ML20076M731

Third Reload Submittal

ML20076M731
Person / Time
Site:	Monticello
Issue date:	12/11/1974
From:	NORTHERN STATES POWER CO.
To:
Shared Package
ML20076M729	List: ML20076M725 ML20076M731
References
NUDOCS 9102120439
Download: ML20076M731 (43)

Text

,,.

t NORTitERN STATES P(MER C(NPANY MCllTICELLO NUCLEAR Cl2iEHATING PIANT

)

W )

G Dl1RD RELMD SUlWITTAL DECIXBER 11, 1974 910212043'1 7 41 U l 1 PDR ADUCK ObOOOW23 I'lif(

p Prepared By Information Supplied By llorthern States Power General Electric Company Ccupany

1 LIST OF ILLUSTRATIONS Figure Titic lage 2-1 Monticello Reload 3 Reference Core Loading 2-2 31 Diagram of a Finger Spring 32 6-1a Peak Cladding Temperature Versus Planar Average Exposure 6-3 6-1b Maximum Average Planar Linear Heat Generation Rate Versus Planar Average Exposute 6-3 6-2 Doppler Reactivity Coefficient Versus Average Fuel Temper-ature as a Function of Moderator Condition 6-4 6-3 Accident Reactivity Shape Function 6-5 6-4 Scram Reactivity Function 6-6 6-5 Scram Reactivity - Monticello roc 4 6-11 6-6 Turbine Trip Without Bypass 6-14 6-7 Inad/ertent Start of HPC1 Pump 6-16 6-8 MCPR Versus Rod Position 6-17 6-9 RBH Response to Cor. trol Rod Motion for Rod Withdrawal Error - Limiting Case, Channel A + C for Reload 3 6-18 6-10 RBH Response to Control Rod Motion for Rod Withdrawal Error - Limiting Case, Channel B + D for Reload 3 6-18 6-11 HSIV Closure, Flux Scram 6-20 6-12 Core Reactivity Stability 6-22 1

)

e 0

4 y

1.

INTRODUCTION This document provides the supplemental information for Reload 3 at the Monticello Nuclear Generating Station.

The technical bases, generic design in-formation, and safety analyses are given in Reference 1.

The design reference core loading is based on the use of 80 8x4 bundles having an enrichment of 2.50 wt I U-235.

Also available will be up to 40 initial cort 7x7 fuel bundles (average enrichment 2.25 vt ! U-235) discharged at the and of Cycle 2 with an average exposure of 13,000 mwd /t. The objective of this out-age is-to reduce. plant offgas and to load the reactor core to ensure sufficient reactivity to operate the plant at its licensed power level for approximately a 9-month cycle.

O 1-1 l

a 2.

SUHKARY The design reference core configuration for this submittal consists of bun-dies shown in Table 2-1.

The relative locations of the 80 nev 8x8 fuel bundles and the fuel from previous reloads is shown in Figure 2-1.

Table 2-1 FUEL TYPE AND NUMBER Fuel Type N umber Initial (7x7) 268 Reload 1 (7x7) 20 Reload 2 (8xB) 116 Raioad 3 (8x8) 80 Total 484 2-1

12 2l 3l3 l2 50 l2 l

2l 48 2l 3l 3l l

l3 l3 l2 2'

2' 3l, 3l,'2@

'2 e-l3 l3@#'l3 46

'2 3,

,2 44 2l

..e..

% ll32l @2!

8 l2@ l2 l2

'2 2

2' 42 3lo' l

l3 l3 l3 l2 3l 40 2l 3l 3l

'2@

'2 l @2l 2l 2l 2

2' 3s l2 l2 l2 i

3l 3l @3l 3;

3l ili l3 l3 l3 l>C l' 3

36 34 l2 l2* l2 l2@ l2 l @2l 2 l o l,

2l al 2

i i

32 2,

3, 3,

3, 3,

i,, i i

,3

,3

,3

,3

,2

..e..

30

'2 la l2 l2@ l2* l

>l 2l @2; al 2'

2' 0l l3 3l @ l

,i i,

i,,

i,i

,i 2:

i i

i

,3 i,

i,i

,i

,i

@i, es 3,.

24 l2 l2 l2b,l2 l2 l2 l

2l 2l 2l %2l 2l 2l C ll l3@ l3 l3 l3 l2 22 2l 3:

3l 3l 3l @ -l -

ili 2

2l 2l 2l

>l 20 l2 l2

2 l2 l2 l

l3 l3 l3 l3 l6l3 ili 18 3l 3l 3l 3l 3l 2l 2;

l 2l 2l is l2 l2 l2 i 12 l2

....i....

0 l,7 0 l,3 0 l,3 l3

!?

3 l

l3 14 -

2l 3l @3l l 3l 2,

2, 2,

2,,

12 l2

,2

,2 10 2l 3l 3l 3l l

l3 l3

!3 l2

,2 2,

2, 2,

.. i.2...,.2 Os i

Os 2l 3

3!

!3 l3

!2 8

04 l2 l @2,,

02 2l 3l3 l1 01 03 05 07 09 11 13 15 17 19 21 23 25 27 29 31 33 35 37 39 41 43 45 47 49 $1 5,

47 O LPRM LOCATION (LETTER INDICATES TIP M ACmNU c3

$ LPRM LOC ATION ICOMMON LOC ATION POR ALL T P MACHINESI 3g

@ IRM LOCATIONS 36 6 SRM LOCATIONS

% SOURCE LOCATIONS 27 i = RELOAD i 23 2

• R E LOAD-2 is

3. RFLOAD 3 lb ii 7

3 2 6 10 14 is 22 26 30 34 38 42 46 60 Figure 2-1.

