+ 20%

l Total Seconcary Side Loco Power l

Uncertainty jI(e)2 Primary Sice Enthalpy Tg (Electrcnics)

RTD Calibration DVM Accuracy g Instrumentation Error fI(e)2 T

T Temoerature Streamino Error H

gTemeratureErrorfI(e)2 T

34b

TABLE 3b (Cont)

CALORIMETRIC RCS FLOW MEASUREMENT UNCERTAINTIES Flow Comoonent Instrument Error (1)

Uncertainty

+a.c TC (Electronics)

RTO Calibration DVM Accuracy

~

InstrumentationErrorft(e)2 TC Pressurizer Pressure (Electronics)

Pressure Cell Calibration Sensor Temperature Effects Sensor Orif t Rack Calibration Rack Termerature Effects DVM Accuracy Total Pressurizer Pressure Error (I(e)2 Pressurizer Pressure Error Assumed Tg Pressure Effect Total Error fI(e)2 TH T Pressure Effect C

l Total Error j!(e)2 T C Total ah Uncertainty g/ (e)2

. Primary Sice Loop Flow Uncertaintyft(e)2 Total RCS Flow Uncertainty j [I(e)2]/N where N = 4 loops 2 1.5%

3 loops 2 1.757.

2 loops 2 2.h 35b

NOTES FOR TABLE 3b 1.

Measurements performed within 7 days af ter calibration thus Rack Drif t,

~

and where possible Sensor Drif t, eff ects are not included in this analy-sis.

2.

Conservative assumption for value, particularly if steamline pressure

~

+ 100 psi is assumed value. Oncertainty f or steamline pressure noted in steam.enthalpy..

3.

To transform error in percent ap span to percent of feedwater flow at 100% of nominal feedwater flow; multiply the instrument error by:

[l/2 Sean of feedwater flow transmitter in cercent of nominal flow \\ 2

(

/\\

100

/

In this analysis the feedwater flow transmitter span is assumec to be

[125]+a,c % of nominal flow.

4 Reading error for multiple readings of a Barton gauge.

5.

Conservative assumption for instrumentation errcr for this analysis.

6.

Maximum allowed moisture carryover to protect HP turbine.

7.

Calibratier accuracy of [

]+a,c span of [

]+a.c which equals

[

j+a, c,

8.

Credit taken for tne 3 tap scoop RTD bypass loco in recucirg uncertain-ties due to temperature streaming.

9.

Convoluted sum of Tg Temperature Error and TH Pressure Eff ect.

10. Ccnvoluted sue cf TC nstrumentation Error anc TC Pressure Eff ec*.

11. Convolutec sum of is Total Error and TC Total Error.

1 i

30D

results in a Primary $1de Loop Flow Uncertainty of [

3**'# % loop flow. The RCS flow uncertainty is the statistical combinstion of the primary side loop flow error and the nunter of primary side loops in the plant. As noted in Table 3b, the RCS Flow uncertainty for N loops is:

+ 1.5 % flow N=4 uncertainty

=

+ 1.75 % flow 3

+ 2.1 % fl ow.

2

=

For ITDP, credit is taken for the increased knowledge of RCS flow and the values noted above are used in the ITDP error. analysis, i.e., the standard deviation for RCS flow, at the 95+5 probability level is:

+a,c 5 flow N=4 e

=

% flow 3

=

5. flow 2

=

5.

USE OF AN LEFM If a plant uses a Leading Edge Flow Meter (LEFM), from the Oceanics Division of Westinghouse, for the measurement of feedwater flow, several changes are made ir *.he calorimetric power and flow uncertainty analy-I ses. The following are typical LEFM uncertainties in mass flow (1bs/hr):

a.

A nominal accuracy of [

3+"'C fl ow.

This is based on a feedwater temperature uncertainty of [

1**'C and a feedwater pressure uncertainty of [

3+a,c, b.

For each [ ]+a,c increase in Feedwater temperature uncer-tainty, the mass flow uncertainty increases by [

1+a,c, c.

For a feedwater pressure uncertainty greater than

[

3+a,c but less than [

]+a,c, the mass flow uncertainty increases by [

3+a,c, 37b r

< - - + ~ -

9

_. _ a.

Thus, for a typical LEFM installation with a feedwater temperature 1

uncertainty of [

J+"'C and a pressurt uncertainty less than

[

]+"'C, the mass flow uncertainty is i l***C fl ow.