Monticello Reload 3 Reference Core Loading 2-2 p

3.

KECitAfi1CA1. DE31G4 The Reload-3 fuel which will be employed has the same mechanical configura-tion and fuel bundle enrichment as the 8D250 fuel assembly described in Refer.

ence 1.

The design criteria, models, and results from design. evaluation presented in Section 3 of Reference 1 and the specific bundle-related information presented in Table 3-1.1 and Figure 3-2.1 of Reference 1 are applicable to the Reload-3 fuel.

The Reload-3 fuel incorporates finger springs for controlling nederator/

coolant bypass flow at the interface of the channel and fuel bundle lower tie plate.

This mechanism of bypass flow control has been used in General Electric's initial core and reload fuel for all BWR 4/5 plants, and for one BWR-3 plant.

The finger springs employed in the Monticello Reload-3 fuel are identical in de-sign to those that have been previously used.

The integrity of this design has been verified through the inspection of more than 900 fuel assemblies f rom operating plants employing finger springs.

Figure 3-1 is a diagram of the finger spring. Channel deflection due to creep increases the area available for leakage flow between the channel and the lower tie plate.

The finger spring is designed to expand and remain in contact with both channel and lower tie plate as inservice channel deficction occurs.

This design feature keeps the channel-to-lover tie plate flow nearly constant over a wide range of channel deflections.

Tests have been performed to determine the hydraulic characteristics of the channel-to-lower tie plate flow path.

These tests were performed at reactor operating temperature, with and without finger springs, for several channel deflections (mechanically inducad prior to test).

Thess tests have verified the ability of finger springs to restrict flow for relatively large deflections and the analytical model (non-finger spring) used to calculate the effect of channel deflections on bypass flow.

11

l f

i O

O i

'l

\\:!;ll

.' i '

t!

.i t 1 N1l4!i!:p:

bO i

1 e

V O

l l".I!

r O

g g

n n

i i

r r

p p

V S

Ol<I

' ' ii 1!

t S

r r

e e

O g

g g

n n

n i

i i

F

i F

r T

O r!1:

i,iihf.! 1l:ib!

y f

f F

o o

c w

w s

O e

e g

i i

a L

v

j v

t G i i o f'i R:: d 0 P pd ! i P

I p

t o

n a

3 T

or i

F e

O

)

m f!

, ' i

ll\\

ii

)

a l

l b

c rg a

iD O

1 3

erug iF

/

l i

f ed uo e

r FR i

e T e g

t ne

{

i N

ra ic el Fal wF o

f p I

o x1/i m

n wi e

igvn

\\p

'~

i

/

dr np 1

ES

/

)

a len r.

n e

a gm h

yN nr C

io Fs i,{

l

4.

TRERMAL-HYDRAULIC CRARACIERISTICS 4.1 METHODS, LICENSING CRITERIA. AND ANALYSIS RESUL"S The thermal hydraulic methods for calculating pressure drop and flow distri-bution and the licensing criteria employed for 8x8 fusi are discussed in Section 4 of Reference 1.

Table 4-1 of Reference i presents results of analyses performed to fully assess the effect of 8x8 fuel on core thermal-hydraulic characteristics during the transition from 7x7 to 8x8 fuel.

These results are applicabis to this reload with the exception of the core total bypass flow caployed in the analysts.

This specific consideration is discussed in Subsection 4.4.

4.2 STATISTICAL ANALYSIS ne statistical analysis of the reactor core was performed using the uncer-tainty inputs described in Table 4-1.

The results of the analyses show that at least 99.9% of the fuel rods in the core are expected to avoid boiling transition if the EPR is 1.06 or greater.

4.2.1 Fuel Cladding Integrity Safety Limit Based on the results of the statistical analysis, the Fuel Cladding Integrity Safety Limit is a HCPR of 1.06.

4.3 ANALYSIS OF ABNORMAL OPERATIONAL TRANSIENTS ne results of the abnormal operational transient analyses are summarised in Table 4-2; the transient analyses are described in detail in Section 6.

The most severe transient from rated conditions is a turbine trip without bypass which has a maximum ACPR of 0.27 for 7x7 fuel and 0.30 for 8x8 fuel.

a n

4 4-1 1

l

f t

t i

Table 4-1 DESCRIPTION OF UNCERTAINTIE,$

Standard Deviation Quantity

(% of Point)

Comment Feedwater 1.76 This is the largest component of total reactor Flow power uncertainty.

[

Feedwater 0.76 This is anothersignificant parameter in Temperature core power deter 41 nation.

Reactor Pressure 0.5 This is another significant parameter in core power determination.

Core Inlet 0.2 Affect quality and boiling length. Flow is not Temperature measured directly, but is calculated from jet Core Total 2.5 pump P.

The listed uncertainty in total core Flow flow corresponds to 11.2% standard deviation in i

each individual jet pump flow.

Channel Flow 3.0 This accounts for manufacturing and service in-Area duced variations in the free flow area within the channel.

l Friction 10.0 Accounts for uncertainty in the correlation Factor representing two-phase pressure losses.

Multiplier Channel 5.0 Represents variation in the pressure loss char-Friction acteristics of individual channels. Flow area Factor and pressure loss variations affect the core l

Multiplier flow distribution, influencing the quality and boiling length in individual channels.

TIP Readings 8.7 These sets of data are the base from which gross power distribution is determined.

The assigned uncertainties include all electrical and geo-metrical components plus a contribution from the analytical extrapolation from the chamber loca-tion to the adjacent fuel assembly segment. Also included are uncertainties contributed by the LPRM system. LPRM readings are used to correct the power distribution calculations for changes which have occurred'since the last TIP survey.