The effect of the use of an LEFM is seen primarily in the measurtment of Reactor Power. The following table provides a comparison of the uncer-tainties for a power calorimetric using a feedwater venturf 'and an LEFM. It is assumed for these calculations that a measurement device (either a venturi or an LEFM) is in the feedwater line to each steam generator.

t e

i o

l l

30b l

i

~

j i

TABLE 4b COWARISON OF VENTURI VS. LEFM POWER CALORI!ETRIC UNCERTAINTIES Venturi

LEFM Reactor Power

+a,c Feedwater Temperature Feedwater Flow Feedwater Enthalpy Steam Enthalpy Loop Power Uncertainty Total Loop Power Uncertainty Total Secondan Power Uncertainty k 4 loops

+ 1.2% RTP

+ 0.4% RTP 1.4% RTP

+ 0.4% RTP 1

3 loops Z loops

1.7% RTP

+ 0.55 RTP from Table 2 due to [

]+"'C assumption The impact of the LEFM on RCS Flow measurement is considerably less (pHmadly due to the f

]+a,c feedwater temperature error al ready being assumed and the pHme error contHbutor* being T and T #0" H

C pHmary side ah). However, the following table notes the differences between the two measurements for an RCS Flow calorimetHc measurement. For these calculations it is assumed that a measurement device (either a' venturi or an LEFM) is in the feedwater line to each steam generator.

I 39

TABLE Sb COMPARISON OF VENTURI VS. LEFM FLOW CALORIMETRIC LNCERTAINTIES Venturi

LEFM RCS Fl ow -

---" + a, c Feedwater Flow Feedwater Enthalpy Steam Enthalpy Secondary Loop Power Uncertainty Total Secondary Power Uncertainty Primary Enthalpy Primary Loop Flow Uncertainty Total RCS Flow Uncertainty 4 loops 1

3 5% flow 31.45% flow 3 loops 3 1.75% flow 21.7% flow 2 loops 12.1% flow 12.05% flow f rom Table 3b due to [

]+a c assumption Theref ore, if a plant has installed an LEFM to measure f eeowater flow credit would be taken in the ITDP error analysis for the lower uncer-tainty in Reactor Power, but no credit would be taken in RCS flow.

6.b NORMALIZED ELBOW tap 5 FOR RCS FLOW.YEASUREMENT Based on the results of Table 3b, in order for a plant to assure opera-tion within the ITDP assumptions an RCS flow calorimetric would have to be perf ormed once every 31 EFPD. However, this is an involved proceoure which requires considerable staff anc setup time. Therefore, many plants perform one flow calorimetric at the beginning of the cycle anc normalize the loop elbow taps. This allows the coerator to quickly cetermine if there has been a significant reduction in loop flow on a snif t basis anc to avoic a long contnly procecure. Tne elbow ta:s are 205

-me

forced to read 1.0 in the process racks af ter performance of the full power flow calorimetric, thus, the elbow tap and it's ap cell are seeing normal operating conditions at the time of calibration / normal-ization and 1.0 corresponds to the measured loop flow at the time of the measurement.

For monthly surveillance to assure plant operation consistent with the ITOP assunctions, two means of determining the RCS flow are available.

One, to read the loop flows from the process computer, and two, to mea-sure the output of the elbow tap Ap cells in the process racks with a DVM. The uncertainties for both methods and their convolution with the calorimetric uncertainty are presented below.

Assuming that only one elbow tap per loop is available to the process computer results in the following elbow tap measurement uncertainty:

%Ap span

% flow

%ap span

% flew RCA{

+a c

a,c pg PEA RTE SCA RD SPE ID

~

STE A/D 50 Readout 1

ao span is converted to flow on the same basis as provided in Note 3 of Table 3L for an instrument scan of [

]+a,c Using Eo. 2 results inaloopuncertaintyof[

]+a c flow per 1000. The total uncer-tainty for N locos is:

+a c

= 4 f)g, N

3 2

The instrument / measurement uncertainties for nor alized elbow ta::s anc the flow caloriretric are statistically incecencent and are 95+5 oret-ability values. Therefore, the statistical co-tinaticn of the standard deviations results in the following total flow uncertainty at a 95+5 j

probability:

elt

4 loops

+ 1.7% flow 3 loops 1 2.0 2 loops 1 2.3 Another method of using normalized elbow taps is to take DYM reaoings in the process racks of all three elbow taps for each loop. This results in average flows for each loop with a lower instrumentation uncertainty f or the total RCS flow. The instrumentation uncertainties.for this 1

measurement are:

%4p scan

% flow

%ap span

% flow pg

+a,c

+a,c SD PEA RCA SCA RTE SPE RD STE OVM Readout ap span is converted to flow on the same basis as provided in Note 3 of Table 3b for an instrument span of [

]+a,c Using Eq. I results in a channel uncertainty of [

]+a,c flow. Utilizing three elbcw taps (which are independent) results in a 1000 uncertainty of [

]+a c flow per loop. The total uncertainty for N loops is:

N = 4

+a c f) c, 3

2 The calorimetric and the above noted elbow tap uncertainties can be statistically combinec as noteo earlier. The 95+% probability total flow uncertainties, using three elbcw taas per locp are:

4 locps

+ 1.c% flow 3 locos

1. 8 2 loc: s 2.2 i

Ihs f ollowing table summarizes RCS flow measurecent uncertainties.

l I20 l

TABLE 6b TO AL FLOW MEASUREMENT UNCERTAINTIES Loops 4

3 2

Calorimetric uncertainty *

+ 1.5

+ 1.75

+ 2.1 Total' uncertainty 3 elbow taps / loop

+ 1.6

+ 1. 8

+ 2.2 Total uncertainty 1 elbow tap / loop

+ 1. 7

+ 2.0

+ 2.3 Calorimetric uncertainty noted assum[s feedwater measurement with a venturi, however, use of an LEFM for feedwater measurement results in essentially the same value.

IV.

PROSABILITY JUSTIFICATION As noted in Section III, it is Westinghouse's belief that the total uncertainty for Pressurizer Pressure, Tayg, Reactor Power, and RCS Flow are normal, two sided, 95'+% probability distributions. This sec-tion will substantiate that position with a comparison between three approaches, the first being that noted in Section II, the second involves determination of the variance assuming a uniform probability distribution for each uncertainty and then determination of the 95%

probability value assuming a one sided normal distribution, and the third involves determination of the variance assuming a normal, two sidec probability distribution for each uncertainty and then determina-tion of the 95% probability value assuming a two sidea nomal distribu-tion.

Table 7b lists the results of the three approaches. Column 1 lists the values noted for CSA on Table Ib whicn are determined through the use of equations 1, 2, or 3, whichever is applicable to that particular f unc-tion. Column 2 lists' the variance for each function assuming the uncer-tainty for each of the parameters listed in Section 2 is a unifom prob-ability cistribution. For this assumptior.,

I 41b l

2 2

R 77r-Eq. 9 e =

where R equals the range of the parameter. The variance for the func-tion equals the arithmetic sum of the parameter variances. From a safety point of view deviation in the direction of non-conservatism is important. Themfore, Column 3 lists the one sided 955 probability values based on the variances provided in Column 2, i.e., the one sided 955 probability value.for" near normal distribution can be reasonably approximated by: 1.645

,2, Column 4 lists the variance for each function assuming '.he uncertainty for each of the parameters listed in Section 2 is a near normal, two sided probability distribution. Efforts have been made to conserva-tively determine the probability value.for each of the parameters, see Table 8.

For example, [

]+a,c The corre-spending I value listed on Table 8 is from the standard normal curve wt ere:

I = (x - u)/o Eq.10 The variance for a parameter is then the square of the uncertainty divided by its Z value:

2

,2,

uncertainty q,

r 1,

~

i The variance for the function equals the aMthmetic sum of the parameter variances. From th'e variance the two' sided 955 robability value for a normal distHbution can be' calculated: 1.96 e2, To sumarize; Column 1 is'the results of Equations 1, 2, and 3.

Column 2 is the total vaHance assuming uniform probabilty distributions, i.e.,

N A

+... = (2.uncj (2 unc 2

1 + 2 2

+

Eq.12 o=

Column 3 is 1.645 i

Column 4 is the total vaHance assuming near normal pmbability, distri-butions, i.e.,

c, g

s.

g = (unc3)2

( unc )E Eq. 13 g

+

+...

o Column 5 is 1.96 1

A comparison of Columns 1, 3, and 5 will show that the approach used in Section 2 results in values more conservative than those of Columns 3 and 5.

Thus, it can be concluded that the results presented in section 3 are total uncertainties with probabilities in excess of 955.

f Confidence limits are applicable only to a particular data set, which in 4

this case not availabl,e. Therefore, based on the relatively small num-ber of reports indicating large values of deviation, i.e., the number of instances where a channel fails a functional test is very small as com-

- pared to the many thousands of functional tests performed, WestinrJhouse believes that the total uncertainties presented on Table Ib are 95% preb-ability values at a high confidence level.

\\

m.

e-

,m, m

,.r.

,,-,-c,

,-,-w----

,.--.-,.-,--r-c-,

+ - - - - -

j V.

CONCLUSIONS The pmceding sections provide what is believed to be a reasonable means of accounting for instrument and measurement errors for four parameters used in the ITCP analysis. The assumptions used in this response are generic and conservative. It is the intent of this response to generi-cally resolve any concerns with the measumment and control'of Reactor Power, RCS Flow. Pressurizer Pressure and T as they are applied to ayg ITDP. As such, plant specific responses will provide only that infoma-tion which indicates that,1) the instrument and measurement uncertain-ties for that plant are consistent with or conservative with respect to those presented here, or 2) specific instrument and/or measurement uncertainties for that plant are not consistant with those presented.

In the second case the impact of the inconsistency on the four param-eters will be provided with corresponding new total uncertainties if the impact is sufficiently large.

D

}

G

I.

TADLE 7b I

f COMPARISON OF STATISTICAL K Til0DS

'l 2

3 4

5 Yarlance 951 Probabilfty Variance 95% Probability Method 1 Hethod 2 Method 2 Method 3 Method 3 I

Pressurizer Pressure - Control ta.c T,yg - Control Steamline Pressure - Computer l

Feedwater Temperature - Computer i

p, feedwater Pressure - Computer feeduater Ap - Computer Pressurlier Pressure - DVH Steauline Pressure - DVH

-l Feedwater Temperature - DVH l

Tgg - DVH l

l TC - DVH I

Notes for Table 7 h l

1.

Uncertainties presented in columns 1, 3, and 5 are in 1 span.

2.

While values noted are listed to the second decimal place, values are accurate only to the first 8

decimal place. Second place is noted for round-off purposes only.

~

TABLE 8 UNCERTAINTY PROBABILITIES.

Two Sided Two Sided Normal Probability (%)

Normal, Z Value

+a,c PMA PEA SCA SD STE SPE RCA RD RTE DVM ID A/D CA e

a

O REFERENCES 1.

Westinghouse letter NS-CE-1583, C. Eicheldinger to J. F. Stolz, NRC, dated 10/25/77.

2.

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

3.

Westinghouse letter NS-TMA-1837,' T. M." Anderson to S. Yarga, NRC, dated 6/23/78.

4.

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

5.

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

6.

NUREG-0717 Supplement No. 4, Safety Eval'uation Report related to the operation of Virgil C. Summer Nuclear Station, Unit No.1 Docket 50-395, A.ugust, 1982.

7.

NRC proposed Regulatory Guide 1.105 Rev. 2, " Instrument Setpoints",

dated 12/81 for implementation 6/82.

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

9. ANSI /N719 ISA Standard 567.04, Draft F, 5/22/79, "Setpoints for Nuclear Safety-Related Instrer.entation Used in Nuclear Power Plants".

10. Scientific Apparacus Manufacturers Association, Stancard PMC-20-1-1973, " Process Measurement and Control Terminology".

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v t e Letter Sequence Other
TAC:49969, (Approved, Closed) TAC:49970, (Approved, Closed)
Initiation Request, Request, Request, Request, Request Acceptance... Supplement Administration Questions Answers Results Approval Other: 05000301/LER-1983-007-01, /01T-0:on 830523,higher than Normal Levels of I-131 in Primary Coolant Prompted Investigation of Fuel Assemblies.Probably Caused by Fretting at Bottom Spring Clip Grid Elevation Optimized Fuel assemblies,ZD3, ML20069D934, ML20077M425, ML20077M437, ML20093E118, ML20094A054, ML20095F318, ML20107D684, ML20108A768
MONTHYEAR1984-10-05 [Table View]

ML20095F318

Text

Dear Mr. Denton:

Reference:

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