The assigned uncertainty affects power distribu-tion in the same manner as the base TIP reading uncertainty.

5 l

l 4-2 4

y

• ,,v r

v.,v-

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

.--_..-~.,-r-.%---.--,--__,,_.--.-wm-w.

,-..-...e

.--....--+,w-%.-,e-c%,,...-r-

%v_m.,ue-,_.-e,--e.-

4 R Factor 1.6 This is the last of the three power distribution related uncertainties.

It is a function of the uncertainty in local fuel rod power and is dis-cussed in detail in Reference 1.

Critical 3.6 Uncertainty in the Gr.KL correlation in terns of Power critical power.

I l

l 4-3

Table 4-2

SUMMARY

OF RESULTS SIGNIFICANT ABNORMAL OPERATIONAL TRANSIENT Lowest HCPR During Transient Description Initial MCPR The Transient Turbine trip without bypass 1.33 (7x7) 1,06 1.36 (8x8) 1.06 i

4 4-4

~ -.

i I

Addition of these ACPR's to the Safety Limit MCPR gives the minimun initial MCPR for each fuel type to avoid violating the Safety Limit MCPR during the most severe transient from rated conditions.

4.3.1 Operating Limit MCpR q

Based on the Fuel Cladding Integrity Safety Limit and the results of the ab-normal operational transient analyses, the Operating Limit MCPR is 1.33 for 7x7 fuel and 1.36 for 8x8 fuel.

4.4 CONSIDERATION OF INCREASED BYPASS FIM Ceneral Electric has maintained a continuing survei' lance program of fuel channels in operating reactors to monitor in-core perforpance. As a part of this surveillarme program, measurements have been made of channel wall deflections re-sulting from creep deformation at operating conditions.

These measurements, in combination with analytical techniques, hr o been used to formulate a model to predict the performance of channels during future operation. Results of this evaluation program indicate that channel creep rate is greater than previously predicted.

Increases in channel wall permanent deflection at the lower-tie-plate lead to inc reased bypass flow through the channel to lower-tie-plate flow path. For reactors of tne BWR-2/3 design this flow path contributes a significant portion (N80%) of the total core bypass flow. Therefore, changes in the flow through this path will affect the totc1 core bypass flow. Results of core hydraulic analyses for Monticello using the prediction model for channel wall deflection indicate that the total core bypass flow will be above the design value of 10%

of total core flow at the end of the next operating cycle.

The total bypass flow will be reduced by the use of finger springs on the Reload-3 fuel assemblies. The prediction model, with this leakage control, indicates that total core bypass flow will be <12% of the total core flow at the 4-5

4 and of the next operating cycle.

The local and total core ef fects of this increased bypass flow have, been included in both the steady-state and transient analyses. Transient effects considering the change in bypass flow are addressed in Subsection 6.3.

e e

4-6

~. _.. _.

4 5.

NUC1. EAR CHARACTERISTICS The bundle characteristics, analytical methods, and model descrintiona pre-sented in Subsections 5.1 through 5.4 of Reference 1 are applicable to this re-load. Results of specific reload core calculations are given below.

5.1 NUCLEAR CHARACTERISTICS OF THE CORE This section presents the results of the calculation on:

(1) Reactivity control characteristics (2) Core average reactivity coefficients The core characteristics were calculated using the design reference loading pattern shown in Figure 2-1.

This loading pattern was designed to accornnodate 80 Reload-3 fuel bundles by discharging 80 initial core fuel bundles.

5.1.1 Core Effective Multiplication. Control System Worth. and Reactivity Coefficients A tabulation of the typical nuclear characteristics of the Reload-3 core is given in Table 5-1.

The nuclear characteristics of the Reload-3 fuel bun-dies are similar or identical to those previously loaded. Therefore, the total control system worth and the temperature and void dependent behavler of the reconstituted core will not differ significantly from those values previously reported.

5.1.2 Reactor Shutdown Martin The Reload-3 core fully meets the established technical specification criteria in that it is maintained subcritical by at least 0.0025 ok in the most 5-1

~..

Table 5-1 NUCLEAR CHARACTERISTICS OF THE DESIGN REFERENCE CORE Core Effective Multiplication and Control System Worth (No Voids, 20*C)

BOC K

U$kbntrolled 1.109 Fully Controlled 0.947 Strongest Rod out 0.984 R. Maximum Increase in Core Reactivity 0.0017 with Exposure Into Cycle. Ak Reactivity Coefficients Range During Operating Cycle

-4 Steam Void Coefficient at 40% Voida;

- 12.7 x 10 ~

(ak/k)/4V, 1/% Void to - 11.6 x 10 Power coefficient at 1670 MWt and

- 0.056 525 Btu /lb Inlet Enthalpy to - 0.046 (ak/k)/(AP/p)

-5 Puel Temperature coef ficient at 650*C

- 1.15 x 10 (ak/k)/AT, 1/*F to - 1.22 x 10' 5-2

~

reactive condition throughout the subsequent operating cycle with the strongest control rod fully withdrawn and all other rods fully inserted.

A minimum shutdown margin of 0.014 ok is calculated throughout the subse-quent operating cycle. This minimum shutdown margin is calculated at 1000 mwd /t Cycle 4 exposure.

The maximum increase in reactivity with exposure into the cycle, R, is 0.0017 ok.

(As discussed in the Technical Specifications. 0.0004ak must be added to this value for potentially inverted poison tubes in control rods.)

5.1.3 Liquid Poison System A boron concentration of 900 ppm in the moderator water will bring the re-actor subcritical by >0.03 ok at 20'C, xenon free.

5.1.4 Reactivity of the Fuel in Storage The basic criteria for the storage of the fuel are satisfied provided the uncontrolled k, of the bundle is 11.263 at 65' C.

The reload bundle has a k, 11.263 for all exposures.

e t

5-3

4 6.

SAFETY ANALYSIS

6.1 INTRODUCTION

The safety analysis for reloads consists of three categories:

(a) generic safety analysis, which is applicable to all reloads; (b) bounding analysis; and (c) specific analysis applicable only to the current reload. Wherever a bounding analysis is applied for an accident or transient, the key parameters need only to be compared with the worst case and, if tney are within " bounds "

all limits and margins applicable to the accidents or transients will be met.

6.2 K) DEL APPLICABILITY TO 8x8 FUEL Information oa the applicability to the 8x8 design of existing models used for safety analyces is given in Reference 1.

6.3 RESULTS OF SAFETY ANALYSES 6.3.1 Core Safety Analyses The analyses given in Reference 1 on fuel damage limits, operating limits, and abnormal conditions are applicable to this reload.

6.3.2 Accident Analyses 6.3.2.1 Main Steam Line Break Accident The analysis given in Reference 1 is applicable to this reload.

6.3.2.2 Refueling Accident f

The analysis given in Reference 1 is applicable to this reload.

l 6-1 l

l

6.3.2.3 Control Rod Drop Accident The technical bases (bounding analyses) which are presented in Referenem I were used to verify that the results of a rod drop excursion in the reloaded core would not exceed the design criteria.

For application to Monticello Reload 3, the actual Doppler coef ficient, accident reactivity shape functions, and e ram reactivity functions are compared with the technical baoas in Figures 6-2 tnrough 6-4. Because the maximum values of the parameters following this reload will be well below the boundary value, the consequences of a rod drop excursion from any insequence control rod would be below the 280 cal /gm design limit.

Further, the radiological ceasegeences will be no greater than those evaluated in Reference 1.

6.3.2.4 Loss-of-Coolant Accident The analyses given in Reference 1 are applicable to this reload. These analyses were performed for the Monticello Reload-3 type 8D250 fuel utilizing both the 10CFR Part 50 Appendix K and IAC-GEGAP III models.

Figures 6-1A and 6-1B show the variation of Maximum Average Planar Linear Heat Generation Rate (FRPLHGR), peak cladding temperature (PCI), and maximum ox-idation fraction versus exposure for the Monticello 8D250 fuel. It should be noted that the calculated PCT ranains below tiie' IImits of 2300 F and 2200 F for the IAC and Appendix K respectively.

Figure 6-lb is therefore what has traditionally been called an ot MAPLHGR curve in IAC analyses; that is, there is no PCT limitation on MAPLRGR which is more restrictive than the LHGR lbnit of 13.4 kw/ft. Appendix K analyses have traditionally referred to the band of exposures where LHGR is more limiting than PCT as having "no ECCS limit." For the Reload-3 fuel the "no ECCS lbmit" band includes the entire range of exposures.

A loss-of-coolant accident (LOCA) analysis using IAC models and GEGAP III densification assumptions results in a MAPLHCR curve identical to that for the Appendix K analysis.

This is because for both analyses there is no MAPLHGR limit O

6-2

m_.__1._.__

~ --

e e

2.10o PE AK CL ADDING TEMP (R ATURE FOR APPENDIX K g ""**"" "

c

/

t"

/p'

.5

~

^

l

/%s%

W5 5

5 r 1,900 PE AK CLADDING TEMPERATURC SOR

..0 u

o APPENDIX K MAPLHGR$ AND IAC,

8 by GE GAP 111 MODE LS (OM EG A CURVll 8

IE Ir w

,o g g 5 1.soo z9 b<

f e

I I

I o

1,700 o

6.000 10.000 15.000 20,000 25.000 30,000 PLANAR AVERAGE EXPOSURE (mwd /t)

Figure 6-1a.

Peak Cladding Temperature Versus Planar Average Exposure f

13 12,5 t

R 12.20 f

m 12.1

<w ZE 11.8 5=

52

'5 E5 r

6*E

>5 0

3k IIS APPENDIX K MAPLHGR$

Z 11 hE 5

23 I

I I

I I

to O

6,000 10,000 15poo 20,000 25,000 30,000 PLANAR AVERAGE EXPOSURE (mwd /tl Figure 6-1b.

Maximum Average Planar Linear Heat Generation Rate Versus Planar Average Exposure 6-3

t f

-4A

/

SOUNDING VALUE

/

2eocd4

/

coLo siARTur ja HOTSTARTUP

/

/

/

/

_ i.o

/

,/

/

/

/

l/

P 4

/

=

/

/

/

5

/

s N ~

~

uoNTacELLO CYCLE 4 I

0

?

I

/

T = 20 C i

/

/

/

I

/

/

I

/

[

MONTICELLO CYCLE 4

' '"*c j

-,.E

/

/

I AVERAGE EXPOSURE OF

/

4 BUNDI ES AROUND DROPPE D ROD

/

I l

i I

-ta o

1000 2000 3000 sono AVER AGE FUEL TEMPER ATURE (*Cl T534 Figure 6-2.

Doppler Reactivity Coef ficient Versus Average Fuel Temperature as a Function of Moderator Condition 6-4

,._-r, e----

---y

e s

Q,016 HOT ST ARTUP. O

• 0.005, P = 1.3 g

4014 SOUNDING VALUE FOR 290 cel/g 0.012 -

0.010 b_

3 0.008 U

5 E

0.006

$+a0066 I

Pt = 1.37 gg MONTICELLO CYCLE

• WORST CASE ROO DROP gg MODERATOR TEMPER ATURL - oc-c 1

I I

I l

0 2

4 6

8 10 12 ROD POSITION (ft W1THDR AWN)

TS34 Figure 6-3.

Accident Reactivity Shape Function 6-5

-..... - - _ ~ _ _ - -.. _

6 e

-0.15

- 061 0 7

E I

I MONTICELLO CYCLE 4 SCRAM FUNCTION 0

MODERATOR TEMPERATURE

• 180 C

-0,06 BOUNDING VALUE FOR 290 cal /g I

I 1

I O

1.0 2,0 3.0 4.0 5.0 LO ELAPSED TIME AFTER SCR AM SIGNAL (seel TBS 4 Figure 6-4.

Scram Rtactivity Function 6-6 w

--n

l 4

as a result of emergency core cooling system (ECCS) considerations.

For exposures

)

greater than 20,000 mwd /t, no credit has been taken for fuel depletion, so the 1

PCTs would be lower than shown.

Because the Reload-3 8D250 fuel is not ECCS (MAPLHCR) limited, changes to Technical Specification 3.5.J will not be necessary.

6.3.2.5 Loading Error Accident 6.3.2.5.1 Event Description A loading error for the reference core configuration is defined ass (1) a reload bundle is rotated 180 degrees in a location near the center of the core or a bundle is inserted in an improper lo-cation; and (2) the. error is not discovered in the subsequent core verifica-tien ard the reactor is operated.

Since two independent errors are assumed to occur, the single error criter-ion is violated, so the event is not classified as an abnormal operating transient.

The following are the results and consequences for a worst case error.

6.3.2.5.2 Results and Conocquences Analysis of the loading error accident results in a peak linear heat gener-ation rate (LHGR) of 16.5 kW/f t and a minimum critical power ratio (MCPR) greater than 1.0 in the misplaced reload bundle. The peak LHCR is less than the damage limit es db11shed Ior the fuel'.

Since the MCPR 18 61.0 'thc Iuel ' oes ' 'o't gi)

~

~~

~

d n

~

into boiling transition and the results of this accident are far less severe than the major accidents.

Should the loading error involve one of the irradiated assemblies, previous loading error analyses results (no fuel damage) would apply.

Fuel bundles adjacent to the misplaced bundle are insignificantly affected by the presence of the misplaced bundle.

6-7

6 6.3.2.6 One Recirculation Pump Seizure Accident 6.3.2.6.1 Event Description s

This event is the result of the instantaneous stoppage of the pump motor shaft of one recirculation pump. The reactor is assumed to be operating at design power. The resulting rapid decrease in core flov is due to the large hydraulic resistance introduced by the stop.ned rei.us.

Core coolant flow and MCPR rapidly reach their minimum values before fuel surface he.it flux begins dropping and restores operating thermal margins. Nucleate boiling is maintained thus assuring that the MCPR does not decrease below the safety limit MCPR. No scram occurs and the initial pressure regulator maintains pressure control as the re-actor settles out at the final, lower power condition.

The primary sys tem pressure decreases throughout the event thus the nuclear system process barrier is not threatened by overpressure.

6.3.2 6.2 Results and Consequences This transient describes the instantaneous stoppage of one pump motor shaft resulting in the most rapid decrease of core flow. The plant is assumed to be operating at the specified conditions (Table 6-1) with the exception of the void coefficient which was adjusted to -6.75c/~"g to introduce conservatism for a de-creasing flow transient. A sharp decrease in drive flow and diffuser flow in loop 1 is caused by the seized rotor. Core flow reaches a miniums value at about 1.4 seconds. The ACPR is only 0.06 and, therefore,10 not the limiting transient and the safety limit is not violated.

6.3.3 Abnormal Operatina Transients 6.3.3.1 Inputs and Results 6.3.3.1.1 Analysis Basis _

This subsection contains the analyses of the most limiting abnormal opera-tional transients for Monticello Cycle 4.

All transients which are the basis of l

6-8 I

l

--n

:1 r

cm

__m._

4 the existing license were reviewed, and those transients which have been limiting in the past with respect to safety margins and are significantly sensitive to the core transient parameter deviations were reanalyzed.

The input patameters and functions which characterine this cycle are ana-lyzed tmder end-of-cycle (EOC) conditions because the response to the limiting transients is most severe at this condition.

6.3.3.1.2 Input' Data and Operating conditions The input data and operating conditions are shown in Table 6-1 and represent the nominal basis for these analyses. Each transient is considered at these con-ditions unless otherwise specified.

6.3.3.1.3 Transient Summary A summary of the transients analyzed and their consequences is provided in Table 6-2.

6.3.3.2 Transient Descriptions The abnormal operating transients which are limiting according to safety criteria and which also are sensitive to nuclear core parameter changes have been analyzed and are evaluated in the following narrative.

6.3.3.2.1 Turbine Trip With Bypass Failure - Trip Scram This transient produces the most severe reactor isolation. The primary characteristic cf this transient is a pressure increase due to the obstruction of steam flow by the turbine stop valves. The pressure increase causes a signi-ficant void reduction which yields a pronounced positive void reactivity. The net reactivity is sharply positive and causes a rapid increase in neutron flux until the net reactivity is forced negative by scram initiated f rom 90% open 6-9

4 I

Table 6-1 TRANSIENT INPUT PARAMETERS Thermal Pcuer (MWt) 1670 100%

6 Rated Steam Flov (1b/hr) 6.78 x 10 100%

6 Rated Core Flow (1b/hr) 57.6 x 10 100%

Dome Pressure psig 1025 Turbine Pressure psig 980 RV Satpoint psig 1080 + 1%

RV/ Capacity (at Setpoint) f/%

6/68.8 RV Time Delay (msec) 400 RV Stroke Time (msec) 100 Void Coefficient (c/% Rg)

-9.37*

Void Fraction

(%)

36.6 Doppler Coefficient (c/%T )

-2.50

• p Avg. Fuel Temperature

('F) 1284 Scram Reactivity Cete EOC 4 *

(Figure 6-5)

Scram Worth

($)

-40.63

• The values given for void and Doppler coefficients are the inputs used in the transient analysis. They consist of the actual values times conservative multipliers of 1.25 and 0.90, respectively.

The EOC 4 scra reactivity curve shown in Figure 6-5 is the actual value. The transient analysis was done with this conservatively reduced by a multiplier of 0.80.

G 6-10

=

l 1.1 44 1.0 40 u

0.8 30 0.7 5

{ 0.6 5

E t

5:

s F,

5u n

3<

0.4 0.3 10 0.2 0.1 I

I O

0 0

1 2

3 4

TIME AFTER SCRAM TRIP (sed TG1 Figure 6-5.

Scram Reactivity - Monticello EOC 4 6-11 a

Table 6-2

)ONTICELLO CYCLE & TRANSIENT DATA

SUMMARY

Core Power Flow Q

Q/A Psi Py Transient it)

_(%)

g3 (peig)_ (psig)_ ACPR Turbine Trip Without Bypass-100 100 430 116.7 1200 12.*4 0.30 Trip Scram Inadvertent Start-HPCI Pump 100 100 116 113.4

<0.18 l

i

(

i l

l 6-12 i

1 l

switches on the turbine stop valves and by a void incraase after the safety /re-lief valves have automatically opened on high pressure. Figure 6-6 illustratas this transient.

The parameters of concern are the peak vessel pressure margin to the vessel code limit and the peak average surface heat flux. The peak steamline pressure margin to the spring safety valve setpoint is no longer a concern with the re-moval of the spring safety valves.

Because the vessel pressure margin to the vessel code limit is usually substantial it is therefore not the limiting para-meter acct.uing to safety criteria.

Consequently, the peak heat flux correlated to HCPR is the limiting parameter.

Neutron flux, the precursor of heat flux, rises to a peak of 430% with a corresponding peak heat flux of 117%. The resulting 6CPR f rom this transient is 0.30 which determines a design basis operating critical power ratio (CPR) of 1.36.

The maximum power level which may be realized is influenced by core man-agement policy and, therefore, no maximum power limit is specified.

The operating MCPR of 1.36 for Monticello should impose no operating power level restriction below 100%.

The peak steamline pressure is limited to 1200 psig as a result of the high-pressure actuation of the six safety / relief valves. The corresponding peak vessel bottom pressure is 1224 psig which provides a 151 psi margin to the vessel code limit.

6.3.3.2.2 Loss of a reedwater Heater The loss of a feedwater heater is normally analyzed in FSAR's and other sub-mittals because it constitutes the most limiting cool water injection transient.

However, the Monticello ple.nt design ic such that a feedwater heater cannot be inadvertently removed from service because bypass capability does not exist.

For comparison only, an analysis for this transient was made, the conse-quences of which are mild, showing a ACPR of 0.18.

6-13

---.--_m.

4 h

T' f

.(e(a B

2 a

/

6 a

a Iq q

[!!f

[

t.tv l

b!!

[/ f 3!!

/

,s 3,* y 9,

y lE a

ljf l\\

l\\

i.E_

i e

y x

'l N

x y

l yi j'

i

~

N._ _.. l:

i e

a e

=

4 2

=

sunca h

e a

( /"

I

^

E!4 l

fMJ -(

u yg

@,j 1..Am5 4p 5

i y

a I,!gk I

!$iE

-\\ y )

,_a n

~

n..

)!//

g 1

. /,) i

,i

\\

[

j I

J L Jl T,

fw

\\xl.

/

. _..m M!,

9

..[dda l

a a

d e

a 8

IC31th n IN3JtGd) 6-14

l The following transient, Inadvertent Start of an HPCI Pump is the most limit-ing transient of this type.

6.3.3.2.3 Inadvertent Start of an HPCI Pump This transient considers the ef fects of the reactor receiving cooler water f rom the feedwater spargers.

The reactor is assumed to be operating at maximum power condition and in manual flew control which is more severe than the auto-matic flow control mode.

Figure 6-7 shows the response to this transient.

The cooler water received by the reactor increases the core inlet subcooling which forces'a reduction in void fraction and results in an increase in power level.

The heat flux increases to only 113%. The resultant ACPR will be less than 0.18.

6.3.3.3 Rod Withdrawal Error Assumptions and descriptions of rod withdrawal error are given in Refer-ence 1.

Figures 6-8 through 6-10 show the results of the worst case condition for Monticello Reload-3. The rod block monitor (RBM) setpoint is selected to allow for failed instruments for the worst situation.

This case demonstrates' that even if the operator ignores all alarms during the course of this transient, the RBM will stop rod withdrawal while critical power ratio (CPR) is still greater than the 1.06 MCPR safety limit.

6.3.4 ASME Vessel Pressure Code Compliance All Main Steamline Isolation Valve Closure-Flux Scram The pressure relief system must prevent excessive overpressurization of the primary system process barrier and the pressure vessel to preclude an uncontrolled release of fission products.

6-15

I i

e

'l l

W' I

q i

f8415

- c,

_a g

7 y

g c

M s

J; 4

W

=l

~

,,b N

H a

1

.o

!aa1 1 a 1 th a

a 1 a1

(

8 I

d d

Y h

"0 E

w roo B

W g

I f

w I

a,

==

=n I

N W

gg g

l W

4 g

g c

(

c

-d J

(

N j

A

~

t I

5.k i., e/

,2 l

....i.

i e

v

/

s t

wum.c vaxw 6-16 i

_ _ -. ~... _

-.- _. - ~._. _

1A L

1A 1.4 1.3 9

4e e

1 i

N 5

g Fuel Assembly

{

Location g

(23,30)

I 1.1 i

(19,34) g gl e

03 a

O l

l l

l l

0 8

18 24 32 40 40 IN NOTCH POSITIONS REMOVED OUT Figure 6-8.

MCPR Versus Rod Position I

(

6-17

^

134 CHANNEL A+C MONTICELLO - CYCLE 4 4

130 H

129 12 2 20.29 A

H 120 C

118 E

o D

o Failed LPRM Strings 110 A

NONE F AILED e

ROD BLOCK LINE 8

' 'M f

C MM 106 D

20J7 E

20J7 AND 20,29 I

i I

i I

I I

I I

1 I

I I

O 1

2 3

4 5

6 7

8 9

to 11 12 13 14 15 CONTROL ROOS Ift WITNDR AWN)

Figure 6-9 RBM Response to Control Rod Motion for Rod Withdrawal Error - Limiting Case Channel A + C for Reload 3 3

136 W

CHANNELS 4D lAONTICELLO - CYCLE 4 4

tu?

20J's

+

i Failed LPRM Strings A

NONE F AILED g3 e

E E$

E1 D

20J7 E

20,37 AND 20,29 120 E

v o

115

=

C D

110 ROD SLOCK LINE 106 I

t I

I t

i I

I t

I t

i 100 0

1 2

3 4

5 6

7 8

9 to 11 12 13 14 18 CONTMOL ROO4191 W1THORAWN)

Figure 6-10.

RBM Response to Control Rod Motion for Rod Withdrawal Error - Limiting Case, Channel B + D for Reload 3 6-18

The Monticello pressure relief system includes only dual function relief valves located on the main steam lines within the dry well between the reactor veosel and the first isolation valve.

These valves provide the capacity to limit nuclear system overpressurization.

The ASME Boiler and Pressure Vessel Code requires that each vessel designed to meet Section III be protected from the consequences of pressure in excess of the vessel design pressures (a) A peak allowable pressure of 110% of the vessel design pressure is allowed (1375 peig for a vessel with a design pressure of 1250 psig).

(b) The lowest qualified safety valve setpoint must be at or below vessel design pressure.

(c) The highest safety valve setpoint must not be greater than 105% of vessel design pressure (1313 psig for a 1250 psig vessel).

Monticello's safety / relief valves are all set at a nominal pressure of 5 1080 psig which satisfies requirements (b) and (c) above. Requirement (a) is evaluated by considering the most severe isolation event with indirect scram.

The event which satisfies this specification is the closure of all main steamline isolation valves with indirect flux scram. The initial conditions assumed are those specified in Table 6-1.

Figure 6-11 graphically illustrates this event.

An abrupt pressure and power rise occur as soon as the reactor is isolated.

Neutron flux reaches scram level in about 1.75 seconds, initiating reactor shut-down. The safety / relief valves open to limit the pressure rise in the steamline at the valves to 1233 psig and at the bottom of the vessel to 1269 psig.

This response provides a 106 psi margin to the vessel code limit of 1375 psig.

Thus, requirement (a) is satisfied and adequate overpressure protection is provided by

~

the pressure relief system.

6-19

,,-v u JY5

/

-Q

^

E I

Ba t

Il l

Mli

~

-~..

b a

N l

1. i I

e

)1 y i

% ('.

f

't

/;

s

. r

....i....

curoo c

h E

i I

h

\\

n.

s h(.-f" II t

z e

b d

\\t c

L,.5

/

s pg

=

t ib2 f

l

~

J t

f E

j

/ '!

t

/

S.

L h

\\

h 1

d o

r

_;a q

rg

)

i i

.... i....:

i

.m i....:j j

wmm o imm 6-20

l 6.3.5 Thermal - Hydraulic Stability Analysis 6.3.5.1 Channel Hydrodynamic Conformance to the Ultimate Performance Criteria The channel performance calculation for Honticello.Teload 3 yields decay ratios as presented below:

Extrapolated Rt.1 Block -

Channel Hydrodynamic Performance Natural Circulas, ion Power Decay Ratio, X !A 2 0 8x8 Channel 0.40 7x7 Channel 0. 2 '-

At this most responsive condition, the most responsive channele are clearly within the bounds of the ultimate performance criteria of <1.0 decay ratio at all attainable operating conditions.

6.3.5.2 Reactor Conformance to Ultimate Performance Criteria The decay ratios determined f rom the limiting reactor core stability condi-tions are presented in Figure 6-12.

The most responsive case is again the extra-polated rod block - natural circulation condition.

Extrapolated Rod Block -

Reactor Core Stability Natural Circulation Power Decay Ratio, X /X 0.47 2 O

(

These calculations show the reactor to be in compliance with the ultimate performance criteria, including the most responsive condition at extrapolated rod block - natural circulation power, i

e e4 l

6-21

m. __ _.

e I

Axl AL POWER DISTRIBUTION 2.0 I

y 1.0 G

5 E

0.0 INLET ExlT AxlAL OtST ANCE 1.0 ULTIMATE STABILITY LIMIT 02 g

0.s k

a y

NATUR AL CIRCULATION 6

E 0.4 DESIGN GUIDE 0.2 R ATED POWER-FLOW LINE I

I I

I 0.0 0.0 0.2 0.4 0.6 0A 1A FR ACTION OF R ATFD POWEH T63 2 Figure 6-12.

Core Reactivity Stability l

4 l

1 6-22 I

I

+m.--

j I

i 6.3.5.3 Channel Hydrodynamic Conformanco _to the Operational Desian Cuide f

4 i

Channel Hydtr. dynamic Rated low and of Flow Control Forformance conditions Ranae (591 power /405 flow)

Decay Ratio, K /I 2 O i

8x8

<0.01 0.07 7x7

<0.01 0.05 The most responsive channel la well in conformance with the operational de-sign guide of 3.0.5 decay ratio.

j 6.3.5.4 Reactor Core Conformance to Operational Design Guide The calculated value of the decay ratio of the reactor power dynamic re-l sponse for rated operating conditions and for the 17w end of the normal flow f

control renge (591 power, 40% flow) are presented below.

Low End of Normal Reactor Core Rated Flow Control Range Performance Conditions (591 Power /401 Flow) 0.25 Decay Ratio 1,0.01 i

Aa noted earlier, Figure 6-12 describes the variation of decay ratio over the entire power flow range.

i t

O f

6-23 1

4

4 I

l i

7.

TECHNICAL SPECITICATIONS There are no Technical Specification changes necessary for operation in Cycle 4 with the reload core as analysed in this submittal.

The analysis of abnormal operational transients was done using the CETAB/CEXL heat flux correlation discussed in Ref erence 1.

This analysis serves as the basis for a new way of expressing safety limits which will be incorporated into the Technical Specifications at a later date. Operation in Cycle 4 under the existing Technical Specifications is discussed in Appendix A.

At the time the ICCS evaluation was performed, it was required by 10 CFR, l

Part 50, Appendix K that compliance to both the Interim Acceptance Criteria and the Final Acceptance Criteria be demonstrated. This has been done for Reload 3 fuel in the same manner that it has previously been done for other fuel types in the reactor. Results of ooth the IAC and FAC calculations for Reload 3 fuel are reported in Section 6.3.2.4 showing tnat no limitations are imposed by ECCS criteria which are more restrictive than the existing Technical Specar cation linear heat generation rate limitation of 13.4 kw/f t for 8x8 fuel.

L O

7 1

4 8.

REFEllNCEE 1.

2/BlGt Generio Reload Licensing Apptioation for Sz8 Fuel, Novensber 1914 (NEto-20360. Rev. 1).

2.

Letter f rcan L 0 Mayer (NSP) to J F O' Leary (USAEC',, April 10, 1974 i

4 fj f

a s

e 4

t I

t 8-1

APPDiDlX A EVAL 11ATiiti 0F AB!iolMAL OPERATIONAL TRA!4SIDit CA14LTIM104S APPLICABLE TO CYCLE 4 OPERATI(ti Section 6.3.3.2.1 reports the results of abnormal operational transient calculations based on the CETAB concept.

I:xttnsive relevant analyses have also been done for previous cycles using the llench Levy correlation which serves as the basis for existira Technical Specifications. An evaluation of the calculations shows that Cycle 4 is bounded by those analyses, l

1he only significant differences between Cycle 3 and Cycle 4 parameters are the scram reactivity function and the void coefficient.

The end of Cycle 4 scram reactivity insertion rate exceeds that of Cycle 3 tending to make Cycle 4 a conservative calculation.

The Cycle 4 calculation was based on a void coef ficient that was more negative than Cycle 3.1his change in void coef ficient is not directly related to the reload fuel.

The severity of a transient can be seen in the dynamic pressure response and the average fuel surface heat flux response. As reported in Reference 2, a turbine trip from rated power with failure to bypass at the end of Cycle 3 was calculated to result in a peak vessel pressure of 1237 psig and a peak average surface heat flux of 120%. Section 6.3.3.2.1 reports the same calculation for Cycle 4 to result in a peak vessel pressure of 1224 psig and a peak average surface heat flux of 1171.

Frcnn this one can conclude that the improvement in the Cycle 4 scram reactivity curve over the Cycle 3 curve more than compensates for the change in void coefficient and that Cycle 4 is therefore bounded by the Cycle 3 llench Levy results even at the higher level of conservattem in the void coef ficient.

A-1

Reference 2 states that even though MCitPR limits are met, power would be reduced at the end of Cycle 3 unless additional analysis was done. As stated, this condition axisted because of concern that McPR may have been limiting if MCPR calculations were perfonced.

Since Cycle 3 calculations bound Cycle 4 conditions, the MCHPR in Cycle 4 is more conservative than the acceptable value calculated for Cycle 3.

Also, because Cycle 4 analysis of MCPR has shown acceptable results (see Section 6.3.3.2.1), there is no need to consider an end of cycle power restriction in Cycle 4 as was discussed for Cycle 3.

A aumerical comparison of Cycle 3 and Cycle 4 exists in the GETAB calculations.

Analysis of the turbine trip with failure to bypass, the most severe of the abnormal operational transients was done for Cycle 3 from a starting MCPR of 1.201 the dCPR was found to be 0.27.

Cycle 4 operation is based on a starting MCPR of 1.36, with the transient dCPR allowance of 0.30 to a safety limit of 1.06.

We relationship between the starting MCPR and the transient dCPR is such that had the Cycle 3 calculation been done from a starting MCPR of 1.36 rather than 1.20, the dCPR would have been about 0.37 rather than 0.27.

Wese results support I

the above conclusion that the most limiting abnormal operational transient in Cycle 4 (dCPR of 0.30) is less severe that in cycle 3 (6CPR of 0.37).

Cycle 4 operation may therefore proceed under the existing Technical Specifications with full assurance of conservative adherence to all limits.

uis conclusion can i

be drawn from extensive analysis using both the Hench-Levy and the GETAB/GEXL haan fit :orrelations.

O g

l A-2 j

.. -